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+[
+    {
+        "image_filename": "designv11_60_0002923_978-94-011-3514-6-Figure4.6-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0002923_978-94-011-3514-6-Figure4.6-1.png",
+        "caption": "Fig. 4.6 Locus of H as a function of 0 (~ 0.8)",
+        "texts": [
+            "39) - 66 - For the point HD (corresponding to the phase, velocity and power resonances ~D = 1 ), the real part is zero when the imaginary part is ao = ;~ (= - ~ D) For the point H2 (amplitude resonance ~2 = ~ ), the modulus is equal to the maximum (4.12) and one obtains 2 (1 - 0 2 ) -1 11 - 20 2' 20(1 - 02 ) When the damping factor tends to zero, a2 tends towards 1/2. Thus, the locus of the points H2 , shown in figure 4.7, has a vertical asymptote. Finally, one can prove that the imaginary part b1 of H1 is very close to a maximum. Let us now choose the damping factor as a parameter, the relative angular frequency being assumed to be constant. One can show that the point H describes a semi -circle in the complex plane (figure 4.6), corresponding to the equation (4.40) (4.41) - 67 - When the damping factor increases, the semi-circle is traced out from the point Q (~ = 0) to the point 0 (~ =~) . For ~ = 0 , the diameter is equal to unity and the interior of the half-circle is inaccessible. ~u Curves of the complex frequency response H with the damping factor ~ and the angular frequency a as parameters. The point .!!Z corresponds to the maximum of the modulus of .!!. The regions of the complex plane which are inaccessible to "
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0001051_phm.2016.7819861-Figure7-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0001051_phm.2016.7819861-Figure7-1.png",
+        "caption": "Figure 7. Flexibility of the sun gear: (a) MNF file, and (b) the flexible body in ADAMS.",
+        "texts": [
+            " The procedure of generating the MNF file is briefly described as follows: 1) Define the material properties and element types; 2) Generate mesh; 3) Create rigid regions and interface points which turn to be joint points between the rigid body and the flexible body in ADAMS; 4) Export to ADAMS. Generally, the more models the MNF file includes the more accurate the computation will be, however, more time will cost. Thus, it is necessary to select appropriate models. Through repeated attempts, the first 20 order models are chosen to be exported which can meet the requirement of precision. Fig. 7(a) shows the generated MNF file and Fig. 7(b) shows the flexible body of the sun gear. III. DYNAMIC SIMULATION OF PLANETARY GEAR SYSTEM WITH GEAR TOOTH CRACK Based on the results in Section II, a rigid-flex coupling model of the planetary gearbox is then established and its dynamic responses for different cracks are analyzed in this section. A rigid-flex coupling model which combines the advantages of the analytical model and the finite element model has been widely used to investigate the dynamic characteristics of planetary gear system"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0000758_978-981-10-0471-1_34-Figure6-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0000758_978-981-10-0471-1_34-Figure6-1.png",
+        "caption": "Fig. 6 (a) Double-sided tape and (b) fabricated magnetic mounting experiment setup for an attachment comparison",
+        "texts": [
+            " 5 Apparatus of the test rig Both methods were applied to attach an AE sensor to the stepper motor of 3D printer and then signals were acquired from normally working conditioned 3D printer. After that, these signals would be compared by mainly focusing on the amplitude and details of the signal. Effectiveness of both attachment methodologies would be compared in order to discover more suitable methodologies for the main experiment to assure well signal transmission for further analysis. Experiment setup for both attachment methodologies is shown in Fig. 6, (a) for double-sided tape and (b) for fabricated magnetic mounting. The main experiment of this study was the study of AE signals under different conditions of 3D printer. The experiment consisted of three conditions which were filament normally feeding condition, filament stop feeding through the extruder condition, and no filament supply for the extruder condition. For the first condition, an AE sensor was attached to a properly working 3D printer to acquire signals as a reference signal. For the second condition which is filament stop feeding condition, in order to simulate filament stop feeding through the extruder, a C-clamp was attached to the filament, which was feeding to the extruder, in order to block the filament from feeding into the extruder"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0002119_12.148097-Figure8-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0002119_12.148097-Figure8-1.png",
+        "caption": "Fig. 8 Experimental rate loop response.",
+        "texts": [
+            " However, it can take on any desired value between 0 and 1 , depending on the resonance Qencountered. Because the mechanical break w1 in Eq. (17) is also eliminated when y = 1, or close to it, the closed-loop bandwidth in this simple system is primarily controlled by GA (amplifier and compensation dynamics), as shown in Eq. (20). 4.1 Application Test results for one of several applications using state equalization to eliminate resonance constraints in motion stabilization loops, a unit in a shipboard pointing system, are presented in Fig. 8 to illustrate the effects of applying this technique. The plant consists ofa geared dc motor-tachometer and amplifier. Tachometer loop bandwidth was restricted to about one octave below the antiresonance frequency to maintam adequate stability margin. Without state equalization, the frequency response ob- '18\" tamed at nominal signal amplitude is shown in Fig. 8. Antiresonance was about 30 Hz with a Qof about 4, and phase loss at 10 Hz (budgeted to 36 deg maximum) was about 78 deg. When the state equalization technique was applied, as shown in Fig. 7 (with 'y equal to about 0.9), the severe anti(19) resonance effect virtually disappeared, as shown in Fig. 8. OPTICAL ENGINEERING / November 1993 / Vol. 32 No. 11 / 2809 (kJL)l/2 QAR= b (15) Comparing Eqs. (15) and (13) then, the damping ratio is - AR'/2QAR . (16) the closed rate loop transfer function becomes \u00b0M KAKMGA E(s/w)F[l ('%/KyH)(KAGA/R)I +(1 + KAKMKRGA) (17) and y=1 Downloaded From: http://opticalengineering.spiedigitallibrary.org/ on 05/19/2015 Terms of Use: http://spiedl.org/terms BIGLEY 2810/OPTICAL ENGINEERING / November 1993 / Vol. 32 No. 11 Downloaded From: http://opticalengineering.spiedigitallibrary"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0002589_s003390051093-Figure1-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0002589_s003390051093-Figure1-1.png",
+        "caption": "Fig. 1. Schematic of the cryogenic set-up of the LT STM with the STM body hanging in its concentric copper housings at the bottom of the cryostat: (1) LHe dewar, (2) LN2 dewar, (3) concentric copper cylinders with doors and windows, (4) STM block, (5) three springs, and (6) tubes for springs. Not shown are the door mechanisms and the stainless steel wires to firmly pull up the STM body and press it against the bottom of the LHe dewar for cooling down or tip and sample exchange",
+        "texts": [
+            " The apparatus containing the new LT STM consists of a custom-designed cryostat chamber and a preparation chamber connected by an Omniax x-y-z sample translator [9] with fast-load lock for sample introduction. A custom-built LHe flow cryostat provides cooling capabilities to the sample translator with full 360\u25e6 of rotation. The two main chambers are pumped by ion-getter and Ti-sublimation pumps to base pressures in the 10\u221211 mbar range. The cryostat chamber contains the liquid-helium (LHe) bath cryostat surrounded by a liquid-nitrogen (LN2) dewar (see Fig. 1). In the preparation chamber, tips and samples can be heated and sputter-cleaned as well as be exposed to various evaporation sources and gases. They can then be put inside a storage carousel in the cryostat chamber, which is cooled to LN2 temperatures. On the sample translator tip and sample can be cooled to 40 K prior to their insertion into the STM head. This novel concept is the key to the very short turn-around times of less than 1 hour, before thermal equilibrium is reached and drift effects in the STM get below 0.2 nm/hour. A schematic of the cryogenic section is shown in Fig. 1 with the LHe dewar (1) and a cross-section of the LN2 dewar (2) surrounding the helium dewar. Both cryostats are mounted on CF flanges and get filled from above. They are independently elongated at their respective bottoms by concentric copper cylinders (3) with door mechanisms (not shown in Fig. 1) for tip/sample exchange and infrared-filter windows for optical access. The STM block (4) is hanging down from 3 springs (5), each of which is running within a tube (6) through the LHe dewar and is mounted to the top of the LHe cryostat. During cool-down a pull-up mechanism employing 3 stainless-steel wires running through the suspension springs (not shown in Fig. 1) allows to firmly press the STM body against the bottom of the LHe dewar. In this position tip and sample can be exchanged by a wobble-stick which reaches into the copper containers, after their doors have been opened. When the minimum temperature has been reached, the pullup mechanism is relaxed and the whole STM body lowered by about 5 mm to hang freely from the springs. The bottom plate on which the whole STM body is mounted has copper wings which radially stick into a concentric periodic magnet structure for eddy-current damping. The schematics of the STM itself are also shown in Fig. 1. Its sample stage is made form a copper block and designed to ensure a good thermal coupling of the whole stage and the sample. Since the whole STM is placed within the copper housing which is at 4.2 K, a uniform temperature distribution is achieved resulting in a very low STM drift rate, measured to be less than 0.2 nm/hour (see Applications below). The scanner is mounted on the 3-axis piezo coarse motor (range: 5 mm\u00d75 mm\u00d710 mm) by employing OMICRON\u2019s micro piezo slide principle [10]. The scan range of the tube scanner [11] is temperature dependent with values of 10 \u00b5m, 4"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0003013_a:1008115522778-Figure18-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0003013_a:1008115522778-Figure18-1.png",
+        "caption": "Figure 18. Experimental results for misalignment compensation: initial error = 7 mm.",
+        "texts": [
+            " Misalignment compensation was accomplished at the beginning of the second corrective motion, because the difference between the inferred motion and the actual motion until misalignment compensation is actually accomplished is not large in the first corrective motion m1. However, there is quite large difference in the second corrective motion m2. This is due to the same reason as em = 1 mm in Figure 16 because the bottom of the peg was moved near to the hole by m1. These successful compensation does not guarantee the success in more general tasks such as high speed or high precision assembly. However, the success rate of misalignment compensation in such tasks will be increased by considering more parameters in a neural net-based inference system. Figure 18 shows the experimental results when em = 7 mm and cr = 4, 7, 10 mm. Both the measurement error by the sensing system and the error in the inferred corrective motion are not large. When cr = 7, 10 mm, two times of corrective motion was required to compensate for the lateral misalignment. Misalignment was compensated at the beginning of the second corrective motion. These results are similar to those in Figure 17. Figure 19 shows the experimental results when em = 7 mm, cr = 7 mm and \u03c6m = 45\u25e6. The \u03c6h denotes the direction of lateral misalignment"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0002923_978-94-011-3514-6-Figure5.3-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0002923_978-94-011-3514-6-Figure5.3-1.png",
+        "caption": "Fig. 5.3 Steady state beats (a) X' 0.7 X\" 0.3 (b) X' = X\" = 0.5",
+        "texts": [
+            "5 0.2 (8.37) -155 - One part of the energy of the system is exchanged alternatively between the variables X1 and X2 , whilst. each mode conserves its own energy. For a symmetrical system, for example those of figures 8.2 to 8.4, the coefficients ~ 1 and ~2 take the vales and-1 respectively. If one chose the initial conditions such that X1 = X2 ,the displacements X1 and X2 would pass through zero values, at zero speed, for each half period 1/2 = n/cr . (This is analogous to the situation shown in figure 5.3 case (b), page 90.) - 156 - CHAPTER 9 THE FRAHM DAMPER A Frahm damper is a device which attenuates the vibrations of a mechanical system over a specified range of frequencies. It consists of an oscillating system, known as the auxiliary system, whether dissipative or not, which is attached to the main system, augmenting in this way the number of degrees of freedom and therefore the number of resonances of the complete system. The att.enuation of the vibrations of the main system is achieved by the transfer of these to the auxiliary system at the desired frequencies"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0000878_978-3-319-47614-8-Figure6.14-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0000878_978-3-319-47614-8-Figure6.14-1.png",
+        "caption": "Fig. 6.14 Boundary conditions",
+        "texts": [
+            "6 MPa, inner diameter is di = 5170 mm, cylinder length of pressure vessel is L = 8000 mm and plate thickness is t = 30 mm. Mean Radius = di 2 \u00fe t 2 \u00bc 2:6 m rl \u00bc pr 2t \u00bc 69:33 MPa rh \u00bc pr t \u00bc 138:667 MPa reqv \u00bc ffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi rl\u00f0 \u00de2 \u00fe rh\u00f0 \u00de2 rl rl q \u00bc 120:08 MPa The geometric model with quarter-symmetry is shown in Fig. 6.12. The FEA model is given in Fig. 6.13. The boundary conditions applied are given in Fig. 6.14. The Equivalent (Von Mises) Stress obtained is given in Fig. 6.15. The result obtained in Worked Example 2.3 is 120.08 MPa and it agrees well with the Finite element result 122.94 MPa above. This is an example where the result could be obtained by a simple equilibrium model approach for the required section; however, in complex geometry systems, the commercial solvers with large elements help the designer. 6.8 General Structures by Commercial Solvers 191 192 6 Bending of a Beam 6.8 General Structures by Commercial Solvers 193 6"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0003319_s0003-2670(00)01261-7-Figure3-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0003319_s0003-2670(00)01261-7-Figure3-1.png",
+        "caption": "Fig. 3. Average peak height at 605 nm vs. KOH concentration for low-end analysis using 10 ml injections of 10 mM BCG. The two sets of data were taken 2.5 months apart. The error bars connote \u00b11 S.D. (n = 6). The solid line constitutes a best fit to the model represented by Eqs. (3)\u2013(8). Inset: typical system performance for 605 nm wavelength using 10 ml injections of 10 mM BCG indicator with 6 min between injections. KOH concentration (in ppm) are A = 1.5, B = 1.9, C = 2.9, D = 3.8, E = 13.4, F = 15.9, G = 19.0, H = 21.2.",
+        "texts": [
+            " About 10 ml of 2 mM solutions of each of the three indicators were initially tried and did not produce satisfactory results. Using a 10 mM solution of BPB (10 ml injected at 8.33 ml/s), the most easily ionized of the three, as the indicator, very low levels of KOH could be easily detected but indicator saturation occurred by the 10 ppm KOH level. With BTB, the least acidic of the three, KOH concentrations below 15 ppm could not be measured. BCG allowed for a usable measurement range of 1.5\u201320 ppm KOH. The inset in Fig. 3 shows typical system performance at 605 nm for 1.5\u201320 ppm KOH. Fig. 3 itself shows the data from two disparate runs (with \u00b11 S.D. error bars on each measurement, n = 6) taken 2.5 months apart. For a theoretical prediction of the above response, we invoke here the Franklin\u2013Marshall solvent system theory of acids and bases and assume that both the proton and hydroxide are solvated by polyol (POH) to produce the characteristic cation (POH2 +) and anion (PO\u2212) of the solvent. The charge balance equation for a system containing the acid form of the indicator HIn, KOH and the polyol can be written as [POH2 +] + [K+] = [PO\u2212] + [In\u2212] (3) where it is understood that K+ and In\u2212 are likely also solvated by polyol",
+            " The proton transfer reaction with the indicator itself can be written as HIn + PO\u2212 = In\u2212 + POH (4) We define an equilibrium constant Kp [In\u2212] [HIn][PO\u2212] = Kp (5) where Kp will be the analog of KIn/Kw in water. Recognizing that the total indicator concentration CIn is given by CIn = [In\u2212] + [HIn] (6) We obtain [PO\u2212] = [In\u2212] Kp(CIn \u2212 [In\u2212]) (7) Putting Eq. (7) into Eq. (3) (with [POH2 +] neglected) results in the equality [K+] \u2212 [In\u2212] [ 1 + 1 Kp(CIn[In\u2212]) ] = 0 (8) Based on the best fit of all the data in Fig. 3 and invoking a least squares minimization routine, we compute the best fit value of Kp for BCG in polyol to be \u223c650. In water, this would correspond to an effective pKIn of 11.2, suggesting that the ionization of BCG is depressed by \u223c7 orders of magnitude in polyol. The corresponding predicted absorbance response is plotted in Fig. 3 in the form of the solid curve. An increase in the injected indicator volume was attempted to cover the higher range. A linear range of 15\u201385 ppm could be obtained with an injection of 30 ml 10 mM BCG (injected at a rate of 10.4 ml/s). Since it was desirable to extend the upper linear range to higher values, two alternatives were investigated. The first involved the injection of an indicator solution both greater in volume and concentration than those used in previous trials and the second involved the addition of a mineral acid to the indicator solution"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0002633_00423119808969458-FigureI-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0002633_00423119808969458-FigureI-1.png",
+        "caption": "Fig. I . Wheelset Coordinate Systems.",
+        "texts": [],
+        "surrounding_texts": [
+            "RAO V. DUKKIPATl\n2. CREEPAGE AND SPIN QUANTITIES\nAccording to Kalker's empirical theory [3,4], the creep force F and moment M i n terms of creepage and spin quantities v,, v, and 4, are given by:\nF/pN = f , ( ~ ) e , + f2 (7 )e2 , when T i 1,\n= e2 when T 1 1 ;\nand\nwhere\nG is the shear modulus of the materials of the rollers and wheels; a and b are the lengths of semi-axes of the contact ellipse in the directions of x and y, x being the rolling direction; Cij are the creepage and spin coefficients and depend on the ratio a/b and Poisson's ratio u; N is the normal load and is the coefficient of friction. (Creepage and spin are defined as the tangential components of the translation velocities and normal component of the angular velocity of the wheelset relative to the roller at the center point of the contact area divided by the rolling velocity V.)\nTo determine the creep forces and moments we therefore need, besides p, N and G, the quantity c. the. creepage coefficients C,, and the creepage and spin quantities v,, v, and +. The determination of c and the creepage coefficients Cij for the CTS have been already documented in [5,6]. Thus there remains the problem of determining the creepage and spin quantities.\nTo accomplish this we shall introduce a convenient set of coordinate systems for describing the surfaces of rollers and wheels and determine the conditions that the coordinates must satisfy at the point of contact. The analysis follows .closely the procedure adopted 'by de Pater [7]. However in our analysis, we shall include the initial canting and the yaw motion of the rollers.\n3. GEOMETRICAL CONSIDERATIONS\nLet Ow denote the center of mass of the wheelset. With Ow as the origin we introduce the coordinate system Ow x , yw z, such Ow y, coincides with the\nD ow\nnl oa\nde d\nby [\nU ni\nve rs\nita t P\nol it\u00e8\ncn ic\na de\nV al\n\u00e8n ci\na] a\nt 0 3:\n12 2\n9 O\nct ob\ner 2\n01 4",
+            "DYNAMICS OF A WHEELSET ON ROLLER RIG\nwheelset axis of rotation, z, axis points downward and x , axis completes the right-handed coordinate system. For the description of the surfaces of the wheels and the positions of the contact points we introduce another coordinate system, shown in Figure 1, with 0; (j = 1 for the right wheel and j = 2 for the left) as the origin and E,,, qwj, c w j as axes. The origin 0; coincides with the central position of the wheel. In the wheelset axes, 0, x, y, z, its coordinates are given by (0, + b, r) where 2b is the distance between the centers of the wheels and r represents - their radii. Following de Pater [7], we describe the profiles of the wheels by the relation\nThus for the surface of the wheels we can take\nwhere\nTo see that Equation ( 2 ) adequately represents the surfaces of the wheels, notice that for 6 , = 0 it reduces to Equation (1) and for a fixed -qwj it reduces to a circle with radius r + fwj(llwj).\nWith respect to the reference system Ow x , y, z,, the coordinates of the point of contact on wheel j are ([,, k ( b - T,), r + twj).\nIn a similar fashion, for the description of the surface of roller j, we introduced the right-handed coordinate system Orj x rj yrjzrj such that the yrj axis coincides with the axis of rotation of the roller j and zrj axis is directed downward. The origin Oij of the local coordinate system trj-qrjLrj, shown in Figure 2 for j = 1, coincides with\nD ow\nnl oa\nde d\nby [\nU ni\nve rs\nita t P\nol it\u00e8\ncn ic\na de\nV al\n\u00e8n ci\na] a\nt 0 3:\n12 2\n9 O\nct ob\ner 2\n01 4",
+            "RAO V. DUKKIPATI\nFig. 2. Right Roller coordinate Systems.\nthe central position of the point of contact and has coordinates (0, 0, -r,) in the roller axes xrjyrjzrj where rx is the radius of the roller.\nLet the profile of roller j be described by the relation\nFor the surface of the roller we can therefore take\nwhere\nThe coordinates of the point of contact j in the reference system Or, x,y,z, are (tri, T ~ r i > -'x + (ri).\n~ e t the wheelset be placed symmetrically on the rollers in such a way that OLj coincides with 0; and be the contact point. In this central position we take the center of the wheelset as the origin of the inertial coordinate system. For the y-axis we take the axis of rotation and z-axis points downward (see Figure 3). As the\nD ow\nnl oa\nde d\nby [\nU ni\nve rs\nita t P\nol it\u00e8\ncn ic\na de\nV al\n\u00e8n ci\na] a\nt 0 3:\n12 2\n9 O\nct ob\ner 2\n01 4"
+        ]
+    },
+    {
+        "image_filename": "designv11_60_0000804_jahs.61.042006-Figure7-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0000804_jahs.61.042006-Figure7-1.png",
+        "caption": "Fig. 7. Face-gear dynamics model: spinning flexible disk and a meshing load unit with prescribed radial and circumferential movements.",
+        "texts": [
+            " When the time is between time B and time C, only one meshing pair engages and the contact centroid overlaps with this contact point (see Fig. 5(b)). The positions of contact points and contact centroid during two adjacent meshing cycles are plotted in Fig. 6. The radial and circumferential positions are shown in Figs. 6(a) and 6(b), respectively. Face-gear model Based on the assumptions and analyses above, the face-gear is modeled as a homogeneous, isotropic, flexible Kirchhoff plate with uniform thickness and it is rigidly clamped at the inner boundary (r = b) and free at the outer edge (r = a), illustrated in Fig. 7. The concentrated meshing load is modeled as a unit comprising pinion mass, mp , pinion bearing stiffness, kb, and damping, cb. This unit is normal to the disk with inplane prescribed movements, including both radial and circumferential periodic oscillations. The transverse movement of the spinning disk is derived from Kirchhoff\u2013Love plate theory, and the meshing load unit is added to the model by coupling the transverse displacement of the disk and the restoring force of the unit. Following the work by Pei et al"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0003331_12.474658-Figure1-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0003331_12.474658-Figure1-1.png",
+        "caption": "Figure 1. 6-legged Stewart Platform(Hexapod). fPg is a Cartesian coordinate frame located at, and rigidly attached to, the payload's center of mass. fBg is the frame attached to the (possibly moving) base, and fUg is a Universal inertial frame of reference.",
+        "texts": [],
+        "surrounding_texts": [
+            "Most failures can be characterized as either soft failures or hard failures depending on whether the failed struts are present or not. This paper extends the work of Ref. 6 by including hard failures. Soft failure is caused by an abnormal but present strut. It is soft in nature. There are two types of soft failures: position failures and torque failures."
+        ]
+    },
+    {
+        "image_filename": "designv11_60_0001996_s0736-5845(98)00026-x-Figure18-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0001996_s0736-5845(98)00026-x-Figure18-1.png",
+        "caption": "Fig. 18. Global configuration planning method used for torso washing.",
+        "texts": [],
+        "surrounding_texts": [
+            "The programming methods, path planning, configuration planning and motion planning will now be used for offline-programming the world\u2019s largest robot, the mobile aircraft cleaning manipulator SKYWASH SW33. The SW33\u2019s redundant kinematic consists of 5 main axes, 5 axes denoting the SW33\u2019s wrist and one axis denoting the angle of the SW33\u2019s rotating brush. The SW33\u2019s radius of action is 33 m. The first task to be programmed is moving the cleaning brush along a path on the hull of an aircraft. The path length on the front part of the aircraft\u2019s torso is 65.0 m. The 1350 data points representing the path are ordered in a relative distance of 0.05 m (see Fig. 14). The second task to be programmed is moving the cleaning brush along the aircraft\u2019s engine with a path length of 45.0 m. The 900 data points representing the engine\u2019s path are also arranged in a relative distance of 0.05 m (see Fig. 15). Fig. 15. Path of datapoints on the aircraft\u2019s engine. First a curve is fitted to the data points by a piecewise cubic polynomial used in the robot control. The tolerance of the fitted curve to the data points is 0.05 m in position and 0.5\u00b0 in orientation. Using the path planning method the number of knots on the torso\u2019s path is 27, the number of knots on the engine\u2019s path is 38, with respect to the required tolerance (see Figs. 16 and 17). The SW33\u2019s configuration will be calculated in knots having specified geometrical characteristics. The constraint for maximum pressure is 2.7]107 N/m2 and for minimum distance between SW33 and aircraft is 1.5 m each main axis and 0.5 m each wrist axis. Using the global configuration planning method we get the configurations in the six specified knots in the case of torso washing and in the 10 knots in the case of engine washing. The profit of putting manipulability and desired configuration into the optimization criteria show Figs. 18 and 19, because two globally planned configurations in order do not pass singularities. Energy optimal configurations in the rest 27!6\"21 and 38!10\"28 knots will be planned within the Fig. 16. Knots of the spline curve fitted to the torso\u2019s path. Fig. 20. Local configuration planning method used for torso washing. Fig. 21. Local configuration planning method used for engine washing. optimal configuration space obtained by the global method. Figs. 20 and 21 show how the method works between a pair of globally planned configurations. Controlling the collision distance between SW33 and aircraft in all knots results 1.72 m minimum distance of the main axes and 0.54 m minimum distance of the wrist axes. Finally, the time optimal velocity profile for the SW33\u2019s motion along the specified paths was calculated, where the following constraints are given. With respect to the constraints in Figs. 22\u201424 the SW33\u2019s time optimal velocity profile gained by the motion planning method leads to a complete motion time of 179 s along the torso path and 281 s along the engine path. Driving nearly always with maximum endeffector speed of 0.4 (m/s) at the torso path shows the effect of progressing manipulability and energy cost towards an optimum within the configuration planning method (see Fig. 25). The process speed of cleaning the engine\u2019s path is limited by 50% of the maximum speed, Fig. 26 shows that this speed is nearly reached along the whole path. Analyzing the dynamics of the flexible subsystem we get the joint torque trajectories q (t) of the wrist axes, with and without the average joint torque each, by Figs. 27 and 29. Figs. 28 and 30 show the FFT-analysis of the robot, washing the aircraft\u2019s torso and washing the aircraft\u2019s engine. In Figs. 27 and 29 1 (kNm) denotes 1000 (Nm), in Figs. 28 and 30 F(- )\"1 is equivalent to q\"3 (Nm). Knowing that the robot\u2019s natural frequencies are beginning at 2.0 Hz, Figs. 28 and 30 ensure that there is no Fig. 29. The robot\u2019s joint torque trajectory during washing the torso. vibration stimulation by the programmed motion, because the amplitude F (- ) at - 0 \"2.0 (1/s) is far below the influence of motion damping by joint friction."
+        ]
+    },
+    {
+        "image_filename": "designv11_60_0002743_0301-679x(90)90005-a-Figure5-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0002743_0301-679x(90)90005-a-Figure5-1.png",
+        "caption": "Fig 5 A schematic of the load system",
+        "texts": [
+            " The test apparatus was designed to test four 23024 EASK MB-C3 spherical roller bearings simultaneously under the test conditions shown in Table 1. The method of loading the test bearings was modified from the Sullivan and Lyman 3 design in that the loading system simulated the load on actual drier roll bearings. The configuration of two bearings supporting a shaft was chosen, with the load applied to the shaft by a large spring through a third larger, more robust 'load' bearing mounted between the test bearings (Fig 4). The mechanism for applying the load (Fig 5) consisted of a spring, hydraulic jack and a load cell. The hydraulic jack compressed the spring to the desired load and the frame was tightened. Then, the hydraulic jack was released and the frame transferred the load to the shaft through the load bearing. The load cell provided an accurate value for the load applied by the spring. TRIBOLOGY INTERNATIONAL 321 S. D. Fysh et aI--Experimental simulation of the tribology of large spherical roller bearings Each shaft assembly was mounted on two C-sections which formed a box beam"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0003014_aim.2001.936474-Figure1-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0003014_aim.2001.936474-Figure1-1.png",
+        "caption": "Fig. 1 PLANE VIEW OF THE 4WS/4WD VEHICLE.",
+        "texts": [],
+        "surrounding_texts": [
+            "Keywords- Mechatronics, Wheeled Vehicles, Tractability, Efficiency\nI. INTRODUCTION\nobjective of the work described below is to apply THE mechatronics to increase the stability, tractability and efficiency of wheeled vehicles operating in difficult conditions (e.g. on steep and/or slippery slopes). Under these conditions the vehicles steering system and drive system will tend to fail. Under ideal conditions there is very little movement of the contact patch of the tyres relative to the ground. ScufEng results when the contact patch of the tyre moves parallel to the wheel axis. Skidding results when the contact patch of the tyre moves at right angles to the wheel axis. Wheel spin is regarded as negative skidding.\nA . The two basic methods of steering a wheeled vehicle\nIt has been shown previously [I, 2,3,4] that there are two basic methods of maneuvering a vehicle. One method is to turn one or more steerable wheels. The other is to drive one or more left-hand drive wheels independently of one or more right hand drive-wheels. If the two systems produce different radii of curvature for the path of the vehicle conflict will result in a compromise between the paths that would be produced by either system alone. Such conflict result in scuffing and skidding which wastes fuel, inflicts damage on the ground and reduces tractability and manoeuvrability. In a traditional road vehicle the above mentioned conflict is avoided by disabling the steering effect of the drive wheels by incorporating a differential in their drive train. Conversely in zero turn radius mowers or robots, steering is affected by driving the left hand and right drive wheels independently. In this case, the steering effect of the nondriven wheels is eliminated by allowing them to turn to any angle (ie. they are turned into castors).\nThe authors are with Gippsland School of Engineering, Monash University, Churchill, Vic. 3842, Australia. Email: Yousef.IbrahimQeng.monash.edu.au\nB. Problems posed by the traditional methods\nIn diflicult conditions disabling one of the basic steering systems leads to traction and stability problems. For example the differential is the Achilles\u2019 heel of the traditional tractor when traversing steep slopes. Similarly the inabdity of castors to transmit side way loads limits the ability of zero turn radius mowers (or robots) t o traverse steep slopes [1,2,4]. In diiEcult conditions traction and stability can be improved by allowing both steering systems to operate so that one system overpowers the other. For example, in four-wheel motorbikes the differential is often eliminated, resulting in increased traction and stability at the expense of introducing a scuffing problem. Conversely in skid-steer vehicles the steering effect of the wheel directions is overpowered by the steering effect of driving the left-hand drive wheels independently of the right hand drive wheels. In this case traction and Stability is improved at the expense of a severe scuffing problem on turning. A much better method of negotiating dScul t terrain is to allow both steering systems to operate but to integrate them so they both tend to produce the same radius of curvature. Although one system is extremely redundant, when one system fails the other system takes over. In other words the two systems mutually reinforce each other. The hypothesis to be tested by the work below is that if an on-board computer is used to integrate both steering systems, then the vehicle produced will combine the non \u2018LscutEng\u201d advantages of traditional road vehicles with the non \u201dskidding\u201d advantages of skid steer vehicles. In traditional road vehicles, rotation and translation are generally linked. Tkanslation of the vehicle along a curved path generally involves rotation. Also, rotation of the vehicle always involves translation. As a consequence rotation and translation in a conked space can be a problem. Vehicles steered by independently driving the left and right hand wheels have improved manoeuvrability since they have been made to rotate about their own centre. This is a pure rotation. Manoeuvrability can be further increased by allowing translation in any direction without the need for rotation. This pure translation is sometime referred to as crab steering. Furthermore, if any centre of curvature can be selected by the driver, manoeuvrability of the vehicle is improved further. The vehicle can execute either pure rotation or pure translation or any combination of both translation and rotation. The preferred means of driver control of the four-wheel steering/four-wheel drive (4WS/4WD) vehicle is by means of rotatable joystick. In this means of driver control, the\n0-7803-6736-7/01/$10.00 0 2001 IEEE 320",
+            "2 IEEE /ASME\ndirection of translation of the vehicle is determined by the direction of displacement of the joy stick from it\u2019s neutral position. In this case the rotation of the vehicle is determined by the degree of the rotation of the joy stick. The amount of displacement of the joystick determines the root mean square of the four wheel speeds. Pure translation occurs when the joystick is displaced but not rotated. Pure rotation occurs when the joystick is twisted as far as it will go. This system is similar to the system used to control a lame aircraft.\nthe radius curvature of the vehicles\u2019 path will be infinity, and the vehicle will move in a straight line parallel to the direction of displacement of the joy stick. If the joystick is twisted as far as it will go in a clockwise direction, the radius of curvature will be zero and the vehicle will rotate clockwise around it\u2019s own centre. Between these extremes the radius of curvature of the path of the vehicle R will be given by Equation 1:\nWhere t is the track of the vehicle, 0 is the rotation of the joy stick and e,,, is the maximum rotation of the joy stick. The driver selects the direction of the centre of curvature by displacing the joy stick at right angles to this direction. The centre of curvature of the path to the vehicle is now specified by the two components Rz and Ry. The driver then selects the root mean square of the four-wheel speeds (RMSWS) by the amount of displacement of the joy stick. The control system then rotates the four drive wheels to the following angles, Equation 2:\nt an+3=& ; tan#J3=\nwhere b is the wheel base of the vehicle, & is the displace Although a four-wheel steering/four-wheel drive ( 4 w S / 4 m ) ment of the centre of curvature forward of ihe centre of the vehicle will be initially described, it will be shown later vehicle. Likewise, R, is the,displacement of the centre of that two-wheel steering/four-wheel drive ( 2 w s / 4 ~ ) ve- curvature to the right of the centre of the vehicle. hicle and two wheel steering/ two wheel drive (2WS/2WD) The amount of displacement of the joy stick d determines vehicles are special cases of the general case. It will also be the root mean square of the four wheel speeds (RMSWS) shown that six wheel drive vehicles (SWS/SWD) and eight according to Equation 3: wheel drive (SWS/SWD) vehicles can be easily developed from the four-wheel steering/four-wheel drive (4WS/4WD) (3) vehicle. (RMSWS)2 = + + w: + = KZd2 4\n11. THEORY\nplacement hydraulic pumps 2 and 3 which in turn drive hydraulic motors 4 and 5 mounted in the steerable front and rear right hand wheeh 6 and 7 respectively. The internal combustion engine 1 also drives left hand variable displaced pumps 8 and 9 which in turn drive hydraulic motors 10 and 11 which are mounted in steerable front and rear left hand wheels 12 and 13 respectively. Note that positive rotation of the front and rear wheels are clockwise and and anti-clockwise respectively. The angle of the wheels 6 , 12, 7 and 13 are slowed as G2, a3 and a4 respectively. The rotational speed of the wheels 6, 12, 7 and 13 are w1, w2, w3 and w4 respectively. The driver controls the vehicle by selecting the radius of curvature of the vehicle\u2019s path and the sense of rotation by rotating the joy stick. If the joy stick is not rotated Where: Where the root mean square radius (RMSR) in equation 4 is given by:",
+            "SPARK A N D IBRAHIM: I N T E G R A T E D MECHATRONICS SOLUTION\n1 2 RMSR = - (R: + @ + Rg + R;)'\nNote that when the rotation of the joy stick 8 is a maximum, the radius of curvature will be zero and the direction of the displacement d of the joystick will be immaterial. It would be natural for the driver to push the joystick forward to commence rotation. If the driver pulls the stick straight back, then the wheels will rotate in reverse. A disadvantage of the vehicle described above is that four independent steering systems and four independent drive systems are required. It will be shown below that under special conditions, the number of systems required can be reduced. If % = 0, the eight control equations become:\nb tan41 = & , tan42 = b \\\nand,\nWhere:\nIn this case only two wheel angle control systems and two wheel speed control system are required. Figure 2 shows that the relationship between the position of the rotatable joy stick and the wheel angles an wheel speed for a 4WS/4WD vehicle. The figure is schematic insofar as the rotatable joystick is enlarged for the sake of clarity. Figure 2(a) shows the general case where the vehicle rotates about a point not located on either axis of the vehicle. In this case the four-wheel angles are different and the fourwheel speeds are different. Figure 2(b) depicts the vehicle when Ry = 0. In this case d1 = 43, 42 = 44, w1= w3 and w2 = w4. Figure 2(c) depicts the vehicle when it moves straight ahead. In this case the wheel angles are zero and the wheel speeds are the same. Figure 2(d) depicts the vehicle when it moves straight to the left. In this case 41 and 4 2 are -90'. On the other hand,\n3\n43 and $4 are +goo. In this case all the wheel speeds are equal. Figure 2(e) depicts the vehicle when it moves in a diagonal direction. In this case 41 and 4 2 are -45O. On the other hand, 4s and 44 are +45'. Also, in this case all the wheel speeds are equal. Figure 2(f) depicts the vehicle when it executes pure rotation in a clockwise direction. In this case (if b = t ) 41 and 43 are +135O and 4 2 and 44 are +Go.\n111. SYSTEM EXPANSION TO MULTI-WHEELED VEHICLES\nThe system illustrated in Figure 2 can be further extended for different number of steering and driving wheels. Figure 3 shows how 2WS/4wD, 2WS/2WD, 4WS/6WD and 8WS/8WD vehicles can be developed from the 4WS/4WD vehicle. If 4 = -: then the eight control equations become:\nWhere:\nR: = b2 + (R, - 4)' Rg = b2 + (R, + 5)' R: = (R, - $)2 R: = (R, + $)2\ni\nIn this case no steering system is required for the rear wheel since 43 and 44 are always zero, Fig. 2(b). The vehicle is further simplified if either the front or rear wheels are undriven (ie. they are free wheels) so that only two speed control system are required see Fig. 2(b) and 2(c). Although the same equations apply to the 2WS/2WD vehicle as apply to the 2WS/4WD vehicle, there is no control imposed on the speed of the free wheels in the former case. In the 2WS/2WD case the speed of the free wheels could be ignored for the purpose of calculating the root mean square wheel speed. If the front wheels are free wheeling; the RMSR for the reax driving wheels will be as expressed in Equation 10, See Fig. 2(b).\nR M S R = ( R 3 +- :)+ If the rear wheels are free wheeling, the RMSR for the front drive wheels will be as expressed in Equation 11, See Fig. 2(c)."
+        ]
+    },
+    {
+        "image_filename": "designv11_60_0000962_tmag.2016.2636208-Figure2-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0000962_tmag.2016.2636208-Figure2-1.png",
+        "caption": "Fig. 2. Electromagnetic torque evaluation by surface integral over S.",
+        "texts": [
+            " Let us consider electromagnetic force/torque exerted in the conductive region of an electromechanical converter. It could be a moving part, e.g., carriage and rotor. Force/torque can be evaluated by surface integral over the surface S placed in the nonconductive region out, which surrounds the conductive region . The outer region out can be the gap of an electromechanical converter. The nonconductive region out usually does not involve electromagnetic force or torque. The integration surface S in the gap out can be placed in a different way, i.e., its radius may vary (Fig. 2). Mostly, for electromechanical converters (e.g., rotating electric machines), the electromagnetic torque value does not depend on the radius of surface S placed in the gap. The independence of surface integral results from magnetic feature of the gap, which is usually the air gap (isotropic gap). However, the air gap out (which surrounds the conductive region , e.g., carriage, cylindrical, or spherical rotor) could be magnetically anisotropic, i.e., could be filled with ferrofluids [9], [11]. The second theorem introduces a sufficient condition for surface-integral representation for different magnetic properties of nonconductive region out (e"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0001081_icarcv.2016.7838776-Figure13-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0001081_icarcv.2016.7838776-Figure13-1.png",
+        "caption": "Figure 13. Angle of robot hand",
+        "texts": [],
+        "surrounding_texts": [
+            "IV. JUSTIFICATION FOR USE OF RUBBER STICK For humans, local and speedy inversion motions of the wrist (snap motions) are important for glockenspiel playing. Thus, Hiro performs the striking motion holding an RS shorter than its arm (Fig. 6, Table 1). That is, the robot can perform snappy motions by controlling the RS speedy inversion motion. Because Hiro uses trapezoidal speed control, a stiff stick (e.g., plastic or wood) would stop the glockenspiel vibration by bearing down on it. Thus, RS characteristic vibration is needed.\nwas used to express the motion equations\n(a) 20 hTt \u2264\u2264 ,\n.0cossin 1 2 1 2 1 2 1 2 =\u22c5\u22c5\u2212++ hhhh tAmlklclml \u03b8\u03c9\u03c9\u03b8\u03b8\u03b8 (2)\n(b) tTt h << 2,0 ,\n.02 2 2 2 2 2 =++ hhh klclml \u03b8\u03b8\u03b8 (3)\nWhere , Th, and are the input angular frequency [rad/s], input power cycle [s], and robot input angle [\u00b0], respectively,\nand h1, h2 are the RS swing angles [\u00b0]. The simulated values given by (2) and (3) are shown in Fig. 9, where the input power is the phase coordinate for the wooden ball vibration. This is a sympathetic vibration, which increases the vibration at the RS top. Further, Hiro\u2019s percussive motion is considered so as to prevent double striking of the object via proper damping. For example, the glockenspiel is located where the stick amplitude is displaced from z1 to z2. Thus, Hiro can autonomously achieve effective motion by shaping the input of a flexible RS via sympathetic vibration.\nFurther, Fig. 10 shows the vibration modes of key C6, while Fig. 11 shows the corresponding acoustic information for",
+            "the impact points a and b in the former figure. There are the differences between the impact points a and b. Therefore, Hiro can recognize an object by impact point as well as frequency. In addition, the robot can recognize the vibration mode in the case of low background noise.\nVI. RUBBER-STICK VIBRATION MODEL DURING MOVEMENT TO TARGET POSITION\nWe studied the robot movement used to play the glockenspiel with a tool. An appropriate tool was chosen based on its vibration analysis. Note that the RS vibrates when the robot arm moves to a different key rather than performing the impact task. This vibration was analyzed by examining the RS in the robot\u2019s right hand. The initial position was defined as key C7. Figs. 12 and 13 show the relationship between the stick and keyboard and the angle of the stick-holding robot hand \u03d5,\nrespectively. First, the robot hand engaged in z-axis rotation by \u03d5 [\u00b0]; then, the key-to-key movement Y [mm] results were obtained.\nAs \u03d5 increases, the RS vibration in the y-axis movement (zaxis) direction is suppressed (increased), because the external force causing the movement is divided. Thus, Y sin\u03d5, which causes the z-axis direction vibration, increases in the 0 \u03d5 90 [\u00b0] range. We considered this vibration model and attempted to control it.\nFig. 14 shows the RS coordinate system. The wooden-ball vibration was modeled as a transient vibration with displacement y0. The quarter sine",
+            "<< \u2264\u2264 = )4,0(.0 )40(.sin 0 tTt TttY y\nm\nm\u03c9 (4)\nwas used to express the motion equations\n(a) 40 mTt \u2264\u2264 ,\n.0)cos(sin 1 2 1 2 1 2 1 2 =+\u22c5\u22c5\u2212++ \u03d5\u03b8\u03c9\u03c9\u03b8\u03b8\u03b8 mmmm tAmlklclml (5)\n(b) tTt h << 4,0 ,\n.02 2 2 2 2 2 =++ mmm klclml \u03b8\u03b8\u03b8 (6)\nwhere is the input angular frequency [rad/s], Tm is the input power cycle [s] and m1, m2 are the RS swing angles [\u00b0]. To travel to B6 from C7, the robot hand moves in the y-axis direction only. The corresponding RS vibration was again observed by a HSVC. The simulated values given by (5) and (6) are shown in Fig. 15. The differences between the theoretical and experimental values are less than 5 [mm] at all amplitudes. Thus, this model is thought to be appropriate.\nVII. RUBBER-STICK VIBRATION SUPPRESSION USING INPUT SHAPING DURING MOVEMENT TO TARGET POSITION\nWe examined a movement method for controlling the unnecessary vibration during key-to-key movement. The conventional method (Pattern 1 in Fig. 16) reaches the target position in one movement. In contrast, Pattern 2 divides the movement into two components.\nThe RS in the robot\u2019s right hand was considered (Fig. 12) and key C7 was taken as the initial position. The robot moved in the y-axis direction to key B6 and the stick tip vibration was observed using a HSVC. The robot moved at 100% motion velocity. For Pattern 2, (Y1, Y2) = (15, 15), (20, 10), and (25, 5) (Y1 and Y2 are defined in Fig. 16). The robot moved the RS from key C7 to B6 using each movement method. Fig. 17 shows the absolute value of its vibration in the y-axis direction for each case. The vibrations are more obviously reduced for the Pattern-2 movement method. However, the degree of vibration restraint varies with (Y1, Y2). We then considered (Y1, Y2) = (20, 10), which restrains the vibration most significantly.\nFigure 17. Vibrations for each movement method\n-15 -10 -5 0 5\n10 15 20 25\n0 0.2 0.4 0.6 0.8 1\nA m\npl itu\nde [m\nm ]\nTime [s]\nInput power Y Input power Y Simulation result\nt t2\nFigure 18. Relationship between vibration and input power"
+        ]
+    },
+    {
+        "image_filename": "designv11_60_0000822_s11015-016-0331-6-Figure1-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0000822_s11015-016-0331-6-Figure1-1.png",
+        "caption": "Fig. 1. Schematic of the CRE process (a) and billet profi le within the deformation zone (b): 1) billet; 2) rolls; 3) die (half deformation zone shown).",
+        "texts": [
+            " The combined rolling\u2013extrusion (CRE) process employing two rolls, rather than one wheel as in CONFORM, has been developed in recent years (see [8\u201310] for detailed descriptions). Our goal here is to study the parameters of the CRE process with a porous stock material. A porous material is used because of the necessity of processing powders, granules, chips, and other loose materials [11]. At the fi rst stage of processing, the loose material is usually compacted until being capable of holding its shape. The next stage is extrusion of a porous compact to produce the desired cross-section. The process is schematized in Fig. 1a. Billet 1 is fed to the pass formed by two rolls 2. On its way, the billet meets die 3. The rolls create friction stresses that push the billet through the die. Figure 1b shows the distortion of the billet within the deformation zone. To solve the problem, we will use the following parameters and boundary conditions. The geometry of the deforma- tion zone is shown in Fig. 2. The effective roll radii R1 = 53.5 mm and R2 = 40.5 mm. The pass width b = 15 mm, the minimum 1 PLM Ural Company Group, Ekaterinburg, Russia; e-mail: eaa@plm-ural.ru. 2 Ural Federal University, Ekaterinburg, Russia; e-mail: j.n.loginov@urfu.ru. 3 Siberian Federal University, Krasnoyarsk, Russia; e-mail: kafomd_1@mail"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0000509_s40815-016-0202-0-Figure5-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0000509_s40815-016-0202-0-Figure5-1.png",
+        "caption": "Fig. 5 Mass\u2013Spring\u2013Damping system",
+        "texts": [
+            " Then, the simulation results are shown in Figs. 1, 2, 3, and 4, where Fig. 1 and Fig. 2 show the trajectories of xi\u00f0i \u00bc 1; 2\u00de and their estimates x\u0302i\u00f0i \u00bc 1; 2\u00de, respectively; Fig. 3 expresses the trajectories of control input ur\u00f0r \u00bc 1; 2\u00de; Fig. 4 shows the trajectory of switching signal r. From the simulation results, it is clear that the proposed output feedback control method can guarantee the stability of the closed-loop switched fuzzy system. Example 2 Consider the mass-spring-damping system [41] shown in Fig. 5 and according to Newton\u2019s law, it follows as \u2018\u20acx\u00fe Ff \u00fe Fs \u00bc u where \u2018 stands for the mass of the spring, Ff and Fs are the friction force and the restoring force of the spring, where the variables are the nonlinear or uncertain terms. u denotes the external control input. Assume that the friction force Ff \u00bc t1 _x 3 with t1 [ 0 and the hardening spring force Fs \u00bc t2x\u00fe t3x 3 with constants t2 and t3. Then, the dynamic equation can be written as \u20acx \u00bc \u00f0t1=\u2018\u00de _x3 \u00f0t2=\u2018\u00dex \u00f0t3=\u2018\u00dex3 \u00fe \u00f01=\u2018\u00deu where x stands for the displacement from a reference point"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0000734_cscwd.2016.7565955-Figure5-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0000734_cscwd.2016.7565955-Figure5-1.png",
+        "caption": "Fig. 5. The finite element mesh of IWRC6*36WS wire rope",
+        "texts": [
+            " Because of the complex structure of wire rope, the contact relationship between each wire and the adjacent steel wire is analyzed by using the finite element method to analyze the complex contact problem. The explicit dynamic method does not exist in the problem of convergence and requires less computer memory. So the wire rope model for this complex structure, the finite element software Abaqus/Explicit solver of Abaqus is for analysis. In the ABAQUS/Explicit unit, C3D8R is used to generate the finite element mesh. A pitch of IWRC6 * 36WS wire rope is divided into 1186304 units, with a total of 1530819 nodes. As shown in Fig. 5. Analysis of the default hourglass stiffness uses Abaqus/Explicit system. The contact relationship between each wire in wire rope is set by the general contact setting of Abaqus/Explicit, the normal contact characteristic is defined as \"Contact Hard\", which can be separated. Taking into account the tangential friction between the steel wires, the friction coefficient between the wires sets 0.15. The two reference points RP-1 and RP-2 which have a distance of 20mm to the two rope end are created at the wire rope"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0003861_s0167-8922(08)71083-6-Figure2-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0003861_s0167-8922(08)71083-6-Figure2-1.png",
+        "caption": "FIG. 2- Dependence of total pin wear on pin hardness, HV (kgf.rnrn-21",
+        "texts": [
+            " 2 5 pm Nuclepore filters proved most successful and also provided a mounted sample on a flat substrate for scanning electron microscopy (SEM). Typically, 0 . 5 to 2 ml of test oil were filtered after dilution of hexane. Radiotracer monitoring of wear debris from the martensitic materials showed that over 90% of the activity was retained on the filters, indicating that most of the worn material was present as solid debris. A radiotracer technique was employed to 3 . WEAR MEASUREMENTS The variation of the total pin wear with Vickers hardness (Hv) is shown in Figure 2 . Wear of the tempered martensite pins varied from 31 pg for the hardest pin (A) to over 9000 pg for the softest pins (N and J). The maximum standard deviation for radiotracer counting of total pin wear results of up to 100 pg was approximately 2 pg, increasing in the case of pin X to 4 p g . These are rather less than errors from other sources, e.g. collection of wear debris. The wear of the two softest pins N and J , was several times the thickness of the activated layer. The loss of mass was therefore estimated from the change of length of the pin, which was measurable within about 2% accuracy",
+            " Conventional models for the prediction of wear are usually defined in terms of two wear coefficient.\" wear coefficient is defined as The dimensional k - V/SL (mm3 mm-' N-l) . . . (1) Where V is the wear volume in mm3 and L is the applied normal load in N. The dimensionless wear coefficient, S is the sliding distance in nun, K - 3H.k . . . (2) where is usually expected to have a constant value for a given combination of materials and environmental conditions. Values for k for most of these experiments (Fig 2 ) were reasonably typical of the lubricated mild wear regime as defined by Childs', on the basis of other work, as less than about 10-l'. SEM examination of the worn surfaces and wear debris indicated the occurrence of at least two processes for the generation of wear particles, different processes being predominant at the extremes of the hardness H is the material hardness in Nnui2, 345 range studied. The softest surfaces (Fig. 5a) suffered severe plastic deformation, forming small particles with a thin platelet morphology consistant with the lamellar folds of material extruded at the trailing edge",
+            " The latter represents the lower end of the ran e predicted by Challen et all6 10- ) for low cycle fatigue wear at the mean combined profile slope of 1.5 degrees prevailing in the present work. Values of A /Ac and %/A were obtained from Eqn. 16 for tie average vayue of the surface topography parameter ( u /Rs)\u2018 of 0.06. The deviations from the rigid-plastic predictions, represented by chain-dotted line in Figure 9, are indicated by curves of k, and $. reasonably good agreement with the wear behaviour of the tempered martensites (Fig. 2 ) . The wear of the softest pins, N and J is underestimated in Figure 9 because of the limitations of the elastic contact model already discussed. The inclusion of a more realistic elastic-plastic model would be expected to result in a more abrupt transition between the two processes, giving better agreement with the experimental observations. Changes in the surface topography also have an important influence on k (Figure 10). The N and J pins both had final values of (us/Rs)\u2019 in the region of 0"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0000188_0954406214560420-Figure5-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0000188_0954406214560420-Figure5-1.png",
+        "caption": "Figure 5. Equivalent schemes of a RR limb.",
+        "texts": [
+            " Obviously, the point p0 is not at the given motion path. In other words, the point p0 is beyond its workplace. Also, the mobile platform can be induced to exceed its valid workspace by the Jacobian matrix of the coupled fewer DoF mechanisms, leading to the failure of the FK solution. Equivalent model of the limb with extremely displacement singularity Planar limbs are used to illustrate how to remove the external displacement singularity, which can be extended to spatial limbs. Just as shown in Figure 5, the workspace of the end-effector is 2D, restricted in the circle by the fixed lengths of two links. Ceccarelli23 studies the chain equivalence to deduce new architectures for parallel manipulators. But it is not enough to remove the extremely displacement singularity. In the figures, the blue position is the target position, and the pink position is the position in process, which exceed the workspace. The generalized coordinates of the RR limb is q \u00bc \u00f0\u20191 \u20192\u00de T. In Figure 5(a), the second link is liberated. A virtual prismatic (P) joint is added along the second link. The generalized coordinates of the equivalent mechanism is q0 \u00bc \u00f0\u20192e le\u00de T, where \u20191 is used to determine the position of hinge point a. So, the RR limb is converted to an RP limb, whose workspace is not restricted. The constraint equation is transformed to the joint variable of the virtual P joint. The generalized coordinate corresponding to the target position is q0m \u00bc \u00f0\u20192 lap\u00de T, where lap is the length of the second link, a fixed value. Also, the virtual P joint can be added to the first link, just as Figure 5(b) shows. The generalized coordinates of the equivalent mechanism is q0 \u00bc \u00f0\u20191e le\u00de T. The target position of the equivalent mechanism is q0m \u00bc \u00f0\u20191 loa\u00de T, and \u20192 is used to keep the orientation of the second link with respect to the first link. Just including P joints is not sufficient to remove the extreme displacement singularity. A PR limb is shown in Figure 6. Although there is a P joint, its workspace is a ribbon area, whose width is two Figure 4. Coupled motion of a rigid body. at Univ Politecnica Madrid on January 14, 2015pic"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0003124_70.88068-Figure6-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0003124_70.88068-Figure6-1.png",
+        "caption": "Fig. 6 . Wrench axis representation and intersection with object A , and A 2 .",
+        "texts": [
+            " 4, AUGUST 1989 Point Contact with respect to the compliance frame The contact force vector is expressed as a 6 x 1 column vector W = { W l r w29 w3, w4, w51 W6}T* ( 5 ) Normalizing the above by the following: we have where The components of { a } can be viewed as Plucker line coordinates of the wrench acting along the wrench axis having direction cosines of ( l , m, n ) and moment lever arm L with respect to the compliance frame L = ( a x ao) (9) Therefore, the above information can be used to uniquely determine the wrench axis corresponding to the contact configuration. A more general parametric representation of the wrench axis is as follows [8]: Y = L, + mt Z = L,+ nt. (10) The above is called the freedom-equation of a line in terms of the parameter t , where t is the magnitude of the line. The intersection between L and the wrench axis (L,, L,, L,) are given by L, = np - Ir L ,=lq-mp. (1 1) Contact points can be determined by finding the intersection of this wrench axis and the surface representing the object. For example, in Fig. 6 the intersection of the wrench axis with the cylindrical representation of the manipulated object can be found by substituting (10) into the cylinder equation which yields (b2l2+aZm2)t2+2(b2Lxl+azL,m)t + ( b2L:+a2LF- a2b2) = 0. (12) The above is quadratic in t and therefore the wrench axis will intersect the quadratic surface at two locations. The two intersection points will coincide if the wrench axis is tangent to the cylinder surface. Line Contact In general, line contact between the object and the obstacle environment can be expected along the surface of the quadratic representation of the object, along the edge of the polyhedra representation, or on the surface of polyhedra located at the edges of the obstacle environment (Fig"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0003471_20.11540-Figure1-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0003471_20.11540-Figure1-1.png",
+        "caption": "Fig. 1 - A generic triclinic element: a prism of rectangular cross section truncated by a plane inclined obliquely at any angle relative to the axis of current.",
+        "texts": [],
+        "surrounding_texts": [
+            "1544 IEEE TRANSACTIONS ON MAGNETICS, VOL. 24, NO. 2, MARCH 1988\nNEW ANALYTICAL FORMULAS FOR CALCULATING MAGNETIC FIELD\nC. F. Weggel and D. P. Schwartz Magnet Ana lys i s f o r Government, I n d u s t r y , and Col leges\n643 R u s h v i l l e S t r e e t La J o l l a , C a l i f o r n i a 92037\nA b s t r a c t\nSimple a n a l y t i c a l formulas have been d e r i v e d f o r c e l - c u l a t i n g t h e magnetic f i e l d generated by a ve ry genera l and v e r s a t i l e u n i f o r m c u r r e n t - d e n s i t y f i n i t e element: a r e c t a n g u l a r - c r o s s - s e c t i o n p r i s m te rm ina ted a t each end by an a r b i t r a r y o b l i q u e end-plane. These new formulas enable t h e user t o c a l c u l a t e e f f i c i e n t l y t h e exac t magn e t i c f i e l d generated by any system b u i l t o f s t r a i g h t , rec tangu l a r - c ross -sec t i on , u n i f o r m - c u r r e n t - d e n s i t y conduc to rs . The formulas may a l s o be used t o c a l c u l a t e t h e magnetic f i e l d generated by any system c o n t a i n i n g curved elements such as c i r c u l a r c o i l s , he l i xes , toruses, t o r - sat rons, saddle c o i l s , and y in -yang c o i l s by approximati n g t h e curved elements by concatenated sho r t , s t r a i g h t , t rapezo id91 segments. Even a f a i r l y coarse mesh o f such f i n i t e elements has proven s u f f i c i e n t t o y i e l d an accur a c y of 1 p a r t p e r m i l l i o n , o r b e t t e r . The new formulas possess t h e advantage over e x i s t i n g e l l i p t i c i n t e g r a l formulas i n t h e i r g r e a t e r g e n e r a l i t y and s u p e r i o r accuracy, e s p e c i a l l y f o r f i e l d p o i n t s w i t h i n t h e c u r r e n t - c a r r y i n g volume o r on i t s boundary, where even t h e bes t e l l i p t i c i n t e g r a l formulas e x h i b i t l o g a r i t h m i c s i n g u l a r - i t i e s and d i s c o n t i n u i t i e s .\naccuracy and s i m p l i c i t y o f t h e new formulas have been m e t i c u l o u s l y v e r i f i e d us ing MACSYMA, t h e powerfu l symbol ic man ipu la to r developed a t t h e Massachusetts I n s t i t u t e o f Technology. Using MACSYMA t o success i ve l y d i f f e r e n t i a t e and s i m p l i f y t h e formulas, we have reder i v e d a l l o f t h e pub l i shed s imp le r spec ia l cases, o f t e n i n more compact o r more symmetric equ iva len t forms. The accuracy o f t h e new formulas f o r curved geometr ies has been v e r i f i e d by comparison w i t h e x i s t i n g codes f o r c i r - c u l a r c o i l s . These comparisons a re i l l u s t r a t e d i n an ex tens i ve s e t o f p l o t s o f t h e f i e l d e r r o r s .\nI n t r o d u c t i o n\nThe\nNo a v a i l a b l e computer code f o r computing t h e magnetic f i e l d generated by u n i f o r m - c u r r e n t - d e n s i t y elements i s y e t e n t i r e l y sa t i s fac to ry - - some a re i nexac t ; some a r e slow; some a r e bo th . Worse, t h e problems o f e r r o r and i n e f f i c i e n c y a r e most severe i n j u s t those cases most o f t e n o f g r e a t e s t i n t e r e s t : whenever t h e magnetic f i e l d w i t h i n t h e c u r r e n t - c a r r y i n g r e g i o n o r on i t s boundary must be c a l c u l a t e d .\nWhat i s t h e cause o f these shortcomings? A l l p r e v i - ous computer codes must o f n e c e s s i t y r e s o r t t o numer ica l i n t e g r a t i o n t o c a l c u l a t e t h e magnetic f i e l d generated by a l l except a few ve ry s imple u n i f o r m - c u r r e n t - d e n s i t y elements (such as an or thorhombic \" b r i c k \" ) . By d e f i n i - t i o n , such numer ica l i n t e g r a t i o n i m p l i e s approx imat ing each three-d imensional c u r r e n t element w i t h some s e t o f one- o r two-d imensional c u r r e n t elements such as c u r r e n t f i l a m e n t s , c u r r e n t loops, c u r r e n t sheets, o r i n f i n i t e l y - t h i n so lenoids. The i n t e g r a l i s then approximated by some j u d i c i o u s l y - c h o s e n weighted sum o f these 1 umped elements.\nThe accuracy which can be achieved w i t h such numeric a l i n t e g r a t i o n can be p e r f e c t l y s a t i s f a c t o r y as l o n g as one i s c a l c u l a t i n g t h e magnetic f i e l d a t p o i n t s which a r e e x t e r n a l t o t h e c u r r e n t - c a r r y i n g reg ion . But whenever one c a l c u l a t e s t h e f i e l d a t p o i n t s i n t h e i n t e r i o r o f t h e windings o r on i t s boundaries, such numer ica l i n - t e g r a t i o n becomes inaccurate--sometimes d r a m a t i c a l l y s o .\nWhy? A l l such one- and two-dimensional elements have\nze ro volume. Therefore, t o c a r r y a f i n i t e cu r ren t , such elements must c a r r y i n f i n i t e c u r r e n t d e n s i t y . Un fo r tu - n a t e l y , t h e magnetic f i e l d i s d i scon t inuous on t h e two s ides o f any i n f i n i t e - c u r r e n t - d e n s i t y c u r r e n t sheet o r so leno id . Even worse, t h e magnetic f i e l d i s i n f i n i t e ad jacen t t o any c u r r e n t f i l a m e n t o r c u r r e n t l oop .\nApprox imat ing t h e d i s t r i b u t e d c u r r e n t elements w i t h a f i n e r s u b d i v i s i o n o f i n f i n i t e - c u r r e n t - d e n s i t y elements can reduce, b u t can by no means e l im ina te , t h e det r iment a l e f f e c t o f t h e s i n g u l a r i t i e s . Such mesh ref inement , o f course, i s o f t e n bo th time-consuming and u n r e l i a b l e , s i n c e t h e convergence may be g l a c i a l l y slow.\nSuch c a l c u l a t e d d i s c o n t i n u i t i e s i n magnetic f i e l d a re e n t i r e l y f i c t i t i o u s , o f course--an i nhe ren t consequence o f t h e numer ica l i n t e g r a t i o n process. E labo ra te p r o - gramming techniques t o cope w i t h these s i n g u l a r i t i e s a re t h e h e a r t o f many computer codes. Such techniques, however, a r e bo th i n e l e g a n t and may be e r ro r -p rone .\nE x i s t i n q Uni form-Current -Densi ty Formulas\nU n t i l now, t h e most genera l a n a l y t i c formulas a v a i l - ab le f o r c a l c u l a t i n g t h e magnetic f i e l d generated by a u n i f o r m - c u r r e n t - d e n s i t y element a re t h e formulas t h a t my tw in b r o t h e r Robert and I d e r i v e d i n 1965 w h i l e we were bo th work ing a t t h e Massachusetts I n s t i t u t e o f Techno1 - ogy F ranc i s B i t t e r Na t iona l Magnet Laboratory . These a n a l y t i c formulas c a l c u l a t e t h e magnetic f i e l d generated by a f i n i t e - l e n g t h , rec tangu la r -c ross -sec t i on , un i fo rmc u r r e n t - d e n s i t y Drism. o r or thorhombic \" b r i c k \" . These r e s u l t s were G b l \\ s h e d . i n Solenoid Maqnet Desiqn by D r . D. Bruce Montgomery i n 1969 [ I ] .\nI have now extended t h e g e n e r a l i t y and usefu lness o f these r e s u l t s by d e r i v i n g s imple a n a l y t i c formulas f o r c a l c u l a t i n g t h e magnetic f i e l d generated by a f a r more genera l and v e r s a t i 1 e u n i f o rm-cu r ren t -dens i t y e lement- -a p r i s m w i t h a r e c t a n g u l a r c r o s s - s e c t i o n te rm ina ted a t bo th ends by a r b i t r a r y o b l i q u e end-planes. The new f o r - mulas enable t h e use r t o c a l c u l a t e e f f i c i e n t l y t h e exact magnetic f i e l d generated by any system composed e n t i r e l y o f s t r a i g h t , rec tangu la r -c ross -sec t i on , un i fo rm-cu r ren t - d e n s i t y conductors . O f course, t h e formulas can a l s o be used t o c a l c u l a t e t h e magnetic f i e l d generated by any system which con ta ins curved elements, such as c i r c u l a r c o i l s and h e l i x e s , t o ruses and t o r s a t r o n s , saddle c o i l s and y in -yang c o i l s ( a l l w i t h r e c t a n g u l a r c ross -sec t i on ) , by approx imat ing t h e curved elements by concatenated sho r t , s t r a i g h t , t r a p e z o i d a l segments.\nThe accuracy and s i m p l i c i t y o f t h e formulas were met i c u l o u s l y v e r i f i e d u s i n g MACSYMA, t h e powerfu l symbolic man ipu la to r developed a t t h e Massachusetts I n s t i t u t e o f Technology. Using MACSYMA t o success i ve l y d i f f e r e n t i a t e and s i m p l i f y my new formulas, I have v e r i f i e d a l l o f t h e pub1 i shed s i m p l e r spec ia l cases, a l though f r e q u e n t l y t h e r e s u l t s were i n s impler , more compact, o r more symmetric forms.\nF i e l d o f \"Orthorhombic B r i c k \"\nI n p a r t i c u l a r , I have d iscovered t h a t t h e a n a l y t i c formulas f o r t h e magnetic f i e l d generated by an \"o r tho - rhombic b r i c k \" which Robert Weggel and I d e r i v e d i n 1965 can be expressed i n an e q u i v a l e n t form i n v o l v i n g a mere t h r e e terms i n s t e a d o f t h e f i v e terms pub l i shed i n 1969. The s imp le r e q u i v a l e n t formula f o r t h e x-component o f magnetic f i e l d f o r t h i s case i s as f o l l o w s :\n0018-9464/88/0300-1544$01.0001988 IEEE",
+            "1545\nBx = l . E - 0 7 * Jy * (1)\n(AMX * Atanh (BMY / ROOT) t BMY * Atanh (AMX / ROOT) - CMZ * Atan [ (BMY * AMX) / (CMZ * ROOT)])\nwhere the electric current density Jy (in amperes per square meter) is defined to flow parallel to the y-axis,\nand AMX = a - x, (2) BMY = b - y, CMZ -- c - z, and ROOT = Sqrt (AMX**2 t BMY**2 t'CMZ**2)\nwhere (x, y, z) are the Cartesian coordinates of the and (a, b, c) are the Cartesian coordinates of each \" f i el d poi n t \" ;\nof the 8 corners of the element.\ncase of the newly-derived finite element--a monoclinic wedge. Most importantly, for example, such elements can be concatenated to generate a polyhedral coil. The magnetic field generated by such a polyhedral coil can provide an excellent approximation to the field generated by an \"equivalent\" circular coil.\nAccuracv of the New Formulas\nThe accuracy depends on the degree to which the polygonal coil accurately models the circular coil; i. e . , the accuracy depends on the number of sides of the polygon. To ascertain the minimum number of sides required to ensure the desired accuracy, an extensive parameter analysis has been performed. The results of this parameter analysis are presented in Figures (3a through 39). As can be seen from these graphs, an accuracy of 1 part per million is readily obtainable with only a comparatively small number of finite elements.\nBy symmetry, the corresponding formula for the z-component of the magnetic field is gotten by interchanging AMX and CMZ in the formula for the x-component of the maqnetic field, and then changing the algebraic sign of the result.\nGeneric Triclinic Element\nFigure (1) illustrates an example of the most general finite element for which I have derived the analytic magnetic field formulas.\nSoecial Case: A Monoclinic Element\nFigure (2) illustrates an especially useful special\nY\nX\nERROR IN B, QENERATED BY A CIRCULAR COIL Approxlmated by an N-Blded PolyQOn\nIn the Mldplane 01 the Coll (1 0 ) 100\n10 Radial Coordinate Of Field Points (Meters) ,4\n.s 0 1 - 0 1 B . I - 1.0\nlo\" I . , - 0.0\n& 0 I - 1.6\n* r - 1 0 x ,-I1 0.1\nMI0.N A w l . 0- 1 -.-\n0.01 C q .\n10 20 40 BO I20 Wi? Number of Sides of Po/ygofl WUI\nERROR IN 8, QENERATED BY A CIRCULAR COIL Approxlmated by an N-Slded Polyoon In the Endplan. 01 the Coll (2 - tl.0)\n100 1 I\n10 20 40 BO Number of Sides of Polygon\nRadial Coordinate Of Field Points hlatersl . I-0.0 0 , - o s . 1-10 0 , - I 6 I I - 2 0 x r - 2 6\nMI0.N A d W k G-. I nbrm.md\nCorOr I20\nw\n/ '\nz Figs. 3a and 3b - The error in calculating the magnetic\nFig. 2 - A monoclinic element, in which the truncating field generated by a uniform-current-density circular plane is rotated around only one of the two permissible coil approximated by an N-sided polygonal coil, as a axes (z - axis as shown here). function of the number of sides of the polygon.",
+            "1546\nERROR IN B, GENERATED BY A CIRCULAR COIL Approximated by a 36-Slded Polygon\n6-\n4 -\n3- -c s 2- P\n-31 -4 I\n0 0.6\nc j - Coil Boundaries\nAxial Coordinate of Field Points (Meters)\n2-0.0\n0 z - 0.6\n8 2 -1.0 0 z - 1.6\nA 2 - 2 . 0\nM m a A for G m m n i I r r * ~ Y . n d c olopw\nI I 7.6 2 2.6\nRedid Coordinate of Field Points (Meters) 1yo2/85\nERROR IN B, T B,(O,O) GENERATED BY A CIRCULAR COIL Approxlmated by a 36-Slded Polygon\n1.5 - Coil Boundaries -\n0.5 1 1.5 2 2.5\nRadial Coordinate of Field Points (Meters)\nAxial Coordinate of Field Points (Meters)\n2 - 0.0 0 z - 0.6\n2 - 1.0 0 2 - 1.5 A 2 - 2.0\nM a p e 1 A nalyrir for G overnrnsnt. I nduruy. and c 011ege.\nllIOU86"
+        ]
+    },
+    {
+        "image_filename": "designv11_60_0001235_b978-0-08-099425-3.00005-4-Figure5.55-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0001235_b978-0-08-099425-3.00005-4-Figure5.55-1.png",
+        "caption": "FIGURE 5.55",
+        "texts": [
+            " suitable for very small obstacles that the tyre might completely envelop. This is clarified by Davis (1974) where it is stated that the wave length of surface variations in the path of the tire should be at least three times the length of the tyre to ground contact patch. The other and most basic limitation of this type of model is that the simulation is restricted to straight-line motion and would only consider the vertical and longitudinal forces being generated by the terrain profile. An example of a radial spring tyre model is shown in Figure 5.55. The work carried out by Kisielewicz and Ando (1992) describes how two different programs have been interfaced to carry out a vehicle simulation where the interaction between the tyre and the road surface has been calculated using an advanced nonlinear finite element analysis program. More recently, advances in computational budget have allowed the development of a number of tyre models that begin to approach a causal description of the tyre structure and its interaction with terrain. A number of high quality terraininteraction models are now available such as FTire and CDTire"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0000558_asjc.1349-Figure1-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0000558_asjc.1349-Figure1-1.png",
+        "caption": "Fig. 1. Example of multi-agent system topology.",
+        "texts": [
+            " Then, for any initial state x(1), the kf th time averaging variation \ud835\udeff ( kf ) satisfies P (\u2016\u2016\u2016\ud835\udeff (kf )\u2016\u2016\u2016 \u2264 \ud835\udefc\u2016\ud835\udeff(1)\u2016 + \ud835\udefd ) \u2265 1 \u2212 \ud835\udefe, (27) where P is the probability measure on noise sequence. Theorem 8. For any initial state x(1), the time average ?\u0302?(k) for \ud835\udf07(k) satisfies E[?\u0302?(k)] = \ud835\udf07(1), Var[?\u0302?(k)] \u2264 v2(k \u2212 1)(2k \u2212 1) 6k(\ud835\udf06(b)N )2 . In this section, we show two numerical examples and illustrate the statements in theorems. We first consider a multi-agent system with a directed graph, which is shown in Fig. 1(a). The graph Laplacian of the network structure of the system is defined by L = \u23a1\u23a2\u23a2\u23a2\u23a2\u23a2\u23a2\u23a3 1 0 \u22121 0 0 0 \u22121 1 0 0 0 0 0 \u22121 2 0 0 \u22121 0 0 \u22121 1 0 0 0 0 0 \u22121 1 0 0 0 0 0 \u22121 1 \u23a4\u23a5\u23a5\u23a5\u23a5\u23a5\u23a5\u23a6 , \u00a9 2016 Chinese Automatic Control Society and John Wiley & Sons Australia, Ltd where the largest singular value ?\u0304? of L is 2.814 and the second smallest eigenvalue \ud835\udf062 of (L + L\u22a4)\u22152 is 0.3820. The measurement noise wij(k) was supposed to be generated according to the standard normal distribution, that is, E[wij] = 0 and Var[wij] = v2 = 1",
+            "8, this result does not invalidate the statement of Theorem 1. Meanwhile, the empirical mean of ?\u0302?(16,450) was 4.990 and the sample variance of ?\u0302?(16,450) was 12.61. Furthermore, the convergence speed of \ud835\udeff(k) between the noisy case and the case without noise is almost same in Fig. 3. The empirical mean is close to \ud835\udf07(1) = 5 and the sample variance is less than its theoretical upper bound of 12.76. That is, the numerical results were consistent with Theorem 2. We next consider a multi-agent system with an undirected graph, which is shown in Fig. 1(b). The graph Laplacian of this graph is given by \u00a9 2016 Chinese Automatic Control Society and John Wiley & Sons Australia, Ltd L = \u23a1\u23a2\u23a2\u23a2\u23a2\u23a2\u23a2\u23a3 2 \u22121 \u22121 0 0 0 \u22121 2 \u22121 0 0 0 \u22121 \u22121 4 \u22121 0 \u22121 0 0 \u22121 2 \u22121 0 0 0 0 \u22121 2 \u22121 0 0 \u22121 0 \u22121 2 \u23a4\u23a5\u23a5\u23a5\u23a5\u23a5\u23a5\u23a6 . The largest eigenvalue is \ud835\udf066 = 5.2361 and the second smallest eigenvalue is \ud835\udf062 = 0.7639. All the other settings were the same as in the directed graph case. Nevertheless, we could set kf = 1,029, based on Theorem 3, in this case. Then, we executed the consensus algorithm of (1) and (2) 10,000 times"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0003646_s0997-7538(01)01203-7-Figure1-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0003646_s0997-7538(01)01203-7-Figure1-1.png",
+        "caption": "Fig. 1.",
+        "texts": [
+            " The described system moves in the vertical plane, while the bar A slides without friction along the vertical guide. Show that the described system satisfies the conditions for the existence of a generalized cycloid at the brachistocronic motion, determine that cycloid and determine the time of the brachistocronic motion. At the moment t0 = 0, when the system was at rest, it holds that y0 = 0, x0 = a 2 , \u03b80 = \u03c0 3 , and at the moment of the and of the brachistocronic motion, t = t1, y1 = a, x1 = 3a 2 , \u03b81 = \u03c0 6 . Solution. The values x, \u03b8 and y (see Fig. 1) represent Lagrange\u2019s coordinates of the system q1, q2, q3, respectively. The kinetic energy of the material system is T = 3 2 m0y\u0307 2 +m0x\u0307 2 + 2 3 m0a 2 \u03b8\u03072 +m0a cos \u03b8x\u0307\u03b8\u0307 \u2212m0a sin \u03b8y\u0307\u03b8\u0307 (p1) and the position of the centre of inertia of the material system (see (5)) is determined by the relation yC = \u03d5( q1, q2, q3) = y + a 3 cos \u03b8 \u2212 a 6 , (p2) where from \u2202\u03d5 \u2202q1 = \u2202\u03d5 \u2202x = 0, \u2202\u03d5 \u2202q2 = \u2202\u03d5 \u2202\u03b8 = \u2212a 3 sin \u03b8, \u2202\u03d5 \u2202q3 = \u2202\u03d5 \u2202y = 1. According to (p1) it follows that (see (2) and (3)) a11 = 3m0, a12 =m0a cos \u03b8; a13 = 0, a22 = 2m0, a23 = \u2212m0a sin\u03b8; a33 = 3m0, (p3) so that the following equalities are obtained: \u2202aij \u2202q3 = \u2202aij \u2202y \u2261 0 (i, j = 1,2,3)"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0003840_acc.2000.879454-Figure4-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0003840_acc.2000.879454-Figure4-1.png",
+        "caption": "Fig. 4 Feedback Control System of Traction CVT",
+        "texts": [
+            " The ratio change is realized by varying the tilt angle 6. of the roller axis with respect to the Z-axis. The swing of the roller is indirectly realised by slightly offsetting the roller axis in the direction of X-axis. The amount of offset, which is controlled by computer, creates a side slip at the diskroller contact, and this side slip induces a side slip force that turns the roller about the swing center to change the ratio. 4.2 RATIO CONTROL SYSTEM The feedback control system for the traction drive CVT is illustrated in Fig. 4. The roller is supported on the piston shaft, which moves axially and rotates about its axis with the roller. The traction forces are dynamically balanced with the piston force created by the hydraulic pressure difference. In this control system, the rotation of the piston shaft, i.e., the roller swing motion, is indirectly induced by the side slip forces at the points of contact on the input and output sides of the roller. During a ratio change, the onboard computer signals the step motor to create displacement xi , which is picked up by the spoon of the servo valve through the lever",
+            " The displacement of the roller along the piston shaft, x,,, offsets the roller axis with respect to the axes of the input and output disks. This offset induces side slip forces at the contact points on both sides of the roller, as formulated in Section 3. The moment made by the side slip forces swings the roller about the piston shaft. The axial displacement x,, and the angle 4 of the roller swing are then both transferred by a cam at the end of one of the piston shaft and feedbacked to the servo valve spoon through the lever. 4.3 MODELLING OF THE CONTROL SYSTEM The system shown in Fig. 4 uses one servo valve to control the four pistons since the- four rollers are mechanically synchronised. The feedback on the roller axis offset, i.e., the piston displacement Q, is realised by a lever, and the feedback on the roller tilt angle @ is realised electronically through a sensor. 4.3.1 FEEDBACK The displacements, x, xi, and xo, of the valve spoon, the push rod and the piston respectively, are related by the following equation, C k,,,,,x, = kixi - k,x, x , =-xi -- c + d c + d 10) d where, c and d are lever constants, k,,, is a factor dependent of cam profile"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0000577_978-3-319-33714-2_5-Figure3-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0000577_978-3-319-33714-2_5-Figure3-1.png",
+        "caption": "Fig. 3 Cylindrical joint j",
+        "texts": [
+            " Axes of the coordinate systems UjVjWj and XkYkZk, attached to the ends of this binary link, are chosen as follows: the Wj and Zk axes are located along the axes of rotation and translation of the cylindrical joints j and k; the origins Oj and Ok of the coordinate systems UjVjWj and XkYkZk are located in points of intersection of the Wj and Zk axes with the common perpendicular tjk between these axes; the Uj and Xk axes are located along the common perpendicular tjk; the Vj and Yk axes are completed the right-hand Cartesian coordinate systems UjVjWj and XkYkZk. At such choice of the coordinate systems UjVjWj and XkYkZk nonzero parameters of the matrix Tjk are ajk and \u03b1jk. Then from the matrix (1) we obtain a matrix of the binary link jk of type CC GCC jk = 1 0 0 0 ajk 1 0 0 0 0 cos \u03b1jk \u2212 sin \u03b1jk 0 0 sin \u03b1jk cos \u03b1jk 2 664 3 775, \u00f02\u00de where parameters ajk and \u03b1jk are constant, and they characterize the geometry of the binary link jk of type CC. Nonzero parameters of the cylindrical joint j shown in Fig. 3 are \u03b8j and sj. Then from the matrix (1) we obtain a matrix of the cylindrical joint j where sj\u2014a distance from the Xj axis to the Uj axis measured along the directions of the Zj and Wj axes; \u03b8j\u2014an angle between the positive directions of the Xj and Uj axes measured counterclockwise about the positive directions of the Zj and Wj axes. Parameters sj and \u03b8j are variable, and they characterize relative translation and rotation motions of the j-th cylindrical joint elements. Choosing the coordinate systems UVW and XYZ, as shown in Figs"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0001570_pime_proc_1944_151_047_02-Figure1-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0001570_pime_proc_1944_151_047_02-Figure1-1.png",
+        "caption": "Fig. 1. Sectional Drawing of Complete Machine",
+        "texts": [
+            " To assess this heat balance, therefore, it requires on the one side a knowledge of the power losses within the box, and on the other side a knowledge of the ability of the box to transfer heat by convection radiation and conduction. (It is unnecessary to separate these three forms of heat transfer and they will be collectively referred to as \u201cheat dissipating capacity\u201d.) at University of Birmingham on June 8, 2016pme.sagepub.comDownloaded from THE THERMAL R A T I N G O F WORM GEARBOXES 327 APPARATUS USED The test bed shown in Fig. 1 was used for obtaining many of the observations on which the results are based. Although not primarily designed for the present purpose, it will be seen from the description which follows that it lends itself to the type of test required. The gearing arrangement comprises a closed circuit train of two pairs of worm gears and a chain drive. The worm gear A is a standard unit with wheel shaft and worm shaft mounted on dualpurpose ball bearings; one extension of the worm shaft is coupled through a flexible coupling to a motor, and the other end is connected by chain drive to the worm shaft extension of the worm gear B"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0003124_70.88068-Figure5-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0003124_70.88068-Figure5-1.png",
+        "caption": "Fig. 5 . Expected contact configuration. (a) Relative orientation of two coordinate frames represented by Euler angles. @) Symmetric cylindrical object and flat obstacle [i) line contact, ii) point contact]. (c) Object represented by polyhedral surfaces [i) surface contact, ii) line contact, iii) point contact].",
+        "texts": [
+            " where c = cos, s = sin, or and the expected goal location and orientation with respect to the base frame can be given as For objects modeled by quadratic surfaces with an axis of symmetry, it is only necessary to consider the relative orientation of the axis of symmetry aCb and the surface normal of the goal location a,b. If the object is represented by polyhedral surfaces, it would be necessary to consider the axes of the compliance and goal frames (i.e., neb, o c b , a c b and ngb, O g b , a g b ) . Fig. 5(b) and (c) demonstrates this procedure for predicting the likely contact configuration between two objects. It was shown in previous sections how the expected contact configuration is determined. Using this information together with the measured force vector, it is possible to fully define the point or line of contact with respect to the compliance frame of reference. The force vector refers to the three forces and three torques obtained using any robot wrist force sensing device such as those provided by Lord and Barry Wright"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0000559_978-3-319-44156-6_24-Figure1-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0000559_978-3-319-44156-6_24-Figure1-1.png",
+        "caption": "Fig. 1 Construction process of a complex eight-link mechanism",
+        "texts": [
+            " They can modify any input data in real time changing values in a numeric way or using the computer mouse to manipulate links and vectors while the mechanism is moving and showing the results. This powerful tool does not only show the results in a numeric way by means of tables and diagrams but also in a visual way with scalable vectors and curves. The drawing scale can be changed with the computer mouse. In this work, those mechanisms with more than four links are called complex mechanisms. These can be built with WinMecC with a base mechanism, a series of dyads and the option to add points to any link. An example of how to assemble a complex mechanism using WinMecC is shown in Fig. 1: (a) First, a base mechanism of four links is inserted. This is integrated by an input link articulated with a dyad (see Fig. 1a). (b) Secondly, point C is added to link 2 (Fig. 1b). (c) The next step consists of adding a dyad linked to points B and C (Fig. 1c). (d) Later point E is added to link 6 (Fig. 1d). (e) Finally, a new dyad is joined to point E and a fixed vertical path whose position is defined by point F8 (Fig. 1e). The Cartesian coordinates of fixed points O2, O4, O8 and F8, the length of links 2, 3, 4, 5, 6 and 7, the path angle of slide 8 and the position of points E and C with respect to their links can be defined with the mouse or by entering values with the keyboard. The tools needed to assemble the mechanism in Fig. 1 are part of a wide set of options that are available in the WinMecC lateral toolbars shown in Fig. 2. Figure 2a shows the available base mechanisms. Once one of them has been chosen, it allows the user to enter different structures that can be added to the mechanism as shown in Fig. 2b, c. One of the main characteristics of WinMecC is the huge quantity and wide range of results obtained for the kinematic and dynamic analysis of a mechanism. All of these can be shown as numeric results and/or graphic visualization"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0000346_15325008.2016.1148080-Figure7-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0000346_15325008.2016.1148080-Figure7-1.png",
+        "caption": "FIGURE 7. Pictures of the test rig: (a) overview of the experimental platforms and (b) enlarged view of the motor shaft and the backup bearing.",
+        "texts": [
+            " The radial displacements of the rotor were measured by eddy-current type transducers and then regulated by PID controllers to obtain the levitation-force references that were input to the analog-digital conversion (ADC) of TMS320C2812. The rotor angle position was detected by three opto-couplers, which were installed on the motor shell with a eight-pole toothed disc mounted on the rotor shaft. The three output signals were input to the TMS320C2812 DSP controller to calculate the rotor speed. Accordingly, the proposed control algorithm was programmed and implemented by the controller. It is worth pointing out that a backup bearing was installed to avoid the collision between the rotor and stator poles. Figure 7 shows the pictures of the test rig. In Figure 7(a), the DSP controller processes the control algorithm to output pulsewidth modulated (PWM) signals to the power converters. The asymmetric half-bridge type power converters energize the motor windings to produce the required winding currents. As a result, the required average torque and average levitation forces are produced to rotate and levitate the BSRM. In Figure 7(b), the backup bearing is installed to protect the shaft, and there exists a gap between the shaft and the backup bearing. Figures 8 and 9 show the experimental results when the prototype was operated at the speed of 1000 and 3000 r/min, respectively. It can be seen that the rotor displacements at the two directions are both less than 50 \u03bcm. The four currents of phase A are different to generate unbalanced magnetic pulls. In particular, the average values of the four currents changed continuously, caused by the required F\u2217 avg; in other words, the radial displacements of the shaft force the winding currents to change continuously to provide the required levitation forces"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0002109_1.2919444-Figure3-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0002109_1.2919444-Figure3-1.png",
+        "caption": "Fig. 3 Machine coordinate system",
+        "texts": [],
+        "surrounding_texts": [
+            "V. Kin American Sykes Co., Beavercreek, OH 45432 Assoc. Mem. ASME\nComputerized Analysis of Gear Meshing Based on Coordinate Measurement Data\nA mathematical model is proposed for the reconstruction of gear tooth surface features based on the manufacturing process deviations measured with the help of a gear inspection {or coordinate measurement) machine. The model is extended to facilitate the computer simulation of meshing between the measured gear and a theoretical master gear, or between two measured gears. This would allow one to predict the transmission errors and bearing contact which are otherwise only possible to investigate using the combination of single-flank testing and a time-consuming work-in and examination procedure.\nThe developed model is applied to the determination of tooth-to-tooth trans mission errors of parallel-axis gears and can also be utilized to determine bearing contact and contact stresses and be useful in tooth surf ace fatigue life calculations.\nIntroduction TCA has been originally developed by The Gleason Works [1] as a way of analyzing the gear-tooth action for spiral bevel and hypoid gears. The properties of meshing that are analyzed include the bearing contact and the transmission errors of such gears based on the machine tool settings for spiral bevel gear generators. The techniques were later extended to other the oretical point- [2], [3] as well as theoretical line-contact [4], [5], [6] geometries. The basis for the bearing contact and trans mission error predictions were again the machine tool settings for the machines utilized in the production of the correspond ing gears.\nEfforts have also been made in the area of gear metrology and optimization of the machine tool settings based on meas urement data [6], [7]. According to this approach, the settings are changed so the resulting deviations of the gear tooth surface from the ideal are minimized. A procedure has also been worked out to determine the first n harmonics of the transmission error function and to thus deduce the properties of the corresponding noise spectrum [8], [9] for relatively simple gear geometries. It does not, however, generalize easily to more complicated cases and even for the simpler ones runs into problems if one wishes to consider the zone of contact to be dependent on the geometric deviations of each tooth.\nA method has been suggested in [7] to reconstruct the real surface of an arbitrary gear from the inspection data by fitting a second-order polynomial surface to the error surface using linear regression. The reasons for this approach were cited as relative simplicity as well as the lack of confidence in the assumption of the validity of every data point.\nContributed by the Power Transmission and Gearing Committee for publi cation in the JOURNAL OF MECHANICAL DESIGN. Manuscript received Dec. 1991; revised Feb. 1993. Associate Technical Editor: D. G. Lewicki.\nIn this paper we propose an approach of fitting a tensorproduct surface to the measurement data. This reconstructed surface is then utilized for the purposes of simulation of mesh ing between the measured gear and another gear (which can also be a measured gear or a theoretical master). While we recognize that the approach does add to the mathematical complexity of the solution, it does not asymptotically increase its computational complexity (i.e., the expense of obtaining this solution does not grow faster with the increase in the number of data points as it does for the previously mentioned method). We also feel highly confident in the ability of a modern generative CNC-controlled gear inspection machine to accurately and repeatedly measure gear surfaces (with mi cron accuracy on certified involute and lead masters). In this situation, tensor-product surfaces offer greater freedom of analysis since they accurately model local as well as global surface deviations, whereas only global trends are revealed by a regression approach.\nMeasured Gear Equations We consider two coordinate systems\u2014Sx and S\u201e,\u2014rigidly connected to the measured gear (gear 1) and the machine, respectively (Fig. 1). The origin of Sm is at the corner of the machine gage block (Figs. 2 and 3), xc[ and j>ci are coordinates of the gear axis in Sm, z0/j is the Sm offset at which the gear is mounted, and <p is the angle of rotation of the rotary table. The origin of the part coordinate system Sp is at the base of the rotary table (Fig. 4). The unusual sign convention employed in Sm is due to that of the machine (Fig. 5) and has to be complied with in order for its numerical control to function properly.\nIn order to correctly measure the gear at point r,\u201e , the probe has to be positioned at the point r,\u201e + pnm , where p is the ,o(0 \u201eo(!)\n738 / Vol. 116, SEPTEMBER 1994 Transactions of the ASME\nCopyright \u00a9 1994 by ASME Downloaded From: http://mechanicaldesign.asmedigitalcollection.asme.org/pdfaccess.ashx?url=/data/journals/jmdedb/27620/ on 01/29/2017 Terms of Use: http://www.asme.org/about-asme/terms-of-use",
+            "Fig. 1 Machine (Sm) and part (S\u201e) coordinate systems\nFig. 2 Machine gage block\nradius of the spherical probe tip. The locus of the probe po sitions in coordinate system Sm is then given by\n, . ( />). (1)\nwhere the coordinate transformation matrix M,nl from Sj to S,\u201e is expressed as\nMml =\nand L,\u201ei is the 3 x 3 left upper submatrix of M,\u201e|. The value of <p when measurements are made must be such that the gear is positioned with respect to the probe so the latter deflects approximately normal to the gear surface.\nThe procedure for the determination of <p and for the correct interpretation of probe deflections depends on (/) equations of the gear tooth surface, and (//) the exact position of the gear coordinate system S\\ relative to the part coordinate system Sp. Item (//) is a considerably more complicated matter than it may appear on the surface because for the purposes of inspection a gear is often mounted differently than it is when the same gear is used to transmit power. A common example would be the case when a gear is mounted on an external spline of a shaft when it is installed in a gearbox. During inspection, however, this gear would usually be mounted between centers or on a magnetic chuck. In either case, the Z\\ and Zp axes will not necessarily coincide, and a complicated procedure is there- ' fore necessary to determine the transformation from Sp to S\\. The procedure also has to take into account the effect of this phenomenon on probe deviations and correctly compensate for the changes in the gear position as well as those in the location of the point of contact between the gear tooth and the probe tip. A procedure properly addressing these concerns is a lengthy and complicated one and space limitations do not permit its inclusion into this paper.\nFig. 4 Part coordinate system\nf ^ l\nFig. 5 Machine axes and sign convention\nSurface E, of the ideal measured gear, as well as the devia tions Ee from this surface are given in S, as\nW- M + r l (2)\nwhere the subscript denotes the coordinate system, the super script stands for gear 1, t - for the theoretical component of the surface, and e for the error component. The error com ponent is measured in the direction of the normal to L, which is acceptable due to small changes of the normal as a result of surface deviations. Measurement points on surface of the measured gear are then expressed as:\n$ 1 (\u00ab/.! Ai) = ri (\" (\"u AA) + * ( \" / A i ) ni(\u00b0 (\"i.i At) (3) where the unit normal to the theoretical surface is defined as:\nN<\" = - dr\"> dr_ du, X dQ (0 (4a)\nJournal of Mechanical Design SEPTEMBER 1994, Vol. 116 / 739\nDownloaded From: http://mechanicaldesign.asmedigitalcollection.asme.org/pdfaccess.ashx?url=/data/journals/jmdedb/27620/ on 01/29/2017 Terms of Use: http://www.asme.org/about-asme/terms-of-use",
+            ",('). Ne>\nIN C)| (46)\nHere and throughout the paper we employ the notation in which the first subscript designates the position in an array and the second - the coordinate system in the case of vectors and the gear number (1 or 2) in the case of curvilinear coor dinates.\nEquations (3) represent only the topologically rectangular in x n array of points r'j ' (\u00ab,., 0,-,]), not the entire surface, since e(u,, 0,) is so far only defined for the members of this array.\nA number of methods are known that would permit the surface to be reconstructed from this array [10], [11], [12]. We reject Bezier's surfaces [10] because the reconstructed surface in this case is not guaranteed to (and probably will not) pass through the points of measurement and will then represent the trends rather than the local phenomena of the surface. We would also be reluctant to use B-Spline surfaces [12], but for quite an opposite reason - a local phenomenon observed for a 4 x 4 subarray of the measurement points will have no effect on the reconstructed surface outside of the bounds imposed by this subarray. This is not a natural property of smooth well-finished gear tooth surfaces but may be applicable in certain cases, for instance for finish-hobbed gears. For a formal description of these surfaces and further references, the reader is referred to [13].\nApart from the already rejected regression techniques we are left with parametric spline surfaces [11], [13]. The order of the spline is in many cases suggested by the theorem of the sufficient conditions of envelope existence [14], which requires the generating surface to be of Class C2. Parametric splines of order 3 have this property and the order is not excessive for an efficient computer implementation.\nPoint \u00a3(1) (\u00ab!, 6{) on the cubic parametric spline surface of deviations is represented by:\ne(1)(wi', fli') = U1 'CHuQ1K /.,C r*1 ' (5)\nwhere the indices i and j are chosen so that\nM , - l < M 1 < M / + l i i\nThe normalized curvilinear coordinates are defined in terms of panel sizes as:\nhi\nh\n(6a)\n(6b)\nand the panel sizes are:\nhi \u2014 W/+i,i \u2014 W/,i\nhi = 0/+u ~0/,i The matrices in (5) are expressed as [13]\nV{=[lu{u?u[3]\n1\\\n* i ' 1 ~ \\a>2 1\n\u00bbi'V\nH/.i =\nK;.i =\n/ I\n1 \u00b0 \\o\n/ I\n1 \u00b0 \\o\n0 1 0 0\n0 1 0 0\n0 0\nhi 0\n0 0\nhi 0\n0 0 0\nhi\n0 0 0\nhi\nThe first two columns of Q, are\n/ e(,,(u;,i,fl!/.i) \u00ab< 1 )(\u00ab/, i ,0 / + l l , ) \\\nQV: ew(ui+l,u0jA) e l , , (\u00ab /+ i . i , 0 ;+ i . i ) .('V\n^WM AD, \u00abi(w;,i!0/+i,i)\n\\4 ' , )(\u00ab/+ i,i.0/,i) e\u00abi'(\u00ab/+ i,i.0y+1,1)/\nand the second two are\nQ.(r) = 41,)(\u00ab/.i,e/.i) 4>,-+1,1,0,-,,) eit1BJ(.Ul,ll9j,l)\n4>/,i,0/+i,i) 4>;+i,i,0/+i,i) e^s,(w;,iA'+i,i)\n{^u^^Ui+l.ltdj.l) \u20ac\u201e1fl,(M/+ 1,1,0/+ l,l)/\nand the matrix itself is thus expressed as\nSurface Ei is thus defined by the equation:\nrSl)(W\u201e01) = r,(')(\u00ab,,01) + \u00a3(1)(tt1,0i)\ntogether with (5)-(6). The unit normal to Ei is defined by:\nni'> =\nN^-\nIN^I\ndu, ddx\nwhere\ndux du\\\n(0 - + U1'\u201eCH,-,1Q,K;,1C 7'*,'\nU,'\u201e = \u2014 [ 0 12M/3U,'2] hi\n9r<\" ft-'\" 90i ~ 90, + U,'CH;,,Q,Ky-,,C 7'*,'e\n* i ' \u00bb\nh,i\no l\n20/ \\30,'2/\nWe note here that the computational effort involved in carrying out (5) [and thus computing (7) and (8)] is quite low since the orders of all the matrices are constant and the only operation where the computational expense depends on the number of data points is the determination of indexes / and j . In the case of unequally-spaced data, the numbers of operations to de termine these and on the order of log m and log n, respectively. They are, however, constant for equally-spaced data.\nThe determination of the elements of Q, is another story altogether because all of these can be computed before (4) has\n740 / Vol. 116, SEPTEMBER 1994 Transactions of the ASME\nDownloaded From: http://mechanicaldesign.asmedigitalcollection.asme.org/pdfaccess.ashx?url=/data/journals/jmdedb/27620/ on 01/29/2017 Terms of Use: http://www.asme.org/about-asme/terms-of-use"
+        ]
+    },
+    {
+        "image_filename": "designv11_60_0000289_1350650115607553-Figure3-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0000289_1350650115607553-Figure3-1.png",
+        "caption": "Figure 3. Geometry of the coil.",
+        "texts": [
+            " The required gap area Ag is first calculated using values of saturation capacity for different values of pole pairs as shown in Table 1. Fmax \u00bc sB 2 satAgn 2 0 \u00f01\u00de The gap area in terms of the dimensions of the pole as Ag \u00bc wp:lp \u00f02\u00de Step 2. The dimensions of the journal may be determined using suitable aspect ratio g. The dimensions of the journal may be determined as rj \u00bc \u00f0rr \u00fe fswp\u00de; Ag \u00bc 2 wp\u00f0rr \u00fe fswp\u00de \u00f03\u00de Step 3. Generally linear range of actuator is set equal to dynamic capacity. Hence, selection of bias point b depends on stator configuration Fdy Fmax \u00bc 2 s \u00f04\u00de Step 4. From the geometry of coil shown in Figure 3, one may write the following expression tan \u00bc tc \u00fe wp 2 rp \u00f05\u00de The pole tip radius and average coil length per turn may be expressed as rp \u00bc rj \u00fe lg and lc \u00bc 2 lp \u00fe wp \u00fe 2tc \u00f06\u00de The thickness of the coil may now be expressed as tc \u00bc rp tan np wp 2 \u00f07\u00de Available coil area may be expressed as Ac \u00bc Bsatlg J 0 ffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi 1\u00fe Fdy s Fmax 2 s \u00f08\u00de Referring to coil geometry shown in Figure 3, the inner radius of the stator may be expressed as rc \u00bc ffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi \u00f0h\u00fe rp\u00de 2 \u00fe wp 2 \u00fe tc 2r \u00f09\u00de Referring Figure 2, the pole length, stator axial length, and the volume of electromagnet are Lp \u00bc rc rp; ls \u00bc lp \u00fe 2tc; Ve \u00bc Ae Lp \u00f010\u00de Referring Figure 3, Ac \u00bc tc rc rp ; rs \u00bc rc \u00fe fswp; Vc \u00bc Aclc \u00f011\u00de Step 5. Amplifier capacity may be expressed in terms of dF dt max VmaxImax \u00bc dF dt max 4lg np \u00bc 5:49 106 4lg np \u00f012\u00de After selecting peak current value, diameter of coil wire may be calculated as dw \u00bc ffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi 4Imax= Jmax p \u00f013\u00de The number of turns Nt may be calculated as NtImax \u00bc Bsatlg 0 \u00f014\u00de The total loss may be expressed as PT \u00bc Pcu \u00fe Pe \u00fe Ph \u00fe Pw \u00f015\u00de Stator copper resistance loss The total copper loss for varying poles may be expressed as Pcu \u00bc I2maxRcu \u00f016\u00de at UNIV OF PITTSBURGH on October 10, 2015pij"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0003804_iros.2001.973345-Figure2-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0003804_iros.2001.973345-Figure2-1.png",
+        "caption": "Figure 2: The three different types of contact between-&e palm and the object, shown with the friction cones used to represent the contact forces.",
+        "texts": [
+            " The main assumption that we make is that the object is in a configuration that is statically stable, i.e. if there is no palm motion then the object will remain at rest. We further assume that the palm is required for this static stability. The consequence of this assumption is that a manipulation plan may be stopped at any point. We consider three types of contact between the palm and the object: between the palm face and an object edge; between the palm edge and an object face; and between the palm face and an object face. These are illustrated for a box in Figure 2. Although our examples show only the first type of contact, our analysis only needs to know the contact point locations and normals, so any of these contact types will do. 2.2 Kinematic analysis of fixed contacts We first perform a kinematic analysis to determine which motions (and associated contact modes) are consistent with the fixed contacts such as with the floor and wall in our example task. We do this by applying Reuleaux\u2019s method for the given object configuration. For each contact, we consider every point in the plane as a possible velocity center for a motion and label the point with the signs of rotation that do not violate the contact constraint"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0002033_(sici)1097-4563(199710)14:10<729::aid-rob3>3.0.co;2-w-Figure5-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0002033_(sici)1097-4563(199710)14:10<729::aid-rob3>3.0.co;2-w-Figure5-1.png",
+        "caption": "Figure 5. Position error of pseudo center with respect to So9 .",
+        "texts": [
+            ")velocities of the mobile robot according to the posiThen, the relation between the velocity at the contacttion and orientation errors with respect to the desired point and the linear and rotational velocities at thetrajectory expressed in the coordinate frame fixed on pseudo center is given approximately by Eq. (2). Con-the robot. The control rule also guarantees that the sequently, if we set the velocity at the contact pointactual trajectory of the robot converges uniformly Ovc asasymptotically to the desired trajectory. Let SO9 denote the coordinate frame with its origin Ovc 5 B21[vo9 , u \u02d9 ]T. (11)at the pseudo center and its X axis parallel to the line that joins the contact point and the center of friction we can expect to make the object follow the desired(see Figure 5). Then, we consider the position error trajectory. The control variable, that is, the arm tipof the pseudo center expressed in SO9 with respect to velocity Uvc realizing the above Ovc is given bythe desired trajectory. The error is denoted as O9pe 5 [O9xe , O9ye]T. Denoting the actual position of the pseudo Uvc 5 URO Ovc . (12)center expressed in SU as Upo9, the desired position as Upo9d , and the rotational matrix from SO9 to SU as URO9 , the error O9pe is given by 4.2. Limitation of Arm Tip Velocity O9pe 5 URT O9(Upo9d 2 Upo9)"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0002006_s0003-2670(98)00705-3-Figure7-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0002006_s0003-2670(98)00705-3-Figure7-1.png",
+        "caption": "Fig. 7. The scheme of the proposed flow-injection analytical system enabling quantitative determination of the analyte concentration by standard addition method or by creation of a calibration curve. All the components are the same as in Fig. 1",
+        "texts": [
+            " By applying triangle (a) or half-triangle (c) injection mode, the maximal detector signal for the added standard cannot be obtained, and for that reason, the slope of the obtained curves (the increase of the detector signal caused by the increase of the rotation speed of the pump delivering standard solution) cannot be properly evaluated. One of the most important bene\u00aets of the proposed sample-injection/sample-propelling (I/P) system is its applicability for on-line sample dilution in combination with the on-line creation of the calibration curve or quantitative determination of the analyte by standard addition method. For both the quanti\u00aecation procedures of on-line quanti\u00aecation one additional computer-programmable peristaltic pump should be included in the analytical system as schematically shown in Fig. 7. The mixing point T1 was a T-piece. Experimentally, it was veri\u00aeed that no contamination of a sample occurred by the standard solution if the peristaltic pump PP3 (Fig. 7) was switched off. The operation principles of all the three peristaltic pumps during sample dilution/quantitative determination by standard addition or by creating a calibration curve are presented in Fig. 8. In Fig. 9 an example of a sample dilution and quanti\u00aecation based on the creation of a calibration curve is presented. The proposed integrated sample-injection/sampledilution/sample-propelling system for \u00afow injection analysis is rather robust and very \u00afexible. It enables the injection of the original sample, its on-line dilution and formation of a calibration curve"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0000860_978-3-319-44735-3_1-Figure1.37-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0000860_978-3-319-44735-3_1-Figure1.37-1.png",
+        "caption": "Fig. 1.37 Moving carrying a weight",
+        "texts": [
+            " On the other hand, when the altitude is zero and the vertical speed is negative it means the hexacopter is falling down. In this condition, the controller must to set output to a value higher than the ZERO just to make the hexacopter stop the falling. The work has been validated through a case study by means of simulation. For that, a model of a hexacopter has been created in the V-REP robotics simulation environment. A free payload has been attached to the hexacopter forming a pendulum as depicted in Fig. 1.37. In the simulated environment, the hexacopter weighs 980g (mass = 0.1) and the payload weighs 49g (mass = 0.005). The hexacopter model used is one that is already available on V-REP. Such a payload weight was defined in 5% of the hexacopter weight due to limitations on the rotors model that cannot provide enough thrust to allow the hexacopter takeoff. For simulation the V-REP has been configured with \u201cDynamic engine\u201d as \u201cBullet\u201d, the \u201cDynamics settings\u201d as \u201cVerry accurate\u201d and \u201cSimulation time step\u201d as \u201cdt = 10ms\u201d"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0002551_980827-Figure2-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0002551_980827-Figure2-1.png",
+        "caption": "Figure 2. the clutch\u2019s scheme construction",
+        "texts": [
+            " The contact force becomes to . Acceleration: the outer rim continues rotation until its angular velocity is the same as that of planet wheel. The process is known as dynamic locking. Working: the outer rollers, planet wheel are connected as a part. At this time, no relative slippery, transmitting torques and movement. Release: the contact force changes into 0, the process symmetry with that of locking process. THE ANALYSIS OF LOAD AND TIME IN LOCKING PROCESS \u2013 The overrunning clutch\u2019s scheme construction is shown in Fig.2(a). It contains mainly: outer rim, rollers, inter rim (planet wheel) and springs. The outer rim is rigidly connected with the output rocker of front of mechanism transmission, and the inter rim is rigidly connected with output shaft. It working principle can be described as follows, the rollers are wedged in between inter rim and outer rim, when the rotor rotates in a given direction to drive the rocker into one way rotary output. If the driver rotates in the opposite direction or if the driven member attains a faster angular velocity, the inter rim and outer rim become disengaged. According to Eq. (1), we can obtain the relative angular velocity between inter rim and the outer rim, (Eq. 2) from Fig.2(b), we know (Eq. 3) according to initial condition: can be given by Eq. (3). so, Eq.(3) can be written as: (Eq. 4) According to Fig. 2(b), when outer rotates angular relative to the planet wheel, the contact points between roller and outer rim and between roller and planet wheel change a, b into a' , b' respectively. Thinking of the process as absolute roll. thus aa'=bb'=R . According to quadrilateral aa' bb', the compressive deformation is: (Eq. 5) substitution of Eq. (4) into (5) (Eq. 6) According to Bochmann experimental conclusion, we know line contact deformation is connected with linear power of load , so, we obtain the formula form Kunert & Lundberg "
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0001747_s0167-8922(08)70478-4-Figure1-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0001747_s0167-8922(08)70478-4-Figure1-1.png",
+        "caption": "Figure 1. Blow-up picture of the mechanical parts for the experimental set-up, from Lindqvist et.al. [49].",
+        "texts": [
+            "2 Split-Hopkinson bar test rig Below a new way of measuring the dilatation, and the density\u2019s dependence on pressure at transient loading is presented. A modified split-Hopkinson pressure bar set-up is used to determine the dilatation (relative volume change)-pressure relation of lubricants. The set-up makes it possible to test oils and greases under conditions similar to those found in a real ehl contact; loading-unloading times of 100 and 300 ps respectively and pressures up to 1.9 GPa. A schematic picture of the experimental set-up can be seen in Fig. 1. Details can be found in Lindqvist et. al. [49]. The tested lubricant is confined in a cylindrical hole (a8 mm) in a pre-stressed container made of cemented carbide and steel. It is pressurized by two axially movable pistons that are in contact with the two rods. The radial clearance between piston and hole is between 1 and 2 pm thus minimizing leakage but still giving a low friction force. One of the rods is axially impacted by a projectile generating a compressive elastic wave. Strain gauges are used to measure the strains in the rods as a function of time at two positions on each rod"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0001000_1.4035601-Figure4-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0001000_1.4035601-Figure4-1.png",
+        "caption": "Fig. 4 Herringone grooves",
+        "texts": [
+            " However, the focus of these investigations is on a radially mounted seal. Advantages for this orientation include being more applicable to turbine split line designs, having to cope with less movement perpendicular to the flow direction, and being easier to scale at various diameters. There are many possible groove shapes, several of which originate from bearing design. These include wedges (flat and tilted) (Fig. 1), circumferential pockets and Rayleigh steps (Fig. 2), inclined grooves (Fig. 3), and herringbone grooves (Fig. 4). Wedges are reported by Dhagat et al. [32] and the tilted variety is recorded by Galimutti et al. [33]. Pockets are also shown by Dhagat et al. [32] and Rayleigh steps are described by Cheng and Wilcock [34]. Herringbone grooves are assessed by several authors although there is no general agreement on whether the center land region should be included or not. The center land is omitted in Dhagat et al. [32] and Liu et al. [35] but included in Cheng and Wilcock [34] and Proctor and Delgado [29]. This list is by no means exhaustive; there is a vast array of literature encompassing these groove types"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0000938_iemcon.2016.7746308-Figure1-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0000938_iemcon.2016.7746308-Figure1-1.png",
+        "caption": "Fig. 1. Two main types of quadrocopter configuration.",
+        "texts": [
+            " Thus the adaptive controller can be wrapped around an already-stable closed-loop system [4], adding performance and robustness in the face of plant variations. It is also easy to predict the time-delay margin using standard linear systems analysis, and this margin has been confirmed experimentally. Finally, output-feedback L1 is relatively easy to implement in practice as will be seen in the experimental sections [5]. II. MODELING OF QUADCOPTER DYNAMIC A quadcopter is an under actuated aircraft with fixed pitch angle four rotors as shown in Figure 1. Modeling a vehicle such as a quadcopter is not an easy task because of its complex structure. The aim is to develop a model of the vehicle as realistically as possible. A typical quadcopter have four rotors with fixed angles and they make quadcopter has four input forces, which are basically the thrust provided by each propellers as shown in Figure 1. There are two possible configurations for most of quadcopter designs \u201c+\u201d and \u201c\u00d7\u201d. An X-configuration quadcopter is considered to be more stable compared to + configuration, which is a more acrobatic configuration. Propellers 1 and 3 rotates counter clockwise (CW) , 2 and 4 rotates counter-counter clockwise (CCW). So that, the quadcopter can maintain forward (backward) motion by increasing (decreasing) speed of front (rear) rotors speed while decreasing (increasing) rear (front) rotor speed simultaneously, which means changing the pitch angle"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0001163_jsee.2016.06.12-Figure1-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0001163_jsee.2016.06.12-Figure1-1.png",
+        "caption": "Fig. 1 Planar engagement geometry",
+        "texts": [
+            " Also, the comparisons between partial ICG law and ANTSMC in [30] are presented. A skid-to-turn missile with cruciform control deflection surfaces is considered here. The motion of this type of missile can be decomposed into two perpendicular channels. Consequently, in each channel, the guidance and control problem can be treated in a plane. In this section, we present the full nonlinear kinematic and dynamic equations which will serve for the subsequent partial IGC law design. The standard planar engagement geometry is shown in Fig. 1. The kinematics between missile M and target T can be denoted by r and \u03bb, where r is the relative range along line-of-sight (LOS), and \u03bb is an LOS angle. The missiletarget engagement kinematics are given [29] by the following differential equations: \u23a7\u23aa\u23aa\u23aa\u23aa\u23aa\u23aa\u23aa\u23aa\u23a8 \u23aa\u23aa\u23aa\u23aa\u23aa\u23aa\u23aa\u23a9 r = Vr Vr = V 2 \u03bb r + ATr \u2212 sin(\u03bb \u2212 \u03b3M )nL \u03bb = V\u03bb r V\u03bb = V\u03bbVr r + ATr \u2212 cos(\u03bb \u2212 \u03b3M )nL (1) where V\u03bb and Vr are the relative velocity between missile and target along and orthogonal to the LOS, respectively; ATr and AT\u03bb are the projections of target acceleration along and orthogonal to the LOS, respectively; \u03b8M is the missile flight path angle; \u03b3M is the flight path angle of missile M ",
+            " It can be verified that the closed-loop system error variables \u03c30j , \u03c3i satisfy the same differential inclusions as (9), so the closed-loop system converge to the nominal case as (12) after a finite time transient process. By applying Lemma 3, the system will converge to the origin after a finite time instant. Thus, the finite time stability of the closed loop system can be established. The partial IGC law is designed in this section to realize the zero miss distance intercepting a maneuvering target. The design principle is shown in Fig. 1. From which, we can see that the whole system is divided into two loops, which are the outer loop and the inner loop. The controllers of these two loops are designed respectively. As stated in the introduction, the partial IGC law is different from the traditional separation design method. The partial IGC method takes full advantage of the inherent time-scale separation property, which not only reduces the system order, but also avoids the drawback of separation design method of guidance loop and control loop"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0000897_978-3-319-50472-8_2-Figure5-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0000897_978-3-319-50472-8_2-Figure5-1.png",
+        "caption": "Fig. 5. Case 3.4.",
+        "texts": [
+            "2: - If the occupancy rate is equal on both sides then the robot makes one hop movement to the side with closer neighboring occupied nodes (Fig. 3(a)) or any of the sides if there is a tie (Fig. 3(b)). \u2022 Case 3.3: - If the occupancy rate is more on the counter-clockwise direction but the clockwise string is nil then the robot does not make any movement (Fig. 4(a)) else it makes one hop movement to the counterclockwise direction (Fig. 4(b)). \u2022 Case 3.4: - If the occupancy rate is more on the clockwise direction then the robot makes one hop movement to the clockwise direction (Fig. 5). 2-Node Problem: The main concern which may act as a thorn on the path of gathering is the 2-node problem, where two nodes in the ring are occupied by the robots. In fact the inclusion of the property of chirality in the algorithm is just to tackle the 2-node problem otherwise N 2 visibility range alone would be enough for the gathering to complete. So we put a special condition in case 3.3 under step 3 to counter the 2-node instability. If a robots finds more occupied nodes on the counter-clockwise direction but the clockwise string is nil then the robot does not make any movement else it makes one hop movement to the counter-clockwise direction"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0003512_135065002760364831-Figure1-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0003512_135065002760364831-Figure1-1.png",
+        "caption": "Fig. 1 Journal bearing con\u00aeguration and coordinate system",
+        "texts": [],
+        "surrounding_texts": [
+            "Some experimental studies show that the properties of the lubricants of bearings can be improved when some high-polymer additives are added to mineral oils. For example, Oliver [1] found that the presence of dissolved polymer in the lubricant increases the load-carrying capacity and decreases the friction coef\u00aecient. Spikes [2] observed that base oil blended with additives can stabilize the behaviour of the lubricants in elastohydrodynamically lubricated contacts and consequently reduce friction and surface damage. Because the stress tensor in these kinds of \u00afuid (named complex \u00afuids) is antisymmetric, their accurate \u00afow behaviour cannot be predicted by the classical Newtonian theory. Currently there exist several theories to describe the \u00afow of complex \u00afuids. Although these various theories are quite different in their applicability, they all take into account the polar nature of the continuum, exhibiting the asymmetric stress tensor. In order to examine the simplest generalization of classical theory that would allow for polar effects, the so-called couple stress theory of \u00afuids was presented by Stokes [3] in 1966. This model de\u00aenes the rotation \u00aeeld in terms of the velocity \u00aeeld; i.e. the rotation vector is equated to one-half the curl of the velocity vector (i.e. vorticity). Owing to its relative mathematical simplicity, the couple stress \u00afuid model has been widely used to study the performance of various bearing systems, such as hydrostatic thrust bearings [4], rolling bearings [5], sliding bearings [6], journal bearings [7] and squeeze \u00aelms [8, 9]. However, the above models do not include the important effects, such as the thermal effects and cavitation effects in the lubricating \u00aelm. Today with the increasing trend towards higher speed, higher performance but smaller size, machinery has pushed the operating conditions of the bearings towards their l\u0300imit design\u2019. Hence, for reliable prediction of the performance of such bearings, a model which accounts for all the operating conditions as closely as possible is becoming increasingly important. In this paper, the Stokes couple stress \u00afuid model is used to investigate the effects of couple stress in lubricants. Meanwhile, the effects of heat and cavitation on the performance of a journal bearing are included."
+        ]
+    },
+    {
+        "image_filename": "designv11_60_0003578_c39860000328-Figure1-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0003578_c39860000328-Figure1-1.png",
+        "caption": "Figure 1. Cyclic voltammogram showing the first 2 potential cycles of a polymer-coated electrode in 0.1 mol dm-3 tetrabutylammonium perchlorate-acetonitrile. Scan rate: 20 mV s- 1 .",
+        "texts": [
+            " Stable polymer coats with good adhesion characteristics were prepared by evaporation of a small quantity of a solution of (2) from a platinum foil. In a typical experiment a solution (7.5 pl) containing 6 mg of (2) in CH2C12 (5 ml) was coated onto a small piece of platinum foil (ca. 4 x 4 mm). Solvent was allowed to evaporate off and the electrode left to dry. The electrode was then warmed to 90 \u201cC in an oven for about 30 min. Cyclic voltammograms were taken in acetonitrile containing tetra-n-butylammonium perchlorate (0.1 mol dm-3) as background electrolyte. Figure 1 shows a typical voltammogram obtained at a scan rate of 20 mV s-1. The initial anodic scan produces a voltammogram very different from the second and subsequent scans which are essentially identical (showing negligible desorption of the polymer in either its oxidised or reduced states) and display a sharp, approximately Gaussian peak at 1.18 V (vs. Ag/AgCl). The shape of this peak is indicative of an electrochemically reversible one-electron Br I Q f CH -CH2-3-;; I Q Br Br Br oxidation of an immobilised species and Tafel analysis of the initial portion of the curve produces slopes close to 60 mV"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0001195_9781782421955.43-Figure2-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0001195_9781782421955.43-Figure2-1.png",
+        "caption": "Fig. 2: Random position in meshing simulation",
+        "texts": [],
+        "surrounding_texts": [
+            "significantly from the theoretical line of action of an involute gear. The correct line of action for an improved load distribution calculation results from a meshing simulation of the actual tooth flanks from the foot-radius beyond the tip radius. Then it is possible to calculate the influence of different tip reliefs, tip radii on the load distribution on the teeth as well as the pressure distribution. Therewith, an accurate detection of the point of the first and last contact of gear pairs with profile modifications and cratering wear is possible.\n3. LOAD DISTRIBUTION - LOCAL PRESSURE DISTRIBUTION Once the path of contact is derived within a meshing simulation, the load distribution across the contact lines can be computed for a certain number of engagement positions by using contact influence figures (5). If the tooth profile has profile corrections or wear, the contact normal forces are not rectified and parallel anymore. This has got to be taken into consideration while formulating the load conservation equation. For a certain amount m of teeth in contact, discretized with a specified number of contact-points n, the discrete load conservation equation arises to equation (1). T = \u2211 \u2211 r , , F , , (1) By knowing the parameterized tooth profile, the effective radius rw arises to equation (2).",
+            "r , , r , , = r , , cos arctan \u2212 , , ,, , , + arctan \u2212 , , ,, , , (2) Because of the fact, that the contact area as well as the load distribution across the contact line and therewith the local Hertzian deformation is unknown, the contact problem has to be solved within an iterative process. Resulting from the parameterized profile the local radius of curvature arises to equation (3) taking profile corrections and wear into consideration. Based on the local radius of curvature, \u03c1 / r , , = , , , , , ,, , , , , , , , , , , , (3) the load distribution resulting from equation (1) across the path of contact and material constants, the local Hertz contact pressure results from equation (4). P , , = ,\u2206 , (4) \u03c1 , = , , , ,, , , , (5) 4. WEAR-SIMULATION Wear on tooth flanks occurs as a result of high sliding speeds, high contact pressure and low oil film thicknesses. Assuming unlimited oil supply and neglecting thermal influences, the local oil film thickness can be computed using equation (6) with regard to Dowson (6), where G is a material coefficient, U a velocity coefficient, W a load coefficient and  the replacement radius of curvature obtained by equation (5). h = 2,65 , ,, \u03c1 (6) On the basis of accurate local contact pressures, sliding speeds, specific sliding, local oil film thickness and material\u2019s hardness an empirical equation for wear on tooth flanks such as equation (7) can be formulated in accordance with Archard and Holm (6, 7). = k \u2219 k \u2044 1 \u2212 ,\u2211 \u2211 (7) Herewith the factor kS considers the chemical characteristics of the lubricant and the factor kW includes the material properties of the gears. Solving the differential equation (7) for a specific amount of load cycles dN leads to a certain amount of wear that is transferred onto the tooth profile. Carrying out a new meshing simulation with the newly worn out teeth leads to a changing path of contact in terms of the path of contact of unworn teeth and therewith to an altering load distribution throughout the wear simulation. Therewith, it is possible to simulate the micropitting test with reference to FVA 54 (9). The factor kS depends on the oil type and the chemical additive. For each type of lubricant this factor must be governed by an experimental test. Therefore may be used the FZG gear test rig according to DIN 51354-1."
+        ]
+    },
+    {
+        "image_filename": "designv11_60_0002430_jsvi.1996.0773-Figure4-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0002430_jsvi.1996.0773-Figure4-1.png",
+        "caption": "Figure 4. Numerical model 2: a laboratory test model (dimensions in m).",
+        "texts": [
+            " From Figure 3 it is evident that the proposed method is far more efficient than the direct method and can save more time as the number of degrees of freedom becomes large. Although the proposed method necessitates solutions of eigenvalue problems, no significant computational burden is caused because all the eigenvalue problems are self-adjoint and required to be solved only once. The present example confirms that the proposed method can significantly reduce the computation time without resulting in any errors. In the present example a typical re-analysis problem for a rotor\u2013bearing model is considered, as shown in Figure 4. The detail specifications of the rotor are given in Table 2. The number of nodal points in this example is 13, so that two 52\u00d752 complex dynamic stiffness matrices need to be solved for unbalance response analysis in the direct method. However, the number of bearings is just two, and the proposed method requires inversion of two 4\u00d74 matrix equations. Three types of bearings are considered in this example: four-pad and five-pad tilting pad bearings, and a two-axial-groove bearing. The bearing characteristics are given in Table 3"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0002614_s0094-114x(97)00067-0-Figure3-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0002614_s0094-114x(97)00067-0-Figure3-1.png",
+        "caption": "Fig. 3. An equivalent plane mechanism.",
+        "texts": [
+            " 4(b)\u00d0Caption on opposite page arbitrarily, provided that the \u00aerst-order in\u00afuence coe cient matrix of the branch is nonsingular. In order to verify the calculated result, we give a special example. The input displacement, velocity and acceleration are y={48 mm (\u00ff2 mm/s)t 0 0 (2p/180/s) t 0}T, _y f0 \u00ff2 mm/s 0 0 (2p/ 180)/ s 0} and y={0 0 0 0 0 0}T, respectively. In this case, the central point oh of the hand moves only in the YOZ plane. Vohx=Aohx= _jhy= _jhz= jhy= jhz=0. It is equivalent to a plane mechanism shown in Fig. 3. The position Pohy, Pohz, velocity Vohy, Vohz, acceleration Aohy, Aohz, angular displacement jhx, angular velocity _jhx and angular acceleration jhx of the hand for the prosthetic arm are shown in Fig. 4. The results are veri\u00aeed by the plane mechanism shown in Fig. 3. CONCLUSION In this paper, the kinematics of the prosthetic arm composed of a complex series\u00b1parallel mechanism is studied. The position, velocity and acceleration analysis of the prosthetic arm with less than six-DOF are discussed. The \u00aerst- and second-order in\u00afuence coe cient matrices of the arm are established using in\u00afuence coe cient by means of the method of hypothetic mechanism and all formulas of velocity and acceleration are derived by explicit expressions of in\u00afuence coe cient matrices"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0001051_phm.2016.7819861-Figure4-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0001051_phm.2016.7819861-Figure4-1.png",
+        "caption": "Figure 4. Crack propagation path along the tooth width.",
+        "texts": [
+            " 3(a), point Q on the tooth is the start position of the crack where the intersection angle of the tangent line and the gear tooth center line is \u03b2; and the point P is the vertex of the parabola representing the crack propagation path along the crack depth. The intersection angle between line PQ and the gear tooth center line is marked as \u03bd. As shown in Fig. 3(b), the propagation path along the crack depth can be defined by \u03b2 and \u03bd which are set as 30\u00b0 and 75\u00b0 in this paper according to Ref. [4]. The propagation along the tooth width can be simulated by another parabola shown in Fig. 4, in which Wc is the crack length along the tooth width. To simulate the crack with different levels, along the tooth depth, the crack is set to 10%, 20%, 50%, and 80% of the maximum crack depth, and each of which has three widths. Fig. 5 shows four depths of the crack on one tooth of the sun gear. Here, the crack depth value is calculated as the ratio of the current length QR and the maximum length QPM. According to the specification of the sun gear shown in Table I, the maximum length of the crack depth QPM is 1"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0002945_bf01129781-Figure7-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0002945_bf01129781-Figure7-1.png",
+        "caption": "Fig. 7. Determination of Cu 2 + in tap water by differential staircase stripping voltammetry",
+        "texts": [
+            " But as we shall explain in another communication, small pulse amplitudes proved to be superior in regard to both the signal to background ratio and detection limits. Table 1 reveals that the detection limit is in DPV and in differential staircase voltammetry about the same. Like in all these sophisticated voltammetric techniques, the detection limit is sensitive to the instrumental performance. Thus, the absolute detection limits may change with other instrumentation. Figure 6 shows the differential staircase vol tammogram of a Cu 2+, Cd 1+, and Ni 1+ containing solution. The method was used for the stripping determination of copper in tap water (Fig. 7). Acknowledgement. The authors express their thanks to Dr. C. Kuhnhardt for helpful comments. References 1. Bond AM (1980) Modern polarographic methods in analytical chemistry. Dekker, New York 2. Wang J (1985) Stripping analysis - Principles, instrumentation, and applications. VCH, Deerfield Beach 3. Barker GC, Gardner AW (1960) Fresenius Z Anal Chem 173 : 79 4. Barker GC (1960) Adv Polarogr 1 : 144 5. Mann CK (1964) Anal Chem 36:2424 6. Christie JH, Lingane PJ (1965) J Electroanal Chem 10:176 7"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0002751_31.45697-Figure2-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0002751_31.45697-Figure2-1.png",
+        "caption": "Fig. 2. Region for pole-assignment",
+        "texts": [
+            " Since the problem of robust controller design has been formulated as an optimization problem, by the techniques of nonlinear optimization we can solve the design problem systematically. This will be illustrated by an example. IV. AN ILLUSTRATIVE EXAMPLE Consider the system described in (lo), where the nominal matrices are \u2018\u201c1, with X = 0.5, -0.4 A = [ -0.1 -0.5 and subject to perturbations IAal1l< 0.01 IAa,,l Q 0.02 lAblll Q 0.012 while the other elements of AA and A B are zero. The poles of the system are required to lie in the region 9 which is the common region of open half-planes H,, H2 and H,, determined by L, , L, and L,, respectively, as shown in Fig. 2. First consider the nominal system. If state-feedback K = [ - 8, - 51 is used, the nominal closed-loop poles, X = - 1.0, - 1.5, are located in the desired region 9. To ensure the perturbed poles within the half-planes H I , H 2 , and H 3 , respectively, the corresponding upper bounds are calculated with p, = 0.000413, p 2 = 0.001504, p 3 = 0.001504. Because ( A U ? ~ + A u ? ~ + Ab;l),,,a = 0.01\u2019 + 0.022 + 0.0122 = 0.000644 > min (0.000413, 0.001504, 0.001504}, we cannot conclude by condition (4) of Theorem 1 that the poles of the perturbed system will lie within the region 9"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0002878_6.1989-1200-Figure2-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0002878_6.1989-1200-Figure2-1.png",
+        "caption": "Fig. 2 Force and moment resultants acting on helical wire.",
+        "texts": [
+            " 1): D ow nl oa de d by F re ie U ni ve rs ita et B er lin o n M ay 1 4, 2 01 5 | h ttp :// ar c. ai aa .o rg | D O I: 1 0. 25 14 /6 .1 98 9- 12 00 Fig. 1 Helical spring subjected to end force and torque. where s is the length along the spring wire center line, N' and N are the normal and binormal shear forces, T is the tensile force, G is the binormal bending moment, H is the torsional couple, and X, Y and Z are external forces, and K', K and 8 are external moments per unit length of the wire central line in the normal, binormal and tangential directions, respectively (see Fig. 2). The normal bending moment G does not exist in the present case due to the kinetic symmetry of the wire cross section. The R and t i n the equilibrium equations are the curvature and twist of the deformed wire central line. In the case of small deformations, they can be replaced by the initial curvature n and twist T. The curvature and twist of the wire central line can be found from the helical angle a and the helix radius r: where k = tan a The deformed curvature it and twist 7 from the deformed helical angle & and helix radius f by similar expressions"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0000739_978-1-4471-4976-7_91-1-Figure2-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0000739_978-1-4471-4976-7_91-1-Figure2-1.png",
+        "caption": "Fig. 2 AAU 5-dof robotic arm: (a) a prototype and (b) coordinate systems",
+        "texts": [
+            " 25 and 26 leads to: M FT q Fq 0 \u20acq l \u00bc n c (27) Equation 27 can solve simultaneously the accelerations and the Lagrange multipliers. Two examples are included to apply the equations of motion to robotic manipulators. In the first example, a lightweight robotic arm consisting of five revolute joints developed at Aalborg University (AAU), Denmark, is considered. In the second example, the dynamics of a spherical parallel manipulator is presented. Example I: A 5-dof Lightweight Robotic Arm The lightweight robotic arm is demonstrated in Fig. 2a. Following the D-H convention, Cartesian coordinate systems are attached to each link of the manipulator, as shown in Fig. 2b. The D-H parameters of the manipulator are listed in Table 1. The kinematics of the robotic arm can be found in (Zhou et al. 2011). This chapter includes here only the dynamics formulation. Jacobian matrix The joint angular velocity can be calculated with the Jacobian matrix _u \u00bc J 1vef (28) where _u\u00bc _y1, _y2 . . . , _yn T h denotes an n-dimensional (n denotes the number of dof) vector of the joint angular velocities, J is the Jacobian of the robotic arm, and vef the velocity of the end-effector"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0002377_elan.1140040207-Figure21-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0002377_elan.1140040207-Figure21-1.png",
+        "caption": "FIGURE 21. Mass transport corrected Tafel analysis of the current-voltage curves shown in Figure 20.",
+        "texts": [],
+        "surrounding_texts": [
+            "Figures 20 and 21 show thc computed voltammograms and Tafel plots for the case a*&* = 1 and for four different values of the Leveque-limit normalized rate constant: Subcase (1,): k(-,,zi(A] >> k;JY/. K*** = k,J(~~i,~*{4h\u2018x,\u2018dL/9V~D~\u201d\u2019 (31) Again, the parameters of the EC\u2019 mechanism were used in the computations. The behavior is entirely analogous qualitatively to the corresponding subcase of the EC\u2019(dispA) mechanism. As with that subcase\u2019s calculations on experimentally realistic conditions, the results are just outside the kveque limit and, hence, full computations are demanded for mechanistic analysis. Figure 22 displays the shift in halfwave potential withK*** which is also similar to the analogous EC\u2019( dispA) subcase. While the magnitudes ofthe shift-both anodic and cathodic-are smaller than for the EC\u2018(dispA 1) subcase, the shapes of the plots reflect the close similarity of the two mechanisms. Suhcme (2) k,,{YJ >> k(-,JA/. The results here are identical t o those generated for subcase ( 2 ) of the EC\u2019(dispA1) mechanism, and are presented in Figures 17-19, except that P*/2 replaces K** in the legends to these diagrams. The arguments relating to the waveshapes observed are the same, except that reactions (vii) and (ix) have the overall effect of turning one B molecule into one A molecule."
+        ]
+    },
+    {
+        "image_filename": "designv11_60_0003019_iros.1996.568948-Figure3-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0003019_iros.1996.568948-Figure3-1.png",
+        "caption": "Fig. 3 Grasp with four frictionless contacts",
+        "texts": [
+            " Therefore, the number ofjoints and contact constraints are determined by m = 2, n = 3. For the given grasp configu- ration, the grasp matrix W and the Jacobian matrix JT are determined in the reference fiame as 1 W = [ t -+ -% 77 0 0 is determined by finding an optimal value of the joint torque. This process ensures that a solution can be obtained by adjusting the joint torques. We see in the following examples that these features can always be obtained by Eq. (15). 4.2 Case 11: fpl = 0, fh2 = 0 In the grasp shown in Fig. 3a, each finger contacts the object at two points. m = 2, n = 4 and the corresponding matrices are -1 -- Jz h A A 0 - 2 0 0 JT=[o 0 2 01 Since rank(W) = 2, rank(JT) = 2 and rank(A) = 4, by using Eq. (6), we obtain Ni = 1, N, = Oand dimrpl = 0, dimrp2 = 2, dimrhl = 2, d imrh2 = 0. Therefore, the contact force is decomposed into f = fp2 + fh1 and can be determined as Fig. 3b shows the meaning of the decomposition, The contact force can also be formulated with respect to the external force and the joint torque Note that, in this case G1 = 0 , which means the grasp can only passively resist the applied external forces. 4.3 Case III: fpl = 0 In the grasp shown in Fig. 4a, the fingers contact the object at five points, m = 2,n = 5. The corresponding matrices are determined as I J = o [ 0 2 4 7 0 0 - 2 0 0 0 In this case, rank(A) = 4 is less than the full rank. Hence, there exists uncontrollable components ofthe contact force"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0001900_156855395x00076-Figure8-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0001900_156855395x00076-Figure8-1.png",
+        "caption": "Figure 8. Inclinometer.",
+        "texts": [
+            " We control the D ow nl oa de d by [ U ni ve rs ity L ib ra ry U tr ec ht ] at 0 2: 10 1 7 M ar ch 2 01 5 184 inclination of the upper surface using {A} and {F}, and the coordinate transformation is written as follows: 4.2. Relation between Euler angles and legs This robot moves by sliding the two frames and supporting the body by each frame, so that inclination control is necessary for each frame. Three legs of each frames form an equal triangle. We show how to control the inclination of the A-frame on which the manipulator is to be fixed. As shown in Fig. 8, an inclinometer that can measure two Euler angles (p, y) is fixed on the upper surface. FD1, FDz, FD3 are the legs in Fig. 8, and FXo and Yo are unit vectors. Their position vectors are presented in equations (8)-(17). D ow nl oa de d by [ U ni ve rs ity L ib ra ry U tr ec ht ] at 0 2: 10 1 7 M ar ch 2 01 5 185 The height of three apexes, (h 1, h2, h3) is written as follows: From the equation (19) and (20), the relation between (h 1, h2, h3 ) and the Euler angle is written as follows: I .., I 4.3. Equation and inclination of the upper surface Let the center of the A-frame set at the origin 0 of {A}, we can express the equation of the upper surface as follows: The coefficients (l, m, n) are calculated as follows: D ow nl oa de d by [ U ni ve rs ity L ib ra ry U tr ec ht ] at 0 2: 10 1 7 M ar ch 2 01 5 186 Figure 9"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0000269_9781782421955.540-Figure2-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0000269_9781782421955.540-Figure2-1.png",
+        "caption": "Fig. 2 \u2013 Aronhold\u2019 first theorem: involute tooth profiles and return circle.",
+        "texts": [
+            " In particular, the Euler-Savary equation for envelopes is given by 0 0 1 1 1 1cos l lP P r r\u03c3 \u03bb \u03b1 \u239b \u239e \u2212 = \u2212\u239c \u239f \u03a9 \u03a9\u239d \u23a0 (1) where \u03b1 is the polar coordinate of point P with respect to the ordinate axis of the canonical frame (P0 t n), P0\u03a9\u03c3 and P0\u03a9 l gives the positions of the centers of curvature \u03a9\u03c3 and \u03a9 l with respect to P0, while r\u03bb and rl are the radii of curvature of \u03bb and l, respectively. Thus, \u03a9\u03c3 is not only the center of curvature of the profile \u03c3, but also the center of curvature of \u03a9 l , and viceversa, during the relative motion between the conjugate profiles \u03c3 and l that is represented by the centrodes \u03bb and l. Camus\u2019 theorem can be also applied to the generation of the involute tooth profiles of non-circular gears in the form of the well-known rack-cutter method. In fact, referring to Fig. 2, when the auxiliary centrode \u03b5 becomes the tangent line t to both centrodes and its attached curve \u03b7 becomes also a straight line, a pair of involute conjugate profiles can be generated as envelope of \u03b7, while their normal line N at their contact point P envelops the corresponding base curves, which are also the loci of the centers of curvature of each involute tooth profile. According to Aronhold\u2019 first theorem, the return circle R is the locus of the centers of curvature \u03a9\u03c3 for all envelopes \u03c3 whose generating curves are straight lines, as the line \u03b7, which is attached to the auxiliary straight line centrode \u03b5"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0003063_1.428360-Figure2-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0003063_1.428360-Figure2-1.png",
+        "caption": "FIG. 2. Ball bearing elements and coordinate systems.",
+        "texts": [
+            " To the author\u2019s knowledge, the effect of the outer ring ovality and the shaft elasticity of systems incorporating radial ball bearings under a mass unbalance has not been studied before. This work simulates the outer ring ovality effect on vibrations of a rotor supported on nonlinear radial ball bearings and driven by a mass unbalance. Here we present a general formulation and a rigid shaft response. Figure 1 shows a rigid disk of a circular shape mounted through its center to an elastic shaft which in turn is mounted on two elastic radial ball bearings. These supporting bearings, however, are mounted into their rigid housings. The bearing details are shown in Fig. 2. In Fig. 1, the triad XYZ is a global coordinates system with its origin at the geometrical center of the shaft left bearing, where the X axis coincides with the shaft bearing\u2019s center line in the nonworking ~zero speed! position of the system. The orientation of the de- 852awzi M. A. El-Saeidy: Rotating machinery dynamics simulation /content/terms. Download to IP: 129.101.79.200 On: Fri, 29 Aug 2014 09:52:47 Redistr flected rotor element in space ~Fig. 3! is monitored using Euler angles ~Fig. 4",
+            " and gyroscopic matrices, respectively. m j d is the mass, and Ip j d and Id j d are as defined before but for the disk. Let ab j d , ac j d be the disk mass center eccentricities in the b, c directions, then its generalized unbalance force vector is Fu j d 5~Q\u0304s! TFm j dVs 2I2 0\u03042 0\u03042 0\u03042 G @ab j d ac j d 0 0#T. The advantages of the bearing model14,16 compared to other existing analyses ~see references listed in Ref. 14! are discussed in Ref. 14. This model is extended herein to account for outer ring ovality. Figure 2 depicts a ball bearing system without outer ring ovality and with uniform radial clearance where the global coordinate system XYZ has its origin at the bearing center with the X axis coinciding with the bearing axis. The frame x\u0304by\u0304 bz\u0304b is a rotating coordinate system spinning with the bearing cage angular speed, Vc ~rad/s!, where the x\u0304b axis coincides with the bearing axis. The bearing inner ring is lightly fitted on its shaft and is modeled as an integral part of it and thus rotates with the angular speed Vs "
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0000921_iecon.2014.7048563-Figure4-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0000921_iecon.2014.7048563-Figure4-1.png",
+        "caption": "Fig. 4. Output voltages of two inverters and stator phase voltage.",
+        "texts": [
+            " Since zerosequence impedance consists of only small leakage inductance, zero-sequence voltage may cause large zero-sequence current. Hence, the reference stator phase voltage v\u2217s has to be on the plane vs0 = 0 to suppress the zero-sequence current. Thus, the reference voltages of the two inverters are required to have the same zero-sequence component. For open-end winding IM control, the reference stator phase voltage v\u2217s is converted into the reference output voltages of the two inverters, v\u2217i1 and v\u2217i2, based on the relation among v\u2217i1, v\u2217i2 and v\u2217s illustrated in Fig. 4. First, the phase difference between the two inverters is set to a certain value \u03b8d. The ratio between the amplitudes of v\u2217i1 and v\u2217i2 is fixed to one. Then, the electrical angles of Inv1 and Inv2, \u03b81 and \u03b82, are derived from the stator electrical angle \u03b8 and the phase difference \u03b8d as below. \u03b81 = \u03b8 \u2212 \u03c0 \u2212 \u03b8d 2 (13) \u03b82 = \u03b81 \u2212 \u03b8d (14) The magnitudes of the reference voltages of the two inverters in d-q plane are calculated as below. \u2016v\u2217i1dq\u2016 = \u2016v\u2217i2dq\u2016 = Kv\u2016v\u2217s\u2016 (15) Kv = (2 sin \u03b8d 2 )\u22121 (16) From (16), the stator phase voltage is maximum when the phase difference is 180\u25e6"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0001781_s0167-8922(08)70009-9-Figure5-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0001781_s0167-8922(08)70009-9-Figure5-1.png",
+        "caption": "Figure 5 - Short Bearing Mobility Map for a Cavitating Film Showing",
+        "texts": [
+            " The first step can be termed a 'squeeze' step (although not pure normal squeeze) and the second step a 'whirl' step. (b) For the squeeze step, if M is the vector of (ME) and (My), then from equation (1 2) representing V, as the vector sum of ( E ) and (E@) we have Now the scalar magnitude of the vector M can be calculated at any point ( E , w ) from equation (10) using either equations (8) or (9). The direction associated with the vector is given by the tangent of the ratio of the resolved mobility components and is termed the squeeze path. (c) Thus in Figure 5 the clearance circle, a circle representing the possible range of movement of the shaft centre and of radius equal to the radial clearance (c), is shown. The magnitude of mobility M at any point in the clearance circle can be determined and the squeeze path direction calculated. Lines of constant mobility may be determined. If the shaft is at a given point (A) as shown its movement over a time period (dt) will be given by the 'Squeeze' Component in the direction of the squeeze path. (see Figure 5). Clearly this will be scaled. (d) Now the angular velocities can be taken into account. movement due to these will be a component (see equation 12(b)), The shaft directed tangentially. Here a sign convention is necessary and anticlockwise rotations have usually been taken as positive. This component can be vectorially added to the squeeze component to give the resultant motion, as shown in Figure 6. Strictly speaking since the value of the average mobility number of the resultant will be different from that used in calculating the squeeze step, an iteration to some acceptable tolerance will be required"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0002923_978-94-011-3514-6-Figure4.2-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0002923_978-94-011-3514-6-Figure4.2-1.png",
+        "caption": "Fig. 4.2 Harmonic steady state; graph of rotating vectors.",
+        "texts": [
+            " In the steady state x = X cos (wt - \"') and these forces have the value f(t) external force, in advance of the displacement x by the [Jhase '\" f(t) F cos wt fm(t) force of inertia, in phase opposition to the displacement m x - w2 m X cos(wt - \"') w2 m X cos(wt - '\" + n) (4.14) - 55 - fc(t) viscous frictional force, in advance of the displacement by the phase of v/2 (in quadrature) c x - w c X sin(wt - ~) w c X cos(wt - ~ + u/2) fk(t) elastic returning force, in phase with the displacement k x k X cos(wt - ~) These forces are equal to the projections onto an axis of vectors rotating at the same angular velocity w (fig. 4.2). The unknowns of the problem are the amplitude X and the phase ~ which the graph enables one to calculate easily (k X - w2 m X)2 + (w C X)2 F2 => X F l/(k - w2 m)2 + w2 c 2 tg ~ w C k - w2 m One clearly finds again the results (4.5) and (4.6). - 56 - A more efficient method of calculation than the above, although based upon the same basic idea, consists of replacing the vectors by complex numbers. This method, developed originally by electrical engineers, will often be used in the following text"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0001097_ecce.2016.7854980-Figure1-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0001097_ecce.2016.7854980-Figure1-1.png",
+        "caption": "Fig. 1 1-shunt inverter with SPMSM",
+        "texts": [
+            ". INTRODUCTION The phase current information is required for the vector control of the AC motor. Generally, hall-effect type current sensors or three shunt resistors are used to measure the phase current. For the cost reduction, a DC link shunt resistor can be used to measure the phase current, as shown in Fig. 1 [1]-[2]. When an active voltage vector is applied by the 1- shunt inverter, a phase current flows through the shunt resistor, as shown in Table 1. Thus two phase currents can be sampled from the shunt resistor because two active vectors are applied in a switching cycle. The other phase current is calculated by using the following equation: (1) However, if the duration of the active vector is shorter than the minimum time to sample the DC link current accurately, it is impossible to obtain the phase current from the shunt resistor"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0001235_b978-0-08-099425-3.00005-4-Figure5.31-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0001235_b978-0-08-099425-3.00005-4-Figure5.31-1.png",
+        "caption": "FIGURE 5.31",
+        "texts": [
+            " The lateral force that arises due to an inclination of the tyre from the vertical is referred to as camber thrust. The SAE definition of positive camber angle is taken Plotting aligning moment versus slip angle. for the top of the tyre leaning outwards relative to the vehicle. The fact that this differs from one side of the vehicle to the other does not lead to consistency when developing a tyre model. For understanding it is useful to remember that the camber thrust will always act in the direction that the tyre is inclined as shown in Figure 5.31. For the SAE system shown here a positive camber angle g will produce a positive camber thrust for all tyres on the vehicle modelled in that system. If the tyre is inclined at a camber angle g, then deflection of the tyre and the associated radial stiffness will produce a resultant force, FR, acting towards the wheel centre. Resolving this into components will produce the tyre load and the camber thrust. An alternative explanation provided in Milliken and Milliken (1998) compares a stationary and rolling tyre"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0000533_vss.2016.7506908-Figure4-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0000533_vss.2016.7506908-Figure4-1.png",
+        "caption": "Fig. 4: Schematic diagram of the magnetic levitation system.",
+        "texts": [
+            " The last plot shows the sequence (\u2212wk/\u03c4) (output of the discrete integrator in the controller) which tracks the estimation error whenever quasi sliding mode is reached. Fig. 3 shows the precision of the state variables when simulating with \u03c4 = 0.01 s, \u03c4 = 0.001 s, \u03c4 = 0.001 s. The results agree with the theoretically obtained precision in Corollary 2. The coefficients \u03bd1 = 2.5, \u03bd2 = 53 and \u03bd3 = 5813 are obtained accordingly. This Section outlines the application of control schemes providing for a certain \u03c3-accuracy, to a magnetic levitation system. A schematic drawing of the magnetic levitation system is provided in Fig. 4. The system input is the supplied voltage u, the position of the steel ball is considered as output. The disturbance d is artificially added to the control channel. The state vector is defined as x := [ y y\u0307 i ]T . In [21] the mathematical model x\u0307 =  x2 gc \u2212 c m x2 3 x2 1 \u2212R Lx3 + 2c L x2x3 x2 1 + 0 0 1 L  (u+ d), y = x1 (33) is suggested to describe the dynamics of such a system. The output y has relative degree r = 3 with respect to the input u. Using the local diffeomorphism z = t(x) = [ x1 \u2212 r x2 gc \u2212 c m ( x3 x1 )2 ]T (34) the model is expressed in new coordinates yielding the system description1 z\u0307 =  z1 z2 f(z) + g(z)(u+ d), y = z1"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0000510_htj.21229-Figure2-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0000510_htj.21229-Figure2-1.png",
+        "caption": "Fig. 2. Whole machine finite element model.",
+        "texts": [
+            " The heat source of the main spindle system is the spindle motor and bearing friction (four angular contact ball bearings in this spindle system). The motor drives the spindle and test bar to rotate at high speed to generate forced convection cooling using the air. The main shaft sleeve and bearing sleeve will have a contact cooling influence on the spindle system. Therefore, the contact thermal resistance between parts should be taken into consideration. The whole machine finite element model is shown in Fig. 2. The main structure of the machine tool includes the spindle system, bed, bed saddle, workbench, pillar, guide rail, spindle box, and cooling channel of the spindle box (with an outside baffle, refer to Section 5 for detailed structure). In the machine tool, the material of the bed, bed saddle, workbench, pillar, and spindle box is HT300, the material of the spindle and guide rail is 40 Cr, the material of the bearing is GCr15, the material of the accuracy test bar is T10A, while the material of the spindle sleeve and bearing sleeve is structural steel"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0000812_epecs.2015.7368542-Figure1-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0000812_epecs.2015.7368542-Figure1-1.png",
+        "caption": "Fig. 1. Windings layout. (a) Full-pitch concentrated winding. (b) Singletooth winding.",
+        "texts": [
+            " For a single-tooth winding, the required number of slots to generate a 2-pole flux distribution when a fundamental current sequence is applied is 10. Hence, the same 10-slot stator is wound with five-phases with full-pitch concentrated coils. This case is the simplest case for a distributed winding with a number of slots/phase/pole equals 1.The numbers of turns per coil for both windings are selected to produce same fundamental MMF component, as explained in [12]. Both winding layouts are shown in Fig. 1. In five-phase systems, there are four available sequences comprise two theoretically decoupled sequence planes and one zero sequence [14]. A set of five-phase currents with a phase shift of 720 represents the first sequence and when applied to a five-phase winding a fundamental flux distribution is produced. For this sequence, the MMF distributions for both windings are shown in Fig. 2. It is clear that with this current sequence, a fundamental component is produced while the lowest order harmonics are the 9th and 11th"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0002837_ip-b.1991.0017-Figure2-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0002837_ip-b.1991.0017-Figure2-1.png",
+        "caption": "Fig. 2 V, = 0.24 pu ~ calculated -0- measured V curves of experimental hysteresis motor",
+        "texts": [
+            " The air-gap length of the motor was made long (2 mm) to decrease the influence of space harmonics. A circumferential-flux type hysteresis motor was used, although the following can also be applied to a radial-flux type hysteresis motor. Fig. 1 shows the change in stator voltage considered. The motor was accelerated to the synchronous speed at a stator voltage Vo . The stator voltage is slowly reduced to a value V, after synchronisation. It is assumed that the load torque remains unchanged when this occurs. Fig. 2 shows the relation between the stator current after synchronisation and the voltage ratio VJV, (overexcitation ratio). The calculated result shown in this Figure is explained later. V, was constant (0.24 pu), and Vo was changed in the range from 2V, to S V , . The power factor of the motor is unity at the point of the minimum 137 stator current. Points to the right of the minimum stator current correspond to a leading power factor. Points to the left correspond to a lagging power factor. The curves in Fig. 2 are V curves of the hysteresis motor which correspond to those of a conventional synchronous motor with a field winding. 0 4 r 3 Analytical investigation The V curve of the hysteresis motor is derived by using the method proposed in the previous paper [SI. 3.7 Magneric stare of roror As a basis for deriving the V curve of the hysteresis motor, the magnetic state of the rotor hysteresis material before and after reduction of the stator voltage was analysed. The following assumptions were made to simplify this analysis : (a) The magnetic flux density distributions in the air-gap and in the hysteresis material are sinusoidal",
+            " When the stator voltage is reduced from Vo to V, , only 4 is iterated to give the required torque TL at a peak flux density, which is gradually reduced. In the final stage of computation, both B, and 4 are iterated to give the required voltage V , and torque TL. In the program, 360 electrical degrees were divided into 64 segments. The magnetic state at each segment in the rotor hysteresis 139 material was given as the data of the loop. The values of el, E , and AB,,, in Fig. 5 are very small and are given as E~ = 0.001 Nm E , = 0.001 v AB,,, = 0.02 T 3.3 Comparison between measured and calculated results Fig. 2 shows the comparison of the measured and calculated performance of the experimental motor. The calculated result shows, like the measurements, that the relation between stator current and overexcitation ratio, V,/V,, is a V curve. 4 Discussion of results The V curve of a hysteresis motor has been clarified both by experimental result and by calculations. This phenomenon is discussed below. 4.1 Interpretation by equivalent circuit Fig. 6a shows a simplified equivalent circuit of the hysteresis motor obtained from the equivalent circuit in Fig"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0003946_a:1013325731855-Figure1-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0003946_a:1013325731855-Figure1-1.png",
+        "caption": "Figure 1. Orbital\u2013rotational variables.",
+        "texts": [
+            " The potential acting on the rigid body S may be represented up to the second order by means of the well-known MacCullagh\u2019s formula V = \u2212G mM r \u2212 G M 2r3 [(I1 \u2212 I2)(1 \u2212 3\u03b12)+ (I3 \u2212 I2)(1 \u2212 3\u03b3 2)], (10) where G is the gravitational constant, r is the distance between the centers of mass O and C, and (\u03b1, \u03b2, \u03b3 ) the direction cosines of the unit position vector u = x/r expressed in the body frame B, that is to say, u = x/r = b1\u03b1 + b2\u03b2 + b3\u03b3. These direction cosines suggest to choose an orbital moving frame O(u, v,w), where w is in the direction of the orbital angular moment (* = x \u00d7 x\u0307), and v = w \u00d7 u. In this way, the two bases O and B are related by several compositions of rotations involving as arguments both the Serret\u2013Andoyer ( , g, h) and Polarnodal (r, \u03b8, \u03bd) variables (see Figure 1). In matrix form, there results (\u03b1, \u03b2, \u03b3 ) = R (1, 0, 0), with R the rotation obtained by the composition of seven rotations R(\u2212 ,b3) \u25e6 R(\u2212\u03c3,m) \u25e6 R(\u2212g,n) \u25e6 R(\u2212\u03b5, ) \u25e6 R(\u03bd \u2212 h, s3) \u25e6 R(I,N) \u25e6 R(\u03b8,w), where N and I are the node direction and inclination of the orbital plane (uv) with respect to the space plane (s1s2). Once obtained the direction cosines \u03b1, \u03b2 and \u03b3 , it is necessary to square their expressions and replace them in the Equation (10). By doing so, in the potential function will appear expressions of the type \u220f a,b sina cosb, that is to say, combinations of products of powers of trigonometric sines and cosines of the involved variables"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0000510_htj.21229-Figure5-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0000510_htj.21229-Figure5-1.png",
+        "caption": "Fig. 5. (a) Temperature and layout of displacement sensors, (b) Layout of displacement sensors T1 to T8.",
+        "texts": [
+            " Besides the spindle system, the thermal deformation of the other parts is small. To validate the accuracy of the simulation, this paper conducts thermal characteristics experiments on the machine tool. Thirty-two temperature sensors were used to measure the temperature field of the machine tool. An eddy current displacement sensor was used to measure the thermal deformation in the Z direction. Two laser displacement sensors were used to measure the thermal deformation in the X and Y directions. The sensor distribution is shown in Fig. 5(a). Temperature sensors T1 to T8 were installed on the spindle front end and flange cover, as shown in Fig. 5(b), while T23, T24, and T25 were installed on the other side of the spindle box, in symmetry with T26, T27, and T28. Sensor T32 was located in the machine cover. Measuring the environment temperature, the spindle speed was 6000 rpm. Initially testing the temperature field and thermal deformation of the machine tool, and thermal characteristics curve in steady-state, we then turn off the machine. Then, compare the test value and simulation value in steady-state. As shown in Table 4, there is a comparison of the test values of spindle box front temperature sensor T1, spindle box sensor T13, spindle box surface sensor T21, and spindle box side sensor T26 and the simulation value"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0003624_3.56200-Figure1-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0003624_3.56200-Figure1-1.png",
+        "caption": "Fig. 1 The spacecraft model.",
+        "texts": [
+            " The minimization of the energy consumed by the reaction wheel motors is clearly a worthwhile goal. Although the integral of power over the maneuver time would seem to be the ideal performance^ index since it represents total mechanical work, it does not yield a unique torque history. Moreover, the total work criterion rewards braking, or negative work, as much as it penalizes positive work. Hence, the index treated herein is the integral of power squared over time. Minimum Power Formulation for Slewing Maneuvers Consider the rigid-body configuration of Fig. 1, where \u00a3 denotes the principal axis about which the maneuver is to occur, / is the spacecraft mass moment of inertia about \u00a3, / is the reaction wheel axial moment of inertial about _\u00a3, 0 and </> are wheel and spacecraft inertia angular displacements, and u is the motor torque exerted by the motor on the wheel. The problem to be considered is the determination of u ( t ) such that the performance index P2dt 0) (where P is power) is minimized while boundary conditions on the spacecraft and wheel states at the initial and final times are Received March 2,1982"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0002002_s0094-5765(99)00125-3-Figure3-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0002002_s0094-5765(99)00125-3-Figure3-1.png",
+        "caption": "Fig. 3. The planar truss structure considered for dynamics simulation.",
+        "texts": [
+            " After the assembly of the nodal mass and sti ness matrices, the eigenvalue problem is solved using routines of the EISPACK library. A speci\u00aeed number of modes to be used in the simulation are retained and the corresponding eigenvectors are ortho-normalized. These are used in eqn (4). Integration of the equations of motion (28) is carried out in the code using either the Gear routine from the IMSL library, Adams\u00b1Moulton method or fourth order Runge\u00b1Kutta integrator. Simulation results were obtained for the planar truss structure shown in Fig. 3, placed in a zero-g environment and attached to a \u00aexed base. This con\u00aeguration was adapted from the 3D space crane concept of Wu et al. [5] and employs similar system parameters. The extended length of the structure is 95 m and the width is 5 m. The total mass of the structure is 1106 kg, which includes 300 kg of additional mass (as a platform and additional hardware situated along the end batten). The crane consists of three articulating truss booms, which are articulated via the joint con\u00aeguration shown in Fig. 3. This arrangement allows for a robust 908 planar articulation of each boom, by virtue of the two actuators per joint. The density and Young's modulus of the longeron, diagonal, batten and end-batten members are 8000, 4000, 12,000, 8.49 106 kg/m3, 412.5, 412.5, 618.7 and 618.7 GPa, respectively. Cross-sectional area is 2.36 10\u00ff4 m2. The dynamic model consists of 19 individual links; 7 trusses, 6 cylinders and 6 piston-rods. Initially, each actuator piston-rod has zero extension, however, to e ect the motion of Fig. 3(b), the \u00aerst two actuators are commanded to extend during the 20 second maneuver per the following trajectories, d t Dd T t\u00ff T 2p sin 2p T t , 67 where, T = 20 s, and Dd = 3.832 m. The remaining four actuators are commanded to have zero extension throughout the maneuver. For times greater than 20 s, actuators 1 and 2 maintain the extended position of d = 3.832 m. It should be noted from the trajectories in eqn (67), that actuators 1 and 2 are speci\u00aeed to yield zero velocity and acceleration at the initial and \u00aenal maneuver times. However, based on the acceleration trajectory, a jerk will be imparted to the structure at t = 20 s. The \u00aenal actuator extension of 3.832 m was simply pre-com- puted based upon the joint geometry of Fig. 3, corresponding to the \u00aenal orientation. The actuator forces to associated with the maneuver and as obtained from the inverse dynamics, are given in Fig. 4. For the joint con\u00aeguration modelled in this space crane, the magnitude of the force between adjacent actuators will be similar, by virtue of the symmetry of their placement and of the small inertia contribution of the three member truss links interfacing the actuator pairs. The actuator force pro\u00aeles of Fig. 4 can be used to perform the simulations of the forward dynamics"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0000250_phm.2014.6988149-Figure3-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0000250_phm.2014.6988149-Figure3-1.png",
+        "caption": "Fig. 3. The position of bearings and load",
+        "texts": [
+            " \u2022 Symmetry: I(X,Y) = I(Y,X) \u2022 Self-information: I(X,X) = H(X) \u2022 Boundedness: ( ) ( ) ( ){ } ( ) ( )( ) ( ) ( ){ } ( ) ( ) ( ) min 2 max I X ,Y H X ,H Y H X H Y H X ,H Y H X ,Y H X H Y \u2264 \u2264 + \u2264 \u2264 \u2264 + When the bearing is working in a fault-free state, the correlation of vibration signal at two different periods of time is big, and the value of mutual information is big, too. However, when faults occur in bearings, the correlation of vibration signal at two different periods of time is smaller than working well, so the value of mutual information is small either. DIAGNOSIS To validate the effectiveness of the proposed approach for fault diagnosis, we use conical roller bearings to conduct the experiment. The position of the bearing and loading in the experiment are shown as Fig. 3. In the experiment, 1# and 4# conical roller bearings are for experiment. While 2# and 3# bearings are loading bearings, and they are cylindrical roller bearings, as shown in Fig. 3. The basic size and other parameters of two different kinds of bearings are shown in table I. Results of the experiment are as followings, the timedomain and frequency-domain signals are calculated, as shown in Fig.4-(a) and Fig.4-(b). TABLE I. THE BASIC SIZE OF BEARINGS Type Bearing inner diameter (mm) Bearing outer diameter (mm) Rolling diameter (mm) Rated dynamic load rC (KN) Limiting speed (r/min) tested bearing 55 120 16.25 152 4300 loaded bearing 60 130 19.1 212 5600 0 0.5 1 1.5 2 2.5 3 -8 -6 -4 -2 0 2 4 6 8 Time(s) A m pl itu de (g ) (a) 0 1000 2000 3000 4000 5000 6000 0 1000 2000 3000 4000 5000 6000 Frequency(Hz) A m pl itu de (g ) (b) Fig"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0002215_0094-114x(95)00011-m-Figure2-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0002215_0094-114x(95)00011-m-Figure2-1.png",
+        "caption": "Fig. 2. The geometric mapping from output-space to control-space.",
+        "texts": [
+            " As will be seen, when particular examples of planar path tracking systems are considered, equation (26) is considerably simplified. Knowing n and n' allows for the system to be controlled (coordinated), by the second-order Taylor series, /a = n2 + \u00bdn'22. (27) Alternatively, the inverse of this series may be used for the coordination. Equations (14) and (26) are purely geometric mapping of the first- and second-order differential properties of the output-space trajectory into the control-space trajectory. This mapping is illustrated by Fig. 2. 3. S Y N T H E S I S OF PLANAR TWO DEGREE-OF-FREEDOM N O T I O N S The Curvature Theory of planar two degree-of-freedom motion will be applied here to the synthesis of a planar two degree-of-freedom motion. Specifically, the motion synthesis problem is that of determining the differential control variables necessary to achieve specified differential output variables. For ths specific example of motion synthesis, equations (14) and (26) are applied to two examples of planar two d.o.f, path-tracking systems",
+            " 8 0 Therefore, the coordinating Taylor series from equation (27) is, It = n2 + \u00bdn'22 = -2.332 + \u00bd(- 1.80)22. Clearly, coordination is always possible via Curvature Theory, i.e. it is non-singular. This example illustrates a robustness to singularities. Differential output-space trajectory geometry can be successfully mapped into differential properties of the control-space trajectory geometry even at a point on the output-space trajectory at which the system is singular. The curvature Theory mapping, illustrated by Fig. 2, is possible even in a singular configuration. Also apparent from Curvature Theory is that in a singular configuration the controlled point B will be at a momentary dwell on its trajectory, since it has come into coincidence with the Pole. Singularities are not encountered in control of the trajectory geometry. This paper has developed a generalized planar two-degree-of-freedom Curvature Theory and demonstrated its application to a two-degree-of-freedom motion synthesis problem, which is the trajectory generation problem of planar path tracking systems"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0003622_bf01589377-Figure1-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0003622_bf01589377-Figure1-1.png",
+        "caption": "Fig. 1. Glucose requirement for pyruvate uptake. Cells were placed in the standard uptake mixture with the following modifications: (O--O) no omissions, (O--Q) no glucose, (\u2022215 no MgSO4 and NaCI, (A--A) no MgSO4, and (O--[]) no NaC1.",
+        "texts": [
+            " The mixtures without antibiotic were used as control. The kinetics of the utilization of pyruvate was followed by determining the residual pyruvate in the supernatants of the SUM at different times. Uti l izat ion o f pyruvate . The charac te r i s t i cs o f the p roces s o f p y r u v a t e up take in Lac tobac i l lu s case i subsp, r h a m n o s u s A T C C 7469 were inves t iga ted with cel ls g r o w n in a syn the t i c m e d i u m with 33 m M glucose and the addi t ion o f 100 m 3 / / p y r u v a t e . As Fig. 1 shows , the re is no p y r u v a t e up take in the a b s e n c e o f g lucose . In the p r e s e n c e o f this sugar, p rac t i ca l ly the who le o f the e x o g e n o u s pyru- va te is c o n s u m e d af ter 90 min. We can state that p y r u v a t e uses an ac t ive t r anspor t that depends on the p r e s e n c e o f g lucose as a source o f energy. Salts such as 2.3 m M m a g n e s i u m sulfate or 14 m M sod ium ch lo r ide do not af fec t up take , w h e t h e r they are used t oge the r or separa te ly "
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0000617_s1064230716030126-Figure4-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0000617_s1064230716030126-Figure4-1.png",
+        "caption": "Fig. 4. Possible location of CCD cameras on the working tool model.",
+        "texts": [
+            " It is necessary to compensate for the inaccuracies of the environment model induced, as has been mentioned above, by the difference between the positions in the base coordinate system of real objects and the positions of their models. This deviation is easily compensated by the corresponding displacement of the working tool by the same magnitude. The possible correction method is based on the use of images of the reference points obtained using charge-coupled devices (CCD cameras) associated with the full-scale model of the working tool (Fig. 4). Using a special software of the simulator interface, at least three reference points are selected on the image of the environment in the vicinity of the full-scale model of the working tool. The two-dimensional vectors of image positions of these reference points on the sensitive surfaces of each of the CCD cameras are magnitudes proportional to the projections of the vectors of the reference points on the plane x1, O, and x2 formed by the coordinate axes x1 and x2 of the coordinate system associ- JOURNAL OF COMPUTER AND SYSTEMS SCIENCES INTERNATIONAL Vol"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0000469_02640414.2016.1206207-Figure3-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0000469_02640414.2016.1206207-Figure3-1.png",
+        "caption": "Figure 3. The orientation of the hand. (a) The hand reference system and the sweepback angle (\u03a8) calculated considering the projection of the hand\u2019s displacement vector onto the xy plane of the hand (dp); (b) the attack angle (\u03b1) calculated considering a vector perpendicular to the hand\u2019s plane (n) and the velocity vector (v).",
+        "texts": [
+            "00 m (width), with 25 control points, was used. This frame was placed so that the X-axis represented the direction of the body\u2019s motion if not tethered, while the Y-axis represented the body\u2019s vertical direction and the Z-axis represented the body\u2019s lateral direction (Figure 2). Based on Sanders\u2019 (1999) study and to facilitate the visualisation of the landmarks, reflective tapes were placed around the distal ends of the third fingers, around the proximal ends of the second and fifth fingers, around the wrists of both hands (Figure 3(a)) and on the right greater trochanter. To improve the contrast between the tapes in relation to the environment, black lipstick was applied to outline them. To measure the force, a load cell (ZX 250 Alfa Instrumentos Eletr\u00f4nicos Ltda, up to 2500 N), previously calibrated, was used with the aid of a data acquisition system (Miotec Equipamentos Biom\u00e9dicos Ltda; 2000 Hz). This system allowed the synchronisation of the cameras and the force data with a trigger. With steel rings and carabiners, one end of the load cell was fixed to the pool wall, and the other end of the load cell was attached to a stiff 6",
+            " The hand\u2019s plane was defined by two vectors: (1) the wrist to the distal end of the third finger and (2) the proximal end of the second finger to the proximal end of the fifth finger. The cross product of these vectors defined a new vector (n). The velocity vector (v) was calculated by using the finite difference of the coordinate data of the mid-point between the four points, and these velocity data were also smoothed using the same procedures described previously. The attack angle (\u03b1) was calculated in accordance with Equation (1) (Figure 3(b)): \u03b1 \u00bc 90 cos 1 v vj j n nj j (1) where \u03b1 is the attack angle, v is the velocity vector of the hand and n is a vector perpendicular to the hand\u2019s plane. The sweepback angle (\u03a8) was calculated by considering the midpoint\u2019s displacement vector (d) and a local reference system of the hand (Figure 3(a)), where the y-axis represented the vector from the wrist to the distal end of the third finger, the z-axis was equal to the cross product between the y-axis and the vector from the proximal end of the second finger to the proximal end of the fifth finger and the x-axis was equal to the cross product between the z- and y-axes. The sweepback Figure 2. Overhead view of the experimental set-up; the thicker arrow represents the displacement of the participant\u2019s body if not tethered. D ow nl oa de d by [ L a T ro be U ni ve rs ity ] at 1 1: 29 1 1 Ju ly 2 01 6 angle was calculated in accordance with Equation (2) (Figure 3(a)): \u03a8 \u00bc tan 1 dpy dpx (2) where \u03a8 is the sweepback angle, dpx and dpy are the projection of the vector displacement onto the XY plane of the hand. To follow a counterclockwise direction (Figure 3(a)), if the result of the Equation (2) was negative, 360\u00b0 was added to the angle. The attack and sweepback angles were input into a Fourier series, as described by Sanders (1999), to calculate the force coefficients. Each force (drag force and the two components of the lift force) of both hands was calculated by using the following general equation (Sanders, 1999): Fdir \u00bc 1 2 \u03c1s2CvA\u00fe Da\u03c1A axj j (3) where Fdir is a force that acts in one direction, \u03c1 is the density of water (996 kg\u00b7m\u22123), s is the hand speed, Cv is a force coefficient that Sanders (1999) named as the velocity coefficient, A is the surface area of the palmar side of the hand,Da is a force coefficient that Sanders (1999) named as the acceleration coefficient and ax is the acceleration of the hand in the direction of its motion"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0003972_a:1020178613365-Figure1-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0003972_a:1020178613365-Figure1-1.png",
+        "caption": "Figure 1. The mutual reference frame: the orbit of the planet lies on the (x, y) plane; the x axis is along the mutual nodal line and is oriented towards the ascending mutual node, marked with asterisks in the figure.",
+        "texts": [
+            " In this framework there are two mutual nodes, the ascending and the descending one: they differ in the change of sign of the z component along the asteroid orbit (negative to positive in the first case and vice versa in the second). The mutual reference frame (x, y, z) is a heliocentric reference system such that the x-axis is along the mutual nodal line, directed towards the mutual ascending node; the y-axis lies on the planet orbital plane, so that the orbit of the planet lies on the (x, y) plane (see Figure 1). We shall use the further convention that the positive z-axis is oriented as the angular momentum of the planet. Let \u03c9M,\u03c9\u2032 M be the mutual pericenter arguments (the counter-clockwise angles between the x axis and the pericenters) of the orbit of the asteroid and of the planet respectively, and let IM be the mutual inclination between the two conics. We define as mutual elements the set of variables {a, e, a\u2032, e\u2032, \u03c9M,\u03c9\u2032 M, IM}. We can express \u03c9M,\u03c9\u2032 M, IM as functions of the Keplerian elements \u03c9, , I, \u03c9\u2032, \u2032, I \u2032, defined using a fixed heliocentric reference frame",
+            " We can study the case of only one perturbing planet: the perturbation of all the planets that we want to include in the model, up to the first order in the perturbing masses \u00b5i (the ratio between the mass of the planet i and the mass of the Sun), will be obtained by the sum of the contribution of each planet. From now on we shall consider a system of three bodies: Sun, planet, asteroid. The quantities related to the planet will be distinguished by a prime from the same quantities for the asteroid. In the mutual reference frame, as described in Figure 1, we can write the equations of the two osculating orbits of the asteroid and the planet P(u) :   x = a[cos\u03c9M(cos u \u2212 e) \u2212 \u03b2 sin\u03c9M sin u], y = a[sin\u03c9M(cos u \u2212 e) + \u03b2 cos\u03c9M sin u] cos IM, z = a[sin\u03c9M(cos u \u2212 e) + \u03b2 cos\u03c9M sin u] sin IM, (7) P \u2032(u\u2032) :   x\u2032 = a\u2032[cos\u03c9\u2032 M(cos u\u2032 \u2212 e\u2032) \u2212 \u03b2 \u2032 sin\u03c9\u2032 M sin u\u2032], y\u2032 = a\u2032[sin\u03c9\u2032 M(cos u\u2032 \u2212 e\u2032) + \u03b2 \u2032 cos\u03c9\u2032 M sin u\u2032], z\u2032 = 0, (8) where \u03b2 = \u221a 1 \u2212 e2, \u03b2 \u2032 = \u221a 1 \u2212 e\u20322 and u, u\u2032 are the eccentric anomalies of the asteroid and the planet, respectively. The positive distance D between a point on the orbit of the asteroid and a point on the orbit of the planet is defined by its square D 2(u, u\u2032) = \u2223\u2223P(u) \u2212 P \u2032(u) \u2223\u22232 = a2(1 \u2212 e cos u)2 + a\u20322(1 \u2212 e\u2032 cos u\u2032)2\u2212 \u2212 2aa\u2032{[(cos u \u2212 e) cos\u03c9M \u2212 \u03b2 sin u sin\u03c9M]\u00d7 \u00d7 [(cos u\u2032 \u2212 e\u2032) cos\u03c9\u2032 M \u2212 \u03b2 \u2032 sin u\u2032 sin\u03c9\u2032 M] + + cos IM[(cos u \u2212 e) sin\u03c9M + \u03b2 sin u cos\u03c9M] \u00d7 \u00d7 [(cos u\u2032 \u2212 e\u2032) sin\u03c9\u2032 M + \u03b2 \u2032 sin u\u2032 cos\u03c9\u2032 M]}"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0001235_b978-0-08-099425-3.00005-4-Figure5.27-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0001235_b978-0-08-099425-3.00005-4-Figure5.27-1.png",
+        "caption": "FIGURE 5.27",
+        "texts": [
+            " For both tractive and braking cases the relationship between longitudinal force and slip ratio is such that the wheel behaviour converges for slip ratios smaller than those at which peak force is produced. However for larger slip ratios the wheel behaviour diverges rapidly. For spin in particular, angular velocity increases very quickly until torque is reduced. The generation of lateral force and aligning moment in the tyre result from combinations of the same mechanisms and are thus treated together here. As a starting point it is helpful to consider Figure 5.27, which is adapted from the sketches for forces and torques provided by Olley (1945). Figure 5.27 is particularly useful for relating the sign convention for the lateral forces and aligning moments plotted and discussed throughout this chapter. From Figure 5.27 it can be seen that for a tyre rolling with a slip angle a at zero camber angle the lateral force generated due to the distribution of shear stress in the contact patch acts to the rear of the contact patch centre creating a lever arm known as the pneumatic trail. This mechanism introduces the aligning moment and has a stabilising or \u2018centring\u2019 effect on the road wheel. This is an important aspect of the steering \u2018feel\u2019 that is fed back to the driver through the steering system. Similarly it can be seen from Figure 5.27 that for a tyre rolling with a camber angle g at zero slip angle, the lateral force generated is called camber thrust. Due to the conditions in the contact patch the camber thrust acts in front of the contact patch centre creating a mechanism that creates a moment. Although this is referred to here as an aligning moment it has the opposite effect of the aligning moment resulting from slip angle and is sometimes called the camber torque as there is no resultant aligning action on the road wheel. The importance of the camber-induced moment is small for passenger cars but can be large for motorcycles"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0003684_20.996240-Figure2-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0003684_20.996240-Figure2-1.png",
+        "caption": "Fig. 2. Field distribution in an open circuit operation = 460 rpm, t = 63; 76 ms, and = 41 .",
+        "texts": [
+            " The movement can be imposed by an input velocity or evaluate by the mechanic equation of the system associated with Maxwell Stress Tensor [9], used to calculate the electromagnetic torque. In order to illustrate the electric circuit influence in the field calculation, some simple examples were analyzed. A permanent magnet generator feeding three different loads was simulated. The prototype characteristics are presented in Table I. This motor has nonskewed both the stator teeth and rotor magnets. First of all is shown the field distribution for an open circuit operation (Fig. 2). The rotor velocity is imposed at rpm. The instant shown is ms, corresponding to an angular displacement . In the second example, the generator feeds a balanced threephase load of 15 phase resistances in Y-connection. The rotor velocity is the same used in the open circuit case (Figs. 3 and 4). In this case, the electric machine is loaded by a three-phase full-wave rectifier with a resistive load of 11, 25, . Also here, the velocity used is the same as the previous examples (Figs. 3 and 4). These simple examples illustrate the need for a strong coupling between fields and circuit equations to better reproduce the influence of nonlinear sources and nonlinear electromagnetic devices in the behavior of the whole system, that way a more robust and trustworthy tool of electrical machine analyses may be developed"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0000439_icecube.2016.7495252-Figure3-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0000439_icecube.2016.7495252-Figure3-1.png",
+        "caption": "Figure 3. forces acting on inverted pendulum",
+        "texts": [
+            " . (g) Acceleration due to gravity 9.8 . (U) Force applied to the cart (x) Cart position coordinate (\u03b8) Pendulum angle from vertical (down) The force is applied which moves cart and stabilizes pendulum by making \u03b8 = 0. While designing mathematical model air resistance and pendulum cart friction have been ignored while it is assumed that the whole system is rigid. Following are the parameters of inverted pendulum used which are taken from [5]. The forces acting on inverted pendulum are shown in figure 3. Balancing the horizontal forces acting on cart + = (1) Summing the forces in the horizontal direction, we have the following expression for N. = + cos \u2212 \u0307 sin (2) Combining these equations we get. ( + ) + cos \u2212 \u0307 sin = (3) In order to find second equation we have to balance the perpendicular components of forces acting on pendulum. sin + cos \u2212 sin = + cos (4) The sum of moments around the pendulum is \u2212 sin \u2212 cos = (5) By comparing these equations we get. ( + ) + sin = \u2212 cos (6) Before introducing back-stepping controller we need an introduction of Lyapunov theory which is used to find stability of system and thus extensively used in nonlinear control design to evaluate stability of system"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0000114_0954406216640806-Figure1-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0000114_0954406216640806-Figure1-1.png",
+        "caption": "Figure 1. Schematic diagram of tube hydroforming (THF) with axial feeding: 1-left pusher; 2-left locating ring; 3- tubular blank; 4-deformed tube, 5-right locating ring; 6-right pusher; and 7-fluid hole.",
+        "texts": [
+            " The former is based on the geometries of wrinkles and plastic incremental theory, while the latter is based on the plastic mechanics and energy conservation law. In this section, an analytic formula to predict the useful wrinkles is derived based on the plastic incremental theory and the state of the strain increment d\"t in the thickness direction at the valley of a wrinkle. In this paper, the prediction method is named the geometry-based prediction method, or GPM. A tubular blank is located in two locating rings and deformed under both the hydraulic pressure P and the axial force Fy, this is the THF with axial feeding, as shown in Figure 1. Wrinkling may occur in the bulging zone lb of the tubular blank. If a wrinkle can be flattened by the following application of hydraulic pressure, it becomes a useful wrinkle; otherwise, it is a harmful one. A wrinkle generally evolves in three stages: onset, growing, and collapse.9 As shown in Figure 2, if the strain increment d\"t at valley V of the wrinkle in the thickness direction x maintains d\"t\u00bc 0 during the whole deformation, or the wall thickness at the valley neither increases nor decreases, the valley is in the state of pure shear strain",
+            " Research on bursting behavior in hydroforming of double-cone tube. Adv Mater Res 2011; 154: 678\u2013685. 11. Gao T, Zhang H, Liu Y, et al. The influence of lengthdiameter ratio in forming area on viscous outer pressure forming and limit diameter reduction. J Brazi Soc Mech Sci Eng 2015; 37: 1\u20136. 12. Gupta NK. Some aspects of axial collapse of cylindrical thin-walled tubes. Thin Wall Struct 1998; 32: 111\u2013126. at UNIV OF CINCINNATI on May 23, 2016pic.sagepub.comDownloaded from d0 (mm) initial outer diameter of a tubular blank (Figure 1) d\"e ( ) equivalent strain increment (equation (5)) d\"t( ), d\"y( ), d\" ( ) strain increments at valley V of the wrinkle in the thickness, axial and hoop directions, respectively (equation (5)) f (c/mm) fluctuation intervals of the hydraulic pressure per axial feeding (Figure 8) Fd (KN) axial force on the cross section at the wrinkle valleys (Figure 2) Ff (KN) friction force between tube and locating ring (Figure 2) FP, FP\u2019 (KN) generated forces on the pusher and the inner wall of deformed tube by the hydraulic pressure, respectively (Figure 2) Fy (KN) push force from the pushers (Figure 1) Fyc (KN) critical value of push force Fz from the pushers (Figure 1, equation (24)) hc(mm) critical geometry parameter for geometry-based prediction method (GPM) in this paper (equation (11)) hV, hF (mm) bulging height at the wrinkle valley and peak, respectively (Figure 2) l (mm) length of the straight yield line FV of the wrinkle (Figure 3) l0 (mm) Initial length of a tubular blank (Figure 1) lb (mm) bulging zone length of a tubular blank (Figure 1) lg (mm) locating zone length of a tubular blank (Figure 1) ly (mm) Projected length of the straight yield line FV of the wrinkle on y-axis (Figure 3) Mp (MPa/mm) plastic modulus per unit length (equation (12)) P (MPa) hydraulic pressure in tube (Figure 1) Pc (MPa) critical hydraulic pressure to get a wrinkle a useful one (equation (24)) S (mm) axial feeding distance on the end of the tube (Figure 1) t (mm) wall thickness at the wrinkle valleys (equation (2)) t0 (mm) initial wall thickness of a tubular blank (Figure 1) T total bulging time (Table 3) Ti (s) sampling time in total bulging time (Table 3) V (mm3) volume of the merging region between the peak and valley of a wrinkle (equation (17)) Wb, Wc, We (KJ) the bending strain energy of the yield line AB, plastic strain energy in hoop deformation, external work (equations (12), (15), and (18)) Wp, WFy (KJ) external work by hydraulic pressure P and external work WFy by axial force Fy, respectively, (equations (19) and (20)) Dh (mm) displacement of the valley of the wrinkle along z-axis in each time (Figure 3) P (MPa) fluctuation amplitude of the hydraulic pressure per axial feeding (Figure 8) V (mm3) volume difference when the yield line moves from FV to F\u2019V\u2019 (equation (19)) \"e( ) equivalent strain (equation (16)) ( ) Friction coefficient between the tube and the locating rings, 0"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0003645_i2002-10009-1-Figure8-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0003645_i2002-10009-1-Figure8-1.png",
+        "caption": "Fig. 8. Densities of edge dislocations of the slabs smectic layers.",
+        "texts": [
+            " In the language of dislocations, one would say that the curvature is relaxed by a density of infinitesimal edge dislocations [17] (a concept first introduced by Nye [27] for metals), whose line directions and Burgers vectors stand both parallel to the layers. This effect, which is present in a cholesteric, is already represented in eq. (4a) by the K\u0303-term. The second case specializes to a true TGBA phase (not a NL* phase). The (|\u03c31| = 0, |\u03c32| = 0) curvature of the slab is relaxed by a surface density |\u03c32|/d of edge dislocation lines perpendicular to the vortex lines, carrying an energy density of the order of Bdd2|\u03c32|/d = Bdd|\u03c32|, see Figure 8. Bd is a typical smectic elastic compression modulus, larger than Bp or Bb in the cases of interest. Therefore one expects that this term which conveys the anisotropic properties of the slabs does not play a role if Bdd|\u03c32| Bb, Bp, i.e. if the condition |\u03c32| q\u03ba2 2 is satisfied (we have used the relations qlbld = d \u2248 \u03be): |\u03c3| q\u03ba2 2. (6) The simplest DD is a circular cylinder, Figure 6c. The related developable domain is translation-invariant along the axis of the cylinder. Consider therefore an orthogonal section"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0003887_978-3-642-73890-6-Figure3\u00b71-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0003887_978-3-642-73890-6-Figure3\u00b71-1.png",
+        "caption": "Figure 3\u00b71: Fixturing layouts for four supports",
+        "texts": [
+            " Fixturing Layout Evaluation Excessive stress level induced at any location within the sheet metal part will cause the part to defonn plastically from its original shape. Three and four fIXturing elements at different mdii from the drilling point were considered. The stress distribution within the workpart for different layout configumtions , were investigated. In addition, several plate curvatures were examined. Specifically, nine different mdii of curvatures are investigated: \u00b125 in,\u00b150 in,\u00b175 in,\u00b1 100 in, and infmity. The fixture elements are located along the diagonal axes as shown in Figure 3\u00b71. The nonnal stresSes in 11, 22 directions and shear stress as a function of fixture support locatinn and the applied forces were analyzed. 1be simulations on yielding were performed for the following cases: flat plates. shells with positive mdius of curvature (convex shells) and shells with negative radius of curvature (concave shells). (19). Figures 3\u00b72, 3\u00b73111d 3-4 show the normal and shear stresses for the flat plate. The normal stresses in II and 22 direction at the loading point iDcIQIe willi the increasing radia1 distance of fixture elements from the center of the plate"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0003041_robot.1993.292161-Figure3-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0003041_robot.1993.292161-Figure3-1.png",
+        "caption": "Figure 3: Basic operation of the microchannel spatial light modulator (MSLM).",
+        "texts": [
+            " The inherently parallel nature of the optical calculation makes it especially well suited for handling large, highresolution workspaces. 2: Optical calculation of potential fields High speed analog optical computation of field maps suitable for robotic path planning can be performed using a microchannel spatial light modulator (MSLM). The MSLM is a device used to produce two dimensional optical images with gray-scale intensity values [9]-[Ill. Commercial versions of the MSLM [12] have been utilized in a variety of optical computing applications. 2.1: MSLM operation A diagram of the MSLM is shown in Fig. 3. The device consists essentially of a photocathode, a microchannel plate (MCP), and an elecuooptic crystal. An input light intensity distribution is converted by the photocathode to a spatial electronic current distribution, which is then amplified by the MCP and deposited on the crystal, typically LiNbO3. The electrooptic crystal is coated with a thin dielectric mirror. The deposited surface charge density, os, creates an electric field across the electrooptic crystal. The polarization of a coherent readout light beam which passes twice through the electmoptic crystal is rotated by this electric field via the Pockels effect [13], and the spatial distribution in os is converted to an amplitude distribution in the readout beam through the use of a polarizer"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0000924_red-uas.2015.7441001-Figure1-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0000924_red-uas.2015.7441001-Figure1-1.png",
+        "caption": "Fig. 1. Experimental platform",
+        "texts": [
+            " III. EXPERIMENTAL PLATFORM The aerial vehicle used to realize the experiments was a mini quadcopter. The attitude system is a 3DM-GX3 IMU by Microstrain, it sends \u03b8, \u03c6, \u03c8 attitude data and \u03b8\u0307, \u03c6\u0307, \u03c8\u0307 angular velocity, to pitch, roll and yaw angles respectively, these data was read through the device serial port. In the communication system was used a X-bee serial communication modem, the power supply is a variable voltage regulator of 30 A. All is integered on a Rabbit core RCM4300 device. The figure 1 show the physical platform described in this section. The system to get the position of the aerial vehicle is a Optitrack Flex 3 vision system by Natural Point, it has 12 cameras, a frame rate of 100 fps. It has a resolution of 640x480 VGA, this system can follow the markers located on the vehicle, just under millimetric movements. The information obtained is showed in the software Motive. In order to realize the optimal control law of finite horizon we are processing the data obtained from Motive interface the control law was calculated to track the trajectory in x and y axis, the control obtained is sent from the PC to the platform by Xbee modem"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0001100_icpeices.2016.7853099-Figure2-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0001100_icpeices.2016.7853099-Figure2-1.png",
+        "caption": "Fig 2: Phenomenological Model of TRMS",
+        "texts": [
+            " The control objective of SMC to make TRMS stable in significant cross coupling due to highly nonlinear system. III. MATHEMATICAL MODELlNG OF TRMS The Dynamic equation of TRMS system is obtained by using Newton's second law of motion. As TRMS model consists of two rotors, a thrust is created by bl ades which moves physical mass of the system. Due to laws of conservation, a rotor must experience a force opposite to the accelerated air. And is implemented using MA TLAB Simulink. A model of the system is followed by nonlinear static characteristics as shown in figure.2. The equation for vertical movement is expressed as: [2] 11, Yv = M1 - MFG - MByv - MG (1) Where the non linear static characteristic equation is: M1 = a1\u00b7 1'i + b1\u00b7 1' 1 (2) The gravity momentum equation is defined by: MFG = Mg.sinyv (3) The friction forces momentum are evaluated as, M - B . 0.0326 . 2 . 2 Byv 1yv' Yv - -2-sm Yv' Yh (4) and the gyroscopic momentum is, MG = kgy.M1\u00b7Yh\u00b7COSYv (5) The first order approximation of motor and electrical control circuit has been implemented in Laplace domain"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0002154_jsvi.1996.0870-Figure1-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0002154_jsvi.1996.0870-Figure1-1.png",
+        "caption": "Figure 1. Co-ordinate system and unit vectors.",
+        "texts": [
+            " In order to combine these two kinds of models together and smoothly switch the calculation of forces from the squeeze film model to the solid contact model and vice versa, a robust transition model is required. The development of a dynamic transition squeeze film model and its implementation into VIBIC are the objectives of this work. The present model extends an earlier primitive model [25]. The instantaneous squeeze film force, which is related to the instantaneous position, velocity and acceleration of the tube within its support hole, was formulated based on a 2p short length, cylindrical squeeze film model [7]. The force is resolved into normal and tangential components. As shown in Figure 1, the co-ordinates and directions are defined as follows: the origin point is located at the centre of the sleeve; the normal vector is a unit vector from the sleeve centre to the instantaneous tube centre; and the tangential vector is a unit vector perpendicular to the normal vector. Y- and Z-axes are the global Cartesian co-ordinates on which the tube position or motions in its support hole are described. The angle from the positive direction of the Y-axis to the normal direction is denoted as the instantaneous angular displacement C of the tube centre"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0002299_s0043-1648(96)07481-9-Figure2-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0002299_s0043-1648(96)07481-9-Figure2-1.png",
+        "caption": "Fig. 2. Test bearing details.",
+        "texts": [
+            " Load is applied downward to the bearing, by means of dead weights (6), applied through a linking mechanism (7) Journal: WEA (Wear) Article: 7481 (8) (9) (10). All seven wheels in the load mechanism are equipped with ball bearings, in order to reduce friction. The journal is driven by a 2 kW commutator motor (11). The power from the motor is transferred via a V-belt (12) to a support shaft (13), and further to the main shaft via a universal shaft (14) with flexible joints. The support shaft has a slipring system (15)mounted on its end, for transfering thermocouple signals from the rotating journal. Fig. 2 shows a section of the test bearing used.Thenominal diameter of the journal is 100 mm, the L/D ratio is 0.5, and the mean radial bearing clearance is 70 mm at room temperature. The bearing is equipped with 9 copper\u2013constantan thermocouples, marked TC1 to TC9, which measure bearing surface temperature at bearing midspan. The thermocouples are inserted in small bores drilled from the outside of the bearing, and fixed with adhesive. The tip of each thermocouple is placed within 0.5 mm under the bearing surface"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0001487_ire-i.1956.5007018-Figure2-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0001487_ire-i.1956.5007018-Figure2-1.png",
+        "caption": "Fig. 2",
+        "texts": [
+            "INE-COSINE ANGULAR POSITION ENCODERS By Carl P. Spaulding G. M. Giannini & Co., Inc. Pasadena, California I. SUIARY Sine-Cosine Angular Position Encoders are driven by a rack and pinion gear. (See Figure 2) useful in coupling an analog device to a digital However, with present practice it is possible to computer. The desired result is increased ac- build a sine-cosine angular position encoder that curacy of the analog equipment and simplification is considerably more accurate than an analog of the digital equipment. mechanism of the same size and inertia. An estimte is made of the accuracy that can These advantages, however, are gained at be obtained from sine-cosine angular position en- the cost of additional mechanical complexity of coders"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0000188_0954406214560420-Figure11-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0000188_0954406214560420-Figure11-1.png",
+        "caption": "Figure 11. Diagram of 4-UPS/UPR parallel mechanism.",
+        "texts": [
+            " The former method converts the fewer DoF mechanisms into the 6 DoF mechanisms. It extends the workspace of the fewer DoF mechanisms. The latter one restricts the motion of the mobile platform in the valid workspace strictly through the modified Jacobian matrix. The FK problem of a 4-UPS/UPR mechanism is taken as an example to illustrate the above two methods. at Univ Politecnica Madrid on January 14, 2015pic.sagepub.comDownloaded from 4-UPS/UPR mechanism is a typical coupled fewer DoF mechanism. As shown in Figure 11, 4-UPS/ UPR mechanism includes four driving limbs (UPS) and an intermediate constraint limb (UPR). Coordinate system o-xyz is attached to the fixed platform. Axis xo, yo is parallel to a3a4, a4a1, respectively, and zo is perpendicular to the platform. Coordinate system p-xyz is attached to the mobile platform, and its initial orientation is parallel to coordinate system o-xyz. The revolute (R) joint of UPR limb is attached to the mobile platform, and its axis is parallel to axis yp. UPR limb determines the motion characters"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0001473_978-3-319-09834-0_10-Figure10.1-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0001473_978-3-319-09834-0_10-Figure10.1-1.png",
+        "caption": "Fig. 10.1 Schematic drawing of a BDD microelectrode. Copyright (2007) American Chemical Society",
+        "texts": [
+            " The surface morphology and crystalline structure of the BDD thin film was determined using scanning electron microscopy (SEM, JOEL JSM 5400) and Raman spectroscopy (Renishaw System 2000). The tungsten wire coated with the BDD thin film was then connected to the coated metal wire using silver paste and then dried. This BDD wire was then inserted into the capillary for insulation through a pre-pulled glass capillary (using capillary puller, Narishige, Tokyo, Japan) followed by resin infusion. Resin was soaked up by capillary action. After drying overnight, the fabrication of BDD microelectrode reached completion. The details are illustrated in Fig. 10.1. CF electrodes were prepared via the same procedure as shown in Fig. 10.1. The electrode length can be easily controlled by adjusting its tip until it had the same length than the BDD microelectrode. A SEM image of the fabricated BDD microelectrode shows that the tip diameter is distinctly small (about 5 \u03bcm) with the polycrystalline diamond grain size being approximately *2 \u00b5m (Fig. 10.2a). The tip size was almost the same as the conventional CF electrode used in this experiment shown as example (Fig. 10.2b). Based on the experimental trial for in vivo measurement the average tip length was set at 250 \u00b5m (Fig"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0003787_iros.1996.570634-Figure9-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0003787_iros.1996.570634-Figure9-1.png",
+        "caption": "Figure 9: Position Control (I)",
+        "texts": [
+            " In the case of the feedback control, 8 converged to the desired direction, and the link barely rotated at the end of the trajectory. Fig.7 shows the result when the initial error of 8 was O.l(rad). The length of the trajectory was 0.3(m). The feedforward control and feedback control were compared also in the rotational trajectory (ii) from 0 = 0.524 to 19 = 1.571 (Fig.8) 8 came closer to the desired angle by the feedback control. Next, position controls by the composite trajectc ries were performed. Fig.9 shows positioning of the free link from [0.35, -0.15, 0.7851 to [0.35, 0.1, 0.7851. The black circle represents C.I., and the white circle represents the passive joint. The feedforward control caused large error of the link motion from the planned trajectory, and the link rotation at the end of the trajectory (b). The final configuration was [0.349, 0.100, 0.8211. Though 5 and y coordinates of the passive joint were controlled exactly, 19 deviated from the desired configuration. On the other hand, the free link almost halted nem the desired configuration by the feedback control (c)"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0000267_iccke.2014.6993354-Figure2-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0000267_iccke.2014.6993354-Figure2-1.png",
+        "caption": "Fig. 2. DDMR free body diagram",
+        "texts": [
+            " In addition, R is radius of wheel. The heading of mobile robot and center position can be obtained by subsequent equations: 2/))(cos( 21Rx (2) 2/))(sin( 21Ry (3) LR 2/)( 21 (4) And velocity of robot can be accessed by: cosvx (5) sinvy (6) Therefore, the general matrix for location and heading is like as following: v y x 10 0sin 0cos (7) Finally, kinematics model of DDMR is acquired as: l r LRLR RRv 2/2/ 2/2/ (8) Generally, in dynamics model, the torque needed for moving to particular position will be computed. Fig.2 exemplifies the FBD (free body diagram) of robot wheel which is subjected to moment on the robot center due to wheels rotation. By solving the dynamic equilibrium conditions, the required torque will be acquired as bellow: 2 2 2 2 1 2 2 1 1 33 rr l l rmMl l rmM (9) 2 2 2 2 1 2 2 1 2 33 rr l l rmMl l rmM (10) In this part, conventional tuned PID controller was integrated by AFC technique. The target of this hybridization is that stability of system will be increased when faced to internal and external disturbances [10]"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0001627_8.480-Figure4-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0001627_8.480-Figure4-1.png",
+        "caption": "FIG. 4.",
+        "texts": [
+            " The steps in their application are as follows: (1) Calculate the \"shaft\" critical in the usual way (assuming no coupling) with the formula \u2022\u2014 (r.p.m.)s?iafl \u2014 o) shaft oU V k shaft _ W-L shaft * shaft Ij 1 /%shaft (a) in which Fs are moments of inertia in lbs.in.sec.2, k's are stiffnesses in in.lbs./radian (as shown in Figs. 3 and 4) and p is the ratio of the propeller and engine speeds. This ratio p is usually smaller than one; for spur gears it is negative: p = \u2014r2/rh while for a planetary system it is positive p = R/R + r. For a 2 to 1 planetary system the arrangement of Fig. 4 is replaced by bevel gears, but the analysis still holds; the formulas may be applied with p = 0.5. The last term in the denominator of Eq. (a) comes in only if the central sun gear of the planetary system is coupled to the frame with some flexibility. In most constructions this coupling is stiff and in such cases ki = \u00b0o and the term simply disappears. (2) Calculate (or estimate) the \"frame\" critical: \u2014 (r.p.m.)/m7;ie V k \"'mount Icvl. (b) in which IcyL is the moment of inertia of the stationary parts in which the pistons and upper rod parts are supposed to be frozen in the mid-stroke position",
+            " (3) by the term containing the symbol T for the tooth pres sure, which is positive if tending to turn both gears clockwise. With these notations, Eqs. (1), (2), and (3) are the Newton equations for the three inertias; the fourth and fifth equations express the assumption that the gears have no inertia, so that the elastic torque from the shaft on each gear equals the counter torque due to tooth pressure. Finally Eq. (6) expresses a geo metric relation regarding the gear rotations. The unknowns in this set are the five <p's and the pressure T. For an engine with planetary gears as illustrated in Fig. 4 the equations are very similar. Since the sun gear is assumed non-rigid, but rather attached to the frame through a flexibility &4, there is one more equa tion than in the previous case: I\\<pi + ki(<pi \u2014 (pi) = 0 (8) I2(p2 + k2((p2 \u2014 (Pb) ~ Irec.W = 0 (9) IzW + h<P3 + ki((pS - (pQ) ~ Irec.2<P = 0 (10) Tr + h((ps - <?6) = 0 TR + k2((p2 - (p^ = 0 T(R + r) - hfa - (Pi) = 0 ((Pi \u2014 (pb)R + (^i \u2014 (po)r = 0 The ratio of the propeller and engine speeds is R/R + r. The physical meaning of the various equations is very much the same as discussed with the expressions (1) to (6) for spur gears"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0000212_1350650115576943-Figure2-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0000212_1350650115576943-Figure2-1.png",
+        "caption": "Figure 2. (a) 3D visualization of the out-of-contact lubricant channeling by the slider; (b) schematic representation of the outof-contact lubricant channeling by the slider.",
+        "texts": [
+            " The micro amounts of oil lubricant for starved conditions were carefully measured and injected on the track by a digital pipette with high accuracy. In all experiments, before starting the actual measurements, the contact has been operated for 2\u20133min to ensure a uniform and homogeneous distribution of lubricant on the tracks of ball and disc. The flow of lubricant through the limited slot is incompressible (lubricant density \u00bc const) and isoviscous. Therefore, the modification of flow is limited to the geometric distribution of lubricant on the surface of ball. As shown in Figure 2(a), in the outlet zone of Hertzian contact, lubricant is pushed aside by the rotating ball and the depleted track is banded by two layers of lubricant on sides. Introducing a mechanism (the slider with confined exit) to scrape the displaced lubricant imposes a modification of at WEST VIRGINA UNIV on June 28, 2015pij.sagepub.comDownloaded from lubricant flow. Figure 2(a) depicts, by means of 3D visualization, how the channeling of lubricant changes the widths G1 and G2 and the thicknesses t1 and t2 of the lubricant layers. Where G1 and t1 represent, respectively, the width of the depleted track and the average thickness of displaced lubricant layers, while G2 and t2 represent, respectively, the width of the depleted track and the average thickness of channeled lubricant layers. Such flow of lubricant is subjected to the mass conservation law (continuity) under steady-state condition. Thus, the two-dimensional incompressible flow in Figure 2(b) can be expressed mathematically as the following @Qx @x \u00fe @Qy @y \u00bc 0 \u00f01\u00de where Qx, Qy are mass flow rates per unit of time in x, y directions. By neglecting the flow in y direction, equation (1) is reduced to @Qx @x \u00bc 0, Qx \u00bc const \u00f02\u00de Equation (2) shows that the mass flow rate is constant in the direction of sliding. Therefore, one can conclude that the amount of lubricant before the slider (Qx1) is equal to the amount of channeled lubricant after the slider (Qx2). Also, the speed of lubricant does not change due to the channeling, where the speed of lubricant before the slider (Ux1) and the speed of lubricant after the slider (Ux2) are equal. In Figure 2(b), b1 and b2 represent the width of displaced and channeled lubricant layers, respectively. The channeled lubricant after passing the slot of the slider has a depleted gap of width (G2) which is less than the width of the depleted track (G1) before the channeling. On the other hand, the width (b2) of banding layers for channeled lubricant is less than the width of displaced lubricant layers (b1) for nonchanneled lubricant. However, the algebraic sum of G2 and 2*b2 is in the scale of the slot width. By assuming the symmetric distribution of lubricant on the surface of ball as shown in Figure 2(b), the following simple equation can be accepted Qx1 \u00bc 2 b1t1Ux1 Qx2 \u00bc 2 b2t2Ux2 ! Qx1\u00bcQx2, Ux1\u00bcUx2 t2 \u00bc b1t1 b2 \u00f03\u00de at WEST VIRGINA UNIV on June 28, 2015pij.sagepub.comDownloaded from From equation (3) it is clear that t2> t1 since b2< b1 regardless of the values of gaps G1 and G2. However, the geometry and distribution of lubricant for natural and channeled flow depend on the viscosity of lubricant, surface tension forces, and dimensions of slot. Indeed, a slot with a small width makes the channeled lubricant approaching more the centreline of depleted track but the entrained amount through the slot could be less, particularly, for oils with high viscosities or greases due to the block of flow through the slot"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0003070_1.1468862-Figure5-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0003070_1.1468862-Figure5-1.png",
+        "caption": "Fig. 5 \u201ea\u2026 A portion of the switching surface together with the switching curve; \u201eb\u2026 the corresponding portion of the slab together with the sets defining behavior of CPTO controller",
+        "texts": [
+            " Since, in actuality, ideal time-optimal control is unlikely, more practical controllers that are nearly time-optimal and that exhibit good behavior, particularly near the origin of the phase space, should be considered. We, therefore, introduce a CPTO controller where the ideal switching surface for the third-order system becomes a \u2018\u2018slab\u2019\u2019 of finite thickness in the neighborhood of the ideal switching surface, and the ideal switching curve becomes a tube in the neighborhood of the ideal switching curve @8#. The tube, lying within the slab, encloses the origin, as in Fig. 5. The form of the CPTO control law is inspired as follows: Similar to the case of double integrator plant, first, we try a continuous controller U8'U*, where U8~x !5sat$k1@x12X1~x2 ,x3!#%. (37a) We note from ~34!\u2013~37! that when x2>0, x3>0, and x1 ,1/k1 , U8(x) becomes U8(x)>k1x1 . Substituting u5U8(x) Transactions of the ASME x?url=/data/journals/jdsmaa/26301/ on 07/30/2017 Terms of Use: http://www.asme.org/about-asme/terms-of-use Downloaded F 5k1x1 into the state equations ~32!, we obtain x\u0302 11Kk1x150",
+            ", ~x12X1>e12T2 if T2P~2e1 ,e1!!, (40) ~x12X1>0 if T25e1!% S25$xPR3:~x12X1<0 if T252e1!, ~x12X1<2e12T2 if T2P~2e1 ,e1!!, (41) ~x12X1<22e1 if T25e1!% S5$xPR3:~0,x12X1,2e1 if T252e1!, ~2e12T2,x12X1,e12T2 if T2P~2e1 ,e1!!, (42) ~22e1,x12X1,0 if T25e1!% T5$xPS:x12X11k2~x22X2!1e1 sat~k1k2x3!P~2e1 ,e1!%, (43) L5$xPT:x12X11k2~x22X2!1k3x3P~2e1 ,e1!%, (44) We note that: R35S1\u00f8S\u00f8S2 ; T,S; L,T; S1 and S2 are closed; S, T, and L are open. The control operation takes place as follows. Without loss of generality, we assume K51. From Fig. 5, note that T,S and T divides S into two portions; one portion is on the right side of the tube, denoted SR , and the other portion is on the left side of the tube, denoted SL . The switching surface V1 forms the upper JUNE 2002, Vol. 124 \u00d5 257 x?url=/data/journals/jdsmaa/26301/ on 07/30/2017 Terms of Use: http://www.asme.org/about-asme/terms-of-use Downloaded F boundary ~in positive x3 direction! of SR , the lower boundary of SL and passes through T. That is, V1 is not centered within the slab. We know from the time-optimal control theory that V1 is reached in finite time from any finite state xPS1 when u(t) [11"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0000710_b978-0-08-100072-4.00007-1-Figure7.13-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0000710_b978-0-08-100072-4.00007-1-Figure7.13-1.png",
+        "caption": "Figure 7.13 The architecture of a wireless capsule containing temperature and pH sensors at the front, followed by application-specific integrated circuit (ASIC) and batteries. E.A. Johannessen, L. Wang, C. Wyse, D.R. Cumming, J.M. Cooper, Biocompatibility of a labon-a-pill sensor in artificial gastrointestinal environments, IEEE Transactions on Biomedical Engineering 53 (2006) 2333\u20132340.",
+        "texts": [
+            " The system contains several sensors with analog readout circuits, digital microcontroller 167 W ireless biosensors for PO C m edical applications circuits, a radio transmitter, and a battery. The sensors convert the physiological parameters to electrical parameters. The controller circuits manage and process all the sensor data. The data from the GI tract is transmitted to an external device for monitoring and recording. The main feature of the system is the integrated multisensor of a pH, pressure, and temperature sensor for real-time signal monitoring of the GI tract abnormalities. The architecture of a wireless capsule system named the lab-in-a-pill (LIAP) is presented in Fig. 7.13 [113]. It consists of pH and temperature sensors and a custom-made application-specific integrated readout circuit. The pH sensor is a microfabricated ISFET with Ag/AgCl reference electrode. The temperature sensor is an n-channel silicon diode. The system consumes 15.5 mW. The circuit has a power saving feature to operate it for 42 h. A wireless capsule can provide an invasive method for the diagnosis of the GI tract. The sensor systems for a wireless capsule are composed of mechanical sensors for pressure and position measurement, chemical sensors for pH, conductivity, and dissolved oxygen measurement, and biosensors for bleeding and pathogens detection"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0000457_978-81-322-2671-0_78-Figure1-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0000457_978-81-322-2671-0_78-Figure1-1.png",
+        "caption": "Fig. 1 H-airframe structure design (all dimensions are in cm)",
+        "texts": [
+            " proposed a PID control system with Ziegler-Nichols tuning method for the longitudinal and lateral directional dimensions, which include angular rate, attitude, altitude, and navigation control which are able to stabilize an unstable system or to improve the system response [7]. This paper is organized as follows; Sect. 2 explains the various subsystem of UAV quadcopter design. Section 3 explains the functional block diagram of proposed UAV transreceiver. Simulation and experimental results are discussed in Sect. 4 followed with the conclusion. Airframe of quadcopter uses H-shape structure and mounting is provided to fix Brush Less DC (BLDC) motor at the end. Figure 1 shows the construction of proposed H-shape airframe structure with 4 BLDC motor. Two motors M1and M3 will rotate in clockwise direction, while the other two motors M2 and M4 rotate in anticlockwise direction. Two motors adjacent to each other are always in the opposite direction of rotation. Thrust produced by motors should be twice that of the total weight of the quadcopter. If the thrust generated by the motors is too little, the quadcopter does not take OFF. However, if the thrust is more than the design level, the quadcoptor might become too nimble and hard to control"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0000709_012002-Figure3-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0000709_012002-Figure3-1.png",
+        "caption": "Figure 3. Tooth profile in normal section. Figure 4. Teeth generating surface.",
+        "texts": [
+            " To ensure continuity of the transfer function and to improve the performances of precessional transmission under multiplication it is necessary to modify teeth profile with the diagram error value \u03943 by communicating supplementary motion to the tool. In this case the momentary transmission ratio of the manufactured gear will be constant. Usually, in theoretical mechanics the position of the body making spherical-spatial motion is described by Euler angles. The mobile coordinate system OX1Y1Z1 is connected rigidly with the satellite wheel, which origin coincides with the centre of precession 0 (figure 3) and performs spherical-spatial motion together with the satellite wheel relative to the motionless coordinate system OXYZ. The elaboration of the mathematic model of the modified teeth profile is based integrally on the mathematic model of teeth profile, previously developed by the authors. With this purpose it is necessary to present the detailed description of teeth profile without modification and, then, to present of the description of modified profile peculiarities. 3.Description of teeth profile designed on sphere"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0001619_004051755002000402-Figure4-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0001619_004051755002000402-Figure4-1.png",
+        "caption": "FIG. 4. Diagrant showing arrangement used for registering deflection of film resulting from the impact of a waterdrop.",
+        "texts": [
+            " The shivering of the drop into droplets occurs much later, when the whole drop has become flat. The second drop, which falls on a wet fabric, loses its spherical shape more quickly because the contact angle is lower. From the calculations of the pressure we get an expression for the deflection of the material during the impact as a function of time. This deflection can also be determined experimentally. A small strip of paper was fastened to the lower side of the acetate film, and the film was fixed in the frame. The waterdrop was adjusted to fall just above the strip (see Figure 4). The movement of the strip during the impact in relation to a fixed scale was photographed with the camera described above. From the photographs the deflection could be measured as a function of the time. The result is given in connection with the mathematical treatment of the deflection. Mathematical Treatment of Waterdrop Impact Definition of Symbols E = modulus of elasticity g = acceleration of gravity h = height of the deformed drop p = pressure between the drop and the foundation po = pressure required to force water through fabric \u2019 P = total force due to drop = ~ X area y = radius of the drop s = depth of penetration of water into fabric t = time at EMORY UNIV on April 19, 2015trj"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0002220_s0263574797000544-Figure1-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0002220_s0263574797000544-Figure1-1.png",
+        "caption": "Fig . 1 . BIPMAN .",
+        "texts": [
+            " The purpose of our study is also to contribute to a better understanding of the dynamic behaviour of a standing man facing specific tasks . More generally , we look for \u2018\u2018optimal\u2019\u2019 strategies adapted to body motion under perturbation ef fects . In this way , we aim at determining the optimal distribution of the forces exerted by the limbs on the trunk of a biped structure with respect to a specified task . 12 , 13 This mechanical structure is composed of 11 links with 12 joints pneumatically actuated and has been called \u2018\u2018 BIPMAN \u2019\u2019 , acronym for \u2018\u2018 BI omechanical and P neumatic MAN \u2019\u2019 (Figure 1) . This paper is organized as follows : In section 2 we briefly recall the overall control architecture and especially the mathematical formulation of the optimization problem . In section 3 , we present the results of simulations with respect to dif ferent optimization criteria and constraints . According to these results , the RTCA approach is illustrated in section 4 for biped dynamic stability under external perturbations . The control architecture is composed of three levels : The upper level , called the \u2018\u2018 Super y isor \u2019\u2019 level determines the general behaviour of the biped , according to the environment and to the specifities of the task to be performed "
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0000919_epepemc.2016.7752145-Figure3-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0000919_epepemc.2016.7752145-Figure3-1.png",
+        "caption": "Fig. 3. Rotors with different configurations of drilled rotor bars with schematic.",
+        "texts": [
+            " ( )1Sd Sd mR S Sd S mR Sq S S di v di i i dt R dt \u03c3\u03c4 \u03c3\u03c4 \u03c9 \u03c3 \u03c4+ = + \u2212 \u2212 (11) ( )1Sq Sq S Sq S mR Sd S mR mR S di v i i i dt R \u03c3\u03c4 \u03c3\u03c4 \u03c9 \u03c3 \u03c4 \u03c9+ = \u2212 \u2212 \u2212 (12) 23 2 m el mR Sq R LT p i i L = (13) Fig. 2 shows the block scheme of the simplified model of IM with broken rotor bars resulting from the previous equations. The proposed model has been tested on a machine, whose data are given in the Appendix. The machine has at its disposal multiple interchangeable rotors, whose bars have been drilled in order to emulate breakage of individual bars (Fig. 3). In following tests, eight rotors have been used, from a healthy one up to the one with seven broken bars. Although the situation with rotors with more broken bars is hardly to be encountered in real life, this heavy fault has been observed in order to demonstrate the consistency of the whole approach. Fig. 4 shows the schematics of the setup. Apart from the tested IM, another machine serves as a controlled load. At each individual experiment the machine has been supplied from the grid-like voltage at 50 Hz, while the stator current has been measured and its spectrum calculated at steady state"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0003462_1.2096645-Figure1-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0003462_1.2096645-Figure1-1.png",
+        "caption": "Fig. 1. Working surface of Clark electrode",
+        "texts": [
+            " The major problem limiting the widespread application of this technology is the incompatibility with whole blood and the difficulties involved in long term in vivo applications. An additional problem is that glucose concentrations in blood can exceed the oxygen concentration. The sensor response can then be sensitive to oxygen as well as glucose. Materials and Methods The pH 7.00 buffer is 0.05M phosphate with 0.1M sodium chloride. Aqueous test solutions are prepared in this buffer. Whole blood samples were spiked with glucose and were assayed with a Yellow Springs Instruments Model 23A glucose analyzer. BIOSTATOR | (Miles Inc.) electrodes (Fig. 1) have been used for the experimental sensors. They consist of a central platinum disk (2 mm diam) surrounded by a broad silver ring (8 mm diam). Epoxy isolates the platinum from the silver and holds both in a stainless steel shell. These * Electrochemical Society Active Member. 1 Present address: Dow Coming STI, Kendallville, Indiana 46755. electrodes were modified by machining the platinum and silver surfaces from the original hemispherical surface to a planar surface. The curved surfaces originally present made it very difficult to apply uniform layers to the sensor surface",
+            " A number of general reviews that cover fuel-cell theory and technology have been published (1-3). More specific to the methanol fuel cell (MFC), Baker (4) has addressed hydrocarbon-fuel-cell progress prior to 1965. Hampson and Williams have recently reviewed methanol-oxidation mechanisms pertinent to the MFC (5), and Weeks et al. have * Electrochemical Society Active Member. examined various carbon-containing porous electrodes for a methanol fuel electrode (6). A schematic illustration of an MFC is shown in Fig. 1. The function of the solid polymer electrolyte (SPE) membrane is to transport hydrogen ions that are generated at the anode by the following overall reaction H20 + CH~OH --> CO2 + 6H + + 6e- [1] The hydrogen ions are consumed at the cathode to complete the electrical circuit. For complete reduction 3 6e- + 6H + + - - 02 --~ 3H20 [2] 2 The resulting overall cell reaction can be obtained by adding the two half reactions 3 - - 02 ~- CH3OH -+ CO2 + 2H20 [3] 2 ) unless CC License in place (see abstract)"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0001815_s003390051131-Figure1-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0001815_s003390051131-Figure1-1.png",
+        "caption": "Fig. 1. Schematic diagram showing implementation of feedback",
+        "texts": [
+            " In this paper we explain why the implementation of such feedback systems in AFM and interfacial force microscopy (IFM) is far from trivial, discuss the necessary re- quirements to truly enhance the stiffness of the force sensor and present preliminary results demonstrating the feasibility of the system. The feedback scheme consists of a force-sensing lever which converts the force into a displacement. This displacement is detected and converted into a compensating force according to the schematic shown in Fig. 1. The response of the force sensor is described by the complex transfer function Gl(\u03c9). Disregarding higher order modes, one has Gl(\u03c9) = C\u03c92 0/(\u03c9 2 0\u2212\u03c92+ i\u03c9\u03c90/Q), where C is the stiffness of the lever and \u03c90 is the resonance frequency of the free lever. For an ideal feedback amplifier with a gain of K the effect of the feedback is to increase the lever stiffness by a factor (K +1) and hence to shift the resonance frequency by a factor of \u221a (K +1). The new resonance frequency corresponds in essence to the frequency at which the loop gain |Gl(\u03c9k)|\u2217K is equal to 1"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0002388_20.706517-Figure2-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0002388_20.706517-Figure2-1.png",
+        "caption": "Fig. 2. (a) Instantaneous normalized source (thin full) and eddy currents (bold), e.m.f. (dashed). (b) Instantaneous normalized magnetic flux density at the three marked nodes in Fig. 1-b, both for the case Wn = 3.",
+        "texts": [
+            " The results were obtained for B, = 0.5 T, H,, = 478.8 k\u2019. The space discretization adopted for Wn = 3 is represented in Fig. 1-b. The adopted space discretizations are characterized by a mesh refinement for greater values of Wn . ..\u2018L... ,,I j$$Ji ... .. .. .. ... ......... . . A complete understanding of the problem may be achieved by studying the field dependence on the space and time. In this way, the presented results describe: - The normalized source current, i s , and the eddy currents ie , as a function of 6 , Fig.2-a (1 1) where Hn, and Hn, are, respectively, the normalized magnetic field strength (Hn=H/Hov) on the subpath s2 and on the sample axis 0 (Fig. 1-a); - The instantaneous normalized magnetic flux density (Bn=B/Bov) for the three nodes indicated in Fig.1, Fig.2-b; - The instantaneous eddy current lines at the instants 5 =0.2,0.4,0.6, Fig.3; - The magnetic hysteresis curves for two nodes, one on the subpath s2 and the other on the sample axis, Fig.4-a. i, = Hn, , i e = Hn, - Hn, , 5 = u t / n Fig. 3. Instantaneous eddy current lines occurring at the instants = 0.2 (a), 0.4 (b), 0.6 (c), for the case W,, = 3. Finally, the losses, as a function of Wn (lo), are represented in Fig.4-b, decomposed into eddy current losses and hysteresis losses, both normalized by Hzv I c "
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0002865_978-94-015-9064-8_16-Figure1-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0002865_978-94-015-9064-8_16-Figure1-1.png",
+        "caption": "Fig. 1: Double-Triangular planar manipulator",
+        "texts": [],
+        "surrounding_texts": [
+            "The kinematic structure of most contemporary robots is an open kinematic chain structure (known also as serial manipulators). Only relatively few commercial robots are composed of a closed kinematic chain (parallel) structure. However, the increasing interest in parallel robots points to the potential embedded in this structure which has not yet been fully exploited. The advantages of parallel robots as compared to serial ones are: \u2022 higher payload-to-weight ratio since the payload is carried by several links in parallel, \u2022 higher accuracy due to non-cumulative joint error, \u2022 higher structural rigidity, since the load is usually carried by several links in parallel and in some structures in compression-traction mode only, \u2022 location of motors at or close to the base, \u2022 simpler solution of the inverse kinematics equations. Conversely, they suffer from smaller work volume, singular configurations and a more complicated direct kinematic solution (which is usually not required for control purposes). Different structures of parallel manipulators are given, for example, in [Hunt, 1983; Innocenti and Parenti-Castelli, 1994; Lin et aI., 1992; merIet, 1994; Pierrot et aI., 1991; 155 J. Lenarcic and M. L. Husty (eds.), Advances in Robot Kinematics: Analysis and Control, 155-164. @ 1998 Kluwer Academic Publishers. Tahmasebi and Tsai, 1995; Ben-Horin and Shoham, 1996; Waldron et aI., 1989]. A comprehensive atlas of parallel robots has been composed by Merlet and can be found in the web site [http://www.inria.fr/prisme/ personnel! merletl merleCeng.html]. This paper presents a new type of a parallel robot that is a modification of and an extension to six-Degrees-Of-Freedom (DOF) of the three-DOF planar parallel robot presented by Mohammadi, Zsombor-Murray and Angeles,[1993]. The structure presented in the above mentioned paper is of two triangles, one stationary and one moveable, connected at three points, one at each side, by a combination of revolute and prismatic joints. The prismatic joints are actuated while the revolute joints are passive. This constitutes three-DOF of planar motion - two translational and one rotational - of the movable triangle (see Fig. I). Observing the work envelope of this structure, it seems that the area covered by the moveable triangle's center (output link), especially when a combination of translational and rotational motion is required, is relatively small. Fig. 2 depicts the work volume of the Double-Triangular manipulator for a translational motion and for a combination of translational and rotational motion. If, for example, a rotational motion of 55\u00b0 is required, then the useful work envelope becomes too small. In this paper a modification of this structure is suggested that increases considerably the planar work envelope. The planar manipulator is described in the next section and a combination of two such designs which enables construction a six-OOF parallel robot, is described in section 3."
+        ]
+    },
+    {
+        "image_filename": "designv11_60_0000691_0954410016664926-Figure1-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0000691_0954410016664926-Figure1-1.png",
+        "caption": "Figure 1. Comparison of the dimensions between the baseline tilt-rotor TR-40 and its WE variant.2",
+        "texts": [
+            " Although the tilt-rotor can take off and land without a runway, the endurance and range of the tilt-rotor are inferior to those of other fixed wing aircraft. In order to increase these performance aspects of the tilt-rotor aircraft, a concept of nacelle mounted wing extension (WE), previously called as a nacelle-fixed auxiliary wings (NFAW), has been studied by the Korea Aerospace Research Institute (KARI).1,2 Similar research also has been conducted to improve the performance of the tiltrotor by NASA.3 The comparison of the geometric configuration of the TR-40 original tilt-rotor and its WE variant is shown in Figure 1.2 In the early stage of the Smart unmanned aerial vehicle (UAV) development program, the design of the control law for the tilt-rotor aircraft was identified as the riskiest element for the full scale development. In order to reduce the risk of the control law design, the 40% scaled tilt-rotor aircraft, named TR-40, was developed as a flight demonstrator to prove the control law of the full-scale Smart UAV, named TR-100. The development of the flight demonstrator was based on the conventional proportional, integral, and 1Future Aircraft Systems Division, Korea Aerospace Research Institute, Daejeon, Republic of Korea 2Department of Aerospace Engineering, Korea Advanced Institute of Science and Technology, Daejeon, Republic of Korea 3Research Institute of Marine Systems Engineering, Seoul National University, Seoul, Republic of Korea 4Department of Aerospace and System Engineering, Gyeongsang National University, Gyeongnam, Republic of Korea Corresponding author: Min-Jea Tahk, Department of Aerospace Engineering, Korea Advanced Institute of Science and Technology, Daejeon 305-701, Republic of Korea"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0000379_012035-Figure1-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0000379_012035-Figure1-1.png",
+        "caption": "Figure 1. \u201cRugby 700\u201d hohlraum with higher case-to-capsule ratio than \u201ccylinder 575\u201d but 10% larger surface.",
+        "texts": [
+            "art of the NIF 2014-2015 High Foot Campaign was carried out with the same \u201cRugby 700\u201d hohlraum (figure 1, waist diameter 7.0 mm) used in 2013 for 3 Low Foot shots [1]. Due to the high neutron yield obtained with the high foot pulse in a cylindrical hohlraum [2], it was of interest to test this laser pulse in a parabolic hohlraum geometry such as the rugby one. Content from this work may be used under the terms of the Creative Commons Attribution 3.0 licence. Any further distribution of this work must maintain attribution to the author(s) and the title of the work, journal citation and DOI. Published under licence by IOP Publishing Ltd 1 Moreover the last low foot rugby shot, N131011, which was the first shot in the CEA/LLNL rugby collaboration, showed that repointing the outer cones inwards by 500 \u00b5m could effectively achieve a round hot spot (figure 2) in a 1"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0001945_ip-cta:19982047-Figure7-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0001945_ip-cta:19982047-Figure7-1.png",
+        "caption": "Fig. 7 Two-link manipulator",
+        "texts": [
+            " A critical point arises due to the mutual influence of K, and Voi upon each other, as can be seen in eqn. 16. This is resolved simply by following an iterative procedure defined below. (i) Select a V o / ~ j > I initial guess; (ii) Evaluate K, from eqn. 16; (iii) Evaluate Voi = (KJe,? + @,;)/2 (iv) Calculate = K,E,;/~; (v) Goto (i). This process can be shown to converge. But in favour of brevity the proof is omitted here. 312 4 Example simulations Concept verification is shown by a simulation presented in this Section. A two-link manipulator (Fig. 7) is taken as the dynamic system to be controlled. The perturbations considered consist of unknown viscous and coulomb type joint frictions, uncertainties in the inertia matrix elements (. lo%), and unknown Coriolis and centrifugal terms. The manipulator model utilised in the simulations is given in [7]. where 8 = [e, is. the vector of the joint. angular positions; b,B =,[b,, e,, bv2 @,IT and p, sgn(8) = [pC1 sgn(O,), p,, sgn(8,)lT are the viscous and coulomb friction force vectors, respectively, and M is the inertia matrix",
+            " 50 to provide an upper bound for A*. To obtain this, it is also necessary to find an upper bound for B [u(t) - u(t - $1. The term in the brackets is the vector containing the control increments from one time step to another. If this quantity is bounded by some known A U , ~ . ~ = [Au!~,,Au~,,,]~, for instance, due to actuator's practical limitations, it is possible to find an upper bound for A*. For simplicity Aulmax = A u ~ ~ ~ ~ = Au,,, (actuator's saturation value) is assumed. Thus, for the two link manipulator (Fig. 7): where a' = II2Jlm - I'3m2 cos2(&) 2 I'2mI'l,m - Therefore the upper bounds of the elements in eqn. 52 are given by: where z is the sampling time. Eqn. 48 is directly used in the estimated perturbation vector which is: *est @calcula ted - f - Bu(t - T ) (49) Eqn. 48 is a good acceleration estimation for fast sampling rates and low velocity measurement noise. On the other hand, in the presence of discontinuous perturbations such as coulomb friction, it may provide a very poor estimation. Comparing eqns"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0003804_iros.2001.973345-Figure1-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0003804_iros.2001.973345-Figure1-1.png",
+        "caption": "Figure 1 : One of our \u201cPalmBots.\u201d",
+        "texts": [
+            " The main advantages are that it requires only a simple manipulator and it permits manipulation of a wide variety of objects. (A gripper limits the size and shape of objects that can be manipulated.) On the other hand, the robot must have some understanding of the task mechanics, and we may need multiple nonprehensile manipulators to accomplish a task. The challenge in nonprehensile manipulation is to formulate robust algorithms for manipulation planning and execution; the focus of our research is to develop such algorithms for nonprehensile mobile manipulators. Figure 1 shows a picture of one of ourrobots. It is a simple differential drive robot equipped with a 2-DOF \u201cpalm ma- nipulator.\u201d This manipulator consists of parallel arms that can be raised or lowered and a plate at the end (the \u201cpalm\u201d) which can rotated. Our eventual goal is to create a multiple-robot system that can pick up an object on the floor, carry it to a goal location, and set it down in a specified orientation. To complete this task, the robots may need to pick up the object, reorient it, and transfer it to other robots"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0000942_j.renene.2016.11.054-Figure16-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0000942_j.renene.2016.11.054-Figure16-1.png",
+        "caption": "Fig. 16. The photograph of bearing damage (a) inner-race damage and (b) outer-race damage.",
+        "texts": [
+            " The rolling bearing running data were acquired from a faulty generator of Spectra Quest Inc.\u2019s mechanical structure damage simulator shown in Fig. 14(a). And Fig. 14(b) shows the vibration test site of the faulty generator and the corresponding complete bearing damage simulation system is shown in Fig. 15. CoCo 80 data acquisition system is selected to collect the running condition vibration data of experimental generator bearing. The weak pitting damages are generated by pneumatic grinding pen on bearing inner-race and outer-race respectively, shown in Fig. 16(a) and Fig. 16(b). The test bearing supports the motor shaft. Condition vibration data was collected using accelerometers, which weremounted on to the housing, with the sampling frequency of 25600 Hz for output end bearing experiments. The experimental bearings with weak damages used in this test are the deep groove ball bearings with the type of NSK 6203 and the parameters of this type of bearing are displayed in Table 2. In this validation experiment, themotor rotating speedwas 2343 r/min (39.05 Hz, 0.0261s) and the length of data analyzed was 2560 points"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0003547_0954406001523074-Figure3-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0003547_0954406001523074-Figure3-1.png",
+        "caption": "Fig. 3 Position and direction of the conjugate lines generated in Fig. 2, shown as the pitch outlines rotate through a full cycle: (a) and (b) representation of the positions and orientations taken by the conjugate line using the PAT mode of locating that line, where (a) is for one-half of the rotation of a \u00aerst-order ellipse and (b) is for the second half; (c) positions taken by the conjugate line for the PAC mode of de\u00aening the angle of that line; (d) two involute teeth in contact are shown with the conjugate line aligned with the centre-line, for which condition the teeth cannot transmit any torque (they would come out of mesh, that is, slide past each other)",
+        "texts": [
+            " The angle of the conjugate line, CA, measured from the LC is given by CA TA\u00ff PAT The pressure angle PAT and the tangent angle TA are indicated in Fig. 2b. Since these are non-circular pitch outlines, the common tangent at the pitch point will oscillate about the vertical as the pitch point moves to and fro between the gear centres. As a consequence of this, a line that makes a \u00aexed angle to this tangent will also oscillate as it moves with the pitch point. For the single-order ellipses in Fig. 2, the largest di erence in the direction of the conjugate line will be at the mid-point of the line of centres. In the half-cycle shown in Fig. 3a, at the mid-point the conjugate line will be at the maximum angle: CAmax TAmax \u00ff PAT In the next half of the cycle (Fig. 3b), again at the midpoint, the conjugate line angle CA will be minimum: CAmin TAmin \u00ff PAT C02197 \u00df IMechE 2000 Proc Instn Mech Engrs Vol 214 Part C at OhioLink on November 7, 2014pic.sagepub.comDownloaded from If CAmin reaches 0 , which is depicted for the two teeth shown in Fig. 3d, the gear teeth will come out of mesh. In practice, to avoid this problem the variation in gear ratio is kept to less than about 3:1 and the pressure angle PAT to about 15 or less. These limitations are indicative of the di culties encountered in the generation and use of non-circular gears. In Figs 2a and c the direction of the line of action is \u00aexed with respect to the line of centres (LC). The conjugate line then maintains the same orientation in a \u00aexed frame as it moves to and fro between the two gear centres, following the pitch point as the gears rotate. This con\u00aeguration will be designated as the PAC. Schematically, the range of locations of the line of action is shown in Fig. 3c. As the gears rotate through a complete cycle, the conjugate line will move twice between its extreme locations on the line of centres. For this de\u00aenition of the pressure angle: CA PAC In summary, if the pressure angle from the tangent is de\u00aened, conjugate lines oscillating about the vertical are obtained. If the pressure angle with respect to the line of centres is de\u00aened, the conjugate lines will remain parallel to each other as the gears rotate. These will be referred to here as rotating and parallel conjugate lines"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0002074_1.1359772-Figure4-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0002074_1.1359772-Figure4-1.png",
+        "caption": "Fig. 4 See-through labyrinth with honeycomb-stator",
+        "texts": [
+            " @4# indicate that a 16-cavity TOR seal does not improve the stability significantly. Unfortunately the number of cavities, the honeycomb cell width and the pressure levels differ and therefore a direct comparison with the actual results is not possible. When the reference seal ~TOR29! is used with a honeycomb-stator, the leakage rate is reduced to 60 percent of the initial value. As to isolate the effects, the seal is first used without swirl brakes. The main dimensions of the honeycombs are mentioned in Fig. 4. As mentioned before, the seal geometry and the guide vane ring can easily be removed and changed. The preswirl, which is of outstanding importance for the generation of the exciting forces, is varied by using guide vanes with different nozzle angles. The magnitude of the entry swirl is estimated out of the metered massflow through the seals and the nozzle area and angle. This is maybe only an approximation of the real entry swirl, but the inflow situation in the real axial turbomachinery is inhomogeneous, too, and therefore probably very similar"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0000963_detc2016-59644-Figure6-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0000963_detc2016-59644-Figure6-1.png",
+        "caption": "FIGURE 6: Surface temperature contours overlaid on mass-modified domain for three separate scan cases, \u03b4 = r0, \u03b4 = 2r0 and \u03b4 = 4r0. Note the presence of significant thermal accumulation for the close-pitched cases, and the mass accumulation on the overlapping regions at the turn areas.",
+        "texts": [
+            " (7) These conditions are meant to reflect the actual process in that solid powder will only melt and be incorporated with a preexisting melt pool present in the powder substrate. It should be noted that this implies that the powder jet diameter is, in this case equal to the laser spot size, though this is not compulsory, and may be adjusted accordingly. Furthermore, the melting of incident solid powder is not taken into account when calculating the local melt pool temperatures, and a uniform mass flow rate over a circular cross-section jet is assumed. Three separate laser scan pitches were examined including \u03b4 = r0, \u03b4 = 2r0 and \u03b4 = 4r0. Fig. 6 shows surface temperature distributions for all three cases of \u03b4 for simulations conducted with local phase transformation being taken into account. Each surface plot shown was taken at its respective scan completion times which vary based upon the total scan length. The first noteworthy observation from this figure, is the fact that the simulation captures the naturally occurring mass accumulation at the turn points of each track since immediately after the right angle turn of the beam at the end if each track, there is a part of the powder injection cross-section that passes twice over the same area. This phenomenon is easily observed for the case of pitch= r0 where the overlap occurs along the entire length of the track. The second observation from Fig. 6 is the fact that as the scan pitch decreases, thermal accumulation becomes significant, and manifests as a persistent elevated temperature in regions adjacent to the active scan path. Figure 7 provides further insight into this scan pitch effect as it reveals the temperature evolution for all three scan pitches at four discrete points (Y1 \u2212Y4) located at the midpoint of each major scan (see Fig. 5). These plots reveal that in addition to persistent elevated temperatures after the laser has passed, there is a pre and post-scan effect whereby adjacent points that have yet to be scanned directly, or adjacent points that had just been scanned exhibit local temperature rises"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0002552_bf02947169-Figure1-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0002552_bf02947169-Figure1-1.png",
+        "caption": "Fig. 1",
+        "texts": [
+            " The purpose of the present work is to obtain exact solutions for out -of -p lane deflections, twist angles, angles of rotation, bending moments and twisting moments in a curved beam due to torque based on the classical and shear deformable beam theories, and to demonstrate the application of the differential quadrature method (D.Q.M.) to obtain accurate approximate solutions. The exact solutions are compared with those obtained by the D.Q.M. for the case of a circular arc of circular cross section with clamped and simply supported boundary conditions. 2. Theoret ical Consideration and Closed-Form Solut ions 2.1 Classical beam theory The curved shaft considered is shown in Fig. 1. The equilibrium equations for out -of -p lane bending and twisting of a thin circular arc can be expressed as follows (Volterra, 1952): M~'+ M~=O ; - M x + . M ~ = O (1) where Mx and Mz are the respective bending and The State Sensitivity Analysis of the Front Wheel Steering Vehicle 245 Geometry of curved beam. twisting moments at a given circumferential angular position r and a prime denotes differentiation with respect to r The constitutive equations for small deflections and rotations are EIx v\" where EI~ and G] are the respective flexural and torsional rigidities, R is the center-line radius of the member, v is the out-of-plane deflection, and ~ is the twist angle"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0003947_rtd2002-1642-Figure4-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0003947_rtd2002-1642-Figure4-1.png",
+        "caption": "Figure 4: Axle box detailed model.",
+        "texts": [],
+        "surrounding_texts": [
+            "The ADAMS detailed model of the axle-box is described in the following figure. The model is composed of 7 rigid bodies: \u2022 Bogieframe: the bogieframe is connected to ground through a traslational joint, which allow only the vertical motion (Z). \u2022 Lenoir-Link (2 bodies): The Lenoir-link has been built as two separate parts, one linked to the bogie and the other to the spring holder, both with two revolute joints. The two parts are then linked each other with a translational joint and a single force, which act as a unilateral bumpstop. When a force is applied to the spring holder by the spring the two parts of the link are moved away and the bumpstop operate such that the force is transferred to the bogie. \u2022 Spring holder: The Spring holder keeps the inner spring in the left side of the axle-box; it is connected to the LenoirLink as shown above. The Link inclination split the force supplied by the spring in two components in the X-Z plane. The spring-holder is connected to the left side of the pusher with a bumpstop so that the force given by the Lenoir-Link in the X direction is transferred to the pusher itself. \u2022 Pusher: the pusher is connected to the bogieframe with a traslational joint which allow only the relative motion in the X direction. The right side of the pusher is connected to the axle box with a force vector, this element model the first friction surface. In the X direction the force vector act as a bumpstop, the value of the X force is then used as the Normal force for the bi-dimensional friction force implemented in the Y and Z direction. \u2022 Axle-Box: The primary friction surface is located in the left side of the Axle-Box in the Y-Z plane. This surface has been modeled using four force vectors, one to each vertex of the surface. Each Force vector is modeled as the one on the pusher, so that the total friction force is distributed among the four forces depending on the normal force acting on each vertex. The vertical load is transferred from the Axle-Box to the bogie frame thought four springs, the outer springs have a gap which is closed only in the laden condition, so that in the tare load condition only the inner springs support the load. Moreover the right inner spring act thought the Link as described above. The axle box is connected to the axle with a revolute joint. \u2022 Axle: The axle in this model is connected to ground with a planar joint. All the bumpstop elements have been modeled with a stiffness of 1\u22c5108 N/m and the damping of 1000 N\u22c5s/m."
+        ]
+    },
+    {
+        "image_filename": "designv11_60_0003124_70.88068-Figure9-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0003124_70.88068-Figure9-1.png",
+        "caption": "Fig. 9. Special cases of line contact. (a) Special line contact between two bodies. (b) Wrench system representation of special line contact.",
+        "texts": [
+            " This can be achieved by slightly changing the direction of motion of the manipulated object, while maintaining the contact configuration. This gives us another point ( - b4/bzl 0, b6/b2) which will be the locus of pencil lines on the plane perpendicular to the distance vector. The location and orientation of the wrench axis can be uniquely determined by fitting a line through a series of these points [9]. In cases where the object is modeled as a quadratic surface or polyhedra with expected line contact, the wrench axis in the wrench coordinates can be written as (Fig. 9) (18) or the Plucker line coordinate of the wrench axis in the compliance frame can be written as w=(w1, w2, w3. 0, 0, 0) ul = wI cos 8- w2 sin 8 u2= w1 sin 8 + w2 cos 8 u3= w3 u4= w3d sin 8 US = w3d COS 8 (I6 = 0 (19) where for both cases of Fig. 9(a), d is a known value. Therefore, the exact location of the wrench axis can be obtained using either of the following equations: CONTACT LOCATION If the wrench axis intersects the object surface at more than one point, a criterion must be used to select the most probable solution. This criterion is based on the evaluation of the angle between the object surface normal and the obstacle (fixture) surface normal at the contact points. When two bodies are in contact, the normal unit vectors at the point of contact have to be in opposite directions [lo]"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0002942_0076-6879(87)41053-7-Figure4-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0002942_0076-6879(87)41053-7-Figure4-1.png",
+        "caption": "Figure 4 shows a device that meets the above requirements and has proven useful in a number of different applications (see below). The vessel made of transparent Plexiglas (polymethylmethacrylate) was designed for an optimal volume of 0.6-1.2 ml. An external water bath controls the temperature. Additions to the reaction mixture can be made with a microsyringe through the top or the side opening (Fig. 4B). The round shape of the reaction compartment makes the stirring by the round magnetic disk (diameter 10 mm; Radiometer, model 912-036) very efficient and the small knobs on top of the stirring disk produce the turbulence essential for fast mixing while the knobs on the bottom of disk have been removed. This results in a mixing time which is less than 0.1 sec while the noise level is less than 0.05 mV. The close fit of the disk in the vessel restricts lateral movements and thereby reduces electrical noise. The disk is driven by an adjustable magnetic stirrer (e.g., Cenco Instr., Breda, Netherlands;",
+        "texts": [
+            " 2a) which is then carefully fitted to the PVC tubing (Fig. 2b and c). Membrane and tubing are bonded by applying light suction to the tube. Light sucking and blowing into the tube is continued until the tetrahydrofuran has evaporated. Thereafter, any excess membrane material is removed with a razor blade (Fig. 2d). Now the electrode is filled with the degassed reference solution (10 mM CaC12) using a syringe. Finally, the minielectrode is completed by plugging the PVC tubing with the ionselective membrane to the electrode head (Fig. 4A) which contains the connector and the chlorinated Ag wire. The Ag/AgCI wire may be prepared from a 0.5-mm Ag wire as follows: First the wire is alternatively dipped into concentrated HNO3 and NH4OH. The clean wire is then connected via a 1.5-kll resistor to the anode of a 1.5-V cell while the cathode is connected to a Pt wire. Chlorination proceeds for 2 hr while both wires are immersed in 0. I M HC1. Reference Electrode. A suitable reference electrode may be made by simply inserting a diaphragm (porous ceramic plug) directly into a piece of PVC tubing and introducing it into the vessel in the same way as the Ca 2+- selective electrode. Alternatively, the reference half-cell of a combination pH electrode (Fig. 4B) introduced through the top opening may be used (e.g., Philips type CA 14/02). Care should be taken to ensure that the diaphragm stays clean and is not clogged by protein coating or AgS deposition. Electronics (Fig. 3). The electrode potential (Eca) may be amplified with any high impedance (-> 10 ~z 12) voltmeter. It may be convenient to use a standard pH meter (e.g., Philips PW 9409/09) provided that its response is fast enough. In this case the Ca2+-sensitive half-cell is connected to the glass electrode input with a shielded wire",
+            "ll Thermostatted Incubation Vessel 12 Optimal application of CaZ+-selective minielectrodes in biochemical experimentation with a need for high sensitivity requires a tailored vessel which should meet the following conditions: (1) small incubation volume, (2) short mixing time, (3) constant stirring without wobbling of the stirring bar, which would cause severe electric noise, (4) water jacket for temperature control, (5) optimal electric shielding, (6) possibility to make addi- n V. Madeira, Biochem. Biophys. Res. Commun. 64, 870 (1975). t2 H. Affolter and E. Sigel, Anal. Bioehem. 97, 315 (1979). tions to the medium during the experiment, (7) mechanical stability, and finally (8) such a vessel should permit the simultaneous measurement of other parameters such as oxygen, pH, or other ion activities. model 34534, max. 3000 rpm). The CaZ\u00f7-selective electrode is introduced into the side wall of the vessel through a hollow screw (Fig. 4B). When this part is screwed tight, an O ring holds the PVC tubing of the electrode in position. The electrode head is shielded and fixed in position by the assembly support which also holds the vessel. The vessel also enables measurements of multiple parameters ~2 and an example of such an application is shown in Fig. 4B. The pH of the solution may be monitored simply by connecting the glass electrode output to a second pH meter. The reference output of the combination glass electrode may in this case be used for both the Ca 2+ and the pH measurement. A K\u00f7-selective electrode 12,13 may be prepared in an analogous way to the Ca 2+ electrode and introduced into the vessel by the second side port. Likewise, a tetraphenylphosphonium-sensitive electrode 14 can be made and used to measure membrane potentials and charge translocations"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0001235_b978-0-08-099425-3.00005-4-Figure5.73-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0001235_b978-0-08-099425-3.00005-4-Figure5.73-1.png",
+        "caption": "FIGURE 5.73",
+        "texts": [
+            " The more recent embodiments of the Magic Formula, as well as switching to the normalised formulation shown in Table 5.5, have added simple scaling factors such as lKya (referred to as LKY inside the property files) to allow just such manipulation. All fitting of the model to measured data is done with these scaling factors set at unity. This allows the intriguing possibility of taking high quality flat-track data, with elaborate fits for all the comprehensive slip character of the tyre and so on, and scaling it according to measured data on a vehicle or tyre test trailer (See Figure 5.73) using high quality load wheel instrumentation to reflect the same tyre on a variety of different surfaces. While still in its infancy, this process shows good promise even with the very noisy data sets that might be acquired on a rolling vehicle, as can be seen in Figure 5.73. A fundamental limitation of most point followers including the Magic Formula concept is that they are limited to steady state calculations of the tyre forces. Additional layers of calculation can be added to the scheme to permit the representation of transient delays in force build-up. The most commonly described method is the so-called \u2018relaxation length\u2019. When considering a tyre to which a step change in, say, slip angle is applied, a finite length must be travelled over the ground before the tyre carcass assumes its steady state shape"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0002669_jsco.1998.0200-Figure1-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0002669_jsco.1998.0200-Figure1-1.png",
+        "caption": "Figure 1.",
+        "texts": [
+            " Polynomial systems, where the variables are interpreted as trigonometric functions of unknown angles, are quite ubiquitous, arising, for instance, in electrical networking and in molecular kinematics. Here, our applications will be taken from the field of robot kinematics. Besides referring to the many situations described in Kova\u0301cs (1993), for the sake of being self-contained, we will outline a few examples of the role of sine\u2013cosine systems in robotics. Example 1.1. Given a robot arm with six revolute joints, i.e. a 6R robot (see Figure 1), a typical problem is finding the values of the different joint angles (with respect to some 0747\u20137171/98/070031 + 40 $30.00/0 c\u00a9 1998 Academic Press standard way of measuring them) that place the tip (or hand) of the robot at some desired position and orientation. This issue, known as the inverse kinematics problem, amounts to solving a polynomial system where the unknowns are the sines and cosines {si = sin(\u03b8i), ci = cos(\u03b8i), i = 1, . . . , 6} of the six joint angles {\u03b8i, i = 1, . . . , 6}. For general robots the solution of such systems is quite involved, as noted in the next example"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0003120_156855387x00057-Figure9-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0003120_156855387x00057-Figure9-1.png",
+        "caption": "Figure 9. Verifying grasping and measuring the end of the rope.",
+        "texts": [
+            " The left half image corresponds to the image viewed from the left eye and the right half to that from the right eye. The end points of the rope and the central point of the ring are detected in both images, and the three-dimensional position in the base coordinates of the camera is calculated. To test whether the points searched independently in both images correspond, vision checks whether the points are on the epipolar line and whether the three-dimensional distance between the object and the eye is valid. 4.5. Verifying the grasp operation and planning the trajectory Figure 9 shows the operation to verify grasping by vision. To determine whether or not the rope is grasped, the robot tries to move the hand (SWEEP). The vision system finds the rope and checks whether the orientation of the rope is the same as one of the fingers. Then the robot rotates the hand in order to aid measurement of the rope length D ow nl oa de d by [ H er io t- W at t U ni ve rs ity ] at 0 3: 23 0 7 Ja nu ar y 20 15 51 (ROLL). Vision measures the vector Vin C, which is from the grasp point to the end point of the rope"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0003887_978-3-642-73890-6-Figure2\u00b73-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0003887_978-3-642-73890-6-Figure2\u00b73-1.png",
+        "caption": "Figure 2\u00b73: Physical model of sbcc:l metal drilling",
+        "texts": [],
+        "surrounding_texts": [
+            "752\nThe objective of this paper is to present an automated manufactwing system for the drilling of sheet metal parts. All stand-alone systems such as the robot, a set of reconfigurable flXtores. and the CAD/CAM wOItstation have been integrated into a Flexible Manufacturing System. This system analyzes and evaluates a given fixturing layout from a stress-slrain point of view and sets the reconfigurable flXtores automatically using a robot manipulator. The developed software consists of a database system and four major modules: a workpart geometry display. a routine for flXturing analysis. a fIXture sequential control software. and a program to configure the fixtores automatically.\n2. Recontigurable Fixtures and The Drilling of Sheet Metal Parts\nThe hardware for tile proposed reconfigurable fixture was designed and manufactured at M.I.T [1.2.18.19.6]. The design takes into accOlDlt tile automated assembly of the fixtores by a robot manipulator. The reconfigurable fixtwing system consists of a fixture bed (baseplate). vertical supports (fIXture elements) and two locating pins. The prototype of the fixturing system with a typical workpart is shown in Figure 2-1. The main idea of the robotic reconfigurable fixtwing concept is to support the sheet metal from beneath with vertical support flXturing elements on a universal flX:ture base. The base consists of a flat plate with evenly spaced machined T -slots. as shown in Figure 2-1. Vertical supports are stored initially in a fIXture magazine and then the robot grasps elements from the magazine and positions them onto the baseplate. Each slot of the baseplate has a tapered entrance to allow easy insertion of the fixturing elements. A typical element shown in Figure 2-2. is adjustable in x. y and z directions.\nA typical sheet metal part, shown in Figure 2-3. has two tabs as a pair of reference points to locate the part which is a common practice in fuselage manufacturing. The sheet metal is placed and located accurately relative to the base by two pins. which constrain the workpart in x-y plane. Only one of the pins is fIXed. and therefore its position is known. The second pin originally is located at its \"home\" position in the sliding track. It prevents the workpart from rotating about the first pin and is adjusted to the fmal location by the robot. During the assembly of the fIXture. the vertical supports are inserted in the appropriate slot, slid along the desired position and oriented appropriatJy. Therefore. the automatic assembly of these fIXtures require the following parameters: slot number. sliding distance. orientation and the supporting shaft height. The assembly of the fIXturing elements must take into account the collision and interference problem since the elements can share the same or adjoining slot. However. in dealing with flexible parts. simply constraining the six degrees of freedom is not sufficient to prevent undesirable deformations while loads are applied. Redundant supports are required so that the part will not distort from its original shape. even wben no machine loads are present. This is especially important in a manufacturing environment requiring high accuracy. The design and assembly of the reconfigurable fixtures must address this inherent characteristic of sheet metals: the fixture must define the workpart shape completely. The possibility for the part to yield or buckle under the drilling loads must be checked.\nThe drilling process can be summarized as two distinct actions: one at the chisel and the other at the lips of the drill bit. The entire process can be modelled as a net thrust T and a torque M by assuming symmetry about the drill point. Several ways were developed to evaluate the thrust and the torque. The equations to determine the drilling thrust and torque along with the parameter definitions and values can be found in reference [141.\nT = 2 Kd Ff F, B w + Kd d2 J W M = Kd Ff F,. A W\n3. Fixturing Analysis\n3.1. Finite Element Procedure Failure analysis constitutes a major element in the fixturing analysis. A major type of failure is the distortion of the part from its true shape after machining due to the inherent low stiffness of the sheet metal. This failure situation can be avoided by performing stress analyses of the system. Finite element analysis is used to check whether the workpart will sustain permanent deformation or buckle under the drilling loads. and to evaluate the fixturing layout considered.\nIn the model of Figure 2-3 the vertical supports are assumed to have point contact with the workparL They only constrain the part in the Z direction. Two locating pins are considered as reaction forces to restrain the part in the X -Y plane. An eight-node double curved shell elements (S8R) \u2022 with four integration points. was used in modelling the sheet metal parts. The machining fOICeS are modelled as forces and moments applied at the corresponding nodes. The vertical supports and the",
+            "753",
+            "754\nlocating pins are introduced as part of the boundary conditions. The degree of freedom in the z-direction is constrained at the nodes where the fIXtures are located.\n3.2. Fixturing Layout Evaluation Excessive stress level induced at any location within the sheet metal part will cause the part to defonn plastically from its original shape. Three and four fIXturing elements at different mdii from the drilling point were considered. The stress distribution within the workpart for different layout configumtions , were investigated. In addition, several plate curvatures were examined. Specifically, nine different mdii of curvatures are investigated: \u00b125 in,\u00b150 in,\u00b175 in,\u00b1 100 in, and infmity. The fixture elements are located along the diagonal axes as shown in Figure 3\u00b71.\nThe nonnal stresSes in 11, 22 directions and shear stress as a function of fixture support locatinn and the applied forces were analyzed. 1be simulations on yielding were performed for the following cases: flat plates. shells with positive mdius of curvature (convex shells) and shells with negative radius of curvature (concave shells). (19).\nFigures 3\u00b72, 3\u00b73111d 3-4 show the normal and shear stresses for the flat plate. The normal stresses in II and 22 direction at the loading point iDcIQIe willi the increasing radia1 distance of fixture elements from the center of the plate. These stresses are compressive when die fixtures are near the center of the plate because the applied forces are directly transferred to the fIXtures resulting in small deflections. In this case the workpart fibers are compressed. However. the stresses are tensile as the fIXture elements move away from the center of the plate. Due to the locating pins constraint, the plate stretches resulting in large deflections. Both normal stresses 11 and 22 at the fixtures increase slightly when fixtures are near the center of the plate. 'but stresses increase drastically as fiXtures near the edges of the plate. The stresses are always tensile because material is pulled toward the loading poinL The normal stresses at the pins are always compressive. and they decrease with the increasing mdius of fixtures from die center of the plate. The shear stresses distribution at the point of pins and applied for are relatively constant and low. but the shear stresses at the fixtures increase as the radius of fixture from the center increase. After reviewing all the plots. the best fixturing layout is at a radiusllength ratio of 0.55. Length and radius were defined previously in Figure 3-1. Several simulations have been conducted for convex and concave shells in reference [19). The best fixturing locations for all positive-curvature shells are given by a radiUS/iength ratio ranging from 0.45 to 0.65. As for the concave shells. any fIXturing layout sbouId be feasible as long as the fixture elements are within a 0.65 radiUS/iength ratio. As for the three support fIXturing system. the fixtures are placed 120 degrees apart and six fixturing layouts for each radius of curvature have been studied. All the collected data show that the stress profiles for all curvatures have more or less the same pattern as the four-point fixturing. However. the three-point support case bas higher stress than the four-support fixturing as expected. The stresses at different locations. which are loading point, locating pins. and fixtures for different curves are given in reference [19).\nThe yielding analysis cannot ensure that the workpart will not defonn from its true shape. Permanent defonnation. which is still possible. can be caused by buckling where the workpart can suddenly snap dIrougb. The buckling phenomenon will happen when forces are applied to the convex faces. In order to analyze the buclding of convex shells. the critical load has to be determined. This can be done by controlling the displacement of the node at the center of the shell to move down a certain"
+        ]
+    },
+    {
+        "image_filename": "designv11_60_0000878_978-3-319-47614-8-Figure4.2-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0000878_978-3-319-47614-8-Figure4.2-1.png",
+        "caption": "Fig. 4.2 Bars of variable cross-section",
+        "texts": [
+            " At the left support the strain is a maximum with a value Pl EA which decreases in a linear manner to zero at the free end. The stress likewise is also a maximum at the left end Pl A and reaches zero at the free right end. 4.1 Simply Supported Bar 105 The structure in Fig. 4.1 is not only idealized as a one-dimensional structure in place of an actual structure like a pillar which is a three dimensional structure. Rarely such uniform structures occur in practice. A structure idealized by violating Poisson\u2019s law may be in practice a non-uniform structure and is considered in the class of one-dimensional structures. Figure 4.2a shows a tapered bar and its area of cross-section can be expressed as a function of x. But in Fig. 4.2b we have a complex structure whose cross-sectional area is not amenable for expressing as a function of x. There can be stepped bars with area of cross-section changing at specified positions along the x axis. In the absence of present day computational facilities, these structures were dealt by equilibrium methods to determine average stresses at any cross-section by balancing the internal force given by average stress multiplied by area A with the external load. Today we adopt a process called Finite Element Method, which is not entirely new but not implemented until recently with the availability of high performance computing"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0003547_0954406001523074-Figure5-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0003547_0954406001523074-Figure5-1.png",
+        "caption": "Fig. 5 Involutes arrived at by tracing a point on a cord as it is",
+        "texts": [
+            " The selection was made on the basis that this outline, while demonstrating signi\u00aecant di culties, might provide manageable challenges. A broader examination did not appear necessary at this stage. The \u00aerst concept applied was that of carrying out the numerical equivalent of tracing points on tangent cords as those cords are unrolled from the base lines. The results of the PAC-E outline (Fig. 4a) are not shown because its discontinuities gave rise to similar outcomes as the PAT-E outline (Fig. 4b). Figure 5 shows the results of tracing points on inextensible tangential cords, as they are unrolled from each of the remaining base outlines. The cords were necessarily on the convex side of the outline, at the tangent point. No reasonable results seemed to be produced by relaxing or reinterpreting the requirement that the cord be tangent at the point of contact. The involutes in Fig. 5a have discontinuities where the tangential cords generating them encounter discontinuities in the base outline. With the type of construction used for Fig. 5 it is not possible to demonstrate directly whether the matching gear faces will mesh successfully. For the PAT-E base line (Fig. 5a) it is clear that problems will arise at least for those portions of the base outlines that protrude into the opposite pitch outlines. The PAC-N base lines (Figs 4c and 5b) do not generate discontinuities and only give rise to concavities if the pitch outline itself is in part concave. For the pressure angle PAC of 60 used here (when averaged over a cycle, this is comparable with a normal pressure angle PA of 30 ), these base outlines extend outside the pitch outlines. With an external base line internal teeth may be expected, but the algorithm used for this \u00aegure does not reproduce this situation",
+            " 4a) is not included because it would show similar features to PAT-E: (a) PAT-E, (b) PAC-N and (c) PAT-N Proc Instn Mech Engrs Vol 214 Part C C02197 \u00df IMechE 2000 at OhioLink on November 7, 2014pic.sagepub.comDownloaded from such pressure angles would create relatively large normal forces at their point of contact and large bearing forces, with associated large frictional losses. There is also the possibility that the teeth will ultimately not slide (at locations away from the pitch points) but wedge and lock the gears together. Figure 5c shows unrolled involutes for the case PATN. An angle PAT or simply PA of 30 is used here, and it can be seen that in this case there is the smallest amount of excursion by the base outline, and interesting involute pro\u00aeles are generated. In all the \u00aegures shown here, the same number of points were calculated for each involute. Where the base outline is closest to the gear centre, the outline becomes concave and involutes are produced on the inside of the base line. Some of the teeth pro\u00aeles in this region appear to be short, but this is because, part way up the involute, they reverse the direction in which they have grown. Each internal tooth will eventually cease to extend inwards; it will then begin to progress outwards. Rotating the outlines in Fig. 5c generates the least problems thus far, but they nevertheless test the imagination if complete revolutions are required from gears. An exercise that can be done for any one of the pitch outlines (with bases and involutes) shown here is to make two identical photocopies of the \u00aegures on to transparent \u00aelm. Flip one over to give the mirror image about the X axis and superimpose the transparencies so that the centres and axes of the two \u00aegures coexist. Photocopy these two transparencies on to another single transparency",
+            " (b) Enlargement of the region about the pitch point, showing that normal conjugate action is taking place between the opposite-handed involutes of action and base outlines of the same second-order ellipse shown above is used. (b) Enlargement of the region about the pitch point, showing that normal conjugate action is taking place between the oppositehanded involutes Proc Instn Mech Engrs Vol 214 Part C C02197 \u00df IMechE 2000 at OhioLink on November 7, 2014pic.sagepub.comDownloaded from contact point. This angle, larger than that used for Fig. 5c, shows that the base outline can be contained within the pitch outline if a su ciently large pressure angle is used. In a coordinate system in which the pitch point is stationary, the instantaneous line of action would rotate about a mean direction, which here would be at 45 to either the X or Y axis. Figure 8, like Fig. 7, shows involutes that meet the requirements of Buckingham's basic law for teeth pro\u00aeles (Sections 3.1 and 3.2 above). It can be seen that, in the region enlarged in Fig. 8b, it would be easy to make excessively long teeth so that they would clash with their opposite numbers"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0001777_951293-Figure8-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0001777_951293-Figure8-1.png",
+        "caption": "Figure 8. Dynamic Hysteretic Process",
+        "texts": [
+            " The result of a full static cycle consisting of loading and unloading phases is the loop OQ,Q,O,O which represents an idealization of the loop OQO in Fig. 6. When immediate changes of the load are replaced with fast but finite ones, the shape and size of this loop depend on the manner of loading conditions especially on their speed and timing. Fig. 6 can be an example of the actual static hysteresis loop presenting actual, static friction processes. b) Dvnamic Hvsteresis Loop The dynamic elastic hysteresis can be understood as an elastic hysteresis of a mechanical system when the system is not able to keep pace with load changes. Fig. 8 shows the situation when loading (OQ) and unloading (QO,) processes are rapid and the load is removed immediately after the loading phase is completed without any time for relaxation. In the triangle OQO,, there is no elastic delay but the curve 0,O represents the reverse elastic delay which is still present in the system. If the load is applied in the opposite direction immediately after the state 0, is reached, the process of deformation and relative motion are running along the path O,Y,T as shown in Fig. 9. Fig. 9 consists of a continuation of the process shown in Fig. 8. It can be noticed that to remove the displacement 002 in this fast (dynamic) process, a negative load OY, is necessary. Let us assume that, when at state T, the next phase of motion starts without any delay and that both unloading and loading processes are similar to those between Q and T. The representation of this phase is a curve TO,Y,Q. The closed line QO,Y,TO,Y,Q is called a dynamic hysteresis loop. ) Mechanical Pro~erties of Elastic Hvsteresis In the case of dynamic hysteresis such as in Fig"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0001627_8.480-Figure3-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0001627_8.480-Figure3-1.png",
+        "caption": "FIG. 3.",
+        "texts": [
+            "50 which, by the figures, gives the criticals 2290 and 872 and it is seen that the shaft critical is raised by the gear effect. An inspection of the figures shows under which circumstances the gear effect is particularly large. In the first place, the shaft and gear critical speeds should be close to each other, so that one moves on the extreme curves of the family. In the second place, the coupling should be large, which means not only a large gear ratio, but more important, a small frame or cylinder inertia. Fig. 3 shows a schematic view of the engine fitted with a spur-gear reduction. The problem is determined by the following six equations: D ow nl oa de d by U N IV E R SI T Y O F M IC H IG A N o n Fe br ua ry 2 1, 2 01 5 | h ttp :// ar c. ai aa .o rg | D O I: 1 0. 25 14 /8 .4 80 Ilipl + kifoi \u2014 (pi) = 0 I2(p2 + k2((p2 \u2014 (fb) \u2014 Ircc.<PZ ~ 0 h<pz + km + T(Ti + f2) \u2014 Irec.'<P2 = 0 Tr\\ = ki(<pi \u2014 (pi) Tr2 = k2(<fb \u2014 <P*) ((Pi \u2014 <pa)ri = ((ps \u2014 (ph)r2 ( i ) (2) (3) (4) (5) (6) In all of these the angles <p are counted positive if clockwise seen from the right of Fig. 3. The last terms of the second and third equations express the \"inertia coupling\" between pistons and cylinders; the Irec% being the moment of inertia of half the re ciprocating masses concentrated at the crank radius r: Ei (7) The term Irec,(pz expresses the clockwise torque exerted on the crank, if the crank is held still and the cylinders be given a clockwise acceleration <ps This causes the pistons to accelerate in and along the cylinders, giving forces in the connecting rods, which add up to form the torque indicated"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0000077_icoin.2016.7427086-Figure2-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0000077_icoin.2016.7427086-Figure2-1.png",
+        "caption": "Fig. 2. The Velocity Obstacle VOA|B",
+        "texts": [
+            " The obstacle also has fixed radius rB, current position pb, and constant velocity vB which are unknown to the UAV due to the limitation of its sensors. III. VELOCITY OBSTACLE We use the method of using velocity obstacle [16] in choosing the velocity that the UAV would take. We will briefly discuss the concept of velocity obstacle and how it is applied to this paper. The velocity obstacle VOA|B is the set of all velocities of A that will result in a collision to B at some future time assuming that B maintains a constant velocity vB. A geometric interpretation of VOA|B is shown in Fig. 2. Let A B be the Minkowski sum of A and B, let \u2013A denote A reflected in its p ,v) be the ray starting at position p with direction v: A B = {a + b | a A, b B} (1) \u2013 A = {\u2013 a | a A} (2) p, v) = {p + tv | t We now define velocity obstacle as: VOA|B = {vA pA, vA \u2013 vB) B \u2013 } (4) which means that choosing vA inside VOA|B will result to collision between A and B at some future time. While selecting vA outside VOA|B would avoid collision with B. IV. COLLISION AVOIDANCE WITH LIMITED OBSTACLE INFORMATION As discussed in the previous section, defining the velocity obstacle requires the obstacle\u2019s radius, position and velocity"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0003787_iros.1996.570634-Figure4-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0003787_iros.1996.570634-Figure4-1.png",
+        "caption": "Figure 4: Positioning of Link (I) Figure 5: Positioning of Link (11)",
+        "texts": [
+            " Any scalar function that satisfies the boundary conditions and is twice differentiable with respect to time t can be s(t) in (i) or @(t) in (ii). In the trajectory of (i), the angle of the link Q remains constant. In (ii), B completely follows @(t) , and the rotation of the link stops when $(t) = 0. 3.2 Composite Trajectory The trajectories which start from an arbitrary initial configuration and reach an arbitrary desired configuration can be composed of the trajectory segments of Section 3.1. Positioning between the initial configuration A:[zo, yo, do] and the desired configuration B:[xd, yd, dd] is considered (Fig.4). The manipulator starts at t = 0 from rest and stops at t = T. The positions of C.I. at the initial and desired configurations are denoted by PO and P d . RA and Rg are circles with radius A, centered at PO and Pd . PO and Pd are connected by a straight line. C and D are the intersection points between the circles and the straight line. (C and D are on the same side in relation to PO and Pd.) C.I. remains at rest as far as the joint is accelerated or decelerated along these circles. Therefore, the joint can move from A to C without shifting C"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0000860_978-3-319-44735-3_1-Figure1.1-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0000860_978-3-319-44735-3_1-Figure1.1-1.png",
+        "caption": "Fig. 1.1 The plant to be controlled: (i) X axis angle is the Roll rotation, (ii) the Y axis is the Pitch rotation, and (iii) the Z axis is the Yaw rotation. The arrow direction means positive",
+        "texts": [
+            " The reminder of this chapter is organized as follows: Sect. 1.2 provides an overview of the control problem; Sect. 1.3 presents the proposed stabilization and movement multi-layer fuzzy controller; Sect. 1.4 provides details on each fuzzy controller that comprises the proposed controller; Sect. 1.5 discusses the conducted experiments and the obtained results; finally, Sect. 1.6 draws some conclusions and presents future work directions. This section describes briefly the system under control, i.e. the hexacopter, as a control plant. Figure1.1 shows the hexacopter, which is composed by six rotors organized as \u201cHexa +\u201d topology.1 By activating the rotors accordingly, it is possible to control the hexacoptermaneuvering through the X, Y and Z axes. In general, each plant must be analyzed to discover the interaction of each force. In this case, the thrust force produced by speeding up or slowing down some rotors leads the hexacopter toward the desired direction (on each axis), i.e. the vectors of the forces acting on the plant. To understand the movements performed by the hexacopter it is worth to take a look at the forces acting in frame"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0000617_s1064230716030126-Figure8-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0000617_s1064230716030126-Figure8-1.png",
+        "caption": "Fig. 8. Recommended kinematics diagram of the manipulator mechanism.",
+        "texts": [
+            " In the course of the experiment it was also revealed that the manipulator mechanism reproduces the position/orientation of the space manipulator model with some errors, which increase during the transition from a quasi-stationary to a more dynamic trajectory defined by the human operator for the real robot using the arm. It was established that it occurs because of the insufficient stiffness of the structure of the manipulator\u2019s mechanism. Apparently, in order to increase the stiffness of the mechanism\u2019s structure, it is advisable to replace the anthropomorphic kinematic diagram that was used in the experiment with the one that provides greater stiffness. A good option is the kinematic diagram shown in Fig. 8. Its first three degrees of freedom provide the movement along three mutually perpendicular axes, with the first two forming a f lat plate, and the third being perpendicular to the plane. The other three movements are rotational. They are provided by the Euclidean hinge with three mutually perpendicular axes. In this paper, methods for remotely controlling a space robot using interface simulators in order to train human operators to interact with the space robot and remotely control the robot in the master-slave mode with reflected force were discussed and theoretically and experimentally proved"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0000914_ijcnn.2016.7727893-Figure3-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0000914_ijcnn.2016.7727893-Figure3-1.png",
+        "caption": "Fig. 3 Results from a synthetic experiment. The purple points show data collected at a given pose. The link on the left hand side corresponds to the end effector. The base joint is at the right hand side. The blue, green and red points show home cloud points labelled according to the learned labels (1, 2 and 3) and transformed according to the corresponding learned transformations. The two figures (a) and (b) show results from two different poses at the corners of the joint space. Since the collected and transformed points essentially overlap, the color of the rigid parts looks blended.",
+        "texts": [
+            " The three joints of the articulated structure have parallel revolution axes such that end effector moves only in a 2D plane. Joint poses are uniformly sampled within a predefined 3D joint space. This data was constructed by creating a home cloud containing three segments, and transforming them according to a known kinematic model. During learning, the point labels and the transformation weights were initialized randomly. After the learning process converges, the segmentation of the home cloud as well as examples of the point cloud alignment obtained with the learned transformation are shown in Fig. 3. The transformed point cloud aligned well with the target point cloud. To quantitatively measure the performance of the algorithm during training, we computed the averaged testing error for the synthetic experiment is defined as: , 1 1( ) ( )i T P l T q i q i q iT e t p F t h C \u2208 = = \u2212 \u22c5 Q (16) 5252 2016 International Joint Conference on Neural Networks (IJCNN) where TQ is a set of testing configurations, P is the number of points in the home cloud, TC is the total number of points tested (P times the number of testing positions, and ,q ip is the point in pose q corresponding to point i in the home pose, which is known since the data set was created by transforming home cloud points",
+            " Since this data is synthetic, if the transformations were learned perfectly, the error would be identically zero. The testing error trend is shown in Fig. 4. The final averaged testing error of 42.62 10\u2212\u00d7 is much smaller than the distance between adjacent 3D points in the home cloud ( 21.00 10\u2212\u00d7 ). To infer the kinematic structure of this synthetic articulated body, we calculate ,k js according to equation (13) - (15). The values are listed in Table 1. Following Algorithm 1, we find [1, 2,3]J = and [1,3,2]S = . This is consistent with the coloring shown in Fig. 3, indicating that the algorithm has correctly inferred the structural relationships between the links and joints. For the experiment with real data, we use a robot built from SCHUNK PowerCube modules as shown in Fig. 1(a). In the experiment, the robot controls three of its joints (3DoF) so that the arm moves in front of a fixed depth sensor. The background is static. This dataset is significantly more challenging than the synthetic data set, since points collected at pose q do not necessarily correspond to points from the home pose"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0000790_cca.2016.7587858-Figure1-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0000790_cca.2016.7587858-Figure1-1.png",
+        "caption": "Fig. 1. a) The center of gravity (CG), the pivot point (P) and the hand position h. b) Relative velocities in the NED frame.",
+        "texts": [
+            " The model (26) is nonlinear and it is not possible to directly apply the distributed controller derived in Section V to a group of vehicles described by (26). Therefore, we will now apply the results of [20], which achieves a LTI external dynamics and an internal dynamics which is asymptotically stable for bounded states of the external dynamics. In particular, in [20] the concept of hand position hi = [xhi , yhi ]T = [xi+di cos(\u03c8i), yi+di sin(\u03c8i)] T , where di > 0 is constant, [xi, yi] T is the pivot point and \u03c8i the yaw angle (cf. Figure 1), has been applied to under-actuated marine vehicles. Using hi as output, it has been shown that the model is output-feedback linearizable. Then applying the change of coordinates T (x) =  \u03c8i=\u03c8hi ri=rhi xhi =xi+d cos(\u03c8i) yhi =yi+d sin(\u03c8i) x\u0307hi =uri cos(\u03c8i)\u2212vri sin(\u03c8i)\u2212rid sin(\u03c8i) y\u0307hi =uri sin(\u03c8i)+vri cos(\u03c8i)+rid cos(\u03c8i)  =  \u03d51i \u03d52i \u03d53i \u03d54i \u03d55i \u03d56i  (27) and the controller[ \u03c4ui \u03c4ri ] = [ cos(\u03c8i) \u2212d sin(\u03c8i) sin(\u03c8i) d cos(\u03c8i) ]\u22121 [ \u2212F\u03d55i (\u03d51i ,\u03d55i ,\u03d56i )+\u00b5\u03d55i \u2212F\u03d56i (\u03d51i ,\u03d55i ,\u03d56i )+\u00b5\u03d56i ] (28) for each vehicle we obtain \u03d5\u03071i =\u03d52i (29a) \u03d5\u03072i =f(\u03d51i , \u03d52i , \u03d55i , \u03d56i , \u00b51i , \u00b52i) (29b) \u03d5\u03073i =\u03d55i + Vxi (29c) \u03d5\u03074i =\u03d56i + Vyi (29d) \u03d5\u03075i =\u00b5\u03d55i (29e) \u03d5\u03076i =\u00b5\u03d56i (29f) where f(\u00b7) is a nonlinear function given in Appendix I"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0001197_978-3-658-12701-5-Figure3.1-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0001197_978-3-658-12701-5-Figure3.1-1.png",
+        "caption": "Figure 3.1: Subsystem motor, gear, arm",
+        "texts": [
+            "3) referred to as kinematic chain in [2], each subsystem\u2019s auxiliary velocities \u0307yu\ufffd can be alternatively expressed using the generalized velocities s\u0307 and the predecessor\u2019s auxiliary velocities \u0307yu\ufffd\u22121, \u239b\u239c\u239c\u239c\u239c\u239c\u239c\u239c\u239c\u239c \u239d u\ufffdv0,u\ufffd u\ufffd\ud835\udf140,u\ufffd u\ufffdr\u0307u\ufffd u\ufffd\ud835\udf14u\ufffd \u239e\u239f\u239f\u239f\u239f\u239f\u239f\u239f\u239f\u239f \u23a0 = \u239b\u239c \u239d ( Au\ufffd,u\ufffd\u22121 O O Au\ufffd,u\ufffd\u22121 ) \u22c5 ( I u\ufffd\u22121 \u0303r\u22ba u\ufffd\u22121 I u\ufffd\u22121r\u0303 \u22ba u\ufffd\u22121 O I O I ) O O O O O O O O \u239e\u239f \u23a0\u23df\u23df\u23df\u23df\u23df\u23df\u23df\u23df\u23df\u23df\u23df\u23df\u23df\u23df\u23df\u23df\u23df\u23df\u23df Tu\ufffd,u\ufffd\u22121 \u239b\u239c\u239c\u239c\u239c\u239c\u239c\u239c\u239c\u239c \u239d u\ufffd\u22121v0,u\ufffd\u22121 u\ufffd\u22121\ud835\udf140,u\ufffd\u22121 u\ufffd\u22121r\u0307u\ufffd\u22121 u\ufffd\u22121\ud835\udf14u\ufffd\u22121 \u239e\u239f\u239f\u239f\u239f\u239f\u239f\u239f\u239f\u239f \u23a0\u23df\u23df\u23df\u23df\u23df\u23df\u23df \u0307yu\ufffd\u22121 +Fu\ufffd \u0307su\ufffd. In order to clarify the method described above, the synthetic modeling approach will be utilized in the following example. The subsystem consisting of motor, gear and arm as depicted in Figure 3.1 is considered mounted on a predecessor body or subsystem at point 0. A frame of reference (subscript \ud835\udc5b) is defined to describe the subsystem\u2019s motion. This frame of reference moves with the translational velocity v0 and with the rotational velocity \ud835\udf140 with respect to the inertial origin. For the subsystem the auxiliary velocities 3.2 Modeling example: subsystem motor, gear, arm 19 are defined as \u0307yu\ufffd = \u239b\u239c\u239c\u239c\u239c\u239c \u239d v0 \ud835\udf140 \ud835\udf14a \u239e\u239f\u239f\u239f\u239f\u239f \u23a0 . The relative translational velocity r\u0307c and the first two components of the relative rotational velocity \ud835\udf14r are zero and therefore omitted for the sake of compactness",
+            "2) on page 17 and the derivations above, the synthesis of the equations of motion can be completed using the full functional matrices F1 = (\u2202 \u0307y1 \u2202s\u0307 ) \u22ba = \u239b\u239c\u239c\u239c\u239c\u239c \u239d 0 0 0 0 0 0 1 0 0 0 0 0 0 0 0 0 0 0 0 0 0 \u239e\u239f\u239f\u239f\u239f\u239f \u23a0 \u22ba F2 = (\u2202 \u0307y2 \u2202s\u0307 ) \u22ba = \u239b\u239c\u239c\u239c\u239c\u239c \u239d \ud835\udc591 sin (\ud835\udc5e2) \ud835\udc591 cos (\ud835\udc5e2) 0 0 0 1 0 0 0 0 0 0 0 1 0 0 0 0 0 0 0 \u239e\u239f\u239f\u239f\u239f\u239f \u23a0 \u22ba F3 = (\u2202 \u0307y3 \u2202s\u0307 ) \u22ba = \u239b\u239c\u239c\u239c\u239c\u239c \u239d \ud835\udc591 sin (\ud835\udc5e2 + \ud835\udc5e3) + \ud835\udc592 sin (\ud835\udc5e3) \ud835\udc591 cos (\ud835\udc5e2 + \ud835\udc5e3) + \ud835\udc592 cos (\ud835\udc5e3) 0 0 0 1 0 \ud835\udc592 sin (\ud835\udc5e3) \ud835\udc592 cos (\ud835\udc5e3) 0 0 0 1 0 0 0 0 0 0 0 1 \u239e\u239f\u239f\u239f\u239f\u239f \u23a0 \u22ba . Analysis of the manipulator\u2019s characteristics shows that it consists of a series of three subsystems with identical structure, i.e. a link that is connected to a motor via gears, see Figure 3.1 on page 18. It is practical to choose a modular approach by creating a dynamic model of the subsystem first and then synthesizing the entire model. In Chapter 3, the Projection Equation is derived in a formulation suitable for the present subsystem class. Synthesis of subsystems As the kinematics have been covered in the previous section and the subsystem matrices for the dynamic system can be obtained from Chapter 3, only the respective predecessor velocities and the relevant model parameters need to be substituted, \ud835\udf140,1 = \u239b\u239c\u239c\u239c\u239c\u239c \u239d 0 0 \u0307\ud835\udc5e1 \u239e\u239f\u239f\u239f\u239f\u239f \u23a0 , \ud835\udf14arm = \u0307\ud835\udc5e1, \ud835\udc5aarm = \ud835\udc5aarm1 \ud835\udc36c,arm = \ud835\udc36c,arm1 \ud835\udc60arm = \ud835\udc60arm1 \ud835\udc5amot = \ud835\udc5amot1 \ud835\udc36c,mot = \ud835\udc36c,mot1 \ud835\udc56G = \ud835\udc56G,1 \ud835\udc40mot = \ud835\udc40mot,1 \ud835\udc51 = \ud835\udc511 \ud835\udf140,2 = \u239b\u239c\u239c\u239c\u239c\u239c \u239d 0 0 \u0307\ud835\udc5e1 + \u0307\ud835\udc5e2 \u239e\u239f\u239f\u239f\u239f\u239f \u23a0 , \ud835\udf14arm = \u0307\ud835\udc5e2, \ud835\udc5aarm = \ud835\udc5aarm2 \ud835\udc36c,arm = \ud835\udc36c,arm2 \ud835\udc60arm = \ud835\udc60arm2 \ud835\udc5amot = \ud835\udc5amot2 \ud835\udc36c,mot = \ud835\udc36c,mot2 \ud835\udc56G = \ud835\udc56G,2 \ud835\udc40mot = \ud835\udc40mot,2 \ud835\udc51 = \ud835\udc512 \ud835\udf140,3 = \u239b\u239c\u239c\u239c\u239c\u239c \u239d 0 0 \u0307\ud835\udc5e1 + \u0307\ud835\udc5e2 + \u0307\ud835\udc5e3 \u239e\u239f\u239f\u239f\u239f\u239f \u23a0 , \ud835\udf14arm = \u0307\ud835\udc5e3, \ud835\udc5aarm = \ud835\udc5aarm3 \ud835\udc36c,arm = \ud835\udc36c,arm3 \ud835\udc60arm = \ud835\udc60arm3 \ud835\udc5amot = \ud835\udc5amot3 \ud835\udc36c,mot = \ud835\udc36c,mot3 \ud835\udc56G = \ud835\udc56G,3 \ud835\udc40mot = \ud835\udc40mot,3 \ud835\udc51 = \ud835\udc513"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0002030_20.560098-Figure10-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0002030_20.560098-Figure10-1.png",
+        "caption": "Fig. 10. (a) Input nonsinusoidal voltage. (b) Flux. (c) Estimated distorted hysteresis loop. (d) Sinusoidal hysteresis loop.",
+        "texts": [
+            " The flux density during time interval t l - t 2 represents the ascending path of the hysteresis loop, and the flux density during the interval t3-44 represents the descending path. 2) During the time intervals ( t 2 - t 3 ) and (t4-t5), the flux density B has its maximum value which causes the core to be saturated. The field intensity H changes, depending on the current, during these intervals. The result of applying a nonsinusoidal excitation waveform can be explained as follows. Consider a saturated flux-density waveform as shown in Fig. 10 whose peak values are sufficient to saturate the core. The change in B as a function of H during one cycle produces the hysteresis shown in Fig. 10(b). The underlying process can be explained as follows. 690 IEEE TRANSACTIONS ON MAGNETICS, VOL. 33, NO. 1 , JANUARY 1997 @) 0 (a) where N is the effective number of turns in the excitation coil. Substituting for A 4 from (20) into (19), the flux density can be calculated for different time intervals. The value of H corresponding to the flux density B is determined using the simple hysteresis loop of Fig. 10(a) for the time intervals ( t l - t2) , (t6-t7), and (t7-43). For the time intervals (t3-t4) and (t5-t6) where the minor loops occur, the value of H cannot simply be detennined from the hysteresis loop of Fig. lO(a). Consequently, the need for a satisfactory method to estimate the value of N during these intervals arises. Assuming that the curvature between points 1 and 2 of the minor loop is the same as the curvature between points a and b on the main loop, i.e., during time interval (t4t 5 ) , the curve between points 1 and 2 is similar to that occurring during the interval (t 8-t9) between points a and b"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0000571_j.proeng.2016.06.743-Figure1-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0000571_j.proeng.2016.06.743-Figure1-1.png",
+        "caption": "Fig. 1 \u2013 Physical picture of the process of loading the material with a gathering-arm",
+        "texts": [
+            " But the consideration of the form of the working element movement trajectory in the pile is difficult in the connection with the absence of the convenient index describing this trajectory and suitable for using at in formulas. For evaluation of the trajectory quality and working out the corresponding criterion it is necessary, basing on the investigations conducted [12-14] on determination of the physical picture of the loading process, to analyze the behavior of the loaded material pile at its interaction with the loading organ. In such interaction the pile volume deformation takes place which can be characterizes by the form and size of the volumes displaced. At the movement of the working element in the pile (fig. 1) it influences some volume of the material forming the prism of displacement. The working element influencing the material lumps before itself consolidate the material which causes protruding some lumps of the material upward from the apron and on the apron along the working element [13]. The part of the material rising along the lines of sliding runs over the working element. The material grasped by the working element is given to the conveyer, the volume of the pile active zone becoming smaller and the material descending and \u201cdeviating\u201d from the loading organ",
+            " The working element trajectory should be as close to ideal one [19]: in the phase of intruding the working element moves in parallel to the receiving conveyer axe entering the pile into the given depth; in the phase of forming the grasping area the working element moves perpendicularly to the receiving conveyer axe in the direction to it, and the phase of scooping-pushing the working element displaces the material by the shortest way to the conveyer. Maximum productivity in the given conditions is reached if the material is displaced most of the part of its way to the receiving conveyer. It is expedient to choose the average vector of the displacing direction of the load grasped as the evaluation. But as it is seen in given fig. 1, at the influence the working element to the material its lumps get displacing in various directions because of the irregularity of distributing their forms and sizes. That is why, taking into consideration that the general direction of the grasped material displacement corresponds to the working element displacement, we can restrict our consideration by the direction of velocity vectors of the working element itself. The more precise the vectors of velocities of the working part of element are directed to the center of the receiving conveyer device, the higher will be the productivity at other equal conditions. So, two types of velocity vectors of each point are considered: the vector of \u201can ideal direction\u201d and the real vector of velocity which compose the angle between them, the counting of the angle being taken from the ideal vector. For example, for the point B (fig. 1) an ideal vector is B ideal, a real vector \u2013 B, the angle between them is B < 0. The condition B > 0 in fig. 1 is done for point F*. It is natural to suppose that individual parts of the working element make an unequal contribution to forming the grasping material volume. Weight coefficients are introduced for definition \u2013 normalized meanings of velocity vectors of the working element calculated points: max i iw (2) The average angle of the trajectory direction can be calculated as the average arithmetical of the weighed-average sums of the angles for every calculated point of the working part of element at the whole period of scooping and pushing (for the mechanism in fig. 1 at [-2.18, 0.78]): m j n i ji n i jiji w w m 1 1 , 1 ,,1 , rad, (3) where n \u2013 the quantity of the calculated points on the working section of operating element; m \u2013 the quantity of the calculated working element positions at the period of the material scooping and pushing. The possible meanings of the quality criterion received: 1) K > 0 \u2013 the working element forms a small volume of grasp and moves it more to the left of the point G, in this case there is a danger of the material concentration on the driving disk and its returning back to the pile; 2) K < 0 \u2013 the working element forms a larger volume of grasp but moves it more to the right of the purposeful point G, in this case the part of the material is moved to the zone of work of the opposite working element, so some volume of the material is oscillate between the working elements; 3) K = 0 \u2013 the working element moves the grasped material to the receiving conveyer with minimum losses, it is the preferable variant"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0002486_s0005-1098(99)00044-8-Figure1-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0002486_s0005-1098(99)00044-8-Figure1-1.png",
+        "caption": "Fig. 1. An example of membership functions covering the input domain of the fuzzy controller with n\"2 and N 1 \"N 2 \"7.",
+        "texts": [
+            "9 j ]; one possible choice is fuzzy sets with the following triangular membership functions: k A1 j (e j )\"k (e j ;!R, a1 j , a2 j ), (6) k Ap j (e j )\"k (e j ; ap~1 j , ap j , ap`1 j ) (7) for p\"2, 3,2, N j !1, and k ANj j (e j )\"k (e j ; aNj~1 j , aNj j ,R), (8) where a.*/ j \"a1 j (a2 j (2(aNj~1 j (aNj j \"a.!9 j , and the triangular membership functions are de\"ned as k(x; a, b, c)\"G x~a b~a if x3[a, b), c~x c~b if x3[b, c], 0 if x3R![a, c] (9) and in the special cases a\"!R we de\"ne (x!a)/ (b!a)\"1 and c\"R we de\"ne (c!x)/(c!b)\"1. Fig. 1 shows an example with n\"2 and N 1 \"N 2 \"7. The fuzzy controller u is constructed from the following N 1 2N n rules: Rui12in : IF e 1 is Ai1 1 and2 and e n is Ain n , THEN u is Bi12in , (10) where i 1 \"1,2, N 1 ,2, i n \"1,2, N n , and Bi12in are fuzzy sets to be determined. (The design of fuzzy controller now boils down to the determination of Bi12in. ) Using product inference engine, singleton fuzzi\"er and center average defuzzi\"er (Wang, 1997), we obtain the fuzzy controller u\"u (e Dh) \" +N1 i1/1 2+Nn in/1 yN i12in(<n j/1 k Aij j (e j )) +N1 i1/1 2+Nn in/1 (<n j/1 k Aij j (e j )) \"hTm(e), (11) where yN i12in is the center of Bi12in, h is the collection of y6 i12in's for i 1 \"1,2, N 1 ,2, i n \"1,2, N n , and m(e) is the collection of the corresponding fuzzy basis functions",
+            " Step 2: Apply fuzzy controller (11) to system (1) and use the adaptation law (3) to adjust the parameters h with the initial parameters h(0) set to zero. Starting from some initial condition e(0) and run the closed-loop adaptive fuzzy control system until the error e(t) converges to around zero. Store the \"nal values of the parameters h. Step 3: Repeat Step 2 with a new initial condition e(0) which is chosen to be the center of the domain of rule Rui12in whose corresponding parameter yN i12in was not updated in the previous runs. Repeat the above runs from di!erent initial conditions until all the N 1 2N n para- meters yN i12in are updated. Fig. 1 illustrates an example where 12 runs are carried out to cover the domains of all the 49 rules. Step 4: The fuzzy sets Bi12in in Ru i12in of Eq. (10) are chosen through the following. First, \"nd the minimum and maximum values of the stored \"nal parameters h among all the runs (let them be u .*/ and u .!9 , respec- tively), and de\"ne N u fuzzy sets B1, B2,2, BNu over [u .*/ , u .!9 ]. For example, we may choose k B1(u)\"k (u ;!R, b1, b2), (12) k Bj(u)\"k (u; bj~1, bj, bj`1) (13) for j\"2,2, N u !1, and k BNu (u)\"k (u;bNu~1, bNu,R), (14) where b ",
+            "9 in the determination of Bi12in according to Eqs. (16) and (17). In this section, we use the method in Section 2 to design a fuzzy controller for the chaotic Du$ng system yK\"!0.1yR !y3#12 cos t#u (t), (18) which is a nonlinear time-varying system and shows chaotic behavior when u (t)\"0 (Wang, 1993). Our objective is to control the system output y (t) to follow a sequence of time-varying setpoints: y $ (t)\"G 2 if t3[0, 8], 0.5 if t3 (8, 16], 1.2 if t3[16, 24]. (19) We use seven fuzzy sets as shown in Fig. 1 to cover each e 1 \"y!y $ and e 2 \"yR !yR $ with a.*/ 1 \"a.*/ 2 \"!2, a.!9 1 \"a.!9 2 \"2, and the membership functions equally spaced. We choose c\"200 and Q, k j 's in such a way that p n \"p 2 \"[5, 5]T. Fig. 2 plots the y(t) versus y $ (t) for a single run of the adaptive fuzzy control system starting from initial condition e(0)\"[!2, 2]T, and Fig. 3 shows the trajectory of the same run in the e 1 }e 2 phase plane. We performed a total of 8 runs to cover the domain of interest and the trajectories of these runs are plotted in Fig"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0000630_acc.2016.7524998-Figure1-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0000630_acc.2016.7524998-Figure1-1.png",
+        "caption": "Fig. 1. a barbell with switching mass at the end",
+        "texts": [
+            " From Corollary 14 and letting Z\u03c3 = K\u03c3 (I +\u039bQ\u03c3 ), we have that \u03a6\u03c3 is stable and \u2016\u03a6\u03c3\u2016l\u221e-ind < \u03b3 if and only if there exist \u03b4 ,\u03b3i > 0, for i\u2208 {1,2,3,4,}, a positive integer M, Q\u03c3 \u2208S M O , and Z\u03c3 \u2208S M+1 O (or Z\u03c3 \u2208S M+1 IO ) such that \u03b31 +\u03b4\u03b32 < 1, (27) 1 \u03b4 \u03b32 + \u03b33 < \u03b3, and sup \u03c3\u2208\u039e \u2225\u2225A\u0302\u03c3 (I +\u039bQ\u03c3 )+ B\u0302u \u03c3 Z\u03c3 \u2212Q\u03c3 \u2225\u2225 l\u221e-ind < \u03b31, (28) sup \u03c3\u2208\u039e \u2225\u2225A\u0302\u03c3 (I +\u039bQ\u03c3 )\u039bB\u0302w \u03c3 + B\u0302u \u03c3 Z\u03c3 \u039bB\u0302w \u03c3 \u2212Q\u03c3 \u039bB\u0302w \u03c3 \u2225\u2225 l\u221e-ind < \u03b32, (29) sup \u03c3\u2208\u039e \u2225\u2225C\u0302z \u03c3 (I +\u039bQ\u03c3 )+ D\u0302zuZ\u03c3 \u2225\u2225 l\u221e-ind < \u03b33 (30) sup \u03c3\u2208\u039e \u2225\u2225C\u0302z \u03c3 (I +\u039bQ\u03c3 )\u039bB\u0302w \u03c3 + D\u0302zuZ\u03c3 \u039bB\u0302w \u03c3 + D\u0302zw \u03c3 \u2225\u2225 l\u221e-ind < \u03b34. (31) We emphasize again that the mappings in (28)-(31) are input-output LSS and their l\u221e induced norm can be computed via LP. Example 16: Consider the barbell of length l illustrated in Figure 1. There is mass of size m = 1kg that jumps from one end of the barbell to the other end. The actuator torque (control input) and the disturbance torque are labeled as u and \u03c4 , respectively. After letting l equal to the gravitational constant and discretizing the model at 2Hz we obtain xt+1 = A\u03c3t xt +B\u03c4 \u03c3t \u03c4t +Bu \u03c3t ut where A1 = [ 1.001 0.050 0.050 1.001 ] ,A2 = [ 0.999 0.050 \u22120.050 0.999 ] , B\u03c4 1 = B\u03c4 2 = Bu 1 = Bu 2 = [ 0.001 0.050 ] . We want to design a state-feedback controller, K : x\u2192 u, that stabilizes the closed-loop and study the l\u221e gain of the closed-loop from the disturbance torque \u03c4 to the states and control input, i"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0002092_0094-114x(96)84603-9-Figure8-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0002092_0094-114x(96)84603-9-Figure8-1.png",
+        "caption": "Fig. 8. Position of concentrated load on worm thread.",
+        "texts": [
+            " The investigations have shown that the fillet stresses are propor t ional to the distance o f the loaded point to the too th root. Accordingly, the influence factors o f the position o f the loaded point in the rad ia l d i rec t ion o f the w o r m th read /gea r t oo th can be expressed by the fol lowing l inear equa t ions flw)(hF) = 0.1128 + 0.8872hFr (11) f]g~(hE) = 0.0287 + 0.9713hEr (12) where her = o The curves in Fig. 10 represent the stress d i s t r ibu t ions a long the w o r m thread for four different load positions in the radial direction of the worm thread, which are defined in Fig. 8. These curves can be interpolated by the following function: (w) f2 (?D, hF) = a + bV7 + cV) '2 (13) where V 7 = 17D - YF[; the angles VD and 7F have to be substituted in degrees. For load positions with hvr = hv/hp >>-0.5, the values of coefficients a, b, and c are given in Table 1. For load positions with hvf < 0.5, the coefficients a, b, and c have to be calculated by the following expressions a = a, + b, hF, + c, h2v,"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0000477_j.ifacol.2015.09.726-Figure4-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0000477_j.ifacol.2015.09.726-Figure4-1.png",
+        "caption": "Fig. 4. Function H (Pk, Q\u03c9) construction",
+        "texts": [
+            "K and K = 49 and be the set of all values taken by Pk and based on experimental knowledge centered on the ideal value. Let H be a function that formalizes the noisy signal that has been generated by simulation. In order to get closer to reality, a white gaussian noise has been added to the signal so that measurement device is sightly noisy. Therefore, H (Pk, Q\u03c9) is the simulated flux probe signal that corresponds to the fault \u03c9 (\u03c9 = 1...W ) and the kth set Pk (k = 1...K) of parameters. The detailed process of this function is shown in Figure 4. Figure 5 summarizes the principle of the short-circuit detection. It consists in measuring by the flux probe into the airgap the voltage induced by the variation of the rotor magnetic flux due to the excitation current when the rotor is moving (though a prime mover in the case of alternator operating). The flux issue from each rotor pole can be identified through the voltage induced in the probe. In the healthy case, the signal is periodic according to the electrical period while this is no more the case if an inter turn short circuit occurs in a rotor pole",
+            "K and K = 49 and be the set of all values taken by Pk and based on experimental knowledge centered on the ideal value. Let H be a function that formalizes the noisy signal that has been generated by simulation. In order to get closer to reality, a white gaussian noise has been added to the signal so that measurement device is sightly noisy. Therefore, H (Pk, Q\u03c9) is the simulated flux probe signal that corresponds to the fault \u03c9 (\u03c9 = 1...W ) and the kth set Pk (k = 1...K) of parameters. The detailed process of this function is shown in Figure 4. Fig. 4. Function H (Pk, Q\u03c9) construction Figure 5 summarizes the principle of the short-circuit detection. It consists in measuring by the flux probe into the airgap the voltage induced by the variation of the rotor magnetic flux due to the excitation current when the rotor is moving (though a prime mover in the case of alternator operating). The flux issue from each rotor pole can be identified through the voltage induced in the probe. In the healthy case, the signal is periodic according to the electrical period while this is no more the case if an inter turn short circuit occurs in a rotor pole"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0002055_ip-epa:19960203-Figure1-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0002055_ip-epa:19960203-Figure1-1.png",
+        "caption": "Fig. 1 Currentfed coupled coils",
+        "texts": [
+            " It emerges that the essence of the vector control technique can be appreciated from the stationary coil analysis. This is a surprising but welcome finding. The coupled circuit results are then applied to the locked rotor model of the induction motor, and the mechanisms of transient torque control are explored. Finally, the results are extended to deal with torque control when the motor runs at constant speed. 2 Behaviour of coupled coils 2. I Outline of approach The rudimentary \u2018transformer\u2019 model Fig. 1 is often used to introduce the theory of the induction motor, but it is usually analysed with the primary supplied from a constant voltage and frequency source to reflect the traditional pattern of operation. However, vector 59 IEE Proc-Electr. Power Appl., Vol. 143, No. 1, January 1996 control schemes always use closed-loop stator current control, so it makes sense to treat the input current as the independent variable from the outset. An approach based on self and mutual inductances is followed in preference to the traditional approach via magnetising and leakage reactances. The last is well suited to the analysis of tightly coupled circuits fed from a voltage source because it avoids the requirement to specify the coupling coefficient to a high precision [8]. In the current-fed case, however, the coupling coefficient is not a critical parameter, as shown. e I r - I S - - - - - - + e 2.2 System equations In Fig. 1, suffices s (stator) and r (rotor) are chosen for the primary and secondary in anticipation of the later use of the model to represent the induction motor. The equations governing the coupled circuits are $, = L,i, + M i , (1) = Mi, + L,i, (2) = -Ri, * dt ( 3 ) With current is treated as the independent variable, the dependent variables in the s-domain are given by where 2.3 Step change of primary current An important assumption underpinning all of the discussion is that the primary current is can be stepped instantaneously"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0002537_a:1007917003845-Figure1-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0002537_a:1007917003845-Figure1-1.png",
+        "caption": "Figure 1. A multiple cooperating robot system grasping an object rigidly.",
+        "texts": [
+            " In Section 3, the Dual method of NLP with its implementation algorithm is explained and analyzed. Numerical examples and conclusion are shown and discussed in Sections 4 and 5, respectively. JINT1410.tex; 18/02/1998; 16:12; v.7; p.3 Various objective functions and constraints have been applied to the optimal force distribution problem of multiple cooperating robots by many researchers [1\u201313]. We formulate a general problem with a quadratic objective function and linear and quadratic constraints. Consider n cooperating robots grasping a single object rigidly, as shown in Figure 1. It is assumed that the robots do not pass through singular positions and that the manipulated object is rigid. Force and moment vectors of the ith (i = 1, . . . , n) cooperating robot exerted on the object are represented as if and in, respectively, as shown in Figure 1. The dynamic equation for the ith robot is i\u03c4 = iM(i\u03b8) i\u03b8\u0308 + ih(i\u03b8, i\u03b8\u0307) + ig(i\u03b8) + iJ(i\u03b8)T ix = iu(i\u03b8, i\u03b8\u0307, i\u03b8\u0308) + iJ(i\u03b8)T ix (1) where i\u03b8, i\u03b8\u0307, i\u03b8\u0308 joint displacement, velocity, acceleration vectors of robot i [m\u00d7 1] i\u03c4 actuator torque vector of robot i [m\u00d7 1] iM(i\u03b8) inertia matrix of robot i [m\u00d7m] ih(i\u03b8, i\u03b8\u0307) vector of Coriolis and centrifugal effects of robot i [m\u00d7 1] ig(i\u03b8) gravity vector of robot i [m\u00d7 1] iJ(i\u03b8) jacobian matrix of robot i [6\u00d7m] ix 4 = [ifT , inT ]T force/moment vector of robot i [6\u00d7 1] while m denotes the degrees of freedom of the robots which are assumed to be equal for simplicity"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0003013_a:1008115522778-Figure3-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0003013_a:1008115522778-Figure3-1.png",
+        "caption": "Figure 3. Misalignment between mating parts. (a) translation error; (b) orientation error.",
+        "texts": [
+            " The elastic deformation curve to be obtained from Equation (2) is called an elastica, its mathematical solution can be obtained by the trial and error method. However, its solving process and solution are very complex and time-consuming. Accordingly, it is not practical to apply its mathematical solution to actual assembly tasks. This paper deals with an assembly method that does not require its mathematical solution. The assembly errors, namely, misalignments between a hole and its mating part can be defined as shown in Figure 3. They are divided into translation error and orientation error. As shown in Figure 3, the em, \u03c6h, \u03c6v belong to the translation error, which is the error between the hole center ch and the center cp of the bottom of a part. And the \u03b8v, \u03b8h belong to the orientation error, which is the error between the center-line Lch of a hole and the tangent line Lcp to the center-line of a part. In other words, the translation error can be expressed by ch \u2212 cp. And the cosines of the orientation error are given by the direction cosines between Lch and Lcp . The orientation error \u03b8z about the z-axis does not appear in cylindrical parts assembly due to their symmetry with respect to the z-axis"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0002252_s0890-6955(98)00006-6-Figure1-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0002252_s0890-6955(98)00006-6-Figure1-1.png",
+        "caption": "Fig. 1. Model of the chamfered main cutting edge nose radius tool f . R,(R\u00de0).",
+        "texts": [
+            " According to Fuh and Chang [13,14], a basic model of three-dimensional cutting process, which can accurately predict the formation of shear planes for the case of chamfered main cutting edge, must have not only normal rake angle an, inclination angle i, side cutting edge angle Cs, and end cutting edge angle Ce, but also the cutting depth d, feedrate f, and cutting speed V. The cutting forces can be obtained from calculation, if the shear plane areas are determined accurately. However, experimental jobs as well as empirical equations [10,11] are necessary to calculate the shear plane areas, such as ae = sin21(sinaS2\u00b7cosab + sinhc\u00b7sinab) (1) where ae is the effective rake angle, ab is the parallel back rake angle (Fig. 1) and aS2 is the second positive normal side rake angle (Table 1). The angle hc is the chip flow angle, defined on the tool face as an angle between the normal to the main cutting edge and the direction of chip flow. All the angles are expressed in radian for digital computation. In the following, the process for deriving the shear plane areas is presented for two cases: (i) tool wear and (ii) no tool wear. 2.1. Shear areas in the cutting process with a chamfered main cutting edge nose radius (R) tool without tool wear The calculations of shear area A and projected area Q fall into one of the following categories depending on the relationship between nose radius, feedrate and the depth of cut: 1. the sharpness of the tool is such that its radius equals zero (R = 0), the calculations of shear area A and projected area Q, as shown in reference [13]. 2. nose radius of the tool (R) is smaller than the feedrate (f), R\u00de0 R , f, as shown in Fig. 1. The shear plane area A includes the area of (1) besides the cylindrical area of the tool nose [14,16]. 3. nose radius of the tool (R) is larger than the feedrate (f), R\u00de0 R . f, according to the depth of cutting, which can be subdivided into three parts: (a) d . R, (b) d = R and (c) d , R, as shown in reference [14]. In this section the case of small radius, that is case 2, R\u00de0 R , f, will be evaluated. The shear plane A and the projection area of the cutting cross-section Q, for the nose radius of the tool with a chamfered main cutting edge is shown in Fig"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0002075_bf02512517-Figure3-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0002075_bf02512517-Figure3-1.png",
+        "caption": "Fig. 3",
+        "texts": [],
+        "surrounding_texts": [
+            "The SOS/ISFET glucose sensor sensing region, used for glucose measurement, is shown in Fig. 4a. Both immobilised membranes were 150#m wide and 500 #m long. These membranes were placed in the ISFETs gate region. Fig. 4b is a cross-sectional view of the sensor sensing region. The sensor was composed of an enzume field-effect transistor (ENFET) and reference field-effect transistor (REFET). The Fig. 1 Systern block diagram SEF collecting cell: (a) photograph; (b) cross-sectional view diagram Medical & Biological Engineering & Computing May 1994 243 glucose concentration was measured by the ISFET glucose sensor, and results were recorded on the personal computer. After each measurement, the sensor was washed with buffer solution. It took about 3rain to take one sample measurement."
+        ]
+    },
+    {
+        "image_filename": "designv11_60_0003222_ac00284a081-Figure1-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0003222_ac00284a081-Figure1-1.png",
+        "caption": "Figure 1. Base line evaluation for the reverse scan in a cyclic voltammetrlc experiment.",
+        "texts": [
+            " I t is based on the same approach used also in ref 4 which takes advantage of the fact that the current exhibits typical diffusion controlled decay independent of potential, beginning from a definite potential sufficiently past the peak. This statement implies that when any further potential shift does not change the electrolysis conditions, the current-voltage curve obeys effectively the simple Cottrell equation (1) i ( t - t?'Iz = const where t ' is the hypothetical origin for the current-time curve (in the linear sweep experiment scale) which matches the diffusion part of the stationary electrode voltammogram. In particular, with reference to Figure 1, this equation will be obeyed by the current iA, recorded at the switching potential E,, as well as by the current i, relative to the potential E , chosen in such a way that (2) This last current represents the actual contribution of the forward process at E p b and hence it must be added to (ipb)' for a correct computation of the peak-current ratio i p b / i p f = [(ipb)' + ixl/ipf (3) Thus, we can write i x ( t x - t')l12 = i,(t, - tq1Iz (4) Since experimental curves are recorded in current-voltage coordinates, it is more convenient to replace times with PO- 0003-2700/85/0357-1503$01",
+            " 2 i12 - i2\u2019 12 The insertion of this equation, written by choosing E2 = EA, in relationship 8 allows one to obtain which represents an effective and rigorous tool for the calculation of the ratio ipb/ipf by quantities easily drawn from a single experimental voltammogram, provided that the potentiostatic condition (1) applies to the currents relative to the potential values El and EA. Therefore, it is of major importance to individuate where this condition is satisfied, that is, where the voltammogram is coincident with the chronoamperometric curve (see Figure 1). In order to locate the \u201cconvergence point\u201d, several voltammograms concerning both uncomplicated and complicated electrochemical reactions were considered. With this purpose, the voltammetric curves relative to different mechanisms already available in the literature (including EE, EC, ECE, proportionation, and disproportionation processes, all characterized by different reversibility degrees for the charge transfer steps) were simulated (10). The relevant programs, available on request, were written in FORTRAN IV and the calculations were carried out on a MINC-11/23 minicomputer (Digital Equipment Corp"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0000134_ecce.2014.6954068-Figure3-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0000134_ecce.2014.6954068-Figure3-1.png",
+        "caption": "Figure 3. Cross section of a hysteresis IPM motor",
+        "texts": [
+            " A hysteresis interior permanent magnet (IPM) motor is a self-starting solid rotor hybrid synchronous motor. Its rotor has a cylindrical ring made of composite material like 17%/36% cobalt steel alloy, special Al-Ni-Co, Vicalloy, etc. with high degree of hysteresis energy per unit volume [8-21]. The rare earth permanent magnets are buried inside the hysteresis ring and the ring is supported by a sleeve made of non-magnetic materials like Aluminum which forces the flux to travel circumferentially inside the rotor ring [8-21]. Fig. 3 presents the cross section of a hysteresis IPM motor which depicts the position and the orientation of permanent magnets. The inclusion of permanent magnets creates rotor saliency without changing the length of the physical airgap and acts an additional permanent source of excitation in the rotor. The induced magnetization of the hysteresis material inside the rotor ring always lags behind the time varying magnetic field. This time lag produces a torque, called hysteresis torque. Also, the induced eddy currents in the hysteresis ring generate some additional starting torque called the eddy current torque"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0002923_978-94-011-3514-6-Figure8.6-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0002923_978-94-011-3514-6-Figure8.6-1.png",
+        "caption": "Fig. 8.6 Relative values of the natural angular frequencies and of the angular frequencies at zero coupling",
+        "texts": [
+            "8) by replacing p by jw, then by dividing the first line by m, and the second by mz k, + kJ - wZ kJ ml ml 0 kJ kz + kJ wZ mz mz The first term is a function of wZ and one can write it, after expansion f(wZ) o One introduces the angular frequencies for zero coupling, defined as follows (figure 8.5), QZ , QZ Z angular frequency of m, when m2 is blocked angular frequency of mz when m, is blocked (2) m, is coupled to zero (8.24) - 147 - The term characterizing the elastic coupling is designated by Q12 so equation (8.24) becomes o The function f (w 2 ) is a parabola (figure 8.6) cutting the horizontal axis at the points w~ and w~ corresponding to the two natural angular frequencies of the system. The intersection of the parabola with the horizontal f{w 2 ) gives the two angular frequencies for zero coupling. These geometrical considerations lead to the following inequalities, which it is easy to establish algebraically Another geometrical representation can be proposed; it is inspired by Mohr's circles used for the study of the state of stress. In order to do that, let us calculate the roots of equation (8"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0001987_bf02347203-Figure2-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0001987_bf02347203-Figure2-1.png",
+        "caption": "Fig. 2. The fractions of the species containing H + (~i) present in the organic phase of the water - HNO 3 - Ca 2+ (microamounts) - 15C5 - nitrobenzene - H+B - system as a function of log c(L). c(HNO3) = 0.1 mol/l, Ca=0.01 mol/I. 1 5 (H+), 2 B (HL+), 3 ~ (HL2\u00a7 The curves were calculated for the log K o =-0.66, log Kex(HL +) = 3.61 and log Kex(HL2 +) = 5.00 constants",
+        "texts": [
+            " 2 The reliability interval of the constants are given - in agreement with Reference 15 - as 3s(K), where s(K) is the standard deviation of the constant K. These values are expressed in the logarithmic scale using the approximate relation log K _+{log [K+l.5s(K)]-log [K-1,5s(K)]}. For s(K)>0.2K, the previous relation is not valid and then only the upper limit is given in the parenthesis in the form log K (log [K+3s(K)]). ** The error-square sum U = Z(log Dcatc-log Dexp). 2 *** The extraction constants log Kex(HLorg +) and log Kex(HL2,org +) were obtained by minimization from the data in Table 1. Furthermore, Fig. 2 presents the part icipat ion o f the Horg +, HLorg + and HL2,org + particles on the total acidity o f the organic phase. F rom Fig. 2 it fo l lows that the complex HL2 + in the n i t robenzene phase is present in significant concentrat ions only at re lat ively high concentrat ions o f the 15C5 ligand in the system under study. On the other hand, ca lc ium is present in the ni t robenzene phase as CaL 2+ and CaL22+ complexes so the concentrat ion o f the cat ion Ca 2+ in this phase is negligible. The \"sandwich\" type CaL22+ complex prevails i f the condi t ion c(L)>c B is fulfil led. Our results are similar to that o f IWACHIDO et al"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0003765_tmag.2002.802692-Figure7-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0003765_tmag.2002.802692-Figure7-1.png",
+        "caption": "Fig. 7. Load-carrying capacity of the finite meniscus ring between parallel planes.",
+        "texts": [],
+        "surrounding_texts": [
+            "Dynamic characteristics of infinite-width meniscus and finite meniscus ring between parallel solid surfaces are analyzed by solving the time-dependent Reynolds equation considering the Laplace pressure of the meniscus. It is clarified by adopting the first order linearization that the dynamic pressure and force have three terms, i.e., the squeeze term generated by the viscosity of the liquid, the spring term by the dynamic Laplace pressure and the static Laplace pressure or the static meniscus force by the static Laplace pressure."
+        ]
+    },
+    {
+        "image_filename": "designv11_60_0003021_iemdc.2001.939271-Figure11-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0003021_iemdc.2001.939271-Figure11-1.png",
+        "caption": "Fig. 11. Current vector.",
+        "texts": [],
+        "surrounding_texts": [
+            "process, the voltage vector 2 is provided. If Id_t > Id--, the vector 2 is nearer to the N-pole than the vector 1. Therefore, 8, is 30 elec.deg., and Id_mor is Id_2. Next, the vectors 3 and 4 are provided, and 0, and Id_\" are similarly updated. However, when Id < Id-,,, the second process is completed because the previous vector is nearer to the Npole.\nOn the other hand, if Id-* < Id mllx, the vectors 12, 11, and 10 are provided, and 0, and Id-mor are similarly updated.\nWhen 8, is 180 elec.deg. in the first process, the vector 8 is provided and the similar way mentioned above is implemented. In this process, the estimation error must be smaller than *I5 elec.deg. The flow chart of the first and second processes is shown in Fig. 5.\nWhen the angle of the actual rotor position is about 90 or 270 elec.deg., the difference between Id-, and Id_, is quite small in the first process because the effect of the magnetic saturation is slight. If this difference is under the value determined in advance, the vectors 4 and 10 shown in Fig. 4 are provided instead of the second process, and Id_4 and Zd-,, are measured. In this study, this value is 0.03 A and it is determined by considering the accuracy of the current sensor. The estimated angle is obtained by similar way to the first process. Next, vectors 13 and 14 shown in Fig. 6 are provided after the vector 4 or 10 is provided once again to remove the effect of the residual magnetism [7]. This process is omitted in Fig. 5.\n3) Third process Three kinds of voltage vectors whose angle are 0, - 7.5, e,, 0, + 7.5 elec.deg. as shown in Fig. 7(a) are provided to the motor based on the result of the second process. The angle 8, in Fig. 7(a) represents the angle estimated by the second process. The estimated angle is updated by similar way to the second process. Thus, the estimation accuracy is *3.75 elec.deg. in this stage. Furthermore, the process mentioned above is repeated twice as shown in Fig. 7(b) and Fig. 7(c) to improve the estimation accuracy.\nSince the initial rotor position is estimated by using the current response for the voltage vector, this method is not affected by the set errors and the variation of the motor parameters. The theoretical maxi-\nmum accuracy of the rotor position estimated by this method is M.9375 elec.deg.\nIn. DECISION METHOD OF OPTIMAL VOLTAGE VECTOR\nThe amplitude and output time of the voltage vector are important because the estimation is implemented based on the variation of current response caused by the magnetic saturation for the voltage vector. The rotor of the motor may rotate during the estimation process if the large voltage vector is provided to the motor. Therefore, the optimal voltage vector, which can estimate the initial rotor position without the mechanical lock, must be determined before the estimation process.\nFig. 8 shows the experimental system to determine the optimal voltage vector. The current control is performed on the x-y frame shown in Fig. 1. In order to set the N-pole in the direction of the xaxis, e*, the following x-axis current reference i=*, and y-axis current reference i,* are applied to the current controller:\n[;:I = [ :] where the direct current I equals to the rated current or the current below that of the motor.\nTo decide the amplitude of the voltage vector, the modulation factor m is introduced and it is defined as follows:\nm = Vm/Ed (3) where V,,, is the amplitude of the voltage vector and Ed is the DC source voltage of the inverter. The output time Ton is initially set to 50 ps when the performance of the DSP controller is considered.\nIn this study, the voltage vector whose amplitude is V,,, is provided to the motor for Ton whenever all gate signals of the inverter are turned off during three times of Ton to force each phase current to zero.\n1) First step Fig. 9 shows the flow chart of the first step. The rotor position is set to 0 elec.deg. by giving 8* = 0 elec.deg. to the current controller. Next, the voltage vector 1 is provided to the motor for Ton to remove the residual magnetism [7], and then the d-axis current Id_/ is measured when the vector 1 is provided once again. Furthermore, the voltage vector 7 is applied and Id is measured similarly. The modulation factor m of the voltage vectors 1 and 7 are gradually increased from 0.1. If AId( = Id_/ - Id_, ) is smaller than the value determined in advance AI, m is increased. If m is larger than the maximum",
+            "modulation factor mmx which corresponds to the modulation factor for 11 0% of the rated voltage Vc of the motor, the output time is increased and the operation mentioned above is repeated as shown in\nstep. Fig. 10 shows the actual angle of the rotor and the measured currents ix and iy when the rotor position is changed from 0 elec.deg. to 60 elec.deg. by giving 8* = 60 elec.deg. to the current controller. Obviously from this figure, the currents iA and iy vary while the rotor moves. Therefore, the rotation of the rotor can be detected without a position sensor by observing ix and i,,. Fig. 1 1 shows the current vector which consists of ix and iy. The angle error A8 between the current vector and x-axis is obtained by the following equation:\nTherefore, the rotor position OMe is obtained as follows:\n(4)\n( 5 )\nwhere n is a sampling number. Fig. 12 shows the comparison of the estimated position OM, obtained by Eq. ( 5 ) and the actual position em by the position sensor when the rotor is set to -10 elec.deg. In this figure, the estimated and actual position show a good agreement.\nwhen the rotor is at\nstandstill. This result was obtained according to the following process. First, the rotor is set to 0 elec.deg., and next, the third step is carried out. Therefore, the rotor should be at standstill. From this figure, the accuracy of Eq. ( 5 ) is approximately +I .O elec.deg. Therefore, if <: 1.0 elec.deg., the voltage vector obtained by the first step is determined as the optimal voltage vector which should be used in the initial rotor position estimation. On the other hand, if 10MJ > 1.0 elec.deg., the modulation factor m is reduced according to the following equation:\nFig. 13 shows the estimated position",
+            "0.0 - -2.0\n'9 -4.0\n25 -6.0 3 - B Q -z .J:,\"\n-12.0\n2 0.20 >4 0.15 h\n-\n0.0 0.2 0.4 0.6 0.8 1.0 Time [SI\nA\n/\nThe experimental system for laboratory test is illustrated in Fig. 14. The test motor is a 400-W, I .9-A, 4-pole surface PMSM as shown in Table 1. High performance DSP controller is used to implement the proposed estimation algorithm. The stator currents of the motor are detected with LEM-modules (CT5-T), and are transformed into the digital signals through a 12 bit A/D converter. The actual rotor position is detected with a rotary encorder (R.E) mounted on its shaft for monitoring, and the resolution of the R.E is 4000 counts per revolution. DC source voltage of the inverter E, is 270 V.\nB. Decision of Optimal Voltage Vector\nWhen Ton < 200 ps, Lv, (= Id_, - Id_7 ) < AI (= 0.2 A), and the condition of the first step is not satisfied. Fig. 15 shows the relation between the modulation factor m and Lv, in the first step when the output time of the voltage vector Tor is 200 ps. As is obvious from this figure, the first step is completed because dl, > AI when the modulation factor m is 0.55.\nFig. 16 shows the estimated position OMe and actual position Om when the rotor is set to 0 elec.deg. after the second step is executed. In this figure, OMe is within + l . O elec.deg., and this means that the rotor is at standstill during the second step. Therefore, the amplitude of the optimal voltage vector is 148.5 V from Eq. (3), and the output\nE d\nPWM ~ n Drive Circuit Inverter . . .\nA/D\nFig. 14. Experimental system.\nTABLE I RATED VALUES OF TESTED MOTOR\n400W 152V 1.9A 5OHz 3000rpm 4poles\n0.25 - 1\n4\" 0.10 a\" v II. 0.05 l=yJzq 0.00\n0.10 0.15 0.20 0.25 0.30 0.35 0.40 0.45 0.50 0.55 Modulation factor m\nTime [SI Fig. 16. Comparison between actual rotor angle and estimated angle in the third step."
+        ]
+    },
+    {
+        "image_filename": "designv11_60_0001860_0094-114x(95)00022-q-Figure2-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0001860_0094-114x(95)00022-q-Figure2-1.png",
+        "caption": "Fig. 2. The Oldham Coupling: an indirect passive coupling with the same (planar) freedoms and constraints as the active coupling of an electric motor.",
+        "texts": [
+            " The modified coupling network must be over-constrained otherwise the couplings of the original machine would have been incapable of transmitting action. By confining attention here to planar coupling networks it is sufficient to replace active couplings by planar passive couplings to create over-constraint in the planar sense. The passive planar equivalent for an active coupling transmitting torque T, but not forces U, V, is the indirect passive coupling often referred to as the Oldham coupling shown in Fig. 2. The Oldham coupling is an indirect coupling comprising two prismatic couplings in series that permits translational velocities u, v but not rotation t. For an active coupling transmitting force, as in the punch/metal interaction, the replacement passive coupling is a binary link with two R couplings. Circuit actions attributable to active couplings 1005 Consider intuitively now the task of closing the circuits of the equivalent over-constrained network. The active coupling of the electric motor, represented by a dashed line in the coupling graph of the machine shown in Fig"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0000872_j.triboint.2016.11.008-Figure2-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0000872_j.triboint.2016.11.008-Figure2-1.png",
+        "caption": "Fig. 2. The schematic diagram of a ball in contact with the inner and outer raceways.",
+        "texts": [
+            " The contact angle, contact area, maximum contact pressure, numbers of balls, normal contact force (per ball), angular speeds \u03c9br (a ball\u2019s absolute revolutional speed with respect to the shaft axis), \u03c9i (rotational speed of the inner raceway or shaft), and \u03c9bs (ball spin angular velocity), and maximum sliding velocity of the ball with respect to the inner raceway (shown in Table 1) were obtained from the contact analyses for ball bearings with different groove factors. In the present study, the radii of curvature, r, for the inner and outer raceways were the same. As shown in Fig. 2, points o and i\u00b0 represent the centers of curvature of the outer and inner raceways before the application of the axial load, respectively. In this study, point o remained unchanged in terms of position after the application of a load, whereas point i\u00b0 moved to the new position i. Ri is the distance from point o to the z-axis (the axis of the rotating shaft), and Ro is the distance from point i to the z-axis. If we define D as the ball diameter, di as the diameter of the inner raceway, and d0 as the diameter of the outer raceway",
+            " The bearing diameteral clearance (Pd) of a ball from the inner and outer raceways, respectively, is written as: P d d D 2 = 2 \u2212 2 \u2212 .d o i (4) The initial contact angle \u03b10 before applying of a load is evaluated as: \u03b1 P A = cos (1 \u2212 2 )d o 0 \u22121 (5) Eq. (5) is valid for both the contacts in the inner and outer raceways provided that the rotational speed set forth in this study is used. The difference in the contact angle between these two raceways is caused mainly due to dependence on the centrifugal force. The geometrical analysis for A is written as: 2 0 0 2 (6) The contact angle \u03b1 (in Fig. 2) with respect to the position angle \u03c6 of the bearing system has a constant value around the raceway. \u03b4a in Eq. (6) can be written as: \u03b4 A \u03b1 A \u03b1= sin \u2212 sin .a o o (7) The axial force (Fa) applied to bearing specimens with different groove factors and numbers of balls was determined by the criterion of their having nearly the same maximum contact stress. \u03b1i and \u03b1o are the contact angles in the inner and outer raceways, respectively, obtained under a load of Fa and at a fixed rotational speed. The normal contact force arising in the contact area of the inner raceway and a ball can be expressed as [2]: Q K \u03b4=i i i 1",
+            " Therefore, the contact angle (\u03b1) after applying an axial load has almost the same value for both the inner and outer raceways, it is obtained from the use of Eq. (7). If the number of balls is N, the applied load (Fa) in the axial direction is calculated as: F NQ \u03b1= sin .a (18) The inner raceway acts as the driving surface with a sliding velocity, V \u2192 iB\u2032, with respect to the ball\u2019s rotational speed, \u03c9br; the balls act as the driven surfaces with a sliding velocity, V \u2192 bB\u2032, with respect to \u03c9br. As shown in Fig. 2, the distance vectors, r\u21c0B\u2032 and r\u21c0A\u2032, from the ball center to the contact points in the inner and outer raceways, respectively, are applied to determine the sliding velocity V \u2192 iB\u2032 as [23]:  \u239b \u239d\u239c \u239e \u23a0\u239f \u239b \u239d\u239c \u239e \u23a0\u239fV \u03c9 r d k D \u03c9 \u03c9 d D \u03b1 j \u2192 = \u2192 \u00d7 \u2192 + 2 = \u2212 1 2 ( \u2212 ) \u2212 cos .iB i br B m i br m i/\u2032 \u2032 (19) and the sliding velocity V \u2192 bB\u2032 as [23]: V \u03c9 r D\u03c9 \u03b1 \u03b2 \u03b1 \u03b2 j \u2192 = \u2192 \u00d7 \u2192 . =\u2212 1 2 (cos cos + sin sin )bB bs B bs i i\u2032 \u2032 (20) The slip ratio (SR) is defined as: V V V V S \u2261 2 \u2192 \u2212 \u2192 \u2192 + \u2192R iB bB iB bB \u2032 \u2032 \u2032 \u2032 (21) for the contact between the inner raceway and a ball only"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0000692_s11164-016-2654-0-Figure1-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0000692_s11164-016-2654-0-Figure1-1.png",
+        "caption": "Fig. 1 Schematic of a microbial fuel cell: a detailed structure, b assembled MFC",
+        "texts": [
+            " In this work, a MFC stack comprised of five two-chambered MFCs that were connected in series was developed and integrated into a U-shaped sink drain pipe for kitchen wastewater treatment and electricity production. The performance of the MFC stack and individual cells as well as the substrate degradation were evaluated. The open circuit voltage response to the flow rate and temperature of the substrate was studied. Finally, the stack performance was tested by using wastewater from a beverage as the anolyte. D. Ye et al. Figure 1 schematically depicts the two-chambered MFC used in the stack. The MFC was made of two polymethyl methacrylate (PMMA) tubes (wall thickness 5 mm). The inner one and the interspace were used as the anode and cathode chambers, respectively. A Nafion 117 proton exchange membrane (PEM) was used to separate the two chambers. Two PMMA end plates integrated with the anolyte inlet/ outlet were placed on the top and bottom of the tubes and were assembled together by four M5 screws. The catholyte inlet and outlet were also drilled on the end plates"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0000378_ever.2016.7476344-Figure1-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0000378_ever.2016.7476344-Figure1-1.png",
+        "caption": "Fig. 1: ljI-i curves/trajectory of a SRM",
+        "texts": [
+            " This paper focuses on ERs of three popular control strategies which are open loop, CCC and DITC which is investigated in reference [11-14]. The mathematical statement of SRD's ER will be introduced in section II, which is followed by simulations and experiments based on the aforementioned strategies in section III. Finally, conclusions will be presented at section IV. II. MATHEMATICAL STATEMENT ON ENERGY RATIO OF SRD In a rotor pitch, the energy conversion process in each phase winding of SRM can be illustrated by a series of 1fI i-e curves and trajectories of lfI-i. Fig.1 shows that lfI-i curves at two key points ell' ea and the trajectory of a SRM during a switching period of one phase winding. In the diagram, the motor is magnetized at the maximum reluctance position ell and demagnetized at the point b. According to the electromechanical energy conversion theory, the area oabco means the energy which is converted into mechanical energy during a conduction period of a phase winding; the area ocbdo means the energy which is still stored in phase winding at phase changing point b and will return to the supply after finishing phase-changing by neglecting the loss of converter; the area oabdo means the energy which is offered by the supply from ell to phase-change point b. As is mentioned in above, ER is a rate which can be expressed as following: total en ergy s upplied to machin e ER = __ - e_n _e _rg:::..: y=---r_et_u_ffi_e_d to_s _u-,,-,pP::....I..:..y __ total energy s upplied Wm We (1) Here, Wm represents the active energy, in other words, the mechanical energy outputted from the motor shaft; We represents the net electrical energy transmitted from the source. Considering the meaning of each area in Fig.1, the formula (1) therefore is also described as S oabeo S oabeo + S oebdo (2) Here, Wf represents the exciting energy what is stored by stator windings, and Wm = f\ufffdm Ijf(i, e) di = S oabeo (3) Wf = S;m i(ljf, e) d Ijf = S oebdo (4) It is obvious that Wm is not equal to Wf for the nonlinear magnetic path and the more Wf, the less ER. That is to say a large Wr will lead to a poor performance when SRM operates as a motor because a lower energy ratio will cause the reactive power increase. Ill. SIMULATIONS AND EXPERIMENTAL TESTS ON EFFECTS OF CONTROL STRATEGY ON ENERGY RATIO A"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0000938_iemcon.2016.7746308-Figure2-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0000938_iemcon.2016.7746308-Figure2-1.png",
+        "caption": "Fig. 2. Forces, moments and reference systems of a quadcopter.",
+        "texts": [
+            " So that, the quadcopter can maintain forward (backward) motion by increasing (decreasing) speed of front (rear) rotors speed while decreasing (increasing) rear (front) rotor speed simultaneously, which means changing the pitch angle. This process is required to compensate the action/reaction effect (Third Newton\u2019s Law). Propellers 1 and 3 have opposite pitch with respect to 2 and 4, so all thrusts have the same direction [7]. There are two reference systems that have to be defined as a reference which are Inertial reference system (Earth frame- 978-1-5090-0996-1/16/$31.00 \u00a92016 IEEE XE, YE, ZE) and quadrotor reference system (Body frameXB, YB, ZB). The reference system frames are shown in Figure 2. The dynamics of quadcopter can be describe in many different ways such as quaternion, Euler angle and direction matrix. However, in designing attitude stabilization control reference in axis angle is needed, so the designed controller can achieve a stable flight. In attitude stabilization control, all angle references in each axis must be approximately zero especially when take-off, landing or hover. It ensures that, the quadcopter body always is in horizontal state, when external forces are applied on it [8]"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0002305_1.2833509-Figure1-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0002305_1.2833509-Figure1-1.png",
+        "caption": "Fig. 1",
+        "texts": [],
+        "surrounding_texts": [
+            "Jim Bonvouloir Ferrofluidics Corp.,\n40 Simon St., Nashua, NH 03061\nExperimental Study of High Speed Sealing Capability of Single Stage Ferrofluidic Seals Several different configurations of single-stage ferrofluidic seals were tested on a spindle which was capable of operating at very high rotational speeds (up to 55 KRPM). A dimensionless number based on the ratio of magnetic force to centrifugal force was defined. It was discovered that this ratio is not a good predictor of high speed seal failure. Reynolds number was found to be a better predictor of seal failure; therefore an empirically derived model for predicting seal failure based on Reynolds number is proposed. The data herein may provide a basis for developing new theoreti cal models for ferrofluidic seal failures at high speed.\nIntroduction\nFerrofluidic seals are used in a wide variety of gas sealing applications. In recent years, an increasing number of potential applications have been encountered where a seal is required to operate at high rotational speeds. In order to design a reliable seal for a high speed application, it is important to understand how the various ferrofluid properties and design parameters will affect the high speed performance. To date, there has been very little empirical data available on ferrofluidic seals operating at high speeds, and the mechanism of high speed seal failure has not yet been identified. This test program was implemented in order to develop a good predictive model of high speed seal capabilities and also to gain some insight into the failure mecha nisms.\nApparatus\nFerrofluidic rotary seals have been previously described in the literature (Raj and Moskowitz, 1990). A schematic of the test apparatus is shown in Fig. I. The apparatus is basically a high speed spindle which drives a ferrofluidic seal, which in turn seals a leak-tight chamber. The shaft is supported by a pair of high-speed angular contact bearings and is driven by an integral DC motor at one end. On the chamber end of the shaft is a machine taper for attaching various shaft adapters. The spindle is capable of operating at speeds of up to 55,000 rpm and the maximum power output is 750 watts. The fixture was designed to accommodate three different shaft adapter sizes: 25.4 mm dia, 50.8 mm dia, and 101.6 mm dia. Three different housings were made to attach to the chamber mounting plate depending on which shaft adapter was installed. Each housing contained a single stage ferrofluidic seal.\nAltogether there were 6 different shaft/pole piece combina tions which were tested, and one configuration in which only the magnetic field strength in the gap was changed, for a total of 7 configurations. These are listed in Table 1. (The magnetic field was changed by varying the amount of permanent magnet material in the magnetic circuit.) There were two general classes of seal geometry which were tested, one with a tapered stage on the shaft (configurations A - F ) and one with a tapered stage on the pole (configuration G). These are shown in Fig. 2,\nContributed by the Tribology Division of THE AMERICAN SOCIETY OF MECHANI CAL ENGINEERS and presented at the ASME/STLE Joint Tribology Conference, San Francisco, Calif., October 13-17,1996. Manuscript received by the Tribology Division February 21, 1996: revised manuscript received June 28, 1996. Paper No. 96-Trib-31. Associate Technical Editor: R, F. Salant.\nThe spindle speed was monitored by a tachometer output signal on the motor controller (the speeds were independently verified using a high speed strobe light). The motor controller also provided a monitor for the motor armature current.\nFerrofluids Tested Six ferrofluids with different physical properties were used in the testing. They are referred to as Fluids 1 through 6 and are Usted in Table 2. The fill volumes which were used (shown in Table 1) are between 1.5 and 3 times the nominal gap volume. This is typical for ferrofluidic seals in most practical applica tions. The seals with a nominal 0.154 mm gap, configurations A, B, C, F, and G, were filled to 2 times the nominal gap volume. The fill volumes for configurations D and E, the large and small gap, were adjusted so that the effective wetted stage length would be approximately the same as for the other con figurations. See Appendix A for an explanation of the procedure for calculating effective stage length.\nProcedure\nThere are three parts to the test procedure. The majority of the testing, Part I, consisted of characterizing failure speed at low differential pressure. For this part of the testing all relevant parameters were varied: fluid viscosity, fluid magnetization, fluid density, shaft diameter, radial gap, concentricity of shaft, and magnetic flux density. For Parts II and III only one fluid was used and only configurations B and G were tested.\nI Failure Speed at Low Pressure. The seal was filled with ferrofluid and the static burst pressure was recorded. The seal was cleaned and refilled, and pressurized to 25 percent of its static burst pressure. The shaft speed was slowly increased until seal failure occurred. The failure speed and motor current values were recorded. This was repeated using a variety of fluids and configurations.\nIn this report, \"failure speed\" is defined as the shaft speed at which either (a) the seal splashed fluid outside the gap or (b) the chamber pressure began to decrease (indicating seal leakage). In Part I of the test, the shaft speed was increased slowly until a splash or leak was observed (the pressure was observed carefully so that even very small leaks were considered failures).\nII Failure Speed Versus Applied Pressure. The seal was filled with Fluid 3 and pressurized to some fraction of its static burst pressure. The shaft speed was increased slowly until failure was observed, and the failure speed was recorded. This\n416 / Vol. 119, JULY 1997 Transactions of the ASME\nCopyright \u00a9 1997 by ASME\nDownloaded From: http://tribology.asmedigitalcollection.asme.org/ on 01/30/2016 Terms of Use: http://www.asme.org/about-asme/terms-of-use",
+            "was repeated several times at various pressures from approxi mately 20-90 percent of static burst pressure for configurations B and G.\nIll Burst Pressure Versus Shaft Speed. The seal was filled with fluid 3 and the shaft speed was increased to 50 percent of the maximum failure speed obtained previously at low pressure. The chamber pressure was increased until the seal burst, and the burst pressure value was recorded. This was repeated at 70 and 90 percent of previously recorded low pres sure failure speed, for configurations B and G.\nResults and Discussion\nI Failure Speed at Low Pressure. A tabulation of the test results is shown in Table 3. Note that some of the tests were conducted with the shaft assembled eccentric in relation to the pole piece.\nIt is reasonable to assume that seal failure may occur when the centrifugal force of the fluid overcomes the magnetic force acting to retain the fluid in the gap. The centrifugal force acts in the radial direction and the magnetic force acts primarily in the axial direction; therefore it is not obvious how these forces interact. Nonetheless, comparing the magni tude of these forces can provide some useful insight into the problem. The (maximum) centrifugal force per unit volume is given by:\nF, = iTT^pDN^ (1)\nThe magnetic retention force per unit volume is given approx imately by the following equation (For the tests where the seal\nwas assembled in the eccentric condition, the value of H in the region of Lamx', i.e., the minimum value of// in the azimuthal direction, is used. This represents the region where the seal would be expected to fail.):\nF\u201e,= MH\n(2)\nA dimensionless number can be defined based on the ratio of F^ to F\u201e, called the force ratio, F^:\nFr lir^pDN^L,\nMH (3)\nThe values of F, at failure are listed in Table 3. They range from a low of 0.22 to a high of 2.37. Figure 3 shows a plot of Fc versus F,\u201e for the stage-on-shaft configurations; the dotted line represents a linear least squares curve fit. The coefficient of determination (r^) is 0.38, indicating a weak correlation. Therefore the force ratio alone appears to be a poor predictor of seal failure. This means that varying the strength of the magnetic force which acts on a particular seal (at least within the range of F,\u201e values tested here) has at most a weak effect on the failure speed. Mitsuya et al. (1993) report that magnetic field strength in the gap has an important effect on the critical speed at which splashing begins; however, they worked with much weaker fields (1.6-3.2 X 10^ A/m) and lower fluid mag netization (0.0180 T) . This would correspond to F,\u201e values which are at least an order of magnitude lower than those tested here. Also, it is not clear how they define splashing; tiny local ized splashing (i.e., that which never leaves the gap area) would not be considered failure according to the definition used in this paper.\nN o m e n c l a t u r e\nA-raiio = aspect ratio, dimensionless D = shaft diameter, m\nFc = centrifugal force per unit volume, N/m' FM = magnetic force per unit volume, N/m' F\u00ab = force ratio, dimensionless H = magnetic field intensity, A/m K, = torque sensitivity of motor, N-m/\namp\nLc = radial gap, m Laimix - maximum radial gap, m\nLs = effective stage length, m Lsiwm = nominal stage length, m\nM = ferrofluid magnetization, Tesla A\u0302 = shaft speed, rev/s P = burst pressure of seal, KPa\nRe = Reynolds number, dimensionless Re.f = Reynolds number at seal failure,\ndimensionless\nT = torque, N-m Ta = Taylor number, dimensionless\nTa^ = critical Taylor number, dimen sionless Ta/ = Taylor number at seal failure, di mensionless\np = ferrofluid density, kg/m' T] = motor efficiency, dimensionless fi = viscosity, N-s/m^\nJournal of Tribology JULY 1997, Vol. 119 / 417\nDownloaded From: http://tribology.asmedigitalcollection.asme.org/ on 01/30/2016 Terms of Use: http://www.asme.org/about-asme/terms-of-use",
+            "Further evidence that force ratio is an inadequate model of high speed seal failure lies in the fact that failure speed is apparently related to fluid viscosity. The viscosity values listed in Table 3 are the estimated viscosities of the fluid at the time of failure. The torque due to fluid friction was derived from the motor current as described in Appendix B. For all configurations tested Lg < D; therefore torque transmitted by fluid friction (assuming laminar flow) is given by:\nT = 2La\nTherefore the fluid viscosity is:\nM = 2TLa\n\u2022K^ND^Ls\n(4)\n(5)\nIt was observed that failure speed is roughly proportional to fluid viscosity, and inversely proportional shaft diameter, maximum radial gap, and fluid density. This suggests that a dimensionless number in the form of Reynolds or Taylor number may be a better way to characterize seal failure. The Reynolds number is given by:\nThe values of Ta^ in Table 3 range from 0.7Ta,. to 3.2Ta,; most seals failed in the range where Taylor vortex flow would be expected to occur, but well below the predicted onset of true turbulence. Theory predicts, and empirical studies have shown, that at Ta > Ta^ the actual torque begins to deviate from that predicted in Eq. (4 ) , with the difference increasing as speed is increased beyond Ta^. For the data in Table 3, the average ratio of failure speed to critical speed is 1.37. Because the failure speeds are generally close to the critical speeds the actual viscosities are not expected to differ greatly from those calculated by Eq. ( 4 ) ; therefore, for simplicity the laminar torque relationship Eq. (4) is used to estimate viscosity.\nThere is a significant degree of scatter in the Taylor number values in Table 3 and no discemable patterns in the data, so it can be concluded that Taylor number is also a poor predictor of seal failure. With the exception of configuration E, the Reynolds number values in Table 3 fall into a relatively narrow range, with the values for configuration E being generally higher. Con figuration E is the geometry with the highest aspect ratio, and a plot of Rc/ versus A\u2122\u201e\u201e, Fig. 4, reveals that there is a slight correlation. Re/ can therefore be estimated by the following (stage-on-shaft configurations only):\nRe\nThe Taylor number is:\n(6) Re/=72(A\u2122\u201e\u201e) + 329\nA \u2014 '^ ratio\n(8)\n(9)\nTa 27:'DN'Lh\u201e,^p'\n(7)\nThe Reynolds and Taylor numbers at failure, Rc/and Ta/, for each test are listed in Table 3. Viscous flow in the annulus of two concentric cylinders, such as in the seals tested here, has been studied extensively. It is known that at Taylor numbers above a certain critical value Ta^ (ap proximately 1700), basic laminar flow becomes unstable and gives way to cellular vortex flow, or Taylor vortex flow.\nFerrofluid\nFluid 1 Fluid 2 Fluid 3 Fluid 4 Fluid 5 Fluid 6\nDensity (kg/m')\n1265 1265 1240 1265 2220 1909\nSaturation magnedzation (T)\n0.0449 0.0454 0.0450 0.0465 0.0438 0.0102\nViscosity @ 27\u00b0C (N-s/m')\n0,052 0.037 0.650 1.200 2.670 0.535\nA dashed line representing Eq. (8) is shown in Fig. 4. Because the stage width, L.s\u201e\u201e\u201e,, was constant for all tests, there is no way to separate a change in gap from a change in aspect ratio, so it is not known for certain whether it is\n418 / Vol. 119, JULY 1997\nTransactions of the ASME\nDownloaded From: http://tribology.asmedigitalcollection.asme.org/ on 01/30/2016 Terms of Use: http://www.asme.org/about-asme/terms-of-use"
+        ]
+    },
+    {
+        "image_filename": "designv11_60_0003887_978-3-642-73890-6-Figure2\u00b72-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0003887_978-3-642-73890-6-Figure2\u00b72-1.png",
+        "caption": "Figure 2\u00b72: Modular and adaptable fixture for sheet metal parts",
+        "texts": [],
+        "surrounding_texts": [
+            "752\nThe objective of this paper is to present an automated manufactwing system for the drilling of sheet metal parts. All stand-alone systems such as the robot, a set of reconfigurable flXtores. and the CAD/CAM wOItstation have been integrated into a Flexible Manufacturing System. This system analyzes and evaluates a given fixturing layout from a stress-slrain point of view and sets the reconfigurable flXtores automatically using a robot manipulator. The developed software consists of a database system and four major modules: a workpart geometry display. a routine for flXturing analysis. a fIXture sequential control software. and a program to configure the fixtores automatically.\n2. Recontigurable Fixtures and The Drilling of Sheet Metal Parts\nThe hardware for tile proposed reconfigurable fixture was designed and manufactured at M.I.T [1.2.18.19.6]. The design takes into accOlDlt tile automated assembly of the fixtores by a robot manipulator. The reconfigurable fixtwing system consists of a fixture bed (baseplate). vertical supports (fIXture elements) and two locating pins. The prototype of the fixturing system with a typical workpart is shown in Figure 2-1. The main idea of the robotic reconfigurable fixtwing concept is to support the sheet metal from beneath with vertical support flXturing elements on a universal flX:ture base. The base consists of a flat plate with evenly spaced machined T -slots. as shown in Figure 2-1. Vertical supports are stored initially in a fIXture magazine and then the robot grasps elements from the magazine and positions them onto the baseplate. Each slot of the baseplate has a tapered entrance to allow easy insertion of the fixturing elements. A typical element shown in Figure 2-2. is adjustable in x. y and z directions.\nA typical sheet metal part, shown in Figure 2-3. has two tabs as a pair of reference points to locate the part which is a common practice in fuselage manufacturing. The sheet metal is placed and located accurately relative to the base by two pins. which constrain the workpart in x-y plane. Only one of the pins is fIXed. and therefore its position is known. The second pin originally is located at its \"home\" position in the sliding track. It prevents the workpart from rotating about the first pin and is adjusted to the fmal location by the robot. During the assembly of the fIXture. the vertical supports are inserted in the appropriate slot, slid along the desired position and oriented appropriatJy. Therefore. the automatic assembly of these fIXtures require the following parameters: slot number. sliding distance. orientation and the supporting shaft height. The assembly of the fIXturing elements must take into account the collision and interference problem since the elements can share the same or adjoining slot. However. in dealing with flexible parts. simply constraining the six degrees of freedom is not sufficient to prevent undesirable deformations while loads are applied. Redundant supports are required so that the part will not distort from its original shape. even wben no machine loads are present. This is especially important in a manufacturing environment requiring high accuracy. The design and assembly of the reconfigurable fixtures must address this inherent characteristic of sheet metals: the fixture must define the workpart shape completely. The possibility for the part to yield or buckle under the drilling loads must be checked.\nThe drilling process can be summarized as two distinct actions: one at the chisel and the other at the lips of the drill bit. The entire process can be modelled as a net thrust T and a torque M by assuming symmetry about the drill point. Several ways were developed to evaluate the thrust and the torque. The equations to determine the drilling thrust and torque along with the parameter definitions and values can be found in reference [141.\nT = 2 Kd Ff F, B w + Kd d2 J W M = Kd Ff F,. A W\n3. Fixturing Analysis\n3.1. Finite Element Procedure Failure analysis constitutes a major element in the fixturing analysis. A major type of failure is the distortion of the part from its true shape after machining due to the inherent low stiffness of the sheet metal. This failure situation can be avoided by performing stress analyses of the system. Finite element analysis is used to check whether the workpart will sustain permanent deformation or buckle under the drilling loads. and to evaluate the fixturing layout considered.\nIn the model of Figure 2-3 the vertical supports are assumed to have point contact with the workparL They only constrain the part in the Z direction. Two locating pins are considered as reaction forces to restrain the part in the X -Y plane. An eight-node double curved shell elements (S8R) \u2022 with four integration points. was used in modelling the sheet metal parts. The machining fOICeS are modelled as forces and moments applied at the corresponding nodes. The vertical supports and the",
+            "753",
+            "754\nlocating pins are introduced as part of the boundary conditions. The degree of freedom in the z-direction is constrained at the nodes where the fIXtures are located.\n3.2. Fixturing Layout Evaluation Excessive stress level induced at any location within the sheet metal part will cause the part to defonn plastically from its original shape. Three and four fIXturing elements at different mdii from the drilling point were considered. The stress distribution within the workpart for different layout configumtions , were investigated. In addition, several plate curvatures were examined. Specifically, nine different mdii of curvatures are investigated: \u00b125 in,\u00b150 in,\u00b175 in,\u00b1 100 in, and infmity. The fixture elements are located along the diagonal axes as shown in Figure 3\u00b71.\nThe nonnal stresSes in 11, 22 directions and shear stress as a function of fixture support locatinn and the applied forces were analyzed. 1be simulations on yielding were performed for the following cases: flat plates. shells with positive mdius of curvature (convex shells) and shells with negative radius of curvature (concave shells). (19).\nFigures 3\u00b72, 3\u00b73111d 3-4 show the normal and shear stresses for the flat plate. The normal stresses in II and 22 direction at the loading point iDcIQIe willi the increasing radia1 distance of fixture elements from the center of the plate. These stresses are compressive when die fixtures are near the center of the plate because the applied forces are directly transferred to the fIXtures resulting in small deflections. In this case the workpart fibers are compressed. However. the stresses are tensile as the fIXture elements move away from the center of the plate. Due to the locating pins constraint, the plate stretches resulting in large deflections. Both normal stresses 11 and 22 at the fixtures increase slightly when fixtures are near the center of the plate. 'but stresses increase drastically as fiXtures near the edges of the plate. The stresses are always tensile because material is pulled toward the loading poinL The normal stresses at the pins are always compressive. and they decrease with the increasing mdius of fixtures from die center of the plate. The shear stresses distribution at the point of pins and applied for are relatively constant and low. but the shear stresses at the fixtures increase as the radius of fixture from the center increase. After reviewing all the plots. the best fixturing layout is at a radiusllength ratio of 0.55. Length and radius were defined previously in Figure 3-1. Several simulations have been conducted for convex and concave shells in reference [19). The best fixturing locations for all positive-curvature shells are given by a radiUS/iength ratio ranging from 0.45 to 0.65. As for the concave shells. any fIXturing layout sbouId be feasible as long as the fixture elements are within a 0.65 radiUS/iength ratio. As for the three support fIXturing system. the fixtures are placed 120 degrees apart and six fixturing layouts for each radius of curvature have been studied. All the collected data show that the stress profiles for all curvatures have more or less the same pattern as the four-point fixturing. However. the three-point support case bas higher stress than the four-support fixturing as expected. The stresses at different locations. which are loading point, locating pins. and fixtures for different curves are given in reference [19).\nThe yielding analysis cannot ensure that the workpart will not defonn from its true shape. Permanent defonnation. which is still possible. can be caused by buckling where the workpart can suddenly snap dIrougb. The buckling phenomenon will happen when forces are applied to the convex faces. In order to analyze the buclding of convex shells. the critical load has to be determined. This can be done by controlling the displacement of the node at the center of the shell to move down a certain"
+        ]
+    },
+    {
+        "image_filename": "designv11_60_0003865_ijeee.39.1.6-Figure2-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0003865_ijeee.39.1.6-Figure2-1.png",
+        "caption": "Fig. 2 Equivalent circuit of an armature-controlled d.c. motor.",
+        "texts": [
+            " However, it is important to note that the same controller could be easily implemented by using a digital controller if this laboratory takes place after a course in discrete-time control systems. A mathematical model for the d.c. motor-generator group depicted in Fig. 1 can be obtained simply by considering the equivalent circuit of the armature-controlled d.c. International Journal of Electrical Engineering Education 39/1 at UNIVERSITE LAVAL on July 13, 2015ije.sagepub.comDownloaded from motor of Fig. 2, where va(t) and ia(t) denote, respectively, the input voltage and the current in the armature circuit, w(t) is the shaft velocity and J and f are, respectively, the load inertia and the bearing friction. It is not difficult to show that:5 (1) where te = La/Ra, tm = J/f, Km is the torque constant, Ke is the counter electro-motor force and td(t) denotes the disturbance torque which, in this case, appears when a load is connected to the generator terminals. However, since La/Ra << 1 then for low and intermediate frequencies tes + 1 \u00aa 1 and, therefore, a simpler model for this system is given by: (2) where In order to obtain a complete model for the system, it remains to take into account the effects of the disturbance current (ig(t)) and the tachometer as well"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0000116_s00707-016-1611-8-Figure2-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0000116_s00707-016-1611-8-Figure2-1.png",
+        "caption": "Fig. 2 A rigid body system. a A rigid body. b The equivalent force system",
+        "texts": [
+            " Here, the generalized active force Kr and the generalized inertia force K \u2217 r , respectively, are defined as Kr = \u03bd\u2211 i=1 v\u0303Pi r \u00b7 Fi (r = 1, . . . , p), (8) K \u2217 r = \u03bd\u2211 i=1 v\u0303Pi r \u00b7 (\u2212miai ) (r = 1, . . . , p), (9) Many researches regarded Eqs. (7) as Maggi\u2019s equations or the relevant form of Gibbs\u2013Appell\u2019s equations. 2.2 For a rigid body system Now, we consider a typical particle Pi of rigid body C (formed by \u03bd particles, say P1, P2, . . . , P\u03bd) belonging to a simple nonholonomic system S possessing p degrees of freedom in a Newtonian reference frame N (Fig. 2a). The angular velocity of C , \u03c9C , and the velocity of mass center C\u2217, vC \u2217 , can be expressed as \u03c9C = p\u2211 r=1 \u03c9\u0303C r ur + \u03c9\u0303C t , (10) vC \u2217 = p\u2211 r=1 v\u0303C \u2217 r ur + v\u0303C \u2217 t (11) where the vectors \u03c9\u0303C r , \u03c9\u0303 C t , v\u0303C \u2217 r , and v\u0303C \u2217 t are functions of q1, . . . , qn , and t . \u03c9\u0303C r and v\u0303C \u2217 r are called r th partial angular velocity and r th partial velocity of C , respectively. Using the kinematic equations, the velocity vi and the acceleration ai of Pi are given as vi = vC \u2217 + \u03c9C \u00d7 ri (i = 1, . . . , \u03bd), (12) ai = aC \u2217 + \u03b1C \u00d7 ri + \u03c9C \u00d7 (\u03c9C \u00d7 ri ) (i = 1, . . . , \u03bd) (13) where ri is the position vector fromC\u2217 to Pi ; \u03b1C and aC \u2217 are the angular acceleration ofC and the acceleration of C\u2217. If a set of contact and/or distance forces acting on C is equivalent to a couple of torque TC along with a force FC whose line of action passes through C\u2217 (Fig. 2b), then the generalized active forces and generalized inertia forces for the rigid body [10] are written as CKr = v\u0303C \u2217 r \u00b7 FC + \u03c9\u0303C r \u00b7 TC (r = 1, . . . , p), (14) CK \u2217 r = v\u0303C \u2217 r \u00b7 FC\u2217 + \u03c9\u0303C r \u00b7 TC\u2217 (r = 1, . . . , p) (15) where the resultant inertia force FC\u2217 and resultant inertia torque TC\u2217 , whose line of action passes through C\u2217, are given by FC\u2217 = \u2212MaC \u2217 , (16) TC\u2217 = \u2212\u03b1C \u00b7 I \u2212 \u03c9C \u00d7 I \u00b7 \u03c9C . (17) Here, M is the mass of C and I is the central inertia dyadic of C . Substituting Eqs. (14) and (15) into Eq",
+            " (52) At the same time, the kinetic energy function T , the functions Pr (r = 1, . . . , p), and Gibbs function G can be expressed as T = 1 2 MvC \u2217 \u00b7 vC \u2217 + 1 2 \u03c9C \u00b7 I \u00b7 \u03c9C , (53) Pr = MvC \u2217 \u00b7 d dt ( \u2202 vC \u2217 \u2202 ur ) + \u03c9C \u00b7 I \u00b7 d dt ( \u2202\u03c9C \u2202 ur ) (r = 1, . . . , p), (54) G = 1 2 (MaC \u2217 \u00b7 aC \u2217 + \u03b1C \u00b7 I \u00b7 \u03b1C + 2\u03b1C \u00b7 \u03c9C \u00d7 I \u00b7 \u03c9C ). (55) For brevity, formal proofs of Eqs. (53)\u2013(55) are omitted. 5.2 Scalar forms Let c1, c2, c3 form a right-handed set of mutually perpendicular unit vectors, each parallel to a central principal axis of C , but not necessarily fixed in C (Fig. 2a), and \u03b1i , \u03c9i , Ii , ai , vi , Fi , Ti be defined as \u03b1i = \u03b1C \u00b7 ci , \u03c9i = \u03c9C \u00b7 ci , Ii = ci \u00b7 I \u00b7 ci (i = 1, 2, 3), (56) ai = aC \u2217 \u00b7 ci , vi = vC \u2217 \u00b7 ci , Fi = FC \u00b7 ci , Ti = TC \u00b7 ci (i = 1, 2, 3). (57) Substituting into Eqs. (51)\u2013(53) and Eq. (55) from Eqs. (56)\u2013(57), one can thus generate the explicit forms of functions CKr , A, T , and G as follows: CKr = F1 \u2202v1 \u2202ur + F2 \u2202v2 \u2202ur + F3 \u2202v3 \u2202ur + T1 \u2202\u03c91 \u2202ur + T2 \u2202\u03c92 \u2202ur + T3 \u2202\u03c93 \u2202ur (r = 1, . . . , p), (58) A = F1a1 + F2a2 + F3a3 + T1\u03b11 + T2\u03b12 + T3\u03b13, (59) T = 1 2 M(v21 + v22 + v23) + 1 2 (I1\u03c9 2 1 + I2\u03c9 2 2 + I3\u03c9 2 3), (60) G = 1 2 [M(a21 + a22 + a23) + I1\u03b1 2 1 + I2\u03b1 2 2 + I3\u03b1 2 3] \u2212 \u03c92\u03c93\u03b11(I2 \u2212 I3) \u2212 \u03c93\u03c91\u03b12(I3 \u2212 I1) \u2212 \u03c91\u03c92\u03b13(I1 \u2212 I2)"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0001986_002071799221244-Figure1-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0001986_002071799221244-Figure1-1.png",
+        "caption": "Figure 1. Behaviour of the minimum distance functions.",
+        "texts": [],
+        "surrounding_texts": [
+            "In the following theorem, we derive the minimum distance functions (MDF) that are used as tools for the calculation of stability margins in this paper. Theorem 2: The minimum distance (7) of the nominal parameter vector (6) to (a) the set of parameter vectors a , at which QR a , z 0 9 where QR a, z is de\u00ae ned in (3) and (4), is given by q p z QR a, z m k 0 wk F Rk z q 1 /q 10 (b) the set of the parameter vectors a , at which QI a , z 0 11 where QI a, z is de\u00ae ned in (3) and (5), is given by t p z QI a, z m k 0 wk F Ik z q 1 /q 12 where q p/ p 1 and z is an arbitrary point in the sweeping range. Proof: See the appendix. h The MDFs derived in Theorem 2 can easily be visualized in the entire range of the sweeping parameter. The plots of the functions provide some insight to the interlacing property presented in Theorem 1. Let the entire range of the sweeping parameter z be denoted by z S, z F . Typical plots are sketched in \u00ae gure 1. In the remainder of this section, two important properties are established for the MDFs: Proposition 1: The minimum distance functions q p z and t p z cannot have any local minimum in the range z S, z F other than the zeros z R and z I, respectively and the extreme points z S and z F. Proof: The proposition is proved for the real part and an analogous argument can be followed for the imaginary part. Consider the function q p z in an interval which ends at a zero of QR a, z , namely z R ( \u00ae gure 2). Since q p z R 0, it is trivial that z R is a local minimum of q p z . This function can also have a local minimum at the extreme points z S or z F, if its de- rivative is respectively positive or negative at these points. To proceed with a contradiction argument, assume that there exists a local minimum at z z A. From Theorem 2, the minimum distance of a to the set of the parameter vectors a at which QR a , z A vanishes is given by h q p z A . Then, if the parameter vector a is perturbed by h, there exists at least one perturbed parameter vector ~a at which QR ~a, z has a zero at z z A. On the other hand, since QR a, z is a continuous function in a, continuous increase of the perturbation d p a,~a results in the continuous movement of the zero z R along the z -axis (see Argoun (1987) for a more detailed discussion in a similar case). The movement of z R as the perturbation increases to h is shown by the segment z D z B in \u00ae gure 2. As the value of the perturbation reaches h, there will be at least one zero for the perturbed family at z z A. This implies that a zero of QR ~a, z has jumped from the segments z D z B to the point z A. The latter contradicts the continuity of QR a, z in a. This completes the proof. h Robust controllers for linear systems 269 D ow nl oa de d by [ St on y B ro ok U ni ve rs ity ] at 2 3: 24 2 1 O ct ob er 2 01 4 Proposition 2: There exists at least one z in each interval D at which the following equation holds: G z . q p z t p z 0 13 Proof: The proposition is proved only for the interval z R, z I , but a similar argument can be followed for the interval z I, z R . The function G z is a continuous function in the entire range of z R, z I except for the points where the denominator of q p z or t p z tends to zero. The discontinuity can be removed by assigning some large values to the functions at these points. At z z R, q p z R 0 and G z R t p z R a 1, where a 1 is a real negative number. As z moves along the interval to z z I, t p z vanishes and it is obtained that G z I q p z I a 2, where a 2 is a real positive number. From the continuity of the function G z in the interval z R, z I , it is concluded that there exists some z in the interval at which the function crosses the zero line, where equation (13) is satis\u00ae ed. h"
+        ]
+    },
+    {
+        "image_filename": "designv11_60_0000942_j.renene.2016.11.054-Figure13-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0000942_j.renene.2016.11.054-Figure13-1.png",
+        "caption": "Fig. 13. The schematic diagram of specialization improved nonlocal means denoising method for damage identification of generator bearing.",
+        "texts": [
+            " Specialization improved nonlocal means denoising for damage identification As the description in the introduction, it should be noted that the proposed method can be used for generator bearing damage identification along with envelope demodulation techniques (under constant speed condition) or order tracking analysis method (under variable speed condition) to ensure the damage feature extraction and identification efficiency [20,40]. To sum up, the schematic diagram of specialization improved nonlocal means denoising method for wind turbine generator bearing damage identification can be presented in the flow chart as displayed in Fig. 13. Meanwhile, the procedure of the proposed method for the mentioned engineering task can be summarized as follows: 1) Collect the generator bearing damage vibration signals using acceleration sensors; 2) Perform angular re-sampling if vibration signals acquired in variable speed condition. 3) Determine the neighborhood selection and so on based on the actual collected signals; 4) Carry out the specialization improved nonlocal means denoising method on the vibration signals; 5) Get the denoised coefficients and then output the corresponding Hilbert envelope spectrum or order spectrum; 6) Conduct bearing damage identification through the Hilbert envelope spectrum or order spectrum"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0003855_978-3-7908-1835-2_5-Figure5-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0003855_978-3-7908-1835-2_5-Figure5-1.png",
+        "caption": "Fig. 5. Membership functions.",
+        "texts": [
+            " and Xm is Ai,m Then Yl = Yi,l and Y2 = Yi,2, where i = 1, ... ,n is the rule index, Ai,j is a fuzzy set for the jth linguistic variable in the ith rule, defined on the universe of discourse for the jth vari able, and Yi,l and Yi,2 are real constants. Note that these values generated for the outputs are not simply the consequent parts of rule Ri because for any input condition it is likely that several rules will be firing, and the outputs Yi,l and Yi,2 are generated as the sum of the individual rule consequents. (See Fig. 5 for the case of a single output.) Each membership function for this controller is defined as a Gaussian function as shown in Fig. 5(a), or as a triangular function, which is piece wise differentiable, as shown in Fig. 5(b). The reason for this choice is to permit the development of a gradient based training algorithm. For every input, the universe of discourse is initially covered by a set of uniformly spaced Gaussian, or triangular membership functions. The membership functions are given by /-Lij = e (2) and /-Lij = bij' 'J 2 J 'J 2 { I - 21 xc ai jl for a .. - !!ii < X\u00b7 < a .. + bij o , otherwise, (3) respectively. For the application of the rules, we need to define a fuzzy inference mech anism, and in this case we use a T-norm in terms the product operator (multiplication)",
+            " This means that the firing strength of each rule is given as Ui = J.li,1J.ll,2 '\" J.li,m, i = 1, ... ,n, (6) where J.li,j is the degree of membership of Xj in the fuzzy set Ai,j. The outputs are required to be crisp values and if de-fuzzification is performed the centre of-gravity method, this crisp value is given (for the case of a single output) as n Y = L UiWi, (7a) where i=l U\u00b7 - . Ui = \",n . L.Jj=l Uj (7b) The general structure of this system in the case of two inputs and a single output is shown in Fig. 5. This representation is equivalent to an architec ture in the form of a radial basis function artificial neural network. In this structure, the first layer performs fuzzification for each input variable in each linguistic rule and outputs from this layer are membership values which are fed to the next layer where the T-norm (multiplication) is applied (see (6)). The result is a firing strength for each rule, and in the next layer the firing strengths are normalised (see (7b)). In the last layer, the output is computed as the weighted sum of the incoming signals (see (7a)), that is, a weighted sum of the individual rule consequents Wi"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0000414_978-3-662-46463-2_40-Figure40.2-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0000414_978-3-662-46463-2_40-Figure40.2-1.png",
+        "caption": "Fig. 40.2 The different phases in a walking cycle",
+        "texts": [
+            " Powered prosthesis movement research is mainly composed of terrain conditions prejudgment and knee joint angle control. According to the different terrain conditions, such as level ground, upslope or down slope, stairs ascent or descent [16, 17], different knee angle swing modes are outputted in the prosthesis control. A movement database is established through using the movement data from the healthy knee joint as reference. The database contains the motor\u2019s pulse control signal of direction and displacement in the different gait phase under a variety of terrain conditions. As shown in Fig. 40.2, the gait phase divides a walking cycle into four phases, which are early stance phase, middle stance phase, late stance phase, and swing phase, respectively. The phases can be recognized via two pressure switches mounted under the sole of the prosthesis foot, one under the ball flat and another under the heel. The prosthesis knee joint used in this paper is shown in Fig. 40.3. The four connecting rods driven by the motor make the change of knee joint\u2019s angle. The relationship between prosthesis knee joint angle (h) and motor shaft displacement (L) is described in Eq"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0003731_ias.1995.530303-Figure4-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0003731_ias.1995.530303-Figure4-1.png",
+        "caption": "Fig. 4 (d\u201d-qe), (ds-qs), (de,-q\u201d,) and a-b-c axes for PM machine.",
+        "texts": [
+            " 4 ALGORITHM DEVELOPMENT FOR THE ESTIMATION OF ROTOR POSITION AND SPEED Transforming ds-qs voltages and currents into de-q\u201c reference frame. (40) I e - cos8 L\u2018 +I\u201d sine, 4 I E 4 = - I ; sine, + I S 4 cose, (41) 4.1 ROTOR POSITION ESTIMATION Modified de-q\u201c equivalent circuits shown in fig. 3 can be used to develop a sensorless position and speed measuring scheme for the drive system. Estimation algorithm is developed in de-q\u201c reference frame, so that this could be used effectively in vector control drives. Fig. 4 shows the synchronous and stationary reference frames for the permanent magnet synchronous machine. The initial rotor position is assumed to be equal to 8 and is shown in fig. 4 where e e e =U3 t (44) Three phase to stationary reference transformation results in equations (36)-( 39) and (41) into equation (45) and equating the resulting equation with equation (42), an expression for 8 , can be obtained as shown in equation (48). 21 0 E2+\") (60) 2( AC+UU)+J4( AC+RD)2-4 c2 +D2 -$2(1+)2 ( 2 ( c , ~ 2 + D 2 - $ 2 ( 1 - E ) 2 ) Let x = ( ? ~ + E R ~ ) I ; +[L] +L,,,~( i -E)]pi l ; -13 e (.y - E L , ~ ) V $ (49) w ,= Y = v ; - ( ~ ~ , ~ + E R ~ ) I ~ -p,] + L ~ ~ ~ ~ ( I-E)]PI~ - 1 2 w e (L4 EL^^^) (50) Since the values of d and q axes inductances are in the range of few micro Henrys to mili Henrys, and the value of 0 is in generally in the range of 0"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0002762_s0094-5765(98)00040-x-Figure2-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0002762_s0094-5765(98)00040-x-Figure2-1.png",
+        "caption": "Fig. 2. The coordinate system used in the derivation.",
+        "texts": [
+            " DERIVATION OF KINETIC ENERGY AND POTENTIAL ENERGY For trajectory control of a manipulator, the computed torque method based on the inverse dynamics is often used. To this end, it is desirable to reduce the number of calculations involved [10\u00b1 14]. Here the equations of motion in recursive form are derived for the MDM, and it is shown that they are suitable for the inverse dynamical computations. The model treated here is a variable geometry manipulator interconnected by n \u00afexible slewing links and n \u00afexible deployable links. Note, the system can serve as a model for a manipulator that is free to translate on a space platform. Figure 2 shows coordinate systems, associated with the MDM, used in the derivation. Consecutive numbers are assigned to the links from one end of the manipulator to the other, and the joint i connects the link i\u00ff 1 to the link i. The coordinate system Fi is \u00aexed to the joint i at the link i with the xi-axis coincident with the link direction in the undeformed state. Note, the deployable link 2j and the slewing link 2j\u00ff 1 are in contact at the joint 2j\u00ff 1 only. Direction of the deployable link 2j in the undeformed state coincides with the direction of the slewing link 2j\u00ff 1 in the deformed state"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0002430_jsvi.1996.0773-Figure2-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0002430_jsvi.1996.0773-Figure2-1.png",
+        "caption": "Figure 2. Numerical model 1: a system with n rotors connected by an identical coupling.",
+        "texts": [
+            " Equation (A9) yields the unbalance responses as pf =Ho ffWV2, pb =Ho bfWV2. (26) The present example deals with an artificial system with n-span rotors to investigate the computational efficiency of the proposed method in comparison with the direct computation method. Both the proposed method and the direct computation method are implemented on a PC with Matlab [16]. The model, in which every substructural rotor is assumed to be identical and serially connected to the neighboring rotor with a coupling, is shown in Figure 2. Two bearings support each rotor. The detailed specifications of the T 2 Specifications of the rotor in numerical model 2 Shaft Length 1\u00b72 m Diameter 8\u00b70 cm Young\u2019s modulus 2\u00b70\u00d71011 N/m Density 8000 kg/m3 Number of finite elements 12 (equal length) Disk Mass 20 kg (three identical) Polar moment of inertia 0\u00b7163 kg m2 Diametral moment of inertia 0\u00b7085 kg m2 Location (distance from left) Nodes 5, 6 and 13 (0\u00b74, 0\u00b75, 1\u00b72 m) Bearing 1 Location (distance from left) Node 1 (0 m) Load 29\u00b742 kgf L/D 0\u00b75 C/R 2/1000 Viscosity 9\u00b737 mPas Bearing 2 Location (distance from left) Node 10 (0\u00b79 m) Load 78\u00b784 kgf L/D 0\u00b75 C/R 2/1000 Viscosity 9\u00b737 mPas T 3 Specification of fluid film bearings used in numerical model 2 Bearing type Parameters Two axial groove Oil groove angle=10\u00b0 Four tilting pad Tilting pad angle=80\u00b0 Preload factor=0 LBP type Five tilting pad Tilting pad angle=60\u00b0 Preload factor=0 LBP type substructural rotor is shown in Table 1"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0003713_robot.1994.350930-Figure1-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0003713_robot.1994.350930-Figure1-1.png",
+        "caption": "Figure 1: 2-DOF planar manipulator",
+        "texts": [
+            " When the system is a t rest (q = 0 the system is stable. In the singularity itself (U, = 01 the right hand side of the above equation is zero again, but reduces from 1 to zero. This implies that the position error in the direction of the singular vector U, becomes uncontrollable, while in other directions it remains stable as will be shown by the following example. Let us now consider the application of the above control strategy to the 2-DOF planar manipulator. 3 2-DOF Planar Manipulator Example The planar manipulator (see Fig.1) which was used in our experiments was actually composed of the second and the third degrees of freedom of PUMA-560 arm, while the other joints were locked. the element (m,m) in the matrix I, - &umum T The Jacobian matrix J which relates between joint velocities and end-effector velocities expressed with re- spect to the base frame has the form where 11,12 are the lengths of the links, si = sin qi , ci = cosqi, i = 1,2, s i 2 = sin(q1 + q z ) , ciz = cos(q1 + q z ) . The Jacobian matrix expressed in the 2nd coordinate frame located at the manipulator tip has the form which is much simpler 2J = [ 12 t "
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0002602_jo00297a046-Figure1-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0002602_jo00297a046-Figure1-1.png",
+        "caption": "Figure 1. Rate of autoxidation of Me3N as a function of 0, pressure ([Me3NIi = 6.5 X M, [NaOH] = 0.10 M , and T = 100 \"C).",
+        "texts": [
+            " The rates of formation of Me2NH and HC0,- (12) IMC Methylamines: A Complete Guide. International Minerals Chemical Co.. 1979 parallel that of the disappearance of Me3N. For prolonged reaction, the concentration of Me2NH reached a maximum and then decreased, attributable to further oxidation of Me2NH. At constant oxygen pressure, the reaction followed pseudo-first-order kinetics for about 2 half-lives and then became slower. The pseudo-first-order rate constants were derived from the first 2 half-lives of the plots and are plotted against the O2 pressure in Figure 1, from which the second-order rate constant k is calculated as 8.7 x atm-' s-l.13 -d[Me3N]/dt = k[Me3N]Po, (5) Such second-order kinetics has also been reported by Riley et al.657 for dimethyldodecylamine for which an activation energy of 19.1 kcal/mol has been reported.I4 We have shown that our results for the autoxidation of Me3N at its natural pH are best described by eq 4, which (13) The data clearly demonstrates the rate dependence on O2 prey sure. However, more data are needed to establish how well eq 5 is followed",
+            " In addition, the retention time of Me3N0 was found to be sensitive to the medium and changed from sample to sample. HCOy was analyzed by IR spectroscopy by its uco band at 1607 cm-'. COP and H, in the head gas were analyzed on a 3.5-ft X '/8-in. Spherocarb column isothermally at 70 \"C with argon as the carrier gas and with a thermal conductivity detector. For analysis of the dissolved COP, a sample was treated with concentrated HC1 in a septum vial and the liberated COP was analyzed by GC as described. Autoclave Experiments. All the experiments shown in Table I1 and Figure 1 were carried out on the Berghof reactor, which was slightly modified by shortening the stirring shaft (which houses the thermal couple) for better temperature control. A solution of desired composition was prepared and loaded into the autoclave, which was then sealed. The reactor was then brought to the desired temperature. For stoichiometric studies presented in Table 11, except for experiment 1, no sample was taken during the run and the product analyses were carried out on samples taken from the cold reactor at the end of the reaction"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0000804_jahs.61.042006-Figure9-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0000804_jahs.61.042006-Figure9-1.png",
+        "caption": "Fig. 9. Pinion tooth geometry and root stress.",
+        "texts": [
+            " 15) cr = Nv1(tan \u03b1a1 \u2212 tan \u03b1) + Nv2(tan \u03b1a2 \u2212 tan \u03b1) 2\u03c0 (10) cos \u03b1ai = rbi/rai (i = 1, 2) (11) where \u03b1 is pressure angle, \u03b1ai is addendum pressure angle, rbi is base circle radius, rai is addendum circle radius. Here rai is obtained in the usual way by adding 1/Pd to the pitch radius, where Pd is diametral pitch. Bending Stress of Pinion Tooth The pinion meshing with face-gear is an involute spur gear. The maximum bending stress at the tooth root is examined to determine the geometric limits of the pinion. The stress calculation employs the conventional method that treats the tooth as a cantilever beam with a distributed load on the pitch circle (Ref. 16), shown in Fig. 9. This load depends on input power, P , and rotation speed of pinion, 1. It is calculated by Ft = P 1rp (12) where Ft is tangent load on the pitch circle and rp is the pitch circle radius. Thus, the simplified stress formula is given as \u03c3max = 6FtH LW 2 (13) where \u03c3max is the maximum bending stress on the tooth root, H is dedendum, L is tooth face width, and W is tooth root thickness. The maximum bending stress at the tooth root must be lower than the yield strength, \u03c3Y , with some safety factor, SF"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0002923_978-94-011-3514-6-Figure3.6-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0002923_978-94-011-3514-6-Figure3.6-1.png",
+        "caption": "Fig. 3.6 Pendulum system rolling un an horizontal plane",
+        "texts": [
+            " We will see that such systems possess an infinite number of natural frequencies, the lowest of which is called the fundamental frequency. The fact that one has neglected the mass of the shaft m' results in one over estimating the natural frequency of the system. One can obtain a lower bound by replacing m by - 20 1m + m') in the expressiun for the natural frequency. m' ~ 4 L Q 88,2 kg => m + m' 388 kg => fo 85 Hz The ac:tlldl fundamental frequency, calculated by means of a finite element program, is 90.2 Hz. It is thus well bounded by the two preceding values: 85 < 90.2 < 97. The pendulum system shown in figure 3.6 rolls without sliding on a horizonl~] plane. It consists of a cylinder of mass M, of moment of irif'rt:iii .T, joined toget.her rigidly by a rod to an assumed point mass m The mass of the rod being neg1 igible, let us establish, by eli fferentiating the mechanicil1 energy, the differential ',quation fnr small mllvements ilbout the equilibrium pusition. oC' _ { RR0} - 21 - The velocity of C' is obtained by differentiation, whence .. t Re - L Sine} OA' = -R + L cose -> => VA' [ (R - L cose; eJ - L sine \u2022 e Knowing the vector expression for the velocity, one can calculate the kinetic energy of the system 1 +-m 2 T = 2 82 (J + M R2 + m (R 2 + L2 - 2 R L cose)) 2 (3"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0003042_iros.1993.583851-Figure4-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0003042_iros.1993.583851-Figure4-1.png",
+        "caption": "Fig. 4: an example of grasp planning",
+        "texts": [
+            " GEP(obj,env,pl) = GP(obj,pl) - COP(env,pl) (8) Further more these GP, COP and GEP can be sliced by orientation Oh (rotation around the y-axis) of the gripper and the sliced subsets are called GPO(obj,pl,Bh), COPO(mv,pl,Oh) and GEPO(obj,env,pl, Oh) (see Eq.9). GP(obj,pl) N UGPO(obj,pl,Bh) COP(env,pl) N UCOPO(env,pl , 8h) GEP(obj, env,pl) N U GEPO(obj, env,pl, 8h) Bh Rh eh , G'EPO is calculated from GPO and COPO as follows. GEPO(obj, env,pl, 8h) = GPO(obj,pl,8h) - COPO(env,pl, 8h) GPO, COPO and GEPO are 2-dimensional polygons. GPO and COPO can be calculated by the grasp planning methods [l]. Fig.4 shows an example of calculating GPO and COPO. The object to be grasped is the smaller rectangular parallelepiped object and the gripping plane pll is centered between the face on the visible side of the object and the face of its opposite side. Orientation of the gripper 8hl is illustrated in Fig.4. The areas shown in Fig.4 are the area of the origin of the gripper. The area GPO(object,plI,8hl) is the grasp candidates of the object. The aiea COPO(obstacle,pl~, 8hl) is the collision area between the gripper and the obstacle. The area COPO(table,pll,8hI) is the collision area between the gripper and the table. G and GE are calculated according to Eqs 7 through 10 from all GPO and COPO in every orientation and every gripping plane. In order to check whether the goal obj, can be achieved by a pick-and-place operation or not, Eq",
+            " Sopen is the distance between the outside of the two fingers when the gripper approachs to object, a is distance between the gripping plane and obstacle, and 6 is angle between the gripping plane and the upper surface of the table. (see Fig.9) Finally in the case where the gripping plane and the upper surface of the t.able are not parallel, in order to escape from the situation where object cannot be grasped, obstacle or object is translated out of three edges of the rectangle constructed by Dal, Da2 and Db (see the shaded portion in Fig.10). Dal and Da2 are the sliding distance calculated in the direction of parallel to the cross line, and Db is in the direction of perpendicular. Fig.10 is vertical viewing of Fig.4. Since, when object is translated toward the fringe of the table, there is a risk it falls down, stability must be ensured. - bject (3.5 Planning of Sliding Motions for a Manipulator and Selection of Plans In this section, we describe motion planning of the manipulator in order to slide object or obstacle from where it cannot be grasped to where it can be. A sliding operation is a parallel translation without rotation. Our system plans motions for sliding operations by pushing on the side faces in the following steps as shown in Fig"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0000897_978-3-319-50472-8_2-Figure2-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0000897_978-3-319-50472-8_2-Figure2-1.png",
+        "caption": "Fig. 2. Case 3.1.",
+        "texts": [
+            " 1 the strings generated by r1 are {0, 1, 1, 1} in clockwise direction and {1, 0, 0, 1} in anticlockwise direction. \u2013 Step 2: - The robot checks the occupancy rate on both sides i.e. the number of nodes occupied by other robots on each side. The robot may do so simply by counting the number of 1s and 0s in the strings. \u2013 Step 3: - This is the most important step since in this step the robot makes a decision on its movement. The cases are listed as follows: - \u2022 Case 3.1: - If the occupancy rate is nil on both sides then the robot does not make any movement (Fig. 2). \u2022 Case 3.2: - If the occupancy rate is equal on both sides then the robot makes one hop movement to the side with closer neighboring occupied nodes (Fig. 3(a)) or any of the sides if there is a tie (Fig. 3(b)). \u2022 Case 3.3: - If the occupancy rate is more on the counter-clockwise direction but the clockwise string is nil then the robot does not make any movement (Fig. 4(a)) else it makes one hop movement to the counterclockwise direction (Fig. 4(b)). \u2022 Case 3.4: - If the occupancy rate is more on the clockwise direction then the robot makes one hop movement to the clockwise direction (Fig"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0002354_s0925-4005(98)00088-4-Figure4-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0002354_s0925-4005(98)00088-4-Figure4-1.png",
+        "caption": "Fig. 4. Cross section of a groove showing stripped-core fibre.",
+        "texts": [
+            " Of course, the ray may be deflected in the wrong direction, however the second boundary, also being rough, may deflect the ray suitably. By the same mechanism, a guided ray may be lost during one of the reflections. The overall process is not efficient but is as effective a process as is possible for guiding light into the side of a normal optical fibre. The cladding has to be removed from the embedded section as it is required that the guiding is within the core to ensure the light is guided along the entire fibre and therefore the deviation must occur at the core boundary (Fig. 4). The number of fibres used in the array is determined by the size of the tube and a compromise between detection efficiency and ease of coupling from the fibres to an optoelectronic detector such as a silicon photodiode or avalanche diode. The current research prototype using this design has an array of 12 fibres placed around a 24 mm tube. The cladding is left intact for the first 5 mm of inserted fibre to allow mode settling before the fibres emerge from the adhesive. The total groove length is 50 mm"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0001181_iwc.2016.8068371-Figure4-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0001181_iwc.2016.8068371-Figure4-1.png",
+        "caption": "Figure 4. Calibration Model",
+        "texts": [
+            " The system realizes measurement of train wheel parameters by \u201clight cross image measurement technology\u201d. Laser device shots on the wheel tread section, and CCD captures the image of the light curves which contain the wheel profile parameters information. The positions of laser lights at the wheel tread are in Figure 2. LDs outside rails shots their lights at the up position of the wheel, while the LDs inside rails shot their light at the down position of the wheel. The laser light image is shown in Figure 3. 978-1-4673-9238-9/16/$31.00 \u00a92016 IEEE Figure 4 shows the calibration model of the system. The pixel information of CCDs has been got through extract chessboard corners\u2019 information at different positions. For every camera, a pixel is corresponding to a line in space, which means every point on space line would be corresponding to a certain pixel [8~10]. Chessboard images and laser lights images were taken at several different positions. While a chessboard image and a laser light image must be taken at the same position. The 3-D world model which contains laser light position information has been built by the calibration process"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0003354_0167-4838(89)90136-2-Figure2-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0003354_0167-4838(89)90136-2-Figure2-1.png",
+        "caption": "Fig. 2. Diagrammatic representation of the tetrameric structure of glyceraidehyde-3-phosphate dehydrogenase with four NAD + spinlabeled analogs bound in the active centers. The diagram is derived from X-ray crystallographic data from Rossmann's group [4J, A and N represent the adenine and nicotinamide portions of the NAD + mulecule. The six-membered ring of the nitroxide spin label ( t ) is shown attached to the adenine moiety at either the N 6 or C8 position. The adenine moieties (A) are closely approximated on adjacent monomers related by the R-axis, but widely separated across the P- and Q-axes.",
+        "texts": [
+            " (Biomedical Division) Introduction The NAD + binding and enzymatic catalysis of glyceraldehyde-3-phosphate dehydrogenase can be examined in novel ways by employing EPR spectroscopy coupled with perdeuterated spin-labeled NAD + analogs with the nitroxide ring on the N 6 or C8 position of the adenine moiety (Fig. 1) [1]. A fundamental prerequisite for the use of a spin-labeled analog of NAD + to invesugate the binding of native NAD + is that the presence of the spin label does not significantly alter the binding or catalytic constants. This condition has been conclusively verified with enzymological techniques by Trommer and co-workers [2] and also with molecular modeling based on X-ray crystallographic data by J.J. Birktoft [3]. As diagramatically represented in Fig. 2, the crystallographic data of Rossmann's group showed that tetrametic glyceraldehyde-3-phosphate dehydrogenase could be considered as a 'dimer of directs' related by the molecular R-axis [4]. The binding of DSL-NAD + to a single monomer or two monomers at distant sites (58 A) related by the Q-axis (Fig. 2) produced a simple hyperfine spectrum (1') (Fig. 3A). Two DSL-NAD + on adjacent monomers within a s!ngle dimer related by the pseudo-2-fold R-axis at 13 A or less interacted to produce a complex dipolar spectrum ( ~ ) (Fig. 3C). The distances between interacting nitroxide radicals correlated well with the X-ray data [2,3]. The special advantage of isotopically substituted DSL-NAD + for studies of this dehydrogenase is that the highly resolved EPR spectra can be quantitatively deconvoluted [3,5,6] into three components: (1) freely tumbling, unbound DSL-NAD+; (2) slowly tumbling, hyperfine spectral components and (3) interacting, dipolar components from two adjacent DSL-NAD +"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0000920_s10958-016-3177-3-Figure1-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0000920_s10958-016-3177-3-Figure1-1.png",
+        "caption": "Fig. 1",
+        "texts": [
+            " We perform the numerical analysis of the stress-strain state inside an elastic body, present its 3D -image, construct the level lines of stresses [2, 7, 9], and determine the localization and value of the stress concentration inside the elastic body. Consider the interaction of a rigid punch with an elastic half plane under the conditions of plane deformation. Under the action of a vertical force P , the rigid punch with a base profile f (x) and a contact region \u2212a \u2264 x \u2264 a is pressed into this half plane (Fig. 1). We denote by E , Ep and \u03bd , \u03bdp the Young moduli and Poisson\u2019s ratios of the half plane and the punch, respectively. We also assume that there are no tangential stresses under the punch. 1 L. Ukrainka East-European National University, Luts\u2019k, Ukraine. 2 Pidstryhach Institute for Applied Problems in Mechanics and Mathematics, Ukrainian National Academy of Sciences, Lviv, Ukraine. Translated from Matematychni Metody ta Fizyko-Mekhanichni Polya, Vol. 57, No. 4, pp. 162\u2013167, October\u2013December, 2014"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0000682_actea.2016.7560128-Figure2-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0000682_actea.2016.7560128-Figure2-1.png",
+        "caption": "Fig. 2. Modeling of the process: To the left the beginning of the process and to the right the indentation process",
+        "texts": [
+            " For this sake, we use an in-plane-out-ofplane separation [8], [9]. In fact, using the PGD, any extra variability can be added as extra coordinate to the problem without increasing the computing time dramatically [4], [5]. This model can be used later on to compare the simulation to the experimental results in order to identify different cartilage properties. The micro indentation process of the cartilage is made by a spherical probe attached to a cantilevered beam as illustrated in figure 1. The cantilever beam can be modeled by a spring as shown in figure 2 which illustrates the approach of the probe in the cartilage. An increasing force then indents the cartilage until reaching a steady state. At this point, we can keep the load for a short period of time then remove the probe. The 978-1-4673-8523-7/16/$31.00 \u00a92016 IEEE 141 stiffness \ud835\udc58 of the spring shown in figure 2 can be modeled as: \ud835\udc58 = 3\ud835\udc38\ud835\udc3c \ud835\udc3f3 (1) Where \ud835\udc38 is the modulus of elasticity of the beam, \ud835\udc3c the moment of inertia of the cross section of the beam and \ud835\udc3f its length. One may note that, function of the time, the contact area between the probe and the cartilage is increasing. Therefore while modeling, one should take in consideration the change of radius of the simulated domain. In order to model the process, we suppose that the probe velocity is small enough in order to be able to work using a quasi-static approach"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0000280_asemd.2015.7453519-Figure1-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0000280_asemd.2015.7453519-Figure1-1.png",
+        "caption": "Figure 1. Structure of the TFSPMLM.",
+        "texts": [
+            " INTRODUCTION Flux-switching permanent magnet linear motor (FSPMLM) has the advantages of high thrust density, well fault performance, and so on. It has a wide application prospect in the new energy power generation, electric drive, and so on [1]. Both the winding and PM of the FSPMLM are set in the primary, and it works on the principle of magnetic flux switching. The secondary iron will generate no small part iron loss, which is different from the permanent magnet linear synchronous motor (PMLSM). This paper researched a tubular flux-switching permanent magnet linear motor (TFSPMLM) as shown in Fig. 1. Its primary iron modules are bonded with the annular PMs to form the primary. And the secondary is one whole tubby iron with slot structure. Predictably, the secondary iron loss is relative big. Based on this, the calculating model was established and the loss measurement of a prototype was conducted. Fig. 2 shows the axial section of the secondary. The tooth and the back iron are tubular as shown in Fig. 3. Fig. 4 shows the radial section of the tubular iron, which has many conductor loops. The current in the inner and outer loop is in the opposite direction"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0002132_jsvi.1996.0655-Figure2-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0002132_jsvi.1996.0655-Figure2-1.png",
+        "caption": "Figure 2. The force and moment equilibrium of a disk.",
+        "texts": [
+            " The problem may be overcome by expressing the governing equations in the rotating co-ordinates instead of the stationary co-ordinates. However, an asymmetry in the boundary condition due to non-axisymmetrical bearings is permitted; i.e., the periodically varying coefficients appear in the governing equations for the bearings expressed in the rotating frame of reference. A non-axisymmetrical and rigid disk, having principal axes m and n at an angle h apart from the principal axes U and V of the shaft is shown in Figure 1. The equilibrium of moment and shear force of a rigid disk are shown in Figure 2. In the rotating frame of reference, the equilibrium equations of moments at the disk point can be expressed by (a list of notation is given in Appendix C) Mr =Ml +(Id \u2212Dd cos 2h)a\u0308\u2212Ddb sin 2h +V(Jd \u22122Id)b +V2(Jd + Id)a\u2212DdV2(b sin 2h+ a cos 2h), Nr =Nl +(Id +Dd cos 2h)b \u2212Dda\u0308 sin 2h \u2212V(Jd \u22122Id)a\u0307+V2(Jd + Id)b\u2212DdV2(\u2212a sin 2h+ b cos 2h), (1) where a= 1U/1Z and b= 1V/1Z are components of the deflected angle about the V- and U-axes, Id =(Id m + Id n )/2 and Dd =(Id m \u2212 Id n )/2, and the equilibrium equations of the shear forces can be expressed by Pr =Pl \u2212md(U \u22122VV \u2212V2U), Qr =Ql \u2212md(V +2VU \u2212V2V), (2) where U and V are components of the lateral displacement along the U- and V-axes"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0003497_j100306a047-Figure1-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0003497_j100306a047-Figure1-1.png",
+        "caption": "Figure 1. Schematic drawing of the reactor assembly: 1, electric lead wire; 2, Teflon tube; 3, top closure; 4, top cover; 5, stainless steel plug (pressure transmitter); 6 , opening with filter; 7, high-pressure shell; 8, glass cell; 9, stirrer; 10, outer magnet; 11, motor; 12, bottom closure; 13, bottom cover; 14, O-ring seals; 15, reaction mixture.",
+        "texts": [
+            " conditions at about 3 \"C. Repeated preliminary experiments showed that the enzyme thus stored maintained constant activity, within 4~2% deviation on average. Sucrose was mainly used as a substrate, though D-glucose and D-fructose were also used in order to study inhibition effects. All these materials were commercially obtained (Guaranteed grade reagent) and used without further purification. Apparatus and Procedures. A schematic drawing of the reactor assembly used in the present study is given in Figure 1. The assembly consists of an outer high-pressure shell and an inner glass cell, A movable stainless steel plug with two O-rings is inserted into the glass cell and serves as a pressure transmitter: two O-rings always hold the plug in a right position, and the lower O-ring seal prevents the reacting solution in the cell from mixing with the pressurizing fluid (2-propanol) surrounding the cell. A Teflon tube with an i.d. of 0.5 mm is passed through the center of the plug. One end of the Teflon tube opens in the cell, and the opening is covered with a filter which separates the catalyst powder from the sampling solution"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0002737_cm.970100304-Figure1-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0002737_cm.970100304-Figure1-1.png",
+        "caption": "Fig. 1. Schematic diagram of the pH jump method. A thin layer of reactivation buffer (R) was made in a hole of ring-shaped Millipore filter (M) stuck to the lower surface of a coverslip (C). The coverslip was supported by a U-shaped silicon rubber (SR) making a narrow space between a slide glass (SG) and itself. A rectangular piece of filter paper (F) inserted in the space was soaked with a small amount of acetic acid solution (A) abruptly applied through a solenoid-driven injector (I). Reactivated sperm contained in the buffer and trapped by the air-water interface experienced a rapid change in pH when acetic acid molecules entered through the interface. A part has been schematically severed from the experimental chamber to show the interior. Not in scale.",
+        "texts": [
+            " To test tubule sliding in this range of pH, demembranated sperm prepared under the same conditions were first digested by incubation with added trypsin at a concentration of 2 pg/ml for 2 min at room temperature (with termination effected with 20 pg/ml soybean trypsin inhibitor) and then reactivated with 1 mM ATP. pH Jump Method Reactivated sperm were maintained in a thin layer of the reactivation buffer spread flat in a hole of a ring-shaped piece of Millipore filter (HAWP-013; pore size, 0.45 pm), which was stuck to the lower surface of a supported coverslip (Fig. 1). Buffer pH was lowered rapidly by dissolving acetic acid vapor in the buffer. Vapor was generated by vaporization from underlying filter paper abruptly soaked with 10-50 p1 acetic acid applied through an injector. The injector was driven with a solenoid, and the moment of injection was recorded by sensing the movement of the plunger. Sperm swimming steadily, close to the air-water interface of the buffer, which experienced an abrupt drop in pH as acetic acid vapor molecules entered through the interface, were observed"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0000699_02533839.2016.1215936-Figure2-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0000699_02533839.2016.1215936-Figure2-1.png",
+        "caption": "Figure 2.\u00a0forces on an angular contact ball bearing.",
+        "texts": [
+            " R is positive if convex (1)B \u2212 A B + A = F( ) = 2 ( 1 \u2212 e2 ) e2 E(e) \u2212 K (e) E(e) + 1, Figure 1.\u00a0the curvature radius of the inner ring and outer ring. velocity about its own center and shaft axis. All of these parameters can be calculated as shown in Cao and Altintas (2004). The contact angle of the inner and outer rings changes due to the effects of centripetal force Fck and gyroscopic moment Mgk. Moreover, the inner ring and outer ring undergo relative displacement, but the axial force on the outer ring of bearing is constant, as shown in Figure 2. By decomposing Fck, Mgk, and the axial preload force Fa in their axial directions, the resultant is constant, the equilibrium relationship of equations (10) to (12) are obtained: (10)Fck + Mgk Dw sin ok + Qik cos ik \u2212 Qok cos ok \u2212 Mgk Dw sin ik = 0, (11)Qik sin ik + Mgk Dw cos ik \u2212 Qok sin ok \u2212 Mgk Dw cos ok = 0, (12)Qok = Fa z \u00d7 sin ok , where \u03b8ik is the contact angle of the inner ring; \u03b8ok is the contact angle of the outer ring; Qik denotes the contact force on the inner ring; Qok is the contact force on the outer ring; Fa is the axial preload force; and z is the number of rolling elements in the bearing"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0000912_epepemc.2016.7752167-Figure1-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0000912_epepemc.2016.7752167-Figure1-1.png",
+        "caption": "Fig. 1. Vehicle model axis and constants",
+        "texts": [
+            " The main disadvantage comes from the higher data processing needs and added loops to control the motors [11]. These disadvantages may reduce the reliability as compared to the mechanical differential. This model was created to study the effect of each wheel torque when accelerating and turning of the vehicle. For this, the base model is a 3 Degree of Freedom model which includes longitudinal acceleration, lateral acceleration and yaw rate. The yaw rate is the rate (rad/s) that the vehicle is rotating at its centre of gravity (CoG). As seen in Fig. 1 the model will include a longitudinal velocity vx (m/s), lateral velocity vy (m/s), vehicle velocity v (m/s), slip angle (rad), yaw rate (rad/s), steering angle on the front wheels . Although the front wheels do not turn with the same angle , since the modelled vehicle has rear wheel traction, this difference is therefore despised in this model. Finally, the vehicle dimensions are the total wheelbase lf + lr and total track wl + wr. To include the forces of each tyre in the base model, a tyre model was also created, including a weight transfer model for the vehicle to calculate the normal forces at each tyre from the accelerations in the base model"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0000524_978-4-431-55013-6-Figure3.2-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0000524_978-4-431-55013-6-Figure3.2-1.png",
+        "caption": "Fig. 3.2 Chaos avoidance in the case of the H\u00e9non map: a distribution of the parameters after 100 updates; b the local expansion rate in the same parameter region",
+        "texts": [
+            " Note that minimization of the local expansion rate \u0393 (N , z0, \u03bb) is equivalent to that of \u03c1(DN (z0, \u03bb)), for which the technique of a matrix inequality can be used. In particular, we consider the following optimization problem: minimize \u03b1 in (\u03b1, \u03bb, X) subject to \u03bb \u2208 \u039b, ( \u03b1X DN (z0, \u03bb)T X X DN (z0, \u03bb) X ) O. We try the proposed method with the system based on the H\u00e9non map (3.2). In the (a, b) parameter space, a grid equally spaced by 0.001 is taken. Each grid point is chosen as an initial parameter \u03bb0 and is updated according to the method given in the previous section. Figure3.2a shows the result together with bifurcation parameter sets. The symbols G p and I p denote tangent and period-doubling bifurcations for p-periodic points, respectively. Chaotic behavior is observed in the shaded parameter region. The parameters \u03bb100 obtained after 100 updates are indicated by the small dots. Four typical trajectories are presented by the solid lines with arrows. The ends of each line correspond to the initial parameter \u03bb0 and the final parameter \u03bb100. The updates are made in the direction of the arrows. Figure3.2b shows an overlapped image of the local expansion rate for attractors and the bifurcation diagram of periodic points. The colored contour plot presents the values of the local expansion rate, as indicated by the color bar. Cold color (blue) expresses a small local expansion rate and then high stability. We see that the small dots in Fig. 3.2a are mainly distributed in the region with a negative local expansion rate in Fig. 3.2b. Thus our method operates the system to avoid chaos. In thisChap.3, robustificationof a dynamical system is considered along the approach in the previous chapter and a technique with a matrix inequality is introduced. Although the stability index considered here is not differentiable in general, its optimization can be formulated into a differentiable optimization problem. Since the resulting optimization problem has a nonlinear matrix inequality constraint, we use the penalty function method of Koc\u030cvara\u2013Stingl to obtain an update rule of the parameter"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0003560_nme.1620180606-Figure4-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0003560_nme.1620180606-Figure4-1.png",
+        "caption": "Figure 4. Variations of function A",
+        "texts": [
+            " The shape functions for this type of element are well known and are listed elsewhere.\u2019 Of particular interest are the shape functions for the nodes on the upper boundary of this element, nodes 6-8: N6 = 0 . 2 3 1 -c ) ( 1 + a)(-& + 7 - 1) ( l a ) N7=0*25(1+[)(1+q)(t+q -1) (1b) N8 = O q 1 -[\u2019)(1 + 7) ( 1 C ) Consider a function where hi is the value of A at the node i. The variation of the function A(& 77) given by equation (2) may be interpreted in the following way: A([, 77) first defines a quadratic curve along the upper boundary (Figure 4a); for each point 6 along the upper boundary, the function determines the value of the curve at the point, i.e. A ( & 1) in Figure 4(a); then, the function defines a second quadratic curve in the 7 direction (Figure 4b) based on the value A (6, 1). In order to obtain a better understanding of the two curves, it is convenient to rewrite (3) equation ( 2 ) in the form A([, 77) = NkA6-t N;A,+ NnA; where Equations l(c), 5(a) and 5(b) indicate that the function may be separated into two parts: (1) the contribution of the corner nodes which varies linearly along the 6 direction and quadratically along the 77 direction; and (2) the contribution of the mid-side node which is quadratic in the [ direction and linear along 77. These shape functions are shown in Figures 3(b) and 3(c). If, instead of just three nodel values, the quadratic curve of Figure 4(a) itself is specified along the upper boundary, the variation throughout the element can be taken in the same way as the function A(& 7) by dividing the curve into two parts and then assigning linear and quadratic curves in the 7 direction for the different parts (Figures 4a and 4b). Based on the preceding observations, a set of new shape functions for the transitional boundary may now be developed. In order to accommodate any arbitrary curve a(,$) (Figure 4c) specified by the adjoining element, a node which moves all along the upper boundary is defined (Figure 3d). At every point [\u2019 occupied by this moving node, the nodal value will be equal to the value of the specified curve at the point, a(,$\u2019). Since this node moves from one end of the boundary to the other, any arbitrary curve may be matched. Next a set of shape functions which define a curve in the 77 direction based on the nodal value of the moving node must be generated. Following the methodology used to construct Figure 4(b), the nodal value a(,$\u2019) of the moving node is divided into two parts (Figure 4 ~ ) ; the part which is linear along the upper boundary generates a quadratic curve in the 7 direction and the remaining part produces a linear curve in the 7 direction (Figure 4d). For this division of the nodal value of the moving node, the values of the specified curve at the two corners of the transitional boundary must be utilized. Taking h6 and h7 as n(-l) and f l (+ l ) , respectively (Figure 4c), the new shape functions and the nodel values are defined by (6) N Q = 0*50(1+ 7) (7) h(,$,v) = Nkh6 + N:h7 + Nkh s\u201d (8) Equation (8) varies in a similar way as equation (3), along with accommodating the function n on the upper boundary of the element. Therefore, if and equation (3) are compatible functions, so is equation (8). When the function A(,$, 7) is used to describe displacements, the preceding arguments can be utilized to assert that rigid body motion is properly represented. In a typical finite element formulation, all nodal degrees-of-freedom (e.g. A,) are independent. In equation (8), however, h6, h7 and h ;I are not really independent but depend on the function and are simply convenient tools for the division indicated in Figure 4(c). Thus the nodes 6 ,7 and 8 are not independent nodal points but sub-nodes within a line node. In the remainder of this paper, the term \u2018point node\u2019 indicates an independent node (i.e. nodes 1-5 in Figure 3d) while the term \u2018sub-node\u2019 means a node within a line node (i.e. nodes 6-8 in Figure 3d). Similarly, the shape functions of the sub-nodes should be considered as parts of the shape function for the moving node. Equations (5)-(8), as they have been stated, may be used to specify displacements as well as geometry"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0002252_s0890-6955(98)00006-6-Figure9-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0002252_s0890-6955(98)00006-6-Figure9-1.png",
+        "caption": "Fig. 9. The modified feed fCM, when curve EP is intersected by the curve CM.",
+        "texts": [
+            " The real cutting conditions are shown in Figs 8 and 9. For simplification of the calculations, the simplified model is shown in Fig. 10, in which the actual feed in Figs 8 and 9 have been modified. The length of WW9 can be found from Fig. 8 as: WW9 = ll/sin(uPC + Ce 2 Cs) (17) ll = (l1 + l2 + l3 + l4)sin(Ce 2 Cs) + (h1 + h2)cos(Ce 2 Cs) + R2cos2uR2 2 R3cos2uR3\u00b7cos(Ce 2 Cs) 2 fsinCe 2 R1\u00b7[1 + sin(Ce 2 Cs) 2 2uR1] (18) The modified feed fCM is calculated as: f CM = fcosCs + R1(1 2 cos2uR1) + WW9cosuPC (19) If WW9 , 0 i.e. the condition as shown in Fig. 9, the intersection angle uSS can be calculated as: uSS = Ce 2 Cs + sin21(mm/R1) (20) where mm = (l1 + l2 + l3 + l4)sin(Ce 2 Cs) + (h1 + h2)cos(Ce 2 Cs) + (R2cos2uR2 2 R3cos2uR3)\u00b7cos(Ce 2 Cs) 2 fsinCe 2 R1sin(Ce 2 Cs) (21) Thus yielding modified feed as: f CM = fcosCs + R1(1 2 cos2uss) + WW9cosuPC (22) After the modified feed fCM is obtained, the shear plane area and the projected area on the tool face can be calculated from Figs 4 and 10, as: A = A1 + A2 + A3 + A4 + A5 + As (23) where A1 = 0.25[4a2 1n2 1 2 (a2 1 + n2 1 2 c2 1)2]1/2; (24) A2 = 1/coshc9\u00b7 E 2uR2 0 \u00b7[fCM 2 (l1 + l2) + l3 2 R2sin(2uR2 2 F)]ds + E 2uR3 0 \u00b7[fCM 2 (l1 + l2 + l3) + R2sin(F)]ds; (25) A3 = PC/(2coshc9)\u00b7[2(fCM 2 l1) 2 l2]sin(hc9 + uPC)const1; (26) A4 = E 2uR1 0 [(fCM 2 R1 + R1cosF)/coshc9]ds; (27) A5 = 1/2\u00b7(k1 + i1)j1\u00b7const1; (28) As = 0"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0003566_an9861100671-Figure2-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0003566_an9861100671-Figure2-1.png",
+        "caption": "Fig. 2. Formation of FeI* - 1,lO-phenanthroline complex as a function of tert-butyl hydroperoxide level. Reaction time: 1 , 15 min; and 2, 3 h. 0, Pipette; 0, microsyringe. Broken line, theoretical stoicheiometry 2 : 1",
+        "texts": [
+            "645 0.922 2.058 3.19 5 0.806 1.060 2.366 2.94 Av. 3.19 k 0.16 Fe\"P complex and that of the hydroperoxides actually present. The need to prolong the reaction time to 2 h was also noted by Amin et ~ 1 . ~ 2 6 who applied this method to the determination of hydroperoxides in low-density polyethylene (LDPE). They attributed this phenomenon to an excessively slow - . penetration of the reagent into the films. OG results obtained with model hydroperoxides in benzene solutions are given in Table 1 and Fig. 2 for tert-butyl hydroperoxide and in Table 2 for cumyl hydroperoxide. It follows from Fig. 2 that at higher hydroperoxide concentrations (exceeding 5 x 10-7 mol per 25 ml) a considerable deviation from a straight-line dependence occurs when a reaction time of only 15 min is allowed. On the other hand, no differences in the results is noted at lower hydroperoxide levels whether the reaction was carried out for 15 min or 3 h. This clearly indicates that under the given conditions the effect of oxygen on the results is negligible and that the curvature of line 1 is due to an insufficient time period being allowed for the reaction",
+            " For example, in a 25-ml flask the total amount of Fez+ available for the ROOH reduction is 2.5 x mol, which, in the most unfavourable instance, Le., for 1.0 X 10-6 mol of ROOH to be determined, represents an excess by a factor of only 1.25. This insufficient excess of Fez+ is believed to be responsible for the abovedescribed phenomena (curvature, long reaction time requirements). It can be concluded, however, that in spite of the above-mentioned disadvantages, the 1 ,lo-phenanthroline method is sufficiently reproducible and precise. This follows from Table 1 and Fig. 2, which show the results obtained using different methods of sampling (direct or with a syringe). Unfortunately the working procedure is relatively complicated and time consuming. As the method is based on the measurement of FeII consumption, it is necessary to know precisely the amount of the reagent added initially. The major disadvantage, however, is the fact that the strongest colouration is always observed at zero concentration of ROOH. Therefore, when very low concentrations of hydroperoxides are determined, one has to compare two relatively high absorbances, which obviously reduces the precision of the determination"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0000991_tmag.2016.2645541-Figure1-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0000991_tmag.2016.2645541-Figure1-1.png",
+        "caption": "Fig. 1. Concept of the reference frame method and zero error condition.",
+        "texts": [
+            " Unlike the previously mentioned methods, the effect of the load is not ignored in estimating the error angle. But during the implementation of the transformation matrix, the load terms are cancelled out automatically and ultimately proposed method gives better estimation for the IPP, compared with the previously mentioned works, both under no load and loaded condition. The validity of the proposed method is verified with experiment. A. Principle of two reference frame method in -45\u00ba~45\u00ba frame The principle of the IPP estimation using reference frame method was proposed in [16], as shown in Fig. 1. According to the principle of IPP, there are two d-q reference frames; namely, motor actual reference frames, dmotor and qmotor and control side reference frames, dcontrol and qcontrol. The actual P 0018-9464 (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. pole position is along with dmotor. Reference commands are given along the dcontrol-qcontrol axes"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0003858_841296-Figure1-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0003858_841296-Figure1-1.png",
+        "caption": "Fig 1 Measuring device",
+        "texts": [
+            "2 Experimental setup used In the study described below, friction was mea sured on the liner of a production engine, with minimum transformation work. The actual operational conditions are therfore conserved. The use of production parts facilitates study of the effect of geometrical differences. The spark ignited engine used was obtained from the Renault - 2 litre range. - Engine characteristics : bore : 88 mm stroke : 82 mm compression ratio : 9.5 : 1 connecting rod length : 137 mm - Piston characteristics ; pin offset : 1 mm 2 compression rings 1 scraper ring The test system on which this engine was equipped is shown in figure 1. The liner under test was isolated from the cylinder head and crankcase. Eight horizontal blades and three vertical blades were attached to the band arranged around the liner. The hori zontal blades take up radial forces on the liner due to the piston. The vertical blades, located at 120\u00b0, transmit the friction force to three piezo-electric sensors. This blade arrangement provides high sen sitivity in measuring friction. The three verti cal blades enable mechanical decoupling of pis ton radial forces"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0000804_jahs.61.042006-Figure1-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0000804_jahs.61.042006-Figure1-1.png",
+        "caption": "Fig. 1. Two examples of helicopter transmission with face-gear drives: (a) DS-21 demonstrator gearbox (Ref. 4) and (b) Apache Block III (Ref. 5).",
+        "texts": [
+            " 6) developed the analytical geometry of face-gear drives and numerically simulated the meshing procedure via computer programs. Optimized tooth contact and bending stress analyses have been conducted using finite element analysis tools in Refs. 7 and 8. Also, experimental investigations of face-gear tooth contact pattern, failure modes, and load capacities were conducted in Ref. 9, and fatigue characteristics were explored in Ref. 10. Finally, several lightweight design examples of face-gear drives have been presented in the literature (Refs. 4, 5, 7), as shown in Fig. 1. The reduced sensitivity of face-gears to pinion misalignments translates directly into weight benefits by reducing bearing support stiffness requirements. Litvin et al. (Ref. 6) proposed the meshing of a face-gear with a spur involute pinion, which has smaller number of teeth than the shaper, in point contact so as to tolerate misalignments. However, after comparing contact ratio, maximum load, contact pressure, and transmission error, Guingand et al. (Ref. 11) concluded that the face-gear drive designed for line contact gives better performances than the point contact design"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0002056_0022-0728(94)03820-s-Figure3-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0002056_0022-0728(94)03820-s-Figure3-1.png",
+        "caption": "Fig. 3. Multicyclic voltammetric curves of 2 mM Fe(II) in 0.75 M Na2S20 ). The sequence of cycles is indicated by numbers. The second negative sweep was stopped at -0.8 V and electrode was held at this potential for 5 min. HMDE, l' = 2 V min i",
+        "texts": [
+            " With a solution of high $2 O2- ion concentration, formation of a high cathodic peak C 3 is observed on the cyclic voltammetric curves. The peak A 3 is the anodic counterpart of the C 3 peak (Fig. 2). The following experiment was carried out to investigate the nature of these peaks. Cyclic voltammetric curves were recorded in a solution of 2 mM Fe(II) and 0.75 M Na2S20 3. In the second negative sweep the electrode was kept at - 0 . 8 V for 5 min. Next, the potential was changed in the positive direction. The broad A 3 peak was observed in the anodic curve and a high C 3 peak was recorded in the subsequent negative sweep (Fig. 3). New anodic peaks appear in the voltammogram if the electrode is kept at a more negative potential than - 1 . 4 V after cyclic polarization within the - 0 . 4 to - 1 . 2 V potential range. In addition, the C 2 peak height decreases and a new reduction peak is formed at - 1 . 0 V potential in the subsequent negative sweep (Fig. 4). The anodic curves recorded from S 2- ion solutions have similar shapes. The results of this experiment indicate that free S 2- ions are formed in the electrode vicinity during the H M D E polarization to more negative potentials than - 1 "
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0003773_robot.1997.619353-Figure2-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0003773_robot.1997.619353-Figure2-1.png",
+        "caption": "Fig. 2 . Wheel-ground contact and slip directions.",
+        "texts": [],
+        "surrounding_texts": [
+            "sis 1 that we consider. The numerical simulation results, therefore, exhibit the presence of wheel slip, a conclusion we prove based entirely on an analylical theory. Our primary results are theorems 6 and 8.\n11. Kinematics\nThis section introduces the kinematic constraints imposed by the nature and configuration of various types of wheels of the mobile robot. The following sub-sections consider the form of the specific instances of the kinematic constraints of a wheel type of a mobile robot. The kinematic model we derive is based on a model of a zero width non-deformable planar circle rolling with n o slip on the ground. The subsequent analysis and results of the paper, however, are not restricted by this intermediate step in deriving the model.\nA. Kinemat ics Model of Mot ion of the Base of Wheeled Mo-\nFirst consider S , the plane of motion of the base of the wheeled mobile robot. Let F be a choice of a coordinate system in the plane so that F : S 4 2': p 5 (F1(p). F z ( p ) ) . Let the configuration of the base of the mobile robot, an element of S E ( 2 ) , be denoted X I = [z, y,O)' in the choice of coordinate system F. Let the velocity of the base of the mobile robot in the plane of motion a t the configuration X I be denoted X I = (i, y, e ) . It is easy to verify that t2he velocity of the base of the mobile robot X;\" ip the moving reference frame M is related to X I by x;\" = Rxl, where R = R(x1) is a homogenized orthogonal transformation matrix of threeby-three of the form\nbile Robot\n1 [ o 0 1 Cos(0) S in (O) O R ( 0 ) = -Sin(O) Cos(8) 0\nB. Kinemat i c Constraints Imposed b y Wheels\nOur model of a wheeled mobile robot is a generalized model of such robots considered by Campion et. al. [CBD 931. A wheeled mobile robot has either conventional type wheel or an omnidirectional type wheel. A conventional type wheel has a given axis about which the wheel can rotate and is driven. I t is of the following three categories: (i) fixed, [ii) centered orientable, and (iii) off-centered orientable. A n omnidirectional wheel can rotate about an arbitrary axis of rotation in the plane of motion of the base of the mobile robot and is (usually) driven about one given axis in that plane. The configuration of a mobile robot with an arhitrary combination of wheels is described by the following: X I . the three coordinates of the base. x2. the vector of angular orientations of the plane containing the off-centered wheel, x3 = (41, dc. doc. d o d ) , the angular orientations of the fixed, centered, off-centered, and omnidirectional wheels, respectively. about their driven directions, xq = ( 6 0 d ) . an appropriate choice of angular velocities of the omnidirectional wheels about directions complementary to the directions of\n' I t e m s in ( ) d e n o t e , depending on t h e contex t , e l e m e n t s of a Yector or If f R\" -t 72\" is s m o o t h m a p functional dependency of a m a p o n variables s u c h r h a t f = (f',fZ, , f m ) , t h e n B \" i s J ' \" 2 ' ' ' = n' I Dlf' = O f [ t ] [ J ] . 8 2 1 and O f 3 2 is t h e Jacobian of t h e func t ion f\ntheir drive, and x5, the orientations of the plane containing centered wheel.\nIf the number of fixed, centered, off-centered, and omnidirectional wheels are N,, N,, No,, and Nod, respectively, then the dimension of an element xC = (XI , x2, x?, x4, x,) describing the configuration of the mobile robot is 3 + N,, + ( N j + Consider that the origin of the choice of coordinate system OeJ, e E {f, c , oc, od}, j E { 1, . . . , N e } , the choice of coordinate frame of the j t h wheel of type e with the origin on the wheel axle above the center point of wheel-ground contact for fixed, centered, and omnidirectional wheels and at the pivot of the arm of off-centered wheels is given by (1, Q j in polar coordinates in the choice of coordinate system M and the radial line cy is the x-axis of O,,(see Fig. 1). Similarly, the origin of the choice of slip coordznatefmmes M e , (or t,he origin of M e , j is (cl, p - i) in the choice of coordinate system U,, and the radial line /r - is the x-axis of the frame Me,. Let ;( be the angle that the direction of complementary rolling 4 of an omnidirectional wheel makes with the direction of &, the axis about which the wheel is driven. The three scalar components of the constraints imposed by the wheels are as follows: Nc + No, + N o d ) + Nod + Nc = 3 + Nj + 2 ( N c + N o c + N o d ) .\nCOS(^) sin(dj I s l n ( d - c r ) + d c o s ( u + / 3 t - i - 6 ) J R x 1\n-r sin(. t p - b j & + dj c o s ( - + p - 6) - r ' s i n ( a + p + -/ - 614,\n[ - s i n ( 6 ) cos(6) I c o s ( 6 - o j + d s i n ( u t ; j + 1 - 6 j ] i t x l (1 )\n+dB sin(. + 0 - 6 ) + r cos(. + ,3 - 6 j @\n+ r ' d c o s ( u + /3 + -, - 6 j ,\n[ b (J l]Rxl+,$\nwhere r is the radius of the wheel about the driven direction 4, and r' is the radius of the omnidirectional wheel about the complementary direction & and (5 is a quantity determined by equating\n[ - s i n ( b j c o s ( & ) L c ~ s i 6 - r _ v j + d s i n f c r + R + - , - b , ] F t x i\n+d / J s in i a + B ~ 6 ,\nto zero in this instantiation so that the y-velocity in the slip co-ordinate frame ,UaJ is zero. For fixed wheels d = 0 and 8 is a constant. for centered wheels d = 0. For centered and offcentered wheels 3 is a state variable, a component of x: and x2 respectively. For fixed, centered, and off-centered wheels. the component containing 6 does not appear and -{ = 0. For omnidirectional wheels .9 is a constant.",
+            "The three scalar constraints in Eq. (1) for each wheel restrict the motion of the base of the mobile robot at the center point of wheel-ground contact in the 2, y, and 6' directions of the slip coordinate system Me,. For convenience, the slip coordinate f rame 2-direction will also be called the longi tudinal direction, the y-direction as the lateral direction, and the 8-direction as the rotutional direction. In this terminology, the scalar kinematic constraints for each wheel are also called longitudinal, lateral and rotational constraints due to the rjth-wheel. Let the longitudinal, lateral, and rotational constraints for all the wheels be collected in the form J , X c , J , X c , and J Q X , , respectively, where\nJ o l t 0 0 [ 0 0 0 01 0 0 Jg = J e i c O 0 [ o O 0 0 1 0 J e s c 0 , [ J o i o d o 0 [ o 0 0 01 0 0 J e l o c O J e z o C 0 [ O 0 0 01 0 0 ] (3)\nfor p E { x , y} and constraints in the third column due to the angular velocity of the wheels are further expanded into four subcomponents corresponding to the fixed, centered, off-centered, and omnidirectional wheels. The functional dependency of the terms in the jacobian are: J , I ~ ( x I , X I ) , J , 3 f ( x l , X 1 ) , J r l c ( x l , x l ) ,\nJ . c 3 c ( X l r X 5 r X l ) , J s l o c ( X { 1 , 2 } i X { 1 , 2 } ) r J , z ~ ~ ( x { I , z } , X { ~ , ~ } ) , J s 3 o c ( x { i , z ) , X I ) ) J s l o d ( ~ i , X I ) , J r ~ o d ( x 1 , X I ) , J z 4 0 d ( X l r X I ) , J y l f ( ~ l , X ~ ) , J ~ ~ ~ ( x I , X I ) , J y 1 c ( X 1 , x l ) , J y 3 c ( x 1 , x 5 , x l ) ,\nJ y l n c ( X { l , ? ) , X { I , ~ } ) , Jy?oe (x{1 ,2} i X I ) , . J y 3 o c ( x { 1 , 2 ) r X{I,Z})>\nJ y l o d ( x 1 , X I ) , JyBod(X11 X I ) , J y 4 o d ( X 1 , X I ) . The constraints in the longi tudinal direction impose rolling with n o slip condition on the wheels of the mobile robot. The constraints in the lateral and rotational directions impose nolateral and no-rotational slips, respectively. There are ( N f + N, -+ No, + Nod) constraints of the longitudinal, lateral, and rotational type. The functional dependency of each of the terms in the Jacobians J,, J,, and JR as indicated above are based on a model of the wheels such that they are valid about a small neighborhood of any state. Although, the jacobians are written in a form that is linear in the velocity xc, they include velocity dependent terms arising from the choice of slip coordinate frames M e , with zero y-direction velocity to facilitate traction force transformation in Eq. (9).\n111. Rolling contact of two elastic bodies\nThe theory of frictional rolling of two bodies addresses the problem of determining the traction force at the wheelground contact. A large fraction of this literature is dedicated to the rolling of tires founded in the empirical models of tire mechanics. We, however, limit our study to linearly elastic wheels. In particular, the two bodies in rolling contact are assumed to follow our Hypothesis 1.\nThe analytical theory of frictional rolling of two linearly elastic bodies associates a definite slip called creep associated with the traction forces in the area of contact. We show a certain new symmetry in the creep-force relation. The remainder of this section reviews other symmetries with the elastic quasi-identitfly assumption of Hypothesis 1 given by\nKalker [Kalker 671. Thee,e relations, in effect, allow us to infer the traction forces a t zero slip velocity. Though the new symmetry we show does not require the quasi-identity assumptions, we also need another symmetry that is valid only with the quasi-identii;y assumption. Therefore, in general, our conclusions on wheel slip remain valid only with the quasi-identity assumption.\nA . Creep-Force Re la t ion Problem Defini t ion\nConsider a linearly elastic circular body, denoted e j , rolling on a planar linearly elastic material. Let the velocity of the center of the wheel a.xle, XyeJ, be in the 2-direction of the slip coordinate frame, Me,, defined in Sect. 11-B. Let J { , , , , R p , refer to the j t h row in the e t h type row block of the Jacobians J,, J,, or J Q defined in Eqs. (2) and (3), respectively. The terms J{r,y,e}e3Xc represent the rigid slip of the wheel at the wheel gr'3und interface in the 2, y, and 6' directions, respectively, of the slip coordinate f rames . Define U,,, , the longi tudinal creepage, uye, the lateral creepage, and uoe3, the sp in for the wheels as\n(4)\nwhere V,, = (xceJI is the magnitude of the 2-direction velocity of the point on the axle of the wheel (recall that by the choice of the frame Me, , the y-direction velocity is zero). The creepage and spin are ratios of the rigid wheel slip to the magnitude of the translational velocity of the axle of a wheel. Let the area of contact of the wheel with the ground be denoted Ce,? described in the respective slip coordinate frames. The material in the two bodies in the area of contact deform elas'tically due to the friction-induced tangential traction and the vertical load-induced compression. Let the slip coordinate frames be the choice of the coordinate system to describe the contact area Ce3. Let us add z-axis to the slip frames so that the z-positive direction points into the material oj',the wheel. In this description, the two bodies are approximated as half-spaces with the material on z 2 0 and z 5 0 of the slip coordinate frames. The elastic deformation on these half-spaces due to concentrated normal load in the z-direction and tangential load along the z-axis and the yaxis have been given by Boussinesq (1885) and Cerruti (1882) [Love 441. Let the elastic strain denoted u=,(z, Y, v,,:,, u y e j r ~ 0 . 3 ) = ( ~ z e 3 , ~ y e j , uze3) he the difference in the elastic strains of the rolling wheel and the ground expressed in the slip coordinate frame. Let (X,,, Y,,, Ze3)(z, Y ~ ~ , ~ ~ , ~ ~ ~ ~ ~ u R ~ ~ ) be the 5 , U and 2 cornponents of the traction and the vertical load at a point",
+            "(x, y) E C,, . The relation between the material strain function U and the traction ( X , I: 2) is\nwhere\nA K+\nR = &A2 + A;), A, = x - X I , Ay = y - y*, the modulus Eiw'S1 1 - L ( i - + L) ~ of rigidity G{w3g1 = Z ( ~ + ~ { W , ! ? > ) ' 3 - 2 Gw GS ' G z ( p 1 all + $), 9 and K = :(- - w), with E as the Young's Modulus of Elasticity, D as the Poisson's ratio and superscript 20 and g standing for the properties of the wheel and the ground, respectively.\nDue to the elastic material flow with respect to a reference frame moving with the wheel, the net relative displacement of one body with respect to the other a t a point in the area of contact is the sum of the gross rigid motion component and the relative elastic motion. The net relative velocity of one body with respect to the other a t a point (z,y) E C,, denoted W(x, y , vre j , u y e j , ~ 0 e j ) = ( w z e j , wye,) is\nwhere steady-state assumption on the flow of material is assumed, i.e., 2 = 0 . Then, according to the law of Coulomb- Amontons,\nA Q e j ( v z e j , V y e j r V O e j ) = - X ~ e 3 ( ~ ~ e j , - v ~ e 3 , - - ~ g e g j , verify the creep-force law when the contact area Ck, (2, y) = C,, ( 2 , -y). Proof.[Sh 961.\nCorollary 3: With no lateral and angular slip, the lateral and angular traction disappear when the contact area c e j ( 2 , y ) = C e j ( % , - y ) , i.e., Ayej(~re3r0,0) =\nUnfortunately, the creep-force problem as posed obeys no other symmetry to exhibit any similar conclusion for the longitudinal traction A,. Kalker [Kalker 791, however, considered the following cases of quasi-identity: either the two elastic bodies are elastically similar, i.e., E g = E\", and ug = U\", or both are incompressible, i.e., ng = cW = 0.5! then K in Eq. (5) is zero. Also, when one body, say a rubber wheel, is incompressible, and the other body is relatively rigid, i.e., uw = 0.5, and E\" << Eg, K is close to zero. When K is zero, the elastic strain and traction relations of Eq. ( 5 ) simplify in such a way that the problem of determining the vertical strain U z , and therefore the vertical pressure distribution Z ( x , y) , separates from that of the tangential problem of determining u z , uY and the corresponding X ( z , y) and Y ( z , y) . Let us call the elastic-traction relation derived from those of Eqs. ( 5 ) with K = 0 the quasi-identical elastic-traction law and the corresponding creep-force law problem posed in Eq. (8) as the quasi-identical creep-force law. This separation of vertical and the tangential problem in quasi-identity allows several other symmetries in the creep-force law [Kalker 671 including a specialization of the Proposition 2 we proved earlier. We mention one other:\nProposi t ion 4 (Kalker, 1967) The traction symmetry relations A z e j ( ~ m e 3 , vyejr v e e j ) = - A r e j ( - v z e j , v y e j , ~ e e j ) , Aye3 (-vze3, v y e j , ~ e j ) , Aye3 (Vzej 1 V y e j , V o e j ) A o e j ( ~ z e 3 , ~ y e 3 , v o e j ) = ~ ~ ~ ~ ( - ~ ~ ~ ~ ~ ~ y ~ ~ ~ v e e j ) , verify the quasi-identity creep-force law when the contact area Ce,(z,yj = c e j ( z ) - ~ ) ,\nCorollary 5 (Kalker, 1967) With no longitudinal slip, the longitudinal traction in quasi-identical problem disappears when the contact area C,,(z,y) = Ce,(x,-y), i.e.,\nA B e j j V m e j , 090) = 0.\n-\nAze j (0 , V y e j , v ~ e j ) 0.\nwhere p e , is the coefficient of friction a t the wheel-ground interface of e j t h wheel. Determine (Arej , A,, A e e j ) , the traction forces at the wheelground interface defined as The creep-force law problem is defined as follows:\n( A x e l , Aye], heeJ) = 1 Le3 (xe,, ye, 3 z y e j - y ~ e j )dxdy, (8)\nso that Eqs. ( 5 ) , (6), and (7) are satisfied when the creepages (Vze,, V e e j ) , the net load Ne3 = J Jc., Ze, ( x , y)dx d y , and the translational velocity V,, are known.\nB. S y m m e t r y i n Creep-Force Relat ion\nThe problem of creep-force law as posed in Eq. (8) admits a symmetry relation that enables us to infer the traction forces on a special subset defined by no-lateral and no-angular slip.\nProposit ion 2: The traction symmetry relations A r e j ( V z e j , v y e j , VQ~,) ( V z e j , - p y e j , - Y Q , ~ ) , A y e l ( V z e j , Vyejr ~ ~ e j ) = -Aye, ( ~ z e j , - v y e j , - ~ e j ) ,\nIV. Equation of motion without and with constraints\nThe dynamic model of mobile robot is obtained by EulerLagrange formulation subject to the external forces applied at the actuated joints, and the forces a t the wheel-ground interface. Let the vector of external (generalized) forces be T = (0, T,,, ~ 4 ~ 0 , T ~ ) , where the three degrees of freedom of the base of the mobile robot XI = (z,y,B) and t,he undriven direction of the omni-wheels are not directly actuated. The forces a t the wheel-ground interface are denoted A,, A, and As, in the directions x, y, and 6' of the slip coordinate frame respectively, of each of the wheels.The A's are vectors with (Nf -t N, + N,, + Nod) component,s. = ir + JZAr i- JTA, + J ? X o , where T is the total energy of the system. Expanded into componentas, the equations of motion look like the following: The generalized equations of motion is 2 -"
+        ]
+    },
+    {
+        "image_filename": "designv11_60_0001781_s0167-8922(08)70009-9-Figure1-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0001781_s0167-8922(08)70009-9-Figure1-1.png",
+        "caption": "Figure 1 - Polar Load Diagram for the Big-End Connecting Rod Bearing",
+        "texts": [
+            " The kinematics of the crank and slider mechanism are straightforwardly calculated and the motion of the shaft can be determined by taking the hydrodynamically generated load capacity due to the pressures in the fluid film to balance the applied load. In this approach a consideration of dynamic instabilities is therefore precluded. It is not proposed to develop here the procedures for calculating the loading and kinematics of specific bearings. The general approach for the crank and slider mechanism is widely available in the literature and the design report of Martin (1974) is particularly useful. A polar load diagram for the big-end connecting rod bearing of a 6 cylinder petrol engine passenger car is shown in Figure 1 whilst that for the main bearing number 1 is given in Figure 2. Attention will primarily be given to the development of the understanding of the lubrication equations appropriate to the situation and how these can be evolved into the Mobility Method design approach. The paper will conclude with an assessment of the limitations of the technique thus highlighting future developments which may be anticipated. A,B constants of integration (Appendix) b bearing width c d Shaft diameter e eccentricity ( e = de/dt) f(c,q) polynomial function (Booker (1 969)) h lubricant film thickness M mobility numberhector (with components in E, y, c, q directions) p pressure maximum instantaneous pressure Pmax - pmax maximum pressure ratio (= pmaJ(P/bd)) P total instantaneous load (with components Pa, Pn) r shaft radius R bush radius t time V, squeeze' velocity x,y,z Cartesian coordinates p /dt) E q lubricant dynamic viscosity c,q $ $1 4 2 limits of full film region y attitude angle (+= dy/dt) radial clearance (=(R-r)=half diametral clearance) angle between fixed axis and line of shaft(S) and bush (B) centres (p = dp eccentricity ratio (= e/c) ( E =dddt) rectangular coordinates in load frame (Figure 4)"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0003013_a:1008115522778-Figure15-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0003013_a:1008115522778-Figure15-1.png",
+        "caption": "Figure 15. Initial configurations of the peg used in the experiments for misalignment compensation.",
+        "texts": [
+            " Also, the experiments to compensate for the inclination error of a part and to insert a part into its mating hole were performed. However, the deep study on them is left as a further study. Figure 14 shows the flow chart of the neural netbased assembly method. The end effector of a robot continues to be moved in the direction of misalignment until misalignment compensation is accomplished. It is checked every step of the corrective motion whether misalignment compensation is completed or not. Figure 15 shows the initial rough shapes of the part used in the experiments for misalignment compensation in cylindrical peg-in-hole tasks. The experiments to compensate for lateral misalignment only were performed in nine cases as shown in Figure 15(a)\u2013(c). And the experiments to compensate for inclination error and to insert a part into its mating hole were performed in one case as shown in Figure 15(d). Figure 16 shows the experimental results for lateral misalignment compensation when em = 1 mm and cr = \u22122, 1, 4 mm. The em denotes lateral misalignment, namely, the distance between a hole center and the center of the bottom of its respective mating part. And the cr denotes the center position of the upper surface of a part. When em is small, the error in estimating a hole center using the implemented visual sensing system becomes large because the visible part of the hole is a little [23]. Figures 16(a) and (b) show the large error between the actual prescribed value and the measured value by the sensing system inferred from the neural network"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0000622_b978-0-12-800782-2.00003-8-Figure3.12-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0000622_b978-0-12-800782-2.00003-8-Figure3.12-1.png",
+        "caption": "Figure 3.12 Field for the determination of the unsaturated rotor leakage flux of a 3.4 MW, two-pole, three-phase induction motor. One flux tube contains a flux per unit length of 0.015 Wb/m [19].",
+        "texts": [],
+        "surrounding_texts": [
+            "Numerical approaches such as the finite-difference and finite-element methods [16] enable engineers to compute no-load and full-load magnetic fields and those associated with short-circuit and starting conditions, as well as fields for the calculation of stator and rotor inductances/reactances. Figures 3.8 and 3.9 represent the no-load fields of four- and six-pole induction machines [17,18]. Figure 3.10a\u2013e illustrates radial forces generated as a function of the rotor position. Such forces cause audible noise and vibrations. The calculation of radial and tangential magnetic forces is discussed in Chapter 4 (Section 4.2.14), where the concept of the \u201cMaxwell stress\u201d is employed. Figures 3.11 to 3.13 represent unsaturated stator and rotor leakage fields and the associated field during starting of a two-pole induction motor. Figures 3.14 and 3.15 represent saturated stator and rotor leakage fields, respectively, and Fig. 3.16 depicts the associated field during starting of a two-pole induction machine. The starting current and starting torque as a function of the terminal voltage are shown in Fig. 3.17 [19]. This plot illustrates how saturation influences the starting of an induction motor. Note that the linear (hand) calculation results in lower starting current and torque than the numerical solution. Any rotating machine design is based on iterations. No closed form solution exists because of the nonlinearities (e.g., iron-core saturation) involved. In Fig. 3.13 the field for the first approximation, where saturation is neglected and a linear B\u2013H characteristic is assumed, permits us to calculate stator and rotor currents for which the starting field can be computed under saturated conditions assuming a nonlinear (B\u2013H) characteristic as depicted in Fig. 3.16. For the reluctivity distribution caused by the saturated short-circuit field the stator (Fig. 3.14) and rotor (Fig. 3.15) leakage reactances can be recomputed, leading to the second approximation as indicated in Fig. 3.16. In practice a few iterations are sufficient to achieve a satisfactory solution for the starting torque as a function of the applied voltage as illustrated in Fig. 3.17. It is well known that during starting saturation occurs only in the stator and rotor teeth and this is the reason why Figs. 3.13 and 3.16 are similar. 220 Power Quality in Power Systems and Electrical Machines F2 = 3500 N/m F1 = 3500 N/m F 2 = 3 11 0 N /m F 1 = 3 15 0 N /m F 2 = 1250 N/m F1 = 4340 N/m F 2 = 1770 N /m F 1 = 5160 N /m F 2 = 1400 N /m F 1 = 4110 N /m F 2 = 1 30 0 N/m F 1 = 5 45 0 N/m 1 \u2013B \u2013B 2 3 A A 4 5 \u2013C \u2013C \u2013A \u2013A C C B B 6 7 8 9 10 11 12 (a) 1 \u2013B \u2013B 2 3 A A 4 5 \u2013C \u2013C \u2013A \u2013A C C B B 6 7 8 9 10 N S 11 12 (b) f f 1 \u2013B \u2013B 2 3 A A 4 5 \u2013C \u2013C \u2013A \u2013A C C B B 6 7 8 9 10 N S 11 12 (c) Figure 3.10 Flux distribution and radial stator core forces at no load and rated voltage for (a) rotor position #1, (b) rotor position #2, (c) rotor position #3, Continued 221Modeling and Analysis of Induction Machines 222 Power Quality in Power Systems and Electrical Machines 223Modeling and Analysis of Induction Machines Figure 3.16 Field distribution (second approximation) during starting with rated voltage of a 3.4 MW, two-pole, three-phase inductionmotor. One flux tube contains a fluxper unit lengthof 0.005 Wb/m [19]. 0.1 0.2 0.3 0.4 0.5 0.6 0.7 0.80 0 1.00.9 2 4 6 8 10 12 0 0.4 0.8 1.2 1.6 2.0 2.4 2.8 3.2 I [kA ] ph START T [kNm ] TOTAL START I HAND I NUMERIC MEASURED CURRENT MEASURED TORQUE V ph [pu ] Vnom . ph TNUMERIC THAND Figure 3.17 Starting currents and torques as a function of terminal voltage for a 3.4 MW, three-phase, induction motor [19]. 225Modeling and Analysis of Induction Machines"
+        ]
+    },
+    {
+        "image_filename": "designv11_60_0002619_1.2829168-Figure2-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0002619_1.2829168-Figure2-1.png",
+        "caption": "Fig. 2{b) Coordinate system relations of grinding wheel right-side sur face and dual-lead worm left-side surface",
+        "texts": [
+            " The position vector and unit normal vector of the grinding wheel surfaces can be represented in coordinate system SdXc, Yc, Z^) as follows: R\u201e, \u2014 and Uw = u cos cti cos 6 u COS a,- sin 9 + (bi \u2014 u sin a,) s m (jj COS sin a, sin ( COS a , (1) (2) X. r^tanar Zo>Z,t where fo, = inmj/4) cos Pj + r\u0302 , tan a;. Subscripts i = r and / where r indicates the right-side grinding wheel surface and / represents the left-side surface. Parameter mj (j = r and /) denotes the axial module of the dual-lead worm surfaces, a, and rgi represent the pressure angle and pitch radius of the grinding wheel surfaces, and u is the grinding wheel surface parameter. Figure 2(a) depicts the coordinate system relations of gener ation mechanism for the dual-lead worm right-side surface and the grinding wheel left-side surface. Coordinate system SdX^, Yc, Zc) is attached to the grinding wheel during the generating process, and axis Z<, is the rotation axis of grinding wheel. Coordinate system SfXXjc, Yfc, Zf^) is considered the fixed coordinate system while coordinate system S\\ (Xi, Yy, Z,) is attached to the dual-lead worm, and Z, denotes the rotation axis of dual-lead worm",
+            " 120 / 415 Downloaded From: http://mechanicaldesign.asmedigitalcollection.asme.org/ on 01/28/2016 Terms of Use: http://www.asme.org/about-asme/terms-of-use where pressed in coordinate system Sc, the following equation must be observed (Litvin 1989, 1994): n ( c ) . ^(Ic) ^ jj(c) . ( Y (1) _ y (c)^) ^ Q (5) where V^'\"' represents the relative velocity between worm and grinding wheel, expressed in coordinate system 5^. Since the grinding wheel is attached to coordinate system SdX^, YcZ^), as shown in Fig. 2(\u00ab) and 2(b), the velocity of grinding wheel with respect to coordinate system 5c is V''^' = 0. According to the geometric relation, the rotation of dual-lead worm can be represented by rotating about axis Z/^ with angular velocity ft>}c' = wik/c and a translation velocity V}c' = PjUJikfc along axis Zfc- These two velocities may be expressed in coordinate system Sc by applying the following vector transformation ma trix equation: = [Lc/Jw}c' = wi and V*\"' = [L./jv;i' = w, 0 sin Pj cos j3j 0 Pj sin Pj Pj cos /3j ( 6 ) ILcfc] 1 0 0 0 cos PJ sin PJ 0 \u2014sin PJ cos PJ and Vc\"* denotes the translation velocity VjP represented in coordinate system Sc",
+            "org/ on 01/28/2016 Terms of Use: http://www.asme.org/about-asme/terms-of-use [Mfc^cA = [Mi/J = 1 0 0 0 1 0 0 0 1 0 0 0 cos (/>! - s i n (t>i 0 0 (fwgi + /\"wj + -^4) 0 - A 5 1 sin </>i 0 0 cos (\u0302 1 0 0 0 1 0 0 0 1 ' and [Af,/\u201e] = - m2iZij sin (l>\\ sin y + mjiir^gi + r\u201ej + A4) cos <pi COS y - Asm2i sin </<i sin yjji + u!i[m2iy\\j sin </>i sin y \u2014 m^iXy cos </)| sin -y + rrhdKgi + r,\u201e; + A4) sin yjk,. (22) Similarly, according to the geometries shown in Fig. 1(a) and 1 {b) and coordinate system relationship shown in Fig. 2(a) and {b), the unit normal vector of the dual-lead hob cutter cos 01 - s i n 4>\\ 0 0 sin <\u0302 i cos y cos 4>i cos 7 sin 7 0 - s i n <\u0302 i sin y - c o s <\u0302 i sin y cos 7 0 ('\"ws/ + r^i -f A4) cos <\u0302 i -{r^si + '\u2022\u2022vj + A4) sin 0 -As 1 The angular velocity of the worm gear, as shown in Fig. 3, can be expressed by d<p2. . 0)2 = \u2014-Kf^ = wjk/H,. at (17) By applying Eq. (16), the angular velocity of the dual-lead worm gear can also be represented in coordinate system 5i {Xi, Ki, Z,) as follows: \u00abP' = cos 4>i sin 4>i cos y -s in </>"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0003787_iros.1996.570634-Figure10-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0003787_iros.1996.570634-Figure10-1.png",
+        "caption": "Figure 10: Position Control (11)",
+        "texts": [
+            ", and the white circle represents the passive joint. The feedforward control caused large error of the link motion from the planned trajectory, and the link rotation at the end of the trajectory (b). The final configuration was [0.349, 0.100, 0.8211. Though 5 and y coordinates of the passive joint were controlled exactly, 19 deviated from the desired configuration. On the other hand, the free link almost halted nem the desired configuration by the feedback control (c). The final configuration was [0.366, 0.099, 0.7811. Fig.10 shows positioning between [0.35, -0.15, 1.5711 and [0.35, 0.2, 2.3561 by another composite trajectory and the feedback control. The final configuration was [0.359, 0.200, 2.3561. The link arrived near the desired Configuration. 6 Conclusions Position control of a SDOF planar manipulator with a passive joint under 2nd-order nonholonomic constraint was studied. A trajectory planning method was proposed for positioning between arbitrary initial configurations and arbitrary desired configurations. The trajectory was composed of siniple translational and rotational trajectory segments, considering ins tion of the center of impact of the free link"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0002215_0094-114x(95)00011-m-Figure5-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0002215_0094-114x(95)00011-m-Figure5-1.png",
+        "caption": "Fig. 5. A n o n - h o l o n o m i \u00a2 system.",
+        "texts": [
+            " With these first and second-order speed ratios, the manipulator can now be controlled to develop the instantaneous second-order output-space trajectory geometry by the coordinating second-order Taylor series in equation (27), i.e. /t = n2 + \u00bdn'22 = -3.312 + \u00bd4.6722. Figure 4 shows the trajectory generation which results from a first- and second-order of the coordinating Taylor series, a direct comparison of first- and second-orders of coordination. The accuracy of the trajectory tracking is significantly enhanced by the second-order of coordination, readily provided by Curvature Theory. 3.3. Example 2: a non-holonomic system A schematic of a cart system with a solid axle and two drive wheels is shown in Fig. 5, with its corresponding Polar Line and the location of the canonical system. In this system, 2 and p are the rotations about the axle of wheels A and B respectively. Using the canonical system, the first and second-order instantaneous invariants are, r r a;. = ~ a~ = ~ a;.;. = 0 a~ = 0 a:.~ = 0 r 2 r 2 b:. = 0 b, = 0 b:.:. = 4k b~ = ~-~ b;., = 0 r r 0~= }k 0~=~-~ Oz~=O 0 . . = 0 0~.=0. With these instantaneous invariants, equations (14) and (26) reduce to, r/ - - y p W k and n' = (PJ)x(n - l)Sr (29) yp - - k 4k a ' which involve geometry observable in the output-space"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0003006_881168-Figure6-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0003006_881168-Figure6-1.png",
+        "caption": "Figure 6 - Magnetic lBvitation from Beneath the Track and Linear Induction Motor (LlN) Al Plate",
+        "texts": [
+            " In order to promote the concept commercially, HSST-03 which carried 50 passengers, was placed in operation for public transportation at the 1985 Tsukuba Exposition and has since carried over 610,000 people. In 1986, at the Vancouver Trans portation Exposition (EXro86) the HSST vehicle carried an additional 460,000 passengers during the 160 day event. In total, over 1,000,000 riders have been carried and public acceptance of the HSST concept appears to be enthusiastic. MAGNEl'IC IEVITATION PRINCIPLE - The HSST levitation principle is based on magnetic attraction from underneath the track. As shown in Figure 6, the suspension bogies fonn a horizontal U-shape and trap the track within the U. The track consists of a linear horizontal prismatic steel beam in the shape of a shallow inverted U. HSST PROFUISION CONCEPT - The HSST vehicle is propelled by a linear induction motor (LlN). However, as shown in Figure 6, the rotor is stationary and it consists of a horizontal aluminum reaction plate mounted on top of and cr:inped to the steel track. Tractive force is induced in the secondary aluminum rail by a phased magnetic field from primary traction coils on the vehicle. Forward or reverse is achieved by controlling the phasing of the primary coils. Power is picked up by contacts with a ground and a hot rail under the track. Figure 7 shows that a closed loop vertical air gap sensor and control system maintain a 1 em air gap shove the track by varying the magnetic flux from beneath"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0000734_cscwd.2016.7565955-Figure2-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0000734_cscwd.2016.7565955-Figure2-1.png",
+        "caption": "Fig. 2.The sketch map of the cross section of strand",
+        "texts": [
+            " At this point, the geometry of wire rope of strand j layer can be viewed as the sweeping around the center line of steel wire by a circle(the diameter is j\u03b4 ) , and in the process of sweeping, wire round face always keeps perpendicular to the center line of the wire rope. If P moves an angle of \u03d5 by turn around the z axis, then the corresponding quantity tz of moving along the z axis by the geometric relationship can be calculated. tan j j r tz \u03d5 \u03b2 = (2) In the form, j\u03b2 stands for twist angle of the j layer of wire rope, jr stands for twisting radius of the j layer of wire rope. Suppose the j layer have jn steel wire, then the angle jji ni \u03c0\u03b8 21\u2212= between the first steel wire and the i steel wire is shown in Fig. 2. Available, the trajectory equation of center line of the outer wire rope is as following: ( ) cos( 2 ( 1) / ) ( ) sin( 2 ( 1) / ) ( ) / tan j j j j j j j j x r k i n y r k i n z r \u03d5 \u03d5 \u03b3 \u03c0 \u03d5 \u03d5 \u03b3 \u03c0 \u03d5 \u03d5 \u03b2 = + + \u2212 = + + \u2212 = (3) In the form, k stands for the coefficient of the wire rope twisting direction, 1k = ,that is to say, right-handed twisted, 1k = \u2212 ,left-handed twisted. The center steel wire of the spiral strand makes linear spiral twist around the center line of the wire rope, space curve equation of center steel wire and curve equation of the center line of the linear strand outer steel wire is basically the same, the trajectory equation of the center line of the wire rope in the i layer is as following: 2 1cos 2 1sin tan s g s j g s g ix R n iy R n R z \u03c0\u03d5 \u03bb\u03d5 \u03c0\u03d5 \u03bb\u03d5 \u03b3 \u03d5 \u03d5 \u03b2 \u2212= + \u2212= + + = (4) In the form, sR stands for the radius of spiral strand twist, gn stands for the number of wire rope in the outer strands, g\u03b2 stands for strand twist angle"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0003962_ias.2000.881983-Figure1-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0003962_ias.2000.881983-Figure1-1.png",
+        "caption": "Fig. 1 - Basic principle.",
+        "texts": [
+            " It is consistent with sensorless or non-absolute position transducers based drives, with a wide range of motors (isotropic and nonisotropic, BLDCM and PMSM) and it does not require the knowledge of any of the motor parameters. 11. BASIC PRINCIPLE The basic approach is the well-known method to estimate the rotor position by using the inductance variation due to the magnets position and an impressed stator current. A suitable sequence of voltage pulses is applied to the stator windings at standstill and the evaluation of the peak value of the current leads to the rotor position estimation. The basic principle can be explained with reference to the schematic PM motor shown in Fig. 1 as follows. In the magnetic circuit of phase A , two effects are superimposed: the PM flux ( ~ A , M ) and the phase current flux (YA). If the phase current is positive (iA+), the total flux linkage with the winding is: V A , t o t = V A , M + y A + * (1) If the phase current is suitable high, the total flux saturates Now, if we change the sign of the current, the total flux the magnetic circuit and the winding inductance is low. becomes v A , t o t = V A , M - V A - . (4 In this case the magnetic circuit goes out of saturation and the winding inductance is h i g h "
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0001768_a:1008839925389-Figure1-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0001768_a:1008839925389-Figure1-1.png",
+        "caption": "Figure 1. This figure shows a five-link biped in the sagittal plane. It has seven degrees of freedom, five joint angles, and two horizontal and vertical displacements.",
+        "texts": [
+            " An algorithm is developed that identifies actuators that contribute to the rhythmic movement, to the support forces and to both. The algorithm also quantifies the contribution of the same actuator in supporting the weight, in breaking contact, i.e., lifting of the foot, and in making contact for the coordination of rhythmic movement. In Section 2, the system model is presented. The control strategy is discussed in Section 3. Simulations and a summary are presented in Section 4 and Section 5. A five-link sagittal biped (Fig. 1) is considered for rhythmic movement in this paper. Its parameters (Table 1) are such that they reflect those of a human being. The bottom two links are analogous to the lower leg. The middle two links are analogous to the thigh, and the upper link is analogous to the torso of a human being. The model has seven degrees of freedom: five joint angles, \u03b8i (i = 1, . . . , 5), and two translational degrees of freedom, (x3, y3), arbitrarily selected to be P1: KCU/RKB P2: STR/SRK P3: STR/SRK QC: Autonomous Robots KL465-03 July 1, 1997 15:43 A Control Strategy for Terrain Adaptive Bipedal Locomotion 245 those of the center of mass of the torso"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0003947_rtd2002-1642-Figure5-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0003947_rtd2002-1642-Figure5-1.png",
+        "caption": "Figure 5: The Lenoir link and the exchanged force",
+        "texts": [],
+        "surrounding_texts": [
+            "The ADAMS detailed model of the axle-box is described in the following figure. The model is composed of 7 rigid bodies: \u2022 Bogieframe: the bogieframe is connected to ground through a traslational joint, which allow only the vertical motion (Z). \u2022 Lenoir-Link (2 bodies): The Lenoir-link has been built as two separate parts, one linked to the bogie and the other to the spring holder, both with two revolute joints. The two parts are then linked each other with a translational joint and a single force, which act as a unilateral bumpstop. When a force is applied to the spring holder by the spring the two parts of the link are moved away and the bumpstop operate such that the force is transferred to the bogie. \u2022 Spring holder: The Spring holder keeps the inner spring in the left side of the axle-box; it is connected to the LenoirLink as shown above. The Link inclination split the force supplied by the spring in two components in the X-Z plane. The spring-holder is connected to the left side of the pusher with a bumpstop so that the force given by the Lenoir-Link in the X direction is transferred to the pusher itself. \u2022 Pusher: the pusher is connected to the bogieframe with a traslational joint which allow only the relative motion in the X direction. The right side of the pusher is connected to the axle box with a force vector, this element model the first friction surface. In the X direction the force vector act as a bumpstop, the value of the X force is then used as the Normal force for the bi-dimensional friction force implemented in the Y and Z direction. \u2022 Axle-Box: The primary friction surface is located in the left side of the Axle-Box in the Y-Z plane. This surface has been modeled using four force vectors, one to each vertex of the surface. Each Force vector is modeled as the one on the pusher, so that the total friction force is distributed among the four forces depending on the normal force acting on each vertex. The vertical load is transferred from the Axle-Box to the bogie frame thought four springs, the outer springs have a gap which is closed only in the laden condition, so that in the tare load condition only the inner springs support the load. Moreover the right inner spring act thought the Link as described above. The axle box is connected to the axle with a revolute joint. \u2022 Axle: The axle in this model is connected to ground with a planar joint. All the bumpstop elements have been modeled with a stiffness of 1\u22c5108 N/m and the damping of 1000 N\u22c5s/m."
+        ]
+    },
+    {
+        "image_filename": "designv11_60_0001235_b978-0-08-099425-3.00005-4-Figure5.37-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0001235_b978-0-08-099425-3.00005-4-Figure5.37-1.png",
+        "caption": "FIGURE 5.37",
+        "texts": [
+            " It is also typical, once the driver feels sufficient confidence, to begin applying throttle, and driving forces, during cornering before exiting the bend. For such situations a tyre model must be able to deal with combined tractive and cornering forces, a situation referred to as comprehensive slip. The basic law of friction relating frictional force to normal force can be of assistance when considering combinations of longitudinal driving or braking forces with Pure and combined braking and cornering forces. lateral cornering forces. The treatment here concentrates on lateral forces due to slip angle with camber angle set to zero. Figure 5.37 initially shows a tyre subject to pure braking or cornering force where in each case the slip in the ground plane is such that the tyre force produced is a peak value, this being mFz, the peak coefficient of friction multiplied by tyre load. Plotting lateral force against longitudinal force (friction circle). For pure cornering the peak force will occur at a relatively large slip angle where in Figure 5.37 some lateral distortion of the contact patch is indicated together with a small amount of pneumatic trail. For a tyre running at a large slip angle with additional braking force the resultant ground plane force is still equal to mFz but the resultant force direction opposes the direction of sliding. The longitudinal and lateral forces Fx and Fy are now components of the resultant force. Thus it can be seen that the simultaneous action of longitudinal and lateral slip reduces the amount of cornering or braking/driving force that may be obtained independently",
+            " As driving or braking force is added the maximum resultant force that can be achieved is defined by points lying on a curve of radius mFz referred to as the \u2018Friction Circle\u2019 or sometimes the \u2018Friction Ellipse\u2019 as some tyres will have more capability in traction or cornering leading to an elliptical boundary shape. Figure 5.38 shows, for a typical tyre, the general form of the friction circle diagram for the full range of driving and braking forces. Note that only lateral forces due to positive slip angle are presented and a similar diagram would exist for measurements taken at negative slip angles. Point D represents an example of a position where the tyre is operating at the friction limit for combined braking and cornering, as shown in Figure 5.37 where it is clear that the amount of braking or cornering force that could be produced independently is reduced and that the magnitude of FR is simply 2 y 2 xR FFF += \u00f05:42\u00de It can also be noted that the curves are not symmetric in that lateral forces initially increase slightly as braking force is applied. As discussed earlier the braking force adds circumferential tension to the tyre material entering the contact patch. This stress stiffening effect can be seen to raise the lateral force slightly, while the reversal of longitudinal force to driving leads to a reduction",
+            " As the curves approach the friction limit it can be observed that they turn inwards. For a fixed slip angle the longitudinal slip is increased moving along the curves, for either braking or driving, until the point where both lateral force and longitudinal force reduce, hence causing the curves to bend back. By plotting aligning moment against longitudinal force, it can be seen that the opposite can occur (Phillips, 2000) and that adding braking force reduces aligning moment and adding driving force raises it. Referring back to Figure 5.37 the bottom diagram shows a tyre running at moderate slip angle producing a lateral force Fy along a line of action set back from the centre by the pneumatic trail. The contact patch is shown displaced laterally due to the cornering force so that for simultaneous braking the braking force produces a moment that would subtract from the existing aligning moment due to the product of lateral force and pneumatic trail. From the diagram at the bottom of Figure 5.37 it is also clear that the simultaneous application of a driving force would produce a moment that would add to the aligning moment due to the product of lateral force and pneumatic trail. At higher braking forces the effect may cause the aligning moment to go negative. The friction circle or ellipse is also a way to monitor the performance of a race car driver using instrumented measurements of lateral and longitudinal accelerations, sometimes called the \u2018geg\u2019 diagram. Comparing this diagram with known tyre data it is possible to see how well the driver performs keeping the vehicle close to the friction limits of the tyres"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0000860_978-3-319-44735-3_1-Figure1.4-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0000860_978-3-319-44735-3_1-Figure1.4-1.png",
+        "caption": "Fig. 1.4 The left side rotors have value greater than applied on the right side rotors, the hexacopter moves to the right, or slows down if it is moving to the left",
+        "texts": [
+            "5 The right side rotors have value greater than applied on the left side rotors, the hexacopter moves to the left, or slows down if it is moving to the right The maneuvers to the right and to the left are achieved by applying forces on the side rotors with different proportion. Thus, it makes the hexacopter rotate around the X-axis and a movement over the Y-axis occurs. For moving to the right or to slow down the movement to the left, the left side rotors have value greater than applied on the right side rotors, as shown the Fig. 1.4. For moving to left or to slow down the movement to right, these forces are inversely applied between the left and right side rotors as shown in the Fig. 1.5. To rotate the hexacopter in the Z-axis (the yaw movement) the forces are applied alternating among the rotors, shown in Figs. 1.6 and 1.7. It is important to note that if a rotor is set to rotate clockwise, therefore, the adjacent rotor is set to rotate counterclockwise. The real propellers are built with clockwise twist and counterclockwise"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0000887_b978-0-12-803581-8.09314-0-Figure1-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0000887_b978-0-12-803581-8.09314-0-Figure1-1.png",
+        "caption": "Fig. 1 Schematic diagram showing the anatomy of an electrochemical glucose sensor.",
+        "texts": [
+            " CGM data are used to ascertain real time insulin needs for patients to control their blood glucose levels. Enzymatic amperometric CGMs are the most common commercially available sensing devices, representing a mature technology with many technological manifestations developed for over 50 years. Amperometric sensors monitor electrical currents generated from electron exchanges either directly or indirectly between physiological fluids containing glucose and an electrode poised to perform electrochemistry on glucose. Fig. 1 shows an illustration of the anatomy of an electrochemical continuous glucose sensor. Glucose oxidase (GOx) is commonly used as an immobilized nonmammalian enzyme for CGM glucose sensing due to its high selectivity for substrate glucose. GOx is readily available, inexpensive and has shown good resilience to changes in environmental factors such as pH, ionic strength, and temperature when compared with many other enzymes.2 Given its stability, GOx endures routine handling, immobilization, sensor manufacturing and storage, and patient use conditions required for CGM reliability and clinical use"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0001834_1.2834133-Figure2-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0001834_1.2834133-Figure2-1.png",
+        "caption": "Fig. 2 Finite element model and boundary conditions",
+        "texts": [
+            " Briefly, the two-dimensional model consisted of a cylinder that was bounded by two, 0.0127 m thick and 0.0381 m long steel plates. Layers of coating were incorporated into the model by attaching very thin (1-10 fj,m) rectangular areas to the internal edges of the plates. For this purpose a fine mesh was required to produce adequate element aspect ratios in the coating. Specifically, there were 200 horizontal divisions along each coating layer, making the elements within the coating 0.1905 mm wide. The resulting three-dimensional model used in this work, which is depicted in Fig. 2, was generated by rotating the right half of the 2-D FEM (see Lovell et al , 1996b) 180 degrees about the vertical axis. Due to the tangential friction forces, the model could only take advantage of the symmetry along the balls rolling path (z-axis). The ball, plate and coatings all consisted of eight-noded, three-dimensional solid elements which could be defined with up to nine orthotropic material constants (Eu, \u00a322, \u00a333, G,2, Go, G23, 2\u0302 12, i^u, and ^\u0302 23). At the completion of the three-dimensional solid element genera tion, the model depicted in Fig",
+            " Briefly, Cu is Young modulus (\u00a311) of the coating in the direction normal to the substrate surface, C33 is Young's modulus of the coating in the direction parallel (\u00a333) to the substrate surface, C44 is the shear modulus (G) of the coating, and C^ and C13 are used to determine Poisson ratio of the coating from the relationship: Cn Cn + C,. (2) It is important to note that with the transversely isotropic as sumption for the coatings, microstructural boundary conditions for the solid lubricant films were not required. Hence, in the finite element model depicted in Fig. 2, perfect bonding between the coating layer and the substrate surface was modeled. For the purpose of performing FEM simulations, a constant friction coefficient was implemented for each of the transversely iso tropic coatings. As there are a number of variables that effect the friction coefficient, the values reported in the literature dra matically vary for each lubricant, making it extremely difficult to arrive at a single coefficient. Nonetheless, for uniformity, a single value of the friction coefficient was assumed for an environment of air",
+            "150), and Cadnium (0.215). It should be noted, however, that in unsealed air lay ered hexagonal films may not be good lubricants because they degrade after relatively short time periods (Roberts, 1987). Additionally, a single value of friction is indeed an idealization of a complicated set of variables and was only used to complete the finite element simulations presented in this work. 2.3 Boundary Conditions and Loading. The required boundary conditions for the three-dimensional FEM are also depicted in Fig. 2. Similar to a rolling element bearing, con straints on the model are appUed to the top surface of the upper plate and the bottom surface of the lower plate. On the upper surface of the top plate, all of the nodes have their degrees of freedom ( and MJ coupled. Then by constraining one of the nodes along the top of the upper plate from horizontal displacements, all of the nodes along the top edge of upper plate were held rigid in the x-direction and allowed to undergo displacements in the vertical (Uy) direction"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0000860_978-3-319-44735-3_1-Figure1.33-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0000860_978-3-319-44735-3_1-Figure1.33-1.png",
+        "caption": "Fig. 1.33 Input linguistic variables and their membership functions for: Vertical distance",
+        "texts": [],
+        "surrounding_texts": [
+            "Vertical navigation fuzzy controller is similar to the horizontal navigation controller. However, it controls the movement on the Z-axis. Figure1.32 shows the block diagram of this controller. This controller takes as input the vertical distance to the target point, as well as the vertical speed. The first one is the error in the Euler approximation of Z-axis, while the second one is the difference in distance (Z-axis). Linguistic variables for vertical distance and vertical speed are presented in Figs. 1.33 and 1.34, respectively. Table1.5 shows the 35 rules that compose the vertical navigation fuzzy controller. As the result, this controller sets the omega throttle variable (OThrottle), presented in Fig. 1.35 which is decomposed in the amount of power applied on all rotors, increasing or decreasing the overall lift force making the hexacopter fly on higher or lower altitude. It is worth noting that this presents a smooth control approach similar to the horizontal navigation, i.e. the power applied on the rotors decreases along with vertical speed as the hexacopter comes closer to target altitude. The fuzzy surface control for vertical navigation and hovering is shown in Fig. 1.36. The altitude is maintained by controlling the throttle applied onto the all rotors. The input information is taken from the GPS sensor. The vertical speed is used to avoid the hexacopter to oscillate up and down. It is similar as the acceleration is used to avoid oscillation in the pitch and roll controller. It is worth to note that the ZERO output not means zero value, but the value that the hexacopter is hovering. To realize the relationship between altitude error and vertical speed, suppose the altitude error is zero, the hexacopter is in the target vertical position, but also suppose the vertical speed is positive, perhaps 0.4 or higher. It means the hexacopter reached the target and goes beyond because it is in movement to up. It must be slowed down. Therefore, the controller sets the output value to a value lower than the ZERO, causing the hexacopter to slow down. On the other hand, when the altitude is zero and the vertical speed is negative it means the hexacopter is falling down. In this condition, the controller must to set output to a value higher than the ZERO just to make the hexacopter stop the falling."
+        ]
+    },
+    {
+        "image_filename": "designv11_60_0003130_bf03186079-Figure2-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0003130_bf03186079-Figure2-1.png",
+        "caption": "Fig. 2. Schematic illustration of the restraint/relaxation U-type hot cracking tester.",
+        "texts": [
+            " Region I corresponds to a zone from the surface of the plate to the point where the curvature of the fusion boundary begins to change. Region II is a zone from the boundary of Region I around the necked fusion boundary. Region III corresponds to the zone from the boundary of Region II to the left zone where the fusion boundary is parallel to the direction of thickness. The U-type hot cracking test was conducted to determine the effects of grain size and homogenization heat treatment on liquation cracking susceptibility during EB welding. Fig. 2 shows the configuration and dimensions of the device used to run the U-type hot cracking test. A specimen (100 mml\u00d7 10 mmw\u00d75 mmt) was arranged between both restraint beams on this tester, which was designed to facilitate EB welding. When EB welding was performed in this state, the specimen accordingly underwent expansion due to the thermal stress imposed by the temperature rise. Liquation cracking was induced by the corresponding strain. The adopted EB welding conditions were an accelerated voltage of 45 kV, a beam current of 100 mA and a welding speed of 2"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0000018_2016-01-1468-Figure2-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0000018_2016-01-1468-Figure2-1.png",
+        "caption": "Figure 2. Tire depositing a yaw mark (not to scale).",
+        "texts": [
+            "1 Because the striations are affected by both the heading angle of the tire and the amount of braking, these tire marks offer a glimpse into the actions of the driver when the tire marks were deposited. A full review of tire striations in the literature can be found in a forthcoming SAE publication [2] In previous work, a theoretical model for determining longitudinal tire slip from striation marks was developed [1]. Equation 1 can be used to calculate the longitudinal slip, using the striation angle, \u03b8, and the slip angle, \u03b1. The variables in Equation 1 are depicted in Figure 2. Full scale vehicle yaw testing was conducted to validate the equation. It was found that the model offered insight into the braking actions of drivers at the time the tire marks were being deposited. 1. Throughout this paper, the term \u201cfull braking\u201d refers to the locking of the wheel, when its rotational velocity goes to zero. This condition would be consistent with full brake application in a vehicle that is not equipped with an antilock brake system (ABS). (1) CITATION: Beauchamp, G., Thornton, D"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0000513_ipemc.2016.7512307-Figure3-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0000513_ipemc.2016.7512307-Figure3-1.png",
+        "caption": "Fig. 3. EI core inductor equivalent circuit",
+        "texts": [
+            " (14) From Part A, varying values of self and mutual inductance are needed for the proposed stationary reference frame feedforward terms. According to the nonlinear B-H characteristic of inductor, a look-up table between current and inductance including mutual flux could be acquired through off-line test of the inductor. A three phase inductor can be modeled as a magnetic equivalent circuit with magneto-motive force (MMF), flux and magnetic reluctance, which can be looked upon as analogous to an electric circuit [6]. Equivalent circuit based on above description is shown in Fig.3. A DC bias table of an inductor under consideration at 10 kHz from 0A to 10A is shown in Fig.4, where the relationship between input winding current and corresponding dynamic inductance can be acquired. Therefore, from (15)-(18), flux linkage can be integrated by dynamic inductance, and corresponding flux and magneto-motive force can be calculated. Dynamic permeance, which is inverse of reluctance, is the differential of flux by MMF. Therefore, table regarding current, self-inductance, deduced self-flux, and selfdynamic permeance can be obtained as follows"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0002549_10693067_1-Figure4-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0002549_10693067_1-Figure4-1.png",
+        "caption": "Fig. 4. Graph of movement ranges dependencies represented by an electric circuit",
+        "texts": [
+            " To be bound to the team rules, the agent must first commit to them either by creating a new team with other agent or by joining an already existing team. However now for the simplicity of the presentation we consider a single agent as a team. For this version the mechanism is quite simple and consists in making, by team T, a fusion with any team that has goal consistent with the goal of T. This results in a situation where the teams are expanding and shrinking very fast. For this version of team formation, the mechanism is more subtle. It is based on the idea of electric circuit with resistors. We explain this using Figure 4. Since team A has a resource needed by team H , it is clear that team H has to make a fusion with team A. However, because of the movement ranges of its members, team H cannot get to team A, so that H needs intermediary teams to get there. The graph of relations between the movement ranges of the teams, considered relatively to particular resources needed by a particular team, can be interpreted as an electric circuit with teams as nodes and resistors between the teams having the following resistance: \u2013 1 if the teams can meet via their representatives, \u2013 +\u221e if the goals of the teams are mutually contradictory"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0000364_intelse.2016.7475150-Figure2-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0000364_intelse.2016.7475150-Figure2-1.png",
+        "caption": "Fig. 2. Definition of body heading angle 1/J and wind triangle .",
+        "texts": [],
+        "surrounding_texts": [
+            "The ground velocity V 9 is the sununation of two velocity vectors, the wind velocity vector V w and airspeed vector Va. (1) Consider a level flight with fixed North-East inertial frame, the analytical expression for wind triangle can be derived as [21]: Vg (~~~~) = (:: ) + Va G~~~) , (2) equivalently can also be expressed as Vg (~~~~) - (:: ) = Va G~~~) . (3) At a given altitude, closed loop speed controller or an open loop preset throttle setting is usually required for UAVs to maintain their airspeed Va. But in the presence of wind despite of maintaining the constant airspeed, Vg will be constantly varying depending upon the magnitude and direction of wind. For the derivation of relationship between X and 'i/J using (3), the row vector (- sin X, cos X) is multiplied on both sides and it yields the form W n sin X - We cos X = Va ( - sin X cos 'i/J + cos X sin 'i/J ), (4) which can be written in simplified form as: 'i/J- x = sin- l (:a (Wn Sinx - WeCOsX)). (5) In the presence of wind disturbance, the idea of coordinated turn is explained in detail in [21]. In literature the coordinated turning is explained with the idea of no side force in the vehicle's body frame, thus side-slip angle,8 = O. During flight aerial vehicle minimizes the cross track error by producing lateral accelerations by tilting the component of aerodynamic Definition of Kinematic Variables. lift in the direction of turn. Fig. 4 explains the concept of bank to turn maneuver, in which the vehicle will bank to produce required lateral accelerations. Therefore the control input for o~r case is the reference bank angle commands \u00a2reJ. In FIg. 4, FliJt represents the aerodynamic lift vector, which is further resolved in two components. One component balances the centrifugal force acting on the vehicle due to the turn and other component (Fli Jt cos \u00a2) balances the weight of the vehicle [21]. FliJtcOS\u00a2 = mg, mV2 FliJt sin \u00a2cos(x -1j; ) = --1-. (6) where 9 is the acceleration due to gravity, m represents the mass of the vehicle and R is the radius of turn. From (6) we can have V 2 tan \u00a2 cos(X - 1j;) = -.lL Rg (7) For a steady turn Vg = RX, therefore (7) can be written as: V\u00b7 tan \u00a2cos(x -1j;) = gX. (8) 9 As the intercept course angle the difference of course angle and the desired course angle i.e., XE = X - XR ' we can have: V (. + . ) t an \u00a2cos(x -1j;) = 9 XE XR. (9) 9 Solving (9) for XE we have . gtan \u00a2 . XE = -V- cos(X -1j;) - XR , (10) 9 The above state equation using (5) takes the following form . _ 9 tan \u00a2 ( . -1 ( - W n sin X W e cos X) ) . XE - -V--cos sm + - X 9 Va Va R ' (11) here XR is the rate of change in reference course angle. This term is non zero for following circular arcs but for straight flight XR = O. Cross track error deviation y is the closet distance to path Fig. 3 and can be its rate of change can be derived using the relation (12) The closed loop dynamics from \u00a2reJ to \u00a2 of the autopilot control loop is approximated using a first order filter [21]: \u00a2 1 \u00a2reJ (13) TS + 1 and incorporated into guidance design. Equations (11), (12) and (l3) represent the overall dynamics for the outer loop guidance design problem with y, XE and \u00a2 as state variables and to avert the cross lateral error guidance loop will gen erate \u00a2reJ as the control signal. Previously we showed that coordinated turn condition can be described by X = ~ tan \u00a2 cos(X -1j;) (14) 9 It can also be expressed in terms of heading and the airspeed as in [21]. . 9 'IjJ = V-tan \u00a2 (15) a The coordinated turn expression in (15) holds true in the presence of wind and will be used later for estimation of airspeed Va."
+        ]
+    },
+    {
+        "image_filename": "designv11_60_0000134_ecce.2014.6954068-Figure5-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0000134_ecce.2014.6954068-Figure5-1.png",
+        "caption": "Figure 5. Elliptical approximation of the hysteresis loops",
+        "texts": [
+            " Br is called the remanence which is the residual value of flux density when the applied field becomes zero. Hc is called the coercive force or coercivity of the material which is the negative value of the applied magnetic field intensity, required to force the flux density to zero. Elliptical modeling is a way to approximate the shape of the hysteresis loops of a material. In elliptical modeling, the hysteresis curves of the rotor material are approximated by a group of inclined ellipses of similar shapes [10-12, 21]. The elliptical approximations of the hysteresis loops are shown in Fig. 5. The trajectory of a B-H curve lies on the ellipse if the motion around the ellipse is in the counter clockwise direction. The lag angle between B and H remains constant as long as the direction of the motion remains counter clockwise. When the direction of motion changes to clockwise, it results in a movement between the inner ellipses with different lag angles. This phenomena is illustrated in Fig. 5 for one inner ellipse and one outer ellipse with two different lag angles \u03b4 and \ud6ff , respectively. The flux density B and the magnetic field intensity H in an elliptical model can be expressed as follows [10-12, 21], \ud435 = \ud435 cos(\ud714\ud461 \u2212 \ud713 \u2212 \ud713 ) (1) \ud43b = \ud435 \ud707 \ud450\ud45c\ud460(\ud714\ud461 \u2212 \ud713 \u2212 \ud713 + \ud6ff) (2) \ud713 = tan \ud45f sin\ud6ff \ud45d\ud707 \ud459 \ud45d\ud461 \ud707 \ud45f + \ud45f \ud45d\ud707 cos\ud6ff (3) where \ud435 is the maximum flux density of the rotor material, \u03bc is the permeability of the elliptic hysteresis loop, \u03c9 is the synchronous angular frequency, \ud713 (\ud713 = \ud45d\ud703 ; \ud703 is the mechanical angle of the rotor and p is the number of pole pairs) is the electrical angle coordinate in the stator frame, \ud713 is the phase shift and \ud6ff is the hysteresis lag angle between B and H"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0000739_978-1-4471-4976-7_91-1-Figure6-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0000739_978-1-4471-4976-7_91-1-Figure6-1.png",
+        "caption": "Fig. 6 The movement of the mobile platform and links",
+        "texts": [
+            " 37 on both sides with vi wi yields vi wi\u00f0 \u00de v \u00bc ui vi\u00f0 \u00de wi _yi (38) from which the SPM velocity equation is obtained as Page 10 of 17 Av \u00bc B _u (39) with A \u00bc v1 w1 v2 w2 v3 w3 T (40a) B \u00bc diag u1 v1\u00f0 \u00de w1 u2 v2\u00f0 \u00de w2 u3 v3\u00f0 \u00de w3 \u00bd (40b) where _u\u00bc _y1; _y2; _y3 h iT . Matrices A and B are the forward and inverse Jacobian matrices of the manipulator, respectively. The kinematic Jacobian matrix J of the manipulator can be expressed as follows as long as matrix A is not singular: J \u00bc B 1A \u00bc j1 j2 j3\u00bd T ; ji \u00bc vi wi ui vi wi (41) Equation 38 is thus rewritten for the single joint velocity as _yi \u00bc jTi v (42) The motions of the link and mobile platform are shown in Fig. 6. The angle rates _f \u00bc _f; _y; _s h iT and the angular velocity v are linearly dependent, namely, v\u00bcC _f. Differentiating the equation with respect to time yields _v \u00bc c\u20acf\u00fe _c _f (43) MatrixC is dependent on the rotations. For example, a rotation with the Euler convention ZY Z of Q \u00bc Rz(f)Ry(y)Rz(s f), matrix C is Page 11 of 17 c \u00bc sycf sf sycf sysf cf sysf 1 cy 0 cf 2 4 3 5 (44) The velocity _c of the intermediate joint of ith leg is found by making use of Eq. 37 to eliminate _yi and _xi. Dot-multiplying Eq"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0002215_0094-114x(95)00011-m-Figure4-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0002215_0094-114x(95)00011-m-Figure4-1.png",
+        "caption": "Fig. 4. Comparison of first and second order coordination of the two-revolute system.",
+        "texts": [
+            " 6 7 (PJ)x could also have been graphically determined from the processed vision data using the well-known Bobillier construction [6], averting the computation association with the Euler-Savary Equation. The Bobillier construction amounts to additional vision processing. With these first and second-order speed ratios, the manipulator can now be controlled to develop the instantaneous second-order output-space trajectory geometry by the coordinating second-order Taylor series in equation (27), i.e. /t = n2 + \u00bdn'22 = -3.312 + \u00bd4.6722. Figure 4 shows the trajectory generation which results from a first- and second-order of the coordinating Taylor series, a direct comparison of first- and second-orders of coordination. The accuracy of the trajectory tracking is significantly enhanced by the second-order of coordination, readily provided by Curvature Theory. 3.3. Example 2: a non-holonomic system A schematic of a cart system with a solid axle and two drive wheels is shown in Fig. 5, with its corresponding Polar Line and the location of the canonical system"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0000154_icelmach.2014.6960384-Figure1-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0000154_icelmach.2014.6960384-Figure1-1.png",
+        "caption": "Fig. 1. Overview of the used reference system for describing the rotor displacement in Cartesian coordinates xd, yd and in polar coordinates r, \u03c6. The resulting magnetic force Fd with its components Fr, F\u03c6 in polar coordinates acting on the rotor in case of a rotor displacement are shown in the right figure.",
+        "texts": [
+            " This comparison necessitates a FE model with voltage fed coils, a sliding interface between rotor and stator for the consideration of rotation as well as deformable air gap elements for taking the rotor displacement into account. The FE-based circuit model approach itself and the required workflow for the look-up table creation process has been already presented in [8]. However, the basic essentials of this approach are summarized in this work again to achieve a better understandability and clearness of the presented simulation workflow. Rotor eccentricity is a displacement of the rotor from its ideal position in the center of the stator hole, as shown in Fig. 1. This displacement leads to an additional magnetic force d F acting on the rotor because the magnetic symmetry in the machine is annihilated. The same force with opposite sign acts on the stator too. The radial component of this force has the same direction as the displacement itself. Thus, it acts as an additional magnetic drag that leads to an intensification of the displacement, as shown for example in [10]. Furthermore, an increase of noise and vibration in the machine is caused by the magnetic force",
+            " Thus, any eddy current effects like skin effect or eddy currents in the magnets are neglected. These effects mainly increase the losses of the PMSM but have negligible influence to the electro-mechanical behavior of the machine. The input quantities of the magneto-static FEM model are the angular rotor position Rot \u03b1 , the displacement parameters , Rot Rot r \u03d5 of the rotor and the machine currents , d q i i in the rotor related d-q reference frame. The rotor displacement is defined in polar coordinates as shown in Fig. 1. This set of five input variables is sufficient to define the machine state in case of an axially parallel rotor displacement (Fig. 2.a) uniquely. All look-up tables are parameterized by these five variables. The used look-up table quantities are the phase to phase flux linkages , AB BC \u03a8 \u03a8 , the machine torque T and the force components , r F F\u03d5 caused by the rotor displacement. All chosen look-up table quantities can be evaluated during the FEM post processing. A quint-cubic spline interpolation is used for the evaluation of , , , AB BC r T F\u03a8 \u03a8 and F\u03d5 "
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0002030_20.560098-Figure8-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0002030_20.560098-Figure8-1.png",
+        "caption": "Fig. 8. Schematic diagram showing the \u201cmajor\u201d hysteresis loop and the \u201cminor\u201d hysteresis loops.",
+        "texts": [
+            " The same procedure was used to represent the hysteresis loop for a distorted flux-density waveform with minor loops occurring on the downside, as shown in Fig. 4 . The original and reconstructed versions of the hysteresis loop for the distorted waveform of Fig. 6 are shown in Fig. 7. DISTORTED WAVEFORM This section presents a method to predict the hysteresis loop for a distorted waveform using the hysteresis loop obtained with a sinusoidal excitation. If a core under test is premagnetized to the extent shown by the point X in Fig. 8, and then the direct current producing H is gradually reduced to zero, the flux density in the core will correspond to point V. If H is increased again in the same r I j B C T ) 1 5 I 1c - 1 5 - i o -05 - - . -. . Original curve - produced curve . direction, the value of B will be carried-back to X along Fig. 5 . Original and reconstructed version of a hysteresis curve MOHAMMED e/ al.: FOURIER DESCRIPTOR MODEL OF HYSTERESlS LOOPS 689 Fig. 6 . Distorted hysteresis loop with minor loop. CARTES SW . INFORMATlOh \u2019 RUN \u2019 RETURN c o e f f - 48 ~~ . . -. . . . original curve. - produced curve. Fig. 7. Original and reconstructed version of a distorted hysteresis loop. a path like VWX, as shown in Fig. 8. If the core is magnetized to the level corresponding to the point U , a small value of H is applied and varied between positive and negative values around a small loop like US. The curves are characterized by \u201cminor\u201d loop or \u201csubsidiary\u201d (VX and U S ) hysteresis loops [12]. If a sinusoidal flux density waveform is applied to the test core with a voltage large enough to saturate the core, the sequence of events occurring during the time intervals of one cycle of the waveform shown in Fig. 9 can be explained as follows"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0002292_s0022-5096(05)80019-5-Figure1-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0002292_s0022-5096(05)80019-5-Figure1-1.png",
+        "caption": "FIG. 1. Schematic representation of a chevron fold and associated coordinate systems.",
+        "texts": [
+            " The directional properties of the material on the two sides of P display initially a mirror symmetry with respect to P. The evolution of these directional properties, that is, of the orthotropic (or transversely isotropic) axes, will be investigated. The associated kinematics and stresses are intended to model the development of chevron folds, a term borrowed from a corresponding geological rockformation but given here a more general meaning as mentioned in the Introduction. An illustrative example of this configuration is shown in Fig. 1 by two decks of cards symmetrically disposed in reference to their contact plane P. In the following, material coordinates in relation to ei will be denoted by X, for the reference and xi for current configurations. The motion of the aforementioned halfspaces are restricted by the following conditions (i) the contact along P is perfect. (ii) the orientation of the contact plane P is fixed and parallel to the e~--e 3 co- ordinate plane. (iii) the individual deformation gradients F}7~= Ox~lc3Xj for the two half-spaces x2 > 0 and x2 < 0, respectively, corresponding to c~ = 1 and 2, are uniform"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0002603_0301-679x(88)90122-3-Figure1-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0002603_0301-679x(88)90122-3-Figure1-1.png",
+        "caption": "Fig 1 Configuration of a misaligned journal bearing",
+        "texts": [
+            " 0 7 ( ~ - [ lYu I + 1\u00a5~ I ] )1 /2 (5) where I ru I and t \u00a5~ I refer to the shear stress at the upper end and lower walls respectively. TRIBOLOGY international 0301-679X/88/010015-05 $3.00 \u00a9 1988 Butterworth & Co (Publishers) Ltd 15 Safar and R iad - friction coefficient of misaligned turbulent flow journal bearing The pressure equation is subjected to Giimbel's boundary conditions which are P(o , , 3 , )=o (o2, v) = o , L ~(o,+- b ) = o (6a) ~P 0 aO = 0 2 = 0 (6b) (6c) H(O,'7) = 1 + e cos(0 - 4 ) The eccentricity ratio, e, and the attitude angle, 4, are functions of 7 as shown in Fig 1, and are given by (3') = lko + arctan [ X sin (0 - ~ o) eo + ;~ c o s ( ~ - 4 o ) (7) and l (8 ) The pressure equation is written in finite difference form, and solved numerically by Gauss-Seidel's iteration method. The approach employed by Mokhtar et al 7 is used to represent the fluid film thickness which is varying in both the circumferential and axial directions. The nondimensional film thickness is represented by e ( 7 ) --- [eo 2 + 2Xeo cos(q~- 4o) + )t2l ,/2 (9) where eo is the eccentricity ratio at the mid-plane, and X is the misalignment ratio defined by X ~- \u00b1~V (R/c) (10) 16 February 88 Vol 21 No 1 Safar a n d R i a d - f r i c t i o n c o e f f i c i e n t o f m isa l i gned t u r b u l e n t f l o w j o u r n a l bear ing N o t a t i o n c Radial clearance between the journal and its bearing D Bearing diameter D m Degree of misalignment e Eccentricity f Coefficient of friction (= F/W) fi Increase in coefficient of friction due to misalignment fo Coefficient of friction for an aligned bearing fp Percentage increase in coefficient of friction due to misalignment F Friction force ff Nondimensional friction force (= F/IzNLD (R/c)) H Nondimensional film thickness (= h/c) h Film thickness L Bearing length N Rotational speed P Lubricant hydrodynamic pressure fi Nondimensional pressure (= P/laN (R/e) 2) R Shaft radius R e 'Reynolds number (= Uc/u) U Shaft rotational speed W Load carrying capacity Nondimensional load (= W/laNLD (R/e) 2 ) x Cartesian coordinate y Cartesian coordinate z Cartesian coordinate /3 Misalignment angle 7 Nondimensional coordinate (= z/R) c Eccentricity ratio (= e/c) em Eddy viscosity co Eccentricity ratio at mid-plane r/ Nondimensional coordinate (= y/h) 0 Nondimensional coordinate (= x/R) 01 The angle at which the pressure profile starts 02 The angle at which the pressure profile ends X Misalignment ratio (=/37 (R/c)) Xe Misalignment ratio at either bearing ends (= ~(L/2c)) hm Maximum possible value of he /~ Lubricant viscosity p Kinematic viscosity of the lubricant Misalignment directional angle, ie the angle between the plane of misalignment and the axial plane containing the load vector ~\" Shear stress Nondimensional local shear stress (= rC/laU) Angle between the line of centres and the axial plane containing the load vector fro Value of ~ at mid-plane The severity of bearing misalignment is defined by a parameter, Dm, where Dm = Xe/hm he = (1 - [ eos in (\u00a2- ~o)] 2 } 1/2_ Co cos(~- \u00a2o) (11) Therefore, for a given Co, L/D, D m and ~b, the distributions of the eccentricity ratio and the attitude angle along the axial direction are computed, and consequently the film thickness variation is determined"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0000766_0954406216671839-Figure5-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0000766_0954406216671839-Figure5-1.png",
+        "caption": "Figure 5. Finite element model of bevel gear pair (left) and spur gear pair (right).",
+        "texts": [
+            " Mesh stiffness and STE of gear pairs The mesh stiffness of gear pair is the main internal excitation of gear dynamics. The mesh stiffness of straight bevel gear pair is different from that of the spur and helical gear pair, and cannot be calculated analytically. This may be resulted from: (1) the bevel gear pair has an approximate point contact, and (2) the mesh positions are varying in space. In the present work, the finite element method is adopted to calculate the mesh stiffness and the STE for both bevel gear pair and spur gear pair. The finite element model is shown in Figure 5. The hexahedron C3D8R element is adopted to mesh the gear pairs. The element number for the pinion and gear of bevel gear pair is 88,768 and 144,384, respectively. The element number of the spur gear pair is 141,960 and 195,844. It should be noted that the thin web of gear of spur at CORNELL UNIV on September 26, 2016pic.sagepub.comDownloaded from gear pair is considered in the static analysis. The input speed and power of the transmission system are 17000 r/min and 33 kW. The pre-calculated drag torque is then applied"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0003304_(asce)1084-0702(2002)7:5(300)-Figure9-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0003304_(asce)1084-0702(2002)7:5(300)-Figure9-1.png",
+        "caption": "Fig. 9. Splay band channel No. 4 ~units in millimeters!",
+        "texts": [
+            " In the mockup test, fewer than 100 wires (20 wires per layer35 layers) were erected for each strand. The lateral displacement of wires at the splay band was investigated first, after which the lift test was performed. The test was performed using strand No. 4, which was linked to the strand shoe located to the left of the cable centerline ~Fig. 5!. Its wires were expected to be displaced laterally during cable spinning. The splay band channel that simulated the groove of the splay band is shown in Fig. 9. The channel has horizontal and vertical curvatures at the exit side to accommodate flaring of the strand in both planes. Several schemes to arrange the wires were Fig. 5. Strand shoe placement ~units in millimeters! 302 / JOURNAL OF BRIDGE ENGINEERING / SEPTEMBER/OCTOBER 20 J. Bridge Eng. 200 devised and tried, and modifications to the splay band channel were made during the tests, if necessary. Proposal I. Wire Former Use A \u2018\u2018wire former\u2019\u2019 proved to be an effective tool for the arrangement of the wires at the pylon saddle of the Grand Bridge ~Gil and Choi 2001"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0001777_951293-Figure3-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0001777_951293-Figure3-1.png",
+        "caption": "Figure 3. Friction Force vs. Velocity",
+        "texts": [
+            " 1 shows that only body A is elastically deformed) as a result of the nomial pressure. Let us consider a case when the body B is moving in an oscillatory manner sliding on the surface of A in response to an active, reciprocating force S. Let us assume that the body B is sliding to the right (with a positive velocity v) under the influence of the force S. This state of the system is indicated in Figures 2 and 3. Fig. 2 presents the relationship between the friction force and 482 displacement and Fig. 3 shows the friction force as a function of the velocity of B. At uo, the magnitude of the friction force acting on B is F = p d N . . . (2) and is constant. As far as the sum of the active force S and inertia force of the body B is equal to F, B is sliding in the positive x-direction. With a decrease of S, the velocity of B drops to zero. This state is shown as u, in Figures 2 and 3. At this instant, the deformation of both bodies is elastic and the reaction of deflected foundation A is directed to the left",
+            " Further motion to the left and the corresponding events are represented by states w, to w4. The next half of the loop ~ 0 - w ~ is very similar to that of uo-u,. The only exception is that the directions of active and friction forces as well as motion and deformations are opposite in these two cases. If the elastic deformation of the part of the system defining the contact between bodies is small in comparison to the amplitude of relative sliding motion, the two inclined branches of the loop (see Fig. 3) can be shifted to the middle to create the dry friction characteristic in the form that is presented in Fig. 4a. The curve illustrates the single-value dependence between the friction resistance and the relative velocity of moving bodies. This curve, which describes a case of \"weak\" damping (see [7]), can be compared to the classical friction model with discontinuity at the zero velocity (see Fig. 4b). The finite slope for small velocities introduces some important qualitative changes to the description of dry a) Weak Damping b) Classical Damping friction"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0003173_s0141-6359(02)00117-4-Figure16-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0003173_s0141-6359(02)00117-4-Figure16-1.png",
+        "caption": "Fig. 16. Finite element analysis of moment-carrying coupling made of three-dimensional printing.",
+        "texts": [
+            " Note that ProMECHANICATM automatically distributes a point load over a small circular area; this area is small relative to the size of the components, and does not significantly change our results. Initial models were done on single fingers or finger pairs to save computation time. After these models demonstrated the feasibility of the idea, models of entire couplings were constructed and analyzed. Sample results are shows in Figs. 15 and 16. Fig. 15 is a plot of displacement magnitude in an aluminum model of the perpendicular adjustment moment carrying variation; displacement is expressed in units of milli-inches. Fig. 16 shows strain energy for the three-dimensional printed perpendicular adjustment non-moment carrying variation. These figures show good strain and stress distribution indicating a robust design. Physical models were tested to validate analytical and finite element models of torsional stiffness. Fig. 17 shows a test fixture that was constructed to permit the testing of three-dimensional printed couplings, as well as an aluminum model fabricated specifically for the tests. Except where noted, all structural components were aluminum, and all fasteners and bearings were steel"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0001658_ie50389a020-Figure7-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0001658_ie50389a020-Figure7-1.png",
+        "caption": "FIGURE 7. POTENTIALS ON SURFACE OF TIRE RVNNING ON STEEL DRUM",
+        "texts": [
+            " on the indoor testing drum. A probe was made from a strong polonium-plated foil 3 cm. in diameter, bent into a cone whose base was 1 cm. in diameter. This probe was well insulated and was connected to a low-sensitivity electroscope whose case was grounded. When the probe was held about 1 cm. from the moving tire surface, the electroscope leaf attained a steady value within 1 minute. The reading was recorded after 1.5 minutes a t each position of the tire investigated. The results of this investigation are shown in Figure 7. Aside from the expected result that the tire tread was a t a high potential even though the rim was grounded, the most striking result was the large difference in tread potential which existed between the regions where the tread approached and where it left the drum contact areaa. Obviously, the electric charge was generated when the tire pulled away from the drum. The data shown in Figure 7 indicate that the charge generated a t the tread surface leaks over the side wall to the rim. It is probable that some charge escapes from the rim by leakage over the side wall to the drum contact area, and this would account for the low potential observed on a car with high-conductivity side wall tires. The probable direction of the currents in a tire running on a pavement are indicated in Figure 8. The following derivation of the equilibrium rim voltage is based on the above principles and, in addition, on the as- * The tread potential distribution shown in Figure 7 is in disagreement with the one presented by Cadwell, Handel, and Benson [News Ed. (Am. Chem. Soo.), 19, 1139, Fig. 3 (1941)l. In the discussion following their paper a t the Atlantic City meeting of the AM~RICAN CHEMICAL SOCIETY, W. F. Buase, who has given us permission to quote him, presented data on equilibrium tread potential distributions which also disagreed with that shown in Figure 2 of Cadwell et al. but agrees in general form with the distribution shown in our Figure 7. Furthermore, the data presented by Busse indicated that these tread potential distributions were not greatly affected by introducing into the inner tube either conducting liquids or conducting powders such as the \u201dstatic neutralizer\u201d discussed by Cadwell et al. As mentioned a t that time, these data, comprising measurements made over more than half of the tire circumference, agreed in general with the results of similar tests we had made previous to the meeting. Busse also reported that the introduction of a conducting liquid or conducting powder to the inner tubes of all four tires on a car had no appreciable effect on the equilibrium potential of the car when driven along the road or on the radio static generated by the discharge of this potential"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0000836_s11804-016-1370-x-Figure1-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0000836_s11804-016-1370-x-Figure1-1.png",
+        "caption": "Fig. 1 Earth-fixed and DP vessel-fixed coordinate frames",
+        "texts": [
+            " In Section 5, the proposed approach is illustrated using a practical example of filtering corrector synthesis, and finally, Section 6 concludes the paper by discussing overall results of the investigation. To consider the problems of DP automatic control, we accept the following widely used 3DOF nonlinear robot-like model of a DP vessel (Fossen, 1994; Hassani et al., 2012): ( ) ( ) t     \u039cv Dv \u03c4 d \u03b7 R \u03b7 v  (1) In the equations,  T u v r\u03bd is the generalized velocity vector defined in a vessel-fixed frame, vvvv zyxO ;  T x y\u03b7 is the joint vector relative to an Earth-fixed frame, Oxyz , that includes position  yx, parameters and the heading angle,  (Fig. 1). A displacement of x and a velocity of u determine the surge motion of the vessel; y and v determine the sway motion, and a pair of  , r is referred to as the yaw motion. The vector 3\u03c4 E implies a control action generated by the propulsion system, and vector 3d R reflects an external disturbance of any nature. In addition, the matrices TM M and D with constant elements, are positive definite. The orthogonal rotation matrix, cos sin 0 ( ) ( ) sin cos 0 0 0 1                 R R (2) determines the only nonlinearity of the system (1)"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0000709_012002-Figure2-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0000709_012002-Figure2-1.png",
+        "caption": "Figure 2. Planetary precessional multiplier with satellite gear mounted radially.",
+        "texts": [
+            " If necessary, this shortcoming can be eliminated using as a controlling mechanism the constant Cardan joint (Hooke\u2019s joint), the ball synchronous couplings, etc. This kinematical diagram of the precessional transmission ensures a range of gear ratios i = 8...60, but in the multiplication regime it operates efficiently only for the range of gear ratios i = 8\u202625. As well, in the controlling mechanism W, that operates with pitch angles of the semi couplings up to 3o, power losses occur reducing the efficiency of the multiplier on the whole. To avoid power losses in the multiplier and to widen the kinematical options (figure 2) [2] the conceptual diagram of the precessional multiplier with wide kinematical options was designed. The planetary precessional multiplier comprises the following units: the housing 1, inside which the fixed sun wheel 2 is placed and connected rigidly to the housing cover 3, exterior satellite wheel 4 with the teeth in the shape of rollers, movable sun wheel 5, linked rigidly to the input shaft 6. The satellite wheel 3 is connected kinematically with the sloped flange of the disk 7, connected rigidly with the sun wheel 8, that gears with the interior satellite wheel 9 mounted unbound on the output crank shaft 10, and linked rigidly to the rotor generator 11"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0000878_978-3-319-47614-8-Figure2.20-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0000878_978-3-319-47614-8-Figure2.20-1.png",
+        "caption": "Fig. 2.20 Construction of Mohr\u2019s circle for stresses on other planes",
+        "texts": [
+            " Recall that the normal stresses equal the principal stresses when the stress element is aligned with the principal directions, and the shear stress equals the maximum shear stress when the stress element is rotated 45\u00b0 away from the principal directions. Also remember from Eq. (2.18), the angle on Mohr\u2019s circle is 2\u03b8, see Fig. 2.19. As the stress element is rotated away from the principal (or maximum shear) directions, the normal and shear stress components will always lie on Mohr\u2019s Circle. The angle between the current axes (X and Y) and the principal axes is defined as \u03b8p, and is equal to one half of the angle between the line Lxy and the \u03c3axis as shown in Fig. 2.20; the procedure to obtain stresses on another plane is illustrated in Fig. 2.20 with the following steps. 1. Locate the two points (\u03c3x, \u03c4xy) and (\u03c3y, \u2212\u03c4xy). 2. Draw the line Lxy across the circle from (\u03c3x, \u03c4xy) to (\u03c3y, \u2212\u03c4xy). 3. Rotate the line Lxy by 2\u03b8 (twice as much as the angle between XY and X\u2032Y\u2032) and in the opposite direction of \u03b8. 4. The stresses in the new coordinates (\u03c3x\u2032\u03c3y\u2032 and \u03c4x\u2032y\u2032) are then read off the circle. 2.3 Equilibrium Relations 53 Worked Example 2.5 Show that the sum of the normal stresses acting on perpendicular faces for a plane stress element is constant"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0000860_978-3-319-44735-3_1-Figure1.6-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0000860_978-3-319-44735-3_1-Figure1.6-1.png",
+        "caption": "Fig. 1.6 The different forces applied to the adjacent rotors to rotate the hexacopter clockwise",
+        "texts": [],
+        "surrounding_texts": [
+            "The proposed controller implements a closed loop that comprises the three layers. Data produced as output in one layer is passed as input to the next layer. Fig. 1.9 With low force, the hexacopter goes down The proposed multi-layer fuzzy controller is based on [12] and is depicted in Fig. 1.10. The Control box is composed by a pre-processing phase (first layer), a set of fuzzy controllers (second layer), and post-processing phase (third layer). As one can observe, after the post-processing phase, the control outputs are applied onto the plant by means of the hexacopter rotors that actuate on the hexacopter movement and stabilization. The sensors perceive the changes on the plant controlled variables, and hence, provide the feedback to the controller. The controller, in turn, compares these input valueswith the reference values established as setpoints thereby closing the control loop [8]. The pre-processing phase (first layer) is responsible for acquiring data from the input sensors, process the input movement commands, as well as calculate the controlled data used as input to the fuzzy controllers in second layer. Before the multilayer controller starts its execution, there is an initialization phase that is performed within the first layer. The target position is set as the current position, so that the hexacopter does not move before receiving any command. Gyroscope and accelerometer sensors are calibrated and the GPS sensor is initialized by gathering at least four satellites. During the execution phase, the first layer is responsible to calculate the input variables to the fuzzy controllers: (i) the angular and linear distance (delta error) for X, Y, and Z axes between the current hexacopter position and the target position; (ii) the rotation and translation movement matrices to translate 3 axes movement into the speed related to the ground (i.e. X and Y axis). In addition, it is responsible to convert the input movement commands into setpoints for X, Y and Z positions. Movements commands are composed of three values representing the positive or negative movement along X, Y and Z axes related to the current positions, i.e. a command indicates a relative position. Thus, when a new command is received, the first layer will convert it to a absolute position. Then, when the control system is executing, this layer uses the GPS coordinates to determine the error in the distance from the hexacopter to the target position. These calculated errors in position are the inputs to the fuzzy controllers (Euler X, Euler Y and Euler Z errors). The second layer contains five fuzzy controllers, which act on issues regarding the hexacopter movement, namely hovering stabilization, vertical and horizontal movement and heading. As mentioned, these controllers take as input the data produced in the first layer and generate output for the third layer. The generated outputs represent the actuation on the six rotors for performing pitch, roll, yawmoves for all maneuvers necessary to reach the target position. The fuzzy controllers are discussed in details in the next section. The post-processing phase (third layer) is responsible for coordinating the fuzzy controllers outputs. As mentioned, in order to perform a proper maneuver, the proposed multi-layer controller establishes a priority on movements needed to complete a maneuver. When a new command is received, i.e. a new target point is set, the hexacopter must firstly reach the target altitude. Then, the hexacopter must turn until its front aims the target position. Finally, the hexacopter moves horizontally towards the target position. This layers also performs a threshold limits control by means of output values saturation, in order to keep the hexacopter stability while flying or hovering."
+        ]
+    },
+    {
+        "image_filename": "designv11_60_0000860_978-3-319-44735-3_1-Figure1.13-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0000860_978-3-319-44735-3_1-Figure1.13-1.png",
+        "caption": "Fig. 1.13 Input linguistic variables and their membership functions for the roll angle, Euler approximation error of X-axis angle",
+        "texts": [
+            " The next sections provides details on these five independent fuzzy controller. Roll is the movement obtained through the rotation around the X-axis, i.e. front-toback axis. The fuzzy controller named Roll Stabilization controls the stabilization of the hexacopter while it is performing the roll maneuver. Figure1.12 shows the block diagram of this controller. This controller has two input data. The first input is error in roll angle Euler approximation. The roll angle is calculated through the Euler approximation of the current X angle and the target X-axis angle. Figure1.13 shows the linguistic variable membership function representing the fuzzification of the error in the roll angle Euler approximation. The second input is the perceived movement in Y-axis represented as the acceleration in Y-axis obtained from the accelerometer over the time. Figure1.14 shows the linguistic variables and themembership functions representing the fuzzification of Y-axis acceleration. Its worth mentioning that the \u201cN\u201d and \u201cP\u201d prefixes of variables names stand for, respectively, Negative and Positive"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0000878_978-3-319-47614-8-Figure6.1-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0000878_978-3-319-47614-8-Figure6.1-1.png",
+        "caption": "Fig. 6.1 Technical theory of Euler beam",
+        "texts": [
+            " This is permissible when the cross-sectional dimensions of a structure are small compared to the length dimension. We also make assumptions on the deformation field of the structure, e.g., if we stretch such a one-dimensional structure, we have a bar; if we twist this structure by a pure moment, then we have a rod (shaft in pure torsion) and another most commonly used simple one-dimensional structure is a beam with a load in transverse or lateral direction. We will consider a beam in its simplest form, see Fig. 6.1, and derive a simple theory called technical beam theory from fundamentals. This will also be same as Euler-Bernoulli theory. We recognize the beam to be lying on the x-axis (only one dimension); the cross-sectional dimensions enter into the theory through the derivation. We will see these parameters entering the governing equation to be the area A and second moment of area I of the cross-section. Consider the cross-section at a distance x, this cross-section is in a plane normal to the x axis before bending and after bending, this plane is assumed to remain plane (no warping or distortions) and rotate in such a way that it still remains normal to the bent (elastic) axis. This is what Leonardo Da Vinci has observed long ago and firmed up by Euler and Bernoulli. By approximating the beam and forcing the cross-section at distance x, we are writing an approximate solution rather than solving the 15 equations of theory of elasticity derived before. This approximation is so good it gives us almost correct solution for slender beams when compared with experimental results. Through this approach we have an approximate but working theory. From Fig. 6.1 the displacement field of a point at coordinates x, y, z of the beam can be written as \u00a9 The Author(s) 2017 J.S. Rao, Simulation Based Engineering in Solid Mechanics, DOI 10.1007/978-3-319-47614-8_6 151 ux \u00bc z w;x uy \u00bc 0 uz \u00bc w\u00f0x; t\u00de \u00f06:1\u00de Notice that a comma followed by suffix x as before denotes derivative with respect to x (,x = d/dx). The beam axis has no strain at all and therefore the elastic axis is also a neutral axis. Once the displacement field is assumed we have violated the exact theory of elasticity; however this allows us to obtain a simple solution because we can obtain the strains directly from strain displacement relations"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0003038_robot.1995.525684-Figure1-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0003038_robot.1995.525684-Figure1-1.png",
+        "caption": "Fig. 1 A system of coupled links.",
+        "texts": [
+            " 2 Decoupled NOC Since the natural orthogonal complement (N0C)a matrix that relates the velocities of the links of a mechanical system to its joint rates, and used to reduce the Newton-Euler equations of motion of uncoupled links of the system to an independent set of EulerLagrange equations-in its present form, i.e., as obtained in [$, 91, is not suitable to write the elements of the GIM as expressions, it is derived here as two decozlpled matrices. This allows to perform the symbolic RGE, as done in $4. Now, let us define the 6- dimensional vector of the twist, or the spat ial velocity [4, 51, of the ith link, Fig. 1, as (1) ti E [UT, Ci T T ] where w; and C; are the 3-dimensional vectors of angular velocity and the velocity of the mass center of the ith link, Ci, respectively. Moreover, for the system of coupled links, Fig. 1, wi and Ci are written as wi = w j + Biei ci = c j + w j x rj + wi x di (2.) (2b) where w j and Cj are the angular velocity and the velocity of Cj of the j t h link, respectively, whereas the Cartesian vectors, rj and di, are shown in Fig. 1. Combining eqs.(2a) and (2b), the twist, ti, is expressed as a function of tj and i i , i.e., t i = Bijtj + piBi ( 3 ) where the 6 x 6 matrix, Bij , and the 6-dimensional vector, pi, are given by B . . = [ \"1 and pi z [ ei ] ( 4 ) * I - c i j 1 ei x di 1 and 0 being the 3 x 3 identity and zero matrices, respectively, which, henceforth, should be understood as of dimensions compatible to the size of a matrix where they appear. Moreover, Cij is the 3 x 3 crossproduct tensor, associated to the vector, cij = cj - ci",
+            " A 3 x 3 cross-product tensor, associated to the %dimensional vector, z , denoted by Z, is defined by Z ~ z X 1 ~ - (5) a ( Z x x) ax for any arbitrary %dimensional vector, x. Furthermore, ei is the unit vector parallel to the axis of the ith revolute pair. It is pointed out here that the matrix, Bij, and the vector, pi , have the following interpretations: while the former multiplied to tj gives t i , if there is no i th joint, the latter takes into account the effect of the ith joint motion. Also, from the definition of B Q , eq.(4), and Fig. 1, the following properties are derived: BijBj~, = Bik and Bii = 1 (6) Now, for the manipulator consisting of n links, as denoted in Fig. 2 by #1, . . 8 , #n, coupled by n revolute pairs, namely, 1, . . . , n , the 6n-dimensional generalized twist, t , is defined as t =. [tT, . . ' ) t y (7) where t i , for i = 1, \" * , n, is given in eq.(l). Using eq.(3) and the properties given by eq.(6), vector t is represented as t = TO, where T r TlTd (8) In eq.(8), the 6n x n matrix, T, is the natural orthogonal complement (NOG) of the system at hand, where TI and T d are the 6n x 6n lower block triangular and the 6n x n block diagonal matrices, respectively, i"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0003941_a:1013730210328-Figure6-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0003941_a:1013730210328-Figure6-1.png",
+        "caption": "Fig. 6. Dependence of reorganization energy on the redox probe location during the reduction process. For redox probe located inside the alkyl layer, the value of \u03bb decreases compared to the situation when the redox probe remains in the aqueous medium.",
+        "texts": [
+            " The derivative of kapp with respect to overpotential \u03b7, being a Gaussian function, allowed us to obtain the distribution of density of electronic states for the oxidized form (3) where represents electronic coupling between the redox species and the electrode surface, \u03bb is reorganization energy, and \u03c1 is density of electronic states. The best Gaussian fit of this curve and experimental data allowed us to obtain kinetic parameters for [24]. Reorganization energy of oxidized form of the studied system could be extracted either from the peak position or from the width of the derivative function. The reorganization energy may be considered as diagnostic parameter for determination of the position of the redox probe during the reduction process at the monolayer/solution interface (Fig. 6). Equation (4) relates the outer-sphere reorganization energy to the radius of ion a, distance between the redox center and the electrode r, and optical \u03b5op and static \u03b5s dielectric constants [61, 62]. It gives, therefore, indirect information about the permeability of the monolayer. The reor- ganization energy both for Fe(CN and for is dominated by outer-sphere contribution; therefore, its value is affected only by the dielectric properties of the solvent. Thus we can conclude that the anolyte ions are located at the outer plane of the monolayer and remain in aqueous medium during the reduction process"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0000026_cdc.2015.7402246-Figure2-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0000026_cdc.2015.7402246-Figure2-1.png",
+        "caption": "Fig. 2. Illustration of the model parameters. Particle i is in the velocity alignment state and j \u2208 Q (i) a . The velocity damper produces a force in any direction.",
+        "texts": [
+            " Let Q(i) s = {k| \u2016rki\u2016 \u2264 \u03c1s} denote the set of particles within the perceptual range \u03c1s > 0 of the ith particle, and Q (i) a = {k| \u2016vki\u2016 \u2264 \u03c1a, \u2016rki\u2016 \u2264 \u03c1s} denote the set of particles that are also within interaction range \u03c1a > 0 in the velocity space. We model each force term as follows: F (space) i = c \u2211 j\u2208Q(i) s (1\u2212 x0/\u2016rji\u2016) rji (5) F (align) i = b \u2211 j\u2208Q(i) a vji (6) F (ext) i = \u2212dvi + wi, (7) where wi represents random noise, and c, x0, b, and d are the spring, rest length, damping, and drag constants, respectively. Figure 2 illustrates the model parameters. In order to generate oscillatory motion, the spacing-force connects interacting particles as opposed to connecting each particle to a fixed point, as was considered previously [12] (see (2)). The previous model was valid for swarms that form above a fixed marker on the ground (this behavior is known to occur for only one of the two anopheline genetic types [10]). The new model accommodates swarming above a fixed point by adding a fixed, virtual particle. Noting that \u2016rji\u2016 < x0 results in repulsion and \u2016rji\u2016 > x0 in attraction, F(space) is a dynamical analogue of existing models with attraction and repulsion zones [1]"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0001817_0141-0296(93)90032-y-Figure7-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0001817_0141-0296(93)90032-y-Figure7-1.png",
+        "caption": "Figure 7 Sect ion t h rough typ ica l mas t f oo t i ng",
+        "texts": [
+            " However, they must be able to rotate freely about the footing (pin) to eliminate torque from the masted part of the structure. The angle of inclination of the masts of the guyed portal or the CRS or any of the guyed rigid towers is a fixed angle although the angle of the masts of the guyed V will change with the height. The angle of the masts of a guyed V set of towers will vary about -4-30-4 \u00b0 . Quite large ball and socket castings of aluminum or steel that permit large movements have been used as well as some very simple fittings, as shown in Figure 7. This detail consists of a steel pin, embedded in concrete or directly into rock and set at an appropriate angle, a grout pad as needed, a steel plate to distribute the bearing pressure and a spherical washer to distribute the bearing load from the base of the mast to the washer. The bearing pressure is on a ring of contact and is infinite until the base plate of the mast deforms slightly as the tower is put in place and the guys tensioned. This deformation is then accepted as a permanent and nonthreatening condition and, as long as it is restricted to a few millimetres at most, is negligible compared to the elasticity of the guy system"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0000188_0954406214560420-Figure1-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0000188_0954406214560420-Figure1-1.png",
+        "caption": "Figure 1. Diagram of 6-UPS parallel mechanism.",
+        "texts": [
+            " An equivalent method is proposed to remove the extremely displacement singularity in the limbs in Equivalent model of the limb with extremely displacement singularity section. The FK solutions of two representative 6 DoF mechanisms are given next. FK problem of coupled fewer DoF mechanisms section analyzes a coupled fewer DoF mechanism, and puts forward two methods to solve its FK problem. In the last section, numerical examples are given to validate the theories proposed above. A typical 6-UPS mechanism is taken as an example to illustrate the FK problem of parallel mechanisms. The diagram of 6-UPS mechanism is shown in Figure 1. Coordinate Systems o and p are attached to the fixed platform and mobile platform, respectively. bi0 (i \u00bc 1, 2, . . . , 6) is the coordinate of the upper hinge points in coordinate system p. ai is coordinate of the lower hinge points in coordinate system o. The IK model can be expressed as li \u00bc Ropbi0 \u00fe pop ai \u00f01\u00de Here li represents the length of the linear actuator. Rop is the rotation matrix of coordinate system p with respect to o, and pop is position vector of that. q \u00bc \u00bdl1 l2 . . . l6 T is the generalized coordinates of the parallel mechanism"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0002264_s0039-9140(97)00049-0-Figure7-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0002264_s0039-9140(97)00049-0-Figure7-1.png",
+        "caption": "Fig. 7. Square wave voltammograms of abscisic acid reduction at different concentrations: r=residual current; (a) 75 ng ml\u22121; (b) 125 ng ml\u22121; (c) 150 ng ml\u22121; (d) 250 ng ml\u22121; (e) 300 ng ml\u22121; (f) 350 ng ml\u22121; (g) 400 ng ml\u22121; (h) 450 ng ml\u22121 and (i) 500 ng ml\u22121.",
+        "texts": [
+            " An accumulation potential Eac= \u22120.4 V was applied for 10 s, with an equilibrium time of 5 s. The initial scan potential coincided with Eac, and the final potential was set at \u22121.2 V to permit the full development of the wave. Instrumental variables during the measurement stage were set at, pulse amplitude a=40 mV, scan increment E=5 mV, and square wave frequency f=120 Hz, producing a scan rate of 600 mV s\u22121. There is an increase in the peak intensity with the concentration, shown in the superimposed voltammograms in Fig. 7. This dependence of the peak intensity on concentration follows a linear relationship, which fits the expression ip= \u22121.76\u00d710\u22122+7.89\u00d710\u22124C ; r=0.999 The intensity of the peak is expressed in mA while the concentration is in ng ml\u22121. The relative error of the method ranges between 0.3 and 0.8% in absolute values and the relative standard deviation varies between 2.1 and 7.8%, producing a detection threshold of 30 ng ml\u22121 and a quantification threshold of 58 ng ml\u22121. After optimising the abscisic acid determination method, it was applied to an extract of pears taken straight from the tree at the initial growth stage"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0000160_6.2015-0904-Figure4-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0000160_6.2015-0904-Figure4-1.png",
+        "caption": "Figure 4. System overview in Matlab/Simulink.",
+        "texts": [
+            " The modeling of this system was done in two major steps: in the first step, the modeling of the mechanical system using simMechanics in Simulink was made by importation of the CAD model and importation of the mass and inertia for each mechanical part. The parts are linked together using simMechanics joint blocks. The second step is the modeling of the electrical and power systems for the static and dynamic model in terms of tension and current. The simMechanics modeling provides interesting features such as the import and visualization of the CAD model from 3D design software such as CATIA v5. An overview of the 3D mechanical system can be found in Fig. 4; this interface allows the visual control of the behavior of the system during simulation. After the importation of geometrical parts, mass and inertia data for each part are implemented in the Simulink model. A model of the external forces is also added; the external forces are the weight of system components and the D ow nl oa de d by C A R L E T O N U N IV E R SI T Y L IB R A R Y o n Ju ly 3 1, 2 01 5 | h ttp :// ar c. ai aa .o rg | D O I: 1 0. 25 14 /6 .2 01 5- 09 04 American Institute of Aeronautics and Astronautics 7 aerodynamic forces"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0000739_978-1-4471-4976-7_91-1-Figure5-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0000739_978-1-4471-4976-7_91-1-Figure5-1.png",
+        "caption": "Fig. 5 A general SPM and the coordinate system, where only one leg is shown for clarity",
+        "texts": [
+            " Let the payload at the end-effector of the robot be 5 kg, which implies f6 \u00bc 5gN, n6 \u00bc 0, where g is the vector of gravity acceleration. Through the modeling of inverse dynamics of Eqs. 32g\u201332i, the joint torques are solved, as shown in Fig. 4. Example II: A Spherical Parallel Manipulator In this example, the Lagrange\u2019s equation is illustrated with a spherical parallel manipulator (SPM) (Gosselin and Angeles 1989). The SPMs are of closed-kinematic chain. The Lagrange\u2019s equation of the 1st kind is adopted for the dynamics modeling of this manipulator. A general SPM is shown in Fig. 5a with the parameterized ith leg in Fig. 5b. The ith limb consists of three revolute joints, whose axes are parallel to the unit vectors ui, vi, andwi. All three limbs have Page 8 of 17 identical architectures, defined by angles a1 and a2. Moreover, b and g define the geometry of two triangular pyramids on the base and the mobile platforms, respectively. The origin O of the base coordinate system xyz is located at point O. The z axis is normal to the bottom surface of the base pyramid and points upward, while the y axis is located in the plane made by the z axis and u1"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0001493_pi-c.1956.0019-Figure4-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0001493_pi-c.1956.0019-Figure4-1.png",
+        "caption": "Fig. 4.\u2014Energy fluxes in long rectangular circuit. (a) Energy streams and electric field components.",
+        "texts": [
+            " (32) in terms of energy momentum, we distinguish the components of electric field arising from the positive and negative charges in the conductors. We shall simplify the problem by assuming that the parallel wires have the same uniform section and negligible resistance, so that the electric field between them, except near the ends, is perpendicular to the current flow and the potential difference between them is constant. The end wire then provides a resistance \"load\" on the long transmission line. Fig. 4(a) shows a portion of the circuit remote from the ends, and Fig. A{b) is a cross-sectional view. We consider each conductor to contain the same quantity of stationary positive charge in its structure per unit length. The current in conductor 1 is provided by the motion, with a mean velocity vu of conduction electrons whose charge per unit length is />,, and the current in conductor 2 is provided by the motion, with mean velocity v2, of conduction electrons whose charge per unit length is p2. The direction of motion is, of course, opposite to that of the current, and />",
+            " Then at any point P we recognize four components of electric field, all perpendicular to the wires, namely E\\ from the conduction electrons in conductor 1. E2 from the conduction electrons in conductor 2. \u00a33 from the stationary positive charge in conductor 1. \u00a34 from the stationary positive charge in conductor 2. CULLWICK: ELECTROMAGNETIC MOMENTUM AND ELECTRON INERTIA IN A CURRENT CIRCUIT 163 and this is deflected at the end wire at the rate (6) Electric and magnetic field components. These components are in the direction shown in Fig. 4(a). The components of the magnetic field intensity, arising from the motion of the sources of Ev and E2, are shown in Fig. 4(6). The positive charges, being stationary, carry no energy fluxes. The conduction electrons in conductor 1 carry an energy flux with velocity v\\'. Sx = Exx (Hi + H\u00a3 or Sx = \\E{ x Hi\\ + \\E{ x H2\\ = Ey(Hx + H2cos 0) (33) directed towards the end DA. TT But E, = \u2014-, so that (H2 + HXH2 cos 0)_ v \" l . (34) Since the velocity of this energy flux is vu its volume density is SJvx and its mass per unit volume is m, = v{c 2 (35) Its kinetic energy is therefore \\m<o\\ = f^(H2 + HxH2 cos 0) . . . (36) Similarly, the conduction electrons in conductor 2 carry an energy flux with velocity v2: Hj + H^cose \u20acQV2 away from the end DA, and its kinetic energy is - ^ ( H 2 + H x H z c o s 8 ) "
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0001976_s002490050241-Figure2-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0001976_s002490050241-Figure2-1.png",
+        "caption": "Fig. 2 Calculated time courses of total membrane current, I, and capacitive current, Ic (Cm = 2 nF) of the model system (Fig. 1) at 10 mM cK,l and 400 mM cK,c, for an individual saw-tooth cycle of the command voltage (=Vm) over a voltage range from V1 = )200 mV to V2 = +50 mV, consisting of three linear slopes with dV/ dt = 0.1 V s)1 between control periods (here 200 ms) of Vm = V0 = const. = )80 mV",
+        "texts": [
+            " In our assumed case of activation by negative voltages via a symmetric Eyring barrier, the transitions between a and i occur with the rate constants ka for activation and ki for inactivation: ka k0ae \u00ffu=2 2a ki k0i e u=2 2b where the superscript 0 marks the rate constant at zero voltage. For our model system we choose k0a = 1 s)1 and k0i = 10 s)1. If pa and pi are the probabilities of the pathway to be active and inactive, respectively, these probabilities follow the di erential equations dpa dt \u00ffkipa kapi 3a dpi dt kipa \u00ff kapi 3b The total current through the membrane is then I CmdV =dt paIGHK 4 Figure 2 shows the behavior of the model system in Fig. 1 under voltage clamp with a saw-tooth shaped command voltage. The suggested voltage protocol starts with a short period at a steady state control voltage, V0, which may be close to the resting or equilibrium voltage of the system. Then the voltage moves with a constant slope to a minimum voltage, V1, where the slope changes its sign and the voltage moves to a maximum voltage, V2. Here, the slope reverses again, and the voltage moves back to V0, where it is kept constant for another short time interval. The time course of the clamp current I through the model system upon such a saw-tooth shaped command voltage with a slope of 0.1 V s)1 is shown in Fig. 2 as well. This I(t) tracing exhibits non-linear features which contain all the properties of the model system. In such diagrams, the capacitive current IC = Cm dV/dt can readily be identi\u00aeed. Although it can hardly be seen in the I(t) tracing in Fig. 2, its discontinuous change by 2|dV/dt|Cm = 0.4 nA becomes obvious when the slope changes its sign at V1 and at V2. IC is also plotted in Fig. 2 in an expanded I scale. The V(t) and I(t) data from Fig. 2 are converted to the current-voltage diagram, I(V), in Fig. 3A. Knowing the general shape of the saw-tooth voltage-protocol and |dV/dt| ( 0.1 V s)1 in this case), Fig. 3A provides in principle the complete information from Fig. 2 which can be used to identify the four free parameters (Cm, g0, k0a, and k0i ) of the model system in Fig. 1. In contrast to Fig. 3B and C, the particular |dV/dt| in Fig. 3A was too small to reveal the \u00aerst parameter, Cm, by the capacitive currents. The second parameter g0 = g(V > 0)/cK,l, can be determined by g(V > 0), i.e. the asymptotic straight I/V line (not drawn) through the origin which is approached by the left portion of the curve. The remaining parameters k0a and k0i can be determined by the right part of the graph, e"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0003765_tmag.2002.802692-Figure1-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0003765_tmag.2002.802692-Figure1-1.png",
+        "caption": "Fig. 1. Geometry of solid surfaces and meniscus bridges.",
+        "texts": [
+            " Note that up-and-down vibration of an upper surface is considered as a disturbance from undulation of a disk surface for the sake of simplicity though the disk surface slides. Manuscript received February 14, 2002; revised June 27, 2002. The authors are with the Department of Applied Mathematics and Physics, Faculty of Engineering, Tottori University, Tottori 680-8552, Japan (e-mail: hiro@damp.tottori-u.ac.jp; fukui@damp.tottori-u.ac.jp; b9739@svr01.damp.tottori-u.ac.jp). Digital Object Identifier 10.1109/TMAG.2002.802692. The analytical models and geometry of meniscus and solid surfaces considered in this study are shown in Fig. 1. The infinite-width meniscus [Fig. 1(a)] and the finite meniscus ring [Fig. 1(b)] between flat surfaces are analyzed. The authors adopt the following assumptions: 1) the liquid is incompressible; 2) liquid meniscus is always in equilibrium thermally; 3) the contact angle is always zero because of the smallness of the amplitude of vibration [assumption 4)], in other words, the solid surface is perfectly wettable and there are not the receding and advancing contact angles; 4) the vibration amplitude is small compared with average spacing; 5) the velocity of the small vibration is also small; 6) the solid surfaces are perfectly smooth; 7) the shape of the cross section of the meniscus is always a part of circle; 8) the mass of the liquid in the meniscus is conserved; 9) and (see Fig. 1) are very small. Reynolds equation for incompressible liquid is given by [6] (1) 0018-9464/02$17.00 \u00a9 2002 IEEE where is the pressure generated in the meniscus bridge, is the spacing and is the viscosity of the liquid. At the meniscus boundary, the difference between the liquid pressure and the atmospheric pressure is given by the following Laplace equation [7]: (2) where is the surface tension of the liquid and and are shown in Fig. 2. Note that denotes the radius of curvature of ring for the finite meniscus ring [see Fig. 1(b)] and in case of infinite-width meniscus, becomes infinity. III. INFINITE-WIDTH MENISCUS In the case of the infinite-width meniscus between parallel plane surfaces [see Fig. 1(a)], (1) is rewritten in the following form: (3) The spacing is given by (4) where is the average spacing. From the assumptions of small amplitude and small velocity, 4) and 5), and (5) Solving (3), the pressure is given by (6) where and are constants, and is neglected since it is a second-order term. When the upper surface vibrates, the meniscus boundary varies as shown in Fig. 3. The radius of curvature of the meniscus is given by (7) From the mass conservation of the liquid, the following equation is obtained: (8) From (8), the meniscus boundary is (9) where is the meniscus boundary at the average spacing and is neglected as a second order term",
+            " Integrating the pressure, the load-carrying capacity is given by (11) Equation (11) also includes the squeeze term, the spring term and the static meniscus force by the static Laplace pressure. Examples of the pressure distribution and the load carrying capacity per unit length are shown in Figs. 4 and 5, respectively. The parameters are 100 Hz, 5 mm, 0.5 m, 75 mN/m, 1 mPa s, 1 10 . The pressure and the load-carrying capacity has negative values in this case because the attractive Laplace pressure is larger than the squeeze pressure. In the case of the finite meniscus ring between plane surfaces [see Fig. 1(b)], (1) is rewritten in the following form: (12) Through the same processes as Section III, the following important equations are obtained: (13) (14) (15) The boundary conditions are a) 0 at 0, i.e., the pressure is axisymmetric with respect to axis, b) the pressure at the boundary is given by (16) from (2). Equations (13)\u2013(15) correspond to (9)\u2013(11), respectively. Equation (13) means that the meniscus boundary varies at the ratio of when the spacing varies at the ratio of . Note that it was in the case of the infinite-width meniscus"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0000783_s11071-016-3104-7-Figure1-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0000783_s11071-016-3104-7-Figure1-1.png",
+        "caption": "Fig. 1 Absolute nodal coordinate system",
+        "texts": [
+            " To do this, non-dimensional variables related to the time, length, and force were used. The three-dimensional(3-D) nondimensional model developed in this study was based on the 3-D model developed by Garcia-Vallejo et al. [22]. For the developed non-dimensional model, the efficiency of the non-dimensional EOM was verified by comparing the CPU time according to the number of elements through examples of a cantilever beam and a free-falling pendulumwith revolute and spherical joints. As shown inFig. 1, an absolute nodal coordinate system is expressed with the global position vector r defined in a global coordinate system and the position vector gradients \u2202 r/\u2202x , \u2202 r/\u2202y, and \u2202 r/\u2202z. As shown in Fig. 1, this comprises a fully parameterized element because the position vector gradients in the y- and zaxis directions (i.e., \u2202 r/\u2202y and \u2202 r/\u2202z) are included. Consequently, the physical deformation of the cross section and the deformation in the length direction are considered. As previously noted, because the position vector gradients in the y- and z-axis directions are considered, the analysis time may be increased compared to the gradient-deficient element. As discussed in Sect. 1, the shear locking phenomenon could occur [11,23]"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0003689_robot.2001.932951-Figure3-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0003689_robot.2001.932951-Figure3-1.png",
+        "caption": "Fig. 3. Geometrical interpretation of the grasp feasibility inequality",
+        "texts": [
+            "3 Geometrical interpretation The proposed grasp feasibility inequality handles the frictional force space directly, instead of the contact force space which is dealt with by conventional grasp stability analysis. The meaning of grasp feasibility inequality can be explained as follows: the inequality consists of the decomposed friction cone and the balance-out property. First, note that the decomposed friction cone induces a convex friction polyhedron in the frictional force space of E for a given joint torque r, as shown in Fig. 3. The figure also shows that the bahnce-out property generates a hyperplane in the frictional force sJace, parameterized by the indeterminate frictional Eorce 6, for a given object wrench (and joint torque). Then if the convex friction polyhedron and the hyper-plane have some intersection, then the current grasp is stable even if there exists indeterminate forces. Otherwise, we can say that the grasp is unstable. Note that in case of determinate gra.sp, the hyper-plane reduces B single point. Therefore, the stability can be checked by considering whether the point belongs to the friction polyhedron or not"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0000602_s11771-015-2876-0-Figure4-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0000602_s11771-015-2876-0-Figure4-1.png",
+        "caption": "Fig. 4 Pressure distribution and streamline inside bearing cavity",
+        "texts": [
+            " The temperature sensor is mounted on the bearing inner ring to indicate that the system has achieved a thermal equilibrium state before testing. All the sensors are connected with Bruel & Kjaer (BK) data acquisition system with computer for further analysis. The specific technical parameters of the experimental apparatus are listed in Table 3. Theoretically, the air phase flow pattern inside the bearing cavity is determined by the combination of the geometry features and the motion characteristics. Under the rotating reference frame, the fluid domain flow features are solved with a steady solver. Figure 4 shows the air phase pressure distribution and air flow stream line inside the bearing cavity at the inner ring rotating speed of 104 r/min. It seems that the air flow direction coincides with shaft rotating direction. Since the steel ball rotates at a high speed compared with the stagnant outer ring, a negative pressure zone appears in the contact zone between the ball and the outer ring. Around the negative pressure zone, a high pressure zone is formed, and the pressure is higher at the zone along the direction of the rotation"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0001235_b978-0-08-099425-3.00005-4-Figure5.44-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0001235_b978-0-08-099425-3.00005-4-Figure5.44-1.png",
+        "caption": "FIGURE 5.44",
+        "texts": [],
+        "surrounding_texts": [
+            "In order to obtain the data needed for the tyre modelling required for simulation, a series of tests may be carried out using tyre test facilities, typical examples being the machines that are illustrated in Figures 5.42 and 5.43. The following is typical of tests performed (Blundell, 2000a) to obtain the tyre data that supports the baseline vehicle used throughout this text. The measurements of forces and moments were taken using the SAE coordinate system for the following configurations: 1. Varying the vertical load in the tyre 200, 400, 600, 800 kg. 2. For each increment of vertical load the camber angle is varied from 10 to 10 with measurements taken at 2 intervals. During this test the slip angle is fixed at 0 . 3. For each increment of vertical load the slip angle is varied from 10 to 10 with measurements taken at 2 intervals. During this test the camber angle is fixed at 0 . 4. For each increment of vertical load the slip and camber angle are fixed at zero degrees and the tyre is gradually braked from the free rolling state to a fully locked skidding tyre. Measurements were taken at increments in slip ratio of 0.1. The test programme outlined here can be considered a starting point in the process of obtaining tyre data to support a simulation exercise. In practice obtaining all the data required to describe the full range of tyre behaviour discussed in the preceding sections will be extremely time consuming and expensive. The test programme described here does not, for example, consider effects such as varying the speed of the test machine, changes in tyre pressure or wear, changes in road texture and surface contamination by water or ice. The testing is also steady state and does not consider the transient state during transition from one orientation to another. High Speed Dynamics Machine for tyre testing formerly at Dunlop Tyres Ltd. Most importantly the tests do not consider the complete range of combinations that can occur in the tyre. The longitudinal force testing described is limited by only considering the generation of braking force. To obtain a complete map of tyre behaviour it would also, for example, be necessary to test not only for variations in slip angle at zero degrees of camber angle but to repeat the slip angle variations at selected camber angles. For comprehensive slip behaviour it would be necessary at each slip angle to brake or drive the tyre from a free rolling state to one that approaches the friction limit, hence deriving the \u2018friction circle\u2019 for the tyre. Extending a tyre test programme in this way may be necessary to generate a full set of parameters for a sophisticated tyre model but will significantly add to the cost of testing. Obtaining data requires the tyre to be set up at each load, angle or slip ratio and running in steady state conditions before the required forces and moments can be measured. By way of example the basic test programme described here required measurements to be taken for the tyre in 132 configurations. Extending this, using the same pattern of increments and adding driving force, to consider combinations of slip angle with camber or slip ratio would extend the testing to 1452 configurations. In practice this could be reduced by judicious selection of test configurations but it should be noted the tests would still be for a tyre at constant pressure and constant speed on a given test surface. Examples of test results for a wider range of tyres and settings can be obtained by general reference to the tyre-specific Flat Bed Tyre Test machine. (Courtesy of Calspan.) publications quoted in this chapter and in particular to the textbook by Pacejka (2012). For the tyre tests described here the following is typical of the series of plots that would be produced in order to assess the force and moment characteristics. The results are presented in the following Figures 5.44e5.53 where a carpet plot format is used for the lateral force and aligning moment results: 1. Lateral force Fy with slip angle a 2. Aligning moment Mz with slip angle a 3. Lateral force Fy with aligning moment Mz (Gough Plot) 4. Cornering stiffness with load 5. Aligning stiffness with load 6. Lateral force Fy with camber angle g Lateral force Fy with slip angle a. (Courtesy of Dunlop Tyres Ltd.) 7. Aligning moment Mz with camber angle g 8. Camber stiffness with load 9. Aligning camber stiffness with load 10. Braking force with slip ratio Aligning moment Mz with slip angle a. (Courtesy of Dunlop Tyres Ltd.) Lateral force Fy with aligning moment Mz (Gough Plot). (Courtesy of Dunlop Tyres Ltd.) Before continuing with the treatment of tyre modelling, readers should note the findings (van Oosten et al., 1999) of the TYDEX Workgroup. In this study a comparison of tyre cornering stiffness for a tyre tested on a range of comparable tyre test machines gave differences between minimum and maximum measured values of up to 46%. Given the complexities of the tyre models that are described in the following section the starting point should be a set of measured data that can be used with confidence to form the basis of a tyre model. Cornering stiffness with load. (Courtesy of Dunlop Tyres Ltd.)"
+        ]
+    },
+    {
+        "image_filename": "designv11_60_0001197_978-3-658-12701-5-Figure6.11-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0001197_978-3-658-12701-5-Figure6.11-1.png",
+        "caption": "Figure 6.11: Planar robot consisting of three subsystems",
+        "texts": [],
+        "surrounding_texts": [
+            "F3 = (\u2202 \u0307y3 \u2202s\u0307 ) \u22ba = \u239b\u239c\u239c\u239c\u239c\u239c \u239d \ud835\udc591 sin (\ud835\udc5e2 + \ud835\udc5e3) + \ud835\udc592 sin (\ud835\udc5e3) \ud835\udc591 cos (\ud835\udc5e2 + \ud835\udc5e3) + \ud835\udc592 cos (\ud835\udc5e3) 0 0 0 1 0 \ud835\udc592 sin (\ud835\udc5e3) \ud835\udc592 cos (\ud835\udc5e3) 0 0 0 1 0 0 0 0 0 0 0 1 \u239e\u239f\u239f\u239f\u239f\u239f \u23a0 \u22ba ."
+        ]
+    },
+    {
+        "image_filename": "designv11_60_0000050_iemdc.2015.7409124-Figure16-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0000050_iemdc.2015.7409124-Figure16-1.png",
+        "caption": "Fig. 16. Method of measurement of torque of investi",
+        "texts": [
+            " All een done using es both the motor cted, for this type iency is very low t 12 V and 4% at speed at constant DC speed at constant DC voltage. dustrial tests: (a) general view; (b) rake. V. PORTABLE LABORATORY EQ A portable equipment for fast measur state and dynamic performance of twomotors for computer fans has been built research (Figs 15, 16 and 17). This equ detection of faulty motors and selection o in computer assembly plants. As a load (brake), a 2.5-W, 12 V equipped with external bearing has bee torque has been measured with the aid of scale (Fig. 16). The rotational speed has been measu that operates as a counter. The timer generated by an optical sensor and a m placed on the rotor of the investigated m pulses have been red in 1-s intervals. VI. CONCLUSIONS Although a two-phase fan PM BLDC rotor is one of the most popular motor papers on analysis and performance chara motors have been published so far. The au the presented paper is the first pape extended laboratory tests and their analytical prediction. In this paper: (a) the magnetic circuit of the motor using the 2D FEM; (b) full laboratory steady-state and tr been performed; (c) experimental test results have be analytical calculations \u2013 good agreement h (d) a portable equipment for fast measur state and dynamic performance of twomotor for computer fans has been designe UIPMENT ements of steadyphase PM BLDC as a part of this ipment allows for f the best motors gated motor"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0003995_a:1016795221756-Figure3-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0003995_a:1016795221756-Figure3-1.png",
+        "caption": "Fig. 3. Polarization curves for the oxygen electroreduction in 0.5 M H2SO4 on RDE with active layers of (1\u20133) CPM, (1'\u20133') CoPP + FP-4D, and (1''\u20133'') acetylene black + Nafion, obtained at \u03c9 of (1, 1', 1'') 78.3, (2, 2', 2'') 171.6, and",
+        "texts": [
+            " In the case of acetylene black without the catalyst but with Nafion, only a slope b 2 of 0.150 V is observed. This is connected with a low activity of this carbon support, as a result of which at potentials, where b 1 is observed in the case of CPM, oxygen electroreduction on carbon black does not proceed yet. For electrodes with active layers of approximately identical thickness but with different compositions of the active mass, polarization curves for the oxygen electroreduction are presented in Fig. 3 for different \u03c9 . O 2 To reference electrode 1 2 3 4 5 6 7 8 9 8 Fig. 1. The cell for taking measurements on modified floating electrode of type 2: ( 1 ) Teflon housing, ( 2 , 5 ) rubber gasket, ( 3 ) graphite current lead, ( 4 ) working electrode, ( 6 ) auxiliary glassy carbon electrode, ( 7 ) lid, ( 8 ) tightening pins, and ( 9 ) electrolyte. 846 RUSSIAN JOURNAL OF ELECTROCHEMISTRY Vol. 37 No. 8 2001 TYURIN et al . Curves 1 ''\u2013 3 '' were obtained on the electrode that was covered by a thin layer of acetylene black mixed with a Nafion solution in the same conditions (ratio acetylene black : Nafion, ultrasound treatment, volume, drying mode) as CPM (curves 1 \u2013 3 ). Curves 1 '\u2013 3 ' refer to CoPP that contained a 2-% solution of polytetrafluoroethylene FP-4D as a binder and did not contain a Nafion solution. No pronounced plateau of a limiting diffusion current was observed in all cases. Dependence of cur- rents at potentials of 0.3 V for CPM (curves 1, 3) and 0.1 V for carbon black without the catalyst (curve 2) on is shown in the inset in Fig. 3 (points). Solid lines correspond to theoretical dependences for 2- and 4- electron reactions of the oxygen electroreduction. For CPM and acetylene black with Nafion, values of currents somewhat exceed values calculated for 4- and 2- electron reactions, respectively. It is worth noting that, for CPM, the deviation from calculated values of limiting current increases with \u03c9. For the catalyst mixed with FP-4D, at \u03c9 > 2000 rpm, the experimental current is below the calculated value. Similar phenomenon was observed in [24] for a composite material pyrolyzed mixture (activated carbon + CoTAA) + Nafion at \u03c9 > 3000 rpm"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0003012_iecon.1994.397742-Figure2-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0003012_iecon.1994.397742-Figure2-1.png",
+        "caption": "Figure 2: Analytical model of brushless dc motor.",
+        "texts": [
+            " For these current references, the current regulator calculates the applied voltage v based. on the inverse model of the motor. The estimations of the actual position 8, speed 0 and emf constant I fE are performed as follows. The controller calculates the model current i~ with the estimated position OM, speed 8, and emf constant KEM. Since a current difference Ai between the actual current i and model current i~ is caused by the estimation errors of the position, speed and emf constant, the estimation can be achieved by the current difference Ai. Brushless DC Motor M o d e l Fig.2 shows an analytical model of a brushless dc motor. The three phase voltage equation is given as follows. where, R,, L,; armature winding resistance and inductance, 1f3; emf constant, i,, i,, i,; line currents, vu, vu, U,; phase voltages and p is a short notation of d l d t , respectively. In Fig.2, d - q axis is the actual rotor position and y - 6 axis is the estimated rotor position. The transformation matrix from the three-phase coordinate to 7 - 6 coordinate is given in Eq.(2). Transforming Eq.(l) by using Eq.(2), the brushless dc motor model on y - 6 axis is given in Eq.(3) and the generated torque is given in Eq.(4). . d8 d8.w A 0 = 0 - 8 ~ , 8 = - , % M = - dt dt T = KTiCcos A0 Under the following assumptions; (4) A8 2: 0 pz 21 . i (n + 1) - i (n) T the voltage equation of the actual motor between the sampling points n and n + 1 is given as follows: where, T is a sampling period"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0003801_robot.1986.1087467-Figure2-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0003801_robot.1986.1087467-Figure2-1.png",
+        "caption": "Figure 2. The effect of intersecting the cones arising from two views of an object.",
+        "texts": [
+            " I t also builds up consistent descriptions of unexpected objects allowing them to be manipulated as single objects. 2.3. Updating the Octree In parallel with the above processing is a second process that operates to update the octree representation. It works directly from images of the world, taken from known positions. The 2D silhouettes of objects in the images are projected into the world octree as cones. When an object is seen from several viewpoints, the intersections of the cones constr;rin the position and size of the object (Figure 2). After several views, the representation begins to resemble the true shape of the object. At all times, the representation of an object occupies a volume a t least as large as the object. This is important because the octree is used for path planning as well as spatial representation 171, so the free space represented must be guaranteed to be empty. The algorithm used for the projection is described in [9]. An example of its use is shown in Figure 3. The two paths to constructing the spatial representation complement each other"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0002728_112515.112577-Figure15-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0002728_112515.112577-Figure15-1.png",
+        "caption": "Fig. 15 Motion wnstraint to realize agatitj contact.",
+        "texts": [],
+        "surrounding_texts": [
+            "(5)Fim2: Constraints forfirw arerepeatedly given for two\nCVS, Suppose CFJ and CF2 represent CFS at two CVS. Let tll\u2019,lEl\u2019 and ACI are transformations for CFI and tn\u2019, tm\u2019and Ac2 are for CF2. If the direction of the X axis of\nCF2 in the world is in the negative direction of the X axis of CFI, m\u2018 and tm\u2019 are changed to M * tn\u2019 and M * tE2 \u2018 . ACI = A[&,O,&, &I, C5yl,6ZI] and Ac2 = A [&,0,42,&,&2,&2] where dil = h, & = &, 61 = &2, and & = - &. If the angle between the Y axes of CFJ and CF2 is q, the following constraints are additionally given (see also Fig. 13):\ne= ~,d=21r-r\u2019lc0s6\n(28)\nwhere lx is signed distance from the origin of CF2 to the origin of CFI along the X axis of CF1.\n(6)Fim: Ac = A[&O,d,&,O,O] (see Fig. 14).\n5.2 Inequality Constraints When CVS are touching the extended feature, other vertices of the internal feature must be locate above the extended feature to avoid the interference (see Fig. 3). This condition is interpreted as inequalities.\nSuppose a vertex V of the internal feature and FI = origin({V ) ). FE is determined in the same manner for the equation constraints. If we consider the VF of V as defined in 3.3, we can represent tc of (23) by\ntC = tVE-l * tVI (29)\nwhere tw and tvE express the position of the axes of the VF with respect to the frames of FI and FE.\nWe approximate tvI and tvE by using the form features with nominal geometry and use tvl\u2019 and WE\u2019 instead. Now (29) can be approximated by\ntc = tVE\u2019-l * (Z + Av) * tVI\u2019 (30)\nwhere Av is a differential motion matrix at the VF.\n,, , ,x,,\n. XJ Ft\nCF\nvAc = A[d.,O&&,O,O]\n6. SYNTHESIS OF CONTACT COMBINATIONS A part has several contacts with other parts in an assembly. In this section, we give a method to synthesize all possible combinations of such contacts based on the amlysis of the part motion.\n6.1 Assumptions for Contact Synthesis The following three constraints are assumed for synthesizing the contact combination.\n(l)A part is assumed to be positioned in stable condition.\nWe define a predicate + as re13 e reb meaning retz is a",
+            "stabler contact than reh. Thus, aguinsfl d ugainsn e agahsts and fits] G fits2 G fitsJ. When two sets S i and Sj of contacts are given, we can ignore Si if any contact in Si is smaller or equal in G to its corresponding con-\ntact in Sj.\n(2)A contact is assumed to be realized in a step by step\nmanner. For example, before realizing three vertices contact (againm) between Ff and F\u20192, two vertices contact (agtukm) must be realized, and before realizing two vertex contact, single vertex contact (againstl) must be realized between the same feature pair. In the same way, be fore realizingjitw or fits2, fitsl must be realized.\n(3)The displacement or rotation of a part from its nominal\nposition to achieve a certain contact should be small enough. This assumption restricts the geometry of a feature to make a contact when a certain relative motion is possible. For example, if a flat face FI is permitted to rotate about Ax, itcan realize againm contact with an edge E] by a small rotation (see Fig, 15(a)), however, it must rotate nearly 90 degrees to realize a contact with E2 (b).\n6.2 Primary Contact When a force is applied somewhem on a part, according to where it is applied and its direction, some specific contacts are more easily realized than others.\nIt is a difficult problem to determine proper contacts\nautomatically by evaluating the effect of the force. Instead, we permit the user to specify some contacts cakxl primary contacts explicitly on account of the force effect. The system determines other unspecified contacts automatically taking the primary contacts into consideration.\n6.3 Analysis of Part Motion Our contact combination synthesis is based on the analysis of relative motion of parts in an assembly.\nWhen the initial position of part P is given asp, the\nsmatl change of the part position top + @ can be expressai asp * (1 + AP) by using the differential motion matrix AP. Thus AP can be considered to represent the part motion.\nWhen an assembly composed by n parts Pi (1 S i S n)\nis given, we assign a differential molion matrix A [&i,c$i,dZi,&i, &i,&i] toeach Pi. If there is a pm in the as-\nsembly whose position is fixed in the world, we give null matrix O instead for its differential motion.\nThese parts may have several contacts each other. For each contac~ we can determine FI and FE by using the CVS realizing the contact as explained in 5.1. Based on the equa-\ntion (27), we can derive a linear algebraic constraint between differential motion matrices of parts mwte@O &l OWW(FE). We can ignore small shape errors for calculating the part motion. Therefore, we use the following equation instead of (27):\ntE\u2019 *fE\u2019 * PE\u2019 * AE - tl\u2019 * fi\u2019 * pI\u2019 * AI\n= AC * tI\u2019 *fI\u2019 * JX\u2019 {32)\nwhere fI\u2019 and fE\u2019 represent nominal positions of form features FI and FE in their owner part\u2019s frame. Other variables are the same as the equation (27). All the transformations except A], AE and Ac are calculated by using the nominal geometry of the parts.\nAccordingly, a system of linear equations relating differential motion matrices of all parts in the assembly is constructed. We can eliminate some variables of A1, AE and A c by applying algebraic variable elimination method (for example, Buchberger algorithm [9]) on the system. If we eiiminate such variables as much as possible, we can represent differential motion matrices AP of part P by some limited number of motion variables, which represents the possible motion of P in the assembly.\n6.4 Motion Constraint for Realizing a Contact We approximate the real contact as the contacts of some vertices of the internal feature and its corresponding extended feature. Here we consider a motion constraint to realize such vertex contacts. Suppose Vs are the vertices of the internal feature proposed to approximate a contact between FJ and F2. If F] = origin, we assign FI = FI and FE = F2, otherwise, FI = F2 and FE = FI, According to the assumption (2) in 6.1, we assume that a set of vertices CVS c Vs are now retilzing contacts with the extended feature. As explained in 6.3, by using currently available contacts, which include the contacts realized by CVS, we can compose a set of linear equations relating differential motion matrices of the parts based on (32) and can derive the possible motion of owner(F~) and owner(F@ in the assembly. NOW we select a vertex v ~ (vs - CV~) and consider whether V can achieve an additional contact. If we consider the VF of V, we can derive the following constraint equation based on (3 1).\ntVI?\u2019 *fE\u2019 * pE\u2019 * AE * pi\u2019-l *fI\u2019-l * tvI\u2019-l\n- tv{\u2019 *fi\u2019 * @ * Ar * PI\u2019-l * fI\u2019-l * tvi\u2019-~ = AV (33)\nwhere tvl\u2019, WE\u2019, pI\u2019 and pE\u2019 are the same as the equation (31).fi\u2019 and fE\u2019 are used instead offl andf!?, which represent nominal positions of FI and FE in their owner part\u2019s fkune. These transformations are all determined by using the solid model. AE and Al are differential motion matrices of",
+            "OWW(FI) and owner(FE) and are already derived As the result, a differential motion matrix Av is determined which represents the possible motion of V with respect to the VF.\nBased on the assumption (3) in 6.1, in order to realize a contact near P on the extended feature, V must have a relative motion toward P (see Fig. 16). Therefore, if origirt({ V ) ) is a flat face, A v must have a relative motion\nalong its X axis. If origirt([V}) is a pin or hole, Av must have a relative motion along its Y axis.\n6.5 Contact Combination Synthesis Algorithm Based on the motion constraint for realizing a contact, we developed an algorithm for c#culating atl possible contact combinations.\nIn the foIlowing explanation, c-rel-list is a list of c-rel\nwhich is a structured data with expected contact type (either against or fits) and two nominal features concerning the contact. We can refer to the contact type by a function ctype(c-rel) and two features byfl~(c-ref) andjj2(c-ref).\np-c-info-list and c-info-list are lists of c-inJo which is\nalso a structured data with possible contact type (either againsnm or fits~/2/J) and various contact information (e.g. CV and CL of the contact) required for calculating the possible motion of the parts.\nContact Combination Synthesis Algorithm Outline\ninput: c-rel-list of expected contact relations and p-c-\ninfo-list of primary contacts.\nOutput: All possible contact combinations.\nMethod (l)Prepare a list results with p-c-info-iist as its element. (2)Iterate (3)to(11) for each c-rel in c-rel-list. (3)Prepare null list work. (4)Iterate (5) and (6) for each c-info-list in results. (5)Realize all possible contacts recorded in c-info-lis[,\nthen calculate the part motion in the assembly,\n(6)C\u2019alcuktte all possible ugainstl (if c-type(c-reo is fits,\nfitsl) conracts between fll(c-re~ and fi(c-rel) in consideration of the motion constraint. For each derived contact, prepare a new c-irt$o and record the contact information. Then make a copy of c-i~o-list and push c-\ninfo into the copy, and the copy is pushed into work. (7)Iterate (8) ~Oreach c-injo-lisl in results. (8)Re-exwute (5) and (6) for againstz (if c-type(c-re~ is\nfits, jltsz) contact.\n(9)Iterate (10) for each c-info-list in results. (lO)Re-execute (5) and (6) for ugainso (if c-type(c-re~ is\nfits, jTts3) contact,\n(1 I)Set work to results. (12)Select stable contacts from resulrs according to the\nassumption (1) in 6.1 and return them.\n7. SOLUTION OF POSITION UNCERTAINTIES For each contact combination, we can derive a set of linear equations based on (27) which relates shape errors and differential motion matrices of the parts. If we assign null matrix O as the differential motion matrices of the parts fixed in the world, we can solve the equations by eliminating the variables representing the differential motion in AI, AE and Ac of (27). We developed variable elimination programs based on the Buchberger algorithm [9]. If such variables are successfully eliminated, differential motion matrices of the parts are represented only by shape error variables.\nExpected contacts between Plate and Base are (against\nTop Bottom), (fits HoleI Pin]) and @ts Holez Pinz). RAPT first determines the nominal positions of the parts. Then the system synthesizes possible contact combinations. To evaluate the force effect on Plate, we explicitly assign against3\nFrame for [ downward Plate - ~force\nt P[ate\nBottom\nPosition errors: Perpendicularity errors Parallelism errors:\nHolet: dyhl, dzhl Hole]: ~W, &kI ~OttODl: &b, &b\n\u2018~~ez\u2019 \u2018Jhz\u2019 \u201cm Holez &2, &u Pm!: dwi, &I\nTop: & &\nPiiu: &I, &I Form errors: Pin: 4P2,d.P2\nPinz: &2, &2 HoleI, Hole2: drhi, drhz pinl?im: drP1,drP2\nFig, 17 Nominal geome~ of example parts and shape errors."
+        ]
+    },
+    {
+        "image_filename": "designv11_60_0002128_ma946435m-Figure7-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0002128_ma946435m-Figure7-1.png",
+        "caption": "Figure 7. Schematic representation of the defect used for the modeling: (a) the whole loop disclination with the horizontal and vertical parts; (b) detail of the horizontal part; (d) director; (\u00e2) the tilt angle of the director; (\u03c9) rotation of the director around the core of the defect; (x1) gradient direction; (x2) vorticity axis; (x3) flow direction.",
+        "texts": [
+            "32 As will be seen elsewhere,32 the lines running parallel to the flow direction can be associated with the vertical streak seen by light scattering while the dark spots are associated with the four lobes. The disclination loops found in a thermotropic polymer and described in refs 24 and 25 will be considered as the scattering objects. These studies performed by scanning electron microscopy showed that several types of defects can be present, the most common one being a half-strength twist disclination loop. These loops lie in the shearing plane (x1, x3), and they have a complex structure.25 We will simplify the real structure by considering the object shown in Figure 7a. It bears the main features of the loops found for thermotropics, i.e. lying in the shearing plane, with the director at a certain angle about the loop main axis (x3), which varies between the x3 and the x2 axes with a helical structure. In this case, the orientation of the optical axis d is characterized by two angles representing the tilt and rotation around the core axis \u00e2 and \u03c9, respectively (see Figure 7b). In the case of a helical structure, the angle of rotation is a function of the coordinate along the core, i.e. \u03c9 ) \u03c9(x3). We shall consider here the simplest linear dependence: where p is the number of rotations per unit of length. For example, for px3 ) 1 there is only one turn along the core axis. A further simplification is introduced by considering only the parts of the loops which are oriented along the flow direction. This means that the model will transform the loop into a set of two rigid rods",
+            " A theoretical analysis of a similar effect has been performed recently by considering the perturbation of isotropic polymer suspensions due to the presence of solid particles which also predicted a fourleaf HV scattering pattern.33 The behavior of flowing nematic LCP\u2019s is much more complicated than isotropic liquids, but the physical natures of the perturbations of the stream lines in both cases may be similar. The optical observation of a four-leaf pattern around a large rigid obstacle placed in a flowing liquid crystalline polymer solution34 is predicted by such a hypothesis. We will study the light scattering from the horizontal parts of the disclination loops of Figure 7 which are oriented parallel to the flow direction. First, we will consider the light scattered by a single rod with a spirallike orientation of the optical axis. We will then study the 3D diffraction of a set of such anisotropic rods in order to take into account the experimental oscillation of the scattering intensity in the vertical streak. 5.1. Light Scattering by a Single Rod with a Helical Orientation of the Optical Axis. Let us consider a straight cylindrical rod with a diameter of D and a length L, oriented along the x3 axis (see Figure 7b). The unit vector d of the optical axis has a random tilt \u00e2 around the rod axis which can fluctuate around an average value with a Gaussian distribution function: where \u03c3\u00e2 is the dispersion of the angle \u00e2 and c\u2032 is a normalizing coefficient. Vector d is rotating around the rod axis in accordance with eq 4.1 and can be written as where i, j, and k are the unit vectors along the x1, x2, and x3 axes of the laboratory coordinate system. In the general case of anisotropic media, the dipole moment induced in a rod by the incident light will depend on the effective orientation of the optical axes, determined by the position of the crystallographical axes of both medium and rod, as well as their polarizabilities",
+            " Such a complete study, interesting in itself, was not performed here. The purpose of the present study is to understand the nature of the defects created during the shear flow of liquid crystalline polymers. Thus we will select, on a best guess basis, a range of \u201cadjustable parameters\u201d which have a physical meaning. We will then vary these parameters in order to fit experiments and theory. 6.1. Comparison between Theory and Experimental Observations. The theory shows that if we consider either a single rod (with the rod geometry given in Figure 7) or a set of aligned rods, the scattering patterns are characterized by a very strong intensity maximum around an azimuthal angle of (90\u00b0, similar to the vertical streak found experimentally (there are of course several polarizer positions with selected director arrangements for which there is no scattering). This result was expected since we are dealing with aligned rods. In no case could a pattern with four lobes in HV polarization be predicted. This provides the first result: The four lobes found experimentally cannot be explained with this model"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0003840_acc.2000.879454-Figure1-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0003840_acc.2000.879454-Figure1-1.png",
+        "caption": "Fig. 1 Half Toroidal CVT",
+        "texts": [
+            " This model has been used to simulate the CVT responses to various ratio change commands under specified torque and angular velocity inputs. Parameters associated with CVT control and performance, including torque and speed output, CVT ratio, etc., have been determined as time functions in the simulation. The effectiveness and stability of the control system have been verified by the simulation results. 2. TRACTION DRIVE GEOMETRY A half-toroidal traction drive CVT consists of three basic elements: Input disk, output disk and power roller. As shown in Fig. 1, the geometry and dimension of a traction drive are determined by four design parameters: r+avity radius, r d o r u s radius, (+half cone angle of the roller, and R2-c radius of the roller profile. For a given CVT, the theoretical gear ratio solely depends on the swing angle 4 and is represented by the following equation: l = 1 - . w r3 - I + k , -cos(e-(b) w3 r, l + k , -cos(O+(P) --- where, W and W 3 are respectively the angular velocities of the input disk and output disk, ko is a constant defined as (r, - ro) / rh rl and r3 are the radii of the two disks at the respective points of contact, related to the roller swing angle (#I as: q =r,[l+k, -cos(B+@)] r3 =r,[l+k, -cos@-@)] For a specified CVT ratio i , the corresponding roller tilt angle is solved from equation 1) using half-angle formula in terms of the CVT geometry parameters by the following equation: - ( i + 1) sin e + JX ( i - l)[(k, + 1) + cos 01 Q = 2 tan -' 3) 0-7803-551 9-9/00 $1 0"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0002368_anie.199524091-Figure4-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0002368_anie.199524091-Figure4-1.png",
+        "caption": "Fig. 4. Top: Scheme showing the one-electron enzymatic oxidation of immobilized CcP by w l i i ~ i o n phase cytc. with subsequent electrochemical reduction. Bottom DC cyclic voltamiiiogr;ims (1 : = S m V s - ' ) of immobilized CcP in 0.1 moldn1C3 sod~ii i i i phosphate hutfer. p H 6.3 ( a ) and arrer successive additions of 7 . 0 ~ ~ (b) and 4?pM C y l C ( C )",
+        "texts": [
+            " To this end, ET reactions between soluble CcP,,,, and the modified electrode were measured (Fig. 3 ) . As expected from its biological function, and due to the relatively free accessibility of the immobilized heme redox center, CcP serves to regenerate immobilized cytc,,,, (via compound I), resulting in an enhanced reduction current a t the electrode. as illustrated schematically in Figure 3. In a series of complimentary experiments, CcP was immobilized using DTSSP, and its direct electrochemistry was characterized in deoxygenated electrolyte (Fig. 4). The weighted k'& was calculated as 0.04 s - ' . The lower value for the ET rate constant for CcP (with respect to cytc) is attributable to the longer probable ET pathway between the heme and the gold surface. Assignment of such a pathlength is, however, difficult as there are 23 lysines on the surface of CcP, available to bind with DTSSP. Plots of ips, or ipc vs. 1: were linear [for example, for iTd, y (/nA) = 9.1 x(/niVs-') + 2.7 (/nA), data not shown], thus confirming that the electrochemical reactions were confined to the surface",
+            " Ag/AgCl at 1: = 10 mVs-', relative to the reduction potential of CcP in solution (compound I1 containing Fe\"' ions). estimated as 848 mV vs. Ag/AgCI.[\"' This large shift in the enzyme's redox character is attributed to the substantial change in the local physical and chemical environment of the immobilized heme redox center. Integration of the reduction peak of CcP demonstrated that the surface coverage of the enzyme was 2.8 x lo-\" molcm-2, a lower apparent value than that measured for cytc. On addition of soluble cytc to the Au-CcP electrode, the CcP reduction current is enhanced (Fig. 4). The redox conditions El. within the electrochemical cell are such that the reaction is reversed from that in nature. where CcP acts as a terminal electron acceptor during tlie oxidation of cytc,,,,, (E l ,* cytc = + 40 mV vs. AgjAgCI) .I' 21 This reverse reaction has bccn studied, for example by means of laser radiolysis.[*] Previously. solution-phase components of the ET chain have been shown to communicate with monolayers a t an electrode surface.['3 ' 51 The novel electrode arrangement that we describe here for the electrochemical investigation of iminobilized components of the ET chain offers potential advantages over such homogenous models"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0003803_a:1004727124456-Figure1-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0003803_a:1004727124456-Figure1-1.png",
+        "caption": "Figure 1 Schematic representation of the test apparatus and the image processing system.",
+        "texts": [
+            " Aluminum was selected as the disk material for two reasons namely 1) it is increasingly used in weight saving applications and 2) to minimise the loss or attenuation of X-rays after passing through the contact. The specimen holder was fitted onto a beam structure with provisions for sensing the experimental parameters such as friction force, displacement, temperature close to the contact and the oxygen potential surrounding the testing atmosphere. A FeinFocus X-ray microscope with a beam spot size of 4\u201320 \u00b5m diameter was used in all experiments. The schematic representation of the apparatus and the image processing system is provided in Fig. 1. Steel pins of dimension\u03c6 5 mm\u00d7 3 mm thick specimens were polished using a 1000 grit SiC 0022\u20132461 C\u00a9 2000 Kluwer Academic Publishers 1589 abrasive paper to provide a surface roughness value of Ra 0.03\u20130.05\u00b5m and was fixed to specimen holder designed to facilitate X-ray observation. The counterface disks were made of Al 6061 machined and fine turned to the roughness value of 0.04\u20130.06 \u00b5m Ra value with a skewness and kurtosis of 1.7 and 3.4 respectively. The wear and seizure experiments were carried out at different sliding speeds of 2, 4 and 5 m/s"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0003042_iros.1993.583851-Figure7-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0003042_iros.1993.583851-Figure7-1.png",
+        "caption": "Fig. 7: calculated sliding distances in the parallel direction of the gripping plane",
+        "texts": [
+            " Then in order to escape from a situation where the object cannot be grasped: object is t,ranslat,ed into t,he shaded portion shown in Fig.6. In the case where the gripping plane and the upper surface of the table are not parallel, the cross line between the gripping plane and the upper surface of the table is calculated. Then we solve the problem by means of calculating the sliding distances in the directions both parallel and perpendicular to the cross line. In the case of the parallel direction, we can calculate the sliding distance of either object or obstacle in the same way as describe above. Fig.7 shows the results of sliding distance of object in the parallel direction of the gripping plane pll shown in Figd when GGPO equals the whole GPO. GGPO(object,pl,, Ohl) is slid in width length (x-direction of the gripper) of the fingers more than calculated sliding distance for the stable grasp and it is illustrated in Fig.8. Since a part of GGPO(object,pZl, 6hl) does not overlap with the collision area COPO(env,pll, Ohl), objed can be grasped. In the case of a direction perpendicular to the cross line, the shapes of the collision areas of the obstacle COPO(obstacle,pll, Oh,) or the grasp candidates areas GGPO(object,pl,,Bhl) expected to be graspable change, and the sliding distance cannot be calculated directly from G G P O and COPO"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0000635_015502-Figure1-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0000635_015502-Figure1-1.png",
+        "caption": "Figure 1. The physical situation and the description of the coordinate systems.",
+        "texts": [
+            " 48 (2016) 015502 E I Saad 3 the motion perpendicular to the line of their centers, which is considered in the present study. The results for the resistance coefficients are in good agreement with the available data in the literature for the limiting cases. We consider two spherical particles which are situated external to each other with their centers distant apart is d, and have arbitrary radii, R1 and R .2 The particles can differ in velocities and angular velocities. We assume that the particles of the same material, therefore they have the same slip length. The geometry of the problem is shown in figure 1. Here, x y z z, , , , ,( ) ( )r f and r, ,( )q f denote the Cartesian, circular cylindrical, and spherical coordinate systems respectively. For the quasi-steady situation at low Reynolds number, the velocity field q  and dynamic pressure field p satisfy Stokes equations: q 0, 2.1\u00b7 ( ) = p q 0, 2.22 ( )m -  = where \u03bc is the fluid viscosity. Despite the fact that the present boundary value problems are not axially symmetric, it is possible to reduce the problem of solving the Stokesian field equations and the equation of continuity to one involving three scalar functions that depend only on r and "
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0002975_s0167-8922(00)80164-9-Figure4-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0002975_s0167-8922(00)80164-9-Figure4-1.png",
+        "caption": "Fig. 4 Elliptical EHL conjunction showing the minor axis of Hertzian contact ellipse inclined at an angle 0 to oil entraining vector",
+        "texts": [
+            " The duochromatic interference fringe pattern obtained by a Xenon light source with a flash duration of 20 Its through red and green filters was recorded with a high speed VCR (200 frames per second) and a 35mm camera attached to a microscope. The experiment was conducted under pure sliding conditions where the glass disk was rotated and the spherical roller was stationary; the line of the minor Experimental conditions axis of the contact ellipse was fixed so as to be at a constant angle, 0, to the flow direction. That is, 0 is the angle of the minor axis of the Hertzian contact ellipse measured from the oil entraining vector or the direction of sliding, as shown in Fig. 4; angles of 0 \u00b0 and 90 \u00b0 correspond to entrainment along the minor and major axes of the Hertzian contact el l ipse respectively. The oil used in this experiment was Santotrac 100. The properties are listed in Table 1. The experiments were done under a room temperature of 20 +_ 1 \u00b0C. The ranges of contact load and sliding veloci ty are summarized in Table 2. 3. RESULTS AND DISCUSSION 3.1. Occurrence of dimple Representative interferograms indicating dimples produced in the contact ellipse for a contact load of 39"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0003736_aim.2001.936501-Figure4-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0003736_aim.2001.936501-Figure4-1.png",
+        "caption": "Fig. 4. Geometry of subspace angles",
+        "texts": [
+            "d,in This approach is well defined, easy to understand and ready to be implemented in computer codes. But the \u201cZero Subspace Angle\u201d condition is not complete(necessary but not sufficient) in deciding discontinuity. So sometimes the mixing factor r drops to zero, even in a case where discontinuity is not imminent. If someone has to know exact situation, this approach can fail to give exact information. So we propose a new approach that can handle this situation in next section. B. A New Approach Figure 4 shows a geometry of subspace angle if dimension of matrix A of equation (20) is 2 x 3. As shown in figure 4, the resultant dimension of solution space is 1. Ian\u2019s approach measures subspace angles between solution space and every faces of hypercube. But there is no need to measure all of them. In figure 4, the solution space has no intersection with the face F3, so the subspace angle between the solution space and F3 needs not be calculated in this case. Because of the above reason \u201cZero Subspace Angle\u201d condition is not said to be complete. Now there is one question remained. How can one select subspaces which have an intersection with solution space\u2018? In figure 4, we can find an answer, the solution space has intersection with face Fl and F2, but not with face F3. In this case, the first and second component of optimal solution z* has the same magnitude with ~ ~ z * ~ ~ ~ , while the third one does not. During the running of algorithm, one can easily determine indices of saturated components of z*. Indices of columns which compose matrix A2 used in our algorithm or Cadzow\u2019s algorithm are the indices of saturated elements of z*. So one can easily find the faces that have an intersection with the solution space, and determine the singular index using the corresponding faces"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0002904_s1474-6670(17)52428-3-Figure3-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0002904_s1474-6670(17)52428-3-Figure3-1.png",
+        "caption": "Fig. 3. The AID robot.",
+        "texts": [
+            " Then it comes: We deduce the relation: W2 = W, R,-' R2 thus: B = Rrl R2 (26) (27) (28) which expresses the (c-b) columns of W 2 and 12 as linear combinations of the b independent columns of W 1 and J 1 respectively. The zero rows of B correspond to the independent columns of [WPj. Remark : In practice the orientation error 0 is difficult to measure. so only the equations corresponding to the position error will be taken into account. The base parameters in this case may be different than the general case. they can be obtained using the same procedure and by taking into account the model: D = Jt(q) t.X (29) where Jt contains the first three rows of J. EXAMPLE Let US consider the robot AID. Fig. 3. it has 6 rotational joints. The complete parameters of the robot for a given location of the base of the robot and a given end effector are given in Table 1. Using the previous algorithm for the position only. Eq.29. we find that the number of base parameters is equal to 23. they are given in Table 2. This paper presents a general method to classify the calibration parameters of robots. The robot location and the end effector location parameters are rreated by an approach similar to those defining the link frames"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0000759_fpmc2016-1739-Figure1-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0000759_fpmc2016-1739-Figure1-1.png",
+        "caption": "Figure 1. Dynamic seal tribometer.",
+        "texts": [
+            " The investigated surfaces, the test conditions and the test procedure are the subject of the third section. Section four presents the experimental results, which are discussed in section five and compared with theoretical predictions in section six. Finally, a conclusion and an outlook are provided in section seven. In order to investigate the lubricated line contact of a hydraulic seal in detail, a new test rig was designed and set up at the Institute for Fluid Power Drives and Controls (IFAS). The main elements of the test rig are shown in Figure 1. The test rig is driven by an electric gear motor with an attached frequency converter. A secondary cycloidal gearbox with a transmission ratio of 43 can be included additionally. Thus, a wide range of rotational speeds is possible: 0.2 - 2.4 rpm (rotation per minute) with or 8.5 \u2013 105 rpm without the additional gearbox. The revolution speed of the drive shaft is measured using an incremental angle encoder (90000 increments/revolution, accuracy \u00b1 5 seconds of degree). The rotating cylinder and the seal specimen are located in the test chamber"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0000878_978-3-319-47614-8-Figure2.4-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0000878_978-3-319-47614-8-Figure2.4-1.png",
+        "caption": "Fig. 2.4 Load imposed on a beam of the crane",
+        "texts": [
+            ", the weight of a man as shown in Fig. 2.2 on the structure. This load can be moving also. There can be other imposed loads (or loads acting on the structure), e.g., the wind load as shown in Fig. 2.3. This load is from the pressure of the wind on the roof which therefore is perpendicular to the roof as shown in Fig. 2.3. This is net load shown by integrating the pressures over the surface of the roof. Another example of imposed loads is a weight being carried on a crane that is being hoisted from one place to another see Fig. 2.4, that we commonly see in a workshop. This load acts on the beam of the crane. Yet another familiar example is the piston of an internal combustion engine on which we have a pressure of the burnt gases from combustion, see Fig. 2.5. The pressures integrated over the area of the piston give the load or Force as shown. These forces or loads on a structure cause them to deform and in this process of deformation we determine the effects of deformation (strain and stress) that allow us to design the structure to be safe"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0000160_6.2015-0904-Figure2-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0000160_6.2015-0904-Figure2-1.png",
+        "caption": "Figure 2. Schematics of the mechanical assembly.",
+        "texts": [
+            " Using a linear actuator reduces the necessary torque generated by the motor for a given linear strength. Figure 1 presents a computer-aided design (CAD) overview of the entire system. This representation shows the three points A, B and C to represent the kinematics of the aileron actuation system. The lever arm length is limited by the space available in the wind tunnel; this length was used in the choice of the electrical motor, the motor drive and the linear actuator. A. Mechanical equations The system can be represented by a plain problem in terms of its mechanical modeling; Fig. 2 presents the schematics of the mechanical assembly. This mathematical representation uses the three coordinate systems consisting of a point of reference with corresponding \ud835\udc65 \u00a0and \ud835\udc66 \u00a0 axes. The main reference part is the wing, represented by \ud835\udc65!, \ud835\udc66!. The first coordinate system (A, \ud835\udc65!, \ud835\udc66!,) is thus considered a reference and is fixed about the wing; the second coordinate system (B, \ud835\udc65!, \ud835\udc66!) is attached to the linear actuator. The third coordinate system (C, \ud835\udc65!, \ud835\udc66!) is attached to the main axis of the lever arm",
+            " axis of the lever arm and the axis of reference, \ud835\udc65!. The distances D, \ud835\udc3f! and \ud835\udc3f! are fixed distances and depend on the system design. The \ud835\udc65 \u00a0distance is the parameter of the system and the control algorithm is based on its value. The equation 1 presents the relationship between the linear actuator distances introducing the d distance corresponding to the variable distance of the entire linear actuation. \ud835\udc51 = \ud835\udc37 + \ud835\udc65 (1) A joint diagram is presented in Fig. 3; this representation shows the interaction and joints between the parts. In Fig. 2 and 3 the numbers correspond to the parts listed in Table 1. As shown in Fig. 2 and 3, the joint between the wing and the linear actuator, and the joint between the wing and lever arm are revolute with A and C as points of application and \ud835\udc67! \u00a0and \ud835\udc67! \u00a0 \u00a0as revolution axes respectively. The \ud835\udc67 \u00a0axis is determined such that the coordinate system is orthonormal and direct. The joint between the linear actuator main frame and its arm is defined as a prismatic joint with \ud835\udc65! \u00a0as axis and C as reference point. At point B, there is a revolute joint of the \ud835\udc67! \u00a0axis is the joint between the actuator piston rod end and the lever arm"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0003444_03601218608907529-FigureI-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0003444_03601218608907529-FigureI-1.png",
+        "caption": "Fig. I (a) A planar truss. (b) The graph model.",
+        "texts": [],
+        "surrounding_texts": [
+            "The mathematical model of a planar truss is considered to be a finite graph S. The nodes and members of S are in a one-to-one correspondence with joints and bars of the truss. A similar correspondence exists among the nodes and members of a subgraph Si and the joints and bars of the corresponding substructure, respectively. Trusses are assumed to be supported in a statically determinate fashion, and the effect of extra supports can easily be included for each special case. Trusses are considered to be stable, and critical forms are excluded from our discussion. For two crossing bars an extra node is considered in the crossing point. As an example, the graph model of a planar truss in Fig. l(a) is depicted in Fig. l(b). The dimension of a statical basis of a planar truss is given by where M, N, and yo(S) are the numbers of members, nodes, and components of S, respectively. For a connected truss, yo(S) = 1. A topological invariant, the so-called first Betti number, is defined for a graph as follows:"
+        ]
+    },
+    {
+        "image_filename": "designv11_60_0000771_wcica.2016.7578490-Figure1-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0000771_wcica.2016.7578490-Figure1-1.png",
+        "caption": "Fig. 1. Structure of quadrotor UAV with X-shaped frame",
+        "texts": [
+            " The backstepping robust control 978-1-4673-8414-8/16/$31.00 \u00a92016 IEEE 697 scheme with disturbance observer for the position subsystem of a quadrotor UAV is presented in Section III. In Section IV, an attitude control strategy based on linear filtering method is designed with nonlinear disturbance observer. Next, simulation results and analyses are provided in Section V. Finally, conclusions are drawn in Section VI. II. DYNAMIC MODEL OF QUADROTOR UAV Quadrotor UAVs are basically actuated by four motors. Fig. 1 shows the structure of a quadrotor UAV with X-shaped frame, where \u03c9i(i = 1, 2, 3, 4) denotes rotating speed of the ith rotor, Fi(i = 1, 2, 3, 4) represents the thrust of the ith rotor, Mi(i = 1, 2, 3, 4) is the torque of the ith rotor, ObXbYbZb is the body coordinate system, OeXeYeZe is the inertial coordinate system. Considering external disturbances, the dynamic model of a quadrotor UAV can be expressed as [9] mp\u0308 = U1Re3 \u2212mge3 + dF (1) \u0398\u0308 = f(\u0398, \u0398\u0307) +G(\u0398)\u0393 + d\u0393 (2) where m is the mass, p = [x, y, z]T denotes the position of the quadrotor UAV with respect to an inertial frame OeXeYeZe, \u0398 = [\u03c6, \u03b8, \u03c8]T represents the three Euler angles, called roll angle, pitch angle and yaw angle"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0001817_0141-0296(93)90032-y-Figure4-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0001817_0141-0296(93)90032-y-Figure4-1.png",
+        "caption": "Figure 4 Guyed rigid frames",
+        "texts": [
+            " Either the dynamic pulsating loads of an adjacent span in gallop, or simply the case of an in-cloud ice load on only one of the adjacent spans can apply a torque load that can rotate the otherwise freely pivoting mast. This may cause problems with clearances to electrical jumpers. Awareness of the potential for twisting led to an important change of the wire attachment detail for guyed single masts used for the large line angles of a 735 kV line in Canada. The guys and phase wires were brought to a single point within the mast in order to ensure that neither static nor dynamic unbalances would cause rotation. Guyed rigid frames Guyed rigid frames, as typified in Figure 4 are very popular and are frequently used at voltages from 69 kV to 230 kV and infrequently as high as 500 kV. A single mast and a set of guys and anchors (usually four) replaces the lower tower body and foundations while the upper part of the structure remains practically the same as it would for a totally rigid framing. The significant difference between these guyed rigid frames and the guyed single masts mentioned earlier is that each pair of side guys is attached at widely separated points so that torsional stability is inherent, an imperative with at least three phases attached to the structure. Furthermore, the two separated points of attachment automatically result in equal guy tensions in the absence of externally applied horizontal loads, an important point that is discussed in greater detail later in the paper. Figure 4(a-h) shows some of the framing and guying arrangements that have been used by designers in attempts to attach the guys and, in some cases, to thread the guys between phases, maintaining the required electrical clearances while minimizing the eccentricities of the loads with the guys and the mast. The overall objective is to keep the lines of action of the guy system centric with the centre of effort of the major load combinations and the centre line of the mast. With a significant eccentricity, a very large horizontal reaction Guyed structures for transmission lines: 14. B. White will be applied to the mast footing and large shears and bending moments can be put into the mast. On some projects, the shear and bending loads applied to the single mast of the structures in Figure4(e,g) were large enough to produce a single leg that weighed more than the two masts of a guyed V tower. The oft prescribed demand for torsional strength to resist the possible loads of broken wires or to resist the complex of loads that might be imposed by failure of an adjacent structure (the duty of failure containment) can create a problem as some of the guy systems with crossed guys, as in Figure 4 (e, g) and Figure 6(d), will permit a rotation that reduces the spacings of the resisting guys, the guy tensions and strains then increase with more rotation etc. until failure or snap through may occur. For this reason some designers have used auxiliary or longitudinal secondary crossarms that will ensure that under torque forces, the spacings of the acting guys will move apart and stabilize the problem as shown in Figure 4(h) and Figure 6(e). These extra arms do not add significantly to tower weights as they keep some of the heavier guy loads out of the upper part of the structure but they can and do make the tower lay out and assembly on the ground more difficult. The tower shown in Figure 4(a) was used in Finland as early as 1927, fabricated as a lattice work in steel. In the 1960s, a similar tower of aluminium was being used in North America for helicopter transport and erection on lines of remote or difficult access. Thousands were installed in Western Canada and Alaska in areas of permafrost or unstable ground conducive to frost heave. A structural fuse was placed in one of the guys to prevent crushing of the tower from excessive guy tensions if the mast footing was lifted. Figure 4(b) shows an aluminium version without ground or shield wires that made use of an extruded aluminium tube as a crossarm. The towers of Figure 4(c), also described in References 14 and 15, are guyed rigid towers fabricated of Corten Hollow Structural Shapes, both round and square. More than 4000 have been used at 138 kV in Manitoba and Saskatchewan, Canada, being installed by helicopter in some areas otherwise accessible only when frozen in the winter. The guyed Y of Figure 4(d) has been used for voltages of 230 kV and up to 500 kV but has not proved to be very useful at the higher voltages. The costs of accommodating the large moments that are developed within the frame from the eccentricities of the major forces being one of the costly features. The towers of Figure 4(e,g) are well suited to the aluminium/helicopter marriage, the present author having witnessed the transport and erection of more than 75 towers of 345 kV in one day by one medium sized helicopter. Guyed and hinged (or pinned) masted structures The guyed portal, guyed V and the cross-rope suspensions (CRS) of Figure 5 are the three most used versions of this family, each of them comprising two tripods of a mast and two guys, the apexes of which are held apart by the crossarm of the portal tower, held together by the crossarm of the guyed V and also held together by the tensioned wire rope suspension assembly of the CRS"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0001235_b978-0-08-099425-3.00005-4-Figure5.36-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0001235_b978-0-08-099425-3.00005-4-Figure5.36-1.png",
+        "caption": "FIGURE 5.36",
+        "texts": [
+            " Thus for a given tyre on a given vehicle it is possible (Milliken and Milliken, 1998) to optimise camber angle for a given combination of slip angle and tyre load. Two of the components of moment acting in the tyre contact patch have been discussed. The generation of rolling resistance moment was described while discussing the free rolling tyre in Section 5.4.3. The self-aligning moment arising due to slip or camber angle was discussed in Sections 5.4.7 and 5.4.8. For completeness the final component of moment acting at the tyre contact patch that requires description is the overturning moment that would arise due to deformation in the tyre as shown in Figure 5.36. The forces and moments as computed in the SAE reference frame are formulated to act at P, this being the point where the wheel plane intersects the ground plane at a point longitudinally aligned with the wheel centre. In Figure 5.36 it can be seen that distortion of the side walls results in a lateral shift of the contact patch that may result from either slip angle or camber angle or a combination of the two. The resulting offset tyre load introduces an additional component of moment Mx. Attention to the sign convention associated with the tyre reference frame is again needed if the moment is to be included in a tyre model. In Figure 5.36, to assist understanding, Fz is represented as the tyre load acting on the tyre rather than the negative normal force computed in the ZSAE direction. y Fz P YSAE O Camber angle ZSAE Mx = Fz y Generation of overturning moment in the tyre contact patch. A consideration of the overturning moment is generally more important where relatively large displacements in the tyre occur, as with aircraft tyres (Smiley, 1957; Smiley and Horne, 1960). Overturning effects are also of major importance for motorcycle tyres, particularly in terms of matching the behaviour of front and rear tyres"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0001098_ssci.2016.7850242-Figure2-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0001098_ssci.2016.7850242-Figure2-1.png",
+        "caption": "Fig. 2. Rotation degree of leg",
+        "texts": [
+            " Then we verify the effect of reduction in recalculation cost by obtain walking behavior in rough terrain with obstacles as the target task. II. SIX-LEGGED ROBOT The six-legged robot we used in this study is shown by Fig.1. Legs are out of the body vertically and at equally spaced intervals. All legs move separately at the same time in simulation. Leg number 1 and 2 are positioned in front of the robot. The aim of the simulation is to get the forward walking behavior. The posture of the robot is calculated based on forward kinematics, then reach position P=(px, py, pz) of the leg as shown in Fig.2 is calculated by joint angle and lengths. Where \u03b81 is the rotation angle around the y axis and \u03b82 or \u03b83 is the rotation angle around the x axis. Eq.(1) shows an arithmetic expression of three dimensional coordinate by using coordinate transformation matrix. In the equation, li is the length of i-th leg part, \u03b8m,i means the joint angle of joint i in m-th leg and u is a 978-1-5090-4240-1/16/$31.00 \u00a92016 IEEE normal vector. Symbol Tr and Rot mean translation and rotation matrix each other. (1) III"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0003684_20.996240-Figure8-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0003684_20.996240-Figure8-1.png",
+        "caption": "Fig. 8. Field distribution for the generator/flyback association: ! = 460 rpm; t = 14; 5 ms; and = 40 .",
+        "texts": [
+            " The current and voltage waveforms for each case are shown in Figs. 3 and 4, respectively. In order to validate the implemented procedure, the permanent magnet generator feeding a flyback converter is simulated and tested. The complete circuit topology is shown in Fig. 5. The rotor speed is constant and equal to 460 rpm. The voltage across the transistor and the current trough the secondary winding of transformer at steady state operation (from ms to ms) are given in Figs. 6 and 7, respectively, as well as the corresponding experimented results. Fig. 8 shows the magnetic flux distribution at ms. The complete electric transient can be investigated by the voltage and current waveforms of each circuit element. Figs. 9\u201311 show some simulated transient waveforms. A general method directly coupling field and circuit equations taking into account the rotor movement in electrical machines is presented. A flyback converter fed by a permanent magnet generator was chosen to validate the proposed methodology. This procedure allows the analysis of the electric circuit transient an"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0001235_b978-0-08-099425-3.00005-4-Figure5.42-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0001235_b978-0-08-099425-3.00005-4-Figure5.42-1.png",
+        "caption": "FIGURE 5.42",
+        "texts": [],
+        "surrounding_texts": [
+            "In order to obtain the data needed for the tyre modelling required for simulation, a series of tests may be carried out using tyre test facilities, typical examples being the machines that are illustrated in Figures 5.42 and 5.43. The following is typical of tests performed (Blundell, 2000a) to obtain the tyre data that supports the baseline vehicle used throughout this text. The measurements of forces and moments were taken using the SAE coordinate system for the following configurations: 1. Varying the vertical load in the tyre 200, 400, 600, 800 kg. 2. For each increment of vertical load the camber angle is varied from 10 to 10 with measurements taken at 2 intervals. During this test the slip angle is fixed at 0 . 3. For each increment of vertical load the slip angle is varied from 10 to 10 with measurements taken at 2 intervals. During this test the camber angle is fixed at 0 . 4. For each increment of vertical load the slip and camber angle are fixed at zero degrees and the tyre is gradually braked from the free rolling state to a fully locked skidding tyre. Measurements were taken at increments in slip ratio of 0.1. The test programme outlined here can be considered a starting point in the process of obtaining tyre data to support a simulation exercise. In practice obtaining all the data required to describe the full range of tyre behaviour discussed in the preceding sections will be extremely time consuming and expensive. The test programme described here does not, for example, consider effects such as varying the speed of the test machine, changes in tyre pressure or wear, changes in road texture and surface contamination by water or ice. The testing is also steady state and does not consider the transient state during transition from one orientation to another. High Speed Dynamics Machine for tyre testing formerly at Dunlop Tyres Ltd. Most importantly the tests do not consider the complete range of combinations that can occur in the tyre. The longitudinal force testing described is limited by only considering the generation of braking force. To obtain a complete map of tyre behaviour it would also, for example, be necessary to test not only for variations in slip angle at zero degrees of camber angle but to repeat the slip angle variations at selected camber angles. For comprehensive slip behaviour it would be necessary at each slip angle to brake or drive the tyre from a free rolling state to one that approaches the friction limit, hence deriving the \u2018friction circle\u2019 for the tyre. Extending a tyre test programme in this way may be necessary to generate a full set of parameters for a sophisticated tyre model but will significantly add to the cost of testing. Obtaining data requires the tyre to be set up at each load, angle or slip ratio and running in steady state conditions before the required forces and moments can be measured. By way of example the basic test programme described here required measurements to be taken for the tyre in 132 configurations. Extending this, using the same pattern of increments and adding driving force, to consider combinations of slip angle with camber or slip ratio would extend the testing to 1452 configurations. In practice this could be reduced by judicious selection of test configurations but it should be noted the tests would still be for a tyre at constant pressure and constant speed on a given test surface. Examples of test results for a wider range of tyres and settings can be obtained by general reference to the tyre-specific Flat Bed Tyre Test machine. (Courtesy of Calspan.) publications quoted in this chapter and in particular to the textbook by Pacejka (2012). For the tyre tests described here the following is typical of the series of plots that would be produced in order to assess the force and moment characteristics. The results are presented in the following Figures 5.44e5.53 where a carpet plot format is used for the lateral force and aligning moment results: 1. Lateral force Fy with slip angle a 2. Aligning moment Mz with slip angle a 3. Lateral force Fy with aligning moment Mz (Gough Plot) 4. Cornering stiffness with load 5. Aligning stiffness with load 6. Lateral force Fy with camber angle g Lateral force Fy with slip angle a. (Courtesy of Dunlop Tyres Ltd.) 7. Aligning moment Mz with camber angle g 8. Camber stiffness with load 9. Aligning camber stiffness with load 10. Braking force with slip ratio Aligning moment Mz with slip angle a. (Courtesy of Dunlop Tyres Ltd.) Lateral force Fy with aligning moment Mz (Gough Plot). (Courtesy of Dunlop Tyres Ltd.) Before continuing with the treatment of tyre modelling, readers should note the findings (van Oosten et al., 1999) of the TYDEX Workgroup. In this study a comparison of tyre cornering stiffness for a tyre tested on a range of comparable tyre test machines gave differences between minimum and maximum measured values of up to 46%. Given the complexities of the tyre models that are described in the following section the starting point should be a set of measured data that can be used with confidence to form the basis of a tyre model. Cornering stiffness with load. (Courtesy of Dunlop Tyres Ltd.)"
+        ]
+    },
+    {
+        "image_filename": "designv11_60_0000164_vss.2014.6881123-Figure1-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0000164_vss.2014.6881123-Figure1-1.png",
+        "caption": "Fig. 1 HSM coordinate frames and longitudinal variables",
+        "texts": [],
+        "surrounding_texts": [
+            "The mathematical model of longitudinal dynamics of the rigid body hypersonic missile [3] (see Figs 1 and 2) that is propelled by an air-breathing jet engine (usually a scramjet) is considered as [1], [2] 978-1-4799-5566-4/14/$31.00@2014 IEEE cos( ) sin( ) sin( ) sin( ) cos( ) cos( ) yy T DV g m h V L T gq mV V q Mq I V (1) where ( )V t denotes the forward velocity; ( )h t denotes the HSM altitude; ( )t denotes the angle of attack; ( ), ( )t q t denote the pitch angle and the pitch rate respectively; ( )t denotes the downrange distance (see Fig. 2) measured with respect to the initial HSM position; , yym I denote the HSM mass and the moment of inertia about the HSM body y-axis respectively; g is the gravity acceleration; , ,T D L denote the thrust, drag, and lift forces, respectively; M is the pitching moment about the body y-axis. In the longitudinal dynamics (1) the flexible modes associate with aero-thermo-elastic effects are neglected for simplicity. The thrust T , drag D , and lift L forces, as well as the pitching moment M about the body yaxis are the nonlinear uncertain functions that can be given as [2]-[4] ( , ), ( , ), ( , , ), ( , ) e e e L L D D M M T T T (2) where ( )e t denotes the elevator surface deflection, and is the dimensionless fuel-to-air ratio ( 0,1.5 ). The control vector is , T eu (3) while a vector of the control output is selected as , .Ty h (4) Unlike in the literature on Hypersonic Vehicle control [4]- [6], where the control output , Ty h V is driven to a constant vector after a short transient when the velocity and altitude are commanded to increase from zero to the prescribed constant values, the presented work considers control of Hypersonic Missile (HSM) in the end-game scenario with a different vector-output (4) and time-varying commanded output trajectories [3] (see also Fig. 2) ( ) ( ), ( ) T c c cy t h t t (5) The goal of the considered end-game scenario is to maximize target penetration by means of generating the optimal end-game HSM trajectory (5) for system (1) and robustly following this trajectory by means of control (3) in the presence of the bounded perturbations (2). The optimal command trajectory is obtained in [3] by minimizing the following cost functional: ( ) ( ) ( ) ( ) , , , , , T f c f f c f T J x t x t x t x t x V q h (6) where 5 5 is a positive definite weighting matrix, and ( )c fx t denotes the desired final value of the state x . The terminal conditions on the state x are defined to ensure 090 and the angle of obliquity (AoO in Fig. 2) equal to zero at the impact as ( ) / , ( ) 0 deg , ( ) 0 deg/ , ( ) 0 , ( ) 90 deg f f f f f f V t V ft s t q t s h t ft t (7) The solution of the trajectory optimization problem is obtained numerically [3] and the optimal output command trajectory is used for the robust output feedback tracking controller design. Remark 1. In this work, the lift-elevator coupling, which usually represents the main source of instability of internal/zero dynamics, is neglected. In a case of significant coupling the canards may be added to HSM to compensate it."
+        ]
+    },
+    {
+        "image_filename": "designv11_60_0003692_iros.1990.262430-Figure1-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0003692_iros.1990.262430-Figure1-1.png",
+        "caption": "Fig. 1 Mathematical Model",
+        "texts": [
+            " In this paper, to make the problem clear, we focused on the collision problem of a 1-link manipulator positioned against a flexible wall with a linear state feedback control. Theoretically, we have analyzed this system by the Lyapunov method and proved that the system can be controlled stably. Experimentally, we also showed that the 1-link manipulator can be controlled by the force feedback control even with a collision phenomenon. 2. SYSTEM DYNAMICS A mathematical model was first established to analyze the stability of a 1-link manipulator positioned against a flexible wall by the Lyapunov method as shown in Fig. 1 with following ascumptions. 2.1 Ass\"0Ns The following assumptions are made for the analysis in this study (1) The manipulator consists of a rigid and mass-free link and a lumped mass at the end of the link. A control torque is applied to the joint axis. - 501 - (2) The wall is represented as an equivalent mass-spring-damper system which can describe the local deformation, especially the transicnt vibration. (3) The impact force is assumed to occur by the local deformation both of the manipulator and of the wall when the manipulator contacts and plunges into the wall",
+            "2CASE OF THE CONTACT STATE In this case, U 7 =0 and t > = -l/p, thus fw E ( I+po Hu3n , (I?;k) = l a Hu-lR Q (14) 2 If Eq. (14) is substituted into Eq. (10) then Ak = a HulD From this equation, if then VSO According to the former results, if the following condition ( Eq. (18) and (19) ) is satisfied on the feedback gains and constant a, K , > O , K,>O , (18) ng.2 overview of E x p e r h u e d Setup a s 2 5 (19) the scalar function V is a Lyapunov function. This implies that the system shown in Fig. 1 and including collision can be controlled stably. Moreover, when Eq. (19) is satisfied. if wc compare Eq. (12) with Eq. (15), Eq. (15) is smaller by the value of As this value originated in the collision term, it is estimated that the collision has a kind of stabilizing effects to the link system. * 4. EXPERIMENT We experimentally investigated the possibility of the stable control of the system including collision. This experiment was carried out aker the stability analysis in the former chapter"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0003851_s0167-8922(08)70928-3-Figure7-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0003851_s0167-8922(08)70928-3-Figure7-1.png",
+        "caption": "Fig. 7 : Isotherms and heat flux lines in the f i lm",
+        "texts": [],
+        "surrounding_texts": [
+            "37\nwhere Ks i s the thermal conductivity of the\nshaft.\nA t the ends of the shaft i.e., for 5 = 2 1/2,\na free convection hypothesis i s assumed which\ngives :\nwhere 8.\nshaft.\n= hS L/Ks i s the Biot number for the\nIn the inlet zone, across the film, the\ninlet f i lm temperature %y) is assumed to vary parabolically between Ts and Tb which are\nrespectively the shaft and the bush temperatures\nat the inlet. The third value needed to determine\nthe temperature profile i s calculated from\nconservation equations by the following relation :\nIS\n- -\n3.3 Methods of Solution and Procedure\n3.3.1 Pressure Distribution i n the Fi lm and Heat\nTransfer in the Solids\ninduce some difficulties in the converging process.\nThus the continuity equation i s differentiated with respect to 7 which gives a second order\ndifferential equation :\nThis equation is then integrated using finite\ndifference with the following boundary conditions :\n3.3.3 Energy equation\nThe energy equation (9) looks l ike a differential\nparabolic' equation where e plays the role of the\ntime in the non-stationary problems. An implicit\nf inite difference method i s used to solve this\nequation and the Richtmyer technique [95] i s employed when the coefficient of ?T/a y2 i s\npositive i.e., when the fluid velocity i is\npositive.\nWhen a reversed flow occurs, at large\neccentricities for example, u i s negative i n the\ninlet zone of the f i lm and classical solutions do\nnot converge. I n this case, the fluid f i lm i s divided into two zones : these zones D1 and D2 are shown on Fig. 5. I n the first one which\n-",
+            "38\ncorresponds to the zone where i s positive, the imposed On the moving surface and heat flux energy equation is solved using the iterative continuity conditions are used on the fixed\nRichtmyer technique. The initial temperature surface. The fu l l lines are the isothermal lines values are the boundary conditions at the inlet and the dash lines are the heat lines.\nboundary and a given value at points situated in D2 near the boundary (fig. 6a). I n the second\nzone, the same iterative technique i s used but\nthe computing process i s started from the\nboundary using the init ial temperature at points\nsituated i n Dl near the boundary (fig. 6.b). The\nB O U N D A R Y l r l\nfor hl/h2 = 4\n3.3.4 Shaft and bush bearing deformations\nExperimental results have shown that thermal\ndisplacements of both the shaft and the bush\nshould be taken into account. I n this case, the\nf i lm thickness depends on both the eccentricity\nand the thermal displacements. But the axial\ntemperature variations i n both the shaft and the\nbush i s small and i t s effect on the thermal\ndisplacement can be neglected. I n this case the\nexpression of f i lm thickness can be written :\n- h = 1 + E cos 8 + 6 ( 0 )\nwhere ( 8 ) i s the difference between the\nnon-dimensional displacements of the bush and the\nshaft.\nIn this case, classical thermoelasticity\nrelations are used analytically for the shaft i n\nwhich the temperature i s axisymmetrical due to\nthe rotational speed. In the bush, the two-\ndimensional thermal displacements are determined\nusing finite element method. This computation i s\nmade only in the mid-plane of the bearing.",
+            "39\nTable I : Operat ing conditions 3.3 Procedure\nThe computational procedure is described by t h e flow chart of fig. 8. The global i t e ra t ive scheme is as follows : a n initial value for the Bearing length temperature field is given to calculate t h e fluid 2 0 0 ~ C = 145 Um viscosity at each point along and across t h e film. Reynolds equation is solved and the fluid velocity Lubricant viscosity at Journal radius R = 50 mm External bearing radius R2 = 100 mm L = 80 mm Radial c learance at speed range 1000<N<4500 'pm Load range lOOO<W<1O,OOO N U . = 0.028 Pa.s vector is calculated at al l points of t h e fluid. 40\"c I Energy equation and hea t t ransfer equations a r e k 0 = 3.287 then solved simultaneously in t h e fluid and in t h e k, = 0.777 Viscosity coeff ic ients kl = 3.064\n3 Lubricant density at ' 40\u00b0C p = 860 kg/m\nThe elast ic displacements a r e determined in Lubricant specif ic hea t co = 2000 J/kgOC both the shaf t and t h e bush which give new film Lubricant conductivity KO = 0.13 W/m\u00b0C thickness values. The i terat ive procedure is Air thermal conducstopped when at each point on t h e boundary tivit)' Ka = 0.025 W/m\u00b0C Bush thermal conbetween the film and t h e bush, t h e relat ive ductivity Kb = 250 W/m\u00b0C difference between two successive s teps is less Shaft conthan 0.1 percent. The convergence of this process Convection hea t t ransfer ,.\nsolids thus producing a new tempera ture field.\nductivity KS = 50 W/m\u00b0C\nis fairly rapid.\n50<hb<500 W/mL\"C\nhs = 100 W/m2\"C\nCoeff ic ient of t h e bush Convection hea t t ransfer\nCoeff ic ient of t h e shaf t Inlet lubricant tempe-\nAmbient tempera ture Ta 45\"s Inlet lubricant pressure P = 70.10 P a\nBush expansion coef fic ient x b = 17.10-6/0C Shaft expansion coeffi-\nra ture T = 40\u00b0C\nGroove angle 18\"\nc ien t x = 12.10-6/\"c\n3.4 Resul ts and discussion\nIn order to compare t h e theory and t h e experiments, t h e results were obtained for t h e operating conditions used by Ferron [74] in his experimental study. These conditions a r e given in\nTable 1. As t h e experimental bearing is housed in a protect ive box, t h e ambient tempera ture is above t h e inlet lubricant tempera ture (40\u00b0C). Depending on t h e operating conditions, t h e ambient tempera ture var ies between 43\u00b0C and 47\u00b0C. In these conditions t h e ambient tempera ture is assumed to be constant and equal to 45\u00b0C as shown in Table 1.\nFigures 9 and 10 produce t h e isothermal c h a r t in t h e journal bearing at a relative eccentr ic i ty of 0.8, a rotational speed of 2000 rpm and for a convection hea t t ransfer coeff ic ient of t h e bush of 80 W/m \"C. 2"
+        ]
+    },
+    {
+        "image_filename": "designv11_60_0000269_9781782421955.540-Figure1-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0000269_9781782421955.540-Figure1-1.png",
+        "caption": "Fig. 1 \u2013 Application of Camus\u2019 theorem for the generation of the conjugate profiles s and \u03c3 and the Euler-Savary graphical construction to give the equivalent link \u03a9\u03c3\u03a9s.",
+        "texts": [
+            " This approach could be further extended to non-circular bevel gears by realizing that the return circle becomes a spherical curve of third degree and the line of action takes the form of a great circle of the fundamental sphere (17-18). 2 CONJUGATE PROFILES GENERATION AND RETURN CIRCLE Camus\u2019 theorem can be applied to generate a pair of conjugate profiles from knowledge of the time-varying transmission ratio through the centrodes, which are traced by the instant center of rotation with respect to both rigid bodies that roll with respect to each other by reproducing their relative motion. Referring to Fig.1, when the relative motion of a kinematic pair is assigned through the centrodes \u03bb and l, which are in contact at the instant center of rotation P0 and undergo a pure-rolling motion, both conjugate profiles \u03c3 and s, which are in contact at point P, can be generated by applying Camus\u2019 theorem. In fact, when a suitable auxiliary centrode \u03b5 is chosen as a continuous curve tangent to both centrodes in P0 and attached to another curve \u03b7, both conjugate profiles \u03c3 and s can be generated as envelope curves of \u03b7 during the pure rolling motion of \u03b5 on \u03bb and \u03b5 on l, respectively"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0001197_978-3-658-12701-5-Figure6.20-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0001197_978-3-658-12701-5-Figure6.20-1.png",
+        "caption": "Figure 6.20: Industrial robot with rectangular path in testing cube",
+        "texts": [
+            " Amongst other tasks, paths are specified that are desired to be followed with constant velocity while measuring the path deviation. To select an appropriate task for the application of the method presented in this thesis, a rectangular path from [4] was chosen. As recommended in the standard, the corners of the rectangle are rounded to avoid sudden changes in the end-effector velocity. Clothoides are used to obtain corner roundings that provide continuous slope and curvature and thus enabling joint accelerations and end-effector accelerations to be continuous. Figure 6.19 depicts an example path. As shown in Figure 6.20 on the next page, the desired rectangular path is located in a test cube whose edges are aligned with the inertial coordinate frame of the robot. The plane containing the rectangular path was selected to be in a diagonal plane of the cube. While the position of the test cube relative to the inertial coordinate frame varies between the subtasks, the end-effector orientation will remain normal to the testing plane. The paths are chosen such that the linear axis of the robot can remain at \ud835\udc5e1 = 1.5 m to follow the path. It will be investigated whether enabling the linear axis results in trajectories with shorter end times. The edge length of the testing cube depicted in Figure 6.20 is chosen to be \ud835\udc4e = 0.63 m from the set of permitted sizes in [4]. As required, the dimensions of the rectangle are 80% of the corresponding edge lengths of the diagonal square in the testing cube. Since the standard does not explicitly require clothoides for rounding the rectangle\u2019s corners, the clothoides\u2019 parameters were freely chosen to be \ud835\udf11c = 0.1745 rad and \ud835\udc5fc = u\ufffd 10 = 0.063 m. In [8] the mathematical function describing clothoides is derived. Figure 6.21 on the next page shows one corner rounding p (\ud835\udc60) = \u239b\u239c \u239d \ud835\udc65 (\ud835\udc60) \ud835\udc66 (\ud835\udc60) \u239e\u239f \u23a0 , \ud835\udc60 \u2208 [0, 1] "
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0002865_978-94-015-9064-8_16-Figure7-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0002865_978-94-015-9064-8_16-Figure7-1.png",
+        "caption": "Fig. 7: Skeleton of the six DOF Double Circular-Triangle parallel manipulator",
+        "texts": [
+            " For a given six points, each on one triangles sides, there are at most, four different solutions for this line and, in our case, for the forward kinematics. The solution for the extension of the end-effector along the output link and its orientation, depends on the specific structure of the manipulator (axes of the U joints, location of the lead screw's nut and the screw pitch). The following solution of the direct kinematics assumes that the outer axis of the U joints at both upper and lower triangle are directed along the X axis of the triangle (see Fig. 7). Also, it is assumed that the helical joint (lead screw's nut) is placed at the upper triangle and a prismatic joint at the lower. The lead screw axis direction coincides with the Z axis of the end-effector, eZg , and is given by: (12) where rs = [xo -xg YO - yg H r is the vector from the lower triangle center to the upper one, H is the distance between the two triangles, and superscripts u and d denote for the upper and lower triangle, respectively. The rotation of the end-effector about the line connecting the two centers is determined by the rotation of the lower triangle which is connected through a U and prismatic joint to the output link"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0003219_s0092-8240(87)80026-5-Figure6-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0003219_s0092-8240(87)80026-5-Figure6-1.png",
+        "caption": "Figure 6. An illustration of the boundary layer and a volume element V.",
+        "texts": [
+            " An estimation of the thickness of the boundary layer and an approximated nutrient profile. This estimate is performed by a one-dimensional calculation because firstly only this case is feasible and secondly we expect no qualitative difference from the three-dimensional case. The boundary layer I is shown in Fig. 5 in its dependency on R. l approaches the value 1 very fast with increasing 3.3. A balance equation of the building material C. We try to formulate a balance equation for C in a small volume element V (Fig. 6). As local variables the arclength S and the curvature x are used, and R should be interpreted as 1/to (x is twice the mean curvature in three dimensions). The general material balance equation is dt C dv= ~ + CV. dA, ov (3) where O V is the surface of V and dA is a surface element. 3.4. A derivation of the local expression of the surface motion. The reaction~liffusion equation of C is substituted into the first term of the righthand side of (3) which then reads dt C dv = div grad C dv + tr dv (4) - T f v C d V + ~ o v C V n d A \" Now in addition to the 'adiabatic' approximation in the first equation of (1), we perform a second 'adiabatic' approximation in equation (4), assuming the left-hand side to be zero in the time scale of the slow motion of the surface"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0000648_j.proeng.2016.08.377-Figure2-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0000648_j.proeng.2016.08.377-Figure2-1.png",
+        "caption": "Fig. 2. Schematic of mesh Fig. 3. The experimental system Fig. 4. Result of experiment",
+        "texts": [
+            "47E+5 Tm K 1723 Tl K 1673 Cp J/(kg K) 500 \u03b2 1/K 1.0 E-5 \u03c3m N/m 1.8 Fig. 1. The calculation model k W/(m K) 24.2 At initial moments, in ABG region was the liquid metal and the melting temperature, other region was the solid metal and 300K. Calculation used unstructured grids. After the grid independent verification, grid dimension was 0.1 in V groove area, other area was 0.2. The grid scale refinement in proportion to the rate of 1.2, make sure the melting zone had sufficient number of the grid to describe the flow. Figure 2 represents the schematic of mesh. The grid model unit number is 13567. Use COMSOL[18] to simulation, and the time step was 0.001s. The next time step calculation continued until the relative residues of all physical variables less than 0.01 in the last time step simulation. The temperature of the molten pool was still higher than melting temperature after stop heat, molten pool kept changing, so the computation time is longer than 1.5s. Select the same parameters as the numerical calculation to do experiment"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0003733_isie.1998.711665-Figure2-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0003733_isie.1998.711665-Figure2-1.png",
+        "caption": "Figure 2: Five different bearing configurations",
+        "texts": [
+            " The limiting transient for the bearings is a 30kN sinusoidal disturbance at 4 Hz. A schematic diagram of the rotor is shown in Fig. 1. A small bearing is situated at the 1e:ft of Fig. 1, and its electromagnets have a designed maximum load of 6kN. Its purpose is mainly to keep the rotor steady whilst most of the disturbance forces are absorbed by the larger bearing which is situated close to the rotor's centre of gravity and has a designed maximum load of 30kN. 0-7803-4756-0/98/$10.00 1998 IEEE 573 earing c o n ~ ~ ~ r a t ~ o n s 4 A non-linear AMB model Figure 2 shows the 5 configurations of the bearings that are tested for fault tolerance here, labelled A-E. Each pole of each bearing has a single independently controlled coil wound on it, so the total number of poles and coils are the same. The 12 and 8 pole bearings are wound in two different ways, coupled alternating north south poles (NSNS) and independent pole pairs of alternating orientation (NNSS). Obviously the six pole bearing cannot be constructed in the NNSS configuration as the total number of poles must be divisible by 4",
+            " The inductance model receives voltages from the amplifier model and converts them to the currents in each magnet coil by the relationship, di d t V = Ri + L-, where R = coil resistance, L = coil inductance, i = current in the coil and V = voltage across the coil. Each coil produces a magnetomotive force, Ni, where N is the number of turns on the coil, and this drives the magnetic flux through a particular magnet pole. The poles through which the flux will return to complete the magnetic circuit are dependent on the configuration of the bearing. If the bearing is constructed from alternating north and south poles (eg. Figure 2B), half the flux will return through each neighbouring pole. However, if the bearing is constructed of alternating north, north, south, south poles (eg. Figure 2C), then the flux will return through only one of the neighbouring poles. This is modelled as a matrix mapping, M , performed on a vector of all the magnetomotive forces in the bearing. For the NSNS case, the entry for one pole would be [ O S , 1, O S ] , centered on the matrix diagonal, indicating that the flux has unit effect at its originating pole and half effect at the neighbouring ones, the other elements of the matrix row are zero. The NNSS case would result in a block diagonal matrix of two by two matrices of ones",
+            " The measure of performance loss is taken as the increase in displacement under maximum load with a failure as a percentage of the displacement under maximum load without a failure. 8 Concluding remarks In every experiment, the magnetic bearing system successfully maintained levitation during a single coil failure and re-activation, it also rejected the disturbance of Fig. 4 with the bearing in its degraded condition. Fig. 5 shows a plot of bearing structure against percentage loss in performance during a coil failure for the five bearing configurations of Fig. 2. The dotted line represents performance of the decentralised control scheme and the solid line represents the centralised scheme with failure compensation. As expected, the 6 pole bearing performs least well and the 12 pole bearing exhibits the greatest tolerance to coil failures. The centralised control approach is consistently more effective than the decentralised one, with approximately 50% less movement under maximum load for each bearing configuration. The performance improvement is greater using centralised control for both the NNSS bearings (C and E) than the equivalent NSNS bearings (B and D)"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0000703_chicc.2016.7555034-Figure1-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0000703_chicc.2016.7555034-Figure1-1.png",
+        "caption": "Fig. 1: The inertial and body coordinate system",
+        "texts": [
+            " The rest of the paper is organized as follows: in part 2, a mathematical model is established considering the gyroscopic moment and external disturbance. Part 3 presents the integral backstepping based controller design approach under a few assumptions. In part 4, simulations for different cases are implemented in the MATLAB/SIMULINK environment. Finally, A summary for the whole article is given in part 5. 2 Modeling of the quadrotor During the modeling of the quadrotor, we should consider two coordinate systems: the world coordinate system oe e e ex y z and the body coordinate system ob b b bx y z , as shown in Fig. 1. The world coordinate system can be treated as the inertial coordinate system. Define i as the i th rotor\u2019s speed; f as the total thrust of the four rotors; as the total torque of the four rotors. The rotation of the quadrotor is expressed in the Euler angle and the body angular velocity, which are denoted as ( , , )T , ( , , )Tx y z . The yaw angle, pitching angle and roll angle are denoted as , , . The translation of the quadrotor is expressed using the position and velocity of the body\u2019s mass center, which are denoted as ( , , )Tx y z , ( , , )Tx y zv v v v "
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0001454_20140313-3-in-3024.00185-Figure3-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0001454_20140313-3-in-3024.00185-Figure3-1.png",
+        "caption": "Fig. 3. \u019f1 and \u019f2 are easily calculated as angles from the vertical axis (Z-axis) and angle from the horizontal axis (Yaxis).",
+        "texts": [],
+        "surrounding_texts": [
+            "We considered several methods to calculate joint angles. The first option was to use Inverse Kinematic Algorithm. Since the manipulator is redundant, the algorithms provided analytical solutions to the equations. [1] gives the DenavitHartenberg (D-H) parameters for the human arm and also describes an optimized Inverse Kinematics Algorithm for finding out the joint angles and also suggests an optimization for the same. We computed the transformation matrices and also implemented this algorithm but found it too slow. [2] uses a Jacobian control method for redundant systems which implements a fast closed loop algorithm. The main reason that these algorithms are much slower than ours is that they assumes knowledge of only the end effector position and orientation, whereas with the Kinect Sensor, we also have the 3D global coordinates of each of the joint position, which helps us in easily finding out the joint angles. In addition to being slow, these algorithms employ a fairly complex algorithm for an otherwise simple problem. We researched some other papers to look into other techniques and know more about Inverse Kinematics and the manipulator {[3], [4], [5], [6]}. Our own algorithm derives inspiration from [7]. In this, the authors have tried to simplify the Inverse Kinematics problem for various cases when one of the joint positions is known. Now with our Kinect Sensor, we know all the joint positions, using analytical 3D geometry, we have computed all the joint angles. [8] provides all the information on Amtec PowerCube manipulator dynamics and control."
+        ]
+    },
+    {
+        "image_filename": "designv11_60_0002530_881621-Figure8-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0002530_881621-Figure8-1.png",
+        "caption": "Fig. 8 Measuring method of piston static deformation",
+        "texts": [
+            "l side Cylinder wall 50ym \"Nv/~\\!L0tJm Antithrust side [a) Symbols Cylinder wall (b) Piston lateral motion . _~ r - / v \u2122 . ..A T y ' . \u2022 50pm B T B (c) Skirt deforemation {XTI-XT2) Fig. 6 Typical piston lateral motion diagram (Full load , 2000 r/min, TC=363K) 881621 5 i.e. Xm-j-X,,, and demonstrates that its maximum value was about 30 ym. The deformation data shown in Figure 7(a) were measured approximately when static thrust load was applied to the test piston at T, and A.,, and at T~ and A, by means of the method shown in Figure 8. Based on these data, piston deformation during 2000 r/min fullload engine operation was calculated (Fig.6(b)) and compared with measured data. Since calculated piston clearance at the location T,A, at this operating condition was 20 ym, and since the measured piston deformation was 30 urn, the measured piston clearance (50 ym) agreed with the sum of these two values. It should be pointed out that in order to confirm the accuracy of the measured data, a dynamic calibration was conducted and compared with the static calibration conducted outside the engine"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0001235_b978-0-08-099425-3.00005-4-Figure5.59-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0001235_b978-0-08-099425-3.00005-4-Figure5.59-1.png",
+        "caption": "FIGURE 5.59",
+        "texts": [
+            " The implementation of these computations as a tyre model with an MBS program is best described using the full threedimensional vector approach outlined in Chapter 2. The following description is based on the methods used in MSC.ADAMS but is applicable to any vehicle simulation model requiring tyre force and moment input. As a starting point the tyre can be modelled using the input radii R1 and R2 as shown in Figure 5.58. Using the tyre model geometry based on a torus it is possible to determine the geometric outputs that are used in the subsequent force and moment calculations. Consider first the view in Figure 5.59 looking along the wheel plane at the tyre inclined on a flat road surface. The vector {Us} is a unit vector acting along the spin axis of the tyre. The vector {Ur} is a unit vector that is normal to the road surface and passes through the centre of the tyre carcass at C. The contact point P between the tire and the surface of the road is determined as the point at which the vector {Ur} intersects the road surface. For the purposes of this document it is assumed the road is flat and only one point of contact occurs"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0003689_robot.2001.932951-Figure1-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0003689_robot.2001.932951-Figure1-1.png",
+        "caption": "Fig. 1. Contact kinematics with virtual frictional joint",
+        "texts": [
+            "R2 and the moment mo E RI, by W O = G f , (2) where the grasp matrix G is of the form G = [G1 G2 . . . GI(] E S@x2A'f for GI, E p ' i X 2 n z k . 2.1 Statical decomposition of contact force 1 1 . . . > > K i = 1,. . . , n k ) , its velocity can be described by Denoting each contact position by xkl E E2 ( k = % = [J r] (i) for the joint angle vector q = (q;, &, . . . , qg)T E RN and the frictional degrees-of-freedom p = (pT, p;, . . . , pg)T E RR\". Note that cadi eleinent pk i E E' denotes the displacement to the cont,a.ct point Ck2. from the coordinate frame { a k ( i ) } . (See Fig. 1.) Assuming that the inmtrix in (3) is square which is equivalent to the condition (A2) Ad = N , the total contact force is decomposed into the one due to the joint torque and the other due to the frictional force as follows: f = n,r + q 6 , (4) where [II, II,] = [J !RA' for E k E W k denotes the frictional force associated to the frictional degrees-of-freedom pk. I?] -\" and 6 = ( E T , [T,. . . , E Applying the decomposed force by (4) to the object wrench equation (2) yields W O = fl,r + OtJ, (5) where fl, = GIIT and flt = GIIc"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0000986_iciev.2016.7760019-Figure8-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0000986_iciev.2016.7760019-Figure8-1.png",
+        "caption": "Figure 8. Experimental setup [15]",
+        "texts": [
+            " 7 clearly provide the outer race fault information in rolling element bearing i.e. BPFO is visible at 3.43 Hz (circle with magenta color). Hilbert transforms in Fig. 7(a) provide higher fault magnitude in compare to the square envelope in Fig. 7(b). V. EXPERIMENTAL RESULT ANALYSIS In this section, real experimental fault signal (vibration signal) are analyzed to show the effectiveness of the proposed method. A. Test Rig A self-designed test rig [15] consist of motor, pulleys, belt, shaft, and fan with changeable blade pitch angle is shown in Fig. 8(a). For this experiment, two-pole induction motor (six 0.5 kW, 60 Hz) is used and the motor was operated under normal condition as a benchmark for comparison with other faulty motors. In this paper, bearing outer race fault is considered and faulty part of bearing (a spalling on the outer raceway) is shown in Fig. 8(b). The BPFO of this bearing is 175 Hz and the sampling frequency of the data acquisition unit was 7.68 kHz in this experiment. B. Results Analysis Fig. 9(a) Vibration signal of rolling elements from accelerometer and Fig. 9(b) indicates corresponding frequency spectrum of vibration signal which contains no diagnostic information i.e. BPFO is not visible at 175 Hz. After applying our proposed method, the Frequency spectrum of the vibration signal is shown in Fig. 10. Fig. 10(a) and Fig. 10(b) are square envelope spectrum and Hilbert-based envelope spectrum respectively"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0000963_detc2016-59644-Figure5-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0000963_detc2016-59644-Figure5-1.png",
+        "caption": "FIGURE 5: Numerical domain utilized for analyses including Gaussian spot initially at (x0,y0,0) and serpentine path with pitch, \u03b4 . Y1\u2212Y4 are data extraction points for information shown in Figs. (7), (9) and (10).",
+        "texts": [
+            " It is noted that this investigation includes a comparison between models that took phase transition into account, and those that did not. Temperature dependent properties of solid Ti-6Al-4V are shown in Figs. 2, 3 and 4 according to [28], while molten Ti-6Al-4V properties were obtained from [29]. For non-phase transition predictions, Cp is a function of temperature only. 2 Copyright \u00a9 2016 by ASME Downloaded From: http://proceedings.asmedigitalcollection.asme.org/pdfaccess.ashx?url=/data/conferences/asmep/90679/ on 03/06/2017 Terms of Use: http://www.asme.org/about-asme/terms-of-use Figure 5 details the domain and boundary conditions applied with uniform initial temperature for a Ti-6Al-4V domain. No convection or radiation to ambient is taken into account, with all of the free domain surfaces specified as thermally insulated. Identical eight-node hexahedral element meshes were used in all presented simulations. The discretization and solution of the resulting system of equations was implemented on the \u201cCOMSOL multiphysics\u201d computational system. An important feature that allows the localized mesh deformation to capture the deposited material track due to the powder jet deposition, is the \u201cdeforming geometry\u201d feature enabled in COMSOL",
+            " This phenomenon is easily observed for the case of pitch= r0 where the overlap occurs along the entire length of the track. The second observation from Fig. 6 is the fact that as the scan pitch decreases, thermal accumulation becomes significant, and manifests as a persistent elevated temperature in regions adjacent to the active scan path. Figure 7 provides further insight into this scan pitch effect as it reveals the temperature evolution for all three scan pitches at four discrete points (Y1 \u2212Y4) located at the midpoint of each major scan (see Fig. 5). These plots reveal that in addition to persistent elevated temperatures after the laser has passed, there is a pre and post-scan effect whereby adjacent points that have yet to be scanned directly, or adjacent points that had just been scanned exhibit local temperature rises. These temperature rises are particularly dramatic for the single radius pitched case, detectable in the two-radius and completely absent from the four-radius. Figure 8 shows the profiles of the associated clad lines predicted by the model"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0000007_eleco.2015.7394494-Figure4-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0000007_eleco.2015.7394494-Figure4-1.png",
+        "caption": "Figure 4. Travel axis",
+        "texts": [
+            " (In Figure 7, NH: negative high, NL: negative low, Z: zero, PL: positive low and PH: positive high.) Free body diagram of pitch axis is shown in Figure 3. The pitch model symbol is (}. I (J is the moment of inertia of the system about pitch axis. Mathematical equation of the pitch axis is given in equation (2). .. Lh1(J(} = -MtO--l'-cos\u00ab(} - 8h ) cos Uh Lh+MbO-l'-cos\u00ab(} + 8h ) (2)cos Uh -TJ(J8 +KmLh(Vt - Vb) 2.2. Mathematical Model of Travel axis (3) Free body diagram of travel axis is shown in Figure 4. The elevation model symbol is cp. l<t> is the moment of inertia of the system about travel axis. Mathematical equation of travel axis is given below in equation (3). l<t>CP = -TJ<t>cP - KmLaCVt +Vb)sin (} 3. Design of a Fuzzy PID Controller for 3-Dof Helicopter Fuzzy Pill control mechanism obtained in order to control the system is shown in Figure 5 Three separate fuzzy Pill controllers are used in order to control elevation, pitch and travel axes of the helicopter. As it can be seen in Figure 6, inputs of Pill controllers are determined as position error and angular The input universe of discourse for position error is [-25, +25], hence scaling factor of position error input (Ke) is 0"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0000928_1.4035169-Figure1-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0000928_1.4035169-Figure1-1.png",
+        "caption": "Fig. 1 The mobile robot and its model",
+        "texts": [
+            " Specifically, an approximate inverse solution combined with a trajectory planning method is presented such that the dominant order solution satisfies the no-slip condition, while the first-order solution provides corrections due to slipping. The algorithm is applied to an in-house mobile robot under conditions with varying slipping of the wheels. Experimental results using a planar mobile robot verify the proposed method by showing improved tracking when slipping is included in the model. The planar kinematic model of a mobile robot with two rear driving wheels and a front spherical ball bearing is shown in Fig. 1. The robot moves on and rotates within a horizontal plane surface and thus, ignoring the ball bearing dynamics due to its negligible mass, has three degrees-of-freedom (DOF). In order to represent the robot motion, one may assign a body-fixed reference frame euev to the robot center of mass (CM) and measure its position and orientation with respect to an inertial reference frame XY. The location of the CM with respect to XY is \u00f0 x; y\u00de and h 1Corresponding author. Contributed by the Dynamic Systems Division of ASME for publication in the JOURNAL OF DYNAMIC SYSTEMS, MEASUREMENT, AND CONTROL"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0003645_i2002-10009-1-Figure7-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0003645_i2002-10009-1-Figure7-1.png",
+        "caption": "Fig. 7. Local geometry of the vortex lines for two neighboring TGBs, in a situation of least curvature energy: a) projection of the vortex lines and their associated normal lines from Ti+1 to Ti; b) quadrangle of geodesic lines on Ti+1 or Ti to which Gauss-Bonnet theorem is applied.",
+        "texts": [
+            " In principle, also, TGBs are no longer present in the NL* phase, which should at least partially decouple the curvature of the NL* cholesteric layers from the screw dislocations, a process that can even be improved by a local adjustment of the distance between the dislocations. Consider a distorted sample of the TGBA phase. By condition of curvature of least energy, we mean in the following that any two neighboring TGBs of the sample, Ti and Ti+1, are parallel surfaces and that the vortex lines form on Ti and Ti+1 two sets si, si+1 of curves at a constant distance ld. We require also that the two sets be at a fixed angle \u03c9, that is the angle of the local projection of one set onto the other, along the (common) normal to the surfaces, Figure 7a. In other words, the curved TGBs reproduce locally the geometry of TGBs in the ground state of the TGBA phase. The curvature energy is then limited to the energy of distortion of the SmA layers inside each slab (see below); there is no variation in the density of vortex lines, and no additional edge lines. This is a very special geometry: because each set is formed of equidistant lines. The lines that are normal to si, si+1, are geodesic lines of Ti, Ti+1. Because of the property of conformal projection, which extends from the sets of vortices to the sets of geodesic lines, there are two sets of geodesic lines on Ti and Ti+1, which are at angles \u03c9 and (\u03c0 \u2212 \u03c9). Consider, Figure 7b, a quadrangle \u2202\u0393 built on both sets of geodesics, on Ti or Ti+1, bounding an area \u0393 . Applying Gauss-Bonnet theorem to this figure, one gets\u222b \u2202\u0393 ds/\u03c1g + \u03a3k\u03b8k = 2\u03c0 \u2212 \u222b \u222b \u0393 Gd\u03a3, (3) with 1/\u03c1g being the geodesic curvature, and \u03b8k the inner angles of the quadrangle, equal to \u03c9 or (\u03c0 \u2212 \u03c9), namely \u03a3k\u03b8k = 2\u03c0; G = \u03c31\u03c32 is the Gaussian curvature. Since the edges of the quadrangle are geodesic lines, 1/\u03c1g \u2261 0. Therefore, the Gaussian curvature \u03c31\u03c32 sums up to zero; since the Gauss-Bonnet theorem can be applied to a quadrangle of the same type as small as desired, one gets G = 0 on any TGB of the sample which satisfies the above conditions of least curvature energy"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0000878_978-3-319-47614-8-Figure6.9-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0000878_978-3-319-47614-8-Figure6.9-1.png",
+        "caption": "Fig. 6.9 A tapered beam",
+        "texts": [
+            " Let the assumed solution be \u1ef9 and we substitute this in the differential equation, we get an error e in the differential equation. 6.7 Galerkin Method for Tapered Beams 185 e \u00bc L~y \u00f06:84\u00de Then Zx2 x1 e~ydx \u00bc 0 i:e:;Zx2 x1 L~y\u00f0 \u00de~ydx \u00bc 0 \u00f06:85\u00de This is also called orthogonalization of the error with the assumed solution. Let\u2019s consider the following problem. A cantilever beam of length L carries a uniformly distributed load q(x) with rectangular cross-section tapering down from fixed end to free end as shown in Fig. 6.9 defined by the following relations for breadth b and thickness t with X \u00bc x L b \u00bc b0 1 aX\u00f0 \u00de t \u00bc t0 1 bX\u00f0 \u00de I \u00bc I0 1 aX\u00f0 \u00de 1 bX\u00f0 \u00de3 \u00f06:86\u00de 186 6 Bending of a Beam where I0 \u00bc b0t30 12 The differential equation is d2 dx2 EI d2w dx2 \u00bc q x\u00f0 \u00de Using (6.76) d2 dX2 \u00f01 aX\u00de\u00f01 bX\u00de3 d 2w dX2 \u00bc qL4 EI0 \u00f06:87\u00de As before the assumed solution is taken as w X\u00f0 \u00de \u00bc a 6X2 4X3 \u00feX4 d2w dX2 \u00bc a L2 12 24X\u00fe 12X2 \u00f06:36\u00de Substituting, the error e in the differential equation is e \u00bc d2 dX2 1 aX\u00f0 \u00de 1 bX\u00f0 \u00de3 a L2 12 24X\u00fe 12X2 h i qL4 EI0 \u00f06:88\u00de Performing the differentiation e \u00bc 12a L4 2 1 aX\u00f0 \u00de 1 bX\u00f0 \u00de3 \u00fe 12b 1 X\u00f0 \u00de 1 aX\u00f0 \u00de 1 bX2 \u00fe 4a 1 X\u00f0 \u00de 1 bX\u00f0 \u00de3 \u00fe 6b2 1 X\u00f0 \u00de 1 bX\u00f0 \u00de 1 2X\u00feX2 \u00fe 6ab 1 bX\u00f0 \u00de2 1 2X\u00feX2 2 664 3 775 qL4 EI0 \u00f06:89\u00de Orthogonalizing as per Galerkin method, we have Z1 0 12a L4 2 1 aX\u00f0 \u00de 1 bX\u00f0 \u00de3 \u00fe 12b 1 X\u00f0 \u00de 1 aX\u00f0 \u00de 1 bX2\u00f0 \u00de \u00fe 4a 1 aX\u00f0 \u00de 1 bX\u00f0 \u00de3 \u00fe 6b2 1 aX\u00f0 \u00de 1 bX\u00f0 \u00de 1 2X\u00feX2\u00f0 \u00de \u00fe 6ab 1 bX\u00f0 \u00de2 1 2X \u00feX2\u00f0 \u00de 2 64 3 75 qL4 EI0 8>< >: 9>= >; 6X2 4X3 \u00feX4 dX \u00bc 0 \u00f06:90\u00de There are six integrals in the above, they are evaluated as follows: 6"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0001000_1.4035601-Figure2-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0001000_1.4035601-Figure2-1.png",
+        "caption": "Fig. 2 Circumferential pockets/Rayleigh step",
+        "texts": [
+            " There are strengths and weaknesses for using either of these orientations. However, the focus of these investigations is on a radially mounted seal. Advantages for this orientation include being more applicable to turbine split line designs, having to cope with less movement perpendicular to the flow direction, and being easier to scale at various diameters. There are many possible groove shapes, several of which originate from bearing design. These include wedges (flat and tilted) (Fig. 1), circumferential pockets and Rayleigh steps (Fig. 2), inclined grooves (Fig. 3), and herringbone grooves (Fig. 4). Wedges are reported by Dhagat et al. [32] and the tilted variety is recorded by Galimutti et al. [33]. Pockets are also shown by Dhagat et al. [32] and Rayleigh steps are described by Cheng and Wilcock [34]. Herringbone grooves are assessed by several authors although there is no general agreement on whether the center land region should be included or not. The center land is omitted in Dhagat et al. [32] and Liu et al. [35] but included in Cheng and Wilcock [34] and Proctor and Delgado [29]"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0000524_978-4-431-55013-6-Figure5.1-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0000524_978-4-431-55013-6-Figure5.1-1.png",
+        "caption": "Fig. 5.1 Bifurcation diagram on fixed and periodic points in the H\u00e9non map, MLLE, and blue horizontal and red diagonal curves corresponding to parameter variation without and with control",
+        "texts": [
+            " To evaluate whether the proposed parametric controller is effective, we carried out several experiments for stable periodic points observed in the H\u00e9non map [7]. The dynamics of the H\u00e9non map is described as x1(t + 1) = 1 + x2(t) \u2212 p(t) \u00b7 x1(t) 2, (5.9a) x2(t + 1) = q(t) \u00b7 x1(t), (5.9b) where x1 and x2 are state variables and t is the discrete time. We here assumed that p and q correspond to out-of-control and control parameters, respectively. In the following experiments, we set T = 500, \u03b7 = 0.1, and \u03bb\u2217 = \u22120.2 in (5.4) and (5.8). Before carrying out experiments, we analyzed bifurcations on fixed and periodic points observed in (5.9). As shown in Fig. 5.1, we found a fixed point, n-periodic points (n = 2, 4, 8), and their period-doubling bifurcations where the solid curve with Pn represents the set of bifurcation points of the n-periodic point. The stable fixed point is present in the left-hand-side parameter regions of the curve P1 and the stable n-periodic point exists in the parameter regions surrounded by the curves of P n 2 and Pn . The MLLE on the fixed and periodic points is indicated in color, for example, the color in the parameter regions surrounded by the curves of P1 and P2 shows the MLLE on the stable two-periodic point. The relationship between the MLLE and colors is shown in the right bar graph. We note that these analyses are not necessary to avoid bifurcations using our controller, i.e., it was carried out only to demonstrate whether bifurcation points are avoided in space of system parameters. When we set (p(0), q(0)) = (0.5, 0.3) corresponding to the point \u201ca\u201d in Fig. 5.1 and (x1(0), x2(0)) = (1.43, 0.0), the two-periodic point was observed in a steady state. By decreasing the value of p with 0.0015 every T along the blue horizontal line from the initial point \u201ca\u201d, the two-periodic point bifurcated on the curve P1 and instead the fixed point appeared at t 93T as shown by the blue trajectory of x1 in Fig. 5.2a. To avoid the period-doubling bifurcation, the proposed controller adjusted the value of q so as to keep \u03bb \u03bb\u2217 after t = 42T (Fig. 5.2b). The trajectory of the controlled parameter is also shown as the red diagonal curve branching from the blue horizontal line with the point \u201ca\u201d in Fig. 5.1. Consequently, the stable two-periodic point could be observed for the duration of 0 \u2264 t \u2264 100T without bifurcating. We also analyzed avoiding the period-doubling bifurcation of the stable fourperiodic point. The initial values were set to (x1(0), x2(0)) = (1.04,\u22120.18) and (p(0), q(0)) = (0.95, 0.3) corresponding to the point \u201cb\u201d in Fig. 5.1. When we changed the value of p along the blue horizontal line starting from the point \u201cb\u201d, we observed the eight-periodic points and a chaotic state caused by a cascade of perioddoubling bifurcations across the curves of P4 and beyond (Figs. 5.1 and 5.3a).Hence, the stable four-periodic point bifurcated and vanished at t 70T owing to its perioddoubling bifurcation curve P4. In contrast, the red diagonal curve branching from the blue horizontal line with the point \u201cb\u201d in Fig. 5.1 indicated that the proposed controller was used to avoid the bifurcation curve of P4. As the results, as shown in Fig. 5.3b, we could observe the four-periodic point for the duration of 0 \u2264 t \u2264 100T . In this chapter, we presented a parametric control system to avoid bifurcations of stable periodic points in nonlinear discrete-time dynamical systems with parameter variation. The parameter updating of our controller is theoretically derived from the minimization of an objective function with respect to the MLLE"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0002566_88-gt-92-Figure1-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0002566_88-gt-92-Figure1-1.png",
+        "caption": "Fig. 1 : The meridional profile of the least sensitive impeller with a specific speed of n, = 0.57",
+        "texts": [
+            " The parameter that is of interest is the degradation of overall efficiency with increasing tip clearance which is correlated with specific speed, and the validity of the correlations is examined in light of variables and effects which specific speed does not account for, and which in turn may affect the clearance loss mechanism. In order to isolate and treat the clearance flow as a superimposition on the main channel flow; the impellers, which were initially fully shrouded two of which are shown in fig. 1 and 2, were tested with full shroud to simulate zero clearance before the front shroud was machined off. This performance was compared with the performance extrapolated to zero for the impellers in the half shrouded (open-impeller) configuration, which showed good agreement in most cases. But with increasing clearance, the effect on the performances was not as simple as hoped for. LIMITATIONS OF SPECIFIC SPEED Due to the absence of a unified review of the various methods which have been proposed for the estimation of tip clearance losses it', centrifugal impellers, a likely parameter that may provide the link was sought",
+            " Head and torque are also related since the blade loading across an impeller is directly proportional to the torque. Therefore it was a logical approach to seek a correlation involving the most important consequence of tip clearance, overall efficiency deterioration and specific speed. The efficiency loss for each clearance was obtained by substracting from a hypothetical zero clearance on = na=o- nA, which was in turn found by extrapolation from the efficiency/clearance plot (1) . A sensitivity factor y was defined as the ratio of efficiency loss to clearance ratio, y = An /A . Fig. 1 and 2 show the meridional profile of the two impellers with opposite extreme sensitivities. The impeller with specific speed n s = 0.83, fig. 2, was the most sensitive of the tested series; especially at small clearances it was highly sensitive. On the contrary the impeller with specific speed n s = 0.57 showed the least sensitivity; especially at small clearances it was insensitive. This sensitivity trend is supported by Senoo et al. (18) , who investigated two centrifugal blower impellers of specific speed n = 0"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0002728_112515.112577-Figure12-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0002728_112515.112577-Figure12-1.png",
+        "caption": "Fig. 12 Algebraic constraints fm~isl contact.",
+        "texts": [],
+        "surrounding_texts": [
+            "(3)Againstx AF= A[dc,0,0,0,&,6z]. If Vi k a vertex of the internal feature whose Y and Z coordinates in the CF are\nViY and Viz respectively, then the X coordinate Vi. of Vi in the CF is calculated as follows (see also Fig. 9):\nVix=Viz&-Viy &+& (17)\nTherefore, the following inequalities must be satisfied:\nh(vi. > Vj.) A(vi. -~. < \u20182\u2018TF) Vi j i *j \u2018~-vbuA -Vk -TF<dx STF+\u2014 2 i (18)\n5.1 Equation Constraints In the following explanation, we refer to two features concerning a certain contact by FI and FE. FI k the feature given by origin(CVs) where CVS are CVS of the internal feature. The other feature is referred to as FE. Let fi = ff - pos(FJ), fE = ff-pos(FE), pI = p-pos(owner(Ft)) and pE = p - pos(owner(FE)). fr and fE are defined in consideration of shape errors as explained in section 4 and may include some\ndifferential motion variables representing shape errors. Accordingly, the following algebraic constraint can be de rived\nW *fE*PE*PI_~ *fI-~ =W *tc (22)\nwhere tcexpresses the positions of the axes of FE in the coordinate frame of FI.\nIf we consider the CF, we can represent tcby the fol-\nlowing transformation:\ntC = tE\u201dl * tl (23)\nwhere tI and L?express the position of the axes of the CF with respect to the coordinate frames of FI and FE.\nWe approximate tland t,?by using the form features\nwith nominal geometry. t] is approximated by tfwhich is a transformation expressing the position of the CF with respect to the frame of the nominal feature of FI. tE is also approximated by tE\u2019 which is derived based on the nominal feature of FE. When these approximations are applied, (23) is changed to\ncc = t~\u2019-l* (Z + Ac) * U\u2019 (24)\nwhere Ac is a differential motion matrix at the CF.\nThe position of a part with influences of shape errors can be expressed as small displacements and rotations from its nominal position. Therefore, we can apply differential approximations to replace pr andpE with\np =pI\u2019 * (Z + AI) (25) p,z =pE\u2019 * (Z + AE) (26)\nwhere pI\u2019 and pE\u2019 represent the nominal positions of the parts. AI and AE are differential motion matrices for pt\u2019 and pE\u2019. Consequently, the following equation is derived based on (22) - (26):\ntE\u2019 *fE * pE\u2019 * AE -tI\u2019 *fI*pI\u2019 *AI\n=Ac*tl\u2019*f[*pl\u2019 (27)\nAc is defined according to the contact type:\n(l)Agairm: Ac = AIO,dY,&&,&,&] (see Fig. 10).\n(2)Agaimtz: Ac = A[0,4,&&,0,&] (see Fig. 11).",
+            "(5)Fim2: Constraints forfirw arerepeatedly given for two\nCVS, Suppose CFJ and CF2 represent CFS at two CVS. Let tll\u2019,lEl\u2019 and ACI are transformations for CFI and tn\u2019, tm\u2019and Ac2 are for CF2. If the direction of the X axis of\nCF2 in the world is in the negative direction of the X axis of CFI, m\u2018 and tm\u2019 are changed to M * tn\u2019 and M * tE2 \u2018 . ACI = A[&,O,&, &I, C5yl,6ZI] and Ac2 = A [&,0,42,&,&2,&2] where dil = h, & = &, 61 = &2, and & = - &. If the angle between the Y axes of CFJ and CF2 is q, the following constraints are additionally given (see also Fig. 13):\ne= ~,d=21r-r\u2019lc0s6\n(28)\nwhere lx is signed distance from the origin of CF2 to the origin of CFI along the X axis of CF1.\n(6)Fim: Ac = A[&O,d,&,O,O] (see Fig. 14).\n5.2 Inequality Constraints When CVS are touching the extended feature, other vertices of the internal feature must be locate above the extended feature to avoid the interference (see Fig. 3). This condition is interpreted as inequalities.\nSuppose a vertex V of the internal feature and FI = origin({V ) ). FE is determined in the same manner for the equation constraints. If we consider the VF of V as defined in 3.3, we can represent tc of (23) by\ntC = tVE-l * tVI (29)\nwhere tw and tvE express the position of the axes of the VF with respect to the frames of FI and FE.\nWe approximate tvI and tvE by using the form features with nominal geometry and use tvl\u2019 and WE\u2019 instead. Now (29) can be approximated by\ntc = tVE\u2019-l * (Z + Av) * tVI\u2019 (30)\nwhere Av is a differential motion matrix at the VF.\n,, , ,x,,\n. XJ Ft\nCF\nvAc = A[d.,O&&,O,O]\n6. SYNTHESIS OF CONTACT COMBINATIONS A part has several contacts with other parts in an assembly. In this section, we give a method to synthesize all possible combinations of such contacts based on the amlysis of the part motion.\n6.1 Assumptions for Contact Synthesis The following three constraints are assumed for synthesizing the contact combination.\n(l)A part is assumed to be positioned in stable condition.\nWe define a predicate + as re13 e reb meaning retz is a",
+            "stabler contact than reh. Thus, aguinsfl d ugainsn e agahsts and fits] G fits2 G fitsJ. When two sets S i and Sj of contacts are given, we can ignore Si if any contact in Si is smaller or equal in G to its corresponding con-\ntact in Sj.\n(2)A contact is assumed to be realized in a step by step\nmanner. For example, before realizing three vertices contact (againm) between Ff and F\u20192, two vertices contact (agtukm) must be realized, and before realizing two vertex contact, single vertex contact (againstl) must be realized between the same feature pair. In the same way, be fore realizingjitw or fits2, fitsl must be realized.\n(3)The displacement or rotation of a part from its nominal\nposition to achieve a certain contact should be small enough. This assumption restricts the geometry of a feature to make a contact when a certain relative motion is possible. For example, if a flat face FI is permitted to rotate about Ax, itcan realize againm contact with an edge E] by a small rotation (see Fig, 15(a)), however, it must rotate nearly 90 degrees to realize a contact with E2 (b).\n6.2 Primary Contact When a force is applied somewhem on a part, according to where it is applied and its direction, some specific contacts are more easily realized than others.\nIt is a difficult problem to determine proper contacts\nautomatically by evaluating the effect of the force. Instead, we permit the user to specify some contacts cakxl primary contacts explicitly on account of the force effect. The system determines other unspecified contacts automatically taking the primary contacts into consideration.\n6.3 Analysis of Part Motion Our contact combination synthesis is based on the analysis of relative motion of parts in an assembly.\nWhen the initial position of part P is given asp, the\nsmatl change of the part position top + @ can be expressai asp * (1 + AP) by using the differential motion matrix AP. Thus AP can be considered to represent the part motion.\nWhen an assembly composed by n parts Pi (1 S i S n)\nis given, we assign a differential molion matrix A [&i,c$i,dZi,&i, &i,&i] toeach Pi. If there is a pm in the as-\nsembly whose position is fixed in the world, we give null matrix O instead for its differential motion.\nThese parts may have several contacts each other. For each contac~ we can determine FI and FE by using the CVS realizing the contact as explained in 5.1. Based on the equa-\ntion (27), we can derive a linear algebraic constraint between differential motion matrices of parts mwte@O &l OWW(FE). We can ignore small shape errors for calculating the part motion. Therefore, we use the following equation instead of (27):\ntE\u2019 *fE\u2019 * PE\u2019 * AE - tl\u2019 * fi\u2019 * pI\u2019 * AI\n= AC * tI\u2019 *fI\u2019 * JX\u2019 {32)\nwhere fI\u2019 and fE\u2019 represent nominal positions of form features FI and FE in their owner part\u2019s frame. Other variables are the same as the equation (27). All the transformations except A], AE and Ac are calculated by using the nominal geometry of the parts.\nAccordingly, a system of linear equations relating differential motion matrices of all parts in the assembly is constructed. We can eliminate some variables of A1, AE and A c by applying algebraic variable elimination method (for example, Buchberger algorithm [9]) on the system. If we eiiminate such variables as much as possible, we can represent differential motion matrices AP of part P by some limited number of motion variables, which represents the possible motion of P in the assembly.\n6.4 Motion Constraint for Realizing a Contact We approximate the real contact as the contacts of some vertices of the internal feature and its corresponding extended feature. Here we consider a motion constraint to realize such vertex contacts. Suppose Vs are the vertices of the internal feature proposed to approximate a contact between FJ and F2. If F] = origin, we assign FI = FI and FE = F2, otherwise, FI = F2 and FE = FI, According to the assumption (2) in 6.1, we assume that a set of vertices CVS c Vs are now retilzing contacts with the extended feature. As explained in 6.3, by using currently available contacts, which include the contacts realized by CVS, we can compose a set of linear equations relating differential motion matrices of the parts based on (32) and can derive the possible motion of owner(F~) and owner(F@ in the assembly. NOW we select a vertex v ~ (vs - CV~) and consider whether V can achieve an additional contact. If we consider the VF of V, we can derive the following constraint equation based on (3 1).\ntVI?\u2019 *fE\u2019 * pE\u2019 * AE * pi\u2019-l *fI\u2019-l * tvI\u2019-l\n- tv{\u2019 *fi\u2019 * @ * Ar * PI\u2019-l * fI\u2019-l * tvi\u2019-~ = AV (33)\nwhere tvl\u2019, WE\u2019, pI\u2019 and pE\u2019 are the same as the equation (31).fi\u2019 and fE\u2019 are used instead offl andf!?, which represent nominal positions of FI and FE in their owner part\u2019s fkune. These transformations are all determined by using the solid model. AE and Al are differential motion matrices of"
+        ]
+    },
+    {
+        "image_filename": "designv11_60_0000928_1.4035169-Figure2-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0000928_1.4035169-Figure2-1.png",
+        "caption": "Fig. 2 Transitional and target circular paths",
+        "texts": [
+            " However, we now turn our attention to deriving the radius rt and rate of speed ah for such trajectories such that all the boundary conditions are satisfied. Consider the case where the robot starts from rest at the origin with a given orientation h0 and the target reference trajectory is rotating at constant angular speed Xc around a circle of radius rc centered at (xc, yc). There are only two transitional circular paths that can be tangent to both the target circle and to the robot axis at the origin. These two circles are shown in Fig. 2 where the smaller one called \u201ctransition circle 1\u201d must be selected for a CW rotating target, while the larger circle called \u201ctransition circle 2\u201d is for a CCW rotating target for the given initial orientation of the mobile robot. Since the center of the selected transition circle lies on a line perpendicular to /0, we find xt \u00bc rtsgn\u00f0ah\u00desin /0; yt \u00bc rtsgn\u00f0ah\u00decos /0 (30) In addition, the geometry of the selected transition circle and the target circle leads to the relation xcxt \u00fe ycyt R2=2 \u00bc 6rcrt (31) where R2 \u00bc P2 c r2 c > 0, and Pc \u00bc ffiffiffiffiffiffiffiffiffiffiffiffiffiffi x2 c \u00fe y2 c p is the distance of the target circle center from the origin"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0000364_intelse.2016.7475150-Figure4-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0000364_intelse.2016.7475150-Figure4-1.png",
+        "caption": "Fig. 4. Components of the lift vector during a steady turn of radius R.",
+        "texts": [
+            " For the derivation of relationship between X and 'i/J using (3), the row vector (- sin X, cos X) is multiplied on both sides and it yields the form W n sin X - We cos X = Va ( - sin X cos 'i/J + cos X sin 'i/J ), (4) which can be written in simplified form as: 'i/J- x = sin- l (:a (Wn Sinx - WeCOsX)). (5) In the presence of wind disturbance, the idea of coordinated turn is explained in detail in [21]. In literature the coordinated turning is explained with the idea of no side force in the vehicle's body frame, thus side-slip angle,8 = O. During flight aerial vehicle minimizes the cross track error by producing lateral accelerations by tilting the component of aerodynamic Definition of Kinematic Variables. lift in the direction of turn. Fig. 4 explains the concept of bank to turn maneuver, in which the vehicle will bank to produce required lateral accelerations. Therefore the control input for o~r case is the reference bank angle commands \u00a2reJ. In FIg. 4, FliJt represents the aerodynamic lift vector, which is further resolved in two components. One component balances the centrifugal force acting on the vehicle due to the turn and other component (Fli Jt cos \u00a2) balances the weight of the vehicle [21]. FliJtcOS\u00a2 = mg, mV2 FliJt sin \u00a2cos(x -1j; ) = --1-. (6) where 9 is the acceleration due to gravity, m represents the mass of the vehicle and R is the radius of turn. From (6) we can have V 2 tan \u00a2 cos(X - 1j;) = -.lL Rg (7) For a steady turn Vg = RX, therefore (7) can be written as: V\u00b7 tan \u00a2cos(x -1j;) = gX"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0001948_5326.760575-Figure2-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0001948_5326.760575-Figure2-1.png",
+        "caption": "Fig. 2. Spherical coordinate system of the range finder.",
+        "texts": [
+            " Then, the range finder scans the front part of the pile which the reclaimer is paused. The height and width of the pile are, in general, about 15 and 30 m, respectively. Range data in the spherical coordinate system centered on the optical center of the range finder are acquired by encoding the step angles of the two rotating axes and detecting the depths of the points in the pile. The range data in the spherical coordinate system (r; ; ) are transformed into the Cartesian coordinate system (xs; ys; zs) as shown in Fig. 2. xs = r sin cos ys = r sin sin zs = r cos : (1) The range data in the local coordinate system are transformed to the world coordinate system by geometric translation and rotation. A height map is generated by mapping the range data in the Cartesian coordinate system with the intensity value in the (U; V ) image plane Fig. 3. The unit steps of the (U; V ) axes are selected according to the resolution of 3-D profile data. Since the range data in the height map are expressed as sparsely scattered points from the optic center of the range finder, it is necessary to interpolate the range data to construct a 3-D profile map"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0002215_0094-114x(95)00011-m-Figure6-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0002215_0094-114x(95)00011-m-Figure6-1.png",
+        "caption": "Fig. 6. Control of trajectory path geometry at a singular configuration.",
+        "texts": [
+            " where (P J) is the unknown Inflection Circle diameter. From this information the equation of the Inflection Circle is found to be, (x - 7.85 in.) 2 + (y + 20.00 in.): = (7.85 in.) 2. resulting in (P J) = 2(7.85)in. = 15.70 in. Knowing yp = - 20.00 in. and (PJ)x = (P J) = 15.70 in., equations (29) yield, n = . 2 5 0 and n '= .0345 which allows for a second-order coordination of the system by equation (27), # = n2 + ln'22 = .2502 + \u00bd(.0345)22. An example of the two revolute system in a singular configuration is shown in Fig. 6. The system is at an extreme reach. In this configuration, the system's Jacobian matrix is rank deficient and singular. The inverse cannot be found, and the standard RMRC method becomes complicard [18-20]. The difficulty that arises in resolved motion rate control is a consequence of using the inverse Jacobian matrix, which combines control of the path variable with control of the trajectory geometry. In the singular configuration shown in Fig. 6, the controlled point B is coincident with the Pole. It is well known in Curvature Theory that the point on the moving body coincident with the Pole is a singular point which is at a momentary dwell on its trajectory, i.e. at a cusp in its trajectory. However, it is possible to determine the Inflection Circle at this instant and therefore to develop a coordinating second-order Taylor series. Via Curvature Theory, control of the path variable and control of the trajectory geometry are separated. As will be seen, there are no singularities associated with the inverse path geometry problem, i"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0000477_j.ifacol.2015.09.726-Figure7-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0000477_j.ifacol.2015.09.726-Figure7-1.png",
+        "caption": "Fig. 7. Prototype generation scheme",
+        "texts": [
+            " To improve the fault detection problem, W.card (P) signals are generated. Let Hi be a 80-dimensional vector that represents the 80 first harmonics (under frequency 2 kHz), i.e. Hi = F (H (Pk, Q\u03c9)). Figure 6 shows the harmonics extracted from both simulated signals of both classes : HB4 and HB2B4. Thus, we can construct H the W.card (P)\u00d780 dimensional prototype matrix. Using this notation, Hi,j is the jth harmonic of the ithprototype and H\u2217,j is the jth harmonic vector of dimension W.card (P). Figure 7 presents the matrix H generation. As the features have not the same magnitude, data are normalized by dividing with the maximum of each feature. Using this process, each feature has the same significance and only variation depending on prototypes is relevant. We perform an iterative feature selection based on correlation between features and dispersion of each feature. Correlations between two features are computed, according to Bravais-Pearson coefficient as : corr (H\u2217,l, H\u2217,j) = cov (H\u2217,l, H\u2217,j) \u03c3H\u2217,l\u03c3H\u2217,j (1) where , \u03c3H\u2217,l is the standard deviation of H\u2217,l and cov (H\u2217,l, H\u2217,j) is the covariance of features H\u2217,l and H\u2217,j , 1 \u2264 l, j \u2264 80"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0002802_an9911600369-Figure1-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0002802_an9911600369-Figure1-1.png",
+        "caption": "Fig. 1 Effect of the imidazole concentration and the accumulation potential on the cyclic voltammograms. obtained at an HMDE, for a 3 x 10-7 rnol dm-3 copper(i1) solution in 0.1 rnol dm-3 hydrogen carbonate buffer at pH 8.5 with an accumulation time of 120 s. ( a ) Accumulation at 0.0 V; (b) accumulation at -0.1 V; and (c) accumulation at -0.6V. Concentration of imidazole [Im]: A, 0; B, 1.0 x 10-6; C, 1.0 x 10-5; D , 1.0 x rnol dm-3 for (a) and (c ) . For (b) A, 1.0 x 10-4; B, 3.0 x D , 7.0 x 10-4; E, 1.0 x 10-3; and F, 1.5 X 10-3 mol dm-3",
+        "texts": [
+            "4 As the highest currents were obtained in 0.1 mol dm-3 hydrogen carbonate buffer (pH 8.5), this buffer was chosen for use in further studies. A very small shift (only a few mV) was observed in the peak potentials when the pH was varied from 7.0 to 10.5, showing that there is no loss or gain of protons in the reduction process. The influence of the accumulation potential on the cyclic voltammograms of a 3.0 x 10-7 rnol dm-3 solution of copper(Ii), at various imidazole concentrations, is shown in Fig. 1. The cyclic voltammograms obtained when accumulation was carried out at 0.0 V for 120 s [Fig. l(a)], in the presence of 1.0 x 10-3 mol dm-3 imidazole, gave a single peak at -0.42 V in the cathodic scan. Two small, broad peaks at -0.40 and -0.35 V were observed in the subsequent anodic sweep. No peak was observed when the imidazole concentration was significantly lower. When accumulation was performed at -0.1 V, an increase in the peak current and a shift in the peak potential were observed when the imidazole concentration was increased from 1"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0003166_0168-874x(90)90028-d-Figure1-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0003166_0168-874x(90)90028-d-Figure1-1.png",
+        "caption": "Fig. 1. Two-dimensional shell model of the tire and sign convention for the external loading, generalized displacements and stress resultants.",
+        "texts": [
+            " In the present study a space shuttle nose-gear tire was modeled using a moderate-rotation Sanders-Budiansky shell theory with the effects of transverse shear deformation and laminated anisotropic material response included [1,20]. A total Lagrangian formulation was used and the fundamental unknowns consist of the five generalized displacements, the eight stress resultants, and the corresponding eight strain components of the middle surface. The sign convention for the different tire stress resultants and generalized displacements is shown in Fig. 1. The concepts presented in the succeeding sections can be extended to higher-order shear deformation theories, as well as to three-dimensional continuum theory. Each of the generalized displacements, the stress resultants, and the strain components is expanded in a Fourier series of the circumferential coordinate 0. The discretization in the meridional direction is performed by using a three-field mixed finite element model. The following expressions are used for approximating the external loading, generalized displace- A"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0002258_1.1332396-Figure1-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0002258_1.1332396-Figure1-1.png",
+        "caption": "Fig. 1 Physical model",
+        "texts": [
+            " Other simulations are presented by Olson @11# where the effects of structural inertia are also important, such as in an uncavitated bearing, a bearing with a rapidly varying applied load, two bearing derivatives of a gas engine main bearing, and rigid bearing stability analyses that identify operating regimes where self-excited oscillations ~instability! occur. 001 by ASME Transactions of the ASME s of Use: http://www.asme.org/about-asme/terms-of-use Downloaded F The physical journal bearing model is shown in Fig. 1, and the appropriate inertial Cartesian coordinate system is shown in Fig. 2. 2.1 Equation of Motion for Elastic Sleeve. The general equation for an undamped, discretized elastic solid is given by Cook @13# @M S#$d\u0308 S%1@KS#$dS%5$rS%. (1) An important assumption made is that transverse direction ~x and y in Fig. 3! tractions on the fluid surface can be ignored. Consequently, all degrees of freedom in Eq. ~1! that are not directly associated with the normal direction ~z in Fig. 3! on the fluid surface can be removed through static condensation ~for @KS#"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0001946_0273-1177(92)90261-u-Figure1-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0001946_0273-1177(92)90261-u-Figure1-1.png",
+        "caption": "Fig. 1. An idealized spherical cell. See text for details.",
+        "texts": [
+            " 4 ~ ~ ~, Gravitational Effects on Monocellular Algae (1)35 restored, and if the fluid\u2019s velocity gradient is not too strong, the contraption will travel on a trajectory along which the axis\u2014orienting gravitational torque is compensated by a torque due to viscosity and shear. This locomotion, directed by the joint effect of two physical interactions, is called gyrotaxis /1/. Once again, the orientation of the device\u2019s body is entirely passive and physical, requiring no clever mechanisms for sensing and steering. For small constant shear rate f, the angle which characterizes gyrotaxis of a spheroidal swimmer is given by KVpf sin9 = mgL 1 where 0 is the angle between the symmetry axis and the vertical direction, (See Fig. 1), K is a numerical constant which depends on the swimmer\u2019s shape, V is its volume and m its mass, L is the distance between the center of buoyancy and the center of gravity, p is the viscosity, and f is given in reciprocal time units. For a sphere of radius a, Ky =4~a 3and f is the vorticity of the fluid. For other shapes, the expression becomes more complicated /3/. The idea of gyrotaxis, doubly-passive physical orientation of a cell body by shear and gravity, coupled to swimming along the body axis, has been applied to Chlamvdomonas and other algae",
+            " SIZE LIMITATIONS: ROTATIONAL DIFFUSION The passive orientation by gravity of the cells\u2019 longitudinal axis depends on the displacement of the center of gravity from the center of buoyancy. The accuracy of that vertical orientation is limited by asymmetrical motions of the flagella and by rotational thermal jitter. Large cells with a substantial mass anisotropy have a well\u2014defined axial orientation when their behavior permits it. For smaller ones, Brownian rotation /5/, i.e. thermal angular jitter, limits the accuracy of this deterministic orientation. Which phenomenon dominates? Figure 1 shows an idealized spherical cell of radius a. The center of mass (CM) is displaced by the distance L from the geometric center (C). The cell swims with velocity V~along the axis (CM,C). The axis is offset by o from the vertical direction specified by g, the acceleration of gravity. The torque due to g is 1\u2019=mgLsinQ. This torque tends to decrease 0; torques due to local shear may tend to increase or decrease 0. When gravity torque is equal and opposite to shear torque, dO/dt=O. The viscous torque associated with rotation having angular velocity ~ is r = 8~tpa3~"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0000878_978-3-319-47614-8-Figure2.17-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0000878_978-3-319-47614-8-Figure2.17-1.png",
+        "caption": "Fig. 2.17 Transformation giving maximum shear stress",
+        "texts": [
+            " This is determined by finding the maximum of the shear stress transformation equation, and solving for \u03b8. The result is, tan 2hs \u00bc rx ry 2sxy ) hs \u00bc hp 45 \u00f02:16\u00de Substituting for \u03b8 = \u03b8s in Eq. (2.13) we get the maximum shear stress to be equal to one-half the difference between the two principal stresses. smax \u00bc ffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi rx ry 2 2 \u00fe s2xy r \u00bc r1 r2 2 \u00f02:17\u00de The transformation to the maximum shear stress direction can be illustrated as in Fig. 2.17. Prior to computer days, most of the designs were limited to simplified 2D cases. Instead of using Eqs. (2.14\u20132.17) for determining directions of principal planes and principal stresses and maximum shear stresses, engineers were accustomed for simple graphical methods, thus the Mohr\u2019s circle approach given here is very 2.3 Equilibrium Relations 51 popular. The design departments carry out routine activities by an automated approach and engage average technicians to carry out the routine part. Thus fool proof design procedures are standardized and this is where Mohr\u2019s graphical approach becomes useful rather than analytical approach"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0000077_icoin.2016.7427086-Figure6-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0000077_icoin.2016.7427086-Figure6-1.png",
+        "caption": "Fig. 6. Scenario when obstacle is detected in three sensors",
+        "texts": [
+            " Thus, we define urB as: urB = {rB 1 2 s} (8) s is the angle range of every sensor. Note again that if ds is detected within sensor 1 or 5, we cannot estimate urB since the obstacle can be significantly big. 3) Obstacle is detected in more than 2 sensors: For obstacle detected in more than 2 sensors, it is similar to previous approach however we need to use the distance given by the most right and most left sensors. We define urB as: urB = {rB 1 2 s} (9) s is the angle range of sensor covered by the obstacle. An example is illustrated in Fig. 6. To compute for given in the current time step tc and previous time step tc-1. As shown in Fig. 7, A moves to the left and the dashed line represents the nearest distance between the obstacle and A at tc1 and between the obstacle and A at tc. The representation of the distance is shown this way to show that the nearest distance point can be anywhere along the dashed line. Given the distances on tc and tc-1, we get the center point of the obstacle for a given radius which is based from our estimate"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0002132_jsvi.1996.0655-Figure3-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0002132_jsvi.1996.0655-Figure3-1.png",
+        "caption": "Figure 3. The force components of a bearing model.",
+        "texts": [
+            " Using non-dimensional quantities, equations (5a) and (5b) become ov014u 1z4 \u2212 a 14u 1z2 1t21\u2212 g2ova 12u 1z2 + ab012u 1t2 \u22122g 1v 1t \u2212 g2u1=0 (6a) and ou014v 1z4 \u2212 a 14v 1z2 1t21\u2212 g2oua 12v 1z2 + ab012v 1t2 +2g 1u 1t \u2212 g2v1=0, (6b) where ov =Cv/C, ou =Cu/C, 2C=(Cu +Cv), a= rV2 0L2/E and b=AL2/C. A linear bearing which is considered to be decoupled between the translational and rotational displacements can be modelled by eight coefficients, i.e., two direct stiffness coefficients Kxx and Kyy , two cross-stiffness coefficients Kxy and Kyx , two direct damping coefficients Cxx and Cyy , and two cross-damping coefficients Cxy and Cyx . In the fixed frame of reference, the equilibrium equations at the bearing point, as shown in Figure 3, are P'r =P'l \u2212KxxXl \u2212KxyYl \u2212CxxX l \u2212CxyY l in the X\u2013Z plane, and Q'r =Q'l \u2212KyyYl \u2212KyxXl \u2212CyyY l \u2212CyxX l (7) in the Y\u2013Z plane. Using co-ordinate transformation, the equilibrium equations in the rotating frame of reference can be expressed by non-dimensional quantities, as follows: Pr =Pl \u2212((kxx + gzxy)(1+cos 2gt)/2+ (kxy + kyx + gzyy \u2212 gzxx)(sin 2gt/2) + (kyy \u2212 gzyx)(1\u2212cos 2gt)/2)ul +((\u2212kxy + gzxx)(1+cos 2gt)/2+ (\u2212kyy + kxx + gzxy + gzyx)(sin 2gt/2)+ (kyx + gzyy)(1\u2212cos 2gt)/2)vl \u2212(zxx(1+cos 2gt)/2 + (zxy + zyx)(sin 2gt/2)+ zyy(1\u2212cos 2gt)/2)u\u0307l \u2212(zxy(1+cos 2gt)/2 + (zyy \u2212 zxx)(sin 2gt/2)\u2212 zyx(1\u2212cos 2gt)/2)v\u0307l , Qr =Ql \u2212((kxy + gzyy)(1+cos 2gt)/2\u2212 (kxx \u2212 kyy + gzxy + gzyx)(sin 2gt/2) \u2212 (kxy \u2212 gzxx)(1\u2212cos 2gt)/2)ul +((\u2212kyy + gzyx)(1+cos 2gt)/2+ (kxy + kyx + gzyy \u2212 gzxx)(sin 2gt/2)\u2212 (kxx + gzxy)(1\u2212cos 2gt)/2)vl \u2212(zyx(1+cos 2gt)/2 + (zyy \u2212 zxx)(sin 2gt/2)\u2212 zxy(1\u2212cos 2gt)/2)u\u0307l \u2212(zyy(1+cos 2gt)/2 \u2212 (zyx + zxy)(sin 2gt/2)+ zxx(1\u2212cos 2gt)/2)v\u0307l , (8a) where kij =KijL3/(EC) and zij =CijV0L3/(EC)"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0000984_j.ifacol.2016.10.148-Figure6-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0000984_j.ifacol.2016.10.148-Figure6-1.png",
+        "caption": "Fig. 6. Distance between the source and the formation of UAVs versus time.",
+        "texts": [
+            " Figure 5 shows the resulting trajectories for the source and for the \u201cleader\u201dUAV. As it is seen from the figure, the formation of UAVs is able to reconstruct the source\u2019s path. Although, the source localization is not a primary goal of the proposed approach, the fact that the UAVs approach the actual source of the gaseous release allows to obtain better concentration field estimation. The performance of the proposed approach is analyzed by the behavior of the distance between the source and the center of the flying formation (leader position), shown in Fig. 6. Three different cases are examined in this figure: the case with a single sensor in the estimator (5), Egorova et al. (in print, 2016), the case of multiple sensors with \u0393ii = 1, i = 1, ...,m in Eq. (6), and the case of multiple sensors with \u0393ii = 1, i = 1, ...,m, i = 4 and \u039344 = 4 in Eq. (6). As it is seen from Fig. 6, the flying formation with equal filter gains in (6) is able to track the source of gas release as fast as a single UAV. However, the use of multiple sensors provides more realistic estimation of the concentration gradients at the \u201cleader\u201d-UAV location required for the implementation of the guidance (13). The error of the estimator is evaluated using the root-meansquare deviation: E = ( N\u2211 i=1 (\u3008c\u3009i \u2212 \u3008c\u0302\u3009i)2 ) 1 2 (23) Figure 7 shows the concentration estimation error, determined by (23) versus time",
+            " Although, the source localization is not a primary goal of the proposed approach, the fact that the UAVs approach the actual source of the gaseous release allows to obtain better concentration field estimation. The performance of the proposed approach is analyzed by the behavior of the distance between the source and the 0 30 60 90 120 150 180 210 240 270 300 330 360 0 500 1000 1500 2000 2500 3000 D is ta nc e be tw ee n se ns or a nd s ou rc e (m ) Time (s) Multiple sensors, equal filter gains Multiple sensors, various filter gains Single sensor Fig. 6. Distance between the source and the formation of UAVs versus time. 220 230 240 250 260 270 280 290 300 310 320 330 340 350 360 0.00E+00 2.50E-05 5.00E-05 7.50E-05 1.00E-04 1.25E-04 1.50E-04 C on ce nt ra tio n es tim at io n er ro r (k g/ m 3 ) Equal filter gains Various filter gains Time (s) Fig. 7. Normalized estimation error versus time. center of the flying formation (leader position), shown in Fig. 6. Three different cases are examined in this figure: the case with a single sensor in the estimator (5), Egorova et al. (in print, 2016), the case of multiple sensors with \u0393ii = 1, i = 1, ...,m in Eq. (6), and the case of multiple sensors with \u0393ii = 1, i = 1, ...,m, i = 4 and \u039344 = 4 in Eq. (6). As it is seen from Fig. 6, the flying formation with equal filter gains in (6) is able to track the source of gas release as fast as a single UAV. However, the use of multiple sensors provides more realistic estimation of the concentration gradients at the \u201cleader\u201d-UAV location required for the implementation of the guidance (13). The error of the estimator is evaluated using the root-meansquare deviation: E = ( N\u2211 i=1 (\u3008c\u3009i \u2212 \u3008c\u0302\u3009i)2 ) 1 2 (23) Figure 7 shows the concentration estimation error, determined by (23) versus time"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0001235_b978-0-08-099425-3.00005-4-Figure5.88-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0001235_b978-0-08-099425-3.00005-4-Figure5.88-1.png",
+        "caption": "FIGURE 5.88",
+        "texts": [
+            " Also possible is the use of a dedicated external piece of software for handling, such as Optimum Tire from Optimum G, that can interrogate and render a tire data file with ease as well as handling a large amount of the fitting process described in Section 5.6.5. Attentionmust be given to the orientation as well as the location of tyre attachment frames within the model, particularly when suspension adjustments such as static toe and static camber are intended to be made parametrically. Not all tyre models run symmetrically in both forward and backward directions and so it is often good practice to ensure that all tyres are rotating the same way, as shown in Figure 5.88. One possible approach with full vehicle modelling is to set up a global coordinate system or GRFwhere the x-axis points back along the vehicle, the y-axis points to the right of the vehicle and the z-axis is up. The local z-axis of each tyre part is orientated to point towards the left side of the vehicle so that the wheel spin vector is positive when the vehicle moves forward during normal motion. Note that this is the coordinate system as set up at the wheel centre and should not be confused with the SAE coordinate system that is used at the tyre contact patch in order to describe the forces and moments occurring there"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0000225_1350650115580629-Figure1-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0000225_1350650115580629-Figure1-1.png",
+        "caption": "Figure 1. Schematic of a plain bearing.",
+        "texts": [
+            "16\u201318 Due to the out-of-roundness of its journal, the shaft\u2019s radius of curvature varies at every point of its contour. For this reason, the model entails dividing the contour of an out-of-round shaft into discretization intervals 2 of a certain size. As a result of their wear during shaft revolution, the bearing elements\u2019 initial parameters of contact change and their wear is cumulative. The paper presents the results of investigating the effect of small shaft lobing (ovality, trilobing, tetralobing) on the contact pressures and durability of plain bearings. To a shaft 2 of a plain bearing (Figure 1) rotating with an angular speed !2, a vertical radial force N is applied. The bearing has a radial clearance defined by \" \u00bc R1 R2. Under load, the elements of the bearing contact in the zone 2R2 0, where contact pressures p \u00f0 \u00de occur. The elements of the bearing are made of different materials, and therefore their properties and resistance to wear differ. Due to their small ovality, the elements of the bearing exhibit either single-area contact (Figure 1) or double-area contact (Figure 2). The technological ovality khhRk of the elements (k\u00bc 1; 2) is defined as follows: 1 \u00bc R1 R01, 2 \u00bc R02 R2 (Figure 2). In the course of rotation of the shaft (angle of rotation a2> 0), two patterns of interaction between the bush and shaft can take place: single-area contact and mixed (single- and double-area) contact. Faculty of Mechanical Engineering, Lublin University of Technology, Poland Corresponding author: Myron Chernets, Faculty of Mechanical Engineering, Lublin University of Technology, Nadbystrzycka 36, Lublin 20618, Poland"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0000267_iccke.2014.6993354-Figure1-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0000267_iccke.2014.6993354-Figure1-1.png",
+        "caption": "Fig. 1. Location and heading of DDMR in movement space",
+        "texts": [
+            " Then, refer to [9], the AFC scheme hybridized by artificial neural network technique (named ANN-AFC) predicting estimated inertia to increase its efficiency. II. MODELING OF TWO WHEELED MOBILE ROBOT In kinematics modeling, for simplifications, three assumptions had been considered in this type of robots: 1- All of robot parts, especially wheels, are solids. 2- Robot travels on the plane, smooth and non- deformable surface. 3- Wheels just move by rolling. The mobile robot location on the surface is illustrated in Fig.1. Because the robot moves by two wheels, the robot balancing is provided by a Castor wheel which plays no role on the kinematics of robot. The rotation angle of wheels, \u03c6(t), generate movement and heading (direction of motion) in robot body. In other words, robot position which was shown by (x,y,\u03b8) is depended to \u03c6(t). The rotation angle of right wheel and left wheel are represented by \u03c61, \u03c62, respectively. The robot linear velocity can be calculated by: 2/21Rv (1) Which, \u03c91 and \u03c92 are derivatives of wheel rotations, \u03c61and \u03c62, respectively"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0002352_s0003-9969(96)00058-1-Figure4-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0002352_s0003-9969(96)00058-1-Figure4-1.png",
+        "caption": "Fig. 4. Double reciprocal plots of oxygen evolution.O, without NaSCN at pH 8; O, 1 mM NaSCN at pH 8; A,",
+        "texts": [
+            " Sodium thiocyanate decreased the amount of molecular oxygen evolved under acidic conditions, but did not significantly affect the amount under neutral and alkaline conditions. Figure 3 shows effects of NaSCN concentrations on the oxygen evolution in dialysed saliva. At pH 8, all the concentrations of SCN- tested slightly increased the rate. At pH 5 and 6 the reagent strongly inhibited oxygen evolution. Half-maximal inhibitions were observed at the concentrations below 0.2 mM. Km values of hydrogen peroxide for the oxygen evolution reaction were determined using dialysed saliva (Fig. 4). The values were about 25 and 50 mM at pH 6 and 8, respectively. Sodium thiocyanate (1 raM) had no effect on the Km at pH 8, but increased it to about 80 mM at pH 6. The Vmax was slightly lower at pH 6 that at pH 8 in the absence of SCN-. Sodium thiocyanate (1 mM) increased Vmax at pH 8, but decreased it at pH 6. We then examined whether the shift of pH affects the rates of hydrogen peroxide-induced oxygen evolution in dialysed saliva (Fig. 5). In the reaction mixture of pH 5, preincubated for 10 min without NaSCN, oxygen was produced rapidly on the addition of hydrogen peroxide (trace A), but practically no oxygen was produced when preincubated with 1 mM NaSCN at pH 5 (trace B), as had already been shown in Fig"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0003571_s0141-0229(01)00317-9-Figure2-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0003571_s0141-0229(01)00317-9-Figure2-1.png",
+        "caption": "Fig. 2. Integration of the fibre optic sensor into high pressure equipment.",
+        "texts": [],
+        "surrounding_texts": [
+            "Investigations of enzymatic reactions in supercritical CO2 are often hindered by the high pressure involved in these processes, making reaction monitoring extremely difficult. This paper describes the implementation of a fiber optic based oxygen sensor into a high pressure reactor for supercritical carbon dioxide. The sensor is pressure resistant, working in supercritical carbon dioxide and reusable after depressurisation. The sensor signal is found to be affected by pressure changes, but stable at constant pressure. Oxygen concentration in supercritical CO2 is monitored using the disproportionation of hydrogen peroxide as a simple oxygen producing reaction. \u00a9 2001 Elsevier Science Inc. All rights reserved.\nKeywords: Supercritical carbon dioxide; Oxygen monitoring; Fiber optic sensor\nOver the last years the application of supercritical carbon dioxide (SCCO2) as medium for enzymatic reactions has received growing interest. SCCO2 is an environmentally friendly alternative to organic solvents. It is not toxic, not flammable and a process in which CO2 is used as medium\u2014 not produced as byproduct\u2014does not contribute to the greenhouse effect. So supercritical CO2 has already replaced obnoxious solvents like hexane or dichloromethane in various industrial scale processes [1].\nIn addition to it\u2019s \u201cgreen\u201d characteristics, SCCO2 has further properties that make it an interesting solvent for enzymatic reactions. At the critical point (Tc 5 31\u00b0C, Pc 5 73.8 bar), the density of gaseous and liquid carbon dioxide becomes equal. The liquid/gas interface disappears and both phases merge into a new single phase. The so formed supercritical phase has physical properties (viscosity, diffusivity, solvent strength) that can be varied in a continuum from gas like to liquid like with relatively small changes in pressure [2]. This offers the possibility to adjust the solvent properties to the needs of the particular process. The low critical temperature (31\u00b0C) allows the exploitation of these\nproperties for the application of enzymatic catalysts and many investigations have been carried out in this field [3,4,5,6,7].\nThe oxidation reactions carried out by Hammond et al. [8] and Randolph et al. [9] are particularly interesting, since they make use of another quality typical for supercritical fluids. Due to their partial gaseous character, supercritical fluids are fully miscible with other gases [8]. While oxygen shows only a limited solubility in water or organic solvents (,0.25% [10]), high oxygen concentrations can be achieved in SCCO2. Supercritical carbon dioxide is therefore a promising solvent for the oxidation of steroids, unsaturated fatty acids and other non polar substrates.\nA great hindrance for the investigation of enzymatic reactions in SCCO2 is the high pressure (.70 bar) involved in the process. The main problem is to draw representative samples during the reaction without disturbing the process. The early investigations were done simply by analysing the reactants before and after the experiment [11]. Later, sampling devices were designed, based on the circulation of the reaction medium through a sample loop [12]. A continuous monitoring of an enzymatic reaction was achieved by Bornscheuer et al. [13], who integrated a UV-spectrometer with high pressure flow-cell into the sample loop.\nThe first application of fiber optics for spectroscopic measurement was reported by Zagrobelny et al. [14], who investigated the spectra of pyrene in SCCO2 and fluorescent * Corresponding author. Tel.: 149-511-762-2509; fax: 149-511-762- 3004. E-mail address: scheper@mbox.iftc.uni-hannover.de (T. Scheper).\n0141-0229/01/$ \u2013 see front matter \u00a9 2001 Elsevier Science Inc. All rights reserved. PII: S0141-0229(01)00317-9",
+            "labeled bovine serum albumin in supercritical ethane. The use of fiber optic sensors for measurements of oxygen have been investigated by various groups [15,16]. In this article we describe the implementation of a fiber optic based oxygen sensor into a high pressure reactor for supercritical CO2. The stability of the sensor under high pressure is tested and a simple oxygen producing reaction, the disproportionation of hydrogen peroxide, is monitored in supercritical CO2.\nOptic sensors are based on a change of the optical properties like absorption or luminescence of certain indicator dyes caused by chemical substances. The fiber optic oxygen sensors used in these experiments is then prepared by immobilisation of tris(4,7-diphenyl-1,10-phenanthrolin)-ruthenium(II)-chloride (RuBPP), an oxygen sensitive dye, embedded in a silicone matrix at the end of an optical fiber. This optical fiber\u2014called optode\u2014is connected with the measuring unit. A diagram of the optical sensor as described by Comte [17], is shown in Figure 1.\nThe excitation light (470 nm) is emitted from a LED, reflected by a dichroitic mirror and focused into the optical fiber. The RuDPP dye immobilised at the other end of the fiber emits fluorescent light which is conducted back through the fiber to the measuring unit. The emission light with a wavelength of 580 nm passes through the dichroitic mirror and is detected by a miniaturised photomultiplier.\nThe underlying measurement principle is the change in fluorescence intensity of RuBPP due to fluorescence quenching by molecular oxygen. The resulting reduction of fluorescence with increasing oxygen partial pressure is converted on the basis of Stern-Volmer kinetics [17,18] to give a relative signal that is directly proportional to the oxygen concentration.\nFor the investigation of processes in supercritical CO2, a cylindrical stainless steel vessel of 1 cm wall thickness and an inner volume of 60 ml is constructed. The top of the vessel is fitted with a screw cap and sealed with o-rings of Viton 500 (Otto Gehrkens GmbH, Pinneberg, Germany). 1/16\u201d steel capillaries can be connected to the side and top of the reaction vessel with conventional HPLC-type ferrules of steel or PEEK (Knauer, Berlin, Germany). The reactor is filled with carbon dioxide using a high pressure pump (mini-pump duplex, NSI33R, Milton-Roy, Obertshausen, Germany). The pressure is monitored by manometers (Hensinger & Salmon, Germany) and needle valves (ERC, Altegolfsheim, Germany) are used to open and close connections. Any conventional HPLC components, like columns or filters can easily be integrated. The whole apparatus is contained in a thermostatic oven (Memmert, Schwabach, Germany) to maintain a constant temperature of 45\u00b0C in the whole system.\nTo connect the optical fiber of the oxygen sensor to the high pressure vessel, the coating of the fiber is removed at a length of approximately 5 cm at the end that will later form the sensor head. The 0.6 mm fiber core can now be introduced into a 1/16\u201d steel capillary of 0.7 mm inner diameter. The remaining space between the fiber core and the capillary wall is filled with clear epoxy (type L, R&G",
+            "GmbH, Waldenbuch, Germany), which is allowed to harden for 24 h. Due to the small area of the pressure exposed surface, this connection is pressure resistant up to at least 150 bar. After the hardening of the epoxy, the RuDPP fluorophor is immobilised on the tip of the optical fiber. The capillary holding the optical fiber can now be connected to the high pressure reactor in the usual manner using PEEK ferrules.\nThe fluorophor RuDPP was synthesised and kindly supplied by Prof. Meyer, Institute for Organic Chemistry, University of Hannover (based on [19]). Hydrogen peroxide solution (30%) was obtained from Aldrich, manganese(II) chloride (tetrahydrate) from Sigma. The carbon dioxide used is grade 4.5 from Linde AG (Wiesbaden, Germany).\nTo test the stability and performance of the sensor under high pressure, the optode is exposed to pure carbon dioxide without addition of oxygen. Experiments are carried out at a temperature of 45\u00b0C and pressures corresponding to gaseous as well as supercritical state. The optode signal is measured for more than 24 h before the pressure is increased. The signal during a typical experiment is shown in Figure 3.\nFirst the sensor is incubated for 30 h in carbon dioxide at normal pressure. The signal shows a slight drift to higher values. The increase of pressure to 50 bar leads to a sudden increase of the signal. Further increase of pressure results in another shift of the signal, this time towards lower values. Repeating the experiment, a shift of the signal, either to higher or lower values, is seen for every change in pressure.\nAt constant pressure the signal is also constant, apart from a slight drift to higher values. The drift can be attributed either to an extraction of the dye from the silicone matrix into the supercritical CO2, interaction of the medium with the silicone matrix (swelling effect) or a bleaching effect known to occur with the fluorophore RuBPP [17]. The drift, however, is no more than 2% of the sensors range in 24 h and can be neglected for shorter experiments or eliminated by linear compensation for longer experiments. No loss of activity was observed after depressurisation. A single optode could be reused for various experiments.\nThe monitoring of oxygen in SCCO2 is investigated using a simple, oxygen producing reaction, the manganese catalysed disproportionation of hydrogen peroxide. Since the sensor signal was shown to be affected by pressure changes, the system has to be equilibrated at operating pressure prior to starting the reaction. This can be achieved by adding aqueous hydrogen peroxide solution through a substrate column as shown in Figure 4.\nThe reaction vessel is filled with an aqueous solution containing catalytic amounts of manganese(II)-chloride. The optode is integrated through the screw cap to measure the oxygen in the supercritical phase at the top of the vessel. Using the right connection line, the system is pressurized to 80 bar. Connected to the reaction vessel is the substrate column containing 2 ml of hydrogen peroxide solution. The substrate column is brought to a pressure of 90 bar. Opening the valve between the column and the reaction vessel allows a piston fitted into the column to inject the solution into the reactor. Figure 5 shows the sensor signal during the experiment.\nWhen the reactor is pressurized, the sensor signal shows a shift to lower values during the pressure build up time. After reaching the operating pressure (80 bar) the signal remains constant. When the oxygen production is started, a constant rise of the signal is seen, as the oxygen produced in the solution diffuses through the supercritical CO2 to the top of the vessel (no pressure increase due to oxygen production is seen during the experiment). After 75 minutes the experiment is terminated by releasing the carbon dioxide and reducing the pressure to 1 bar. The signal of the oxygen sensor drops, since the oxygen is released from the vessel together with the carbon dioxide. After depressurisation, a new increase of the signal is seen. Oxygen is still produced from the hydrogen peroxide solution and the sensor now registers the oxygen diffusing through the gaseous CO2.\nA cheap and simple method for implementing an optical oxygen sensor into a high pressure reactor for supercritical carbon dioxide was developed. The sensor proved to be"
+        ]
+    },
+    {
+        "image_filename": "designv11_60_0003042_iros.1993.583851-Figure6-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0003042_iros.1993.583851-Figure6-1.png",
+        "caption": "Fig. 6: calculated sliding distances in horizont,al gripping plane",
+        "texts": [
+            "he table, each grasp candidate GGPO(object, p12, 8 h ) is translated too. The sliding dista.nce of object or obstacle can be calculated using GGPO and COPO. Practically the distances to translate GGPO(object,plz, Oh) or COPO(obstacle,p12, 8h.) where a part of the GGPO is not overlapped with any COPO are calculated for every discrete sliding direction. Fig5 shows an example of calculating GPO and COPO in the case where the gripping plane p12 and the upper surface of t.he table are parallel, and Fig.6 shows the results of sliding dist,ances of object of the gripping plane pl, shown in Fig.5 in the 32 directions when GGPO equals the whole GPO. Then in order to escape from a situation where the object cannot be grasped: object is t,ranslat,ed into t,he shaded portion shown in Fig.6. In the case where the gripping plane and the upper surface of the table are not parallel, the cross line between the gripping plane and the upper surface of the table is calculated. Then we solve the problem by means of calculating the sliding distances in the directions both parallel and perpendicular to the cross line. In the case of the parallel direction, we can calculate the sliding distance of either object or obstacle in the same way as describe above. Fig.7 shows the results of sliding distance of object in the parallel direction of the gripping plane pll shown in Figd when GGPO equals the whole GPO"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0001861_0379-6779(94)90257-7-Figure2-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0001861_0379-6779(94)90257-7-Figure2-1.png",
+        "caption": "Fig. 2. Potential- t ime relations during electrochemical aniline polymerization for a 0.1 M aqueous solution of aniline sulfate in 0.i M sulfuric acid under constant current (0.1 mA cm -2) conditions: (1) without initiator; (2)-(5) MoCI5 concentrations of 2 X l 0 -3 M, 4X10 -a M, 6X10 -3 M and 10 -2 M; (6) 10 -3 M HElrCltl.",
+        "texts": [
+            " considerable enhancement of the potential, on H2PtCI6 and RhCI3 addition to the solution (curves 6 and 7, respectively) is probably brought about by the formation of an oxide film. SSDI 0379-6779(94)02057-6 If the titanium electrode is kept in the H2IrCI6 solution at the potential of 0.8 V, then washed and introduced into the synthesis cell with aniline sulfate solution in sulfuric acid, the induction period is practically not seen (Fig. 1, curve 4). A similar effect has been found to take place on the introduction of MoC15 and CuCI2 under the same conditions. Figure 2 ultimately evidences the decrease in the overvoltage potential (curves 2-5). In addition, it is clear that the addition of HpIrCI6 (curve 6) produces a much more pronounced effect than those of MoCIs. It is of interest that enhancement of the initiating action is not always a post-effect of the increase in the additive concentration. Thus, the addition of CuCI2 (Fig. 3) of rather high and low salt concentrations is not efficient, though there is an optimum value of its concentration, which causes the most pronounced effect",
+            "8 V (Fig. 5) confirm the above observations. Voltammetric characteristics (VACs) of the same area can be obtained after 120 cycles of the synthesis without initiator and 20 cycles in the presence of HpIrC16. Therefore, introduction of a number of transition-metal > ~.6 \"~0.8 0.3 02 i i i i 02 t.0 t.8 Q/m 1, Fig. 3. (a) Maximum value of potential during aniline polymerization under constant current ( I=0.1 mA cm -2) conditions vs. CuCI2 (0.1 M solution) added. Other experiments conditions as in Fig. 2. (b) Potential of the first peak on the VAC of the polyaniline film vs. CuCI2 added. Experimental conditions are the same as for (a). Electrolyte 2.5 M H2SO4; scan rate 20 mV S - I . salts brings about a considerable reduction of the induction period and decrease of the synthesis overvoltage. Figure 5 also shows that the first peak of the VAC of the polyaniline film prepared using initiators is shifted in the cathodic direction. This effect will be discussed in a subsequent paper. As found in our previous work [7], pulse polymerization of aniline on a glassy carbon electrode yields qualitative layers of polymer of high corrosion resistance"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0000622_b978-0-12-800782-2.00003-8-Figure3.36-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0000622_b978-0-12-800782-2.00003-8-Figure3.36-1.png",
+        "caption": "Figure 3.36 Spatial location of four search coils.",
+        "texts": [
+            " Flux densities in the teeth of rotating machines are of the alternating type, whereas those in the yoke (back iron) are of the rotating type. The loss characteristics for both types are different and these can be requested from the electrical steel manufacturers. The motors to be tested are mounted on a testing frame, and measurements are performed under no load or minimum load with rated run capacitors. Four stator search coils withN\u00bc3 turns each are employed to indirectly sense the flux densities of the teeth and yoke located in the axes of the main- and auxiliary-phase windings (see Fig. 3.36). The induced voltages in the four search coils are measured with two different methods: the computer sampling and the oscilloscope methods. The computer-aided testing circuit and program [24] are relied on to measure the induced voltages of the four search coils of Fig. 3.36. The CAT circuit is shown in Figs. 3.36 and 3.37, where eight channel signals (vin, iin, im, ia, emt, eat, emy, and eay corresponding to input voltage and current, main- and auxiliary-phase currents, main- and auxiliary-phase tooth search-coil induced voltages, and main- and auxiliary-phase yoke (back iron)-search coil induced voltages, respectively), are sampled [36]. 249Modeling and Analysis of Induction Machines The flux linkages of the search coils are defined as \u03bb t\u00f0 \u00de\u00bc \u00f0 e t\u00f0 \u00dedt; (3-38) and numerically obtained from \u03bb0 \u00bc 0 and \u03bbi+1\u00bc \u03bbi +\u0394t ei+1 + ei\u00f0 \u00de=2, f or i\u00bc 0, 1, 2\u2026, n 1\u00f0 \u00de: (3-39) The average or DC value is \u03bbave \u00bc 1 n Xn 1 i\u00bc0 \u03bbi "
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0000868_fpmc2016-1767-Figure1-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0000868_fpmc2016-1767-Figure1-1.png",
+        "caption": "Figure 1: Transmission CAD Assembly",
+        "texts": [],
+        "surrounding_texts": [
+            "A preliminary analysis of the circuitry is an essential phase of the project. The tractor\u2019s transmission is an extremely complex assembly composed by hundreds of components. The 2 Copyright \u00a9 2016 by ASME Downloaded From: http://proceedings.asmedigitalcollection.asme.org/pdfaccess.ashx?url=/data/conferences/asmep/90210/ on 02/18/2017 Terms of Use: http://www.asme.org/about-asme/terms-of-use lubrication circuit, reflecting this complexity, has a large number of hydraulic consumers connected by a large network of conduits. Since two separated inlets feed two separate sections of the lubrication circuit and because of the complexity of geometry it is favorable to split the transmission in two parts, each fed by the corresponding inlet. The first part has one inlet and 56 outlets. A collector chamber is placed downstream of the inlet \u201cIN 1\u201d and it distributes the flow to moving shafts: Range 1/2 (A5) and Range 3/4 (A6), PTO Output (A7), Planetary Carrier Output Shaft (A9), Planetary Sun Output Shaft (A10), Satellite Shafts (S1-3), MFD Shaft (A11). Second part has one inlet and 17 outlets. Downstream of the inlet \u201cIN 2\u201d a stationary chamber can be found distributing the flow to shafts A1 (Planetary) and A2 (Forward Shaft). From the collector volume two nozzles (O31, O47) spray directly oil to gears internally of the transmission carter. An important task for this kind of problems is to correctly generate the Fluid Domain. The lubrication circuit is composed by the shafts internal conduits, by piping, by housing internal connections and collector volumes and finally by the volume generated from coupling different mechanical parts for example, bearing, shaft and gears. One of the most critical targets of this study has been the calculation of the flow rate of oil that lubricate gear, bearing and clutches.. These volumes are not easy to obtain for two reasons. First, the assembly model of transmission is composed more than 800 solid parts and the task of locating all fluid passages is not simple. Second, the solid model must not have dimensional, geometric and mating errors because it could generate negative volumes, misaligned geometries and nonconformal surfaces. The methodology adopted for fluid volume obtained is adapted piece by piece. As well as this, a common work flow can be outlined: first the 3D solid model of the transmission has to be simplified by deleting parts not directly interested by fluid flow. Second lubrication passages are generated through the function of subtraction of volume or filling function in mesh generator environment. Solving the problem different passages between mesh generator and solid modeler was necessary because of the very complex mating (i.e. splined coupling), very thin clearances, and complex geometries. The filling and subtraction tasks in general are undertaken for limited subset of assembly after the fixing of geometry errors. At the end of this very time consuming phase all the partial volumes have to be assembled to obtain the various computational subdomain. The subdivision of the computational domain into subdomains is necessary to define different speed rotating domains, which communicate via interface surfaces. In figure 3 the complete fluid domain is colored according to Figure 2 color scheme."
+        ]
+    },
+    {
+        "image_filename": "designv11_60_0003304_(asce)1084-0702(2002)7:5(300)-Figure16-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0003304_(asce)1084-0702(2002)7:5(300)-Figure16-1.png",
+        "caption": "Fig. 16. Splay band channel No. 10 ~units in millimeters!",
+        "texts": [
+            " In order not to hold the spinning wheel, the layers needed to be arranged within that time. It was determined that the wires could be arranged within this time because the arrangement time measured in Table 1 ranged from 14 min 23 s to 12 min 40 s. The tests were performed using the No. 10 strand, which was expected to show wire lift as well as the lateral displacement phenomena. The splay band channel that houses the No. 10 strand is made of cast iron and is installed with an angle of 16.2414\u00b0, corresponding to the main cable angle. As shown in Fig. 16, the channel has vertical and horizontal curvatures at the exit side ~strand shoe side! to accommodate the flaring of the main cable. It was also designed to have a 450 mm long horizontal strand spacer on the floor, which is used to separate the strands inside the splay at the exit side. In the test, the actual degree of the wire lift was first determined and then compared with the analytical values. After that, schemes to control the lift and to arrange the wires were devised. The steel bars and the clamps, which were originally devised to prevent lateral displacement of the wires in the previous test, were attached to the channel"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0000018_2016-01-1468-Figure4-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0000018_2016-01-1468-Figure4-1.png",
+        "caption": "Figure 4. Full Braking, slip angle comparison (not to scale).",
+        "texts": [
+            " Under the full braking condition, \u03b8 = 0 and \u03b1 + \u03ba = 90. The tires in Figure 3 all have the same slip angle. Now consider the effect of changing the slip angle. Under full braking, the angle \u03ba will be largest when the tire slip angle is small, and vice versa. In other words, when the slip angle is small, there is a relatively large angular difference in \u03ba between no braking and full braking. When the slip angle is large, there is a relatively small difference in \u03ba between no braking and full braking. In Figure 4, full brake scenarios for slip angles of 10 and 80 degrees are depicted on the left and right, respectively. In practice, the angle of the striation marks will become more sensitive to braking as the slip angle increases. This sensitivity is illustrated in Figure 5. The longitudinal slip percentage is plotted as a function of the angle \u03ba (Equation 2), for lines of constant slip angles. Maximum deceleration likely occurs at approximately 25 percent longitudinal slip for a typical passenger tire [1]"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0003887_978-3-642-73890-6-Figure11-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0003887_978-3-642-73890-6-Figure11-1.png",
+        "caption": "Figure 11. Basic hardware for the vi bratory insertion Process. The board is vibrated while the robot places the component at the nominal insertion point. The component drops into the board when holes and pins align.",
+        "texts": [
+            " The tube is SUbjected to mechanical vibrations at its natural frequency and in that condition, the fuel pellet easily enters into the tube, though the clearances are very small. In another application [14], vibrations were used to improve the lead insertion reliability during the assembly of nonstandard electronic components on Printed Circuit Boards. There are often problems in inserting rigid-leaded cube components in holes with tight lead/hole clearances because of tolerance stack-up and imprecise positioning by the robot. In the vibratory insertion process, the PCB is vibrated in a tranverse direction as illustrated in Figure 11 [14], so as to create a relative motion between the part and the board. The robot 705 picks up a component from the feeder and brings it to the nominal insertion location on the board (which is being vibrated). Eventual1 y, the board holes \"find\" the component leads. When the holes and leads align, the component drops into the board. The use of vibra tions greatly facilitated the insertion task and insertion reliabili ties in excess of 99.8% were realized under normal operating condi tions. From the time the component first contacts the board surface until the component inserts, the elapsed time was typically 0"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0003280_j.1460-2687.2001.00076.x-Figure8-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0003280_j.1460-2687.2001.00076.x-Figure8-1.png",
+        "caption": "Figure 8 (a) Spin mechanism in pure top spin case. (b) Spin mechanism in general con\u00aeguration with top spin, side spin and spiral. (c) Spin mechanism in pure spiral con\u00aeguration with ball",
+        "texts": [
+            " Just before release, the front air chamber pressure is reduced to vacuum, retracting the suction cup towards the front bearing support arm. When the ball contacts the drive shaft, the vacuum is broken and the cup continues to retract, allowing clearance between the cup and ball for the ball's accelerating trajectory provided by the propulsion system. An alternate method of ball release involves switching the suction cup pressure from vacuum to a positive pressure. Photographs of the spin mechanism in three different spin axis con\u00aegurations are shown in Fig. 8. Figure 8a portrays a pure top spin case, Figure 7 Cross-sectional view of spin rate mechanism with ball attached. \u00d3 2001 Blackwell Science Ltd \u00b7 Sports Engineering (2001) 4, 123\u00b1133 129 while Fig. 8b shows a setting that includes top spin, side spin and spiral. The power for rotating the spin mechanism is provided by a Dayton 83 W (1/9 hp, horsepower) permanent magnet DC motor seen in Fig. 8a. The motor rotation is delivered to the suction cup through a series of timing belts and gears. The X-axis stepper motor is visible in the foreground of Fig. 8b. This motor provides the ability to change the part of the ball angular velocity vector perpendicular to the barrel axis, i.e. to control the spin axis orientation in the range between top spin and side spin. The clear removable acrylic cover for the ball insertion is also shown in Fig. 8c with the ball in place. Here the axis of the spin mechanism is coincident with that of the barrel, resulting in pure spiral. An optical encoder (Hewlett-Packard QEDS-7163) with 256 counts per revolution per channel was attached to the timing belt shaft in parallel with the motor, and used as a tachometer. The performance of the spin control subsystem (with a desired spin rate of 10 Hz) is shown in Fig. 9. A brief period of open-loop control is used to bring the spin above 10 Hz, after which a PD control loop is entered"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0000688_j.asr.2016.08.002-Figure1-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0000688_j.asr.2016.08.002-Figure1-1.png",
+        "caption": "Fig. 1. Illustration of a possible beam deviation problem.",
+        "texts": [
+            "2016.08.002 0273-1177/ 2016 COSPAR. Published by Elsevier Ltd. All rights reserved. \u21d1 Corresponding author. E-mail address: henrique@ene.unb.br (H.C. Ferreira). Please cite this article in press as: Souza, A.L.G., et al. Antenna pointing s dictive control techniques. Adv. Space Res. (2016), http://dx.doi.org/10.1 Also, wind forces and manufacturing imperfections may disturb the control system (Gawronski et al., 2000; Gawronski, 2008). All these events cause pointing deviation, as illustrated in Fig. 1, and sensor calibration procedures for improving the pointing accuracy, as described in Bandikova et al. (2012) and Lee and Yeom (2015), are not always sufficient. Despite that, the control system must be able to point the antenna correctly. In order to have a satisfactory beam pointing, the pointing deviation is estimated during the communication. RF sensing techniques are used for this purpose allowing to estimate the spacecraft position relative to the beam (Dang et al., 1985). There are two main classes of RF sensing techniques, monopulse based techniques (Nateghi et al"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0003124_70.88068-Figure4-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0003124_70.88068-Figure4-1.png",
+        "caption": "Fig. 4. Initial contact between objects and obstacles. (a) Flat fixture in obstacle space. (b) Fifth-order polynomial representation of a fixture.",
+        "texts": [
+            " 3. Equation (1) is versatile and provides adequate mathematical representation of various objects for use in contact reasoning during robot compliant motions. Polyhedra may also be used to represent objects geometry, however, this adds extra steps to the algorithm used for contact reasoning because of the discontinuity of surface normals. MODELING OBSTACLE SPACE The success of any compliant motion path generated by a motion planner depends on the definition of the initial contact Configuration. Fig. 4 illustrates some models of the obstacle space (environment) and the anticipated initial contact location of the manipulated object when it reaches its goal location in the obstacle space. EXPECTED CONTACT CONFIGURATION Execution of compliant motion between objects in contact is further enhanced by the knowledge of initial contact configuration which includes: a) location and orientation of the compliance frame with respect to the fixed base frame, b) geometric model of manipulated object, and c) expected surface normals at the contact location"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0003433_s0378-7796(00)00121-8-Figure1-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0003433_s0378-7796(00)00121-8-Figure1-1.png",
+        "caption": "Fig. 1. Schematic representation of a synchronous machine.",
+        "texts": [
+            " This paper presents a synchronous machine model in phase coordinates where the effect of saturation along direct and quadrature axes can be taken into account by identifying the flux linkages of individual phases in terms of the flux linkages of direct and quadrature axes. The proposed machine model was implemented in SIMPOW\u2122 [13] using the high-level dynamic simulation language (DSL). The computed performance of the proposed model was verified against that of the standard dqo model. The performance of proposed model was found very close to the standard dqo model in SIMPOW\u2122. 2. Modelling details 2.1. Synchronous machine The three stator windings are arranged to form a set of 3-phase balanced system as shown in Fig. 1. The magnetic axes of the 3-phase system abc are displaced from each other by an angle 2p/3. The stator windings of the 3-phase systems abc are identically sinusoidally distributed windings, displaced 2p/3, with resistance ra and armature leakage inductance la. The rotor is equipped with a field winding and three damper wind- ings, one along direct axis and two along quadrature axis since the usual practice is to consider two damper circuits along quadrature axis for round rotor machine and only one circuit for salient pole machine [6]",
+            " The principal justification comes from the comparison of calculated performances based on these assumptions and actual measured performances. Assumption (h) is made for convenience in analysis. With magnetic saturation neglected, we are required to deal with only linear coupled circuits, making superposition applicable. The position of the rotor is specified with reference to the axis of phase a by an angle u. In terms of the flux linkages, the voltage relationships for all the seven circuits in Fig. 1 can be written as follows: 1 vb d dt [c]= [e]\u2212 [R][i], (1) where vb is the base speed and is taken to be equal to synchronous speed vs and the time t is expressed in seconds while all other quantities are expressed in per unit relative to Lad \u2013 base reciprocal per unit system [6]. The other matrices in Eq. (1) are given below: [e]= [ea eb ec efd e1d e1q e2q]T,[c] = [ca cb cc cfd c1d c1q c2q]T,[R] =diag[ra ra ra rfd r1d r1q r2q],[i] = [\u2212 ia \u2212 ib \u2212 ic ifd i1d i1q i2q]T. Since the damper windings are short-circuited, the voltage associated with them i"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0003722_s0921-8890(00)00130-5-Figure2-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0003722_s0921-8890(00)00130-5-Figure2-1.png",
+        "caption": "Fig. 2. Virtual components used by Pratt [23] for steady dynamic walking.",
+        "texts": [
+            " [22] implemented an algorithm called \u201cTurkey Walking\u201d 1 using the VMC 1 The label \u201cTurkey Walking\u201d is used because the algorithm was first applied to the dynamic walking task of a planar biped called \u201cSpring Turkey\u201d. approach. In the algorithm, they used a virtual parallel spring-damper component in the vertical direction to control the height, a virtual rotational spring-damper component for the pitch angle control of the body, and a virtual damper for the horizontal velocity control of the biped (Fig. 2). Such a selection and placement of virtual components was successfully implemented for both level ground walking [22] and sloped terrain walking [4] of the planar biped. The next natural task is to add disturbance adaptation capability to the algorithm. One common disturbance experienced by a service robot is a change in body mass, e.g., when the robot is used to transport external objects between locations. If the biped needs to carry an external load, the original Turkey Walking algorithm may result in a different or even unstable walking gait",
+            " The body height, body pitch, and horizontal velocity controls correspond to keeping z, \u03b1 and x\u0307 close to the desired value zd, \u03b1d and x\u0307d, respectively. In the Turkey Walking algorithm, the body height is controlled by a virtual spring-damper. The body pitch control is controlled by a virtual rotational spring-damper. The horizontal velocity control is con- 3 A singular configuration is one in which the upper and lower links of the legs are completely aligned with each other. trolled by a virtual damper. Fig. 2 shows the locations of these virtual components. These components yield the virtual forces [Fx, Fz, M\u03b1]T computed as follows: Fz = kz(zd \u2212 z) + bz(z\u0307d \u2212 z\u0307), (1) where ki and bi are the spring stiffness and the damping coefficient, respectively, for the virtual components in i (= x, z or \u03b1) coordinate. The virtual forces are then transformed into a set of desired joint torques for the stance leg(s) using two Jacobian matrices, one for the single support phase and one for the double support phase [22]"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0001146_jctn.2015.4303-Figure2-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0001146_jctn.2015.4303-Figure2-1.png",
+        "caption": "Fig. 2. The 1/24 workbench finite element model.",
+        "texts": [
+            " So it obtains the thermoelastic equation by calculating, as shown in formula (4): +2G e r \u22122G ( 1 r z \u2212 r ) \u2212 3 +2G T r = 0 +2G 1 r e \u22122G ( rz \u2212 z r ) \u2212 3 +2G T = 0 +2G e r \u2212 2G r ( 1 r r r \u2212 rz ) \u2212 3 +2G T z = 0 \u23ab\u23aa\u23aa\u23aa\u23aa\u23aa\u23aa\u23aa\u23aa\u23aa\u23aa\u23aa\u23aa\u23aa\u23aa\u23aa\u23aa\u23aa\u23aa\u23aa\u23aa\u23aa\u23ac \u23aa\u23aa\u23aa\u23aa\u23aa\u23aa\u23aa\u23aa\u23aa\u23aa\u23aa\u23aa\u23aa\u23aa\u23aa\u23aa\u23aa\u23aa\u23aa\u23aa\u23aa\u23ad (4) In the formula: r = 1 2 ( 1 r \u00b7 uz \u2212 u z ) = 1 2 ( 1 r \u00b7 ur z \u2212 uz r ) z = 1 2r ( ru r \u2212 ur ) It is known from the analysis of the above equations that if the temperature distribution of worktable is known, the deformation which needs to be calculated can be obtained by taking temperature distribution into the above equations. AND INITIAL CONDITIONS The 3D model of h drostatic bearing table is shown in Figure 1. This paper built the finite element model of worktable by using the software named ANSYS Workbench. Because the rotary worktable has periodicities, it just needs to mesh 1/24 of one single cycle. As shown in Figure 2. The initial conditions and boundary conditions are necessary to solve specific problems. The center of the worktable which is studied by this paper is touching radial cylindrical roller bearing, so it is applied by Fixed Support code to the fixed boundary conditions in inner surface of the circumference. The distribution of corresponding J. Comput. Theor. Nanosci. 12, 3917\u20133921, 2015 3919 Delivered by Publishing Technology to: Chinese University of Hong Kong IP: 117.253.218.180 On: Fri, 26 Feb 2016 00:22:32 Copyright: American Scientific Publishers R E S E A R C H A R T IC L E Fig"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0000363_rjaset.11.2025-Figure1-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0000363_rjaset.11.2025-Figure1-1.png",
+        "caption": "Fig. 1: SBE finite element",
+        "texts": [
+            " In this study, the strain based approach which was recently extended to the material nonlinear analysis of 2-D structures (Rebiai and Belounar, 2014, 2013) is used to examine linear and dynamic behavior (free and forced vibration analyses) of membrane structures through a new strain based element with drilling rotation named \u201cSBE\u201d Strain Based Element. A number of benchmarks with classical conditions and load conditions are considered which were already used in the validation of new finite elements. Formulation of the developed element: The element SBE with three degrees of freedom (Ui, Vi and in plane rotation \u03b8i) at each of the four corner nodes is shown in Fig. 1. In a 2-D analysis the relationship between strains and displacements are given by: x V y U y V x U xy y x \u2202 \u2202 + \u2202 \u2202 == \u2202 \u2202 = \u2202 \u2202 = \u03b3 \u03b5 \u03b5 (1) The strains given by Eq. (1) must satisfy the compatibility Eq. (2) which can be formed by the eliminating U, V from Eq. (1), hence: 0 2 2 2 2 2 = \u2202\u2202 \u2202 \u2212 \u2202 \u2202 + \u2202 \u2202 yxxy xyyx \u03b3\u03b5\u03b5 (2) We first integrate Eq. (1) with all three strains equal to zero to obtain: 3 32 31 a xaaV yaaU = += \u2212= \u03b8 (3) Equation (3) gives the three components of rigid body displacements"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0000021_iemdc.2015.7409254-Figure4-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0000021_iemdc.2015.7409254-Figure4-1.png",
+        "caption": "Fig. 4. Flux paths of PMSM",
+        "texts": [
+            "75th and multiples of this harmonic between all three phases is equal to zero. In the next section, the phase voltage and inductances under eccentricity conditions are discussed to explain the existence of harmonics at triplen mechanical frequency such as 0.75th component. In order to take a closer look at the dynamics of distinctive harmonics, the machine inductance derivations are revisited briefly. The air-gap flux density and inductances derived by applying the Ampere\u2019s law to the daxis current and loop along the flux path in Fig. 4; 0 2 2m core d PM Fe B B B h g Ni \u03bc \u03bc \u03bc + + = (12) where \u03bcPM is the permeability of permanent magnet (PM), lcore is the total length of flux path, hm and g are the thickness of magnet and air gap dimension and N is the number of turns of the d-axis winding. The third term in (12) can be eliminated because it is negligible when compared to the other terms (\u03bcFe =4000\u03bc0). In healthy PMSMs, the d-q axis inductances are defined as; ( ) 2 0 2d q m N A L L g h \u03bc = = + (13) As it is given in (13), stator inductances are inversely proportional with the effective air gap length"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0002619_1.2829168-Figure1-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0002619_1.2829168-Figure1-1.png",
+        "caption": "Fig. 1 (a) Grinding wheel left-side surface parameters and geometry",
+        "texts": [
+            " Essentially, a dual-lead worm has two different axial modules and heUx angles for the right and left worm tooth surfaces. Therefore, two different grinding wheels with some common parameters must be used to produce the respective tooth surfaces of a dual-lead worm. The grinding wheel right-side surface generates the left-side surface of the dual-lead worm, while the grinding wheel left-side surface gen erates the worm right-side surface. 2,1 Family of the Grinding Wheel Surfaces. The left side surface of a cone-type grinding wheel, as shown in Fig. 1 (a) , is used to generate the right-side surface of the ZK-type dual-lead worm. The position vector and unit normal vector of the grinding wheel surfaces can be represented in coordinate system SdXc, Yc, Z^) as follows: R\u201e, \u2014 and Uw = u cos cti cos 6 u COS a,- sin 9 + (bi \u2014 u sin a,) s m (jj COS sin a, sin ( COS a , (1) (2) X. r^tanar Zo>Z,t where fo, = inmj/4) cos Pj + r\u0302 , tan a;. Subscripts i = r and / where r indicates the right-side grinding wheel surface and / represents the left-side surface",
+            " (3) yields the family of grinding wheel surfaces as follows: R\u201e = By sin ip + Cy cos Lp By cos if - Cy sin if Dy - Pj^ (4) where and \u00b1ibi - u sin a,) sin pj + u cos \u00ab,\u2022 cos Pj sin 9, Cy = u cos a,- cos 9 + (rj, -I- r\u201ej + Ai), Dy = u cos a.; sin Pj sin 9 \u00b1 (\u00ab sin a, - bj) cos Pj - A^. 2.2 Mathematical Model of the Dual-Lead Worm Sur face. When two gear surfaces are in mesh, both meshing sur faces should keep in tangency at every contact instant under ideal contact conditions. Owing to the orthogonality of the rela- Fig. 1(/)) Grinding wheel right-side surface parameters and geometry tive velocity \u00a5\u201e'\"* and the surface common normal n'-p, ex- Journal of Mechanical Design SEPTEMBER 1998, Vol. 120 / 415 Downloaded From: http://mechanicaldesign.asmedigitalcollection.asme.org/ on 01/28/2016 Terms of Use: http://www.asme.org/about-asme/terms-of-use where pressed in coordinate system Sc, the following equation must be observed (Litvin 1989, 1994): n ( c ) . ^(Ic) ^ jj(c) . ( Y (1) _ y (c)^) ^ Q (5) where V^'\"' represents the relative velocity between worm and grinding wheel, expressed in coordinate system 5^",
+            " 1 2 0 / 4 1 7 Downloaded From: http://mechanicaldesign.asmedigitalcollection.asme.org/ on 01/28/2016 Terms of Use: http://www.asme.org/about-asme/terms-of-use [Mfc^cA = [Mi/J = 1 0 0 0 1 0 0 0 1 0 0 0 cos (/>! - s i n (t>i 0 0 (fwgi + /\"wj + -^4) 0 - A 5 1 sin </>i 0 0 cos (\u0302 1 0 0 0 1 0 0 0 1 ' and [Af,/\u201e] = - m2iZij sin (l>\\ sin y + mjiir^gi + r\u201ej + A4) cos <pi COS y - Asm2i sin </<i sin yjji + u!i[m2iy\\j sin </>i sin y \u2014 m^iXy cos </)| sin -y + rrhdKgi + r,\u201e; + A4) sin yjk,. (22) Similarly, according to the geometries shown in Fig. 1(a) and 1 {b) and coordinate system relationship shown in Fig. 2(a) and {b), the unit normal vector of the dual-lead hob cutter cos 01 - s i n 4>\\ 0 0 sin <\u0302 i cos y cos 4>i cos 7 sin 7 0 - s i n <\u0302 i sin y - c o s <\u0302 i sin y cos 7 0 ('\"ws/ + r^i -f A4) cos <\u0302 i -{r^si + '\u2022\u2022vj + A4) sin 0 -As 1 The angular velocity of the worm gear, as shown in Fig. 3, can be expressed by d<p2. . 0)2 = \u2014-Kf^ = wjk/H,. at (17) By applying Eq. (16), the angular velocity of the dual-lead worm gear can also be represented in coordinate system 5i {Xi, Ki, Z,) as follows: \u00abP' = cos 4>i sin 4>i cos y -s in </>"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0001625_s0140-6736(00)61292-8-Figure7-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0001625_s0140-6736(00)61292-8-Figure7-1.png",
+        "caption": "FIG. 7-Showing : A, effect of thin soles in stimulating toes to contract when stepping on an irregularity, thus allowing all the weight to come on to the metatarsal heads ; and B, effect of thick sole-i.e., surface is always smooth for foot, however irregular the ground.",
+        "texts": [
+            " The stimulus of a hard or irregular object is a deep sensa- tion and causes inhibition of the interossei muscles, whereas the stimulus of a smooth surface is a superficial sensation which causes the interossei to contract \u2019synergically with the flexor muscles of the toes. I believe this possibly explains why women suffer so much more than men from metatarsalgia. I do not think it is due to high heels so much as thin soles. If the sole of the shoe is thin enough for the foot to be able to perceive the irregularities of the ground, it would stimulate them to flex when walking instead of being straight (Fig. 7). PRACTICAL APPLICATIONS Our appreciation of the importance of paying more attention to the afferent side of the reflex arc led us, at Guy\u2019s Hospital, to be the first teaching hospital in London to appoint a chiropodist to work in our orthopaedic clinic. In the presence of a painful focus, whatever it may be, it is impossible to re-educate the muscles of the feet and toes. The chiropodist is the eradicator of nociceptive impulses, and he is, in my opinion, an essential member of an orthopaedic team"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0001774_1999-01-1767-Figure4-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0001774_1999-01-1767-Figure4-1.png",
+        "caption": "Figure 4. Tooth force calculation",
+        "texts": [
+            " First, gear misalignment resulting from the shaft dynamic response at this time step can be calculated and used when interpolating for transmission error and mesh stiffness from the maps as described above. Second, the vibrational response along the line of action of the tooth force is used to modify the effect of the transmission error on its dynamic component. This coupling of the shaft dynamic response from one timestep into the calculation of the excitation force and thus the dynamic response in the next timestep is summarised in Figure 4. where : Ftc = total tooth contact force Fnt = tooth force due to nominal torque km = mesh stiffness dte = transmission error along the line of action dvib = vibratory displacement along the line of action ct = gear mesh damping vvib = vibration velocity along the line of action The effect of coupling shaft dynamic response is best illustrated by considering a simple example. A positive transmission error will lead to an increase in the tooth mesh force equal to the product of transmission error and mesh stiffness at each time step"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0003547_0954406001523074-Figure7-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0003547_0954406001523074-Figure7-1.png",
+        "caption": "Fig. 7 (a) Involutes arrived at by tracing points as they move",
+        "texts": [
+            " The PAC-N base lines (Figs 4c and 5b) do not generate discontinuities and only give rise to concavities if the pitch outline itself is in part concave. For the pressure angle PAC of 60 used here (when averaged over a cycle, this is comparable with a normal pressure angle PA of 30 ), these base outlines extend outside the pitch outlines. With an external base line internal teeth may be expected, but the algorithm used for this \u00aegure does not reproduce this situation. An algorithm that does was used for Fig. 7, shown later. By using a su ciently large pressure angle, it would have been possible to eliminate the excursions of the base outline. However, upward of an equivalent of a PA of 55 would have to be used to contain the base line in this example. Gear teeth with unrolled from a base outline, maintaining the cord tangent to the base line at the point of contact. The pitch and base outlines used are the same as those for the second-order ellipses shown in Fig. 4, except that the PAC-E base pro\u00aele (Fig",
+            "comDownloaded from all other points on the conjugate line) will have moved down the conjugate line by the distance c1\u00b1c0. By these means incremental displacements can be arrived at for any point as it moves from an initial location such as ci to its terminating location on the \u00aenal base outline. If a point such as c0 is traced separately on two rotating coordinate systems, \u00aexed in each gear centre, then a pair of mating involutes such as those shown on Figs 7 and 8 will result. Alternatively, if c0 is traced in a coordinate system \u00aexed with the gear centres a continuous conjugate line will be generated [15]. Figure 7 shows the outcome of tracing points on the line of action for the case of PAC-N. Figure 7b is an enlargement of the area around the pitch point. In a coordinate system in which the pitch point is stationary, the gear centres will move horizontally as the gears rotate. The instantaneous line of action will remain stationary and will become a continuous line of action, \u00aexed in direction by the angle PAC, but its end-points will move up and down. On the other hand, in a frame of reference \u00aexed with respect to the gear centres, the instantaneous line of action will be straight but the continuous one will not necessarily be so. As may be seen from Fig. 7b, conjugate action is taking place as de\u00aened by Buckingham in Section 3 above. The meshing involutes are tangent to each other and perpendicular to the conjugate line where they cross that line. Where the base line extends outside the pitch line, there are reversals in the involutes, which indicates one of the many practical problems. By coincidence, in Fig. 7, the angle of rotation at which the pitch outlines are shown, the conjugate line, appears to meet both the tangent and normal conditions (E and N). This condition occurs twice per cycle, and Fig. 7 shows just one of them. Figure 8 is similar to Fig. 7, except that here a pressure angle PAT-N of 45 was used. That is, the line of action made a \u00aexed angle of 45 to the tangent at the down the conjugate line from one base outline to the other. The points are traced on rotating coordinate systems \u00aexed in each gear centre. The PAC-N line of action and base outline of the second-order ellipse as in Fig. 4c are used. (b) Enlargement of the region about the pitch point, showing that normal conjugate action is taking place between the opposite-handed involutes of action and base outlines of the same second-order ellipse shown above is used",
+            " (b) Enlargement of the region about the pitch point, showing that normal conjugate action is taking place between the oppositehanded involutes Proc Instn Mech Engrs Vol 214 Part C C02197 \u00df IMechE 2000 at OhioLink on November 7, 2014pic.sagepub.comDownloaded from contact point. This angle, larger than that used for Fig. 5c, shows that the base outline can be contained within the pitch outline if a su ciently large pressure angle is used. In a coordinate system in which the pitch point is stationary, the instantaneous line of action would rotate about a mean direction, which here would be at 45 to either the X or Y axis. Figure 8, like Fig. 7, shows involutes that meet the requirements of Buckingham's basic law for teeth pro\u00aeles (Sections 3.1 and 3.2 above). It can be seen that, in the region enlarged in Fig. 8b, it would be easy to make excessively long teeth so that they would clash with their opposite numbers. It is also clear that, where the base line comes close to the pitch line, it would be possible easily to cut teeth deeper than the base line. The involutes here are in principle the equivalent of those generated by rolling a rack around the pitch pro\u00aele, but it is clear that the addendum and dedendum vary considerably around the pitch outline"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0000743_gt2016-56592-Figure15-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0000743_gt2016-56592-Figure15-1.png",
+        "caption": "Figure 15 SOLID MODEL OF INITIAL (LEFT) AND OPTIMAL (RIGHT) IMPELLERS",
+        "texts": [
+            " Figure 14 THE CONTROL POINTS OF BLADE CAMBER CURVE AT THE HUB (0% SPAN) The meridional geometries of the initial impeller and the two optimal impellers are compared in Figure 13. As can be seen that there is a significant change of the endwalls in the first optimization. Such large modifications to the endwalls would normally completed manually but here it is achieved automatically. The cambers at hub of the three impeller are compared in Figure 14, and the 3D model of the initial and the final impellers is shown in Figure 15. It is worth to mention that the big changes to the endwalls and the camber by the optimisation will greatly modify the mechanical properties of the impeller blades, and a proper mechanical design is usually needed to qualify the impeller for production. However, this work focuses only on the optimization of aerodynamic performance of centrifugal impellers, the mechanical design and impeller manufacture of such impellers are not the subjects of this paper and will not be considered further here"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0002132_jsvi.1996.0655-Figure1-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0002132_jsvi.1996.0655-Figure1-1.png",
+        "caption": "Figure 1. A model of a non-axisymmetrical shaft and disk.",
+        "texts": [
+            " The problem may be overcome by expressing the governing equations in the rotating co-ordinates instead of the stationary co-ordinates. However, an asymmetry in the boundary condition due to non-axisymmetrical bearings is permitted; i.e., the periodically varying coefficients appear in the governing equations for the bearings expressed in the rotating frame of reference. A non-axisymmetrical and rigid disk, having principal axes m and n at an angle h apart from the principal axes U and V of the shaft is shown in Figure 1. The equilibrium of moment and shear force of a rigid disk are shown in Figure 2. In the rotating frame of reference, the equilibrium equations of moments at the disk point can be expressed by (a list of notation is given in Appendix C) Mr =Ml +(Id \u2212Dd cos 2h)a\u0308\u2212Ddb sin 2h +V(Jd \u22122Id)b +V2(Jd + Id)a\u2212DdV2(b sin 2h+ a cos 2h), Nr =Nl +(Id +Dd cos 2h)b \u2212Dda\u0308 sin 2h \u2212V(Jd \u22122Id)a\u0307+V2(Jd + Id)b\u2212DdV2(\u2212a sin 2h+ b cos 2h), (1) where a= 1U/1Z and b= 1V/1Z are components of the deflected angle about the V- and U-axes, Id =(Id m + Id n )/2 and Dd =(Id m \u2212 Id n )/2, and the equilibrium equations of the shear forces can be expressed by Pr =Pl \u2212md(U \u22122VV \u2212V2U), Qr =Ql \u2212md(V +2VU \u2212V2V), (2) where U and V are components of the lateral displacement along the U- and V-axes"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0000622_b978-0-12-800782-2.00003-8-Figure3.26-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0000622_b978-0-12-800782-2.00003-8-Figure3.26-1.png",
+        "caption": "Figure 3.26 Fundamental induction motor (Te1) and load (TL) torques as a function of angular velocity and slip s1.",
+        "texts": [
+            "25, one obtains the current of the hth harmonic eI shTH \u00bc eVshTH RsTH + R0 r sh + jh\u03c91 Ls\u2018TH +L0 r\u2018 : (3-24) Therefore, the electrical torque for the hth harmonic is Teh \u00bc 1 \u03c9sh q1V 2 shTH R0 r sh RsTH + R0 r sh 2 + h\u03c91\u00f0 \u00de2 Ls\u2018TH +L0 r\u2018 2 : (3-25) Similarly, the fundamental (h\u00bc1) torque is Te1\u00bc 1 \u03c9s1 q1V 2 s1TH R0 r s1 RsTH + R0 r s1 2 + \u03c91\u00f0 \u00de2 Ls\u2018TH +L0 r\u2018 2 : (3-26) The fundamental slip is s1\u00bc\u03c9s1 \u03c9m \u03c9s1 (3-27) or s1 \u00bc 1 \u03c9m \u03c9s1 ) \u03c9m \u03c9s1 \u00bc 1 s1 where \u03c9s1\u00bc \u03c91 p=2 is the (mechanical) synchronous fundamental angular velocity and \u03c9m is the mechanical angular shaft velocity. 237Modeling and Analysis of Induction Machines The fundamental torque referred to fundamental slip s1 is shown in Fig. 3.26 where Te1 is the machine torque and TL is the load torque. IfRs\u00bc0 then Te1 is symmetric to the point at (Te1\u00bc0/s1\u00bc0) or (Te1\u00bc0/\u03c9s1). The harmonic slip (without addressing the direction of rotation of the harmonic field) is defined as sh\u00bc h\u03c9s1 \u03c9m h\u03c9s1 ; where\u03c9s1\u00bc (\u03c91)/(p/2) and\u03c91 is the electrical angular velocity,\u03c91\u00bc2\u03c0f1 and f1\u00bc60 Hz. To include the direction of rotation of harmonicmmfs, in the following we assume that the fundamental rotates in forward direction, the 5th in backward direction, and the 7th in forward direction (see Table 3"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0002678_1.2829318-Figure5-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0002678_1.2829318-Figure5-1.png",
+        "caption": "Fig. 5 Chamber",
+        "texts": [
+            " The vertexes of the profiles of the central screw define the helical lines. The part of the cylindrical external side of the central screw that comes into contact with the satellite screws, and the segments of hehcal line, that are also common to the satellite screws determine the seal curves on the satellite screws (see Fig. 8). As well, the seal curves on the central screw can be determined by considering the vertexes of the satellite screws. The pump case and the parts of the screw surfaces between the exhaust zone and the seal curves define the chambers (see Fig. 5) in which the pressure can be assumed constant and equal to the value of delivery pressure. The parts of screw helical surfaces that are walls of the chambers are called sectors of surface. Finally we define as domain the set of surface parameters corresponding to a sector of surface (see Fig. 6). From a purely analytical point of view, once the domain of the surface parameters T and the transformation 3' are defined, the sector of surface S is also defined. The transformation ^ is here represented by three equations for the coordinates x, y, z in the surface parameters \u00a7 and u"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0002351_mchj.1998.1658-Figure1-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0002351_mchj.1998.1658-Figure1-1.png",
+        "caption": "FIG. 1. Mercury(I) benzoate electrode: (A) conductor cable, (B) banana plug, (C) metallic mercury, (D) Pt wire, (E) silicone glue, (F) sensor pellet (graphiteuHg2(Bzt)2uHg).",
+        "texts": [
+            " Part of the resulting solid was transferred to a press mold, being compressed at 20,000 psi, for about 90 s. The black pellet (1.5 mm thick, 13 mm o.d., and 0.6 g mass) was fixed at one end of a glass tube (13 mm o.d., 80 mm long) with a silicone\u2013rubber glue (Rhodiastic, Rho\u0302ne-Poulenc, France) and allowed to dry for 48 h. Sufficient metallic mercury was then introduced into the tube to produce a small pool on the inner pellet surface; electrical contact was established through a platinum wire plunged into the mercury pool and a subsequent conductor cable. The resulting electrode is diagrammed in Fig. 1. The electromotive force (emf) values are read to the nearest 0.1 mV with a Metrohm Model 670 Titroprocessor. The reference electrode was a Metrohm AguAgCl double junction, Model 6.0726.100. The pH of the solutions ions was adjusted with a Metrohm combination pH electrode (Model 6.0234.100). A thermostated titration cell (25.0 6 0.1\u00b0C) was employed. Volume were measured to 6 0.001 ml with Metrohm Model 665 automatic burettes. All experiments were performed in a thermostated room (25 6 1\u00b0C). The following cell was used: 2AguAgCl | [NaCl](aq)5 50"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0000766_0954406216671839-Figure2-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0000766_0954406216671839-Figure2-1.png",
+        "caption": "Figure 2. Coordinate system and mesh point.",
+        "texts": [
+            " The shaft geometric parameters, straight bevel gears and spur gears for the two-stage spur, and bevel gear transmitting system (shown in Figure 1) are listed in Table 1 to 3, respectively. In the following part, the spur and bevel gear mesh models are derived basically and the system model of the gear transmitting system are constructed. Here, it is noted that the lubrication and/or mesh friction have obvious influence on the bevel gear dynamic performance,20,21 but the friction effect is neglected in the following system model. Bevel gear mesh. The coordinate system and mesh point are illustrated in Figure 2. Here, rlx, rly, and rlz are defined as the position vector rl l \u00bc p, g\u00f0 \u00de of the mesh point in the local coordinate system. nlx t\u00f0 \u00de, nly t\u00f0 \u00de, and nlz t\u00f0 \u00de are components of the instantaneous unit vectors nl in the direction of the line of action. The position vector and direction vector are obtained using TCA, which is different from that in Peng and Lim22 and Lim and Cheng,23 using equivalence to Figure 1. Illustration of bevel gear transmission system. at CORNELL UNIV on September 26, 2016pic"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0003858_841296-Figure2-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0003858_841296-Figure2-1.png",
+        "caption": "Fig 2 Gas seal O-ring device",
+        "texts": [
+            " This arrangement considera bly reduces the transverse component of the for ce supported by the sensors. However, accurate setting of the horizontality of the eight blades is indispensable. This is performed with the cy linder head removed, using a lever device which enables the radial force diagram to be applied to the liner. The measurEment system uias calibrated with the engine hot by suspending weights from the conrod, which was disconnected from the crankpin. As far as gas leakage is concerned, the equipressure device shown in figure 2 and used in (7) was inadequate. A level difference o caused by crushing of the cylinder head gasket on tightening, differential expansion between the liner and ring and movement of the cylinder head due to the pressure forces would have ren dered this completely inefficient. Figure 2b shows the solution used. Sealing is provided by a viton ring seal, which offers a minimum surface with respect to hot gases ; this offered longer service life. 174 3.3 Acquisition and Shaping of Signals The sealing device used did not enable balancing of the pressures applied to the liner. This resulted in an error around the explosion point, which was corrected by means of pressure measured in the chamber. The correction force results from the pressure applied to the non-balanced area &s {fig 2b). The engine was equipped with its original mounts system. Due to non-balanced inertial for ces, as a consequence of its design, the engine was subjected to vertical acceleration. Measu rement of this acceleration at the sensors ena bled this to be allowed for. After correction, the friction force is written : F(t) = f(t) - pAS - m y f(t) = friction force measured by load cells p = pressure measured in chamber i s = annular surface of thickness 0.24 mm and diameter (determined by experiment) Y - acceleration measured on crankcase m = weight of setup supported by the force sensors = 2"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0000116_s00707-016-1611-8-Figure3-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0000116_s00707-016-1611-8-Figure3-1.png",
+        "caption": "Fig. 3 A rolling coin",
+        "texts": [
+            " , p) as follows: Pr = Mv1 [ d dt ( \u2202v1 \u2202ur ) + \u03c92 \u2202v3 \u2202ur \u2212 \u03c93 \u2202v2 \u2202ur ] + Mv2 [ d dt ( \u2202v2 \u2202ur ) + \u03c93 \u2202v1 \u2202ur \u2212 \u03c91 \u2202v3 \u2202ur ] +Mv3 [ d dt ( \u2202v3 \u2202ur ) + \u03c91 \u2202v2 \u2202ur \u2212 \u03c92 \u2202v1 \u2202ur ] + I1\u03c91 [ d dt ( \u2202\u03c91 \u2202ur ) + \u03c92 \u2202\u03c93 \u2202ur \u2212 \u03c93 \u2202\u03c92 \u2202ur ] +I2\u03c92 [ d dt ( \u2202\u03c92 \u2202ur ) + \u03c93 \u2202\u03c91 \u2202ur \u2212 \u03c91 \u2202\u03c93 \u2202ur ] + I3\u03c93 [ d dt ( \u2202\u03c93 \u2202ur ) + \u03c91 \u2202\u03c92 \u2202ur \u2212 \u03c92 \u2202\u03c91 \u2202ur ] . (65) To demonstrate how we can use the first-order form, Lagrange\u2019s forms, and Gibbs\u2013Appell\u2019s forms to derive the equations of motion, the classical problem of a rolling coin [10] as depicted in Fig. 3 is considered. In this figure, a circular disk C of radius R in contact with a horizontal plane H that is fixed in a Newtonian reference frame N is shown. Mutually perpendicular unit vectors nx , ny , and nz = nx \u00d7 ny are fixed in N . In addition, b1, b2, and b3 form a dextral set of orthogonal unit vectors (called reference frame B). Here, b1 is parallel to the tangent to the rim of C at the contact point C\u0302 between N and C , b2 is parallel to the line connecting C\u0302 to C\u2217(the mass center of C), and b3 is normal to the plane of C "
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0003560_nme.1620180606-Figure1-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0003560_nme.1620180606-Figure1-1.png",
+        "caption": "Figure 1. Combined finite element mesh",
+        "texts": [],
+        "surrounding_texts": [
+            "PHILLIP L. GOULDS Department of Civil Engineering, University of Houston, Houston, Texas, U.S.A. Department of Civil Engineering, Washington University, St. Louis, Missouri, L'.S A."
+        ]
+    },
+    {
+        "image_filename": "designv11_60_0001774_1999-01-1767-Figure13-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0001774_1999-01-1767-Figure13-1.png",
+        "caption": "Figure 13. Finite element mesh of transmission casing",
+        "texts": [
+            " A dynamic model of the geared shaft system was created, as described earlier, and used to predict the bearing forces. The shaft modes within the frequency range excited by the first harmonic of tooth passing frequency are shown in Table 2. The analysis of five complete revolutions of the input shaft was found to be sufficient to achieve a convergent solution. The dynamic bearing forces were then used to excite a finite element model of the casing with approximately 470,000 degrees of freedom, as shown in figure 13, and a boundary element model was used to predict the sound pressure at each microphone position Table 2. Shaft modal frequencies Mode Frequency [Hz] Speed at which HCR gearset excites mode [rev/min] Speed at which LCR gearset excites mode [rev/min] 1876 1585 1941 8 CORRELATION \u2013 The main objective of the simulation is to predict the effect of design and load changes on noise and vibration. Figure 14 shows the difference in vibration at eleven positions around the casing due to an increase in input torque from 200 Nm to 600 Nm, as measured and as predicted by the model"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0003906_s1474-6670(17)37234-8-Figure6-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0003906_s1474-6670(17)37234-8-Figure6-1.png",
+        "caption": "Figure 6: Structure of one axis of the robot with a pay load (mac;s M kg).",
+        "texts": [
+            " the Xk = Uk(t) The modal value is actually equal 10 the control law uk(t) provided by the kth controller. It is obvious to no tice that all the modal values are known time-varying variables as they are equal 10 the outputs of the fixed controllers. Finally, the crisp output u is obtained: p 3.1. Two-link robot The nonlinear plant to which the switching algorithm will be applied, is a two-link industrial robot. It corre- sponds in fact to the terminal part of a robot with two axes of rotation. The behavior of the robot with one axis as seen in Figure 6 will be studied. Let us consider that 1 ' is the total moment of inertia. 1 Hence: ( J s + J cJ ( J s + M L2 JJ,' = J m + -----;:;:- = Jm + N2 In the same way, the total viscous friction is: Y, = (Ym+ ~2J It is obvious to find the state reprensentation of the ro bot (Same equations as Eq. (1.7\u00bb with: ( Xl (1)= 8m (I\u00bb) X(I) = . xil)= 8m(l) Thus: Xl (t) = x2(t) . 'I, Ke L . (Xl (I\u00bb) X2(t)= - j'X2(t) + j'V(t)-MgNl,SIn ~, , , y(l) = 8s(t) = Xt(I)IN where M is the mac;s of the payload, V(t) is the voltage signal applied to the synchronous motor, L is the dis tance between the center of gravity of the mass M and the axis 8s and g is the acceleration due to gravity"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0000710_b978-0-08-100072-4.00007-1-Figure7.5-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0000710_b978-0-08-100072-4.00007-1-Figure7.5-1.png",
+        "caption": "Figure 7.5 (a) Schematic illustration of the microelectromechanical systems (MEMS) fringefield capacitive pH sensor. (b) Schematic of the multilayered pH sensor fabricated on silicon wafer. (b) Schematic of the IDEs. M.S. Arefin, M.B. Coskun, T. Alan, J.-M. Redoute, A. Neild, M.R. Yuce, A microfabricated fringing field capacitive pH sensor with an integrated readout circuit, Applied Physics Letters, 104 (2014) 223503.",
+        "texts": [
+            " The design of IDEs consists of two comb-like electrodes structure on a wafer and the sensing material on top of the electrodes. The amount of H+-ions, depending on pH values on the sensing materials, change the dielectric permittivity for the fringe electric fields from electrodes. This provides an overall capacitance change for pH change. 161Wireless biosensors for POC medical applications The MEMS pH sensor detects pH levels by measuring pH-induced permittivity changes in the interdigitated electrodes (IDEs) (Fig. 7.5) [47,48]. In Fig. 7.5(a), the schematic of the sensor is illustrated in three dimensions. The IDEs and pads are fabricated on a silicon (Si) substrate covered with a 500 nm silicon oxide layer. As represented in the exploded view of the sensor in Fig. 7.5(b), the top surfaces of IDEs are passivated by 5 nm thickness of silicon nitride (Si3N4) layer to provide a sensing surface for pH buffer solutions as well as to limit Faradaic currents between the electrodes. The IDEs, having each electrode width of 25 \u03bcm and length of 200 \u03bcm, and interelectrode spacing of 10 \u03bcm, are illustrated in Fig. 7.5(c). In the solution bulk, an electric double-layer capacitance and a diffuse-layer capacitance is formed due to the free H+-ions in the solution [94,95]. An electric double-layer of H+-ions forms at the nitride\u2013solution interface that depends on the concentration of H+-ions of the solutions [94]. The surface charges on the nitride layer act as a source or sink for the H+-ions in the solutions. In the diffuse layer, the dielectric properties of the solution are modified predominantly by electronic and orientational polarizations under the influence of the high-frequency electric fields",
+            " The higher concentration of H+-ions attenuates the external electric fields due to the higher local electric field around the ions and, thus, orients the dipolar water molecules in its vicinity. Hence, lower pH values decrease \u03b5pH. Moreover, the changes in H+-ions concentrations yield the changes in conductivity (\u03c3pH) of the solution that produce frequency (\u03c9)- dependent complex permittivity \u03b5 * pH = \u03b5pH + j\u03c3pH/\u03c9\u03b50. Therefore, the change of frequency-dependent dielectric constant for different pH levels is reflected as a change in the capacitance of the sensor. 162 Medical Biosensors for Point of Care (POC) Applications The capacitance of such a multilayered structure, as shown in Fig. 7.5, having low layer thickness is very difficult to study either by modeling or by finite element method analysis [96]. However, if there is a monotonic increase in permittivity of the layers in the direction of the electrodes\u2019 plane to outer medium, the total capacitance of the sensor can be evaluated with the contribution of each layer in series as [96]: Csen = N 2 ( CoxCsi Cox + Csi + CpHCn CpH + Cn ) (7.1) in which Csen is the total sensor capacitance, N is the number of electrodes, and Cox, Csi, CpH, and Cn are the capacitance contributions from silicon dioxide layer, silicon wafer, pH-buffer solutions, and the silicon nitride layer, respectively"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0003173_s0141-6359(02)00117-4-Figure5-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0003173_s0141-6359(02)00117-4-Figure5-1.png",
+        "caption": "Fig. 5. Wood concept model of coupling with circumferential adjustment compliance and with contact faces being independent.",
+        "texts": [
+            " Although part radius is a function of the torque to be transmitted and the materials available, steps can be taken to ensure that for a given part radius, the radius of contact is maximized. Three wood models of coupling concepts were constructed to form physical \u201csketch models\u201d that could be handled as a means to quickly physically evaluate concepts. If the coupling concept was robust, it should be able to accommodate the errors associated with novice woodworking. The first was of the basic matched-compliance concept, as seen in Fig. 3. The second and third addressed compliance to allow for manufacturing errors. In the second, shown in Fig. 5, adjustment is parallel to the working direction, and consists of a new male component for the old female component. The third uses beveled fingers which adjust normally to the working direction, requiring both a new male and female component. This model is shown in Fig. 4. All of the models are of glued-assembly construction to minimize the complexity of machining required. Segments of 25 mm (1\u2032\u2032 nominal) square stock (30 cm long for the fingers, and 10 cm long for core support) were finished and glued together in a grid to form the male and female components"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0000897_978-3-319-50472-8_2-Figure4-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0000897_978-3-319-50472-8_2-Figure4-1.png",
+        "caption": "Fig. 4. Case 3.3.",
+        "texts": [
+            " The cases are listed as follows: - \u2022 Case 3.1: - If the occupancy rate is nil on both sides then the robot does not make any movement (Fig. 2). \u2022 Case 3.2: - If the occupancy rate is equal on both sides then the robot makes one hop movement to the side with closer neighboring occupied nodes (Fig. 3(a)) or any of the sides if there is a tie (Fig. 3(b)). \u2022 Case 3.3: - If the occupancy rate is more on the counter-clockwise direction but the clockwise string is nil then the robot does not make any movement (Fig. 4(a)) else it makes one hop movement to the counterclockwise direction (Fig. 4(b)). \u2022 Case 3.4: - If the occupancy rate is more on the clockwise direction then the robot makes one hop movement to the clockwise direction (Fig. 5). 2-Node Problem: The main concern which may act as a thorn on the path of gathering is the 2-node problem, where two nodes in the ring are occupied by the robots. In fact the inclusion of the property of chirality in the algorithm is just to tackle the 2-node problem otherwise N 2 visibility range alone would be enough for the gathering to complete"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0000622_b978-0-12-800782-2.00003-8-Figure3.15-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0000622_b978-0-12-800782-2.00003-8-Figure3.15-1.png",
+        "caption": "Figure 3.15 Field for the determination of the saturated rotor leakage flux of a 3.4 MW, two-pole, three-phase induction motor. One flux tube contains a flux per unit length of 0.005 Wb/m [19].",
+        "texts": [
+            " No closed form solution exists because of the nonlinearities (e.g., iron-core saturation) involved. In Fig. 3.13 the field for the first approximation, where saturation is neglected and a linear B\u2013H characteristic is assumed, permits us to calculate stator and rotor currents for which the starting field can be computed under saturated conditions assuming a nonlinear (B\u2013H) characteristic as depicted in Fig. 3.16. For the reluctivity distribution caused by the saturated short-circuit field the stator (Fig. 3.14) and rotor (Fig. 3.15) leakage reactances can be recomputed, leading to the second approximation as indicated in Fig. 3.16. In practice a few iterations are sufficient to achieve a satisfactory solution for the starting torque as a function of the applied voltage as illustrated in Fig. 3.17. It is well known that during starting saturation occurs only in the stator and rotor teeth and this is the reason why Figs. 3.13 and 3.16 are similar. 220 Power Quality in Power Systems and Electrical Machines F2 = 3500 N/m F1 = 3500 N/m F 2 = 3 11 0 N /m F 1 = 3 15 0 N /m F 2 = 1250 N/m F1 = 4340 N/m F 2 = 1770 N /m F 1 = 5160 N /m F 2 = 1400 N /m F 1 = 4110 N /m F 2 = 1 30 0 N/m F 1 = 5 45 0 N/m 1 \u2013B \u2013B 2 3 A A 4 5 \u2013C \u2013C \u2013A \u2013A C C B B 6 7 8 9 10 11 12 (a) 1 \u2013B \u2013B 2 3 A A 4 5 \u2013C \u2013C \u2013A \u2013A C C B B 6 7 8 9 10 N S 11 12 (b) f f 1 \u2013B \u2013B 2 3 A A 4 5 \u2013C \u2013C \u2013A \u2013A C C B B 6 7 8 9 10 N S 11 12 (c) Figure 3"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0002486_s0005-1098(99)00044-8-Figure4-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0002486_s0005-1098(99)00044-8-Figure4-1.png",
+        "caption": "Fig. 4. Trajectories of eight runs of the adaptive fuzzy control system which cover the domain of interest.",
+        "texts": [
+            "y $ and e 2 \"yR !yR $ with a.*/ 1 \"a.*/ 2 \"!2, a.!9 1 \"a.!9 2 \"2, and the membership functions equally spaced. We choose c\"200 and Q, k j 's in such a way that p n \"p 2 \"[5, 5]T. Fig. 2 plots the y(t) versus y $ (t) for a single run of the adaptive fuzzy control system starting from initial condition e(0)\"[!2, 2]T, and Fig. 3 shows the trajectory of the same run in the e 1 }e 2 phase plane. We performed a total of 8 runs to cover the domain of interest and the trajectories of these runs are plotted in Fig. 4 in the e 1 !e 2 plane. The \"nal rule base is shown in Fig. 5. The method proposed in this paper achieved the goal of designing fuzzy controller automatically without domain experts; this is a step forward from the existing approach where human experts are needed to specify the rules. The key idea of the approach is to use the feature of adaptive fuzzy controller that control parameters can be automatically generated through adaptation (learning). Simulation results showed that the method can success- fully create a complete fuzzy rule base for controlling a nonlinear time-varying chaotic system"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0003453_cccc19841390-Figure1-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0003453_cccc19841390-Figure1-1.png",
+        "caption": "FIG. 1",
+        "texts": [
+            " 1 Sulfate lignin; 2 alkaline sulfate lignin. Conditions: acetate-phosphate buffer, pH 5'0, 35\u00b0C Collection Czechoslovak Chern. Commun. [Vol. 491 [1984] ----------------_.------------------ Oxidation of Lignins 1393 amide gel. The calibration curves were plotted from the steady state response of the electrode and coincided for the above substrates. The results obtained permit an en zyme electrode for the determination of lignin to be constructed. The lignins under study completely dissolve in dioxane and partially in ethyl alcohol. Fig. 1 shows the rate of oxidation of pine lignin isolated from sulfate liquor and measured in terms of absorption of oxygen as a function of lignin concentration. The maximum rates coincide for lignin dissolved in dioxane and alcohol. However, the concentrations at which no dependence of the rate of the enzymatic reaction on substrate concentra tion can be observed, are lower for lignin dissolved in alcohol than in dioxane. The lignin residue insoluble in alcohol is not oxidized by oxygen (curve 3). The v'nax -value for the lignin samples studied is (7 \u00b1 2) 10- 2 mol O2 min -1"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0002252_s0890-6955(98)00006-6-Figure11-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0002252_s0890-6955(98)00006-6-Figure11-1.png",
+        "caption": "Fig. 11. The flank wear of a worn nose radius tool with chamfered main cutting edge.",
+        "texts": [
+            " (33), (FH)Umin is equated to the principal component of the resultant cutting force Rt, which consists of Ft and Nt, as: (Rt)H = NtcosaS2\u00b7cosab + (Ft)Umin \u00b7sinae 2 (FH)Umin (34) where the frictional force is determined by Ft 2 tssinbcosaeQ/[cos(fe + b 2 ae)sinfe] (35) Therefore, Nt can be rewritten as Nt = [(FH) 2 (Ft)Umin sinae]/(cosaS2\u00b7cosab) (36) The values of FT and FV can be determined from the components of Nt and Ft as: FT = 2 NtcosaS2sinaS2sinab + Ft(sinhccosab 2 coshcsinaS2sinab) (37) FV = 2 NtsinaS2 + Ft(coshccosaS2) (38) where Nt is the normal force on the tip surface with minimum energy; and (Rt)H is the horizontal cutting force in the horizontal plane. The wear of nose radius tool tip is shown in Fig. 11. The coordinates of points O1, C, S1, T1 and P derived by using the data measured from optical microscope, are shown in the following: O1: (XO1, YO1) = (R,R), P: (XP,YP) = (h3 + h5, l1 + l2), C: (XC, YC) = (h3 + h5 + h2, l1) and S1: (XS1, YS1) = (R\u2013Rcos Cs, R\u2013Rsin Cs). The coordinates of point S2 can be obtained by establishing the equations for straight lines O1S1 and PC, from which the point of intersection S2 can be calculated, as, S2: (XS2, YS2)where XS2 = {[l1h2 + l2(h2 + h3 + h5) 2 R1h2]cosCs + R1h2sinCs}/[PCcos(uPC 2 Cs)] (39) YS2 = {[l1h2 + l2(h2 + h3 + h5) 2 R1l2]sinCs + R1l1cosCs}/[PCcos(uPC 2 Cs)] (40) Tool-edge wear in the direction of depth of cut has to be obtained in order to estimate the worn depth dB of the tool edge"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0003149_tpas.1983.317956-Figure8-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0003149_tpas.1983.317956-Figure8-1.png",
+        "caption": "Fig. 8 Alternating static B-H characteristics for Vicalloy used in hysteresis ring.",
+        "texts": [
+            " The mean value of peak flux densities of the inner, outer and centre layers is also shown in Fig. 7 and as expected it is closely linear. I. Fig. 5 Output from I = inner search (b) integrators connected to search coils. coil, 0 = outer search coil (a) supply phase voltage 100 V, vertical scale: 0.4 (T) per division. (b) supply phase voltage 240 V, vertical scale: 0.9 (T) per division. (a) 2747 In order to determine the power loss in the ring due to the alternating hysteresis, a series of static B-H loops are plotted in Fig. 8 for specimens of Vicalloy cut from the same batch as the ring laminations and subjected to the same heat treatment. The specimens were cut along and across the rolling direction. Exact simulation of the B-H loops for the Vicalloy type alloys are done using improved algorithms 12, 19. The curves for the dynamic B-H characteristics at 50/60 Hz do not differ appreciably from the static curves21. It is now also possible to use the measured flux distribution of Fig. 7 and the B-H loop areas of Fig. 8 to compute the motor developed torque with the assumption of alternat-ing fields only, that is, no rotational field effects in the ring. Fig. 9 shows the terminal voltage-power characteristics of the motor at a slip of 0.13. Curve (a) of Fig. 9 shows the measured power across the airgap, which is obtained from input power less the armature copper and slip ring losses. Curve (b) shows the corresponding computed power using only alternating hysteresis effects. While curve (c) shows the corresponding parasitic losses in the ring"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0000323_j.ifacol.2016.03.043-Figure1-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0000323_j.ifacol.2016.03.043-Figure1-1.png",
+        "caption": "Fig. 1. Chaser (3 \u00d7 3 m), manipulator (link lengths 1.3, 1.3, 0.4 m, width 0.2 m), and target (2 \u00d7 2 m, solar panels 1.5\u00d7 1 m). Manipulator is in its stowed pose.",
+        "texts": [
+            " In the section after that we describe how the non-smooth path thus obtained is smoothened and parametrized for time optimization. In the section after that we formulate the minimum time trajectory determination problem and describe how it was solved. In the last section we summarize the major conclusions from the work. We chose the dimensions of chaser and target satellites to be somewhat similar to that of the ETS VII mission (Oda 1994). The chaser, manipulator in stowed pose, and target are shown in Fig. 1, and the relevant planar dimensions are mentioned in the caption of the figure. 4th International Conference on Advances in Control and Optimization of Dynamical Systems February 1-5, 2016. NIT Tiruchirappalli, India Copy ight \u00a9 2016 IFAC 142 Minimum Time Collision-Free Trajectories for Grabbing a Non-Tumbling Satellite \u22c6 T Venkata Bhargava \u2217 K Kurien Issac \u2217\u2217 \u2217 Team Indus, Bengaluru, India (e-mail: venkat.bhargav28@gmail.com) \u2217\u2217 Indian Institute of Space Science and Technology, Valiamala, Thiruvananthapuram 695547 India (e-mail: kurien@iist",
+            " In the section after that we describe how the non-smooth path thus obtained is smoothened and parametrized for time optimization. In the section after that we formulate the minimum time trajectory determination problem and describe how it was solved. In the last section we summarize the major conclusions from the work. We chose the dimensions of chaser and target satellites to be somewhat similar to that of the ETS VII mission (Oda 1994). The chaser, manipulator in stowed pose, and target are shown in Fig. 1, and the relevant planar dimensions are mentioned in the caption of the figure. 4th International Conference on Advances in Control and Optimization of Dynamical Systems February 1-5, 2016. NIT Tiruchirappalli, India Copyright \u00a9 2016 IFAC 142 ini u i e ollision-Free rajectories for rabbing a on- u bling Satellite \u22c6 T Venkata Bhargava \u2217 K Kurien Issac \u2217\u2217 \u2217 Team Indus, Bengaluru, India (e-mail: venkat.bhargav28@gmail.com) \u2217\u2217 Indian I stitute of Space Science and Technology, Valiamala, Thiruvananthapuram 695547 India (e-mail: kurien@iist",
+            " In the section after that we describe how the non-smooth path thus obtained is smoothened and parametrized for time optimization. In the section after that we formulate the minimum time trajectory determination problem and describe how it was solved. In the last section we summarize the major conclusions from the work. We chose the di ensions of chaser and target satellites to be somewhat similar to that of the ETS VII mission (Oda 1994). The chaser, manipulator in stowed pose, and target are shown in Fig. 1, and the relevant planar dimensions are mentioned in the caption of the figure. 4th International Conference on Advances in Control and Optimization of Dynamical Systems February 1-5, 2016. NIT Tiruchirappalli, India Copyright \u00a9 2016 IFAC 142 ini u Ti e Collision-Free Trajectories for Grabbing a Non-Tu bling Satellite \u22c6 T Venkata Bhargava \u2217 K Kurien Issac \u2217\u2217 \u2217 Team Indus, Bengaluru, India (e-mail: venkat.bhargav28@gmail.com) \u2217\u2217 Indian Institute of Space Science and Technology, Valiamala, Thiruvana thapuram 695547 India (e-mail: kurien@iist",
+            " In the section after that we de ib h w the n - mooth path thus obtained is smoothened and parametrized for time optimization. In the section fter th t we formula e the minimum time traj ory determination problem and describe how it was solved. In the las section we summarize the major conclusions from the work. We chose the dimensions of chaser and target satellites to be somewhat si ilar to that of the ETS VII mission (Oda 1994). The chaser, manipulat r in stowed pose, and target are shown in Fig. 1, and the relevant planar dimensions mentioned in the caption of the figure. 4th International Conference on Advances in Control and Optimization of Dynamical Systems February 1-5, 2016. NIT Tiruchirappalli, India Copyright 2016 I AC 142 T Venkata Bhargava et al. / IFAC-PapersOnLine 49-1 (2016) 142\u2013147 143 The three joint angles of the planar 3R manipulator are used to define its configuration space. We assume that the joint motions are restricted to 10o \u2264 \u03b81 \u2264 170o, \u2212160o \u2264 \u03b82 \u2264 160o, \u2212160o \u2264 \u03b83 \u2264 160o, to prevent collisions between adjacent links which are assumed to be in the same plane"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0003658_esej:20000302-Figure6-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0003658_esej:20000302-Figure6-1.png",
+        "caption": "Fig. 6 The rotational sensor",
+        "texts": [],
+        "surrounding_texts": [
+            "a s a support system for the pilot is accepted undci- all cii-cuIiistiiiiccs (iniplying diat it should b the i r s t of the coiistriiction) aci-odyii:imically iiiorc cliicicnt dcsigiis bccoiiic fcasiblc.\nOther good exaiiiples ofuiecliatroiiic system c m be found i n automotive applic:itions, such :is AIS, clcctrunic stahilisatimi systems, active suspctrsion systciiis and antonlared highways.\nIn tlie iiiiiii symposiuiii 'Mechatroilics iii control systim dcsign' nr the Control 'OX confci-encc in Swnnsc:i various applications a n d design issues wcrc pi-escntcd. Among tlicsc wrrr pipers on 'A knowledge-based ~iiechatrooics approaccli to controller design\" arid 'A I'C based uiiiversal control system'.'. I'npers 011 npplications involved 'Visioii-in-thc-loop control applicatioiis in textile mariiifactiire\", ' I h e l o p - iiiriit of a fuzzy hehavioui-al cmiti-ollci- for an autoiioiiious veliiclc\"' and 'I)evelopirieiit of :idiptivc cruise control systciiis for iiiotor velii~lcs'~.\nr l ic syiicrhy ofdiffcrcnr disciplines allows tlic dcsigii of really :idvanccd aRbi-d:iblc products and production niachines.\nMechatronic design\nMecliatronics is a way uf thinking rather than a coniplctcly iicw discipline. It needs the advaiiced knowledge of specialists tiom diffcreiit disciplines meeting i n a riieclinti-onic design tcniii. Mcchatrouics is a design philosopliy It is neither realistic: 1101' necessary to re-invent tlie wheel, especially since the time to market is so itliportarit. Mechati-mic design ofprodtiction machincs can spccd up rcnction to lilnrkct deinands. A flexible pi-o-oduction line that c m be recontigmed using software is niucli easiei- to adapt\nthnn coiivcntioiial lincs tllat rcquire iiieclmiicnl devices to be innnually rccoiifigiircd.\nUy dcvckiping proper took and kiio\\vlcdgc bases, cxistiiig hnowledgc caii be niadc available to less cxperienccd dcsigiicrs. Such knowledge bises should be tilled not oiily with standard soltitioris for iiicch:inicnl cotiipoiiciits, but also with appnipi-iatc CAI) tools :ind mntlicmarical niodels of tliese coiiipmeiih. and with coiitrol structiises fix dcfined clns~cs of problcui. 'I 'hey could also contain srandal-d, wcll tcstcd sokwirc iirodules, ctmbling the aotouiatic genelation of code for a collipliter-hased conti-ollci-. Oiie may doubt whether the dcsigii process could evcibe done automatically, Alrhougll the power of computational intelligence is increasing rqidly human creativity mi iiot yet be bcntcn by a coiiiputer. But\nENGINEERING SCIENCE ANI) El)UC.KI'ION JOUICNAI. JUNE 2000\n107",
+            "providing the lionaan designer with proper tools can considerably increase their productivity\n?bob fur imdellin'q, simulation and mntvuller des@ Simulation can play an iniportant role in the process of designing inechatronic systems. Computer simulation allows alternative designs to be compared arid evalnatcd without the expense of building real prototypes. Sininlation tools used in control engineering arc mostly based on a block diagram rcpresentation of the underlying mathematical model. These models have a direct connection with the transfer functions of the various couipoiients of the system. If necessary, they can be extcuded with uorilinearities. For the design of uiechatronic system transfer functions and block diagrams ai-e often not the iiiost appropriate models. A basic assumption in ablock diagram is that tlie &&rent blocks do not influence each other:? properties, or that any interaction between the blocks has been accounted for in the parameters. This iniplics that they canuot easily be replaced by other system components. Another problem is that thc parameters of various physical components appear in\nvarious combinations and at various locations in the block diagram. Uilless there is a supporting system available that autouutically relates the different parameters of the mechanical systetn to the parameters of the block diagram, investigating the effects of parameter changes becoiiics a tedious job. Iconic diagrams like basic electrical network or mechanical diagrams do not have this problem. Energy-based modelling approaches (e.g. the bond graph approach) can form a link between iconic diagrams and matheuiatical cquatious. Such models can help to increase insight in the design and may suggest alternative solutions (Fig. 3).\nI n the Coiitrol T.aboratory at thc University of Twente a sofhvare package (20-siin)' bas bccn dcvekiped that supports tlie modelling and sinrulatioii with bond graphs, i n addition to the use of equations and block diagrams. Version 3 of this program also supports iconic diagranis and object orientation. The latter allows the process to start with a simple des@ using only basic fiinctions of tlie various components. As the design process proceeds more complex representations ofthe coniponcnt can be incorporated in the iiiodcl and their effect on the system bchaviour can be examined. A iiiodcl of a coniponerit is this not fixed. It can have various shapcs. 'The models arc polyniorphic, i.e. they can have various levels ofdetail. Also viewing the system iu various representations or in iiiultiple views can increase insight into the propcrties of thc systcm (Fig. 4). Aiiiong these various repi-esentations are: rcpresentations in the freqnency domain, time domain, differential equations, bond graph, iconic diagrams aud block diagrams, as well as more fancy representations like stereo views as foond in virtual reality 20-sim 3.0 cm autornatically generate (linear) state space descriptions from the simulation code. This allows the use of tools like Matlah for ftirther atialysis, control system dcsigii and geiicratiug other representations. Demo versions of20-sim arc available from the web'.\nThese concepts and their impact on mechatronic design have been described in the I'hD thesis of De Vrics\". Ureuncsc\"' lias further worked\nthein into a couccpt for a modelliiig aud simulation language. Related work has beeri done, e.g. in the Schcmcbuikler project\". Proper soilware tools ~hould support various represcntations arid allow conversion &om one representation to another.\nI n order to advance the applications of real mcchatronic designs it is essential that design knowledge is foriiialised and brought together in a reusable knowledge base. This base",
+            "support tools to retrieve the knowlcdgc and niakc a new design.\nA coiitrol etigineeriiig cllalleiigc is to introduce iiiodern control illethods to 'standard' iucchatroiiic designs. 111 inany cases, simple PIl>-type coiitrollers arc applied because they can perform rcasonably well witliout too much toning and design e f i m I t is a cliallcngc to develop tools that allow the application o f i i i o ~ advanced controller algorithms with the saoie or cvcii less cffm as required for tuning a PID-controller. Uy developing tools that support such a design for various classcs ofsystcms, this should be possible.\nExamples\nTlicsc 111cc1iatri)nic design cxaniplcs arc fioiii the (hi tsol Laboratory of the Paculty of Electrical Enginccring of the Uiiivcrsity of Twciitc. All thcsc projects wcrc pcrfimiied in the niiiltidisciplinary eiiviroiiinent of the Cornelis J. I)rebbel Institute for Systems Engineering (fornlerly MILCT), a cooperation of the faculties of clectrical eiigineeriiig, mechanical engineering, applied mathematics and coinpiiter eiigineering. The projects indicate that good iiiecliatronic designs requirc attention to the iiicclianical design, computcr iiiiplcnicnta~ion and choice of sensors and control system.\nAlnscn projerl 111 tlic Alasca project a device for placing 1C:s on a priiitcd circuit board has bccn devclopcd. I t s l i o u l d replace older, difilciilt to control pnciumatic cquip~iient by an electric servo system that can rot:itc and translate simultaneously. accul-ately arid Cut. A design team of a mechanical and clrctrical cnginccr wgs Eomicd to design tlie niotoi- and its control (both students fiaoi the 'Mcchatronic Ilcsigncr' postgraduate course). An induction type of iiiotor W:IS developed witti two sets ofwitidings, one to realise the rotation ami another one to realise tlic translation (Fig. 5).\nTo achieve the required accuracy, air bearings were used. T h i s could only be achieved by using contactless sciisors to iiicasure the two motions. The inductive sensor fur measuring the translation w x iiiore or less stnndaiil, although care lind to be taken due to the magiictic fields of the motor. A contactless rotational s e i i s w was developed that could accurately nicasurc tlic rotation even wlieii the actuator peiforiiis trauslatiotial motions. The setisor coiisists of a coiiihiiiatioii of four LEDs at the stator, a slicet of polarising material at tlie rotor/traiislator and four photodiodes at the stator covered with sheets of polarising material whose directions of polarisation arc spread over 120\u00b0 (Fig. (I). One of die LEl)-pIiotodiodc conibinations acts as a reference; the other three yicld a scnsor signal conipatiblc with the signal o f a synchrn A standaid synchi-o-to-digital coiivcrtcr co~ild thus intei-bce bct\\vccn the sensor and the coinputel; yielding a resolutioii of 14 bits.\nllccausc induction motors have n low efficieiicy, Fig. 9 MART robot"
+        ]
+    },
+    {
+        "image_filename": "designv11_60_0002670_0301-679x(90)90059-x-Figure11-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0002670_0301-679x(90)90059-x-Figure11-1.png",
+        "caption": "Fig 11 A gear tilted reladve m end plates",
+        "texts": [
+            " the angle at which the maximum clearance occurs and t the tilt of the relative surfaces which is defined as h ...... - h,, t t7. (3) Before solving the equation given in ( 1 ), it is convenient to transform it into non-dimensional form by introducing the following non-dimensional wlriables: ()301-679X/90/060429-9 \u00a9 1990 Butterworth-Heinenlann Ltd 429 Notation b b d d h hc hd hrnax hmin he hq h q r r c ~c F0 Y t B L P ~c P Pd Q T Half-width of tilted ring bearing = b/ro, the non-dimensional width of the ring Gear displacement (see Fig 11) = d/hd the non-dimensional gear movement Film thickness between relative surfaces of bearing W Central film thickness W Design clearance (see Fig 11) Wa Maximum film thickness Minimum film thickness Wm~ Central clearance Z Bearing misalignment Z = h/ho, the non-dimensional clearance 8 Flow per unit length Radius Central radius 0 = rJro , the non-dimensional central radius Outer radius of bearing and gear fG~x = r/ro, the non-dimensional radius Tilt of the relative surfaces Bearing width Bearing length Pressure Central pressure = ph2/ ,qmr 2, the non-dimensional pressure - ph2?qmr g. the non-dimensional pressure Total flow - h</hd, the non-dimensional bearing misalignment (see Fig 11) Load-carrying capacity of bearing = Wh20/~qm~, the non-dimensional load = W h 2 / ~ a ~ , the non-dimensional end load acting on the gear -:- W h2midTl~or 4, the non-dimensionai load Radial coordinate of tilted ring bearing = Z / ro non-dimensional radial coordinate Distance between two adjacent points Distance between two adjacent points Sector angle Angular coordinate in finite difference method Angle of maximum clearance Dynamic fluid viscosity Rotational speed r = r/ro 7~ = h /ho and -ff = ph2 l~lear 2 where re is the outer radius of the bearing",
+            " TRIBOLOGY INTERNATIONAL 435 The numerical method developed has been applied to the gear ends of high-pressure pumps and motors in order to investigate the sealing and lubrication mechanisms of such fluid power units. By examining the bearing misalignment and hydrodynamically generated pressures between the ends of the gear and casing, the likely position of the gear under any operating conditions is assessed and the likely minimum film thicknesses predicted by this theory are shown to agree with the values measured ~4 Fig 11 shows a gear in a fixed end plate pump, which has a relative misalignment between the two shaft bearings. The magnitude of the bearing misalignment is expressed in terms of the difference in clearance between the maximum and mean clearance at the gear tips, hq in Fig 11. In addition to being misaligned, the gear may also be displaced from the centre towards one end, d in the same figure. In this analysis, the outer radius of the gear, to, and design clearance, hd (the clearance between the end faces of the gears and the end plates, being formed as a thrust bearing, with the gears perfectly aligned) were used as reference values for non-dimensional variables ~4. The anatysis starts with defined values of non-dimensional pressure Pd and gear position. For both ends, the nondimensionalized sealing pressures are obtained in terms of the central clearance, h~"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0003339_s0022-460x(87)81301-9-Figure1-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0003339_s0022-460x(87)81301-9-Figure1-1.png",
+        "caption": "Figure 1. Element of the system.",
+        "texts": [
+            " Detailed analysis of non-linear parametric vibration may be found in Evan Iwanowski's book [6]. This paper is concerned with torsional vibrations of two parallel shafts coupled together by a number of mechanisms. A purpose of this paper is to find the principal sources of vibrations and to study the influence of the system's parameters on its behaviour. The system considered consists of two parallel shafts coupled together by a number of mechanisms. The system may be divided into N elements shown in Figure 1. Here, An' Ell are the mass moments of inertia of the discs with respect to the rotation axes; S,,, D n , H; are the coefficients of viscus damping; MA \" , M a\" are the torques acting on the discs. The rotary motion of the disc An is transformed through intermediate massless links (not shown in Figure 1) into the motion of the disc B\" according to the function w; = 1Jr(rfJn). Man = Mn,,(cPn) and dlfl'n/dcP\" are periodic functions of ep\". The torque M A\" is assumed to be of the form 0022-460X /87 /140221 + 17S03.00/0 MA\" = e\"(n,, -dcPnldt). 221 (1) \u00a9 1987 Academic Press Limited An s\".D\" S. ,,0. , This relation expresses a static characteristic of a motor in a simplified form [6]. nn is the angular velocity for which MAn =0 and -e\" is the slope of the motor characteristic. The shafts are supported in frictionless bearings and the flexural vibrations are prevented"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0003183_00032718908051410-Figure1-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0003183_00032718908051410-Figure1-1.png",
+        "caption": "Figure 1. Schematic representation of the flow injection system. P, HPLC pump; I, injector; D, detector; W, waste; C, chart recorder.",
+        "texts": [
+            " Spectra were recorded utilizing a Perkin-Elmer double beam spectrophotometer (Coleman 124). The output was displayed and recorded using a REC 80 Servograph chart recorder. A Rabbit-HP pump (Rainin Instrument Co, Inc.)\u2019was used to deliver specific amounts of the appropriate reagents at a preset flow rate. Manual injection was achieved using a Rheodyne type manual injector. 0.5 mm internal diameter stainless steel and Teflon tubing was used in all connections after the injection port. A schematic diagram of the flow injection system used in this study is shown in Figure 1. Reagents Solutions DTNB and Triton X-100 were purchased from Sigma Chemical Company. 1% (w/v) solution of DTNB was prepared in acetone (EM Science). Cetyltrimethylammoniurn bromide (CTAB), cetylpyridinium chloride (CPC) and lauric acid (LA) were purchased from Matheson Coleman & Bell. Anhydrous sodium sulfite was obtained from Mallinck- rodt, Inc. A stock solution of sulfite (1 x loe2 M) was made in 1 mM CPC and standardized against acidified potassium permanganate. All preparations were made using doubly distilled deionized water"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0002252_s0890-6955(98)00006-6-Figure5-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0002252_s0890-6955(98)00006-6-Figure5-1.png",
+        "caption": "Fig. 5. Specifications of nose radius tool with wear.",
+        "texts": [
+            " The area of shear plane A and the friction plane Q for case (2) can be obtained as follows: For convenience of calculation, the shear plane must be projected in the plane perpendicular to the speed of cut, which makes the calculations and analysis much easier and saves the time required for calculations. Defining the chip flow angle in this perpendicular section as hc9, the relation between hc9, and hc on the tool face is presented in Appendix Eq. A(6). According to Eq. A(6), the shear plane area can be varied by changing hc9 in small increments. Geometrical wear of the cutting tool on the tool face is shown in Fig. 5 and the type of wear on the tool edge is defined by a curve of radius R1, a straight line and a curve of radii R2, R3 (R2 = R3). The side view of the worn tool is shown in Fig. 6, from which the flank wear and wear land can be readily realized. The measurable values (i.e. l1, l2, l3, l4, h1, h2, h3, h4 and h5) can be obtained by amplifying the profile projector screen by drawing the circumference of the tool edge before the cut, and can draw it after wear has occurred. The geometrical lengths, radius of curve and curve angles of the worn tool are shown in Fig. 5, and can be obtained by the following equations: EP = (l21 + h2 1)1/2; PC = (l2 2 + h2 2)1/2; CN = (l23 + h2 3)1/2 and MN = (h2 5 + l24)1/2 (10) uR1 = tan21(l1/h1); uR2 = tan21(h3/l3); uR3 = tan21(h5/l4) and uPC = tan21(h2/l2) (11) R1 = EP/(2sinuR1); R2 = CN/(2sinuR2) and R3 = MN/(2sinuR3); (R2 = R3) (12) Due to the variations in Cs and the different kinds of geometrical circumferences of tool edge, several different cutting conditions occur when the feed is varied. In order to understand the cutting conditions, a criterion for determining the critical length and critical feed are developed and shown in Fig",
+            " Besides the (FH)Umin force, the plowing force FP, due to the effects of tool edge specification, and wear force FW, from the effects of flank wear, are considered to predict the horizontal cutting force (Fig. 15) [14]. That is: (FH)M = FHH(FH)Umin + FP + FW (47) Lf1 = d/cosCs + (l1 + l2)tanCs 2 h1 2 h2 + (2R1uR1 + PC + 2R2uR2 + 2R3uR3)/cosaS2 + (50) [fCM 2 (l1 + l2 + l3 + l4)]/cosaS2\u00b7cos(Ce 2 Cs)] Lf2 = 2R1uR1/cosaS2 + d/cosCs + (l1 + l2)tanCs 2 h1 2 h2 (51) Lf3 = 2(R2uR2 + R3uR3)/cosaS2 + [fCM 2 l1 + l2 + l3 + l4)]/[cosas2cos(Ce 2 Cs)] (52) where Lf1, Lf2 and Lf3 are the contact lengths between the cutting edge and workpiece, as shown in Fig. 5; HB is the Brinell hardness of workpiece; r is the roundness of tip between the main cutting edge and flank; R is the tool nose radius and dB is the tool worn depth (Fig. 12). VB2 and VB3 are the tool flank wear (shown in Figs 13 and 14). The expressions for sy and ty are [19] sy = HB p ,ty = sy 2 (53) The modified transverse cutting force (FT)M is equal to the theoretical transverse cutting force, FT, obtained from Eq. (37), plus a thrust force caused by tool wearing. This thrust force can be estimated by the worn surface area (Lp\u00b7VB) and the yield strength of the workpiece sy",
+            " The shear force can be obtained by worn surface area (Lp\u00b7VB) and shear strength of the workpiece ty. The expression of (FV)M is given in Eq. (55). (FT)M = FT + sy[PC\u00b7dB\u00b7cos(uPC 2 Cs)/2cosaS2] + sy[(LP1VB2 + LP2VB3)] (54) (FV)M = FV + ty[PC\u00b7dB\u00b7cos(uPC 2 Cs)/2cosaS2] + ty[(LP1VB2 + LP2VB3) (55) LP1 = 2R1uR1\u00b7(uR1 + Cs)/cosaS2 + [d/cosCs + (l1 + l2)tanCs 2 h1 2 h2]sinCs (56) LP2 = 2[(R2uR2 + R3uR3)cos(2uR2 + 2uR3 + Ce 2 Cs)/cosaS2] + [fCm 2 (l1 + l2 + l3 + l4)] (57) /cos(Ce 2 Cs)) where, LP1 and LP2 are the projected contact lengths between the tool and workpiece, as shown in Fig. 5. On the basis of Fig. 16, the final modified cutting force can be rewritten for Cs\u00de0\u00b0 as follows: FHH(FH)M = (FH)Umin (58) FVV = (FV)M\u00b7cosCs 2 (FT)M\u00b7sinCs (59) FTT = (FT)M\u00b7cosCs + (FV)M\u00b7sinCs (60) Cutting tests were conducted on a high speed Vector machine tools type (600 \u00d7 700) lathe. Mild steel rods of 60 mm diameter were machined with Sandvik cemented carbide throwaway inserts (WC 69%, TiC 15%, TaC 8%, Co 8%, HV = 174) [20]. The tool geometries are summarized in Table 1, and the workpiece is mild steel (SAE 1045, HB = 295 NN/mm2)"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0000770_iet-cta.2016.0480-Figure1-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0000770_iet-cta.2016.0480-Figure1-1.png",
+        "caption": "Fig. 1. Relative coordinates for desired periodic formation",
+        "texts": [
+            " The specific position is decided by the initial values x0(0) and y0(0). \u2022 The required physically feasible pattern is defined by di\u22121,i, \u03b2i and \u03b3i, i = 1, \u00b7 \u00b7 \u00b7 , N , where di\u22121,i denotes the desired distance between vehicle i \u2212 1 and i, \u03b2i denotes the desired angle difference between vehicle i\u2019s heading and the line of sight which would take directly towards vehicle i \u2212 1, and \u03b3i is the desired heading difference between vehicles i and i \u2212 1. The physical meanings of these symbols are shown in Fig. 1. Throughout the paper, set di\u22121,i > 0, \u03b2i \u2208 [\u2212\u03c0, \u03c0) and \u03b3i \u2208 [\u2212\u03c0, \u03c0). \u2022 The rotation direction and angular speed of the whole multi-agent system can be controlled by parameter w0. When w0 > 0, all the vehicles rotate in counterclockwise and the desired angle difference for vehicle 1 is defined as \u03b21 = \u03c0 2 ; When w0 < 0, all the vehicles rotate in clockwise and \u03b21 = \u2212\u03c0 2 . Denote the radii of the circles as d0i. 3 IET Review Copy Only Content may change prior to final publication in an issue of the journal"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0000602_s11771-015-2876-0-Figure16-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0000602_s11771-015-2876-0-Figure16-1.png",
+        "caption": "Fig. 16 Comparison of aerodynamic noise model at different distances from air inlet to inner ring: (a) Schematic diagram; (b) Distance of 2.84 mm; (c) Distance of 1.94 mm",
+        "texts": [
+            " It seems that the vibration of the mechanical shaft does not change much when the air is supplied or not supplied. Therefore, it can be concluded that the main source of the noise comes from the high speed external air, moving through the vortex inside the bearing cavity, and impinging on the bearing part such as ball and cage. Also, the noise is more likely to be aerodynamically generated noise rather than to be generated by mechanical vibration. To refrain the aerodynamic noise, it is proposed to change the position of the oil outlet. In Fig. 16, a set of noise predicting model is shown. On the left, the center of the air outlet is 2.84 mm to the inner ring; on the right, the center of the air outlet is 1.94 mm to the inner ring. The Williams and Hawkings model [13\u221214] is employed to calculate the noise generated inside the bearing cavity and the signal receiver is set at the same position. The shaft rotating speed is 3.5\u00d7104 r/min. After the sound pressure data are calculated, the sound pressure level is obtained and shown in an octave band mode, as shown in Fig"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0001256_jahs.60.042007-Figure6-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0001256_jahs.60.042007-Figure6-1.png",
+        "caption": "Fig. 6. Tooth stiffness model; tip loaded cantilever beam with average tooth dimensions.",
+        "texts": [
+            " 29) cr = Nv1(tan \u03b1va1 \u2212 tan \u03b1PA) + Nv2(tan \u03b1va2 \u2212 tan \u03b1PA) 2\u03c0 (12) cos \u03b1vai = rvbi / rvai (i = 1, 2) (13) where \u03b1vai, rvbi, and rvai are, respectively, addendum pressure angle, base circle radius, and addendum circle radius of the formative spur gears. The addendum circle radius is obtained by adding ag (tooth addendum) to the pitch radius. Following the works in Ref. 32, the tooth stiffness is estimated by considering the tooth as a tip-loaded cantilever beam with average tooth dimensions, shown in Fig. 6. The tooth stiffness, kt , along the line of action is calculated as kt = E cp3d 32(ag + bg)3 cos \u03b1PA (14) with gear elastic modulus E, tooth dedendum bg = 1.25ag , face width d , and circular pitch cp. The mesh stiffness is obtained by cascading two engaging tooth stiffnesses, and its equivalent model is shown in Fig. 7. In addition, the total mesh stiffness is further reduced by gear body stiffness, bending stiffness, and bearing stiffness. The average mesh stiffness for each mating tooth pair, km, is given by km = \u03bc 1 1 kt + 1 kt (15) where \u03bc is a correction factor to be determined by experiments or further calculations (Ref"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0003568_20020721-6-es-1901.00889-Figure1-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0003568_20020721-6-es-1901.00889-Figure1-1.png",
+        "caption": "Fig. 1. The Pendubot.",
+        "texts": [
+            "eywords: Robot control, Stability, Lyapunov methods, Energy control, Attractor The Pendubot as shown in Fig. 1 is a t w e degree-of-freedom planar robot with single actuator at the shoulder of the first link; the joint of two links is unactuated and allowed to swing free. In addition to other mechanical systems such as inverted pendulum (Astrom and Furuta, 2000), the Acrobot (Spong, 1995), (Berkemeier and Fearing, 1999), (Olfati-Saber and Megretski, 1998), (Zergeroglu et al., 1999), (Brown and Passino, 1997), and brachiating robot (Nakanishi e t al., 1999), such robot is used for research as an example of underactuated mechanical systems (Kolmanovsky and McClamroch, 1995) and for control and robot education, (Spong and Block, 1995)",
+            " To explain specifically, first we present simple formulae of the energy of the Pendubot when the latter case occurs. Then, we show that if two parameters in the control law of (Fantoni e t al., 2000) satisfy a linear inequality, then the former case will occur. In this way, the characteristics of the solution to closed-loop systems with the energy based control law for swing up phase is illustrated further. We recall the result of (Fantoni e t al., 2000) further for describing our result in the next section. With the notation and conventions shown in Fig. 1, from (Spong and Vidyasagar, 1989), (Fantoni e t al., 2000), the equations of motion of the Pendubot are: D(q)ii + C(q, 4.14 + G(q) = 7 (1) where with The object of control is to swing the Pendubot up and balance it to 41 = X/2, q 2 = 0 (6) with 4.1 = 0, 4.2 = 0 (7 ) where (6) holds in the meaning of modulo 27r. The total energy of the Pendubot is given by E = L4.TD(q)4. + Q4gsinq1 + Qsgsin(q1 + q 2 ) (8) 2 The total of energy when the Pendubot is at rest at the vertical, i.e., (6) and (7) hold, is Etop = Q49 + Q59 (9) Define the following Lyapunov function candidate The main result in (Fantoni e t al"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0000255_0954410016641321-Figure14-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0000255_0954410016641321-Figure14-1.png",
+        "caption": "Figure 14. Cooperative tracking using three UAVs with a switching communication topology. (a) Digraph (G1), (b) Digraph (G2), (c) Digraph (G3).",
+        "texts": [],
+        "surrounding_texts": [
+            "Case 1: Fixed topology Simulation results for the following two scenarios of the graph connections as shown in Figure 3 are presented. The target trajectory is pinned into UAV-1 for both cases. Topology 1: A strongly connected digraph G1 with 3 UAVs is considered as shown in Figure 3(a). The adjacency matrix A1 for this case is A1 \u00bc 0 0 1 1 0 0 1 1 0 2 4 3 5: The graph Laplacian matrix is obtained as L1 \u00bc 1 0 1 1 1 0 1 1 2 2 4 3 5: at University College London on May 31, 2016pig.sagepub.comDownloaded from at University College London on May 31, 2016pig.sagepub.comDownloaded from at University College London on May 31, 2016pig.sagepub.comDownloaded from Since the target UAV is pinned into the UAV 1, one can write the matrix B (Cooperative tracking with fixed communication topology section) as B \u00bc 1 0 0 0 0 0 0 0 0 2 4 3 5. The parameters for the target trajectory vgr\u00f00\u00de \u00bc 12 m=s, _ r \u00bc 0:1 deg=s, r\u00f00\u00de \u00bc 25 , _ r \u00bc 0:1 deg=s, r\u00f00\u00de \u00bc 12 , xr\u00f00\u00de \u00bc 100 m, yr\u00f00\u00de \u00bc 100 m, zr\u00f00\u00de \u00bc 100 m: The actual vehicles are initialized with the following values: for UAV1: x1\u00f00\u00de \u00bc 40 m, y1\u00f00\u00de \u00bc 60 m, z1\u00f00\u00de \u00bc 70 m, vg1\u00f00\u00de \u00bc 8 m=s, 1\u00f00\u00de \u00bc 8 , 1\u00f00\u00de \u00bc 8 for UAV2: x2\u00f00\u00de \u00bc 50 m, y2\u00f00\u00de \u00bc 80 m, z2\u00f00\u00de \u00bc 90 m, vg2\u00f00\u00de \u00bc 11 m=s, 2\u00f00\u00de \u00bc 13 , 2\u00f00\u00de \u00bc 13 , and for UAV3: x3\u00f00\u00de \u00bc 60 m, y3\u00f00\u00de \u00bc 100 m, z3\u00f00\u00de \u00bc 110 m, vg3\u00f00\u00de \u00bc 14 m=s, 3\u00f00\u00de \u00bc 18 , 3\u00f00\u00de \u00bc 18 : The actuator constants and gains for each UAV are, c1 \u00bc 10, c2 \u00bc 10, c3 \u00bc 10, 1 \u00bc 5, 2 \u00bc 5, 3 \u00bc 5, lvg \u00bc 5, l \u00bc 5, l \u00bc 5: The simulation results for the communication topology 1 are shown in Figures 4\u20138, where all UAVs reach consensus and track the target trajectory successfully. at University College London on May 31, 2016pig.sagepub.comDownloaded from at University College London on May 31, 2016pig.sagepub.comDownloaded from at University College London on May 31, 2016pig.sagepub.comDownloaded from Topology 2: Consider a group of 4 UAVs having the digraph G2, which is strongly connected as shown in Figure 3(b). The adjacency and graph Laplacian matrices can be written as A2 \u00bc 0 0 1 0 1 0 0 0 0 1 0 1 1 0 0 0 2 664 3 775, L2 \u00bc 1 0 1 0 1 1 0 0 0 1 2 1 1 0 0 1 2 664 3 775 The simulation results for the communication topology 2 are shown in Figures 9\u201313, where 4 UAVs achieve cooperative target tracking successfully as expected. Case 2: Switching topology First, consider three UAVs wherein the communication topology switches to a new topology every 5 s in the following order: G1 ! G2! G3. The total simulation run time is 15 s. The adjacency matrices for the graphs fG1, G2, G3g are, respectively A1 \u00bc 0 0 1 1 0 0 1 1 0 2 4 3 5, A2 \u00bc 0 1 1 1 0 0 1 0 0 2 4 3 5, A3 \u00bc 0 0 1 1 0 0 0 1 0 2 4 3 5 For all the time intervals, the target vehicle is pinned into the vehicle 1 for all networks i.e. the gain matrices are the same in each case. For this particular case, consider the graphs at each interval as strongly connected graphs, i.e. at each interval it has a directed spanning tree. From Figures 16\u201318, one can easily see that all UAVs reach consensus using the cooperative tracking algorithm since the union of three graphs has a directed spanning tree. Further consider a group of four UAVs shown in Figures 14 at University College London on May 31, 2016pig.sagepub.comDownloaded from and 15 where the adjacency matrices for each switching interval are considered as follows A1 \u00bc 0 0 0 0 1 0 0 0 1 1 0 0 0 0 1 0 2 6664 3 7775, A2 \u00bc 0 0 0 1 0 0 0 0 0 1 0 0 1 0 0 0 2 6664 3 7775, A3 \u00bc 0 0 0 0 1 0 0 0 0 1 0 1 0 1 0 0 2 6664 3 7775: Note that each of the topologies at each interval has a directed spanning tree. One can easily see from the simulation results as shown in Figures 19\u201321 that the consensus of target tracking is achieved successfully since the union of the graphs over the interval has a directed spanning tree."
+        ]
+    },
+    {
+        "image_filename": "designv11_60_0000130_ijsnet.2016.075369-Figure2-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0000130_ijsnet.2016.075369-Figure2-1.png",
+        "caption": "Figure 2 Example of the MLDAC problem (see online version for colours)",
+        "texts": [
+            " Each selected link denotes a node pair that will aggregate the data with cooperative communication scheme. To simplify the network management, each node will participate into the cooperative transmission once at most. Then, each node ui costs the power consumption ie by equations (6)\u2013(11), so that all sensor nodes finally aggregate data to the base station. Assume that the initial power of node ui is Ei. The lifetime maximisation is expressed by: max min , .i i i E u e   \u2200    (12) This section illustrates the MLDAC problem by an example. As shown in Figure 2(a), the value attached with each link denotes the minimum power consumption through this link with direct transmission. Assume that all nodes possess the uniformly initial energy, denoted by 410 . We adopt the LPEDAP algorithm to construct the lifetime-optimal tree, which contains the link set 2 1 1 4 5 4 4 3 3, , , , ,{ , , , , }v v v v v v v v v bl l l l l illustrated by Figure 2(b), for data aggregation under the traditional sense. Under the uniform energy distribution, the network lifetime is inversely proportional to the maximum power consumption of all sensor nodes. Among the network, as node v3 costs a maximum power consumption of 50, the network lifetime is 10,000 50 200.= As shown in Figure 2(c), we select two node pairs (v2, v1) and ( 4 3,v v ) for cooperative data aggregation. Obviously, power consumptions of nodes 3v and 4v are much more than those of nodes v1 and v2. For node 3v , the power consumption consists of two parts. One is for data exchanging with node v4, resulting in the power consumption of 10. The other is for cooperative transmission, which costs the power consumption of 12.5 for node v3 by equation (8). Thus, the maximum power consumption in the network is decreased to 22"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0003034_iros.1996.571046-Figure4-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0003034_iros.1996.571046-Figure4-1.png",
+        "caption": "Fig. 4 Calculation of Stability Measure Using I/2/-Norm",
+        "texts": [
+            " If no sliding contacts exist, we can extend our definition of gravity closure to the dynamic case by adding inertial terms to Qhown. This is almost equivalent to the notion of \u201cdynamic closure\u201d shown in [lo]. 3.2 Calculation of stability measure It is time-consuming to calculate the accurate stability measure shown in (3.4). Here we take 1 -norm: l l ~ l l , = CI(W)il (3.5) or -norm: llKAq11, = maxl(KAq)i( (3.6) as IlAqll instead of 2-norm (3.3). Our approximation by 1-norm or -norm makes z larger or smaller than 2-norm, respectively, but the sign of z never changes. Fig. 4 shows a schematic view in a 2-dimensional case and illustrates that we can calculate z by investigating the inscribed \u201ccontour line\u201d of each norm in the polyhedral convex region, Q. Let us consider the case of 1-norm. We define 1; = K [ O ... 0 1 0 ..- OIT \u20acR6 (3.7) A i ( i = 1, ..., 6) and the following linear programming problem subject to maximize z\u2019 z\u2019l, = e,,,, + Wk,\u2019 z\u2019l6 = ehom +Wk,\u2018 - 2\u20191, = ehom +Wk; - ~\u201816 = ehom +Wk{2 k; , ..., k12 2 0 = JTC.kL J J 1 1 1 (i = 1,. . . ,12, j = n + 1,. "
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0000251_maes.2014.130034-Figure1-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0000251_maes.2014.130034-Figure1-1.png",
+        "caption": "Figure 1. 3-DOF helicopter.",
+        "texts": [
+            " The problem considered in this article is the design and evaluation of the robust control law for a small helicopter, which allows for vertical, pitch, and travel rate dynamics tracking reference trajectories. The article is mainly dedicated to robust control aspects of helicopters, where the evaluation of the H\u221e control is addressed to demonstrate the effectiveness of the performance and the robustness of the proposed control law. HELICOPTER MODELING In this section, we present the model of the 3-DOF helicopter portrayed in Figure 1. This particular helicopter model is restricted to 3-DOF, which are the pitch movement, travel rate, and vertical movement [13]. The 3-DOF helicopter consists of a base upon which an arm is attached. The arm has the two ends: the first one includes two direct current (DC) motors with propellers, and the second one represents the counterweight. The arm can evolve around the pitch axis and can swivel the travel axis (Figure 2). To measure these different movements, three encoders are fastened on the helicopter\u2019s body"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0002163_pime_proc_1992_206_062_02-Figure1-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0002163_pime_proc_1992_206_062_02-Figure1-1.png",
+        "caption": "Fig. 1 Schematic illustration of optimization problem",
+        "texts": [],
+        "surrounding_texts": [
+            "Circumferentially butt-welded pipes are frequently used in power stations, offshore structures and process industries. It is well known that the bead width becomes wider as the circumferential welding of small diameter pipes progresses, if constant welding conditions are maintained over the full joint length. In order to obtain a uniform weld bead over the entire circumference of the pipe, the welding conditions should be adjusted and optimized as the welding proceeds. The extensive use of circumferential gas tungsten arc (GTA) welding in industrial applications has been handicapped due to the difficulty of choosing the optimal process parameters. Until quite recently, only limited information on optimum welding conditions was available in the related literature. In order to solve this problem, many attempts were made to estimate the effect of process variables on the weld bead geometries and to determine the optimum process parameters experimentally. Kikushinia and Katsutani (1) developed a system of generating the optimal process parameter for the circumferential arc welding based on a heat conduction analysis, but an Optimization technique was not used to determine it. Ohji et al. (2) developed an algorithm for the determination of the optimal heat input control, but mainly on GTA welding of plates. They (3) also developed a computerized optimal control system of GTA welding of thin plates based on the developed algorithm. In this study, a transient three-dimensional finite difference model (FDM) of the heat conduction flow in the circumferential GTA welding of pipes (4) was adopted for calculating the temperature field and determining the resultant bead width. A mathematical optimization model was proposed to evaluate the optimal welding The M S M'RS recelued on 8 January 1492 and was accepted for puhhcation on I April 1992. current with a given welding velocity for obtaining a uniform bead width required over the entire circumferential joint. Its solution was obtained by employing the steepest descent method, where the initial value of the welding current was estimated by using the linear complementary problem (LCP), and the welding currents in the middle part of the pipe were interpolated by the least-squares mcthod of the second order. The results were experimcntally verified and discussed."
+        ]
+    },
+    {
+        "image_filename": "designv11_60_0000602_s11771-015-2876-0-Figure1-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0000602_s11771-015-2876-0-Figure1-1.png",
+        "caption": "Fig. 1 B7008C angular contact ball bearing model",
+        "texts": [
+            " Considering the shaft rotating speed ranging from 104 to 3.5\u00d7104 r/min during the calculation, the air phase flow inside the bearing was treated as a turbulent flow, and the standard k\u2212\u03b5 model [11] was adopted. Considering the specific geometric features of the B7008C bearing, the numerical model was developed. The B7008C bearing was modeled to analyze the air flow pattern under the ingenious influence of the bearing motion and different cage structures. The specific geometry structure is shown in Fig. 1, containing the inner ring, outer ring, cage, and steel ball. Also, the geometry parameters are listed in Table 1. The size of the geometry is identical to the practical one. The two axial faces were set to atmosphere pressure with zero gauge pressure. The entire flow domain was set under the rotating reference frame, and the steel ball, cage, inner ring, and outer ring were set as moving wall without sliding condition. The motion of ball, cage, inner ring revolving and ball spinning was considered"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0003173_s0141-6359(02)00117-4-Figure1-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0003173_s0141-6359(02)00117-4-Figure1-1.png",
+        "caption": "Fig. 1. Conceptual hardware verification.",
+        "texts": [
+            " Commonly, they have a higher torsional stiffness, like a metallic coupling, but allow considerable angular and axial misalignment, as would an elastomeric coupling. Couplings with this type of \u201chybrid\u201d behavior are sometimes made of composite materials that perform better in harsh chemical environments than would elastomers. However, there is still the problem of how to achieve blind assembly. Numerous concepts for a backlash-free blind-mating coupling were evaluated before the concept discussed in detail here was selected. Six were selected as the most promising as shown in Fig. 1a\u2013f. Table 1 summarizes the dominant physics, risks, and benefits of each concept. The first was a kinematic coupling between shafts (Fig. 1a); this would guarantee constant contact between the shafts, preventing backlash, and would also allow simple blind assembly. However, it would not accommodate misalignment. A contrasting option was a torque finger connection (Fig. 1b); many elastic fingers on the end of each shaft mesh together to transmit torque. This concept uses elastic averaging to average out errors in the fingers\u2019 geometry. This ensures that for modest misalignments or manufacturing errors, mating fingers are always in contact, preventing backlash. Torque fingers would allow looser manufacturing tolerances and easy assembly, but the torque capacity might not be as high as with other alternatives. Two concepts involving compliance were also considered. The first was a compliant spline (Fig. 1c). Leaf spring-like features would be incorporated into the shape of the spline keys (not shown) or keyways (shown) and would be preloaded when the keys were inserted into the key ways. This preload guarantees constant contact, eliminating backlash. This concept would allow one-step assembly, but the leaf spring features would make the coupling more complex to manufacture than a simple spline. Another alternative used elastomeric inserts (Fig. 1d) in a spline connection similar to the current one. Like the elastomeric couplings discussed previously, these inserts would not eliminate backlash, but they would absorb energy and eliminate rattle. The insert could be a resin injected in the form of a liquid under pressure and allowed to cure. However, such inserts could add considerable complexity to the assembly process and for rework should the coupling require disassembly. A further idea (Fig. 1e) also altered the basic spline concept to ensure constant contact, splitting each spline key into two wedges that can slide against each other. As they slide, the effective width of the key changes. Adding an appropriate preload will maintain a key width sufficient for constant contact. The last alternative concept considered was a friction coupling (Fig. 1f). Two shafts would have mating conically shaped ends with an elastomer bonded to their surfaces. An axial bolt would be tightened to create the desired normal force between the two faces, with frictional force being proportional to the normal force. This concept eliminates backlash, but does require the extra assembly step of bolt tightening. Such a design would have little allowance for misalignment. Fig. 1g evolved by combining the positive features of the other couplings along with the concept of elastic averaging and self-locking tapers; and it is the basis of the new coupling concept. The new coupling uses the principle of matched compliance between radially tapered beams and consists of two pieces, either symmetrical or asymmetrical, engaging through fingers which behave as cantilever beams. The ends of the fingers from one piece engage the roots of the fingers from the other, and vice versa. The effect is to have two sets of rigid-compliant engagement beams"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0003063_1.428360-Figure3-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0003063_1.428360-Figure3-1.png",
+        "caption": "FIG. 3. Rotor deflected element.",
+        "texts": [
+            " These supporting bearings, however, are mounted into their rigid housings. The bearing details are shown in Fig. 2. In Fig. 1, the triad XYZ is a global coordinates system with its origin at the geometrical center of the shaft left bearing, where the X axis coincides with the shaft bearing\u2019s center line in the nonworking ~zero speed! position of the system. The orientation of the de- 852awzi M. A. El-Saeidy: Rotating machinery dynamics simulation /content/terms. Download to IP: 129.101.79.200 On: Fri, 29 Aug 2014 09:52:47 Redistr flected rotor element in space ~Fig. 3! is monitored using Euler angles ~Fig. 4!. The elastic rotating shaft is discretized using a C0 four-node isoparametric beam finite element ~Fig. 5! with four degrees of freedom ~DOF! per node: two translational motions plus two total rotations. The C0 beam finite elements were first used in rotor dynamics in Ref. 15, using a weak formulation, where the authors demonstrated the excellent performance of the C0 elements compared to their counterparts of C1, reported in the literature. The author14 has presented a rotating shaft C0 finite-element model using the finite-element displacement method taking into account translational motions, rotary inertia, shear deformations, gyroscopic moments, and mass unbalance forces. The advantages of our formulation compared to those of Ref. 15 are discussed in Ref. 14. In Fig. 4, XYZ is an inertial coordinate system and abc is a body-fixed coordinate system that rotates with the shaft differential element and represents its principal directions where ia , ib , ic are unit vectors along axes a, b, c ~Fig. 3!. x\u0304 y\u0304 z\u0304 is an auxiliary, moving frame system that initially coincides with XYZ. Euler angles are: ~1! rotation c about the X axis results in Y coinciding with y\u0304; ~2! rotation u about y\u0304 results in the moving frame coinciding with ay\u0304z% ; and ~3! spin f about the a axis results in the moving frame coinciding with frame abc. The angular velocity vector is v\u03045v\u0304aia 1v\u0304bib1v\u0304cic . Its components, expressed in the body coordinate system, are14 @v\u0304a v\u0304b v\u0304c# T5@c\u0307 cos u1f\u0307 c\u0307 sin u sin f 1 u\u0307 cos f c\u0307 sin u cos f2 u\u0307 sin f#T"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0003804_iros.2001.973345-Figure4-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0003804_iros.2001.973345-Figure4-1.png",
+        "caption": "Figure 4: Valid velocity centers for right sliding motion. The velocity centers that produce velocity along a given vector lie on the normal to that vector. Points to the right are labeled negative, those to the left, positive. Sweeping the velocity vector from one edge of the cone to the other produces the regions of positive and negative velocity centers shown.",
+        "texts": [
+            " If the contact mode is right or left sliding, then the velocity of the contact point will be a positive linear combination of the palm velocity and velocity along the palm (in the appropriate direction). The exact velocity will be determined by the actual object motion. At this point, we can only eliminate the kinematically inconsistent contact modes and must wait for the stable equilibrium analysis to determine what motion occurs. We can label the points in the plane with the sign of the rotation (if any) that produces a velocity in the cone defined by the palm velocity and velocity along the palm. Figure 4 gives an example of this set of labeled points for a single SR palm contact. In Figure 5, this labeling is combined with the labels for the fixed contacts from Figure 3. 2.4 Stable equilibrium analysis We now test each kinematically consistent contact mode for quasistatic stability and use the results of this test to determine the object motion. We test for quasistatic stability using the force dual method of Brost and Mason [l]. The advantages of using this method are that any number of contact forces can be easily combined and that both taking a force dual and combining two force duals can be described as simple graphical operations"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0001276_j.egypro.2015.03.267-Figure2-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0001276_j.egypro.2015.03.267-Figure2-1.png",
+        "caption": "Fig. 2. The combat vehicle dynamic model scheme",
+        "texts": [
+            " By changing the parameters x \u0307 = y then the function variables becomes: (3) Letting that tribe (i) in x and y is around xi and yi, then the time increment \u0394t, using taylor series in x and y equations become: (4) (5) From equation (4) and (5) above is taken as the average slope of the first derivative, so that the next derivative can be ignored. After was to make use of the Simpson method, the average slope in the interval \u0394t, was changed to: (6) (7) By using the Runge Kutta fourth order, equation (7), the middle part of the equation into two parts so the equation have four parameters, namely as Table 1 below: In Fig. 2, can be viewed that the combat vehicle with rifle caliber 5.56 mm drawing scheme. The vehicle contains platform, azimuth and elevation part. The combat vehicle system has been presumptions: The vehicle platform does not move when firing, backslash are the non linear system can be ignored. All consider forces and moment act in vertical and the system is symmetrical in right and left side. The force and moment acting in elevation part is the value of impact force shot of weapon. Fig. 2 is the dynamic model of combat vehicle with mobile robot LIPI V2 scheme. The considered system has 11 degree of freedom: (11) where, xe is horizontal displacement of elevation part, ye is vertical displacement of elevation part, is angular displacement of elevation part, ya is vertical displacement of azimuth part, \u03b3 is angular displacement of azimuth part, yp is vertical displacement of platform, is angular displacement of platform, y1is vertical displacement of first wheel, y2 is vertical displacement of second wheel, y3 is vertical displacement of third wheel, y4 is vertical displacement of fourth wheel"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0003685_itsc.1997.660569-Figure3-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0003685_itsc.1997.660569-Figure3-1.png",
+        "caption": "Figure 3: Bond graph for torque phase of 2-3 shift",
+        "texts": [
+            " R c i RSi R T ~ ~ B = ---(Tc2 - -wSi - IciWci) For the speed of the r.s.g. to be zero, in the first gear, 1-2 shift, and second gear, the torque on the r.s.g. by the 1-2 band pressure must be larger than the reaction torque transmitted to the r.s.g. Therefore, an additional dynamic constraint is: T t 2 ~ > RTIZB (6) 2.3 Second-to-thrid shift model 2.3.1 Torque phase model. Until the 1-2 band pressure becomes sufficiently small, the speed of the r.s.g. remains zero. Hence, the 2-3 shift torque phase is bond graphed as in Figure 3. In Figure 3, it can be seen that if the 1-2 band pressure is not decreased sufficiently before the third clutch pressure builds up, the turbine torque can be transmitted through either the i.s.g. or the i.c.g. From the bond graph model, the state equations are: Tc3 - R2R3Ts (7) the kinematic constraints are: W t = wci, q = wci 2.3.2 Speed phase moldel. As the pressure of the 1-2 band clutch is decreased, the reaction torque is also decreased. This implies that the dynamic constraint cannot be satisfied. At this instant, the r"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0000134_ecce.2014.6954068-Figure1-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0000134_ecce.2014.6954068-Figure1-1.png",
+        "caption": "Figure 1. Cross section of a line start IPM motor",
+        "texts": [
+            " INTRODUCTION The line start interior permanent magnet (IPM) motor has become an emerging topic of interest among the researchers due to the recent developments in permanent magnet motor technology [1-7]. The line start IPM motor can be a potential replacement of the conventional squirrel cage induction motors because of its high torque density, smaller size, higher efficiency, etc. [1-4]. The conventional line start IPM motor is an induction start synchronous motor. The cross section of a line start IPM motor is illustrated in Fig. 1. The line start IPM motor is connected to a fixed frequency 3-phase balanced ac supply. It starts asynchronously as a squirrel-cage induction motor with the help of the developed cage torque due to the currents flowing in the short-circuited rotor cage windings. As the IPM motor has magnets inserted inside the rotor, a brake or drag torque is also created during run-up due to the magnetic flux created by the permanent magnets and it opposes the cage torque. Fig. 2 illustrates the asynchronous torque vs"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0000726_med.2016.7535981-Figure1-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0000726_med.2016.7535981-Figure1-1.png",
+        "caption": "Fig. 1. Hexacopter",
+        "texts": [
+            " For this purpose, the command governor augmentation is derived for the existing nonlinear dynamic inversion controller and the practical applicability of the presented approach is illustrated by performing flight tests, which include artificial reduction of the rotor control effectiveness. The remainder of this paper is organized as follows: In Section II, the dynamical equations of the hexacopter are derived. Section III reviews the basics of a NDI attitude controller. In Section IV the command governor is derived. Then, Section V presents the experimental results. Finally, a short conclusion is provided in Section VI. For describing the dynamics of the hexacopter we use an inertial frame I and a body-fixed frame B as in Figure 1, such that the origin is at the center of gravity. The rotational dynamics are given in the body-fixed frame B by the Euler\u2019s equation J\u03c9\u0307 =\u2212\u03c9\u00d7J\u03c9 +Mp +Md , (1) where \u03c9 \u2208 R3 is the angular rate of the body-fixed frame relative to the inertial frame, J \u2208 R3\u00d73 is the moment of inertia, Mp \u2208R3 is the propulsion moment and Md \u2208R3 is the disturbance moment. The disturbance moment is a result of unmodeled aerodynamics, wind, parameter errors, etc. In order to describe the propulsion moment Mp, we consider the configuration as in Figure 1 and assume that the motor dynamics are considerably faster than the rigid-body dynamics and can be neglected, that the propellers rotation axis is parallel to the body-fixed z-axis, that the propulsion moments and forces within the rotor plane are small, and that the propulsion moments and forces perpendicular to the rotor plane can be modeled as Fi,z =\u2212kT \u03c92 i , Mi,z =\u2212sgn(\u03c9i)kM\u03c92 i , (2) i.e. the thrust produced by the i-th rotor and the moment about its rotation axis can be modeled such that they are proportional to the square of its angular rate \u03c9i",
+            " (25) Consequently, the extent to which the disturbance \u0394 and hence e\u0308 can be estimated by the command governor is determined by the bandwidth of the noise filter (17) and the choice of the command governor gain \u03bb . In contrast to standard reference models, the overall command yc cannot be assumed bounded as there is a feedback into the reference system due to the additional command yc,CG. For a sketch of the stability analysis refer to [18], [22]. In this section, the proposed controller is evaluated in flight experiments. As a testbed, the AscTec Firefly [33] is used (see Fig. 1). For the experiments, a second order reference model (9) with a natural frequency \u03c90,re f = 5.24[rad/s] and a relative damping \u03b6re f = 1/ \u221a 2 is chosen for all axes (i.e. roll, pitch, and yaw). The gains of the NDI baseline controller ck,0 and ck,1 are designed such that the tracking error (13) decays with a natural frequency \u03c90,re f = 7.75[rad/s] and a relative damping \u03b6re f = 0.84 for all axes. The baseline controller is augmented by the command governor as proposed in Section IV. The additional parameters in the implemented equations (16) - (18) are the command governor gain \u03bb = 5 and the low-pass filter gain \u03ba = 3",
+            " Note that the baseline controller\u2019s performance could be improved by using a more 2In this context, the desired reference model is the original reference model (9) without additional command as this reference model depicts the desired behavior. sophisticated control structure. However, the paper\u2019s focus is on showing that augmentation of a comparatively simple controller by the command governor framework recovers satisfactory performance even in case of severe degradation. For further analysis, actuator degradation is considered. Fig. 3 illustrates the tracking performance of the augmented controller subject to 20% degradation of rotor 3 (refer to Fig. 1), i.e. the rotational speeds of rotor 3 are artificially reduced to approximately 80% during flight by multiplying the commands to the actuators with the control effectiveness in software, implying a loss of 36% of generated thrust. Although a slight decay in performance can be seen with respect to the full-authority flight in Fig. 2, the tracking is still adequate and superior to the baseline performance. In Fig. 4, rotor 3 is further degraded, resulting in 60% rotational speed degradation or approximately 84% loss of generated thrust",
+            " A different choice of command governor parameters can certainly improve the tracking performance and reduce the oscillatory tendency. However, a specified amount of oscillations may hardly be avoided since the command governor only reacts to measurable effects. This is particularly the case when subject to large disturbances as displayed in this example. Finally, Fig. 5 shows the propulsion moments Mp, which are commanded to the rotors according to Eq. (3). Since rotor 3 is severely degraded, primarily the pitch axis is degraded as the angle to the x-axis is 150 [\u25e6] (see Fig. 1). In order to cope with this uncertainty, a negative pitch moment needs to be commanded throughout degraded flight. Similarly, a negative roll moment of smaller amplitude is required to maintain adequate flight. Fig. 5 confirms this observation. In this paper, a command governor augmentation was applied to a NDI baseline controller to cope with exogenous disturbances and parametric uncertainties. Flight tests on a hexacopter show beneficial behavior, both regarding robustness and performance, of the augmented system without relying on reconfiguration to sustain safe flight when subject to actuator degradation"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0000524_978-4-431-55013-6-Figure8.1-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0000524_978-4-431-55013-6-Figure8.1-1.png",
+        "caption": "Fig. 8.1 Stable and unstable quasi-periodic orbits of the coupledmap lattice (8.4). a The parameters are fixed at (\u03b1, \u03b2, \u03b4) = (0.7, 0.02, 0.5), in which the eigenvalues are |\u03bc1| < 1 < |\u03bc2,3|. The solid circle shows the stable quasi-periodic orbit. The gray dot shows the unstable fixed point. The dashed line shows the synchronization set. b (\u03b1, \u03b2, \u03b4) = (0.79, 0.02, 0.06), in which |\u03bc1|, |\u03bc2,3| > 1. The dotted circle shows the unstable quasi-periodic orbit. The period-2 points are stable in this case (shown as black dots)",
+        "texts": [
+            " The Jacobian matrix J \u2217 at the fixed point is represented as follows: J \u2217 = \u03b3 \u239b \u239d 1 \u2212 \u03b2 \u03b2+\u03b4 2 \u03b2\u2212\u03b4 2 \u03b2\u2212\u03b4 2 1 \u2212 \u03b2 \u03b2+\u03b4 2 \u03b2+\u03b4 2 \u03b2\u2212\u03b4 2 1 \u2212 \u03b2 \u239e \u23a0 , where \u03b3 = f \u2032(x\u2217) = 1 \u2212 \u221a 4\u03b1 + 1. The Jacobian matrix has a real eigenvalue and a conjugate pair of complex eigenvalues: \u03bc1 = \u03b3, \u03bc2,3 = \u03b3 2 \u2212 3\u03b2 \u00b1 i \u221a 3\u03b4 2 . We consider two cases in which the fixed point is destabilized by the distinct settings of the eigenvalues: (a) |\u03bc1| < 1 < |\u03bc2,3| and (b) |\u03bc1|, |\u03bc2,3| > 1. We show the dynamics of cases (a) and (b) in Fig. 8.1a and b, respectively. The dashed line shows the synchronization set (xn = yn = zn) in which all synchronization solutions occur. The fixed point exists in the synchronization set (shown as the gray dot). In case (a), since the absolute value of the complex eigenvalue |\u03bc2,3| is greater than one, the Neimark-Sacker bifurcation occurs. Then, the system is desynchronized and the stable quasi-periodic orbit emerges around the unstable fixed point (Fig. 8.1a). In case (b), a quasi-periodic orbit exists similarly to case (a) because |\u03bc2,3| > 1. However, since \u03bc1 is less than -1, the period-doubling bifurcation occurs at the same time. The quasi-periodic orbit is stable on its stable manifold as shown in Fig. 8.1b. However, the outer states except the stable manifold converge to the stable period-2 points. Therefore, this quasi-periodic orbit is unstable and has saddle-type instability. Using the control methods, we aim to stabilize this unstable quasi-periodic orbit. The external force control was proposed by Pyragas to stabilize unstable periodic orbits [13]. The feedback input is defined by the difference between the current state and the external force that is the unstable periodic orbit itself. The control system is defined by: xn+1 = F(xn) + K un, (8"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0002229_jsvi.1996.0017-Figure8-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0002229_jsvi.1996.0017-Figure8-1.png",
+        "caption": "Figure 8. Poincare\u0301 map for b=5\u00b7590 with a=ac , d=0\u00b71, v=1, G=1.",
+        "texts": [
+            " We now discuss some critical values of b, all for a=ac=0\u00b77338 and d=0\u00b71 at which structural changes in the Poincare\u0301 maps occur. (1a) bq5\u00b7590 (Figure 7). There are no transversal manifold intersections of any kind. The Poincare\u0301 map has two unstable fixed points one at either side of the origin which correspond to two unstable periodic orbits in phase-space. (1b) bQ\u22124\u00b7008. The behaviour is as in (1a) above, but with stable fixed points and associated stable periodic orbits (attractors). (2a) b=5\u00b7590 (Figure 8). This is the first interactive tangential manifold contact, by which we mean that a heteroclinic manifold intersects with a homoclinic manifold or vice versa. If we denote by Wu,s ho,he the unstable (stable) homoclinic (heteroclinic) manifolds, then in this case Wu ho is \u2018\u2018outside\u2019\u2019 Ws he . There are no other transversal intersections. This value marks the beginning of a range of such intersections. (2b) b=\u22124\u00b7008. The behaviour is as in (2a) above, but with Ws ho is \u2018\u2018outside\u2019\u2019 Wu he . (3a) 4\u00b7464QbQ5\u00b7590"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0000168_detc2015-46253-Figure12-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0000168_detc2015-46253-Figure12-1.png",
+        "caption": "Figure 12. SIMPLIFIED FINITE ELEMENT MODEL TO INVESTIGATE THE INFLUENCE OF THE HELIX ANGLE [6]",
+        "texts": [
+            " 11. These stresses result from high deformations in the area of the acute facing edge showing lower stiffness. Consequently, high tooth flank loads can result in unfavourable strain and stress in this area which can cause critical ratios of local stresses and local strains [1] and can finally cause facing edge tooth flank fractures. To investigate influence factors of the helix angle on principal stresses, a simplified finite element model in the form of two cylinders was built up as shown in Fig. 12. Different radii of curvature along the path of contact were neglected. The model is based on a gear geometry with a helix angle of \u03b2 = 29\u00b0. The cylinder diameters correnspond to the radii of curvature of the pitch circles. The angle between the end faces and the rotation axis of the cylinder has been varied from 0\u00b0 to 45\u00b0. This angle corresponds to the base helix angle \u03b2b of the gear. The results of the calculation are shown in Fig. 13. As it can be seen in Fig. 13, the maximum value of principal stress first increases up to the base helix angle of \u03b2b = 20\u00b0 - 25\u00b0 and then decreases again"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0000207_s11771-014-2226-7-Figure1-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0000207_s11771-014-2226-7-Figure1-1.png",
+        "caption": "Fig. 1 Definition of coordinate systems",
+        "texts": [
+            " Inverse speed is a special phenomenon distinguished from normal maneuver, but its essence still has close connection with the mechanics of fluids. In order to describe the movement of AUV and to make dynamics analysis, coordinate system which is suitable for the AUV must be established firstly. Based on the recommendation from International Towing Tank Conference (ITTC) and the system of the Society of Naval Architects and Marine Engineers (SNAME), two coordinate systems were built up as follows. These are the fixed coordinate system E\u2212\u03be\u03b7\u03b6 and the moving coordinate system G\u2212xyz, as shown in Fig. 1. According to dynamics theory of rigid body, AUV\u2019s space movement equations, in six degrees of freedom, could be expressed as [16] ( ) ( ) ( ) ( ) ( ) ( ) x z y y x z z y x m u qw rv X m v ru pw Y m w pv qu Z I p I I qr K I q I I rp M I r I I pq N                               (1) On the horizontal plane, rolling motion could be discarded. Let the roll angular velocity and the lateral velocity be zero (p=0, V=0), and then the equations are simplified as ( ) ( ) y m u qw X m w qu Z I q M            (2) When the AUV is on weak maneuver in the vertical plane, \u0394u, w, q, \u03b4b and \u03b4s have small quantities"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0000868_fpmc2016-1767-Figure4-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0000868_fpmc2016-1767-Figure4-1.png",
+        "caption": "Figure 4: Mesh of the Computational Domain Part 2d",
+        "texts": [
+            " The best compromise between the quality of the mesh (distortion of the elements) and the number of elements was found providing that even in the most narrow gaps a minimum of 6 prismatic layers are generated with a minimum of 15 total elements (prism + tetra). Part 1 circuit was cut in two parts at the orifice R1, similarly Part 2 was cut at the orifice R3. The upstream parts (Part 1u, Part 2u) were previously processed with monophase environment, the downstream parts (Part 1d, Part 2d) were meshed and calculated with a Biphase simulation. In Tab. 2 the final mesh characteristic of total domain is reported. The various circuit of lubrication parts have been individually meshed and then merged in the CFX pre-processor. In Fig 4 a detailed view of a sample of the mesh is reported. A multiphase flow system consists of a number of single phase regions which are bound by moving interfaces. In principle, a multiphase flow model could be formulated in 4 Copyright \u00a9 2016 by ASME Downloaded From: http://proceedings.asmedigitalcollection.asme.org/pdfaccess.ashx?url=/data/conferences/asmep/90210/ on 02/18/2017 Terms of Use: http://www.asme.org/about-asme/terms-of-use terms of the local instant variables relating to each phase and matching boundary conditions at all phase interfaces"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0001489_jrproc.1954.274732-Figure1-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0001489_jrproc.1954.274732-Figure1-1.png",
+        "caption": "Fig. 1-The transducer diagram.",
+        "texts": [
+            " To a great extent their value stems from the graphical nature of the procedure. It enables a quick and clear grasp of the transformation to be obtained, which is not afforded by the numerical methods using complex numbers. The present communication, which deals with a more general problem than that of the uniform transmission line, presents a graphical procedure for transforming impedances through an arbitrary loss-free transducer. For any such transducer there exists a particularly simple diagram (Fig. 1) based on ideas of projective geometry, which, when drawn, allows the input impedance corresponding to any terminating impedance to be quickly deduced. To the knowledge of the writer, the only projective diagram previously published for performing this transformation is that of Weissfloch.' The latter is considerably more cumbersome to use and more susceptible to graphical errors than the one described below. The relation between the two diagrams is demonstrated later. The present diagram was originally described in a restricted wartime report",
+            " Bracewell, \"A graphical means of investigating loss-free transducers,\" Radiophysics Laboratory Report T.I. 226, 1945. Let these be A, B, C, respectively. In practice the terminating reactances would often include 0 and o, i.e., measurements would be made with the transducers short-circuited and open-circuited. Now let the quantities a, b, c, A, B, C, be marked off on two linear scales whose scale factors and origins may be chosen at will to suit the values of the quantities. Lay one scale on the other, at any convenient angle, so that A and a coincide (Fig. 1). Then mark the point P which lies on the intersection of Bb and Cc. The two graduated scales and the point P then comprise the desired diagram. We shall now prove that if the transducer is terminated with any reactance d, then the input reactance is D, where D is the intersection of dP with AB. We shall assume without proof the following proposition from projective geometry. If the points of two straight lines a and v are in (1, 1) correspondence, the general analytical relation between co-ordinates x",
+            ", C + dZ2 zi = t a + bZ2 where a, b, c, d, are the \"linear parameters\" as customarily defined. As a particular case of this, the input impedance to a loss-free transducer terminated in a reactance X2 is itself reactive (=jX,), and the relation between the reactances Xi and X2 is bilinear with three independent real constants. But this is a projective relationship, and therefore if it is made to agree with a particular projective transformation in three independent instances, then it must do so in all. Hence the points d and D in Fig. 1 represent corresponding reactances. Although the data specifying the transducer have been presented in the form of electrical measurements of a certain kind, the transformation diagram may also be set up from any other specification of the transducer such as the equivalent-T impedances, or short and open circuit impedances; and conversely, beginning from the diagram, any other specification may be deduced. It is sufficient to note the simple constructions which give the open and short-circuit impedances Z1, Zo2, Z,1, Zs2 (Fig"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0000996_lars-sbr.2016.21-Figure3-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0000996_lars-sbr.2016.21-Figure3-1.png",
+        "caption": "Figure 3. FFSM with the DEMs that will map the two end-effectors.",
+        "texts": [
+            " W\u0394(s) = \u23a1 \u23a2\u23a2\u23a2\u23a3 W\u0394,1(s) 0 \u00b7 \u00b7 \u00b7 0 0 W\u0394,2(s) \u00b7 \u00b7 \u00b7 0 ... ... . . . ... 0 0 \u00b7 \u00b7 \u00b7 W\u0394,n(s) \u23a4 \u23a5\u23a5\u23a5\u23a6 W\u0394,i(s) = s+ \u03c9bc Mu(\u03b5s+ \u03c9bc) (17) where \u03c9b is the controller band width, Mu is the maximum gain and \u03b5 > 0. We and W\u0394 parameters must be selected in a way that the following conditions are satisfied [8]: ||S(s)||\u221e \u2264 ||W\u22121 e ||\u221e ||T (s)||\u221e \u2264 ||W\u22121 \u0394 ||\u221e (18) where S(s) is the sensibility function and T (s) is the complementary sensitivity function. The control objective is to make the FFSM, initially with both arms bent as shown in Fig. 3, open its arms completely so that the end-effectors are as far as possible from each other. In other words, all joint angles must be zero at the end of the simulation, and with the specified deadline of 10 seconds for accomplishing the task. For the LQR controller it will be inicially studied the nominal case, then disturbance torques and uncertainties in inertia matrix will be inserted so it will be possible to better analyze the resulting tracking errors, torques and joint angles. For the H-infinity controller the scenario with disturbance torques on the joints and uncertainties in the inertia matrix will be considered"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0002974_robot.1994.351104-Figure2-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0002974_robot.1994.351104-Figure2-1.png",
+        "caption": "Figure 2: Definitions for A-type Contacts.",
+        "texts": [
+            " The variables c and s are to be thought of as independent variables and are used to represent the orientation of the workpiece in place of 0, so that the contact constraints may be written as algebraic, rather than trigonometric, equations. This, however, will require the introduction of an additional algebraic constraint, namely, 2 + 2 - 1 = 0. In this spirit, we define the system\u2019s modified Cspace, 2, to be the set of all poesible vectors p, and denote the three contact constraints by Cr(p) = 0 for 1 = 1,2,3. We then define the resulting CF-cell, CF, as follows: C3 = {p E 2 IC\u2019 + s2 - 1 = 0 and G(p) = 0 (4) Consider Figure 2. Let 91 denote the angle between the outward normal to edge 1 and the positive z-axis of W . The signed distance between vertex 1 and the line supporting edge I, the so called Gfunction [6], is given by: (.j) C i ( p ) = V I . (ai - bl) be rewritten M: Di(r, C, 8 ) + El(r, C, 8) t + 4(r, c, 8)y = 0 ; (8) 1 = 1,2,3 c'+s'- 1 = 0 (9) The first two of the three equations represented by equation (9) can be solved for t and y yielding: where the vectors .r , bi, and VI are @ran in [4]. Here 4 ie the jxasition of vertex I , br L the position of the selected point on edge I , aad VI U the outward unit normal of edge I ; all d there quan t i tk u e expressed with tespect to the f r u ~ U",
+            " Notice that we haw labeled the deea of the workpiece and the p o l y g o ~ 10 that contact 1 is between polygon 1 of the manipulator and the line supporting edge 1 of the workpiece. Note .Is0 that we handle the awes in which an edge of the workpiece is to be contacted by more than one verkx of the manipulator by labeling the edge more than once. The Gfunction correaponding to the l r ~ ty e A contact can be written M a function of the con&uration of the workpiece to yield the following system of contact constraint equations: where 1 = 1,2,3. The coefficients 01, )I, dr, er, and f i , illustrated in Figure 2, are functions of the geometry of the bodiee in the aystem (including the workpiece) and the configuration of the manipulator and are given in [4]. Ea& Gfunction, Cr(p), M well aa the unit circle equation (equation 7)), dehee a quadric hypemur- per surfaces defines the set of eomeiricully admissible 3A codgurationa of the woripiece as a function of face in Gspaw. d e intersection of these four hy- meane that contact is maintained along the line sup porting the edge, not the actual physical edge of the workpiece",
+            " If the system under consideration is planar with n, contacts, then the wrench matrix W, can be partitioned into normal and tangential components W, and Wt, which have size (3 x nc). The wrench matrices have the following form: where nl is the inward pointing unit normal vector to edge l , - t l is the tangential unit vector defined so that ni x tr points out of the page, pr is the position of the contact point, and pI A n r is given by p1=nr, - pI,,nr=. Here PI, and P I , are the components of PI. The components of nl are defined analogously. Given the geometric definitions of the coefficients shown in Figure 2, the wrench matrices can be rewritten as explicit functions of the system configuration In [4] it is shown that G and H are related to the determinants of W, and Wt,vcrt, by the following simple formulas: Det(OW,) = C H - S G where the superscript indicates that the matrices are expressed with respect to frame U. Note that this definition of oW;,verfr.uses for the pi, the positions of the vertices of the manipulator polygons designated to contact the workpiece even though contact may be impossible for the manipulator configuration under considerat ion"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0003105_1.1326443-Figure1-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0003105_1.1326443-Figure1-1.png",
+        "caption": "Fig. 1 Main part of EHD tester",
+        "texts": [
+            " Manuscript received by the Tribology Division November 18, 1999; revised manuscript received June 30, 2000. Associate Editor: B. O. Jacobson. 54 \u00d5 Vol. 123, JANUARY 2001 Copyright \u00a9 rom: http://tribology.asmedigitalcollection.asme.org/pdfaccess.ashx?url= stress characterizing the non-Newtonian are obtained and are discussed in relation to kinds of refrigerant and refrigerant concentration in POE. Film thickness and the coefficient of traction were determined with an EHD tester in the autoclave. Its main part is shown in Fig. 1. The principle of measurement of the traction is basically the same as that reported from Gunsel et al. @13# except for the method of giving the sliding speed. The details of the apparatus are described elsewhere @11#. Two examples of the measurements are shown in Fig. 2, in which two curves for each refrigerant are obtained by reversing the rotational direction of the disk. In an experiment the traction curve slightly shifted downwards on the whole as the load was increased. In that case, the traction curve was corrected by shifting the two measured traction curves upward until the absolute values of the coefficient of traction at a slide-roll ratio of 3 percent became the same and symmetric relative to the abscissa"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0000524_978-4-431-55013-6-Figure10.7-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0000524_978-4-431-55013-6-Figure10.7-1.png",
+        "caption": "Fig. 10.7 Cart-trailer system is composed of a heading cart with articulated trailers, which undergoes holonomic constraints due to rigid linkage",
+        "texts": [
+            " Using J+ and J\u2212, we can define the following two branches of Lie brackets: [g1, g2]+(\u03be) := J+(g2)g1 \u2212 J+(g1)g2 = \u239b \u239d \u03b2S0 + \u03b1C0 \u2212\u03b2C0 + \u03b1S0 0 \u239e \u23a0 , [g1, g2]\u2212(\u03be) := J\u2212(g2)g1 \u2212 J\u2212(g1)g2 = \u239b \u239d \u03b2S0 \u2212 \u03b1C0 \u2212\u03b2C0 \u2212 \u03b1S0 0 \u239e \u23a0 . Their values at \u03be = 0 are: g1(0) = \u239b \u239d 1 0 0 \u239e \u23a0 , g2(0) = \u239b \u239d 0 0 1 \u239e \u23a0 , [g1, g2]+(0) = \u239b \u239d \u03b1 \u2212\u03b2 0 \u239e \u23a0 , [g1, g2]\u2212(0) = \u239b \u239d \u2212\u03b1 \u2212\u03b2 0 \u239e \u23a0 , which are consistent with the actual displacements shown in Fig. 10.6. In this section, we consider planar locomotion of multiple rigid bodies connected to each other. Suppose a cart towing trailers as shown in Fig. 10.7(left). Each of the carts 0, . . . , \u2212 1 has a free joint on the center of its wheel axis, which connects the following cart to itself. The length of each connecting link is supposed to be 1. The state vector is \u03be = (x0, y0, \u03b80, . . . , \u03b8 ) T \u2208 X , X := SE(2) \u00d7 T \u22121, where (x0, y0) denotes the position of the truck (cart 0) and \u03b8i denotes the orientation of the cart i for i = 0, . . . , . This system undergoes +1 nonholonomic constraints y\u0307i cos \u03b8i \u2212 x\u0307i sin \u03b8i = 0, i = 0, . . . , \u2212 1 (10.11) and holonomic constraints of rigid linkage as well: { xi = xi+1 + cos \u03b8i+1, yi = yi+1 + sin \u03b8i+1, i = 0, ",
+            " By taking all the constraints into account, the state equation is obtained as \u03be\u0307 = g1(\u03be)u1 + g2(\u03be)u2, (10.12) g1(\u03be) := \u239b \u239c \u239c \u239c \u239c \u239c \u239c \u239c \u239c \u239c \u239d cos \u03b80 sin \u03b80 0 \u2212 sin(\u03b81 \u2212 \u03b80) \u2212 sin(\u03b82 \u2212 \u03b81) cos(\u03b81 \u2212 \u03b80) ... \u239e \u239f \u239f \u239f \u239f \u239f \u239f \u239f \u239f \u239f \u23a0 , g2(\u03be) := \u239b \u239c \u239c \u239c \u239c \u239c \u239c \u239c \u239c \u239c \u239d 0 0 1 0 0 ... \u239e \u239f \u239f \u239f \u239f \u239f \u239f \u239f \u239f \u239f \u23a0 , where u1 is the forwarding velocity and u2 is the heading angular velocity of the truck (cart 0). It is easy to show that this system is also controllable by analyzing its controllability Lie algebra. Now let us turn to consider the discrete counterpart (Fig. 10.7, right). Each cart is placed on the hexagonal cells, thus each joint angle is the difference between adjoining cart orientation, e.g., \u03b8i+1 \u2212 \u03b8i . We also assume that the joint angles are limited to |\u03b8i+1 \u2212 \u03b8i | < 3, i = 0, . . . , \u2212 1. The state vector is \u03be = (x0, y0, \u03b80, . . . , \u03b8 ) T \u2208 X , X := SEH (2) \u00d7 Z \u22121 6 . Control inputs are assigned to the velocity of the trucks, i.e., u1 is the forwarding velocity and u2 is the heading angular velocity of the front cart, respectively: x0C0 + y0S0 = u1, \u03b80 = u2of"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0001142_raha.2016.7931891-Figure3-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0001142_raha.2016.7931891-Figure3-1.png",
+        "caption": "Fig. 3. Isometric odel",
+        "texts": [
+            " End Caps are machined from a thick block of propylene to fit the acrylic tube tightly and a groove is made to fix the O-rings to seal the tube properly. 978-1-5090-5203-5/16/ $31.00 \u00a92016 IEEE Fig. 2. Top View of Model Holes are drilled on it to accommodate for the screw threads to pass through. Polypropylene sheets are used in construction of the ROV outer structure in order to give it high strength, rigidity, durability and less weight. The sheet was cut out and holes were drilled as per design in order to accommodate the hull, battery pod and the motors. The top view and isometric view of CAD model of the ROV are shown in Fig. 2 and Fig. 3 respectively. The completely assembled ROV system with components mounted on it is shown in the Fig.4. B. System Modeling the speed the diameter (0.1 m) of the propeller, the distance between thruster 2 and 3 (0.15 m). For the thrust coefficient Kt = 0.35 (from datasheet using P/D ratio 1.4) and , the viscous factor of water (1.002 * 10-3 Ns/m2), the thrust (1) Hence maximum thrust from a single thruster having four blades with high angles is, (2) (3) Maximum linear acceleration of ROV = 2*a = "
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0000759_fpmc2016-1739-Figure2-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0000759_fpmc2016-1739-Figure2-1.png",
+        "caption": "Figure 2. CAD picture of the test chamber.",
+        "texts": [
+            " A secondary cycloidal gearbox with a transmission ratio of 43 can be included additionally. Thus, a wide range of rotational speeds is possible: 0.2 - 2.4 rpm (rotation per minute) with or 8.5 \u2013 105 rpm without the additional gearbox. The revolution speed of the drive shaft is measured using an incremental angle encoder (90000 increments/revolution, accuracy \u00b1 5 seconds of degree). The rotating cylinder and the seal specimen are located in the test chamber. A detailed view of the test chamber and the force measurement arm is shown in Figure 2. The lateral surface of the rotating steel cylinder is brought into contact with the seal specimen. The cylinders diameter D of 200 mm leads to a relative velocity v at the linear contact of 2 - 25 mm/s or 0.09 \u2013 1 m/s, respectively with or without the cycloidal gearbox. 2 Copyright \u00a9 2016 by ASME Downloaded From: http://proceedings.asmedigitalcollection.asme.org/pdfaccess.ashx?url=/data/conferences/asmep/90210/ on 02/25/2017 Terms of Use: http://www.asme.org/about-asme/terms-of-use The test chamber can be entirely flooded with a lubricant, at environment pressure",
+            " A force measurement arm is used for applying the normal load and measuring the resulting friction force. The force sensor position is adjustable. Therefore, by changing the effective lever arm, the sensor can measure a wide range of friction forces. An adjustable counterweight is used to guarantee a normal force of 0 N when no test weight is applied. Hence the load does not depend on the sensor position. The whole test chamber is double-walled. A heat transfer fluid, provided by an external temperature control unit, flows through a temperature chamber (Figure 2). Thus the tribometer is temperature-controllable between 10\u00b0C and 100\u00b0C. A temperature sensor is located near to the tribological contact in order to verify a constant temperature during the measurements. Three cylinders with different surfaces were used during the experiments. All three specimens lateral surfaces were sandblasted after the last manufacturing process. Thus an ideally-uniform initial condition was produced prior to the following anisotropic surface roughness production process. Afterwards anisotropic surfaces were generated on two cylinders"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0003219_s0092-8240(87)80026-5-Figure8-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0003219_s0092-8240(87)80026-5-Figure8-1.png",
+        "caption": "Figure 8. The eigenvalues 2~ (1 = 0, 1, 2, 3, 4) as a function of the curvature K (cf. Fig. 3).",
+        "texts": [
+            " (15) Then the curvature has the form up to first order oft , ~Co2(/+2) ( l - 1) Y~/(0, ~)e ~t. (16) x \u00b0 ~ x \u00b0 + ~ 2 Equation (16) is substituted into equation (14) and after the linearization one finds the eigenvalue 2 in the following form, Ko2(l+ 2)(l-- 1) 21 = 2(1 + h(tco)- C 1 - KoI(t\u00a2o) ) { - ( y I ' +g ' ) -exg l ( l+ l)h'}, (17) where the prime means the derivative d/dK o . One can see immediately that 2~ = 1 = 0, which corresponds to a translational invariance of the protocell model. All other modes are calculated numerically and an example is given in Fig. 8; again a good correspondence to Fig. 3 is obtained. 5. A Preliminary Numerical Result. A numerical analysis can be performed in the following way. The time evolution of the curvature is given by the geometrical relation, = - ( K2 + A s ) Vn. (18) Equation (18) together with equation (14) enables us to calculate the time evolution of the surface. Here it should be noted that the arclength changes also by the geometrical relation, equation (11). The numerical procedure we have used is the same as in the work of Brower et al"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0001113_apede.2016.7879015-Figure1-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0001113_apede.2016.7879015-Figure1-1.png",
+        "caption": "Fig. 1. Vibration generator - external view",
+        "texts": [
+            " The vibrator consists of two main systems: - moving system; - magnet system. Both systems are joined with attachements. The magnet system comprises: - the magnet path; - two permanent magnets. The moving system consists of: - two coils of copper wire on paper carcasses; - stem pusher, used for transferring the vibration impact; - suspensions. The measured element is located on the one suspension of the moving system, constructed in the form of a radial centering spring. The external view and structure of the vibration generator are shown in Fig. 1 and 2. Vibration generators are fixed at the base of the positioning system in such a way that the axes of stem pushers coincide with the fixture holes of the tested object. The vibration generator is used for producing the speed of vibration-force impact of linear movements. Piezotransducer is made with the plate spring, bearing two piezoresistive strain gages, mounted according to Wheatstone bridge scheme. Plate bending causes the strain gages deformation. Piezoresistive accelerometers, that use resistive sensors with low impedance, are insentitive to external ghost influences and disturbances"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0003840_acc.2000.879454-Figure3-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0003840_acc.2000.879454-Figure3-1.png",
+        "caption": "Fig. 3 Roller Tilting Mechanism",
+        "texts": [
+            " According to the model by Tevaarwerk and Johnson [12], the normalised traction force and side slip force are given by the following equations if the influence of spin is neglected: where, In the formulation above, the fluid film thickness has been cancelled out. The two parameters, namely, the slope of the traction curve, m, and the maximum traction coefficient, pmax, are obtained from experiment as mentioned previously. The creep rate R, and side slip rate R, depend on the CVT geometric and kinematic parameters and their detailed derivation is available in [ 113. 4. CONTROL OF THE TRACTION CVT 4.1 RATIO CHANGE MECHANISM The ratio change mechanism of a traction drive CVT is illustrated in Fig. 3. As shown in the figure, the coordinate system XYZ has its origin at point 0, which is the cavity center. The Y-axis is parallel to the disk axis and the Z-axis bisects the cavity. During steady state CVT operation, the swing center of the roller coincides with the cavity center. The ratio change is realized by varying the tilt angle 6. of the roller axis with respect to the Z-axis. The swing of the roller is indirectly realised by slightly offsetting the roller axis in the direction of X-axis. The amount of offset, which is controlled by computer, creates a side slip at the diskroller contact, and this side slip induces a side slip force that turns the roller about the swing center to change the ratio"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0002252_s0890-6955(98)00006-6-Figure4-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0002252_s0890-6955(98)00006-6-Figure4-1.png",
+        "caption": "Fig. 4. Detailed model of the shear plane area A and friction area Q of Fig. 3.",
+        "texts": [
+            " f) according to the depth of cutting, which can be subdivided into three parts (a) d . R, (b) d = R and (c) d , R. According to the above conditions, case (1) had been discussed in reference [15]. This paper will therefore focus on the case of small radius, i.e. case (2). However, an experiment has been performed to study the case of large nose radius cutting, i.e. case (3), the results of which will be presented in future. Therefore, the shear plane and the projected area of cutting cross-section for a chamfered main cutting edge tool with nose radius, are redrawn in Fig. 4. The area of shear plane A and the friction plane Q for case (2) can be obtained as follows: For convenience of calculation, the shear plane must be projected in the plane perpendicular to the speed of cut, which makes the calculations and analysis much easier and saves the time required for calculations. Defining the chip flow angle in this perpendicular section as hc9, the relation between hc9, and hc on the tool face is presented in Appendix Eq. A(6). According to Eq. A(6), the shear plane area can be varied by changing hc9 in small increments",
+            " 9, the intersection angle uSS can be calculated as: uSS = Ce 2 Cs + sin21(mm/R1) (20) where mm = (l1 + l2 + l3 + l4)sin(Ce 2 Cs) + (h1 + h2)cos(Ce 2 Cs) + (R2cos2uR2 2 R3cos2uR3)\u00b7cos(Ce 2 Cs) 2 fsinCe 2 R1sin(Ce 2 Cs) (21) Thus yielding modified feed as: f CM = fcosCs + R1(1 2 cos2uss) + WW9cosuPC (22) After the modified feed fCM is obtained, the shear plane area and the projected area on the tool face can be calculated from Figs 4 and 10, as: A = A1 + A2 + A3 + A4 + A5 + As (23) where A1 = 0.25[4a2 1n2 1 2 (a2 1 + n2 1 2 c2 1)2]1/2; (24) A2 = 1/coshc9\u00b7 E 2uR2 0 \u00b7[fCM 2 (l1 + l2) + l3 2 R2sin(2uR2 2 F)]ds + E 2uR3 0 \u00b7[fCM 2 (l1 + l2 + l3) + R2sin(F)]ds; (25) A3 = PC/(2coshc9)\u00b7[2(fCM 2 l1) 2 l2]sin(hc9 + uPC)const1; (26) A4 = E 2uR1 0 [(fCM 2 R1 + R1cosF)/coshc9]ds; (27) A5 = 1/2\u00b7(k1 + i1)j1\u00b7const1; (28) As = 0.5\u00b7W2 ecos2as1tanCs/cosabsinfe (as shown in Fig 4). The projected area on the tool face is derived as: Q1 = [d/cosCs + (l1 + l2)tanCs]\u00b7fCM + 1/2\u00b7(l3h3 + l2h4 2 l2h2 2 l1h2) + (l4h3 2 l4h4 2 l1h2) + 0.5R2 1(2uR1 2 sin2uR1) 2 0.5[R2 2(2uR2 2 sin2uR2) 2 R2 3(2uR3 + 0.5f 2 CMtanCs + [h3 2 h4 2 0.5\u00b7(fCM 2 l1 2 l2 2 l3 2 l4)\u00b7tanCs] (29) (fCM 2 l1 2 l2 2 l3 2 l4); Q2 = (0.5\u00b7W2 ecosas1tanCs)/cosab\u00b7(Q3 is the area of triangle DD9Y, Fig. 4) The expressions for a1, n1, c1, e1, g1, const1, h1, i1, F and j1 are given in Appendix B; while ds and hc9 are given in Appendix A. The shear energy per unit time Us and the friction energy per unit time Uf can be determined from the following equations: Us = FsVs = (tsAcosae\u00b7V)/cos(fe 2 ae) (30) Uf = FtVc = ftE B1 0 dbVc = (tssinbcosaeQ\u00b7V)/[cos(fe + b 2 ae)cos(feae)] [13] (31) in which Fs = ts\u00b7A; Vs = (V\u00b7cosae)/cos(fe 2 ae) Vc = (V\u00b7sinfe)/cos(fe 2 ae) ft = (ts\u00b7f\u00b7cosCs\u00b7sinb)/[cos(fe + b 2 ae)sinfe] (32) and U = Us + Uf = V\u00b7(FH)Umin The shear areas A for a worn nose radius tool is calculated according to Eqs (6), (23)\u2013(28), and the friction area Q is calculated from Eqs (9) and (29)",
+            " From Figs 12\u201314, the types of wear can be seen from the views of tool edge and that of the main and front edges and a simple relationship between flank wears VB2, VB3 and worn depth of the tool edge dB can be obtained. dB = S1S2\u00b7cosae(coturef 2 tanae) (44) VB2 = PP9(coturef1 2 tanaS2) (45) VB3 = CC1(coturef2 2 tanae) (46) whereuref1 is the side relief angle on the main cutting edge (Fig. 13)uref2 is the side relief angle on the front cutting edge (Fig. 14) A modified cutting force model included the effects of size, shape and tool edge wear, is presented in this study and it based on the force model in Fig. 4. Besides the (FH)Umin force, the plowing force FP, due to the effects of tool edge specification, and wear force FW, from the effects of flank wear, are considered to predict the horizontal cutting force (Fig. 15) [14]. That is: (FH)M = FHH(FH)Umin + FP + FW (47) Lf1 = d/cosCs + (l1 + l2)tanCs 2 h1 2 h2 + (2R1uR1 + PC + 2R2uR2 + 2R3uR3)/cosaS2 + (50) [fCM 2 (l1 + l2 + l3 + l4)]/cosaS2\u00b7cos(Ce 2 Cs)] Lf2 = 2R1uR1/cosaS2 + d/cosCs + (l1 + l2)tanCs 2 h1 2 h2 (51) Lf3 = 2(R2uR2 + R3uR3)/cosaS2 + [fCM 2 l1 + l2 + l3 + l4)]/[cosas2cos(Ce 2 Cs)] (52) where Lf1, Lf2 and Lf3 are the contact lengths between the cutting edge and workpiece, as shown in Fig"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0000622_b978-0-12-800782-2.00003-8-Figure3.10-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0000622_b978-0-12-800782-2.00003-8-Figure3.10-1.png",
+        "caption": "Figure 3.10, cont'd (d) rotor position #4, and (e) rotor position #5.",
+        "texts": [
+            " 218 Power Quality in Power Systems and Electrical Machines The small-slip approximation results in the torque equation T \u00bc q1 V 2 s s \u03c9s1 R0 r : a) Applying the relation to rated conditions (1) one gets T 1\u00f0 \u00de \u00bc q1 V 2 1\u00f0 \u00de s s 1\u00f0 \u00de \u03c9 1\u00f0 \u00de s1 R0 r : Applying the relation to low-voltage and low-frequency conditions (2) one gets T 2\u00f0 \u00de \u00bc q1 V 2 2\u00f0 \u00de s s 2\u00f0 \u00de \u03c9 2\u00f0 \u00de s1 R0 r : With T(1)\u00bcT(2) and the relations \u03c9 1\u00f0 \u00de s1 \u00bc 2\u03c0f 1\u00f0 \u00de p=2 and \u03c9 2\u00f0 \u00de s1 \u00bc 2\u03c0f 2\u00f0 \u00de p=2 one obtains s 2\u00f0 \u00de \u00bc s 1\u00f0 \u00de V 1\u00f0 \u00de s\u00f0 \u00de2 V 2\u00f0 \u00de s\u00f0 \u00de2 f 2\u00f0 \u00de f 1\u00f0 \u00de \u00bc 0:0367: b) The synchronous speed ns (2) at f (2)\u00bc59 Hz is n 2\u00f0 \u00de s \u00bc 120 59\u00f0 \u00de 6 \u00bc 1180rpm, or the corresponding shaft speed n 2\u00f0 \u00de m \u00bc n 2\u00f0 \u00de s 1 s 2\u00f0 \u00de \u00bc 1137rpm: c) The torque-output power relations are T 1\u00f0 \u00de \u00bc P 1\u00f0 \u00de out \u03c9 1\u00f0 \u00de m and T 2\u00f0 \u00de \u00bc P 2\u00f0 \u00de out \u03c9 2\u00f0 \u00de m : With T(1)\u00bcT(2) one gets P 1\u00f0 \u00de out \u03c9 1\u00f0 \u00de m \u00bc P 2\u00f0 \u00de out \u03c9 2\u00f0 \u00de m or P 1\u00f0 \u00de out \u03c9 1\u00f0 \u00de s1 1 s 1\u00f0 \u00de\u00f0 \u00de \u00bc P 2\u00f0 \u00de out \u03c9 2\u00f0 \u00de s1 1 s 2\u00f0 \u00de\u00f0 \u00de or P 1\u00f0 \u00de out 120f 1\u00f0 \u00de p 1 s 1\u00f0 \u00de \u00bc P 2\u00f0 \u00de out 120f 2\u00f0 \u00de p 1 s 2\u00f0 \u00de : For the reduced output power it follows P 2\u00f0 \u00de out \u00bc 1 s 2\u00f0 \u00de 1 s 1\u00f0 \u00de\u00f0 \u00de f 2\u00f0 \u00de f 1\u00f0 \u00deP 1\u00f0 \u00de out \u00bc 97:65hp: d) The copper losses of the rotor are Pcur \u00bc q1 I 0r 2 R0 r and the output power is Pout\u00bc q1 I 0r 2 R0 r 1 s\u00f0 \u00de s : The application of the relations to the two conditions (1) and (2) yields P 1\u00f0 \u00de out\u00bc q1 I 0r \u00f01\u00de 2 R0 r 1 s 1\u00f0 \u00de s 1\u00f0 \u00de and P 2\u00f0 \u00de out\u00bc q1 I 0r \u00f02\u00de 2 R0 r 1 s 2\u00f0 \u00de s 2\u00f0 \u00de or I 0r \u00f02\u00de 2 I 0r \u00f01\u00de 2\u00bc P 2\u00f0 \u00de out P 1\u00f0 \u00de out s 2\u00f0 \u00de s 1\u00f0 \u00de 1 s 1\u00f0 \u00de 1 s 2\u00f0 \u00de\u00f0 \u00de \u00bc 1:20: The rotor loss increase becomes now P 2\u00f0 \u00de cur \u00bc I 0r \u00f02\u00de\u00f0 \u00de2 \u00bc 1:20: P 1\u00f0 \u00de cur I 0r \u00f01\u00de\u00f0 \u00de2 219Modeling and Analysis of Induction Machines One concludes that the simultaneous reduction of voltage and frequency increases the rotor losses (temperature) in a similar manner as discussed in Application Example 3.1. It is advisable not to lower the power system voltage beyond the levels as specified in standards. Numerical approaches such as the finite-difference and finite-element methods [16] enable engineers to compute no-load and full-load magnetic fields and those associated with short-circuit and starting conditions, as well as fields for the calculation of stator and rotor inductances/reactances. Figures 3.8 and 3.9 represent the no-load fields of four- and six-pole induction machines [17,18]. Figure 3.10a\u2013e illustrates radial forces generated as a function of the rotor position. Such forces cause audible noise and vibrations. The calculation of radial and tangential magnetic forces is discussed in Chapter 4 (Section 4.2.14), where the concept of the \u201cMaxwell stress\u201d is employed. Figures 3.11 to 3.13 represent unsaturated stator and rotor leakage fields and the associated field during starting of a two-pole induction motor. Figures 3.14 and 3.15 represent saturated stator and rotor leakage fields, respectively, and Fig",
+            " It is well known that during starting saturation occurs only in the stator and rotor teeth and this is the reason why Figs. 3.13 and 3.16 are similar. 220 Power Quality in Power Systems and Electrical Machines F2 = 3500 N/m F1 = 3500 N/m F 2 = 3 11 0 N /m F 1 = 3 15 0 N /m F 2 = 1250 N/m F1 = 4340 N/m F 2 = 1770 N /m F 1 = 5160 N /m F 2 = 1400 N /m F 1 = 4110 N /m F 2 = 1 30 0 N/m F 1 = 5 45 0 N/m 1 \u2013B \u2013B 2 3 A A 4 5 \u2013C \u2013C \u2013A \u2013A C C B B 6 7 8 9 10 11 12 (a) 1 \u2013B \u2013B 2 3 A A 4 5 \u2013C \u2013C \u2013A \u2013A C C B B 6 7 8 9 10 N S 11 12 (b) f f 1 \u2013B \u2013B 2 3 A A 4 5 \u2013C \u2013C \u2013A \u2013A C C B B 6 7 8 9 10 N S 11 12 (c) Figure 3.10 Flux distribution and radial stator core forces at no load and rated voltage for (a) rotor position #1, (b) rotor position #2, (c) rotor position #3, Continued 221Modeling and Analysis of Induction Machines 222 Power Quality in Power Systems and Electrical Machines 223Modeling and Analysis of Induction Machines Figure 3.16 Field distribution (second approximation) during starting with rated voltage of a 3.4 MW, two-pole, three-phase inductionmotor. One flux tube contains a fluxper unit lengthof 0"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0001466_978-3-319-17518-8_6-Figure1-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0001466_978-3-319-17518-8_6-Figure1-1.png",
+        "caption": "Fig. 1 Block diagram for the MMAE scheme",
+        "texts": [
+            " The performance of the proposed FTC is validated in two different situations, whose result is given in Sect. 5. Sect. 6 concludes the paper. This section first introduces the model of the quadrotor, which includes the translational and rotational dynamics. The assumptions which are required are also presented. Then, the modeling of actuator and actuator faults are introduced. Finally, the measurement model is given. Define an earth frame {\u03a3e}(Oe,xe,ye,ze) and {\u03a3b}(Ob,xb,yb,zb) in which Ob is fixed to the quadrotor (see Fig. 1). The earth frame is the North East Down (NED) frame in which the ze axis points down. The zb axis of the body frame also points down. The rotation of the body frame with respect to the earth frame is denoted by the following rotation matrix R: R = \u23a1 \u23a3 cos\u03b8 cos\u03c8 sin \u03c6 sin\u03b8 cos\u03c8 \u2212 cos\u03c6 sin \u03c8 cos\u03c6 sin\u03b8 cos\u03c8 + sin\u03c6 sin\u03c8 cos\u03b8 sin\u03c8 sin \u03c6 sin\u03b8 sin \u03c8 + cos\u03c6 cos\u03c8 cos\u03c6 sin\u03b8 sin\u03c8 \u2212 sin\u03c6 cos\u03c8 \u2212sin\u03b8 sin\u03c6 cos\u03b8 cos\u03c6 cos\u03b8 \u23a4 \u23a6 (1) where \u03c6 , \u03b8 and \u03c8 will be defined later. Before introducing the equations of motion of the quadrotor, the following assumptions have to be made [2]: 1"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0001017_cdc.2016.7798601-Figure3-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0001017_cdc.2016.7798601-Figure3-1.png",
+        "caption": "Fig. 3. Adapted nonuniform computational grid",
+        "texts": [
+            " According to this algorithm, the computational grid is stretched in the area of interest (which coincides with the center of the flying formation) and is coarser away from it (the grid nodes away from the center of the flying formation are distributed according to the parabolic law). The total number of grid nodes remains constant during the whole estimation process. The suggested grid adaptation algorithm results in the dynamically adapted computational grid, with the stretched region moving together with the leader-UAV. A snapshot of the computational grid is shown in Figure 3. The new grid is generated every time the leader passes a distance equal to the minimum grid size. The trilinear interpolation method is used to interpolate the solution obtained at the previous time iteration. The approach is tested using aerodynamic characteristics of Aerosonde UAV [13] under the following atmospheric conditions: U = 9 m/s, V = W = 0, KXX = KY Y = KZZ = 20 m2/s. The domain under consideration has dimensions \u2126 = 4km \u00d7 4km \u00d7 1km. The source is stationary, elevated, located at \u0398c = (0"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0001197_978-3-658-12701-5-Figure6.5-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0001197_978-3-658-12701-5-Figure6.5-1.png",
+        "caption": "Figure 6.5: Planar robot \u2014 Inverse kinematics \ud835\udc5er = \ud835\udc5e3, virtual link",
+        "texts": [
+            " From the transformation of the cartesian representation of the end-effector coordinates to polar coordinates range \ud835\udc59 and bearing \ud835\udefc, \ud835\udc59 = \u221a\ud835\udc652 E + \ud835\udc662 E \ud835\udefc = arctan (\ud835\udc66E, \ud835\udc65E) , an expression for the auxiliary angle \ud835\udf05 can be found, \ud835\udf05 = \u00b1 arccos (\ud835\udc592 + \ud835\udc4e2 \u2212 \ud835\udc5923 2 \ud835\udc59 \ud835\udc4e ) , considering that the cosine is an even function. From a comparison of Figure 6.3 and Figure 6.4 on the next page, the joint angle \ud835\udc5e1 can be determined, \ud835\udc5e1 = \ud835\udefc + \ud835\udf05 \u2212 \ud835\udefd. In Figure 6.3 it can be seen that the remaining joint angle \ud835\udc5e3 is \ud835\udc5e3 = arctan (\ud835\udc66E \u2212 \ud835\udc591 sin (\ud835\udc5e1) \u2212 \ud835\udc592 sin (\ud835\udc5e1 + \ud835\udc5e2) , \ud835\udc65E \u2212 \ud835\udc591 cos (\ud835\udc5e1) \u2212 \ud835\udc592 cos (\ud835\udc5e1 + \ud835\udc5e2))\u2212\ud835\udc5e1\u2212\ud835\udc5e2. Due to the assumption that \ud835\udc5e3 is known, the adjoined links 2 and 3 can be reduced to one single, virtual link of length \ud835\udc4e, \ud835\udc4e = \u221a(\ud835\udc592 + \ud835\udc593 cos (\ud835\udc5e3))2 + (\ud835\udc593 sin (\ud835\udc5e3))2, (6.3) as depicted in Figure 6.5. After transforming the cartesian representation of the endeffector position to the polar form, range \ud835\udc59 and bearing \ud835\udefc, \ud835\udc59 = \u221a\ud835\udc652 E + \ud835\udc662 E (6.4) \ud835\udefc = arctan (\ud835\udc66E, \ud835\udc65E) , (6.5) and applying the law of cosines to the triangle with edges \ud835\udc4e, \ud835\udc59 and \ud835\udc591 from Figure 6.6, an expression for the auxiliary angle \ud835\udefd can be found, \ud835\udefd = \u00b1 arccos (\ud835\udc592 + \ud835\udc5921 \u2212 \ud835\udc4e2 2 \ud835\udc59 \ud835\udc591 ) , considering that the cosine is an even function. Now the joint angle \ud835\udc5e1 can be computed, \ud835\udc5e1 = \ud835\udefc \u2212 \ud835\udefd. With the auxiliary angles \ud835\udefe and \ud835\udf05 from Figure 6"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0001051_phm.2016.7819861-Figure1-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0001051_phm.2016.7819861-Figure1-1.png",
+        "caption": "Figure 1. Crack model at gear tooth root [4].",
+        "texts": [
+            " Section II briefly introduced the gear model with tooth crack in [4], and then a modified model is proposed, in which the propagation along both tooth width and tooth depth are considered. Based on this model, a rigid-flex coupling dynamic model is constructed in Section III and the corresponding analyses of dynamic response are made. Finally, the conclusion is drawn in Section IV. II. MODELING OF SUN GEAR WITH TOOTH CRACK For the tooth crack, it is not always throughout the tooth width but grows along the tooth width gradually in the AM model, in which a parabola is suitable to approximately simulate the path along the tooth width [4], as showed in Fig. 1. In this model, the crack propagation path along the tooth depth was simplified as a straight line. In most of resent works, the crack in FE method is also set as a straight line along the crack depth and throughout the tooth width. However, this is a simplified version of the real crack. Experiment and simulation results [6, 7] demonstrated the crack propagation path along the tooth depth can be approximated by a parabola, as shown in Fig. 2. Based on the results in [6, 7], in this paper, the tooth crack in the gear model would be simulated along both the tooth width and the crack depth, each of which is a parabola"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0001081_icarcv.2016.7838776-Figure16-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0001081_icarcv.2016.7838776-Figure16-1.png",
+        "caption": "Figure 16. RS key-to-key movement method",
+        "texts": [
+            " The corresponding RS vibration was again observed by a HSVC. The simulated values given by (5) and (6) are shown in Fig. 15. The differences between the theoretical and experimental values are less than 5 [mm] at all amplitudes. Thus, this model is thought to be appropriate. VII. RUBBER-STICK VIBRATION SUPPRESSION USING INPUT SHAPING DURING MOVEMENT TO TARGET POSITION We examined a movement method for controlling the unnecessary vibration during key-to-key movement. The conventional method (Pattern 1 in Fig. 16) reaches the target position in one movement. In contrast, Pattern 2 divides the movement into two components. The RS in the robot\u2019s right hand was considered (Fig. 12) and key C7 was taken as the initial position. The robot moved in the y-axis direction to key B6 and the stick tip vibration was observed using a HSVC. The robot moved at 100% motion velocity. For Pattern 2, (Y1, Y2) = (15, 15), (20, 10), and (25, 5) (Y1 and Y2 are defined in Fig. 16). The robot moved the RS from key C7 to B6 using each movement method. Fig. 17 shows the absolute value of its vibration in the y-axis direction for each case. The vibrations are more obviously reduced for the Pattern-2 movement method. However, the degree of vibration restraint varies with (Y1, Y2). We then considered (Y1, Y2) = (20, 10), which restrains the vibration most significantly. Figure 17. Vibrations for each movement method -15 -10 -5 0 5 10 15 20 25 0 0.2 0.4 0.6 0.8 1 A m pl itu de [m m ] Time [s] Input power Y Input power Y Simulation result t t2 Figure 18"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0003941_a:1013730210328-Figure2-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0003941_a:1013730210328-Figure2-1.png",
+        "caption": "Fig. 2. The surface-pressure mean molecular area isotherms for ( 1 ) CH 3 (CH 2 ) 17 SH , ( 2 ) CH 3 (CH 2 ) 17 OH , and ( 3 ) HO(CH 2 ) 22 SH on pure water subphase (from [39]).",
+        "texts": [
+            " They are however known to form most stable monolayer coatings on gold and other metal substrates. Therefore, the self-assembly approach involving simple adsorption of the thiol from the bulk solution onto the electrode is a preferred approach. However, if mixed monolayers are to be prepared, this approach does not provide good control of the composition of the monolayer formed on the solid substrate since the thiol exhibits preferential adsorption compared to the other components of the solution. In such a case we proposed to modify the thiol molecules with another terminal OH group (Fig. 2). This way the OH group is responsible for the high stability of the monolayer on the air/water interface, while the thiol terminal group is responsible for stable attachment of the monolayer to the gold surface [39]. RUSSIAN JOURNAL OF ELECTROCHEMISTRY Vol. 38 No. 1 2002 ELECTRON TRANSPORT THROUGH COMPOSITE MONOLAYERS 31 Using an approach similar to Schaefer\u2019s [27] horizontal touching the monolayer can be efficiently transferred to the electrode surface. In both Langmuir\u2013Blodgett and the self-assembly approaches the driving force for the formation of a stable monolayer is the formation of thiolate-gold bonds and favorable lateral interactions among the aliphatic chains of neighboring alkanethiol molecules"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0001560_36.3.302-Figure1-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0001560_36.3.302-Figure1-1.png",
+        "caption": "Fig. 1. Mechanism of production of apical linear shadows. (Roentgen rays projected tan gentially to surface prominences produce sharply defined linear markings.)",
+        "texts": [
+            " An apical structure produces a visible linear shadow on the film only (1) when placed in juxtaposition to another struc ture of strikingly lessened radiographic density, and (2) when so oriented that it presents a distinct surface or marginal prominence, which the roentgen rays may strike tangentially. Practically speaking, apical structures produce visible linear shadows only when seen as distinct pro jections either at the endothoracic or skin surface in contrasting relationship with air or air-filled lung. Roentgen rays projected tangentially to the surfaces of such prominences produce sharply defined linear contours (Fig. 1). Structures which produce no surface saliences add to the general density but cannot be seen on the films as separate and distinct entities. The Supraclavicular Border Shadow. This is a linear shadow two to five milli meters wide in relation to the upper border of the clavicle, representing skin together with a thin layer of subcutaneous connec tive tissue as it crosses the prominence of the clavicle. It is ordinarily visible as far medially as the clavicular attachment of the sterno-mastoid muscle",
+            " Zawadowski offers the following ex planation for this shadow: \"The first rib takes a very oblique course such that its internal border in its upper por tion is at the same level as its external border. This results in the medial border of the first rib becoming countersunk into the lateral surface of the lung apex, the latter rising above it. The boundary between air-filled parenchyma and solid tissue describes at this point a little curve (courbe) which touched tangentially by the rays produces a linear shadow having the density of soft tissues and paralleling the medial border of the rib.\" (See Fig. 1.) Knuttsen (quoted by Pendergrass) and Andrus, on the other hand, attribute this shadow to direct visualization of a con nective-tissue fascial band corresponding to the fascia of Sibson (Sebileau) pre viously described. The Border Shadow of the Second Rib. This shadow is narrower (1 to 3 mm.) than the border shadow of the first rib but considerably longer. It is usually bilaterally symmetrical. The clean-cut lower margin of the shadow forms an shadow to represent the upper limit of the lung. It has been shown, however, that the shadow persists unchanged after pneumothorax"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0001900_156855395x00076-Figure7-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0001900_156855395x00076-Figure7-1.png",
+        "caption": "Figure 7. Coordinate systems.",
+        "texts": [
+            " Inclination control of the upper surface is indispensable in the case of conveying a load. Here we study the inclination control method. This omni-directional walking robot has two frames and they are assumed to be in the same plane. Walking is performed by sliding frames against each other. To determine the position and inclination of the upper surface in the three-dimensional space, we determine the coordinates of the three points on the upper surface. In order to calculate the posture and the inclination mathematically, we define the coordinate systems as shown in Fig. 7. Five coordinate systems {U}, {A}, {B}, {C} and {F} are defined. {U} is the world coordinate system located on the earth and {F} is the base coordinate system fixed on the upper surface. P is the parallel moving vector. {A} is the coordinate system generated by moving the coordinate system {U} by {P}. {B} is the coordinate system generated by revolving the coordinate system {A} about the Za-axis. {C} is the coordinate system generated by revolving the coordinate system {B} about Yb-axis. {F} is the coordinate system generated by revolving the coordinate system {C} about X,-axis. A capital letter means a vector and a small letter means a component of the coordinate system. The large letter located at the left side of a capital letter means in which coordinates the vector is expressed. Trigonometry functions (sin 0, cos 0, tan 0) are sometime expressed as (sO, cO, tO). To express the coordinate transformation from {U} to {F}, we use Euler angles (a, 0, y). The relation between {Xu, Yu, Zu} and [Xf, Yy, Z f} is obtained by ZY-X coordinate transformation, as shown in Fig. 7 and equations (4) and (5). Z-X-Y coordinate transformation is written in [18]. Let us suppose that the angle a is 0, {Xa} and {X y} are in the same vertical plane and the relation between {A} and {F} is expressed in equations (6) and (7). We control the D ow nl oa de d by [ U ni ve rs ity L ib ra ry U tr ec ht ] at 0 2: 10 1 7 M ar ch 2 01 5 184 inclination of the upper surface using {A} and {F}, and the coordinate transformation is written as follows: 4.2. Relation between Euler angles and legs This robot moves by sliding the two frames and supporting the body by each frame, so that inclination control is necessary for each frame"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0002370_70.795796-Figure7-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0002370_70.795796-Figure7-1.png",
+        "caption": "Fig. 7. Three-fingered colinear grasp on a concave object: (a) the grasped object with three contacts, (b) internal force triangle, and (c) decomposition of internal force into parallel and interaction forces.",
+        "texts": [
+            " Some researchers [1], [3], [7], [30] have recognized that the internal force shows compression or tension. As we see, the parallel force at each contact is parallel to the normal of the grasp plane (for coplanar grasps), or perpendicular to the common line of the contact points (for colinear grasps). It shows neither compression nor tension. However, the parallel force does have important effect on 2-D grasps, as will be shown later. In this section, two examples are presented to illustrate the effects relative to the equivalence problem of internal and interaction forces. Fig. 7(a) shows a three-fingered grasp on a concave object. The contact points are colinear and the inward contact normals intersect at point O. = 20 . The coefficient of friction between the fingertip and object surface = 0.2. Now we consider the effects of interaction and internal forces on the closure property of grasps [33]\u2013[35]. Since the three contact points are colinear, the interaction force between any two contact points is along line C1C3. From the given condition, we obtain tg 1 = 11:30 < , which implies that there exists no interaction force situated inside the friction cone at each contact. If we regard the interaction force as the equivalent characterization of the internal force, we may conclude that the grasp is not prehensile [35]. However, there exists such an internal force, whose threedimensional component at each contact is along the inward normal, as shown in Fig. 7(b). Apparently, this internal force satisfies the friction constraints. So, the grasp is actually prehensile [35]. The former wrong conclusion results from the ignorance of the equivalent conditions for internal and interaction forces. Since the three contact points are colinear, the ranks of the internal and interaction force matrices are: rank(W ) = 4, rank(W a) = 2. Namely, rank(W a) < rank(W ). Therefore, the interaction force space is smaller than the internal force space and the interaction force can not be regarded as equivalent to the internal force. As a matter of fact, the internal force space in this case is the direct sum of the interaction and parallel force spaces. The internal force in Fig. 7(b) can be uniquely decomposed into a parallel force and an interaction force, as shown in Fig. 7(c). The quantitative relation is given by fh1 fh2 fh3 = e13 0 e31 + f b1 f b2 f b3 where f bi is the parallel force at Ci, and is a scale factor denoting the magnitude of the interaction force. As we see, the interaction between C1 and C3 shows compression, and the parallel force at each contact shows neither compression nor tension due to its perpendicular property to the common line of contacts. However, the parallel force does have effect on grasp feasibility, as shown in the above analysis. Also note that the resultant of interaction force and parallel force (i"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0002614_s0094-114x(97)00067-0-Figure2-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0002614_s0094-114x(97)00067-0-Figure2-1.png",
+        "caption": "Fig. 2. Diagram of a prosthetic arm.",
+        "texts": [
+            "00PII: S0094-114X(97)00067-0 895 obtain the explicit expressions of its acceleration using other methods. In this paper, the \u00aerstand second-order in\u00afuence coe cient matrices of the arm are established and the explicit expressions of velocity and acceleration are derived. The in\u00afuence coe cient matrix is only the function of dimensional and positional parameters and the formulas of velocity and acceleration have the same form as those of a series manipulator. POSITION ANALYSIS The diagram of the prosthetic arm is shown in Fig. 2. The upper arm consists of a planar o - set guide-bar mechanism and a spatial four-link mechanism. Its forearm consists of a threeDOF series mechanism. The prosthetic arm is connected to the reference frame with three ``Hooke Joints'' at points O, A and F. Points B and D are spherical joints. The \u00aeve inputs are s1, s2, y1, y2, y3. The origin of the \u00aexed coordinate system O-xyz is located at the center of ``Hooke Joint'' O. The x-axis is along OA (from point O to point A). The y-axis is at the plane determined by points O, A and F",
+            " The branches AB and DF can be considered to consist of six basic pairs with oneDOF by means of decomposition of kinematic pair [11]. The expression between the generalized velocity of every pair and the velocity of point P on link OC can be obtained from the velocity relation of series manipulator as in Equation (25). Since there are only two basic pairs with oneDOF in branch OC, the Equation (25) can not be used. But we can add four hypothetic links with basic pair in branch OC (refer to part of dotted line in Fig. 2), so that it becomes a branch with six basic pairs [7]. Then the Equation (25) can be used. Note that the added hypothetic links and basic pairs are needed to ensure that in\u00afuence coe cient matrix used later is nonsingular. For every branch of the parallel mechanism, from Equation (25) we have: _f r GP f \u00ff1 r VP r 1; 2; 3 ; 26 where _f r f _f1 _f2 _f3 _f4 _f5 _f6g r is the generalized velocity of r-th branch, [GP f] \u00ff1(r) is the in\u00afuence coe cient matrix of r-th branch, VP is the velocity of point P on link OC",
+            " Since the acceleration of the central point oh of the hand can be indicated by the following equation [4, 5]: AH GH y y _y T HH y y _y ; 67 substituting Equations (37) and (65) into Equation (67), the acceleration of the hand can be expressed as: AH GH y Gy y y GH y _y T Hy yy _y _y T Gy y T HH y y Gy y _y: 68 The above equation can be written in more compact form as follows: AH GH y y _y T HH yy _y; 69 where HH yy ~G ~H Gy y T HH y y Gy y ; ~G ~H ij f GH y Hy yy ijg: 70 If the acceleration of the hand is given, the input acceleration can be obtained using the follow- ing equation y GH y \u00ff1fAH \u00ff _y T HH yy _yg: 71 Fig. 4(a)\u00d0Caption on p. 907 Equations (69) and (71) are the forward and inverse acceleration formulas of the prosthetic arm. They have also the same form as those of a series manipulator. NUMERICAL EXAMPLE In this section, we will calculate the position, velocity and acceleration of the prosthetic arm shown in Fig. 2 using the formulas derived in the paper. The kinematic parameters of the arm are given in Table 1. For the parallel mechanism, by the decomposition of kinematic pair, the unit vectors of the kinematic pairs on every branch of the parallel mechanism are given in Table 2. On branch OC, the orientations of hypothetic joint axes s1, s2, s3 and s4 may be de\u00aened Fig. 4(b)\u00d0Caption on opposite page arbitrarily, provided that the \u00aerst-order in\u00afuence coe cient matrix of the branch is nonsingular. In order to verify the calculated result, we give a special example"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0000606_978-3-319-44087-3_43-Figure3-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0000606_978-3-319-44087-3_43-Figure3-1.png",
+        "caption": "Fig. 3 Cross-section of the mechanical system of the needle bar a Holding position, b Release position",
+        "texts": [
+            " Kom\u00e1rek (&) Technical University of Liberec, Liberec, Czech Republic e-mail: jiri.komarek@tul.cz \u00a9 Springer International Publishing Switzerland 2017 J. Beran et al. (eds.), Advances in Mechanism Design II, Mechanisms and Machine Science 44, DOI 10.1007/978-3-319-44087-3_43 323 article enables the analysing of the mechanical system of the needle bar and it will help with its optimization. A detailed description of the sewing machine is included in the work [1]. The cross-section of the mechanical system of the needle bar is shown in Fig. 3. The floating needle 20 is held by collets 12 inside the mechanical system of the needle bar. The collets are made up of two balls whose axial movement is controlled by a cylinder 4. The balls are pushed into a conical hole in a shell 2 due to the pressing force of the springs 11, 15. This situation is shown in Fig. 3a. The release of the needle is started in a moment when the control element (assembled from parts 1, 3, 6, 10, 11, 16, 17) collides with the machine frame 18. The impact is absorbed by the rubber pad 10. The influence of stiffness of the rubber pad on absorbing of the impact is described in [2]. After the impact, the shell 2 continues its movement towards the bottom dead centre position, whereby the balls are released in the enlarged area of the conical hole. This situation is shown in Fig. 3b. A more detailed description of the mechanical system of the needle bar is in [1]. There has been carried out an experimental measuring of compression of springs in response to applied force for the purpose of determining the stiffness of the springs. There has been obtained values of stiffness of the spring 11 k2 = 900 N/m and the spring 15 k3 = 690 N/m. There has been also experimentally determined dependence of deformation of the rubber pad in response to the applied force. The dependence is nonlinear and its obtained values has been approximated by a polynomial of third degree (1), where x20 denotes the deformation of the rubber pad and F20 the applied force"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0001522_pi-c.1955.0020-Figure2-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0001522_pi-c.1955.0020-Figure2-1.png",
+        "caption": "Fig. 2.\u2014Primitive machine with stationary axes.",
+        "texts": [
+            " Quantities which are not tensors may arise in one system of reference and become zero in another; in other words, they may arise because of the particular reference frame chosen. This effect occurs especially when there is relative motion between the reference axes. Kron takes a primitive or elementary representative network which has comparatively simple equations; these are written in tensor form and transformed to give the equations of the required system. A wide range of machines can be considered as a group of interconnections of the windings of the primitive machine shown in Fig. 2, and the equations of any of these may be obtained by transformation of the primitive-machine equations. The expression \"reference frame\" denotes the system of measurement from which the variables and parameters are determined. In a static electrical network a change of components from branch to mesh currents is a change of reference frame.2 In the case of a slip-ring induction motor, stator currents will be measured from stationary terminals and rotor currents from axes rotating with the rotor",
+            "sin0 6' \u2014Md sin 0 (LQr \u2014 Ldr) sin 0 cos 0 Lrfr sin2 0 + Lqr cos2 0 M^ cos 0 M 9 s i n 0 M? cos 0 (9) where Lss is the moment of inertia of the rotor. 152 LYNN: THE TENSOR EQUATIONS OF ELECTRICAL MACHINES + [ab,s]iaib (18) to cover only the mechanical part of the range of variables,1 d1xt thus fs = Rsti* + Ls,-fiT where [ob,s] = \u2014 \u2022= Therefore /\u201e = Rtt-r + L,,-^ 2 Id (19) ddnd .. . , Rst-f- \u2014 fnctional torque r d 2 B \u2022 \u2022\u2022 \u2022 *'T2 = i n e r t i a torque and - -^ \u2014 electrical torque The second form of the primitive machine considered here is shown in Fig. 2. The rotating axes a' and b' of Fig. 1 are here Fig. 3.\u2014Primitive machine with axes rotating freely. In synchronous-machine studies the quadrature-axis stator (field) coils are omitted unless amortisseur windings are being considered. Here sinusoidal flux distribution is considered and M'd = Md, etc. Eqns. (20) may be obtained from those of the previous form of primitive machine using the relationship ,-\u00ab' \u2014 (dr c o s Q _|_ [gr s m Q (22) resolved along the direct and quadrature axes. All axes are now relatively stationary",
+            " The following Sections show that the connection Yuvw arises naturally because of the dynamical relationship between the two types of primitive machine, this being quasi-holonomic and non-integrable. Fig. 4 shows the form of the connection Yuvw when written as a matrix in the form of a cube, together with the arrangement of matrix multiplication leading from eqns. (26) and (27) to eqns. (23) and (24). (3) NON-HOLONOMIC TRANSFORMATIONS The currents in the armature axes a' and b ' in Fig. 1 may be resolved along \"d\" and \"q\" axes shown in Fig. 2, the relationship being /<*\u2022 = /a' cos 9 \u2014 ib> sin 9 \"1 r ' (2.8) iv = p' sin 9 + ib> cos 9 j Since the variables in Lagrange's equations are the charges, x\u00b0, eqns. (28) represent a transformation of differentials of the variables, where dxa\\dt = ia, etc. These are equations of constraint among the differentials of the variables xk, and the transformation must therefore be written dxdr = dx?' cos 9 - dx5' sin 9 \"] . y . . . (29) dxv = dxa' sin 9 + dxb' cos 9 J They obtain at a given instant and cannot be integrated to give a relationship among the charges",
+            " The nature of the connection T^, w is examined by ' is \u2014Md sin 8 \u2014Md cos 6 \u2014Md sin 9 2{Lqr \u2014 Ldr) sin 9 cos 8 (Lqr-Ldr)(cos28-sm2d) Mq cos 8 \u2014Md cos 9 {Lqr-Ldr)(cos*8-sin28) -2(Lqr-Ldr)cos8sind \u2014Mq sin 8 Mq cos 9 -Mq sin 9 c\\fl ds ds a' b' qs The equation is therefore This may be written where c\\a ds ds a' ca b' qs b' qs (67) ec = Rcai\u00b0 + Lcapi\u00b0 + Vcai\u00b0p8 + Gcai\u00b0p8 a' b' (68a) (686) qs \u2014Md sin 9 (Lqr - Ldr) sin 9 cos 9 Lqr cos2 9 + Ldr sin2 9 Mq cos 9 \u2014Mdcos9 -Lqr sin2 9 - Ldr cos2 9 -{Lqr - Ldr) sin 9 cos 9 \u2014Mqs\\n9 (68c) <r\\a ds ds a' b' qs \u2014Md sin 8 \u2014Md cos 8 a' (Lqr - Ldr) sin 8 cos 8 -Ldr cos2 8 - Lqr sin2 8 b' Ldr sin2 8 + Lqr cos2 8 -(Lqr - Ldr) sin 8 cos 8 qs Mq cos 8 \u2014Mq sin 8 (6Sd) Considering the transformation to general rotating axes. The (5.1.2) The Stationary Axis Frame, (ii) objects rMUH, and Q^ have the following laws of transformation:2-5 From Fig. 2 it is seen that the transformation from frame (i) r.. \u2022 = r U V , W 1 C\" Cv C^ _i_ r r'w u uv, w^u'^v'^w \"\u0302 wu w'7\\xv' (63) to frame (ii) is given by where and ana 1 \u2022 w - t - - \u00ab u v dx\u00b0 = Cc.mdxm dr qr and are therefore not tensors.The quantities Fwt, w uv w The equations of the electrical machine with general rotating axes are derived in Section 5. (5) THREE REFERENCE FRAMES (5.1) Electrical Equations (5.1.1) The Holonomic Frame, (i) ec = Rcai\u00b0 + pLcai\u00b0 (65) e c = R c a i \u00b0 + L c a p i \u00ab + [ a b t f W ",
+            "1) The 3-Phase Induction Motor A 3-phase induction motor has been analysed by Stanley9 by expressing the equations in terms of an equivalent 2-phase machine. The 3-phase rotor and stator currents, voltages and flux linkages are resolved along two axes, namely the axis of the stator phase A and the common axis of the stator phases B and C in quadrature with phase A. Since both rotor and stator quantities are resolved along the same two axes these are relatively stationary and fixed on the stator as in Fig. 2. The equations for a balanced symmetrical motor with voltages applied to the stator may therefore be written stator appears to rotate synchronously backwards and the rotor to rotate backwards at the angular velocity of slip. The relationship among the currents in the stationary axes and those in the free frame may be written ids \u2014 ;S3 c o s Ql _ /S4 s i n idr \u2014 iS\\ c o s Q{ _ /S2 s m iv = (Si sin 0j + iS2 cos /<?* = iS3 sin 0, + i im = (143) cos 0, or (144) m\\y ds dr \u2022where C\u2122 = qr qs COS 0! sin 0j cos 0, sin0, \u2014sin 0j COS0, -sinfl, COS 0 "
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0003120_156855387x00057-Figure3-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0003120_156855387x00057-Figure3-1.png",
+        "caption": "Figure 3. Arm calibration camera calibration and hand-eye calibration.",
+        "texts": [
+            " The method used is based on the way the hand positions in camera coordinates which are measured by the vision system are related to the hand positions in arm coordinates which are calculated from the joint angles of the manipulator. The vision system is used not only to obtain the reference of manipulator motion but also to calibrate the parameters for coordinate transformations. When the view point of the vision system is changed, the parameters of (3) must be calibrated again. If the calibrations are done automatically by hand-eye coordination, the robot can adapt itself even if the relationship between the eye and arm is changed in operation [5]. The following describes the steps to obtain the transformation matrix. Figure 3 shows the coordinate systems in the hand-eye system. W, A, H and C stand for world coordinates, arm coordinates, hand coordinates and camera coordinates, respectively. The hand position H is described in two ways; hc in camera coordinates (C) and hA in arm coordinates (A). Let the transformation matrix be Tc 1. A , where D ow nl oa de d by [ H er io t- W at t U ni ve rs ity ] at 0 3: 23 0 7 Ja nu ar y 20 15 To perform the calibration, the hand moves a small ball to some suitable position and then the three-dimensional position of the ball is measured by binocular stereo vision"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0002999_peds.1997.627419-Figure1-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0002999_peds.1997.627419-Figure1-1.png",
+        "caption": "Fig. 1 Vector Relationship for PMSM",
+        "texts": [
+            " But owing to the construction property of PMSM, its vector control drive is only equivalent to a DC motor with no-compensation winding, its transient performance is affected, thus some other control strategies are also de~e loped[~~- [~ ] . Otherwise, in the most papers, whether the maximum torque control for PMSM exists and its maximum torque control is the same concept with vector control has not been proposed and analyze. In this paper, the various vector control strategies and maximum torque control for PMSM drives are developed and analyzed under different d and q axes parameters, the emphases on maximum torque and relationships between vector control and maximum torque control are described. Fig. 1 shows vector relationship for PMSM. Each phase torque of PMSM is given by (1) 1 =-E,I,+(L,- L,)I,I, U where Ld , Lq -------- d , q axis inductances; w -------- synchronous angular frequency; I I d , q -------- stator d and q axes currents; E , _ _ _ _ _ _ _ _ e.m.f produced by the rotor flux. Analysis are carried out in the two categories described in the following paragraphs. 7803-3773-5/97/$10.00 0 1997 IEEE Case I -for non-salient rotor PMSM, the magnetic palhs are both equal in the d and q axes , and Ld = Lq The torque is obtained by equation (1) 1 T, =-E,Iq w When the rotor pole axis is selected as flux orientation direction of vector control , PMSM drive is equivalent to a DC motor drive by controlling the stator q axis current I, only"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0002738_bf02455425-Figure2-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0002738_bf02455425-Figure2-1.png",
+        "caption": "Fig. 2. - Critical reduced magnetic coherence length ~c as a function of the reduced thickness ~ of the cell in the hypothesis of strong anchoring at the external wall. For a cell with r0 = 1 mm and d = 20 ~m, if the elastic ratio \u2022 assumes the values ( - 2, 0, 2), and the twist-extrapolation length is ~ 10 ~m, then )~ = ( - 95, - 99, - 103). We point out that the expected Zc-shift corresponding to a variation 5\u2022 = ___ 2 is of the order of ___ 1%. The curve )~ = - oo refers to the case of twist-anchoring strong on both walls.",
+        "texts": [
+            " giv ing the critical value of the reduced cohe ren t l eng th 3r for the ar is ing of in te rna l wall, the b o u n d a r y condit ions t u r n out to be (12) Oo -- O, 70 - - O, (Oz)l -}- ~,Ol O1 -- O, (Yz) J[- )~71 Y1 -- O, be ing ~01 - 1 + 2x + l~ , 2rl -- 1 + l~ , w h e r e lol =-K2e/rlWol, and lrl = Kl l / r lwr l are the r educed ex t rapo la t ion length for tw i s t and splay, respec t ive ly , at the ex t e rna l wall. N o w the crit ical equa t ions wr i t e (13) ~ = Zc {7: - a c tg (~01Sc)}, ~c = exp [ - t gh ~1] - 1. I n fig. 2 the behav iou r of 3r vs. ~ is r e p o r t e d for x = - 2 , 0 , 2 ( the c o r r e s p o n d i n g values of ~0o a re - 95, - 99, - 103) and loo = 10 .2 (if r0 ~ 1 m m the extrapolation length will be - 1 0 ~m, value easily obtained with common NLC and by convenient chemical treatment of the surface (1~)). The curve with ~0o = - ~ is related to strong-anchoring condition at both walls (14). Notice t h a t for given 8 the critical value -~c is a function of ~00 as expected, allowing us to determine K24 by a direct measurement of H~, if r0, the anchoring strength W~o and the twist elastic constant Kz2 are known"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0001046_icmic.2016.7804166-Figure1-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0001046_icmic.2016.7804166-Figure1-1.png",
+        "caption": "Fig. 1. Geometrical representation of the WMR.",
+        "texts": [
+            " PROBLEM FORMULATION The mission assigned to this robot is to move from a start point Pi of configuration qi to an end point Pf of configuration qf while respecting a set of constraints and achieving specific goals. This section is devoted to the establishment of the mathematical formulation of the considered problem. 978-0-9567157-6-0 \u00a9 IEEE 2016 1) Robot Description: The considered WMR consists of a platform equiped with two independent driven wheels. The operational configuration of the system is given in the frame (O, X, Y, Z) by:   T x, y, \u03b8q  where (x, y) represent the robot center position and \u03b8 the robot orientation around the Z-axis (Fig 1). 2) Kinematics:The kinematic model of the robot is given by [6]:  q S q v  with         T cos \u03b8 0 0 v\u00a0\u00a0\u03c9 \u00a0\u00a0,\u00a0\u00a0\u00a0 sin \u03b8 0 0 0 0 1            v S q  v and \u03c9 represent respectively the linear velocity and the angular velocity of the robot. Relations (2) and (3) are equivalent to:    \u03b8 \u03b8x v\u00a0cos , \u00a0\u00a0y v\u00a0sin , \u03b8 \u03c9    The assumption of no slipping along lateral direction of the wheels leads to the non-holonomic constraints given by:   0\u00a0 C q q   where      sin \u03b8 \u2013cos \u03b8 0   C q  From (3) and (4), we find the expression of key kinematic parameters v, \u03b8, \u03b8 and \u03b8 : 2 2 v x y  () y x \u03b8 arctg         2 2 xx+yy v x +y   2 2 xy yx \u03b8 \u00a0\u00a0\u00a0 x y             2 2 2 2 2 x +y x y-yx -2 xx+yy xy+yx \u03b8 x +y   3) Dynamics: The dynamic model of the WMR, satisfying the kinematic constraints (5), can be obtained by using the Lagrange\u2019s equations, i"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0000088_j.ifacol.2015.06.369-Figure4-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0000088_j.ifacol.2015.06.369-Figure4-1.png",
+        "caption": "Figure 4 The 6-SRS parallel robot",
+        "texts": [
+            " And the optimization result is surrounded by non-degenerate singularity surfaces. The relationship between the iteration and the planning results is presented by Figure 3 also, in which branch optimum method is shown clearly to deduce the maximum of max minl l . The iteration step numbers in Figure 3 are not namely equal to the computing time as the time loss of each step isn\u2019t presented here. The mechanical structure and coordinates for all the terminal points of 6-SPS parallel robot are shown in Figure 4 and Table 2.Assume that 0 (0,0,3.5,0,0,0)P  . The boundary of workspace for actuators in M space is 120 five-dimensional hyper surfaces, and there are 64 terminal points. As a regular polyhedron, the lines between terminal points sketch the outline of the actuator workspace. The constraint conditions for the optimization are given as follows: a) Singularity-free condition det( ) 0J  ; b) Homeomorphic condition 2  and should satisfy equation (7) when 6n  . In optimization here, it\u2019s uneasy to present a similar traversal path as shown in Figure 2, and the path is simplified to improve the algorithm efficiency",
+            " The relationship between the iteration and the planning results is presented by Figure 3 also, in which branch optimum method is shown clearly to deduce the maximum of max minl l . The iteration step numbers in Figure 3 are not namely equal to the computing time as the time loss of each step isn\u2019t presented here. Figure 2. Direct view of actuators\u2019 planning Figure 3. Iteration steps against actuators\u2019 length of P0 The mechanical structure and coordinates for all the terminal points of 6-SPS parallel robot are shown in Figure 4 and Table 2.Assume that 0 (0,0,3.5,0,0,0)P  . Table 2 The coordinates for 6-SPS parallel robot Actuators Base-terminals End-effector terminals 1l 1A (-1.9544,0.0481,3) 1B (-1.4269,2.3695,0) 2l 2A (0.9355,1.7166,3) 2B (-1.3386,2.4205,0) 3l 3A (1.0189,1.6685,3) 3B ( 2.7655, 0.051, 0) 4l 4A (1.0189,-1.6685,3) 4B ( 2.7655, -0.051,0) 5l 5A (0.9355,-1.7166,3) 5B (-1.3386,-2.4205,0) 6l 6A (-1.9544,-0.0481,3) 6B (-1.4269,-2.3695,0) The boundary of workspace for actuators in M space is 120 five-dimensional hyper surfaces, and there are 64 terminal points",
+            " In optimization here, it\u2019s uneasy to present a similar traversal path as shown in Figure 2, and the path is simplified to improve the algorithm efficiency. Denote the minimum actuator length to be\u201d0\u201d and maximum length to be\u201d1\u201d, there are two methods to define the traversal order. One ordered by Grey code, which holds the traversal path stands in the surface of a hypercube; the other ordered by B(2,6) De Bruijn sequence, which has the shortest path through all terminals. In this work, we use the Grey sequence in the optimization algorithm. y z y z O O 1A 2A 3A 4A 5A 6A 2B 1B 3B 4B 5B 6B 1L 2L 3L 4L 5L 6L Figure 4 The 6-SRS parallel robot Notice that we also use l instead of u here. The optimization process is shown in Figure 5, and we have the results max 4.717l  and min 2.989l  . INCOM 2015 May 11-13, 2015. Ottawa, Canada The selection of simply connected motion domain in central symmetric parallel robots is easier than in other structure of parallel robot. Once proper motion domain is determined, the constraint conditions can be used in the optimization of singularity-free workspace with different objective function in various structure of parallel robots"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0001235_b978-0-08-099425-3.00005-4-Figure5.5-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0001235_b978-0-08-099425-3.00005-4-Figure5.5-1.png",
+        "caption": "FIGURE 5.5",
+        "texts": [
+            " The effective rolling radius, Re, is the ratio of the linear velocity of the wheel centre in the XSAE direction to the angular velocity of the wheel. While such considerations are easily imagined when the tyre is upright on a smooth surface with a constant load, it may be readily appreciated that discerning the rolling radius for a real tyre with a nonzero inclination angle and on a nonsmooth surface is constantly varying and somewhat difficult to discern in anything resembling real time using the above notions. A more detailed treatment of effective rolling radius is provided by Phillips (2000) based on the representation given in Figure 5.5 that allows the effective rolling radius to be related to the unloaded radius and tyre deflection. For the tyre shown in Figure 5.5, the wheel axle is considered fixed and the road moving such that the relative forward velocity of the wheel is V. If a number of equidistant radial lines are drawn on the tyre the number passing point A in a given time must be the same as the number passing point P in the contact patch. Deformation of rolling tyre. If we take d \u00bc the distance between the radial lines at the tyre outer radius near A d0 \u00bc the distance between the radial lines at the contact patch near P then d' V d R\u03c9 u = \u00f05:2\u00de therefore d d'R \u03c9 VR ue == \u00f05:3\u00de The tread band is subject to a longitudinal compressive strain within the contact patch \u03b5 where d d'd\u03b5 \u2212= \u00f05:4\u00de \u03b5)(1dd' \u2212= \u00f05:5\u00de therefore \u03b5)(1RR ue \u2212= \u00f05:6\u00de Assuming that the strain in the contact line is constant we have (assuming ....... 5! \u03b8 3! \u03b8\u03b8sin\u03b8 53 +\u2212= ) 6 \u03b81 \u03b8 \u03b8sin arcBC cordBC\u03b51 2 \u2212\u2248==\u2212 \u00f05:7\u00de From Figure 5.5 we also have (assuming .......\u03b8\u03b81\u03b8cos 42 +\u2212= ) 4!2! ( ) 2 \u03b8R\u03b8cos1R\u03b4 2 uuz \u2248\u2212= \u00f05:8\u00de From Eqns (5.7) and (5.8) we have u z R3 \u03b41\u03b51 \u2212=\u2212 \u00f05:9\u00de From Eqns (5.6) and (5.9) we have 3 \u03b4RR z ue \u2212= \u00f05:10\u00de If we substitute the loaded radius as Rl \u00bc Ru dz into (5.10) we get 3 \u03b42RR z le += \u00f05:11\u00de Despite the complexity of the approach, the final approximation in Eqn (5.11) is pleasingly simple and useful in time-varying applications. With some knowledge of vertical stiffness, characteristics, it can be used to estimate rolling radius with useful accuracy in real time applications"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0001619_004051755002000402-Figure1-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0001619_004051755002000402-Figure1-1.png",
+        "caption": "FIG. 1. Diagram showing the arrangement used for photographing the course of impact of waterdrops on a fabric surface. -",
+        "texts": [
+            " Under these conditions the velocity of the drops at the time of impact was about 4.6 m. per sec. (determined from the photographs). The photographic film had a length of 30 m., which enabled two consecutive drops to be photographed on one film. at EMORY UNIV on April 19, 2015trj.sagepub.comDownloaded from 216 The first drop fell on the dry sample. The second drop fell on the same place-therefore, on a wet surface. The drops struck the surface at right angles, and the impact was photographed at nearly right angles to the falling direction (Figure 1). Figure 2 shows the course of imp2ct of waterdrops on dry, water-repellent cotton fabric. The time that elapsed between two consecutive pictures was about 2 X 10-4 sec. Figure 3 shows the impact on dry acetate film. The film was cross-ruled, each square being 5 X 5 mm. These squares can be seen on the pictures and used as a scale. The time between two consecutive pictures was 2.0 X 10-4 sec. (scale, 1:1). The pictures reveal some interesting relationships. During impact the upper part of the drop retains its spherical shape while the lower part is pressed outwards as a continuous film"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0000556_s0146411616030056-Figure1-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0000556_s0146411616030056-Figure1-1.png",
+        "caption": "Fig. 1. Phase portrait of the system with a high gain coefficient.",
+        "texts": [
+            " In [31], by analogy with [39], control (4) is called equivalent to robust. 3.1. Specific Characteristic of Motion For a fairly high gain coefficient, short-term rapid motion occurs, which causes the whole phase f low to reach the hyperplane s = 0 almost immediately. The further motion of the system to the origin of coordinates is slow. Taking into account this peculiarity, in a linear case, the system motion can be described by two components as follows: where is the rapid motion (see the trajectory AB in Fig. 1), while is the slow motion along the hyperplane s = 0. Thus, the system motion is eventually localized on the manifold s = 0 and is described (with sufficient accuracy) by a hyperplane equation on the order of n \u2013 1. In [36\u201338], this behavior of the system is called motion localization. Figure 1 shows (for n = 2) a phase portrait of a system with a fairly high gain coefficient. Within the limit, when the gain coefficient K goes to infinity, the rapid motion exponentially decays exp(\u2013Kt). Then, the system is described by the equation of slow motion along s = 0, which also decays exponentially as \u0435\u0445\u0440(\u2013\u0441i t), i = 1, 2, \u2026, n \u2013 1. For n = 2, the hyperplane equation is with the corresponding solution 4. DETERMINATION OF TUNING PARAMETERS Controller (4) can be tuned in two steps as follows: \u2013 the a priori determination of \u0441i, i = 1, \u2026, n \u2013 1; \u2013 the experimental determination of K"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0000743_gt2016-56592-Figure18-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0000743_gt2016-56592-Figure18-1.png",
+        "caption": "Figure 18. In the initial impeller, a high entropy region is seen in Figure 18(a) at section I. It is located at the corner between the SS and the shroud endwall, showing the highest entropy value in the whole passage. The corresponding streamlines are highly twisted due to the flow separation from the SS. These contribute to a significant loss. This high entropy region diffuses when it moves downstream to the impeller exit, forming a wide coverage of high entropy near the shroud that supports the large wake shown in Figure 17(a). In the optimal impeller, Figure 18(b), the entropy and the entropy production region are distinctly reduced compared with the initial one, and the streamlines also behave well.",
+        "texts": [],
+        "surrounding_texts": [
+            "The optimization of an impeller in a single-stage, industrial centrifugal compressor is presented in this section. The geometrical parameters of the impeller and the target values (by CFD) of the optimization are listed in Table 4. stage were firstly calculated and compared with the experimental measurements. The same type of mesh topology of Eckardt impeller and a similar CFD setup were employed. The results are shown in Figure 11. Both the isentropic efficiency and total pressure ratio are slightly over predicted, considering the omission of the volute that was presented in the experiment by the CFD, the CFD results are reasonable. (a) ISENTROPIC EFFICIENCY The target setting for the optimisation is ambitious. This means large modifications to the baseline geometry. Some of the such changes could be quickly done manually if the impeller designer is experienced, but here all the alterations are left to computer. The number of control points of the hub (ZH,i , RH,i) or the shroud (ZS,i , RS,i) are both 10 (see Figure 13). 24 of 40 variables are allowed to vary. The range and step size of the variations and the constraints for the control points are given in Table 5. They need to be planned carefully so that the optimisation can proceed smoothly and may produce desired results. The step size for the first optimization process is equal to (Upper bound - Lower bound) / 10, and the step size for the second optimization process is all set to be 0.1mm. Table 5 RANGES, STEP SIZES AND CONSTRAINTS FOR CONTROL POINTS OF ENDWALLS Variables First optimization process Second optimization process Lower bound Upper bound Step size Lower bound Upper bound Step size Note: the units are millimeter (mm) ZH,3 35 45 1 35 38 0.1 ZH,4 60 90 3 65 70 0.1 ZH,5 85 115 3 87 94 0.1 ZH,6 100 130 3 101 105 0.1 ZH,7 110 130 2 117 119 0.1 ZH,8 124 128 0.4 RH,2 60 65 0.5 RH,3 60 75 1.5 70 75 0.1 RH,4 65 90 2.5 85 90 0.1 RH,5 90 110 2 100 105 0.1 RH,7 143 145 0.2 RH,8 169 171 0.2 7 Copyright \u00a9 2016 by ASME Downloaded From: http://proceedings.asmedigitalcollection.asme.org/pdfaccess.ashx?url=/data/conferences/asmep/89477/ on 02/05/2017 Terms of Use: http://www.asme.org/about-asme/terms-of-use ZS,4 49 50 0.1 49 50 0.1 ZS,5 65 70 0.5 ZS,6 78 83 0.5 ZS,8 93 103 1 100 101 0.1 ZS,9 95 110 1.5 104 107 0.1 ZS,10 95 110 1.5 104 108 0.1 RS,3 131 132 0.1 131 132 0.1 RS,4 128 138 1 134 136 0.1 RS,5 130 145 1.5 RS,6 140 150 1 RS,8 178 179 0.1 RS,9 199 201 0.2 199 201 0.1 Constraints (fixed): (1) ZH,i(i=1, 2, 9, 10), RH,i(i=1, 6, 9, 10), ZS,i(i=1, 2, 3, 7) and RS,i(i=1, 2, 7, 10) are unchanged in the first optimization process; (2) ZH,i(i=1, 2, 8, 9, 10), RH,i(i=1, 2, 6, 7, 8, 9, 10), ZS,i(i=1, 2, 3, 5, 6, 7) and RS,i(i=1, 2, 5, 6, 7, 8, 10) are unchanged in the second optimization process. Table 6 RANGES, STEP SIZES AND CONSTRAINTS FOR CONTROL POINTS OF BLADE CAMBER CURVES Variables First optimization process Second optimization process Lower bound Upper bound Step size Lower bound Upper bound Step size Note: the units are radian (rad) \u03b8H,2 0.08 0.18 0.02 0.15 0.17 0.002 \u03b8H,3 0.15 0.31 0.02 0.26 0.30 0.002 \u03b8H,4 0.20 0.40 0.02 0.36 0.40 0.002 \u03b8H,5 0.28 0.50 0.02 0.46 0.50 0.002 \u03b8H,6 0.34 0.60 0.02 0.56 0.60 0.002 \u03b8H,7 0.40 0.68 0.02 0.62 0.67 0.002 \u03b8H,8 0.46 0.72 0.02 0.67 0.71 0.002 \u03b8H,9 0.53 0.77 0.02 0.71 0.76 0.002 \u03b8H,10 0.59 0.79 0.02 0.74 0.79 0.002 \u03b8H,11 0.65 0.83 0.02 0.78 0.82 0.002 \u03b8S,2 0.07 0.23 0.02 0.18 0.23 0.002 \u03b8S,3 0.14 0.36 0.02 0.32 0.36 0.002 \u03b8S,4 0.22 0.44 0.02 0.40 0.44 0.002 \u03b8S,5 0.30 0.50 0.02 0.47 0.51 0.002 \u03b8S,6 0.38 0.56 0.02 0.51 0.54 0.002 \u03b8S,7 0.44 0.60 0.02 0.59 0.61 0.002 \u03b8S,8 0.50 0.64 0.02 0.59 0.64 0.002 \u03b8S,9 0.58 0.64 0.02 0.58 0.64 0.002 \u03b8S,10 0.60 0.68 0.02 0.62 0.66 0.002 \u03b8S,11 0.64 0.74 0.02 0.65 0.73 0.002 Constraints: (1) \u03b8H,1 and \u03b8S,1 are unchanged in the optimization processes. (2) \u03b8H,i<\u03b8H,i+1(i=1,2\u202610); \u03b8S,i<\u03b8S,i+1(i=1,2\u202610) in the two-step optimization process. The sections of blade camber curve are two at 0% and 100% spans in the optimization, and other camber sections are determined by linear interpolation from these two. The number of control points of blade camber curves at the hub (0% span) \u03b8H,i and shroud (100% span) \u03b8S,i are both eleven (see Figure 14). The ranges, step sizes and constraints for \u03b8H,i and \u03b8S,i are listed in Table 6. The leading edge of blade was set to be fixed values (ZH,2=0mm, ZS,2=0mm), while the trailing edge was unchanged. The blade thickness was unchanged too in the optimization. Figure 12 THE POPULATION DISTRIBUTION OF TYPICAL ITERATIONS IN OPTIMIZATION AND RESULTS Figure 13 ENDWALLS OF INITIAL, MEDIAL AND OPTIMAL IMPELLERS 8 Copyright \u00a9 2016 by ASME Downloaded From: http://proceedings.asmedigitalcollection.asme.org/pdfaccess.ashx?url=/data/conferences/asmep/89477/ on 02/05/2017 Terms of Use: http://www.asme.org/about-asme/terms-of-use The population size was 30 and the maximum iteration number was 25 for each of the two optimization processes. The geometry and CFD prediction results were saved in the database for each iteration. The convergence criterion of the optimization was the isentropic efficiency over 95.00 percent and the total pressure ratio \u2265 2.15:1. The crossover probabilities were 1 0.9cP  and 2 0.6cP  , and the mutation probabilities were 1 0.5mP  and 2 0.1mP  for the first optimization, and 1 0.9cP  , 2 0.3cP  , 1 0.1mP  and 2 0.01mP  for the second optimization. The medial result from the first optimization was obtained after reaching the maximum iterations in about 22 hours. The isentropic efficiency and the total pressure ratio were 93.98 percent and 2.161:1 respectively. The result was then taken as the starting point to begin the second optimization. The final result was obtained after reaching convergence criterion (EFF>95%, TPR >2.15) in about 31 hours (19 iterations). The total computational time required to run the two-step optimization was about 53 hours with 1380 CFD runs. The computational time of the second optimization was longer than the first one because of its smaller global residual requirement in CFD calculations. The population distribution of several typical iterations are given in Figure 12. It shows the trend of population distribution in the optimization processes moving toward the high efficiency and high pressure ratio region. The medial and the final impellers are marked with blue circles, together with the other two candidates A and B that also satisfy the performance target. The final one was selected because of its higher efficiency and of the higher weighting given to impeller isentropic efficiency than to total pressure ratio in the sorting process of the improved NSGA-II. The scattered population distribution of final iteration indicates that the improved NSGA-II is able to find a good spread of solutions efficiently at Pareto-optimal front and keep the diversity of populations. Figure 14 THE CONTROL POINTS OF BLADE CAMBER CURVE AT THE HUB (0% SPAN) The meridional geometries of the initial impeller and the two optimal impellers are compared in Figure 13. As can be seen that there is a significant change of the endwalls in the first optimization. Such large modifications to the endwalls would normally completed manually but here it is achieved automatically. The cambers at hub of the three impeller are compared in Figure 14, and the 3D model of the initial and the final impellers is shown in Figure 15. It is worth to mention that the big changes to the endwalls and the camber by the optimisation will greatly modify the mechanical properties of the impeller blades, and a proper mechanical design is usually needed to qualify the impeller for production. However, this work focuses only on the optimization of aerodynamic performance of centrifugal impellers, the mechanical design and impeller manufacture of such impellers are not the subjects of this paper and will not be considered further here. The impeller performance by CFD is presented in Table 7. It shows a great improvement of both the isentropic efficiency and the total pressure ratio, and reaching the target values. Table 7 THE OVERALL PERFORMANCES OF INITIAL AND OPTIMAL IMPELLERS Parameters Initial Optimal Improvement EFF 90.93% 95.76% 5.3% TPR 1.820 2.194 20.5% The velocity vectors and relative Mach number contours for the design operating condition at 90% span of the two impellers are shown in Figure 16. A large low momentum area can be seen at the suction side (SS) of the initial impeller caused by the flow separation from the suction surface, which is eliminated in the optimal impeller. The flow diffusion in the optimized impeller is much smoother than in the baseline impeller, indicating an improvement of flow field in the former. 9 Copyright \u00a9 2016 by ASME Downloaded From: http://proceedings.asmedigitalcollection.asme.org/pdfaccess.ashx?url=/data/conferences/asmep/89477/ on 02/05/2017 Terms of Use: http://www.asme.org/about-asme/terms-of-use The comparison of meridional velocity contours for the de- sign operating condition at impeller exit (section IV, / 1.00mx S  in Figure 7) is shown in Figure 17. The jet-wake pattern can be seen near the SS in the both initial and optimal impellers, but the wake region of the optimal impeller is significantly reduced from that of the baseline. This is a further indication of flow field improvement after the optimization and the mixing loss in the downstream diffuser will be reduced. The comparison of entropy contours (reference point 94450Pa, 303oK) and streamlines (white curves) at sections I ( / 0.25mx S  ), II ( / 0.50mx S  ), III ( / 0.75mx S  ) and IV ( / 1.00mx S  ) in the design operating condition are shown in in Figure 19. The results show that the performance gains at the 10 Copyright \u00a9 2016 by ASME Downloaded From: http://proceedings.asmedigitalcollection.asme.org/pdfaccess.ashx?url=/data/conferences/asmep/89477/ on 02/05/2017 Terms of Use: http://www.asme.org/about-asme/terms-of-use design condition also extend to other operation conditions except near the choke. The choke mass flow rate of the optimal impeller is reduced by about 3.6% from the baseline, because of the decrease of its inducer flow area (see Figure 13)."
+        ]
+    },
+    {
+        "image_filename": "designv11_60_0001805_10402009808983760-Figure4-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0001805_10402009808983760-Figure4-1.png",
+        "caption": "Fig. 4--Application of slide lines In Interbristle contact pairs.",
+        "texts": [
+            " If the interacting bodies are cylinders/tubes, as in the case presented here, through the use of special slide line contact elements, the relative motion is allowed to be along a culve of contact. In this case, the contact direction is normal to the slide line in the direction of the smallest distance between the surfaces of the cylinders (bristles). For a bristle couple, all the nodes in one of the bristles form the slide line. The corresponding second-order slide line contact elements on the other bristle are defined in such a way that they start with the second node at the free tip, including all the nodes along the bristle up to the third node from the top. Figure 4 illustrates the use of slide lines between the bristles. The model does not include the first node at the free tip in the contact elements to avoid overconstraining the tip node as it is also in contact with the rotor surface. The last two nodes at the top of the bristles are also spared, as they would be overconstrained. Bristles are fixed at this end, and these nodes do not see any sizable sliding. The regular pressure-clearance relationship, which is used for the bristle-rotor interfaces, is not used in bristle-bristle contacts for two main reasons"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0002594_bf00046883-Figure3-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0002594_bf00046883-Figure3-1.png",
+        "caption": "Fig. 3. Construction of the centrode of a 4-bar.",
+        "texts": [
+            ", ABIICD (Figure l(a)). Differentiating again gives nontransversality precisely when d 2 - d 4 . This means that the 4-bar is type II or III and that any parallelogram configuration is E N G (Figure l(b)). To locate L \u00b0'\u00b0 points we use Proposition 2.1(a). Since B and C lie in the moving body and their trajectories are perpendicular to AB and DC respectively, GENERIC PROPERTIES IN EUCLIDEAN KINEMATICS 283 the instantaneous centre, P, must lie on the intersection of the extensions of these two lines (Figure 3). (This also explains why K 2 points occur when the lines are parallel.) It follows that if the centrode has a singularity at P then AB, DC are instantaneously stationary, which means that the motion has an L 2 point rather than an L \u00b0'\u00b0 point, unless P = A or D, and either B = D or C = A. In this case, the 4-bar must be of type II' or III and the motion lie in a branch o[ the residual curve corresponding to one pair of sides rotating about the fixed point A or D (Figure l(c)). These then are E N G configurations"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0000804_jahs.61.042006-Figure8-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0000804_jahs.61.042006-Figure8-1.png",
+        "caption": "Fig. 8. Tregold\u2019s approximation: (a) formative spur gear and (b) mapping from face-gear drives to spur gear drives.",
+        "texts": [
+            " 13, 22, 23) w|r=b = 0, \u2202w/\u2202r|r=b = 0 (6) [ \u22022w \u2202r2 + \u03bd ( 1 r \u2202w \u2202r + 1 r2 \u22022w \u2202\u03b8 2 )]\u2223\u2223\u2223\u2223 r=a = 0 (7)\u23a1 \u23a2\u23a2\u23a3 \u2202 \u2202r ( \u22022w \u2202r2 + 1 r \u2202w \u2202r + 1 r2 \u22022w \u2202\u03b8 2 ) +1 \u2212 \u03bd r2 \u22022 \u2202\u03b8 2 ( \u2202w \u2202r \u2212 w r ) \u23a4 \u23a5\u23a5\u23a6 \u2223\u2223\u2223\u2223\u2223\u2223\u2223\u2223 r=a = 0 Contact ratio calculation The contact ratio of the spur pinion/face-gear meshing is calculated via numerical methods based on differential geometry and theory of gearing (Ref. 6). Since no exact closed-form expression exists, this paper employs so-called Tregold\u2019s approximation (Ref. 14) for contact ratio calculation. In this method, the face-gear is represented by a formative spur gear to engage with the pinion, illustrated in Fig. 8(a). This is a common practice used in bevel gearing. The mapping from face-gear drives to formative spur gear drives is drawn in Fig. 8(b). Here, the solid circles 1 and 2 represent the formative spur gears for the spur pinion and the face-gear, respectively. Rp and Rg are the pitch radii of the pinion and the face-gear, respectively, and Rvp and Rvg are the pitch radii of the formative spur gears; A is the distance from the vertex of pitch cones O to the pitch point P (along the OI line); \u03b31 and \u03b32 are pitch cone angle of pinion and face-gear, respectively; \u03b3 is pitch cone angle of face-gear drive. The relationships between these angles are given in Ref"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0003234_ma00182a011-Figure1-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0003234_ma00182a011-Figure1-1.png",
+        "caption": "Figure 1. Geometry of the reaction chamber and relevant cylindrical coordinates.",
+        "texts": [
+            " We are able to arrive at an exact analytical solution for the distribution. By comparing the approximate with the exact solution we obtain a measure of its error. Finally, we examine the possibility of generating isolated polymers, in the gas phase (as in the case of styrene vapor), large enough to be observed directly by light scattering methods. Formulation and Solution of the Boundary Value Problem, Exact and Approximate As indicated in the previous section we shall deal with a cylindrical reaction chamber. Figure 1 will be helpful. The chamber is of radius a and height 2L. We choose r for the radial coordinate and z for the vertical one and position the chamber such that - L l z l L O l r l a (1) Nonterminated Chain Polymerization 913 The reactions which take place in the chamber are of the type monomer - R', 1 R' + M - RM', RM' + M - RMM', kl[M]P1 k2[M]P2 etc. (2) in which R' denotes a free radical and M a monomer. R', R M , and RMM' are polymers of sizes 1, 2, and 3 whose concentrations are denoted by Pl, P2, and P3, respectively"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0000414_978-3-662-46463-2_40-Figure40.3-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0000414_978-3-662-46463-2_40-Figure40.3-1.png",
+        "caption": "Fig. 40.3 The prototype of powered knee prosthesis",
+        "texts": [
+            " The database contains the motor\u2019s pulse control signal of direction and displacement in the different gait phase under a variety of terrain conditions. As shown in Fig. 40.2, the gait phase divides a walking cycle into four phases, which are early stance phase, middle stance phase, late stance phase, and swing phase, respectively. The phases can be recognized via two pressure switches mounted under the sole of the prosthesis foot, one under the ball flat and another under the heel. The prosthesis knee joint used in this paper is shown in Fig. 40.3. The four connecting rods driven by the motor make the change of knee joint\u2019s angle. The relationship between prosthesis knee joint angle (h) and motor shaft displacement (L) is described in Eq. 40.1. And the motor shaft displacement is 0.0508 mm for each pulse of signal. So the number of pulse signal can be calculated by knee joint angle\u2019s change for establishing the displacement database. h \u00bc 2:188 L \u00f040:1\u00de When the knee joint sways in reverse direction, the direction of the driving motor should change at the same time"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0002863_19.744189-Figure6-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0002863_19.744189-Figure6-1.png",
+        "caption": "Fig. 6. Radii of the filamentary loops at two circular loops with finite conductor radii.",
+        "texts": [
+            " The filamentary conductors for two infinite parallel lines have been discussed by Kraus [6, Figs. 3\u201316, p. 84] and by Ku\u0308pfmu\u0308ller [2, Abb. 11.14, eq. (92) on p. 92]. Here it is assumed that the total conductor current flows through the fictitious thin filamentary loop. The relation gives the radius of the current filamentary loop, where is the radius of transmitting loop and is the conductor radius. The receiving loop with finite conductor radius can encircle a part of a magnetic field with the loop radius . From Fig. 6 the equivalent thin current filament radii of the circular loops L1 and L2 (9a) (9b) The radii of the effective circular surface of loop L1 and L2 (9c) (9d) If the loop L2 has many windings with a fixed geometry, the effective thin filament radius and the effective distance must be defined [3, pp. 17\u201325]. For higher calibration accuracy, the transmitting loop should have only one winding. For Hz, (dc) the current flows uniformly through the conductor cross-section. For this case and . A formula for interpolation between dc and high frequency in terms of the skin depth is still lacking and needs a more accurate calculation of and "
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0003383_acc.2000.879261-Figure1-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0003383_acc.2000.879261-Figure1-1.png",
+        "caption": "Figure 1: Aerodynamic model for air to air missile",
+        "texts": [
+            " Convergence to a local solution with linear rate for any feasible starting point. Furthermore, the approach here is more general since a best y level is computed whereas the Frank & Wolfe technique will require a less efficient dichotomy scheme to minimize y. Through examples, we shall also discuss the efficiency of the approach as compared to the classical D - I< iteration method which does not enjoy good convergence properties. 4.1 Autopilot robust control of missile Consider the missile-airframe control problem illustrated in Figure 1. when the vehicle is flying with an angle of attack (a). The control problem requires that the autopilot generate the required tail deflection (6) to produce an angle of attack. corresponding to a maneuver called by the guidance law. Sensor measurements for feedback include missile rotational rates q (rate gyros) and a. For the problem considered here. it is desired to track step input commands a, with a steady state accuracy of 1% and to achieve a rise time less than 0.2 second, and limit overshoot to be 2% over a wide range of angles of attack 5 2 0 deg and variations in"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0001235_b978-0-08-099425-3.00005-4-Figure5.58-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0001235_b978-0-08-099425-3.00005-4-Figure5.58-1.png",
+        "caption": "FIGURE 5.58",
+        "texts": [
+            " A tyre model, for handling or durability analysis, requires input regarding the position and orientation of the wheel relative to the road together with velocities used to determine the slip characteristics. The implementation of these computations as a tyre model with an MBS program is best described using the full threedimensional vector approach outlined in Chapter 2. The following description is based on the methods used in MSC.ADAMS but is applicable to any vehicle simulation model requiring tyre force and moment input. As a starting point the tyre can be modelled using the input radii R1 and R2 as shown in Figure 5.58. Using the tyre model geometry based on a torus it is possible to determine the geometric outputs that are used in the subsequent force and moment calculations. Consider first the view in Figure 5.59 looking along the wheel plane at the tyre inclined on a flat road surface. The vector {Us} is a unit vector acting along the spin axis of the tyre. The vector {Ur} is a unit vector that is normal to the road surface and passes through the centre of the tyre carcass at C. The contact point P between the tire and the surface of the road is determined as the point at which the vector {Ur} intersects the road surface"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0001235_b978-0-08-099425-3.00005-4-Figure5.17-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0001235_b978-0-08-099425-3.00005-4-Figure5.17-1.png",
+        "caption": "FIGURE 5.17",
+        "texts": [
+            " Other examples where confusion may arise include the use of aligning torque, aligning moment or self aligning moment, longitudinal or tractive force and lateral or cornering force. Tyre forces and moments shown acting in the SAE tyre axis system. The use of the term stiffness can also add confusion to newcomers to the subject area. A traditional static force/displacement approach is used by Moore (1975) to define longitudinal, lateral and torsional stiffness of a tyre. In each case a nonrolling tyre is mounted on a plate and incrementally loaded as indicated in Figure 5.17 until complete sliding occurs. Plotting graphs of force or moment against displacement or rotation allows the stiffness parameters to be obtained from the slopes at the origin. Wewill see later that terms such as cornering stiffness and aligning moment stiffness are associated with a rolling tyre and should not be confused with the lateral and torsional stiffness defined here. The term longitudinal stiffness can be particularly misleading as another definition is commonly used when longitudinal tractive forces due to driving and braking are discussed"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0000685_med.2016.7536020-Figure1-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0000685_med.2016.7536020-Figure1-1.png",
+        "caption": "Figure 1. Gear model with 16-degree-freedom[18].",
+        "texts": [
+            " (12) and (13) respectively: 2 2 k k ka t f t H f t (12) and ( )1 1 2 2 ( ) kk i k H f td t d F t arctang dt dt f t (13) A single stage spur gearbox is considered. It is composed of a specific pinion-gear elements modeled as a pair of rigid disks connected by a spring damper set along the line of contact mounted on a flexible input/output shafts which in turn are supported by bearings introduced as additional stiffness elements, and a load machine [18]. The dynamic system behavior includes 16 degrees-offreedom. The schematic diagram of the model is shown in Fig. 1. This study takes into account the following simplifying assumptions: Resonances of the gear case are neglected. Shaft mass and inertia are lumped at the bearings or the gears. Shaft transverse resonances are neglected. Shaft torsional stiffness is ignored (flexible coupling torsional stiffness is very low). Gear teeth profiles are perfect involute curves, with no geometrical, pitch or run out errors. Inter-tooth friction is ignored. A. Differential equations of motion The resulting equations of motion describing the model are given by Eqs"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0003042_iros.1993.583851-Figure3-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0003042_iros.1993.583851-Figure3-1.png",
+        "caption": "Fig. 3: an example of gripping plane",
+        "texts": [
+            "2 Planning Grasps for a Parallel Twofingered Gripper In this section we explain a method for planning grasps for a parallel two-fingered gripper and checking whet.her the goal can be achieved or not by a pick-and-place operation. The objects considered in this paper are polyhedra. A pair of grasp elements is defined as the two elements of the object (faces, edges and vertices) which two fingers of the gripper contact with when the gripper grasps it in a stable way. \\\\'hen the gripper grasped the object the gripper is constrained to the gripping plane, which is constructed at the center of the grasp elements pair (as shown'in Fig.3). The constraint is for the origin of the gripper to lie on the gripping plane and for the y axis of the gripper to be perpendicular to it. Then the mentioned G(obj) and GE(obj, env) are divided into subsets so that each subset is associated with each gripping-plane. The divided subsets associated to gripping plane p l are called GP(obj.pl) and GEP(obj, env,pl) (see Eq.7). G(obj) = U GP(obj,pl) GE(obj. env) = UGEP(obj , env,pl) (7) Pl Pl Due to the constraint on the gripping plane, the motion of the gripper is described in a 3-dimensional configuration space, which is constructed from 2-d translation on the gripping plane and 1-d rotation around the normal vector of the gripping plane"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0003155_cdc.1987.272906-Figure6-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0003155_cdc.1987.272906-Figure6-1.png",
+        "caption": "Figure 6. All the operations needed to calculate this control law are given in Table 2.",
+        "texts": [],
+        "surrounding_texts": [
+            "for i = l , ..., n iVi = iAi-l (i-lVi-l + i-l X i-l Pi) + oi {i.Zo (30) i w i = iAi-l i-l 0i-1 +'oi 41 Zo (31 1\nThis choice is not only because the recursive procedures have been revealed efficient at the minimization of the computational cost , but especially because wl, ..., wn will be used in the inverse dynamic problem calculations.\nSince w in the Newton-Euler algorithm will be calculated with respect to frame i , then we calculate first ' i n , and then we project it into the desired frame.\nExample: The calculation of and 3v3 , which is equal to 3V6 , for the Stanford Manipulator needs 19 multiplications and 11 additions .This model I S given in appendix 1 .\n5.3 Calculation of ;(q& ;1 The calculat,on of this term by differentiating the jacobian\nmatrix will need a prohibitive number of operations, we propose once more a recursive algorithm whose elements will be used at the dynamic inverse calculation.\nFrom Eq.(18), and since X, = [ i n ;n ITwe conclude that:\ni ( q , i ) 4 is equal to [ i n in IT when { =O\nI.\nTherefore , ;(q& 4 can be deduced from the equations of \\ii and I!O by eliminating from them the terms containing Gi.\nLet i ( q , i ) 4= I :I By the use of Eqs.(5),(6),(7) we get:\ni y I . -iA. 1-1 '-1 Y i-1 +Z i (i ~ i - 1 X ;li zo) (32) . A iu*. - I 0. i si + lyi . A\nI - I (33) io. I = i ~ i - 1 (i-1 0, 1-1 + i- 'u*i-l i - l ~ i ) + 2 q i q - 1 X ((ii 2,)\n(34) we have oYo=O,o~o=O Example : in the case of Stanford Manipulator the calculation of 0 and Y needs 31 multiplications and 15 additions. The model is given in appendix 2.\n5.4 Calculation of J(q)- ly\nThis problem has been treated in the regular case in many references and efficient methods have been proposed [13,14,15,16].\nIn this paper , we utilize the method developed by Renaud [16], which gives the jacobian matrix by the following decomposed relation:\n0 .h J = I\",*i Oni1 I -'Lj,n 13 1 'Jj (35)\nwhere : 13 :the (3~3)identity matrix.\nOllZ1XiLl , j + 01 121 ... On - IZnXILn,j+ . . On 12,\nIJ, = - ... On Qn ' ' 1 (36) It has been shown that, in general the best value of i and j (giving the simplest expressions for IJj) is equal to p and p+l respectively, where p is equal to the integer part of (1112).\nThe inverse of J is give2 as : ,\nJ-1 =iJ,-l J 1 ':In [ I 'to iAo I (37) 0\nTherefore, J-l y will be, equal to: J -1 y = IJ,-1\nJ lyj\nThe calculation of iJj'l iy, will be directly realized without inverting explicitly the jacobian matrix.\nExample: for the Stanford Manipulateur ,we find that 3J3 is the simplest iJ, and L6 3=O. The calculation of 3J3-1 3y3 needs 21 multiplications and 9 additions [21,22].\n5.5 The modified dynamic inverse calculation\nWe will modify the Newton-Euler algorithm described previously, to take into account hat 0 and have been calculated. The equations (5),(6),(7) calculating Ai and ci will be replaced\nby : , * 1V. - io , + ibi\nI - I (40) e ' i 9 = 1yi + iei (41 1\nwhere jbi , iei represent the linear and rotational acceleration when q = 0 ;\nibi = iAi-l (i-l bi-l + i-l ei-1 I-1 Pi) + oi ';i 2, (42)\niei = i ~ ~ - ~ i-lei-l + oi qi 2, we have Obp= g, Oe0=O The matrix IUi will be calculated as:\nA .\n- I*\n(43)\n. . .h Iu. =Iu*. + le\nI I ~ (44) Example : The dynamic inverse calculation of the stanford manipulator is given in appendix 3. The computational cost is 11 0 multiplications and 82 additions . The inertia matrix of the links are supposed diagonal, each first moment vector has one component different than zero, the terminal forces and the motor inertias are taken into account. This is the case denoted by simplified in table 1.\n5.6 The complete computational scheme\nIn the previous sections the calculation of the components needed to the Cartesian dynamic ontrol have been presented, but sqme results are expressed in frame 0, others in frame 3 or 6. The complete block diagram for the Stanford Manipulator, which takes into account this fact is given in",
+            "I . . I\n6. CONCLUSION\nThis paper presents a complete computational scheme for controlling a robot in the Cartesian coordinates. The given algorithm deals with the inverse dynamic and inverse kinematic problems simultaneously in order to eliminate many redundant calculations. The inverse dynamic calculation is based on a Newton Euler formulation which makes use of many variables and elements which have been calculated during the inverse kinematic solution. The proposed method is general and can be used for any robot. A software package for the symbolic modelling of robots \"SYMORO\" [23] has been developed to generate automatically all the elements needed in this control strategy.\nFiaure 6. REFERENCES\n[I] O.Khatib, '' A Unified Approach for motion and force Control: The operational space formulation\", IEEE conference on Robotics and Automation, Sari Francisco 1986. [2] C.S.G.Lee, M.J.Chung, \"An Adaptive control strategy for mechanical Manipulators\", IEEE Transaction on Automatic Control, Vol.AC-29, N\"9, September 1984 pp.837-840, [3] J.Y.S.Luh, M.W.Walker, R.Paul, \"Resolved-acceleration control of mechanical manipulators\", IEEE trans. Auto. Control, vol. AC-25, pp.468-474, june 1980. [4] M.W.Walker, \"Dynamic Cartesian coordinate control of a manipulator\", ACC,1984, San Diego ,pp. 866-871, [5] J.Y.S.Luh, M.W.Walker, R.Paul, \"On-line computational scheme for mechanical manipulators\", ASME Transaction, J. of Dynamic Systems, Measurements and Control, Vol. 102, No\n[6] W.Khalil, J.F.Kleinfinger, M.Gautier, \"Reducing the computational burden of the dynamic model of robots\", IEEE conference on Robotics and Automation, San Francisco 1986,\n[7] J.F. Kleinfinger, \"Modelisation dynamique de robots a chaine cinematique simple, arborescente ou fermee, en vue de leur commande\",doctorat thesis, Nantes,l986. [8] W. Khalil, \"Technique de modelisation et de commande dynamique\", Colloque AD1 SYSCOROB, Paris, Dec. 1986. 191 W.Khalil, J.F.Kleinfinger, \"Minimum operations and minimum parameters of the dynamic models of tree structure robots\", IEEE ,J. Robotics and,Automation, to appear, [IO] D.T.Horak, \" A fast computational scheme for dynamic control of manipulators\", American control conference 1984,\n[ l l ] T. Kanade, P. Khosla, N. Tanaka, \"Real-time control of CMU direct-drive arm II using customized inverse dynamics\", Proceeding of 23 th CDC Las Vegas, 1984, pp 1345. [12]M.Renaud \"Quasi-minimal computer of the dynamic model of a robot manipulator utilizing ther Newton-Euler formalism and the notion of augmented body\", Proceedina of IEEE\n2, 1980, pp.69-76.\npp.525-531\npp.625-630.\nconference on roboticsand automation , Raleigh-bSA,1987, D D . ~ 677-1 682 . . . (131 J. M. Hollerbach, G. Sahar, \" Wrist- partitioned Inverse kinematic Accelerations and manipulator Dynamics\",lnt.J. Robotics Research,vol. 2, pp.67-76,Winter 1983. [14]R.Featherstone,\"Position Velocity ransformations between robot and effector coordinates and joint angles\" Int. J. Robotics Research 2,2 (June 1983)",
+            "[IS] R.Paul, \"Robots manipulators: mathematics, programming and control, MIT press, 1981. [I 61 M.Renaud, \"Calcul de la matrice jacobienne necessaire a la commande coordonne d'un manipulateur\", Mechanics and Machine Theory, vol.15, n c l , pp. 81 -91, 1980, [ i 7]C.H.Liu, Y.M.Chen, \"Multiprocessor-based Cartesian space control techniques for a mechanical manipulator\", IEEE, J. Robotics and Automation, Fev. 1986, pp. 823-827, [I81 W.Khalil, J.F.Kleinfinger, \"A new geometric notation for open and closed loop robots\" IEEE conference on Robotics and Automation, San Francisco 1986,PP.l 174-1 180, [I91 W.Kha1il. A.Liegeois, A.Fournier, \"Commande Dynamique de Robots\", RAIRO Automatique, Systems Analysis and Control, Vo1.13, N\"2, 1979, pp. 189-201. [20]C.Chevallereau,W.Khalil,\" Efficient method for the calculation of the pseudo inverse kinematic problem\", IEEE conference on Robotics and Automation, 1987,pp. 1842-1 848 [21]W.Khali!,C.ChevaIlereau,\" Utilisation du rnodele dy,namique pour la commande des robots\", journees de robotique, INRIA, Sophia- Antipol is, june 1987, France. l22] C. Chevallereau, \"Cornmande des Robots Manipulateurs dans I'espace cartesien\", doctorat thesis ,Nantes 1988, to appear [23] W. Khalil, J.F. Kleinfinger, C. Chevallereau, \" SYMORO: systeme pour la modelisation des robots\", Notice d'utilisation, Note interne LAN , Nantes ,Fev. 1987.\nAPPENDlXl. Calculation of the Cartesian Velocity Notations: 'mi-1 is represented by [ W l l i WI2i w13i 1 T\n'mi is represented by [W l i W2i W3i ] T\n6i is represented by QPi. i -1 + i-1\nVi-1 mi-1 x i-lPi is represented by: [vs l i 2 i VS~~IT\ni V is represented by [V l i V2i V3i]T\nAPPENDIX 2. The calculation of &q,$ q\nNotations: 'Yi is represented by [PSlli PSl2i PS13iIT\n' a i is represented by [PHI11 PH12i PH13iIT\n' - ' 01 -1 + i-lU*i-l i- lPi is represented by:\n[PHIIIi PH112i PH113iIT U E l l i UE12i UE13i\nU'i= UE2li UE22i UE23i UE3Ii UE32i UE33i\nN0tations:The following table gives the meaning of the variables used in the derivation of the model\nVariables of the model ~~~~~\nACCi [EWIi EW2i EW3iIT\n[WPli ~ ~ 2 i WP~~IT U l l i U12i U13i U21i U22i U23i U31i U32i U33i [MXi MYi MZiIT XXi XYi Zi XYi YYi YZi XZi YZi ZZi [VSPI i V S P ~ ~ VSPBIIT [B l i ~ 2 i 3 l T [VPli V P ~ , VP~~ IT [Fl i F2i F3iIT [Nl i N2i N3iIT [SDJIi SDJ2i SDJ3il [El i E2i E3i] [ M l i M2i M3i]\nGAMi"
+        ]
+    },
+    {
+        "image_filename": "designv11_60_0000077_icoin.2016.7427086-Figure11-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0000077_icoin.2016.7427086-Figure11-1.png",
+        "caption": "Fig. 11. The path taken by UAV in scenario 1 with constant obstacle velocity. (a) Using To Goal search (b) Using Maximum Velocity search",
+        "texts": [
+            " Using the mentioned heuristics, we have tested our approach with the following parameters and scenarios: Sensor\u2019s distance range = 7 meters. UAV can detect obstacle within 7 meters. Time step ti increments at every 0.1 second. We generate velocity obstacle and choose UAV\u2019s velocity every 0.1 second. UAV\u2019s maximum speed = 5 m/s. UAV\u2019s initial position= (0, 0). Goal\u2019s position = (0, 13). In this scenario, the obstacle\u2019s position at t0 = (5, 7) with constant velocity (vx = \u20134 m/s, vy= 0 m/s). Applying TG search, as shown in Fig. 11(a), the UAV stops moving at t3 then starts moving at t6 in varying speed until t16. At t17 it moves at maximum speed until it reaches goal at t33 or 3.3 seconds. Using MV search with same velocity and initial position, as shown in Fig. 11(b), UAV maintains its maximum speed throughout the entire path. It moves to the left at t3 then it starts moving to the right at t13. It reaches goal at t38 or at 3.8 seconds which is slower than TG search by 0.5 seconds. For this scenario, the obstacle\u2019s position at t0 = (6, 1). The obstacle is moving with constant velocity (vx = \u20133 m/s, vy= 4 m/s). Using TG search, as shown in Fig. 12(a), the UAV moves with maximum speed at t1 then stops at t2. It starts again to move at slow speed at t7 and moves at changing speed until t21"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0003253_1520-6416(20010130)134:2<36::aid-eej5>3.0.co;2-j-Figure12-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0003253_1520-6416(20010130)134:2<36::aid-eej5>3.0.co;2-j-Figure12-1.png",
+        "caption": "Fig. 12. Experimental setup of torsional system.",
+        "texts": [
+            " In calculating the time response, it is used as a nonlinear element as is; however, when looking at the frequency response, it must be handled as an equivalent gain by using a descriptive function. If the input is A sin Zt, the descriptive function Kl of the dead zone can be given by the backlash are combined, a model of a two-inertia system with backlash can be obtained; however, in addition to the inertias of motor and load, there are two inertias of gear which should be added, resulting in a four-inertia system which is very large. For simplicity in this paper, by considering, in the experimental machine shown in Fig. 12, the distinctive structural feature that the gear is positioned close to the motor, we believe that the model of three inertias shown in Fig. 4 may be sufficient and therefore we will use it.Fig. 2. Two-inertia system. (1) (2) Jm is the sum of motor inertia and gear inertia on the motor side, Jg the gear inertia on the load side, Jl the load inertia, and Ks the elasticity of the shaft. 3.1 Gear torque observer In the backlash model shown in Fig. 3(b), if it is assumed that load torque Ts = 0 and there is no backlash, since we have Eq",
+            " In the adjustment, it is desirable to maintain the original response performance as much as possible. Accordingly, Kd only is adjusted in such a way that the equivalent time constant W Kp /Ki which expresses the basic response in the coefficient graphic method is not changed. Here, since the stable condition of the system is given by Eq. (18), Eq. (19) will also be used taking into account the modeling error. Namely, the PID controller shown in Eq. (16) will be used in the speed controller; however, Eq. (19) will be used as Kd only when the controlled system satisfies Eq. (17). Figure 12 shows the configuration of the experimen- tal setup of the two-inertia system with gear used in this paper. The gear inertia Jg, load inertia Jl, and elasticity coefficient Ks of the shaft can be changed in a certain range; however, the results of the case of the parameters given in Table 1 only are shown in this paper. (17) (18) (19) (16) The backlash angle 'T is set as 1\u00b0; however, this is employed only for the simulation and is not used in the design of the control system. In the simulations and experiments, as the inputs, a speed command of 20 rad/s is given stepwise at t = 0, and a disturbance torque of 1 N is also given stepwise at t = 0"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0000277_12.2218911-Figure3-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0000277_12.2218911-Figure3-1.png",
+        "caption": "Figure 3. (a) Schematic of the normal force test setup, including the electroadhesive device, the adhering surface and the support through which the adhering surface is connected to the load cell. (b) Normal force for a cardboard substrate at 0 or 4 kV applied voltage. In order to detach the adhering surface and its support from the electroadhesive device, the Instron must apply a force greater than the weight of the test surface and support. The plateau force corresponds to the weight of the cardboard adhering surface and its support. The difference between the maximum force measured with and without applied voltage and the plateau weight was taken as the value of the electroadhesive normal force and intrinsic chemical adhesion, respectively.",
+        "texts": [
+            "org/ss/TermsOfUse.aspx from Techspray), for 10 seconds. The silicone elastomer was allowed to cure for 24 hours, after which the surface became non-adhesive to most substrates, except rigid acrylic, as noted in the results section. Electroadhesive devices were constructed specifically to test normal or shear adhesive forces. To test normal adhesion forces, the elastomer was cast on a polyethylene terephthalate glycol (PETG) substrate, which was then cut to size and placed flat onto a rigid support as shown in Figure 3 (a). The electrode made of SWCNTs at a concentration of 7.5 mg/m2 was transferred via stamping. The connections to the power source were made using conductive carbon tape (Ted Pella, Redding CA). To test shear adhesion forces, the process was identical, except that the elastomer was placed flat onto a sheet of Mylar (75 microns thick). Due to the small thickness of the Mylar backing, the adhesive device could be flexed to ensure good contact with the adhering surfaces during shear testing. A diagram describing the shear test is shown in Figure 4 (a)",
+            "aspx Dielectric elastomer actuators were made by pre-stretching commercial acrylic elastomers (VHB 4910) to 200% strain in both x and y planar directions using a biaxial stretcher. Circular electrodes made of SWCNTs were adhered directly onto the pre-stretched elastomer. As described in part 2.1 and shown in Figure 2 (b), adhesive device fabrication was completed by spin coating a layer of acrylic elastomer directly on the DEA. The polymer was then cured under UV light, and the electroadhesive device completed as described earlier. The effectiveness of electroadhesion is quantified relative to the adhesive force measured without any applied voltage, as shown in Figure 3 (b). In order to for the Instron to lift the support and adhering surface off the adhesive device, the force applied must be at least equal to the weight of the support and adhering surface. The value of the weight is taken from the plateau force value, once the adhering surface has been completely peeled off the electroadhesive device. The data presents the adhesion force as the cardboard adhering surface is pulled away from the device in a normal force experiment. The force is measured with and without applied voltage"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0002923_978-94-011-3514-6-Figure11.1-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0002923_978-94-011-3514-6-Figure11.1-1.png",
+        "caption": "Fig. 11.1 Symmetrical triple pendulum",
+        "texts": [
+            " The main use for the Rayleigh quotient is to make an estimate of the fundamental natural angular frequency of a system by using an approximate fundamental eigenvector, determined by physical considerations (an example is given in paragraph Finally, it is ea~;y to verify that the relationship (11.79) can be put into the following equivalent form R(u) This is preferable when one knows the flexibility matrix [crl. ( 11 .87) (11.88) (11. 89) - 208 - The calculation of small movements of a system cnnsisting of three massive, equal and symmetrically coupled pendulums (symmetrical triple pendulum, figure 11.1) is an elementary and classical example to make the concept of natural modes more concrete. T 1 3 . - m L (L xi)2 2 i The potential energy is made up of two parts, that due to the elevation of the masses in the field of gravitation and that which corresponds to the deformation of the springs 3 V m g L L (1 - cos xi) + 1 k lid sin Xl - d sin X2)2 + (d sin X2 - d sin xJ)2] - 209 - !L (QL) + Q.'L 0 i=1,2,3 dt 8xi 8xi one differentiates and then adopts the small angle hypothesis (sin x cos x ~ 1). We get in this way: r m L2 + Xl (m 9 L + k d2) - X2 k d2 X2 m L2 - Xl k d2 + X2 (m gL+2kd2) - XJ k d2 XJ m L2 - X2 k d2 + xJ(m g L + k d2) By comparing these equations with the matrix equation "
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0003887_978-3-642-73890-6-Figure3-2-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0003887_978-3-642-73890-6-Figure3-2-1.png",
+        "caption": "Figure 3-2: II nocmal stresses in flat plate",
+        "texts": [
+            " The actuation causal relationship of a program synchronization point is a directed acyclic AND/OR graph rooted at that point as in Figure 3-1. Physical process PP2 needs to be actuated by both two signals - one is set up by computer process CPI at state SI, and another is set by CP2 at S2. Physical process PP5 needs to be actuated by either PP4 or PP3. Computer process CP at S is waiting for a precondition which needs be set up by CP3 at S3 and PPI, or only by PP5 at certain state. Any solution graph [Nilsson 801 in the AND/OR graph will be an actuation causal chain shown in Figure 3-2. CPI, CP2, and CP3 are called actuating processes of CP at S. If process CP is stuck at S, then CP is called a waiting process and the actuation causal chain is called a stuck chain of CP=S. A stuck chain of CP=S is called an actuated stuck chain of CP=S if all actuating processes of CP at S have activated the physical processes when CP is stuck at S. 4. Diagnostic Reasoning about Robotic Cells 4.1. Diagnosis from Actuation Causal Chains The principle for diagnosis from actuation causal chains is to check whether or not stuck chains are actuated"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0000767_aim.2016.7576819-Figure1-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0000767_aim.2016.7576819-Figure1-1.png",
+        "caption": "Fig. 1: Quadrotor MAV platform views",
+        "texts": [
+            " IV; 4) control performance are analyzed with flight tests results for real scenarios in Section. V and the last section is the conclusion part. A quadrotor platform with gross take-off weight around 200 g with onboard avionics, height detection sensor, infrared obstacle sensors, visual sensors (downward facing and front facing camera) as well as upper-level onboard computer, was developed. The largest dimension, diagonal distance from propeller tip to tip, is 38 cm. With a 1300 mAh LithiumPolymer (Li-Po) battery, flight endurance is up to 15 min. Fig. 1(a) gives the overview of the MAV platform shown in 3D CAD model, where all the electronics, sensors and other parts are imported with their real dimensions and material properties. Detailed design and presentation of all the components in SolidWorks in advance gives a reference to the real system. Force and frequency simulation were done to ensure the reliability of the structure to avoid resonant 978-1-5090-2065-2/16/$31.00 \u00a92016 IEEE 513 mode during flight. Fig. 1(b) gives the real view of the assembled quadrotor MAV. For the onboard electronics, the key components are: 1) flight controller module, which mainly performs attitude loop and position loop control along with the hardware peripheral management such as remote control (RC) input, sensing input, servo output, status display and flight data logging; 2) electrical speed controllers (ESCs), which receive the servo outputs and perform 3-phase brushless motor timing control; 3) optical flow module, which takes in images and other sensor information and estimates velocity in body frame; 4) upper computer and a forward-facing camera for off-board video transmission and motion planning",
+            " This maps variations to change rotors\u2019 angular velocity as: (\u03c9\u03032) = \u0393 \u22121u (30) where (\u03c9\u03032) = [ \u03c92 1 \u03c92 2 \u03c92 3 \u03c92 4 ] and the new motor input can be solved with the relationship between PWM input and the rotor RPM. Fig. 4 gives the step response of the CNF controller using the Matlab Simulink Toolkit by [20] compared to the response by a pole-placement controller. It can be seen from the response that the CNF controller provides faster tracking performance with small overshoot due to the compensation of the nonlinear terms while with the normal pole-placement controller the settling time is relatively longer even the control effort is large. The flight tests were conducted with the platform shown in Fig. 1(b) with manual control in lab environment. The manual pilot gave the perturbing reference in each channel with the highest perturbing angle up to 30\u25e6. The controller designed based on Sec. IV is as follows: v1 = [ \u221219.11 \u2212191.19 \u221211.07 ] x\u0304 +\u03c1(e) [ 0.30 10.46 1.04 ] x\u0304 (31) where \u03c1(e) = \u221210.3e\u2212|eBx|. The linear component of gain sets the closed-loop damping ratio as \u03b6 = 0.4, the natural frequency as \u03c9n = 13.82 rad/s (with the same and \u03c9n and \u03b6 = 0.8 for the pole-placement implementation). The damping ratio of CNF linear component is smaller to improve the tracking speed when tracking error is large"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0003898_oceans.1998.724374-Figure2-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0003898_oceans.1998.724374-Figure2-1.png",
+        "caption": "Figure 2 : the robot, 5 controlled variables",
+        "texts": [
+            "025 Kg m2 1 Volumic mass of water 1025 Kg/m3 1 rd/s I11 - CASE STUDY Second link 8.04 Kg 0.05 m l m 0.134 Kg 2.41 Kg 0.025 Kg m2 1 Viscosity coeff. of water 1.56.10-6 m2/s 1 rd/s We have chosen to develop a simulator to test this control law under MATLAB-SIMULINKTM. Some simplifymg hypothesis were necessary. Characteristics of simulation The robot is a spherical platform with a manipulator constituted by two cylindrical elements. It moves in the (X,Y) plan. We consider that the five state variables are controlled independently (Cf. Figure 2). Geometry Mass Radius length Added Mass Mx MY Iz Drag coeff. Environmt Limitation velocities of joint Vehicle 523.6 Kg 0.5 m 261.8 Kg 261.8 Kg OKgm2 0.4 Gravitational acceleration 10 d s 2 E ible 3 : charact Modelisation The Lagrange method has been used to develop the dynamic model of the robot. We have chosen to consider only three hydrodynamic phenomena : added mass, drag and buoyancy. The calculation of drag forces acting on the elements of the arm have to be integrative along the length of the links [KIE 961"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0003085_1.3267409-Figure10-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0003085_1.3267409-Figure10-1.png",
+        "caption": "Fig. 10 Driving clevis, coupler, and driven clevis",
+        "texts": [
+            "org/about-asme/terms-of-use Fig.8 Oldham\u00b7coupllng type balancer assembly Fig.9 Exploded view 01 the balancer (5) 53.6 mm by 53.6 mm square and 26.1 mm thick and is made of a leaded-bronze bearing material which has a specific gravity of 9.3. This results in a balance mass of 698 g for each coupler. The offset distance between the driving and driven clevises was chosen to be 14 mm. Therefore, the balancing force generated by the two couplers will be Fy = 8 x 0.698 X (0.01412)w2 cos2wt =0.0391w2 cos2wt N Figure 10 shows details of the driving clevis, coupler block and driven clevis, respectively. The driving and driven clevises are each supported on two cylindrical journal bearings that are approximately 25 mm diameter by 21 mm long and 55.2 mm apart (Fig. 9). The clevis shafts are all 25 mm in diameter. Each coupling is lubricated separately. The journal bearings 288/ Vol. 106, SEPTEMBER 1984 which support the shaft of each clevis are end fed by sup plying lubricant to the housing midway between the bearings",
+            "4 mm wide, were made on the unloaded sides of the shaft to direct the lubricant to the supporting journal bearings. The coupler block is also cross drilled to ensure that lubricant is supplied to all of the sliding surfaces. Cone-shaped pickup holes are provided at the en trances of the cross-drilled holes in the coupler block, and on the bottom surfaces of the clevises. These holes are located so that the oil passage in the coupler block will always be con nected to either or both the driving or driven clevis as the coupler block slides between them. Figure 10 illustrates the lubricating passages in the parts. The diametral clearances Transactions of the ASM E Downloaded From: http://mechanicaldesign.asmedigitalcollection.asme.org/pdfaccess.ashx?url=/data/journals/jmdedb/28043/ on 06/07/2017 Terms of Use: http://www.asme.org/about-asme/terms-of-use \u2022 \u2022 Vibtation Occurred,\"Th'''7r 20 ,----------------------, 15 20.010.0 Time - ms 0.0 2.0 -2.0 -4.0 f----'>..L,----,----,--,-----.-----'--,-----,--,.--.--------J 4.0 Fig. 11 Balancing force at 3000 rlmln between the shafts of the clevises and journal bearings are in the range of 0"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0000559_978-3-319-44156-6_24-Figure4-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0000559_978-3-319-44156-6_24-Figure4-1.png",
+        "caption": "Fig. 4 Fixed and moving centroids drawn with WinMecC",
+        "texts": [
+            " Nearly all the variables, whose numeric results can be displayed, can also be represented graphically as a point, vector or curve among others. The user has the option to switch on or off the display of a series of layers which contain the selected variables such as position, velocity and acceleration of points and links, instant rotation centres, Kennedy\u2019s lines, fixed and moving centroids, point trajectories, hodographs, forces between links and so on. To show the potential of WinMecC, we have selected three educational examples. Figure 4 shows an example of a fixed and moving centroid of a double slider mechanism. The instant centre of rotation of link 3 with respect to link 1 (IRC13) is the point of tangency of the fixed and moving centroid. Therefore, the centroids are tangent at point IRC13 and movement of link 3 can be understood as a pure rotation of the moving centroid about the fixed one. The second example displays how WinMecC represents the reaction force on the frame of a slider crank linkage with and without adding balancing masses to the crank"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0002199_hlca.19920750107-Figure5-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0002199_hlca.19920750107-Figure5-1.png",
+        "caption": "Fig. 5. The electron-trap niode of action: a ) Pseudo-spherical arrangement of the outer viologen subunits (V) around the central [Co(corrinato)]entity (Co) in 7b forcing the interaction with an electrode surface to be always ES-V-Co, i.e. independent of rotation of7b. b ) Fixed redox potentials of the inner (Co) and outer (V) electroactive subunits in 76. Intramolecular and heterogeneous electron transfers: Co\"-V:+ in solution with the tunable clcctrode potential E = El = +0.5 V (activation barrier: 710 mV); Co'\"-Vl+ in solution with E = E2 = -0.65 V. c ) Electrode potential dependent rate of all-over Co'\" reduction and Co\" oxidation of a solution of 7b with [Co\"'-V:+] = [Co\"-V:+]; I: V-catalyzed reduction; 11: reduction with potential-dependent activation barrier; 111: oxidation with potential-independent barrier",
+        "texts": [
+            "8 The portion of Col was not determined. t [min] Fig. 4. Concentration vs. time behavior upon potential step showing slow kinetics of Co\" oxidation in 7b us compared to 6c. Potential step -0.2++0.5 V for A : 7b; 0 : 6c/8 1 :5 . 1 electrode independent from each other according to the rules of thermodynamics and kinetics. However, in the (pentaviologen-corrinato)cobalt(II) 7b, only the viologen subunits come into close contact with the electrode, provided that the space-demanding Vf' subunits envelop the corrin moiety tightly (Fig. 5 ) . According to this model, heterogeneous electron transfer to or from Co has to involve a viologen subunit as a 'bridge'. A lower limit for the activation barrier of 710 mV (16.4 kcal/mol) for oxidation results from the relative redox levels of the central Co\" and the outer sphere (V++) retarding the Co\" oxidation dramatically. This barrier is expected to be independent of the electrode potential, and absolute-rate theory predicts a rate constant in the range of 0.1 to 10 s-' (depending on the choice of the preexponentional factor), i"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0002200_s0956-5663(97)00120-6-Figure1-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0002200_s0956-5663(97)00120-6-Figure1-1.png",
+        "caption": "Fig. 1. Structure of the PVC membrane-based ammonium ion sensor with microbial membrane. 1. Terminal, 2. Sensor cap, 3. PTFE body, 4. Reference electrode, 5. Sensor chip, 6. Plasticized PVC membrane, 7. Microbial membrane, 8. Supporter for microbial membrane, 9. Ammonium chloride inner solution, 10. Inner chloride ion electrode, 11. Saturated potassium chloride paste.",
+        "texts": [
+            " 10 ml of the THF solution was added dropwise onto a porous PTFE membrane which was placed on an end of a sensor chip (DKK, Sensor Kit) and the THF was then allowed to evaporate. This procedure was repeated a total of 30 times. Finally, the sensor chip was maintained at 30\u00b0C for 12 h, giving the plasticized PVC membrane on the sensor chip. The concentration of nonactin in the plasticized PVC membrane thus obtained is approximately 2.0 \u00d7 1022 mol l21 of plasticizer. The thickness of the membrane is ca. 0.5 mm. The structure of the ammonium ion sensor, along with the microbial membrane is shown in Fig. 1. The sensor chip with the PVC membrane was filled with 3 ml of 0.01 mol l21 ammonium chloride solution as the inner filling solution. The sensor chip was then screwed into a sensor body consisting of an inner chloride ion electrode and a reference electrode with the porous PTFE liquid junction (Ito et al., 1996). The microbial membrane was fixed on the surface of the PVC membrane so that the two were in close contact. When the ammonium ion sensor with the microbial membrane is immersed in a sample solution containing ammonium ion and organic compounds, organic compounds, which diffuse into the microbial membrane are assimilated by the microorganism in the membrane and are then decomposed"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0001235_b978-0-08-099425-3.00005-4-Figure5.89-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0001235_b978-0-08-099425-3.00005-4-Figure5.89-1.png",
+        "caption": "FIGURE 5.89",
+        "texts": [
+            " The model of the tyre test machine presented here contains a tyre part that rolls forward on a flat uniform road surface in the same way that the tyre interacts with a Overview of the tyre modelling system. Orientation of tyre coordinate systems on the full vehicle model. moving belt in the actual machine. In this model the road is considered fixed as opposed to the machine where the belt represents a moving road surface and the tyre is stationary; modelling a moving belt is surprisingly awkward in a MBS environment. Considering the system schematic of the model shown in Figure 5.89, the tyre part 02 is connected to a carrier part 03 by a revolute joint aligned with the spin axis of the wheel. The carrier part 03 is connected to another carrier part 04 by a Applied force equal to required wheel load Mechanism sketch for a flat bed tyre test machine model. revolute joint that is aligned with the direction of travel of the vehicle. A motion input applied at this joint is used to set the required camber angle during the simulation of the test process. The carrier part 04 is connected to a sliding carrier part 05 by a cylindrical joint that is aligned in a vertical direction"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0002258_1.1332396-Figure3-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0002258_1.1332396-Figure3-1.png",
+        "caption": "Fig. 3 Fluid film coordinate system",
+        "texts": [
+            " Other simulations are presented by Olson @11# where the effects of structural inertia are also important, such as in an uncavitated bearing, a bearing with a rapidly varying applied load, two bearing derivatives of a gas engine main bearing, and rigid bearing stability analyses that identify operating regimes where self-excited oscillations ~instability! occur. 001 by ASME Transactions of the ASME s of Use: http://www.asme.org/about-asme/terms-of-use Downloaded F The physical journal bearing model is shown in Fig. 1, and the appropriate inertial Cartesian coordinate system is shown in Fig. 2. 2.1 Equation of Motion for Elastic Sleeve. The general equation for an undamped, discretized elastic solid is given by Cook @13# @M S#$d\u0308 S%1@KS#$dS%5$rS%. (1) An important assumption made is that transverse direction ~x and y in Fig. 3! tractions on the fluid surface can be ignored. Consequently, all degrees of freedom in Eq. ~1! that are not directly associated with the normal direction ~z in Fig. 3! on the fluid surface can be removed through static condensation ~for @KS#! and dynamic ~Guyan! reduction ~for @M S#!, leaving a much smaller set of coupled equations. 2.1.1 The Condensed Form. After condensation/reduction, the only degrees of freedom left from Eq. ~1! are those associated with the nodes on the surface contacting the fluid in the normal direction. The condensed form is @M #$d\u0308 %1@K#$d %5$r%. (2) rom: http://tribology.asmedigitalcollection.asme.org/ on 08/08/2017 Term 2.1.2 The Modal Form",
+            " requires that the z-displacements be known for the journal at every node of the fluid mesh, so the relationship between the generalized rigid journal displacements $e% and the corresponding displacements at the fluid film nodes $d j% must be known and can be related through the static equilibrium matrix so that $d j%[@G#$e%. JULY 2001, Vol. 123 \u00d5 463 s of Use: http://www.asme.org/about-asme/terms-of-use Downloaded F The finite element form of the fluid model is well developed and its full derivation is given by Booker and Huebner @14# for the 3-noded triangular elements used throughout this work. The appropriate fluid film coordinate system for the current application to journal bearings is shown in Fig. 3. The formulation employs the usual assumptions used to reduce the Navier-Stokes equations to the Reynolds equation \u00b9\"S %h3 12m \u00b9r D5\u00b9\"~%hU\u0304!1 ] ]t ~%h ! with appropriate boundary conditions. In its finite element, discretized form the solution to the Reynolds equation comes from the n coupled equations $q%5@Kp#$p%1$Q%. (6) With @Kp# and $Q% specified for a complete oil film, Eq. ~6! can be solved as a standard finite element problem with mixed boundary conditions ~essential boundary conditions within $p% and natural boundary conditions within $q%"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0003812_iemdc.1997.604176-Figure5-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0003812_iemdc.1997.604176-Figure5-1.png",
+        "caption": "Figure 5: effect of a non-zero voltage application",
+        "texts": [],
+        "surrounding_texts": [
+            "The basic idea of DTC is to choose the best vector of voltage which makes the flux rotate and produce the desired torque. During this rotation the amplitude of the flux rests in a pre-defined band. Thus it is almost constant. With a three phase voltage source inverter, there are six non-zero voltage vectors and two zero voltage vectors whch can be applied to the machine terminals. The stator flux vector can be estimated using measured current and voltage vectors: YS = 5 (Vs - RIs)dt (1) Then torque can be calculated using the components of the estimated flux and measured currents: I- = P W J p -Yp-la) (2) P is the pole pair and the a and p indiced parameters are the Concordia transformation components of the current and flux. As it can be seen, using the equations (1) and (2) to estimate flux and torque, R is the only parameter of machine which is needed to estimate the torque and the flux. To use the switching table, the zone where the flux is situated should be known. There are two two-level hysteresis controllers for the flux and the torque. 0-7803-3946-0/97/$10.00 Q 1997 IEEE. MC34.1 Figure 2 shows the switching table for a two quadrant system (o>O). It can be seen that Vi+, is chosen to increase the torque and flux, while flux is situated in the zone i, and Vi+z to increase the torque without increasing the flux. To decrease the torque, a null vector is applied. The zones, the voltage vectors and the trajectory of the flux are shown in Figure 3. III. ORIGIN OF THE STEADY STATE TORQUE ERROR Using the vectorial model of the machine (Figure 4) the r = p - YsK sin(y ) torque equation can be written as: (3) 1 Ls where Ls is the stator inductance of the machine. Referring to the equation 1, \u2018U, can be controlled by applying the voltage vectors. AOs is a function of the DC inverter input voltage and the initial position of Y, in the zone. A0 is a function of the mechanical speed which can be considered constant during one sampling period. Thus, at low speeds, applying a non-zero voltage vector causes a big Ay, which means an important change of the torque. On the other hand, while the zero vectors are applied to decrease the torque, AOs is approximately zero and Ay equals -AO. Then the torque changes slowly. ~ MC3-4.2 At high speeds, due to big AO, applying a non-zero voltage can\u2019t produce important Ay. Thus, the torque changes are very slow. But applying the zero voltage can produce important changes in the torque. These phenomena in addition to the torque hysteresis controller cause an error in the steady state torque of machine which depends on the speed. As it can be seen in the Figure 7, at high speeds the output torque is less than reference torque (positive error) and at low speeds it is more than the reference (negative error). Figure 8 shows the error as a function of the load for different speeds while the sampling frequency is 50 psec and the torque hysteresis band is 0.05 Nm for all cases. It can be seen that it is almost load independent and a function of the speed. Increasing the sampling fiequency decreases the torque error. But the switching frequency is limited, and the control system needs some data processing time. Thus, the sampling frequency can\u2019t be more than some values. As it can be seen in Figure 8, increasing the sampling frequency decreases considerably the output torque error. IV. MODIFIED TORQUE CONTROLLER To correct the steady state torque the torque reference which will be applied to the direct torque control system (r*ref ) is modified so that the output torque, and in consequence the estimated torque (rea), be as near as possible to the input torque reference (Tref). To have a zero steady state error, the torque error passes through an integrator. The inner loop stabilizes the integrator while its output can be limited by the limiter. As it can be seen in Figure 10 and 11, the torque error for different speeds and sampling frequencies are considerably less than those without modified controller. In addition we can see that even with a lower sampling frequency the steady state error of the torque is still much less than that of the system without modified torque reference. So it seems logical to use the modified torque controller to have more precise torque control system Even if, because of the added calculations, the sampling fiequency should be decreased. On the other hand, the switching frequency can be also decreased and this can be one of the most important advantages of this controller."
+        ]
+    },
+    {
+        "image_filename": "designv11_60_0002396_ultsym.1994.401648-Figure1-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0002396_ultsym.1994.401648-Figure1-1.png",
+        "caption": "Fig. 1 Schematic shape of a stator vibrator for a micro ultrasonic motor.",
+        "texts": [
+            " The high linearity suggests possibility of the low mechanical vibration loss. The deposit process is carried out in high pressure liquid, s o that the base material shape is free. Fortunately, 1994 ULTRASONICS SYMPOSIUM - 549 the substrate material is limited to titanium or titanium oxide. Titanium possesses high tensile strength and high mechanical Q factor. Hence, this substrate material is suitable for a stator transducer. We have fabricated a cylindrical shape stator transducer as shown in Fig.1. The base material is titanium tube and the PZT film is deposited on the side wall. The polarization of the PZT film is the thickness direction. By the transverse effect, the bending vibration of the tube is excited. As same as the other mode rotation type ultrasonic motors, the vibration of the stator transducer is transformed to the rotation motion through the friction force as shown in Fig.2. The rotor turns opposite direction by alternating the mode rotation direction. FABRICATION OF STATOR T ANSDUCER Fabrication process of the stator transducer has four steps"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0002608_1.2833766-Figure2-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0002608_1.2833766-Figure2-1.png",
+        "caption": "Fig. 2 Computation domain and finite difference grid",
+        "texts": [
+            "10 Density-Temperature-Pressure Relation. The lubricant density was determined by the following equation: p, = po[l + G p / ( 1 + Csp) - DAT - To)] (10) where CA, CB and DT are the lubricant dependent constants. 3 Computation Procedure As stated earlier, previous studies show that under heavy loads {po > 1 GPa), the pressure distribution in the EHD contact is close to Hertzian. Hence, the solution domain adopted in this study was similar to the Hertzian contact area, but the inlet and exit boundaries were extended, as shown in Fig. 2(a). Their locations were defined, respectively, by Xm = -a,rJl - y^ (1 s flin ^ 1.2), and JE; = Ue^l'^-y^ with 1 < a\u0302 < 1.12. As the problem was symmetrical about the minor axis {x) of the con tact ellipse, the computation was carried out only in half of the solution domain, as illustrated by the shaded area in Fig. 2 (a ) . The locations of inlet and outlet boundaries of the solution domain or the values of fli\u201e and a^ vary slightly with different operating conditions. Hence, for the convenience of computa tion a so-called \"computational region\" bounded by j r= \u20141.2 and X = 1.12, and y = -\\ and f = Q was used. The computa tional region was divided into A'\u0302 and A', equal spaces in the Xand y-directions, respectively. Across the film thickness, N^ equal spaces were used, as shown in Fig. 2{b). The dimensionless lengths of the grid spaces, denoted by Ax, Ay and Az, in the three directions are shown in Fig. 2(c) . 3.1 Determination of Pressures. Unlike the \"conven tional\" EHL analyses in which the film pressures are obtained by solving the Reynolds equation, the present model employs an empirical approach to determine the film pressures. Consider ing that the pressure distribution in the EHD contact under heavy loads is similar to Hertzian, the pressures in the solution Journal of Tribology OCTOBER 1998, Vol. 120 / 687 Downloaded From: http://tribology.asmedigitalcollection.asme.org/pdfaccess.ashx?url=/data/journals/jotre9/28678/ on 05/03/2017 Terms of Use: http://www",
+            " Details of the approach have been presented elsewhere (Ma, 1997), but a brief summary is given here. (a) In the inlet region of the conjunction, the pressure was approximated by a power-law function which satisfies the zero-pressure and zero-pressure-gradient condition at the inlet boundary. (b) Over the central region, the pressure was determined from the Hertzian distribution. (c) In the outlet region, the pressure was described by a fifth-order polynomial with a set of boundary condi tions being imposed. For example, at the exit boundary (Xe), as shown in Fig. 2(a), the Reynolds boundary condition was adopted. Note that in the previous publi cation by the author (Ma, 1997), the outlet boundary coincided with the Hertzian circle in dimensionless co ordinates. In the present work, however, this boundary was extended. (d) After the corrections in the inlet and outlet regions, if necessary, the pressure was further modified in the vicinity of the two leakage sides. The pressure was represented by a third-order polynomial. Again, a set of boundary conditions had to be implemented in order to determine the coefficients of the polynomial"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0000510_htj.21229-Figure7-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0000510_htj.21229-Figure7-1.png",
+        "caption": "Fig. 7. Cooling channel structure after optimization.",
+        "texts": [
+            " (2) Change the direction of cooling channels A and B, and enlarge the size close to the spindle bearing and motor in the Z direction, to ensure more coolant carries heat from the spindle box at the high-temperature positions. (3) The layout of cooling channels A and B in the spindle box is more complicated. For convenience in sealing, change the width of the cooling channel from 20 mm to 15 mm; change the depth of the cooling channel from 4 mm to 5 mm to increase the contact area between coolant and spindle box. The structure of the cooling channel after optimization is shown in Fig. 7. Recalculating the forced convective cooling capacity of the cooling channel, conducting a simulation on the thermal characteristics of the machine tool, a result comparison of temperature field and thermal deformation simulation before and after optimization is shown in Tables 6 and 7. The temperature rise improvement of test point T26 that is close to the cooling channel is the most obvious. Due to the increased size in the X, Y, and Z directions, the thermal deformation in the Y and Z directions is significantly decreased"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0002594_bf00046883-Figure4-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0002594_bf00046883-Figure4-1.png",
+        "caption": "Fig. 4. Body axode of a moving rigid body generated by inertia ellipsoid rolling on invariable plane.",
+        "texts": [
+            " Such a motion is clearly nongeneric. Suppose on the other hand that /z has nonzero rotation (i.e., nonzero angular momentum). If we ignore for the time being the motion of the centre of mass, Poinsot's description of the motion about the centre of mass is of the inertial ellipsoid E, in the body, rolling on an invariable plane ~r in space, which is orthogonal to the angular momentum vector m, (see [15]). In effect, this introduces a simple mechanism of two components, E and ~', whose motion in contact reproduces the motion (see Figure 4). The curves of contact are known as the herpolhode (in the body and, hence, in E) and polhode (in space and, hence, in ~r). Since points on these curves are instantaneously stationary at the moment of contact, they must lie on the body and space axodes, respectively. Thus, the herpolhode and polhode define directrices for the axodes. The centre of mass is stationary, so it also lies on every generator of the axodes. Since the helpolhode lies in the body it is independent of the motion of the centre of mass. Now, by the conservation of energy, the herpolhode lies in the intersection of the inertia ellipsoid a n d a sphere (surface of constant energy) centred at 0, and the body axode is the cone of this curve over 0, as depicted in Figure 4. We have seen that the herpolhode has a singularity precisely when the axode is cylindrical. Generically, the intersection is an antipodal pair of simple closed curves, as shown in Figure 4. But when the radius of the energy sphere equals one of the principal semi-axes of the inertia ellipsoid the corresponding principal axis acts as a constant axis of rotation of the motion, called a stationary motion. (Such a motion associated with the middle moment of inertia is dynamically unstable, while the others are stable.) In the case of stationary motion Kl ' l /x = R - the axodes are simply a line in the body and a plane in space, respectively. We have established that motions of zero angular momentum and stationary motions of bodies with nonzero angular momentum are nongeneric"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0001768_a:1008839925389-Figure2-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0001768_a:1008839925389-Figure2-1.png",
+        "caption": "Figure 2. These are the leg muscles that were modeled on each leg.",
+        "texts": [
+            " This five-link biped is controlled by sixteen muscle-like actuators which model the muscles in a human being. The actuators are driven by a neural input, R, which is analogous to neural inputs supplied by the brain in humans. For the model to be physiologically valid, both R and F have to be positive. For the values chosen, when R is positive, F is positive. Constraints are added in the computation of the forces from the torques to guarantee that F and R are positive. The eight leg muscles shown in Fig. 2 are included indirectly by noting the effect of the torque on each corresponding limb, E\u03c4 = \u2212\u2202LT \u2202W F (6) where \u2202L \u2202W is a 16 \u00d7 7 matrix (Table 2) containing the moment arm for each muscle at each joint and F is the vector of muscle forces applied. For simplicity, we approximate each moment arm as constant for different joint angles. Including the P1: KCU/RKB P2: STR/SRK P3: STR/SRK QC: Autonomous Robots KL465-03 July 1, 1997 15:43 A Control Strategy for Terrain Adaptive Bipedal Locomotion 247 individual muscle forces, the equation of motion for the system becomes J (W )W\u0308 + B(W, W\u0307 )W\u0307 + G(W ) = \u2202C \u2202W T 0 \u2212 \u2202LT \u2202W F (7) The musculoskeletal system, as formulated above, is shown in Fig"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0003887_978-3-642-73890-6-Figure7.1-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0003887_978-3-642-73890-6-Figure7.1-1.png",
+        "caption": "Fig. 7.1 Micropolar Elastic Cylinder with Semi-Circular Groove",
+        "texts": [
+            " Analytical solutions are given by Gauthier [8), Force Stress: Pz Tzz = - A r-displacement: z-displacement: \u00b7\u00b7\"\u00b7p,Z v= -- e\",A where A is a cross-sectional area and Young's modulus and Poisson's ratio are defined by E\", = (2fL+ K)(3A +2JL+ K) (2JL+2A+K) 150 Both FEM and analytical solution gave the exactly same numbers: Displacemerit: u = - 0.34615386x10\u00b73 (inch), v = 0.11538462xlO\u00b72 (inch), Stress: T .. = 1.0 (psi), T\" = l.Ox1O-IS (psi) The same problem was solved using 4-node and 8-node axisymmetric elements and exactly the same values were obtained. (3) Micropolar Easltic Cylinder with a Semi-Circular Groove An isotropic micropolar elastic cylinder with a semi-circular groove is shown in Fig. 7.1. The diameter of a cylinder is fixed to 0.1 inch. The radius ratio Ie is defined by Ie = ~, where r is the radius of groove, and d = D - 2r. Ie is varied in the range of 0.05 to 0.5. Finite element mesh used is shown in Fig. 7.2 for radius ratio Ie = 0.5. The material properties used for isotropic case are shown in Table 7.1, with characteristic length: l = 8.3333x1O-3 and Poisson's ratio = 0.3. Numerical results of stress concentration factor, K., is plotted in Fig. 7.3. for Isotropic Micropolar Elastic Cylinder with Semi-Circular Groove The same model was used to solve for orthotropic material cases"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0001522_pi-c.1955.0020-Figure1-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0001522_pi-c.1955.0020-Figure1-1.png",
+        "caption": "Fig. 1.\u2014Primitive machine with axes fixed to windings.",
+        "texts": [
+            " This equation is very suitable for certain classes of transformations of co-ordinates, but it has been found that under the conditions of transformation obtaining in electrical machines a modified form of Lagrange's equation must be used. The modified equation was developed by Boltzmann and Hamel13 to cover certain conditions of constraint in dynamical systems and, as Kron has shown,6 the Boltzmann-Hamel equation can be used to form a basis for tensor analysis of electrical machines from the dynamical viewpoint. (2) MACHINE EQUATIONS The first type of primitive electrical machine to be considered is shown in Fig. 1. The rotor is assumed to be smooth and to have on it a symmetrical 2-phase winding sinusoidally distributed. The field is fixed in space and consists of windings ds and qs in the stator\u2014direct and quadrature axes respectively. Iron loss and saturation are neglected. The armature axes a' and b' are fixed to the armature and rotate with it. Threephase machines may be analysed by resolving the resultant armature current and flux vectors along two similar axes.7 The inductance of phase a' of this machine may be written:* Phase a' self-inductance = LA+ LB cos 20 LYNN: THE TENSOR EQUATIONS OF ELECTRICAL MACHINES 151 where LA = (Ldr + Lqr)l2 LB = (Ldr - Lqt)l2 Ldr and Lqr are the self-inductances of rotor phase a' when in the direct- and quadrature-axis positions respectively",
+            " cos 0 (9) where Lss is the moment of inertia of the rotor. 152 LYNN: THE TENSOR EQUATIONS OF ELECTRICAL MACHINES + [ab,s]iaib (18) to cover only the mechanical part of the range of variables,1 d1xt thus fs = Rsti* + Ls,-fiT where [ob,s] = \u2014 \u2022= Therefore /\u201e = Rtt-r + L,,-^ 2 Id (19) ddnd .. . , Rst-f- \u2014 fnctional torque r d 2 B \u2022 \u2022\u2022 \u2022 *'T2 = i n e r t i a torque and - -^ \u2014 electrical torque The second form of the primitive machine considered here is shown in Fig. 2. The rotating axes a' and b' of Fig. 1 are here Fig. 3.\u2014Primitive machine with axes rotating freely. In synchronous-machine studies the quadrature-axis stator (field) coils are omitted unless amortisseur windings are being considered. Here sinusoidal flux distribution is considered and M'd = Md, etc. Eqns. (20) may be obtained from those of the previous form of primitive machine using the relationship ,-\u00ab' \u2014 (dr c o s Q _|_ [gr s m Q (22) resolved along the direct and quadrature axes. All axes are now relatively stationary. This primitive machine has been used by Sabbagh, Stanley, Kron and others7\"11 as the basis of the analysis of many derived machines",
+            " (17), does not give these equations directly since it does not include generated voltages. The following Sections show that the connection Yuvw arises naturally because of the dynamical relationship between the two types of primitive machine, this being quasi-holonomic and non-integrable. Fig. 4 shows the form of the connection Yuvw when written as a matrix in the form of a cube, together with the arrangement of matrix multiplication leading from eqns. (26) and (27) to eqns. (23) and (24). (3) NON-HOLONOMIC TRANSFORMATIONS The currents in the armature axes a' and b ' in Fig. 1 may be resolved along \"d\" and \"q\" axes shown in Fig. 2, the relationship being /<*\u2022 = /a' cos 9 \u2014 ib> sin 9 \"1 r ' (2.8) iv = p' sin 9 + ib> cos 9 j Since the variables in Lagrange's equations are the charges, x\u00b0, eqns. (28) represent a transformation of differentials of the variables, where dxa\\dt = ia, etc. These are equations of constraint among the differentials of the variables xk, and the transformation must therefore be written dxdr = dx?' cos 9 - dx5' sin 9 \"] . y . . . (29) dxv = dxa' sin 9 + dxb' cos 9 J They obtain at a given instant and cannot be integrated to give a relationship among the charges",
+            " Both of these have been analysed by Kron, but the analysis as set out below demonstrates details of the general method of using tensor equations for this purpose. The transient equation of a simple RL series impedance may be written e = Ri + Lpi Under steady-state conditions, with sinusoidal voltage applied, the equation becomes e = Ri + jtoLi which may be obtained from the transient equation by putting P equal to jco. When a 3-phase system is being considered the instantaneous line currents and phase voltages and impedances may be resolved into Clarke components.22 In order to conform to the phase positions and direction of rotation shown in Fig. 1, the current components may be defined as follows: /*' = i(2iA - iB - 1\u00b0) . . . . (119) ia' = -T\u00a3B - i\u00b0) (120) i* ) . . . . (121) where iA, iB, ic are the instantaneous line currents (Miss Clarke uses indices a and jS instead of b' and a' as written here). Zerosequence currents i\u00b0 are those residual currents in the neutral connection to an unbalanced load or point of fault. To simplify the analysis a balanced system will be considered with no zerosequence currents. In a machine wound for three phases these instantaneous components /*' and ia> lie respectively along the axis of phase A and along the common axis of phases B and C in quadrature with phase A. These are the same as the axes b' and a' used in the holonomic primitive machine, Fig. 1, and are stationary with respect to the armature phase windings. The ar and b' components of an external 3-phase network would be connected to the machine axes as shown in Fig. 5(a). It is therefore possible to define for either a machine or stationary network a set of currents iSl and iS2 expressed along axes rotating with uniform angular velocity with respect to the axes of the Clarke components. From Fig. 3 the relationships among such currents may be written = isi c o s cos 6f (121) For a stationary network the holonomic (Lagrangian) equations, in terms of a' and b' components are a' b' ea> eb> w\\y b' b' R + Lp R+Lp a' b' i\u00bb' i*' (122) 160 LYNN: THE TENSOR EQUATIONS OF ELECTRICAL MACHINES where 2iiYj>aW\u00ab = 2Q\u00bb L6a/ a/>0 = Indirect notation C\\ = C and 2Clys = C, ~\u2014r~ Za zb \u20141 __ (a) Thus y\\8 S, Si s2 y\\a - 1 +1 (126) (127) (128) and Thus 2Q \u2014 -L +L \u2014 V = 7"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0003765_tmag.2002.802692-Figure4-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0003765_tmag.2002.802692-Figure4-1.png",
+        "caption": "Fig. 4. Pressure distribution in the infinite-width meniscus between parallel planes.",
+        "texts": [],
+        "surrounding_texts": [
+            "In the case of the finite meniscus ring between plane surfaces [see Fig. 1(b)], (1) is rewritten in the following form: (12) Through the same processes as Section III, the following important equations are obtained: (13) (14) (15) The boundary conditions are a) 0 at 0, i.e., the pressure is axisymmetric with respect to axis, b) the pressure at the boundary is given by (16) from (2). Equations (13)\u2013(15) correspond to (9)\u2013(11), respectively. Equation (13) means that the meniscus boundary varies at the ratio of when the spacing varies at the ratio of . Note that it was in the case of the infinite-width meniscus. The squeeze term, the spring term and the static Laplace pressure or the static meniscus force by the static Laplace pressure are again found in (14) and (15). Examples of the pressure distribution and the load carrying capacity are shown in Figs. 6 and 7, respectively. The parameters are the same as those of the infinite-width case (see Section III). Note that the load-carrying capacity is the net force in this case."
+        ]
+    },
+    {
+        "image_filename": "designv11_60_0002865_978-94-015-9064-8_16-Figure5-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0002865_978-94-015-9064-8_16-Figure5-1.png",
+        "caption": "Fig. 5: The Double Circular-Triangular six-DOF robot.",
+        "texts": [
+            " Each triangle's center contains a double U joint which is connected to the output link at one triangle's center through a prismatic joint and at the other through a helical joint (nut and a lead screw). Rotational motion of the output link about the line connecting the centers is achieved by rotating the first moveable triangle (the prismatic joint) while motion along the line is achieved when the triangles rotate at different angles. The structure of the six-DOF Double Circular-Triangle parallel robot is shown in Fig. 5. 4. Direct Kinematics While the solution of parallel robots inverse kinematics is usually trivial, the direct kinematic is more complicate. In fact, in most cases the solution is reduced to a solution of a high degree polynomial equation. Once the polynom is obtained it means that the kinematics is solved in a closed-form even though the high degree polynom has only a numerical solution. In the present structure a solution of the six-DOF parallel robot is obtained next in closed-form. We start the kinematic analysis by first solving the planar Circular-Triangular case"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0002252_s0890-6955(98)00006-6-Figure8-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0002252_s0890-6955(98)00006-6-Figure8-1.png",
+        "caption": "Fig. 8. The modified feed fCM, when PC is intersected by the curve CM.",
+        "texts": [
+            " In the first case, wherein the straight line PC is intersected by curve , feedrate (f) can be divided into another two cases, depending on the cutting condition, as: 1. when condition f . fc occurs 2. when condition f # fc occurs The present work only concerns the case (i), i.e. f . fc. future work is intended to extend the present analysis to case (ii). The real cutting conditions are shown in Figs 8 and 9. For simplification of the calculations, the simplified model is shown in Fig. 10, in which the actual feed in Figs 8 and 9 have been modified. The length of WW9 can be found from Fig. 8 as: WW9 = ll/sin(uPC + Ce 2 Cs) (17) ll = (l1 + l2 + l3 + l4)sin(Ce 2 Cs) + (h1 + h2)cos(Ce 2 Cs) + R2cos2uR2 2 R3cos2uR3\u00b7cos(Ce 2 Cs) 2 fsinCe 2 R1\u00b7[1 + sin(Ce 2 Cs) 2 2uR1] (18) The modified feed fCM is calculated as: f CM = fcosCs + R1(1 2 cos2uR1) + WW9cosuPC (19) If WW9 , 0 i.e. the condition as shown in Fig. 9, the intersection angle uSS can be calculated as: uSS = Ce 2 Cs + sin21(mm/R1) (20) where mm = (l1 + l2 + l3 + l4)sin(Ce 2 Cs) + (h1 + h2)cos(Ce 2 Cs) + (R2cos2uR2 2 R3cos2uR3)\u00b7cos(Ce 2 Cs) 2 fsinCe 2 R1sin(Ce 2 Cs) (21) Thus yielding modified feed as: f CM = fcosCs + R1(1 2 cos2uss) + WW9cosuPC (22) After the modified feed fCM is obtained, the shear plane area and the projected area on the tool face can be calculated from Figs 4 and 10, as: A = A1 + A2 + A3 + A4 + A5 + As (23) where A1 = 0"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0003666_aim.2001.936770-Figure3-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0003666_aim.2001.936770-Figure3-1.png",
+        "caption": "Fig. 3. Two consecutive scan windows for corner detection",
+        "texts": [
+            " (6) Finally, this single measurement is corrected to account for the deformation of the tool and the robot under the current contact forces. This correction is based on a linear spring model with stiffness kcam. It shifts the contour position to the Corrected Offset Contour absHCoC according to B. Corner detection The corner detection algorithm is based on the first step of the contour measurement. If there is a sudden jump in coordinates of the extracted contour points [X,Y,] or if there is a faulty scan due to the absence of an edge, the scan window is shifted and rotated as shown in figure 3. The image is scanned again. If there are no errors (sudden jump, no edge . . . ) in the new contour measurement, its orientation is compared to the orientation of the previously logged contour. A corner is identified if the difference in orientation exceeds a given threshold. The position of tvhe corner then easily follows from the intersection of two lines. The exact location of the corner, however, is updated afterwards, when the optical axis of the vision systems is again positioned over the contour (as it was before the corner occurred) in which case the contour measurement is more accurate"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0003465_bf01242890-Figure1-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0003465_bf01242890-Figure1-1.png",
+        "caption": "Fig. 1. Diagram of the apparatus used for the determination of aluminium with a bio-membrane based fluorescence sensor: 1 light source; 2 detector; 3 bifurcated fibre optic; 4 injection syringe; 5 light-tight housing; 6 biomembrane with immobilized morin; 7 magnetic stirrer",
+        "texts": [
+            " However, no bio-membrane has been reported so far in the literature for the preparation of fibre optic. In the present communication, a mut ton membrane is suggested as support to make fluorogenic sensor for aluminium determination with satisfactory performance. * To whom correspondence should be addressed A multi-function optic-fibre spectrophotometer (Model 8510, Jiangsu Electroanalytical Instrument Factory), a filter fluorometer (Model F-1A, Xiangdai Electronic Factory) and a spectrophotometer (Model 751, Shanghai Analytical Instrument Factory) were used. Fig. 1 shows the experimental assembly. A 7 mM solution of morin in acetone and 0.01 M solution of aluminium potassium bisulfate were prepared. 4~o solution of formaldehyde was used as immobilizing agent. Buffer solution of pH 4.8 was composed of sodium acetate and acetic acid [9]. All chemicals used were of analytical reagent grade, Doubly distilled water was used throughout. Thin translucent membrane surrounding muscle bundles was peeled off from flesh mutton, Morin Immobilization Procedure and Preparation of the Bio-Membrane Immerse a freshly stripped mutton membrane in a 4~o solution of formaldehyde for ca"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0002530_881621-Figure1-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0002530_881621-Figure1-1.png",
+        "caption": "Fig. 1 Measuring device for piston friction force and piston lateral motion",
+        "texts": [
+            " Based on these experimental results, the relationship between the clearance and the friction force of the piston under engine operating conditions was clarified. Furthermore the causes of the friction increase were disclosed. Also presented in this paper are results of the study of metal-to-metal contact which was found to occure locally on the piston skirt, reduction of the friction force by modifying the skirt profile, and influences of the piston clearance on the lubrication condition. MEASUREMENT INSTRUMENTS MEASUREMENT DEVICE FOR PISTON FRICTION AND PISTON LATERAL MOTION - Figure 1 shows a crosssectional view of the test engine and the measurement device used for the present experiment. The test engine was a single- Numbers in parentheses designate references at end of paper. cylinder gasoline engine. For the piston friction measurement, its cylinder block was modified to floating liner structure. The liner was supported by three piezo-type pick-ups, by which the liner support became reliable. By increasing block rigidity and minimizing its deformation, high measurement accuracy was obtained"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0002235_0924-0136(95)02159-0-Figure1-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0002235_0924-0136(95)02159-0-Figure1-1.png",
+        "caption": "Fig. 1. A scheme of thread rolling with helical-turn rollers and crossed axes.",
+        "texts": [
+            " [1,2] does not provide a precise solution to the problem, and is applicable only to the profiling of round-turn rollers. This paper offers a general solution to the problem of the profiling of rollers with round- or screw-turns used for the rolling of threads, as well as of worms, shots, gear-wheels, or other screw or rotating profile surfaces. 2. Mathematical model The mathematical model is developed for the most general technological scheme of rolling, in which the axes of rotation of the rollers and the half-finished material are crossed at angle p, and their surfaces (1) and (2) are of screw type (Fig. 1). During the rolling process, the rollers and the half-finished material rotate simultaneously with angular speeds of O~o and o~1, respectively, and for time t, they rotate in opposite directions through angles g~o = toot and \u00a2~ = co~t. Over the same time, the half-finished material is also axially displaced by a distance s~. The same motion is carried * Fax: +359 82455145. 0924-0136/96/$15.00 \u00a9 1996 Elsevier Science S.A. All rights reserved SSDI 0924-0136(95)02159-J out also, by the introduced coordinate systems OoxoYoZo, associated with the roller, and 01x~y~z, associated with the half-finished material, towards the fixed coordinate systems Oxyz and O'x'y'z', which are translated one towards the other along the axis Ox by a distance A, referred to as the centre-line distance",
+            "l COS (0 o I% o m2 -- cos tpt sin ~ot 0 - sin ~o~ cos ~pi 0 0 and m 3 ---= li 0 0 cos/t sin/~ -s in/z cos/t are the matrices of the transitions of the coordinates from the mobile coordinate system OoxoYoZo and OlX~ytz~, towards the fixed coordinate system Oxyz, indicating their rotations at angles ~Po, ~Pl and/z. If mt -I, m~ -t and m7 t are designated the opposite matrices of ram, m2 and m3, i.e., the matrices of the opposite transitions from the immobile coordinate system Oxyz towards the mobile, after processing, Eq. (2) acquires the following form: Rou = mi-tm3m2RtM + m/- IS 1 \"11\" m l IRe, I Then, for the vector VOM: (3) VOM = ~ RoM = dl i - tm3m2RiM + mi-lm3fit2RiM + Ihl St +mi- l~t + lhiRo,~ (4) The normal vector NoM in Eq. (1) is determined from the half-finished material surface Eq. (2) (Fig. 1) defined in the coordinate system Otxty~zt, and from the matrices of the transitions towards the coordinate system OoxoYo=o: NOM = m( Im3m2Nim (5) where Nix[ t~R OR and R=R(p~,O~) is the equation of the half-finished material defined surface. After the substitution of Eqs. (4) and (5), Eq. (1) has the following form: VoMNoM = (m~tm;lmlfntm3m2RiM)NtM + (m2 -~ Im~ Im I m~-1m31il2 R ! ~,t )Ni M + (m2 Im3 tml m~- tSl)Ni M + (m2 ImP- Iml m/- 1~1 )Ni M + (m~- imp- tm I ffai- IRe; )Ns u = 0 (6) Eq. (6), together with Eq"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0002870_982639-Figure1-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0002870_982639-Figure1-1.png",
+        "caption": "Figure 1. (a) Schematic diagram of the dynamically loaded journal bearing model. (b) The",
+        "texts": [
+            " In this paper we describe the model and demonstrate the effect of viscoelasticity on the dynamics and stability of a journal bearing. Our numerical results would seem to support the experimental observations of Williamson et al [1] in that the beneficial effects of viscoelasticity only show up when the journal is up against the wall - the bearing wall, i.e., at high eccentricity ratios oils with larger relaxation times can produce larger minimum oil film thicknesses. THE GEOMETRY \u2013 Consider the two-dimensional geometry shown schematically in Figure 1(a). The journal of radius RJ rotates with a constant angular velocity \u03c9 in a stationary bearing of radius RB. Both the journal and the bearing are assumed to be of infinite extent in the axial z-direction. The distance between the axes of the journal and the bearing is given by e. The eccentricity ratio is defined by \u03b5 = e/c, where c = RB - RJ is the average gap so that 0 \u2264 \u03b5 \u2264 1. The region between the journal and the bearing, \u2126, is occupied by a lubricant. The journal is free to move under the action of an applied load which may be variable, its own weight and the reaction force exerted on it by the lubricant",
+            " We make no attempt at modelling rupture, and simply assume that there remains fluid in the cavitated region but that the viscosity quickly but smoothly decreases to an asymptotic value \u03b7min. Ideally, we would wish to choose \u03b7min to be some effective viscosity of the cavitated medium. We have chosen the value \u03b7min = 8.0 x 10-4 Pa s. The effect of this is to quickly reduce the viscosity in the cavitating region to a level somewhat above that of air. SPECTRAL ELEMENT METHOD \u2013 The region \u2126 between the journal and bearing is partitioned into a number of spectral elements, \u2126k, 1 \u2264 k \u2264 K, such that for all k \u2260 l, as shown in Figure 1(b). Let Er and Ea denote the number of spectral elements in the radial and azimuthal directions, respectively. We also assume that the decomposition is geometrically conforming in the sense that the intersection of two adjacent elements is either a common vertex or an entire edge. Each physical element is mapped onto the parent element [-1,1] x [-1,1] on which a Legendre GaussLobatto grid is used. The transfinite mapping technique of Gordon and Hall [11] is used to perform this mapping. The PN - PN-2 spectral element method (Maday et al",
+            " To ensure that large variations in the approximations do not vary across element boundaries we have implemented a fully dynamic positioning of the radial inter-elemental boundaries. This is done in such a way so that, at each timestep, each spectral element is approximately either full of fluid or fully cavitating. The position of the fluid/cavity boundary, due to the small variation in pressure in the radial direction, follows approximately a radial direction. We are thus justified in varying only the radial elemental boundaries. We illustrate this grid in Figure 1b where the radial elemental boundaries corresponding to = and = separate, approximately, the fluid and air regions. The positions of these radial elements are determined by the pressure distribution at the previous time step. The time steps are kept small enough for and to be slowly varying and thus retain the NISDAT philosophy introduced in Gwynllyw et al. [8] which assumes that the positional change in spectral element grid points between successive time-steps is \u2018small'. The results generated in this paper were produced with the following values of the discretization parameters: Er = 1, Ea = 4 and N = 8"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0000516_j.procir.2016.02.025-Figure3-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0000516_j.procir.2016.02.025-Figure3-1.png",
+        "caption": "Fig. 3. Model of the servo mechanism",
+        "texts": [
+            " In general, each algorithm has its own preprocessing method and requirements for data standardization such as centering and scaling. Likewise, analysis and visualization of the results is different. These processes are encapsulated into the data mining algorithms and integrated to the system together. 5. Case study Here is a problem about a servo mechanism. The resonant frequency is one of the assembly performance indicators. In the view of vibration, bearing and gear are the weak point and affect the stiffness of the whole structure seriously. The model of the servo mechanism is shown in Fig.3. The inner frame system includes inner frame, outer rings of two bearings and anti-backlash gear. It is braced by the bearings with the semi axles which are adjusted by the thickness of the shim. The driving moment torque is generated by electric machine and transfers to the inner frame system from anti-backlash gear. In actual assembly process, the resonant frequency, the resonant peak and the amount of shim adjustment depends on the experience of the technician entirely. It increases labor cost and reduces efficiency"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0001860_0094-114x(95)00022-q-Figure5-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0001860_0094-114x(95)00022-q-Figure5-1.png",
+        "caption": "Fig. 5. A modification to Fig. 4 incorporating prismatic couplings that are equivalent to active couplings transmitting torque to the ports of the epicyclic gear (Fig. 3). The modified coupling network is then over-constrained.",
+        "texts": [
+            " The equivalence is only transitory; the linkage cannot permit more than an infinitesimal displacement without losing its freedom. To this mechanism, or the idealised epicyclic gear train, the two active couplings must be added. The action that each can transmit must be identical to that of a passive coupling that will transmit torque T but cannot exert a force U along the x-axis. The appropriate coupling is therefore a prismatic passive coupling allowing sliding in the x direction. These additional prismatic P couplings are shown labelledj and k in Fig. 5. The coupled members I and 4 to the frame member numbered 0. An Oldham coupling is inappropriate because attention is confined to a 2-system of motion screws. The coupling graph G (Fig. 5) can represent either the actuated gear change shown in Fig. 3, or the equivalent linkage shown in Fig. 4 modified in the adaptation shown in Fig. 5 to introduce overconstraint. Graph G has six circuits and 11 edges, two circuits and two edges more than the graph of the mechanism. Each edge and circuit is labelled and provided with a sense represented by arrow heads. The senses are arbitrary decisions. The motion screw matrix [M]j l . 2 and unit motion screw matrix [/fl]~l, 2, presented in transposed form for convenience, are: and v [ t~ % t; ta t,, tr tg th ti 0 0 ] ; [M ]z j l = at,, bt,, ct,. dtd et,, f t t 0 0 0 u~ uk [M ]2j i = b c d e f 0 0 0 1 ' o > "
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0000577_978-3-319-33714-2_5-Figure4-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0000577_978-3-319-33714-2_5-Figure4-1.png",
+        "caption": "Fig. 4 The first limb",
+        "texts": [
+            " Then from the matrix (1) we obtain a matrix of the cylindrical joint j where sj\u2014a distance from the Xj axis to the Uj axis measured along the directions of the Zj and Wj axes; \u03b8j\u2014an angle between the positive directions of the Xj and Uj axes measured counterclockwise about the positive directions of the Zj and Wj axes. Parameters sj and \u03b8j are variable, and they characterize relative translation and rotation motions of the j-th cylindrical joint elements. Choosing the coordinate systems UVW and XYZ, as shown in Figs. 2 and 3, the constant and variable parameters of the PM 3CCC have been obtained. Constant and variable parameters of the first limb ABC of the PM 3CCC are shown in Fig. 4, where \u03b87 and s7 are the generalized coordinates of the active joint A; \u03b82, s2 and \u03b83, s3 are the variable parameters of the passive joints B and C; all other parameters are the constant parameters characterizing the geometry of links. Constant and variable parameters of two other symmetrical legs are defined similarly. In direct kinematics of the PM 3CCC position of coordinate system XPYPZP attached to the mobile platform 3 are defined with respect to the base frame UoVoWo by known constant geometrical parameters of the links and the generalized coordinates si, \u03b8i, (i = 7, 8, 9)"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0001353_icmic.2016.7804262-Figure1-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0001353_icmic.2016.7804262-Figure1-1.png",
+        "caption": "Fig. 1. Helicopter with two degrees of freedom (CE 150).",
+        "texts": [
+            " ( ) / ( ) (76) Considering (47) and (68), we have \u2016 \u2016 \u221a( ( \u0305 ) ) ( ( ) ) (77) The bounds of and  \u0303 , can be similarly shown. This ends the proof of this theorem. IV. SIMULATION RESULTS In this section, two degrees of freedom helicopter CE150 [18] is used in this numerical experiment. The model is a multidimensional, nonlinear system with two manipulated input and two measured outputs (horizontal and vertical angles) and also with significant cross couplings. The system consists of a massive support, and a main body, carrying two propellers driven by DC motors ( ) as shown in Figure 1. The model can be described by the nonlinear state equations with four states, two inputs, which are the control values for main and side propeller motors. The two outputs are the elevation and azimuth angles. The dynamics of the helicopter are given by: ( ) 0 \u0308  \u0308 1 ( \u0307  \u0307) 0 \u0307  \u0307 1 ( ) (78) with ( ) [ ( ) ], ( ) [ ], ( \u0307  \u0307) 6  \u0307  \u0307  \u0307 7, and 0 1. The expressions of inertia and the center of gravity are given by: ( ) ( ) , and ( ( ) ) . The dynamic model of the helicopter CE 150 with input nonlinearities can be rewritten as follows: {  \u0307 ( \u0305 ) ( \u0305 )  \u0307 ( \u0305 ) ( \u0305 ) ( ) (79) where ( \u0305 ) , ( \u0305 )  \u0305 (  \u0305 is an identity matrix), ( \u0305 ) ( ), ( ) ( )-, ( \u0305 ) ( ) and ( ) , ( ) ( )- ",
+            " INTRODUCTION In the modern civil engineering, construction of the underground structure and the building above the ground have become inseparable two parts. Through the recent engineering practice, people found that the construction of underground structure has become the most important factor in determining the quality of engineering. When the excavation depth is deep, the scope is large or the time is limited, gravity structure is the first construction type to choose. Regular types of gravity structure are shown in Figure 1. It can be seen the first type is piled by rock block which costs large building stones and need to design each stone respectively. The second type applies precast hollow tank by reinforced concrete which has good integrity and fast installation speed. The last type adopts hollow steel cylinder which has low cost and fast installation speed, but the corrosion resistance of the hollow cylinder needs to be tested and verified. Because of these advantages, gravity structure is widely used in civil engineering",
+            " Keywords\u2014Anchored; Sheet-pile; Interaction; Deformation I. INTRODUCTION With the fast development of world economy, the international trade keeps accelerating. Water transport has still been the most widely used transportation way in international trade as its low cost and large carryings. Ship usually needs a vertical free face with certain water depth for the process of loading and unloading. Currently there are two forms of vertical free face, one is formed by large concrete block, the other is formed by anchored sheet pile, shown in Figure 1 respectively. The sheet pile scheme is simple, convenient to construct, and costs less construction materials. The economic benefits are obvious. However, the interaction of surrounding soil and sheet pile is still a hard technical issue. Although many scholars in the world has taken a lot of research proposed a variety of computing scheme, but the results are not satisfying because the problem itself is very complex. For the interaction of surrounding soil and the sheet pile, as well as the deformation and internal force distribution of sheet pile, the researchers have done a lot of work"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0002648_1.869570-Figure5-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0002648_1.869570-Figure5-1.png",
+        "caption": "FIG. 5. The axes x and y are aligned in the plane of centers, y is the axis of symmetry of the triangle and z is orthogonal to the plane of centers.",
+        "texts": [
+            " Thus if, for example, the three spheres have initially the same velocities along the main direction of the translational added-mass tensor, the body S (3) always remains a triangle with two equal sides during the course of the deformation motion. Because of the symmetry one has Z\u0302[0, G50. ~6.1! To illustrate the applicability of ~3.10! and ~3.11! for the three-sphere problem, let us consider the influence of a third remote bubble on the stability of the translational motion of a bubble pair in close proximity ~the remote bubble must translates in the same direction ~see Fig. 5!!. All addedmasses depend now on the 2-D set $d% of the deformation variables ~say, the distance between the remote and the second spheres and the angle between them!. Using then ~4.11! and considering only terms of leading order ~i.e., only the first image, which means that the third bubble is far from the other two! the following inequalities hold: tz~$d%!.ty~$d%!.tx~$d%!, ~6.2! where the axes x and y are aligned in the plane of centers, y is the axis of symmetry of the triangle and z is orthogonal to the plane of centers"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0003425_physreve.63.016611-Figure7-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0003425_physreve.63.016611-Figure7-1.png",
+        "caption": "FIG. 7. Rings ten times larger than those in previous figures. Here k50.005, g50.375, Q510, t5185, b50, and z , from top right to bottom left, is 0, 0.03, 0.07, and 0.10.",
+        "texts": [
+            " Figure 5 shows the plot of the twist density gp(s) 1-9 versus arclength s for z50.1 and b50, 0.01, and 0.02. In these plots we see the variation of the local twist density with the arclength s. The variation is smaller where the mass density is larger, but this effect decreases as b increases. Figure 6 shows a rod with b50.05 and g50.75, twice as twisted as before. We see that, now, the symmetry break occurs even at high viscosity, showing that the asymmetric deformation does not depend only on the medium but also on the applied stresses on the rod. Figure 7 shows a much larger ring, with radius R51/k 5200. Here, g50.375, G50.9, and b50. In this case the principal mode is n550 and we show the effect of a periodic oscillation of the density with Q510. For the sake of clarity, only the central curve of the rod is drawn. The case Q510 replicates the effect shown in Fig. 3; each high-density segment of the ring tends to a more flat position while the small 016611 density parts coil in large loops. Figure 8 shows the effect for b50.01 and b50.02. We have also considered the case where the frequency of the density oscillations Q, is much higher than the frequency of the last unstable mode, A11Gg2/k2"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0000188_0954406214560420-Figure12-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0000188_0954406214560420-Figure12-1.png",
+        "caption": "Figure 12. Coupled motion of 4-UPS/UPR mechanism.",
+        "texts": [
+            " Axis xo, yo is parallel to a3a4, a4a1, respectively, and zo is perpendicular to the platform. Coordinate system p-xyz is attached to the mobile platform, and its initial orientation is parallel to coordinate system o-xyz. The revolute (R) joint of UPR limb is attached to the mobile platform, and its axis is parallel to axis yp. UPR limb determines the motion characters. In any configuration, UPR limb provides a constraint force along the R joint, and a torque, which is perpendicular to the universal (U) joint of UPR. As shown in Figure 12, a coordinate system e-xyz is established to describe the coupled motion of the mechanism. Coordinate system e-xyz coincides with o-xyz at the initial state, and its orientation always coincides with coordinate system p-xyz. When the mobile platform rotates around axis xe, the platform will produce adjoint displacements ybp, zbp. This adjoint affection also reflects on the velocity of the mobile platform. The coupled motion equation is gotten through an equivalent motion method,24 but its derivation process is a little complicated"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0002057_rob.4620110202-Figure2-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0002057_rob.4620110202-Figure2-1.png",
+        "caption": "Figure 2. Cartesian transition model.",
+        "texts": [
+            " Compared to the work of the authors mentioned, who have developed this approach of trajectories composed of straight line segments connected by smooth arcs, our originality lies in the fact that we propose a trajectory model still based on this approach but in which the necessary time parameters are all analytically deduced from velocity and acceleration pa- rameters. In our work, time parameters are not supposed to be constant, as the ''taC:' parameter in Paul's work13 or computed from an optimization procedure as in Luh's ~ o r k . ~ , ~ A typical trajectory is shown in Figure 1. Figure 2 shows the details of a typical transition. Two time parameters have been introduced: the motion time duration between the nodes and Pi, noted Ti; and the half time duration of the transition in Pi, noted 7;. The constant velocity magnitude on the segment going to Pi is noted V j and the maximum acceleration magnitude during the transition in Pi is noted Ai. Starting with an assumed acceleration profile that is continuous, zero at both ends, and that reaches a maximum half-way through, one can readily derive the equations of the trajectory by using various continuity constraints on velocities and positions at both ends of the transition"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0003048_iros.1994.407649-Figure1-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0003048_iros.1994.407649-Figure1-1.png",
+        "caption": "Figure 1 The general configuration of a TSS being controlled using a manipulator.",
+        "texts": [
+            " The objective of this paper is to investigate the possibility of using a space manipulator to stabilize the dynamics of tethered satellite systems during station keeping and retrieval. To start with, a dynamical model is developed for the system consisting of a spacecraft-mounted manipulator and a tethered payload. Control laws are then developed using this dynamical model. Finally, computer simulations are carried out to validate the control laws developed. EQUATIONS OF MOTION The system under consideration, shown in Figure 1, consists of a main spacecraft, the orbiter, a two-link spacecraft-mounted manipulator and a subsatellite, connected to the orbiter by a tether. In deriving the equations of motion. it is assumed that the spacecraft and the maiiipulator are rigid and the entire motion is coplanar with the orbital plane. The subsatellite i s modeled as a point mass, while the tether is assumed to be massless and to reiliain straight during the inotioii With the above assumptions. the systeni has four degrees of freedom (DOFs) which are described by the followiiig generalized coordinates: q, , the pitch angle of the spacecraft: c["
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0003785_mhs.1998.745770-Figure1-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0003785_mhs.1998.745770-Figure1-1.png",
+        "caption": "Fig. 1 Concept of Bio-aligner",
+        "texts": [
+            " If the location of the object is not known exactly, the system must measure the location of the object, otherwise automation is impossible. We have studied on the image recognition of the object. Image recognition the position error. If the object is easily fixed at the exact location, it helps to realize automation of the contact manipulation tasks. This also leads to the mass production. However, conventional methods are not suitable for bio-automation. Here, we propose the bio-aligner, which can control position and orientation of the object and fix it firmly even if the external force is applied, and release it fast. Figure 1 shows the concept of the bio-aligner. This unit can be arrayed on the 2D surface. Figure 2 shows the concept of the bio-micromanipulation system where the bio-aligner is set on the microscope. Desirable function of the bio-aligner is summarized as follows: (a) it can fix the object even if the external force is applied; (b) it can inhibit adhesion of the object; (c) it can control the position/ orientation of the object. Generally, many different kinds of forces are acting on the cell in solution such as the gravity, buoyancy, resistance force from viscosity of the fluid, Brownian motion, interactive forces in the microhano world such as Van der Waals force, electrostatic force depending on the surface electrons"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0002926_978-1-4684-6237-1-Figure2.1-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0002926_978-1-4684-6237-1-Figure2.1-1.png",
+        "caption": "Figure 2.1 Multi-modal sensing",
+        "texts": [
+            " In engineering, analysis and observations drawn from existing systems are used to refme and improve the science of the subject. In this way we move from a collection of 'rules of thumb' towards a theory that represents the current body of knowledge. Unfortunately, AI and robotics have not yet reached this point and so there are no fundamental 'text books' on topics such as sensor selection. Nevertheless, we must try to consider the options as scientifically as we can. Consider an example task as a way of illustrating the sensor selection problem. Figure 2.1 shows an insertion process, where a peg is to be inserted a small distance into the hole in a block. Suppose that the peg has plenty of clearance but occasionally (due to undefined errors) the peg is not quite central in the jaws and thus can hit the edge of the hole. Suppose further that we have been asked to design a sensory system that will signal when this error event occurs. The diagram shows five possible sensing methods: It seems that each of these could satisfy the functional requirement of detecting the specific error, and yet they are entirely different in nearly every respect",
+            " It is not easy to ensure that a pressure sensor will only experience the desired stimulus without interference or that a vision sensor will only 'see' a restricted range of objects. Consequently, sensory fields will frequently overlap in some way, especially if a high degree of coverage is aimed for. Consider different manifestations of sensory overlap: Multi-modal overlap - it is common for two quite different sensory systems to be able to detect and report on a particular event This has already been seen in Figure 2.1 and is quite easy to produce in practice. Intra-modal overlap - this occurs when two sensors of similar type have fields that share a common region. The overlap can be simple and direct, as in two cameras that view the same work area, or it may be created through more indirect means, as in internal force sensing in a robot arm where the arm configuration provides a coupling between different sensors. Redundancy - this is the case of deliberate overlap where one sensor duplicates the field of another sensor"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0000153_imece2015-50907-Figure3-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0000153_imece2015-50907-Figure3-1.png",
+        "caption": "FIGURE 3. THE COORDINATE SYSTEMS OF THE NEW CONFIGURATION.",
+        "texts": [
+            " For its special non-spherical wrist, the original point of the base frame is actually at the virtual wrist point which is the projection of the intersection point of axis 5-1 and 5-2 onto axis 4 when the joint angles of axis 4 and 5 are both 0, as shown in Fig. 2. Meanwhile, the link 1 and 2 of the new configuration are the forearm and upper arm of the real manipulator. We also leave out the offsets between Axis 1, 2 and Axis 3, 4 which does not impact on the position of the real base and simplifies the calcula- tion. And because the motion ranges of axis 4, 5, 6 are extensively wide, a 3-DOF spherical joint is designed at the virtual wrist point in addition. Coordinate systems are formulated and noted as S1, S2, S3, S4, S5, S6 which are shown as Fig. 3 . And the corresponding D-H parameters are listed in Tab. 1. To facilitate description and calculation, as shown in Fig. 4 2 Copyright \u00a9 2015 by ASME Downloaded From: http://proceedings.asmedigitalcollection.asme.org/pdfaccess.ashx?url=/data/conferences/asmep/86880/ on 01/31/2017 Terms of Use: http://www.asme.org/about-asme/terms-of-use Z1 Y1 X1 A FIGURE 4. THE EXPRESSION OF LINK 1 BY SPHERICAL CO- the end position of link 1, namely the position of point A, can be expressed by spherical coordinates as P\u0304A =  1xA 1yA 1zA = a3 cos\u03b1 sin\u03b2 a3 sin\u03b1 sin\u03b2 cos\u03b2  (1) in which r is the length of link 1, \u03b1 is the angle between axis X1 and the projection of link 1 on X1OY1, and \u03b2 is the angle between axis Z1 and link 1"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0001672_1.1712746-Figure1-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0001672_1.1712746-Figure1-1.png",
+        "caption": "FIG. 1. Section of a lubricating film in a full journal bearing.",
+        "texts": [
+            " Muskat and F. Morgan, J. App. Phys. 10,46 (1939). 7 F. Morgan and M. Muskat, J. App. Phys. 10, 327 (1939). 142 small, that the pressure variation across the film may be neglected, and that fluid flow will consequently take place only in the x and z directions. With the same assumption of a small film thickness its form may be expressed as: h=c+e cos 0, (2) where the origin of coordinates is taken at the position of maximum film thickness, c is the radial clearance, and e is the journal displace ment (d. Fig. 1). If we now introduce the notation: 'I/=e/c; w=z/r; x=rO, (3) where r is the journal radius, and assume that JJ. is constant everywhere, Eq. (1) assumes the dimensionless form: a { ap } a2 p - (1+'1/ cos 0)3- +(1+11 cos 0)8- ao ao aw2 = -/311 sin 0, (4) where (5) For the problem of electrical conduction in two dimensions the basic differential equation corre sponding to Eq. (1) is ~(/V)+~(/V) =q, (6) ax ax az az where V is the potential, (}\" the conductivity, and JOURNAL OF APPLIED PHYSICS [This article is copyrighted as indicated in the article"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0003124_70.88068-Figure10-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0003124_70.88068-Figure10-1.png",
+        "caption": "Fig. 10. Locating contact point between a cylindrical object and a circular slab.",
+        "texts": [
+            " Therefore, the exact location of the wrench axis can be obtained using either of the following equations: CONTACT LOCATION If the wrench axis intersects the object surface at more than one point, a criterion must be used to select the most probable solution. This criterion is based on the evaluation of the angle between the object surface normal and the obstacle (fixture) surface normal at the contact points. When two bodies are in contact, the normal unit vectors at the point of contact have to be in opposite directions [lo]. This is demonstrated in Fig. 10 where the manipulator hand is holding an object represented by a circular cylinder and the fixture configuration is a circular slab, with contact expected at point A . Let the wrench axis acting in the compliance frame be given by a= (0, m, 0, P, 0, 0) (22) which produces the following parametric equation of the line of action: x= 0 y = t z=zo. 538 IEEE TRANSACTIONS ON ROBOTICS AND AUTOMATION, VOL. 5 , NO. 4, AUGUST 1989 Substituting into the quadratic equation ( x 2 + y 2 = a 2 ) we obtain two points on the cylinder surface with the following coordinates: AI =(O, a, zo) A2 = (0, - a , 20)"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0000386_celc.201600175-Figure7-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0000386_celc.201600175-Figure7-1.png",
+        "caption": "Figure 7. Bioelectroreactor with A) the inner chamber as counter electrode compartment equipped with a cation exchange membrane and B) an illustration of the whole reactor.",
+        "texts": [
+            " After about 7 h the late exponential growth phase was reached and the preculture was used for inoculation of the experiment. Precultures from different parallel batches were combined prior to the experiments to provide the same inoculum for each bioelectroreactor. Conventional 1 L glass bioreactors (Multifors, Infors HT, Bottmingen, Switzerland) were equipped with an upgrade kit allowing the performance of bioelectrochemical studies,[27, 39] these reactors are termed within this study bioelectroreactors (see Figure 7). All experiments in the bioelectroreactors were carried out under potentiostatic control using a multi-channel Potentiostat/Galvanostat (VSP, BioLogic Science Instruments, Claix, France). The electrodes were operated in a three electrode arrangement consisting of working electrodes (WE), an Ag/AgCl reference electrode (SE 11, 0.195 V vs. SHE Sensortechnik Meinsberg, Germany) and counter electrodes (CE). The working and counter electrodes used throughout this study were pairs of graphite rods (CP-Handels GmbH, Wachtberg, Germany) with dimensions of 10 250 mm"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0000251_maes.2014.130034-Figure2-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0000251_maes.2014.130034-Figure2-1.png",
+        "caption": "Figure 2. Moments and forces acting on the helicopter.",
+        "texts": [
+            " HELICOPTER MODELING In this section, we present the model of the 3-DOF helicopter portrayed in Figure 1. This particular helicopter model is restricted to 3-DOF, which are the pitch movement, travel rate, and vertical movement [13]. The 3-DOF helicopter consists of a base upon which an arm is attached. The arm has the two ends: the first one includes two direct current (DC) motors with propellers, and the second one represents the counterweight. The arm can evolve around the pitch axis and can swivel the travel axis (Figure 2). To measure these different movements, three encoders are fastened on the helicopter\u2019s body. The propellers generate proportional forces to the voltage motors. These forces can cause the helicopter body to lift off the ground. Moreover, the aim of the counterweight is to reduce the power applied to the DC motors. To simplify the modeling of the helicopter, we consider the following assumptions: C All angles are sufficiently small, so the linear approximation is valid. C The friction forces are neglected",
+            " In comparison to other robust standard H\u221e [5] controller designs, the selection of the weighting functions in H\u221e loop shaping controller design is not very difficult, because the two weighting functions are chosen according to some well-defined system specifications such as bandwidth and steady-state error requirements. In addition, the H\u221e loop shaping controller can be found with no iterative computation, and its implementation is not very difficult because its order is not very high, unlike standard H2 and H\u221e controllers. In the future, we will investigate the sensorless control of elevation, travel rate, and pitch by using an extended Kalman filter observer. APPENDIX A The elevation dynamic of the helicopter is described by the following differential equations by referring to Figure 2: SEPTEMBER 2014 IEEE A&E SYSTEMS MAGAZINE 11 (1) The elevation torque is controlled by the total force Fm = Ff + Fb. The forces Ff and Fb are generated by the two propellers, where Vm1 and Vm2 are the voltages applied to the front and back motors resulting in forces Ff and Fb, respectively. In addition, Tg is the effective gravitational torque due to the mass differential Fg about the elevation axis. The following pitch dynamic is obtained: (2) The horizontal component of Fg will cause a torque about that the travel axis, which results in acceleration about the travel axis: (3) The overall motion of the helicopter is described by the following equations: (4) The first one characterizes the elevation dynamics, the second one represents the dynamic of the pitch motion, and the last one describes the travel rate motion"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0000602_s11771-015-2876-0-Figure5-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0000602_s11771-015-2876-0-Figure5-1.png",
+        "caption": "Fig. 5 Velocity distribution on radical plane (a) and central axial (b)",
+        "texts": [
+            " Figure 4 shows the air phase pressure distribution and air flow stream line inside the bearing cavity at the inner ring rotating speed of 104 r/min. It seems that the air flow direction coincides with shaft rotating direction. Since the steel ball rotates at a high speed compared with the stagnant outer ring, a negative pressure zone appears in the contact zone between the ball and the outer ring. Around the negative pressure zone, a high pressure zone is formed, and the pressure is higher at the zone along the direction of the rotation. Figure 5 shows the air phase velocity contour and the stream line on both radial and axial planes of the angular contact ball bearing. It can be seen that the maximum velocity is near the inner ring, and the minimum velocity is around the outer ring. This phenomenon can be explained with the momentum conservation near the wall. Also, the velocity change in the axial plane is more obvious than that in the radial plane. In the axial central plane, the air enters the bearing cavity from the position near the inner ring, as denoted by A and B"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0002184_an9952001137-Figure1-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0002184_an9952001137-Figure1-1.png",
+        "caption": "Fig. 1 P1 and P2. polarizers; a. phase angle. Experimental set-up of the split cell: SC, split-type flow cell;",
+        "texts": [
+            " When this compound shows optical activity and rotates the plane of polarized light by an angle (J, I is further changed to I = Ioe-EcfKcos2(a + 6) If the assumption is made that the column effluent flows into both cells at the same velocity, then the output as an absorbance, Abs, from the detector equipped with the split cell can be expressed as Abs = log [cos2(a(, + (3)/cos2(as + (J)] (1) where the subscripts s and r indicate sample and reference sides, respectively. The effects of light absorption are offset in eqn. (l), see Fig. 1. Next, we consider the case when a(, = -as in eqn. (1). When the optically-active analyte is in the split cell, the signal intensity, AAbs, is given by AAbs = -log [COS~(-CX~ + ~)/cos~((x, + p)] (2) because the baseline absorbance is given by log (cos2a,/ C O S ~ ~ , ) . By substitting f ( x ) for log (cosx) in eqn. (2), AAbs can be rewritten as AAbs = -21f(as - (J) - f(a, - (J + 2(3)] Pu bl is he d on 0 1 Ja nu ar y 19 95 . 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 8/ 10 /2 01 4 09 :3 1: 38 ",
+            " 9 res. 3 - 0 10 Time/min Fig. 3 Comparison of sensitivities of the split and normal cells. Conditions: eluent, acetone-water (5 + 3); flow rate, 0.5 ml min-1; column temperature, 70 \"C; detection wavelength, 530 nm; phase angle between two polarizers (a), f0 .9 rad (split) and +0.9 rad (normal); time constant, 10 s for res. 9 and 0.5 s for res. 3. See text for further conditions. A, Normal; B, split. two on the transmitting light side, whose phase angles are the inverse of each other, as shown in Fig. 1. This assembly was inserted into a Shimadzu (Kyoto, Japan) SPD-1OAV UVNIS spectrophotometric detector, and used as the PPD. Regulation of this detector to obtain high PPD sensitivity was described in a previous paper.4 Chromatographic separations of sucrose and glucose syrup were carried out on a 250 x 4.6 mm id TSK-Gel Amide-80 column (Tosoh, Tokyo, Japan) maintained at 70 \"C by a Shimadzu CTO-1OAC column oven. Acetone-water eluents were delivered with isocratic or gradient elution using a Shimadzu LC-4A pump"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0002055_ip-epa:19960203-Figure19-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0002055_ip-epa:19960203-Figure19-1.png",
+        "caption": "Fig. 19 Experimental arrangement for step change in torque",
+        "texts": [
+            " 17 to keep the resultant rotor flux-linkage constant. The results of these experiments are shown in Fig. 18, from which it is clear that, as predicted, the torque depends only on the slip, and not on the absolute speed of the rotor. (The minor variations which occur are probably due to the change of rotor resistance with temperature.) Normalised stator current and torque against slip frequency under Accordingly, all the tests were carried out with the rotor at rest (but not locked), and with torque applied by means of weights as shown in Fig. 19. One weight is attached to the motor shaft via a stiff rope, and applies a constant load torque. (It was important to minimise system inertia, so a large weight near to the axis was desirable.) The other weight is suspended by an electromagnet from the end of a horizontal arm fixed to the motor shaft. The current in the magnet is set so that this weight is only just suspended, so that when the current in the magnet is switched off, the weight immediately falls away and there is a sudden step reduction in the total load torque"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0000617_s1064230716030126-Figure6-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0000617_s1064230716030126-Figure6-1.png",
+        "caption": "Fig. 6. Imaging of reference points.",
+        "texts": [
+            " This procedure is the final one for each control step. After it a signal of the completion of the step is transmitted to Earth. In order to find expressions with which the aforementioned correction to the program vector can be calculated, we will use well-known optical conversion formulas, which localize two-dimensional vectors , , and xi, with the vectors of the positions of the reference points in model and the real environments of the simulator and the space robot, respectively. These formulas are easy to understand using Fig. 6, which illustrates the optical conversion for one of the two TV-cameras located on the space manipulator\u2019s working tool. All notations in the figure corresponds to this conversion: (4.1) \u0399\u0394 rX \u0399\u0394 GX r.p ix r.p ix r.p ix r.p ix \u0399 ( )rX t ix ix ix ix ix ix \u0399\u0394 rX \u0399 rX r.p ix r.p ix \u0399\u0394 rX \u0399\u0394 rX \u0399\u0394 GX \u0399 \u0399+ \u0394( )r rX t X \u0399\u0394 GX \u0399\u0394 GX \u0399 \u0399+ \u0394 + \u0394( )r i r GX t X X I rX I rX \u0394 + \u0394I r GX X \u0399\u0394 rX \u0399 rX r.p ix r.p ix ix r.p ix ( )iX \u2022 \u2022 \u2022 \u2022= \u2212\u03c4 \u03c4 = \u2212 \u03c4 + \u03c4 =r.p ( ) r.p ( ) r.p ( ) r.p ( ) r.p 1,2 r.p 3 r.p r.p 1,2 r"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0002990_901764-Figure2-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0002990_901764-Figure2-1.png",
+        "caption": "Figure 2. Isoparametric plate element (a) Typical solid element",
+        "texts": [
+            " This yields the elements of the j-th column, {Ku, K2j, K3j, . . . , K ~ ~ ) T , of the coefficient matrix F;ijlnxn . Repeating the procedure n times, we can obtain all the coefficients in the matrix [Kijlnxn. Of course, the influence coefficients Kij's include bending, shear, and contact effects (3). Due to the complex geometry of a meshing pair of teeth, the finite element method is adopted for the bending and shear deflections. Isoparametric plate elements are used which have 3 degrees of freedom, one dispacement and two rotations, at each node. Figure 2 shows a typical isoparametric plate element and Figure 3 shows an example of generated meshes on the neutral surface of a gear tooth. Using isoparamemc plate elements, the transverse shear deformation and the rotations are automatically accounted for, whereas the transverse shear deformations are neglected in plate elements and the rotations are not included in three-dimensional solid elements. Moreover, isopararnetric plate elements have the advantages that the size of the stiffness matrix is reduced and the computing time is also reduced by avoiding integration along the thickness (b) Isoparametric plate element after specialization from solid element (c) Degrees of freedom at node i To generate the finite element mesh, the geometries of an involute tooth and contact line, and the positions of moving contact along the line of action with one engagement cycle are defined based on references (2,3)"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0003124_70.88068-Figure3-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0003124_70.88068-Figure3-1.png",
+        "caption": "Fig. 3. Objects geometric representation [(a), (b), and (c) are special cases",
+        "texts": [
+            " It is proposed to use an idealized geometric representation of the object or contact surfaces instead of using elaborate geometric modeling techniques which require relatively long computation time and are not suitable for real-time control systems. Fig. 2 shows that by using the idealized representation of an object combined with the knowledge of task space, the likely contact region or location can be predicted. In most cases the object can be modeled by a quadratic surface represented by [6], [7] + Gx+ Hy + ZZ + K = 0 (1) where A , B, . . . , K are real coefficients and x, y , and 2 are Cartesian coordinates in the compliance frame. Some of these coefficients become zero for specific geometric shapes as illustrated in Fig. 3. Equation (1) is versatile and provides adequate mathematical representation of various objects for use in contact reasoning during robot compliant motions. Polyhedra may also be used to represent objects geometry, however, this adds extra steps to the algorithm used for contact reasoning because of the discontinuity of surface normals. MODELING OBSTACLE SPACE The success of any compliant motion path generated by a motion planner depends on the definition of the initial contact Configuration. Fig"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0001762_a:1018959823016-Figure2-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0001762_a:1018959823016-Figure2-1.png",
+        "caption": "Fig. 2. rtrr-robot.",
+        "texts": [
+            " We fix the virtual normal a\u22121 in the x0-axis line, then fix the (virtual) normal line corresponding to a0 in the same line as the x1-axis (choose x1-direction to be equal with x0-direction), then a rotation around the vertical axis (0) (this axis coincides with axis (1)) is measured by \u03980. Accordingly, the normal a1 and the x2-axis are fixed in the same line being parallel to a0 (note x2 here has opposite direction to x1); a2 and the x3-axis coincide with the line spanned by the link with length l2, similar with the a3 and a4 and the corresponding x-axes, respectively (cf. the convention of fixing local coordinate systems, (D\u2013H)). Please note, we set the parameter \u03982 = 0 (this measures the angle between a1 and a2) in contrast to the picture of the rtrr-robot arm (Fig. 2) where \u03982 = 90\u25e6. Based on this data we can now find the corresponding local transformations T i\u22121 i , then calculate their product and establish the global matrix kinematics relation T 0 n = E from which we derive the polynomial equations in the si, ci. This is presented subsequently. Having the general formulas implemented in a computer algebra module this amounts to just inserting the right parameters and evaluate the results, system respectively. Going into greater details here, we display all the individual local coordinate transformations explicitly"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0002874_881783-Figure6-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0002874_881783-Figure6-1.png",
+        "caption": "Fig 6. Response lag to steering torque {Running pattern Hl",
+        "texts": [
+            " While we are actually operating a motor cycle on a public road, we experience or en counter a variety of situations - we may be changing the lane, passing other traffic and coming back on the original lane J or we may be averting a person or motorcycle that has sud denly jumped into our way. In consideration of such situations these two courses were set up. On these courses the Japanese highway lane widths were used for a lane passing width and an obstacle averting width. Running Patterns (1) In Running Pattern I, the motorcycle ran in 881783 3 1\":.lr the ,-lep<?ndence of the response l::tg on the motorcycle speed, Fig. 5 (Running pattern I) and Fig. 6 (Running pattern II) show the relation of the lag in the test motorcycle response to the motorcycle speed. The fOllowing describes the equivalent characteristics which are represented by the similar dependency of all four test motorcycles on the motorcycle speed. (l) A lag in the lateral acceleration response -\u00b7of the frame in Running pattern I. (2) A lag in the yaw rate response in Running pattern II. (3) A lag in the roll rate response in Running pattern II. (4) A lag in the lateral acceleration response of the frame in Running pattern II"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0001113_apede.2016.7879015-Figure2-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0001113_apede.2016.7879015-Figure2-1.png",
+        "caption": "Fig. 2. Vibration generator - structure",
+        "texts": [],
+        "surrounding_texts": [
+            "Penza State University, e-mail: kipra@mail.ru MULTI-CHANNEL VIBRATION MEASUREMENT MACHINE The article highlights the problems of the efficiency of vibration test methods and means. It gives the brief overview of existing methods. It also gives characteristics of the developed multi-channel vibration measurement machine, its component parts and the benefits of its application. The test results are given that prove that the developed vibration vibration measurement machine allows to reproduce the necessary vibration characteristics similar to real-life vibration exposure during operation life of an object. Achieving technical efficiency of vibration tests for electronic devices (ED) is a complex problem connected with the solution of interrelated tasks [1]. The first task - the choice of adequate vibration test modes - arises due to the fact that the test object (TO) functions in different systems, when used in actual and bench test conditions. In the first case the ED interacts with the external environment (the object equipped with the electronic device), in the second - with the bench stand that has to reproduce the vital influences of the external environment and provide true-to-life conditions of text object operation. In high-frequency subrange during the transmission of the vibration from the table to TO attachment periphery the TO is subjected not only to the intertial influence but to the deformation influence as well, in which the attachment points of the TO make antiphase elastic displacements. The second problem is the precision in rerproducing the vibration test modes. The problem appears due to significant interaction between the test object and the bench stands, and also due to the limited possibilties of the bench stands to reproduce the vibration modes. Research in this area has been conducted only with inertial (non-resonance) loading. Physical reasons of the distortion of test modes have not been analyzed in detail for more general and the most important situation for the ED, when the test object is a complex vibratory system. There is also no simple and efficient means for eliminating these distortions. That is why, the distortions of vibration modes during tests amount to hundreds of per cent. The vibration measurement systems, that are in current use, only take consideration the inertial component when producing a test signal of vibration. This is the reason for insufficient equality of bench tests and real-life operation conditions. [2, 3]. That is why it is necessary to consider the deformation component as well when exciting all natural frequencies (including the even ones) [4, 5]. The multi-channel vibration measurement machine was developed to solve the above-mentioned problems. The machine consists of the postitioning system and four vibrators with current sensors. The positioning unit was built on the basis of Purelogic PLRA4 machine. The positioningl system performs the movement of the vibration instrument along X, Y, and Z axes with max. speed 5000 mm/min. The movement is performed by step motors and limited by end probes on all three axes. The vibrator consists of two main systems: - moving system; - magnet system. Both systems are joined with attachements. The magnet system comprises: - the magnet path; - two permanent magnets. The moving system consists of: - two coils of copper wire on paper carcasses; - stem pusher, used for transferring the vibration impact; - suspensions. The measured element is located on the one suspension of the moving system, constructed in the form of a radial centering spring. The external view and structure of the vibration generator are shown in Fig. 1 and 2. Vibration generators are fixed at the base of the positioning system in such a way that the axes of stem pushers coincide with the fixture holes of the tested object. The vibration generator is used for producing the speed of vibration-force impact of linear movements. Piezotransducer is made with the plate spring, bearing two piezoresistive strain gages, mounted according to Wheatstone bridge scheme. Plate bending causes the strain gages deformation. Piezoresistive accelerometers, that use resistive sensors with low impedance, are insentitive to external ghost influences and disturbances. The pilot model was built on the basis of the suggested structure. The positioning unit was built on the basis of Purelogic PLRA4 machine. The positioning system performs the tested object along the axes X, Y, and Z with maximum speed 5000 mm/min. The movement is performed by step motors and limited by end probes on all three axes. Control of the positioning system is realized inn the AVR microcontroller, model atmega8. Polulu products - step motor driver (Fig. 3) - are used to cope with high-current load. Microchip DRV8825 of TexasInstruments gives the maximum current of the coil as 2.5 A, with source voltage 24V and temperature Microcontroller timing is performed with the external quartz-crystal resonator with 11.0592 mHz frequency. Microchip DRV8825 timing is performed with internal source signal, which allows the driver to accept control signals STEPd with maximum frequency 250 kHz. The unit of step motors control (Fig. 4) supports the step partitioning, microstepping, partitioning coefficients: 1/2/4/8/16/32, which allows to achieve a smooth movement and a more precise positioning of step rotors. Microstep is also one of the means of overcoming such unpleasant phenomenon as resonance of step motor drive, which is evident at low rotation frequencies. In practice, the effect of resonance leads to difficulties in operating at close to resonance frequencies. Resonance can significantly lower the precision of the drive. In systems with low damping there is danger of losing steps or exceeding the noise, when the motor is working close to resonance frequency. Apart form the main microcontroller, the ooperation mode is managed by 9-position DIP switch, which is subdivided into 3 groups with 3 switches closing the corresponding contacts pairs. Table 1 lists the modes of step division, depending on the position of the switches. 0 - the switch is turned off, 1 - the switch is turned on. Microchip DRV8825 has inbuilt overheating and winding short circuit protection. It activates when the current exceeds 3A. The protection works as the emergency stop of H bridges of the chip, logical block of step indexing and control. \u2018FAULT\u2019 signal lights the corresponing diode on the error panel. Through rs232 interface the microcontroller receives commans either from the computer with pre-installed software or from the control panel. Commands are subdivided into two main types: a) rotation clockwise/anticlockwise; b) movement at a set value (can be performed only with the special software). Table 2 describes commands of A-type. Commands of B-type are the 5-byte package with header \u2018send\u2019 0b00000000 (bin), 0 (dec). On receiving the header, the microcontroller waits for the command \u00abto*\u00bb (with axis X,Y, or Z instead of *), that contains 1 byte of information of the data option of the rotation axis, and 3 bytes showing (openly) the number of steps determining the servo drive rotation of the chosen axis. In the last three bytes high-order bit is responsible for the rotation direction, the remaining 23 bits, with maximum step units \u201832\u2019 (6400 steps per rotation) and the ballscrew step 0,4 mm/rot determine the absolute movement for more than 524 mm, which is enough for positioning on A4 format size surface. Fig. 5 shows the graphic interface of the programming module of the positioning system control. The \u2018Control\u2019 tab contains the elements of drives rotation control and elements for assigning the coordinate for movement, in milimeters. The \u2018Settings\u2019 tab contains the option of the serial interface for connection with atmega8. In the drop-down lists \u2018Microstepping\u2019 and \u2018Ballscrew\u2019 the options are chosen for step unit and ballscrew step, for transorming the default values of the number of drive rotatios to flatwise movement of the vibration sensor in milimeters. The drop-down lists \u2018Accelerate\u2019 and \u2018Maxspeed\u2019 allow to adjust the values of acceleration/deceleration and the maximum rotation speed of the rotor of the step motor. \u2018Send\u2019 key sends the set/changed parameters to the controller, \u2018Home\u2019 key transmits command to the machine: determine the absolute zero of the positioning system. As a result the machine movement in coordinates can be controlled. On receiving the movement instructions, the microcontroller sends the control signals STEP/DIR to drivers of step motors. Run-up is a very important moment. During run-up or slow-down it is important to chose correctly the law of speed change and maximum acceleration. The greater the load inertia is the lesser should be the acceleration. The criterion of the correct choice of acceleration mode is performng the run-up till the required speed for specific load in a minimal time. In the developed system the run-up and slow-down with constant acceleration were applied. Use of the rule, governing acceleration or deceleration of the motor, is program-controled, by the main microcontroller. Thus, the developed multi-channel vibration system allows to set the vibration impact, whose dynamic characteristics correspond the actual ones in real operating conditions. That is a key aspect when choosing the methods, means and modes of conducting vibration tests. The use of independent influence of fixing points of the TO allows to create resonance in all its forms in the studied frequencies range. The article has been prepared within the project \u2018Developing methods and means of developing highly reliable components and systems of on-board radio-electronic equipment for new generation machines in aerospace and -19-10037 of 20th of May, 2015) with the financial support of the Russian Science Foundation. References 1. Ostromensky P.I. Vibration Testing of Radio Equipment and Devices / P.I. Ostromensky. Novosibirsk: Novosibirsk University, 1992. 173 p. 2. Golushko D.A. Speed of Frequency Change During Tests of Dynamic Characteristics of the Structure / D.A. Golushko, A.V. Zatylkin, A.V. Lysenko // 21st Century: Past Results and Present Problems Plus. 201 P. 147-154. 3. Tankov G. The Report of Data Exchange of Multichannel Vibration Testing Equipment with the Program Environment of Researches Control / G. Tankov, A. Zatylkin, N. Yurkov // International Journal of Applied Engineering Research. 2015. Vol. 10, 23. . 43839-43841. 4. Dynamic Analysis and Optimization of Parameter Control of Radio Systems in Conditions of Interference / A. K. Grishko, N. V. Goryachev, I. I. Kochegarov, N. K. Yurkov // Proceedings International Siberian Conference on Control and Communications, 2016. DOI: 10.1109/SIBCON.2016.7491674. 5. Golushko D.A. Method of Controllilng the Amplitude of Resonance Frequencies of Structural Elements of Onboard Radio Equipment/D.A. Golushko, G.V. Tankov, N.K. Yurkov//Innovations on the Basis of Information and Communications Technologies. Vol. 1. P. 306-309."
+        ]
+    },
+    {
+        "image_filename": "designv11_60_0000510_htj.21229-Figure1-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0000510_htj.21229-Figure1-1.png",
+        "caption": "Fig. 1. Structure of spindle system.",
+        "texts": [
+            " As for the above problem, this paper, based on a vertical machining center, conducts simulation on the thermal characteristics of machining tools under the condition that the idling spindle has a cooling channel. With the major research on the temperature field and thermal deformation of the spindle system, this paper verifies the effectiveness of the simulation by testing, and then optimizes the weak links of the thermal characteristics of the spindle system. The structure of the spindle system is shown in Fig. 1. The thermal deformation of the spindle system is in themicron grade, and it drives the test object to rotate at high speed; therefore, the surface roughness of the test object should be relatively small, with high requirements for the concentricity and roundness. We take the spindle accuracy test bar as the base object being tested, and using highaccuracy displacement sensors to obtain the deformation in X, Z, and Y directions, we obtain the thermal deformation of the spindle system in the corresponding directions"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0003835_robot.1994.351168-Figure2-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0003835_robot.1994.351168-Figure2-1.png",
+        "caption": "Figure 2: Wheelchair schematic.",
+        "texts": [
+            " For example, being a few inches off of the reference path in a factory corridor may be acceptable for many AGVs, but straying a few inches as a wheelchair passes through a doorway could be damaging or disastrous. Furthermore, the system must provide a smooth ride, with small accelerations, since it carries a human passenger. The accuracy required for the wheelchair application demands very precise estimates of pose (i.e., a very good answer to the question Where am I?\u201d). The experimental wheelchair system is pictured in Figure 1. This wheelchair is driven by two rear wheels which are actuated independently for steering control. For this drive configuration, shown schematically in Figure 2, a set of differential equations which relate the differential wheel movement of the two drive wheels to the differential position and orientation of the wheelchair have been derived [1][2]. These differential equations can be integrated numerically by measuring the wheel motion of the two drive wheels using optical shaft encoders. In Figure 2, the position of a point on the wheelchair is denoted by X and Y, while the orientation of the wheelchair is denoted by d. Estimates of the wheelchair\u2019s pose produced by this numerical integration have been referred to in the literature as \u201cdead-reckoning.\u201d However, if there are any initial estimation errors, modelling errors, or disturbances (such as wheel slippage), then errors in the wheelchair\u2019s pose estimates produced by dead-reckoning will grow as the wheelchair travels throughout its environment"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0002092_0094-114x(96)84603-9-Figure9-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0002092_0094-114x(96)84603-9-Figure9-1.png",
+        "caption": "Fig. 9. Position of concentrated load on gear tooth.",
+        "texts": [],
+        "surrounding_texts": [
+            "a(') 7.3184 F. (0.8489 + 0.2771 #g 0.1260 t&)fz (hv)f2 (?D, hv) Rill = ~ -- 2 (g) (g) mOw.07,5/~_o.0940(,~o.0782(~)-o.0396(~)-o.o,54(__~ ) 0.3554 c2.,=,,(1 + \\ - 0.0471 where /~,,/& = Poisson 's ratios o f worm and gear materials, respectively; flW)(hF), flg)(hF) = influence factors o f the posit ion o f the loaded point in the radial direction o f the worm thread/gear tooth; and j2\u00a2tw)[~'o~.~ , hF),f~g)('~o, ~/F)----stress distribution factors along the worm thread/gear too th (shown in Figs l0 and l 1). The investigations have shown that the fillet stresses are propor t ional to the distance o f the loaded point to the too th root. Accordingly, the influence factors o f the position o f the loaded point in the rad ia l d i rec t ion o f the w o r m th read /gea r t oo th can be expressed by the fol lowing l inear equa t ions flw)(hF) = 0.1128 + 0.8872hFr (11) f]g~(hE) = 0.0287 + 0.9713hEr (12) where her = o The curves in Fig. 10 represent the stress d i s t r ibu t ions a long the w o r m thread for four different load positions in the radial direction of the worm thread, which are defined in Fig. 8. These curves can be interpolated by the following function: (w) f2 (?D, hF) = a + bV7 + cV) '2 (13) where V 7 = 17D - YF[; the angles VD and 7F have to be substituted in degrees. For load positions with hvr = hv/hp >>-0.5, the values of coefficients a, b, and c are given in Table 1. For load positions with hvf < 0.5, the coefficients a, b, and c have to be calculated by the following expressions a = a, + b, hF, + c, h2v,"
+        ]
+    },
+    {
+        "image_filename": "designv11_60_0002923_978-94-011-3514-6-Figure4.5-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0002923_978-94-011-3514-6-Figure4.5-1.png",
+        "caption": "Fig, 4.5 Nyquist graph of H(P) (with 11 0.4)",
+        "texts": [
+            " Thus, the velocity resonance occurs at w = woo x - w2 X cos(wt - ~) - A cos(wt - ~) A w2 X _~:::::::8:2::::::::~ w~ Xs 3A ap o =) ,(1 - ~2)2 + 4 ~2 p2 o Let us call ~3 the solution to this equation The acceleration resonance would appear therefore for an angular frequency W3 greater than wo wo W3 To summarize, a linear elementary oscillator has the following notable angular frequencies : { natural angular frequency without damping, angular frequency at the phase resonance, at the velocity resonance and at the power resonance (4.35) (4.36) - 64 - \",1 \"'0 ~ natural angular frequency with damping \"'2 \"'0 ~ angular frequency at the amplitude resonance \"'0 \"'3 angular frequency at the acceleration resonance They are classified in the following order W2 < W1 < wo < \"'3 not do this, it is in principle to \"'0 to which one refers. The curve of the figure 4.5 shows the complex frequency response H as a function of the relative angular frequency ~,for a constant value of the damping factor . One such curve is a Nyquist graph. According to the relations (4.25) and (4.26), H is given by the expression H (1 - ~2) + 2j~~ with tgtj) The curve approximates to a circle, especially when the damping factor is small (figure 4.7). It is shown plotted for increasing values of ~ from ~ = 0 (point s) to ~ = ~ (point 0) in figure 4.5. (4.37) - 65 - Let us designate the real and imaginary parts of H by a and b H = a + j b They have respectively the values {: 1 - ft2 (1 - ~2)2 + 41l2~2 -2 D ft (1 - ~2)2 + 41l2~2 Since the imaginary part is always negative, only the lower half of figure 4.5 is accessible, with the exception, however, of the half-circle of diameter OS, as we shall see below. The notable angular frequencies, listed at the end of the previous paragraph, correspond to the points !io, HI, Hz and !!3, in the order shown in the figure as a result of the inequalities (4.37). (4.38) (4.39) - 66 - For the point HD (corresponding to the phase, velocity and power resonances ~D = 1 ), the real part is zero when the imaginary part is ao = ;~ (= - ~ D) For the point H2 (amplitude resonance ~2 = ~ ), the modulus is equal to the maximum (4"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0002926_978-1-4684-6237-1-Figure9.10-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0002926_978-1-4684-6237-1-Figure9.10-1.png",
+        "caption": "Figure 9.10 Partially completed assembly",
+        "texts": [
+            " 9 A fr am e ba se d ro bo t s up er vi so r + +~ O B JE C T I , P O IN T , rO IN T T O P O IN T , D E P A R T A C T IV IT Y A C T IV IT Y /I L I C L O SE , , O P E N , , M O V E I S T R A IG H T 186 Intelligent robotics The other technique that is relevant concerns the use of a world model. Many features of the real world will be involved in errors. Such features, including elasticity of components, greasy surfaces and foreign objects in the world cell, are all extremely difficult to model in the idealized world. However, these concepts are necessarily important in diagnosis because we wish to find solutions to the recovery problem. Returning to our typical assembly task in Figure 9.1, we can examine a few of the real world happenings that must be considered. Figure 9.10 shows the assembly nearing completion. Geometric information about the size and shapes of the parts will be necessary in order to calculate whether a given component can be assembled to another component. In addition, some method of dealing with engineering tolerances must be included. If component F is to be able to clear both components E and C, then we must know the range of the tolerances on these three components, in order to know if there is any likelihood of the parts interfering. From our geometric database and a model of tolerances, we will be able to build automatically a network of conditions which determine how one component can constrain another component from reaching its required location"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0003013_a:1008115522778-Figure4-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0003013_a:1008115522778-Figure4-1.png",
+        "caption": "Figure 4. Corrective motion to compensate for misalignment by a robot.",
+        "texts": [
+            " And the cosines of the orientation error are given by the direction cosines between Lch and Lcp . The orientation error \u03b8z about the z-axis does not appear in cylindrical parts assembly due to their symmetry with respect to the z-axis. Consequently, the assembly errors in the assembly tasks of flexible cylindrical parts are composed of the aforementioned five parameters. To compensate for these errors, it has to be determined how much the end effector of a robot has to be moved. The robotic error-corrective motion to compensate for the assembly errors can be defined as shown in Figure 4. It is divided into the translational motion mr = [mre,mr\u03c6h,mr\u03c6v ]T and the rotational motion m\u03b8x , m\u03b8y . As described in Section 3.1, the rotational motion m\u03b8z about the z-axis does not have to be considered in cylindrical parts assembly because the error \u03b8z does not appear. To obtain the robotic error-corrective motion corresponding to the measured assembly errors is to obtain the relations between the five components of the assembly errors and the five components of the error-corrective motion"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0000659_acs.2709-Figure1-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0000659_acs.2709-Figure1-1.png",
+        "caption": "Figure 1. Body-fixed and Earth-fixed coordinate frames.",
+        "texts": [
+            " This section will give a brief overview of the definitions and the general marine vehicle model. Simplifications needed to derive the identification by \u00a9 2016 The Authors. International Journal of Adaptive Control and Signal Processing Published by John Wiley & Sons, Ltd. Int. J. Adapt. Control Signal Process. (2016) DOI: 10.1002/acs self-oscillations will be considered. Finally, propulsion and allocation models required to correctly achieve desired forces are presented. The marine vehicle dynamics and kinematics can be defined inside the inertial and body reference frames, as shown in Figure 1. The inertial or north\u2013east\u2013down frame, denoted as {E}, is the local tangent plane used for navigation. The frame can be treated as inertial owing to slower marine vehicle dynamics. The body-fixed reference frame, denoted as {B}, moves with the marine vehicle. The body-fixed frame origin is often defined at the rotational or gravity centre. Position and orientation, D x y \u00b4 , are defined in {E}, while linear and angu- lar velocities, D u v w p q r , are defined in {B}. Vector defines virtual forces and moments acting on the rigid-body in {B}"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0000622_b978-0-12-800782-2.00003-8-Figure3.31-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0000622_b978-0-12-800782-2.00003-8-Figure3.31-1.png",
+        "caption": "Figure 3.31 Induction motor torque for fundamental, 5th, and 7th time harmonics as a function of angular velocity and slip s1.",
+        "texts": [
+            " 240 Power Quality in Power Systems and Electrical Machines The relation between s1 (fundamental slip) and sh (1) (reflected harmonic slip) is s 1\u00f0 \u00de h \u00bc h\u03c9s1 \u03c9m \u03c9s1 \u00bc\u03c9s1 \u03c9m \u03c9s1 + h 1\u00f0 \u00de\u03c9s1 \u03c9s1 \u00bc 1 \u03c9m \u03c9s1 + h 1\u00bc h 1 s1\u00f0 \u00de: Therefore, the reflected harmonic slip, in terms of fundamental slip, is s 1\u00f0 \u00de h \u00bc s1 + h 1\u00f0 \u00de: (3-31a) For the backward rotating 5th harmonic one gets s 1\u00f0 \u00de 5 \u00bc s1 + 6 thus s1\u00bc s 1\u00f0 \u00de 5 + 6; and for the forward rotating 7th harmonic s 1\u00f0 \u00de 7 \u00bc s1 + 6 thus s1\u00bc s 1\u00f0 \u00de 7 6: Note that s 1\u00f0 \u00de 5 \u00bc 5\u03c9s1 +\u03c9m \u03c9s1 : In general, the reflected harmonic slip sh (1) for the forward and the backward rotating harmonics is a linear function of the fundamental slip s1: 241Modeling and Analysis of Induction Machines s 1\u00f0 \u00de h \u00bc s1 + h 1\u00f0 \u00de f or forward rotating field s 1\u00f0 \u00de h \u00bc s1 + h+ 1\u00f0 \u00de f or backward rotating field: (3-31b) Superposition of fundamental (h\u00bc1), fifth harmonic (h\u00bc5), and seventh harmonic (h\u00bc7) torqueTe\u00bcTe1+Te5+Te7 is illustrated in Fig. 3.31. At\u03c9m\u00bc\u03c9mrated the total electrical torque Te is identical to the load torque TL or Te1+Te5+Te7\u00bcTL, where Te1 is a motoring torque,Te7 is a motoring torque, andTe5 is a braking torque. Figure 3.32 shows the graphical phasor representation of superposition of torques as given in Fig. 3.31. 242 Power Quality in Power Systems and Electrical Machines In this section, the relation between the reflected harmonic slip sh (1) and the harmonic slip sh is determined as s 1\u00f0 \u00de h \u00bc f sh\u00f0 \u00de: The harmonic slip (without addressing the direction of rotation of the harmonic field) is sh\u00bc h\u03c9s1 \u03c9m h\u03c9s1 \u00bc h 1\u00f0 \u00de\u03c9s1 h\u03c9s1 + \u03c9s1 \u03c9m h\u03c9s1 with s1\u00bc\u03c9s1 \u03c9m \u03c9s1 : It follows that sh\u00bc h 1\u00f0 \u00de\u03c9s1 h\u03c9s1 + s1 h : With s 1\u00f0 \u00de h \u00bc s1 + h 1\u00f0 \u00de, one obtains for forward rotating harmonics s1\u00bc s 1\u00f0 \u00de h h 1\u00f0 \u00de or sh\u00bc h 1\u00f0 \u00de\u03c9s1 h\u03c9s1 + s 1\u00f0 \u00de h h 1\u00f0 \u00de h \u00bc h 1\u00f0 \u00de h + s 1\u00f0 \u00de h h h 1\u00f0 \u00de h \u00bc s 1\u00f0 \u00de h h : Similar analysis can be used for backward-rotating harmonics"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0002132_jsvi.1996.0655-Figure7-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0002132_jsvi.1996.0655-Figure7-1.png",
+        "caption": "Figure 7. A model of an unsymmetrical rotor\u2013bearing system.",
+        "texts": [
+            " Thus, unstable regions of the synchronous whirl and the 3\u00d7 whirl are obtained while, in the determination of the truncated 2T-type determinants, the minimum values are obtained to give approximate subcritical speeds of the even n\u00d7 whirls. Thus, the result is that there is no unstable region of subcritical resonances of even n\u00d7 whirls, and subcritical speeds of 2T-type motions reveal that there are response peaks in the free whirl which can be excited at the even n\u00d7 subcritical modes. It is also shown that subcritical speeds of the 4\u00d7 whirl are located at about half the values of the 2\u00d7 whirls. A complex rotor\u2013bearing system shown in Figure 7, having a non-axisymmetrical shaft (o=D/C=0\u00b75) and three axisymmetrical disks, is considered. Two non-dimensional parameters m1 = Id/(rCA) and m2 =md/(rAL), are utilized for the moment of inertia and the mass of these identical disks. The coefficients of both axisymmetrical bearings are assumed to be kxx = kyy = k and zxx = zyy = z, and the other coefficients are zero. When the parameters are k=10, m1 =14\u00b762 and m2 =7\u00b731, the dependence of the shaft asymmetry on unstable regions for various damping coefficients is shown in Figure 8",
+            " In the third case, with m1 =6\u00b767 and m2 =3\u00b733, the unstable region of the first flexural mode is eliminated from z=28 to z=52 as shown by chain curves. Also, the unstable regions of the P2f and P3f modes for three cases are shown by three types of transition curves without hatching. These instabilities can be completely eliminated when the damping coefficient increases to z=15 or z=20. For the modified transfer matrix approach, the accuracy of solutions is independent on the partitioned number of shaft elements. The least number is required. For example, the system as illustrated in Figure 7 is partitioned into four elements. When either the finite element approach or the conventional transfer matrix approach is utilized, the accuracy of the solutions may be improved by increasing the partitioned number. To achieve the same accuracy, the CPU time and memory size of both approaches are much more than that of the modified transfer matrix approach. For steady state solutions, the comparisons of accuracy and CPU times between the FEM and the modified transfer matrix approach are reported by Lee et al"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0003798_fie.1996.568510-Figure1-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0003798_fie.1996.568510-Figure1-1.png",
+        "caption": "Figure 1. A wide range of products were developed in the class. Shown above is the schematic drawing of a latch, cinch, and power lock system for an automobile door.",
+        "texts": [
+            " The class spans three academic quarters and the product development process which begins with a re-examination of the client\u2019s requirements is very similar to the product development process guidelines used in industry [8]. Projects The projects analyzed were taken from the 1992-93 academic year and dealt with a wide range of products including a catheter for gene therapy in the human body, a control mechanism for maintaining the focus of an infrared optical system, and a power locking device for an automobile door (Figure 1). A complete list of the projects and the class grades for the winter and spring quarters is shown in Table 1. Observe that the grades varied between A+ and B. Requirements Document While there was a general guideline for writing the document, there were noticeable variations in the format used by each project team. For example, one project team described the design requirements in a series of paragraphs, while another used a table format, and still another used a combination of both. Also, some teams used numbered lists and others used bullets"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0001235_b978-0-08-099425-3.00005-4-Figure5.48-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0001235_b978-0-08-099425-3.00005-4-Figure5.48-1.png",
+        "caption": "FIGURE 5.48",
+        "texts": [],
+        "surrounding_texts": [
+            "Cornering stiffness with load.\n(Courtesy of Dunlop Tyres Ltd.)\nThe modelling of the forces and moments at the tyre contact patch has been the subject of extensive research in recent years. A review of some of the most common tyre models was provided by Pacejka and Sharp (1991), where the authors state that it is necessary to compromise between the accuracy and complexity of the model. The",
+            "Aligning stiffness with load.\n(Courtesy of Dunlop Tyres Ltd.)\nauthors also state that the need for accuracy must be considered with reference to various factors including the manufacturing tolerances in tyre production and the effect of wear on the properties of the tyre. This would appear to be a valid point not only from the consideration of computer modelling and simulation but also in terms of track testing where new tyres are used to establish levels of vehicle performance. A more realistic measurement of how a vehicle is going to perform in service may be to consider testing with different levels of wear or incorrect pressure settings.\nOne of the methods discussed by Pacejka and Sharp (1991) focuses on a multispoke model developed by Sharp where the tyre is considered to be a series of radial",
+            "Lateral force Fy with camber angle g.\n(Courtesy of Dunlop Tyres Ltd.)\nspokes fixed in a single plane and attached to the wheel hub. The spokes can deflect in the radial direction and bend both circumferentially and laterally. Sharp provides more details on the radial-spoke model approach in (Sharp & El-Nashar, 1986), (Sharp, 1990) and (Sharp, 1993). The other method of tyre modelling reviewed is based on the \u2018Magic Formula\u2019 that will be discussed in more detail later in this section. Another review of tyre models is given by Pacejka (1995), where the influence of the tyre is discussed with regard to \u2018active\u2019 control of vehicle motion. The radialspoke and \u2018Magic Formula\u2019 models are again discussed. More recently, the"
+        ]
+    },
+    {
+        "image_filename": "designv11_60_0000036_ilt-04-2015-0057-Figure1-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0000036_ilt-04-2015-0057-Figure1-1.png",
+        "caption": "Figure 1 The structure diagram of point contact test bench",
+        "texts": [
+            " For friction pair under sliding condition, the average gap between two metal surfaces is closely related to the number of contact asperities and the actual contact area size, and the value of contact resistance can directly reflect the number of contact asperities and the actual contact area size, so, in principle, the lubrication state can be obtained by measuring the contact resistance. The upper sample used the G15 steel ball, polished surface. The lower sample used 45# steel disc specimens. The experiments were conducted on the home-made point contact friction and wear test machine, as shown in Figure 1, and the test parameters are shown in Table I (Wang, 2012). For the above equipment, the load change range is from 1.96 N to 19.6 N, the temperature range is from 100\u00b0C to 220\u00b0C, the speed range is from 300 r/s to 1,800 r/s and the range of resistance measurement is from 0.001 to 6 105 . Assuming electric current contraction region as an ellipsoid field, the following simple relationship between the shrink resistance and the conductive spot size can be obtained according to the contact resistance theory: Lubrication film thickness calculation De-Liang Liu, Shu-hua Cao, Shi-feng Zhang and Jiu-jun Xu Industrial Lubrication and Tribology Volume 68 \u00b7 Number 2 \u00b7 2016 \u00b7 176\u2013182 D ow nl oa de d by U ni ve rs ity o f K en tu ck y A t 1 8: 43 1 4 Ju ne 2 01 6 (P T ) Where: R contact resistance; r conductive spot radius; and resistivity of the contact element material"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0001945_ip-cta:19982047-Figure1-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0001945_ip-cta:19982047-Figure1-1.png",
+        "caption": "Fig. 1 Problematic stuil, =,(I)",
+        "texts": [
+            " However, the frequency of these jolts is finite (as opposed to the chatter phenomenon in SMC) and can be adjusted by proper selection of control IEE Proc.-Control Theory Appl , Vol. 145, No. 3, May 1998 parameters. This point is detailed in Section 3. We use a saturation function type smoothening in the control law (eqn. 6), against these jolts, defined as: where Sj is a boundary thickness. This type of control leads to another problem, however. The sufficient condition of eqn. 9 is not assured inside a problematic strip defined as: Fig. 1 depicts this zone of concern where q, = d 2 . ~ ~ l K ~ and = 425. As a remedy to this obstacle, a control strategy is proposed such that the state (eJ, g j ) is forced not to dwell inside this undesirable zone zjl) more than At; max seconds, a prescribed value. This is done by a strategy which yields a controlled acceleration of the error according to the following rule: (i) if ej < 0, enforce an acceleration e; > U; > 0 (introducing a positive variation in ej; upward direction in the phase portrait, j = 1, "
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0003425_physreve.63.016611-Figure9-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0003425_physreve.63.016611-Figure9-1.png",
+        "caption": "FIG. 9. Right: s versus n for a DNA ring with k50.057, g516k , and G52/3. Left: twist density gp(s) for this ring after a time t570 with Q51, z50.1, and b50.098 ~see next figure, bottom right!.",
+        "texts": [
+            "8,A3 and, therefore, this DNA minicircle is stable. Indeed, Haijun and Zhong-can @25# showed that the kink deformations observed in this DNA are caused by bending, and not by twisting. The effects of inertial forces, however, are better visualized on unstable rings. We therefore consider here an artificial excess of linking number equal to 16, corresponding to 100% of the natural twist, and a zero spontaneous curvature. Such a ring does not represent those in the experiment of Ref. @24#, but is just inspired by it. Figure 9 shows on the left the plot of s versus n for the above parameters. We see that the principal mode is n58. Figure 10 shows the principal mode of the perturbed DNA minicircle for various values of the perturbation amplitude z~0, 0.03, 0.07, and 0.1!. We see that, despite the very low value of Reynolds number, the effects produced by the inertial forces are the same as in Sec. V. The regions with larger mass density deform less than those with smaller densities. Figure 9 also shows on the right the twist density of the perturbed rod gp(s). Again, the amplitude of variation of gp(s) is of the same order as in the examples considered in Sec. V. We finally notice that, although the theory in Ref. @25# does account for the onset of instability of the DNA rings in the experiment of Han et al., it does not explain the asymmetric shapes exhibited by most of the kinked rings shown in Ref. @24#. Our results suggest that this asymmetry is due to nonuniformities in the rings, either intrinsic or induced by the binding of ions"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0002923_978-94-011-3514-6-Figure3.4-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0002923_978-94-011-3514-6-Figure3.4-1.png",
+        "caption": "Fig. 3.4 Oscillator consisting of wire of fixed length and two massless",
+        "texts": [],
+        "surrounding_texts": [
+            "In summary TO Xu ke Xo - mg o Let us now write Newton's law for the total displacement of the mass One has moreover m x' = mg - T' 2 T' x\u00b7 k x' I I k x' 2 2 ? (1(' + x') I 2 (3.11 ) (3.12) (1.13) (3.14) (1. 15) - 15 - and consequently x' => T' ke x'"
+        ]
+    },
+    {
+        "image_filename": "designv11_60_0003166_0168-874x(90)90028-d-Figure2-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0003166_0168-874x(90)90028-d-Figure2-1.png",
+        "caption": "Fig. 2. Characteristics of the space shuttle orbiter nose-gear tire used in the present study (h = h /ho) .",
+        "texts": [
+            " The procedure is based on transferring the anisotropic (nonorthotropic) terms (submatrices [Ka] in equations (8)) to the right sides of equations (12), and adding another level of PCG iterations to account for them. Numerical studies were performed to assess the accuracy of the two-dimensional shell model of the tire, and the effectiveness of the computational procedure described in the preceding section for generating the response associated with different harmonics. Herein, the application of the model and the computational procedure to the space shuttle nose-gear tire are presented. The geometric and material characteristics of the tire are given in Fig. 2. Measurements were made to determine the shape of the tire cross section, and the thickness variation. The cord-rubber composite was treated as laminated anisotropic material. The material properties of the different layers were obtained using the mechanics of materials approach, which has been widely applied to rigid composites [10,22]. Because of symmetry only half the tire cross section was modeled. It was divided into seven segments (see Fig. 2) with different number of layers, different material properties (corresponding to different cord content in the composite), and different fiber orientation. Spline interpolation was used to smooth out the measured data, and to obtain the geometric and material characteristics of the two-dimensional shell model. The outer surface of the tire was chosen to be the reference surface for the two-dimensional shell model. The meridional variations of both the geometric characteristics of the reference (outer) surface and the stiffness coefficients of the shell model are shown in Figs",
+            "K. Noor et aL / Analysis of aircraft tires (nonorthotropic) stiffness coefficients. Ero = 8 \u00d7 10 6 Pa. A.K. Noor et al. / Analysis of aircraft tires ~ = 0 = h i 227 using the geometrically nonlinear shell theory. Twelve finite elements were used in modeling half the cross section (a total of 384 strain parameters, 384 stress-resultant parameters, and 243 nonzero generalized displacements--see Fig. 5(a)). Comparison was made with the experimental data obtained on the shuttle nose-gear tire (see Fig. 2). The results are summarized in Figs. 6 to 8. Close agreement between the predicted deformations and experimental results is demonstrated in Fig. 6. Figures 7 and 8 show the meridional variations of the generalized displacements, stress resultants and strain energy densities. As can be seen from Fig. 8, for the case of inflation pressure, the transverse shear strain energy density is considerably smaller than the extensional/bending energy density. Pa. 228 A.K. Noor et al. / Analysis of aircraft tires A"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0002281_1.2831582-Figure1-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0002281_1.2831582-Figure1-1.png",
+        "caption": "Fig. 1 Winding rolls pressed by nip rollers",
+        "texts": [
+            " This study has application to the prediction of the amount of air entrained in a winding roll. Introduction It is important to exclude entrained air when a web wraps a heating or cooling drum or a surface treatment roller. In winding rolls of plastic-coated papers and other nonpermeable webs, it is necessary to control the maximum and minimum thickness of the film of entrained air. Winding of thin nonpermeable webs at velocities higher than about two meters per second usually requires a roller which compresses the winding roll to prevent excessive entrainment of air, as shown in Fig. 1. A roller used for this purpose is called a nip roller or rider roller or paclsroll. A nip roller may be wrapped by the web, or the winding roll may be wrapped upstream from the nip. For a web passing around an unnipped winding roll, Knox and Sweeney (1971) suggested that the air layer thiclcness may be approximated with the foil bearing equation, R, \\ T J (1) where u is the winding speed or the average velocity of two adjacent surfaces. The effects of a nip roller were reported by Bertram and Eshel (1980) and by Eshel et al",
+            " The purpose of this study is to obtain the air film profile and the pressure profile at the nip, and the total amount of entrained air. The winding roll and packroll are modeled as two rotating rollers as shown in Fig. 3. Note that the presence of web is neglected in this analysis. In application, the effect of the web Journal of Tribology JULY 1996, Vol. 1 1 8 / 6 2 3 Copyright \u00a9 1996 by ASME Downloaded From: http://tribology.asmedigitalcollection.asme.org/pdfaccess.ashx?url=/data/journals/jotre9/28520/ on 03/02/2017 Terms of Use: http://www.asme.org/about-asme/terms-of-use would depend upon the winding configuration (Fig. 1) and the web thickness. Also, it is assumed that the elastic properties of the winding roll and the packroll are isotropic, the process is isothermal so that the density of air is proportional to the abso lute pressure, the viscosity of the air is constant, and the inertia of the air is negligible compared to the viscous force. Procedure The solutions must satisfy both the Reynolds equation and the equation of material deformation. The Reynolds equation, based on the above assumptions, is \u2014 \\ph^ -J- + 6k\u201ep\u201eh dx dx dp dx 12/^M d(ph) dx (8) where the second term in the parenthesis indicates the effect of a slip boundary, which can become important when the air gap at the nip is not much larger than the mean-free-path of the air (Adams, 1989)"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0002678_1.2829318-Figure2-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0002678_1.2829318-Figure2-1.png",
+        "caption": "Fig. 2 Cross section to tlie rotating axis",
+        "texts": [
+            " Definition of the Geometric Elements Before considering the details of the load computation, it would be useful to introduce the different geometric elements that are considered and the suitable reference system, in order to reduce possible misunderstanding about their definitions. We consider that the main parts that compose the examined pump are well known, however it is possible to refer to (Mimmi and Pennacchi, 1995) and (Mimmi and Pennacchi, 1997) also for the mathematical modelling of the rotors, and we define as profile the generic contour line of any cross section to the rotation axis of the screws where the thread is complete (see Fig. 2). It is also important to choose a reference value for the rotor rotation angle ip (see Fig. 2) and we choose the initial value in Fig. 3 as zero. Notice also that with the adopted convention, the rotation of the central screw is counter-clockwise and that of the satellite screws is clockwise. Journal of Mechanical Design Copyright \u00a9 1998 by ASME DECEMBER 1998, Vol. 120 / 581 Downloaded From: http://mechanicaldesign.asmedigitalcollection.asme.org/ on 09/02/2013 Terms of Use: http://asme.org/terms Fig. 1 Screw pump rotors The profiles are composed of different arcs of curves (see Fig. 4). These are in particular arcs of circumference, cycloid or epitrochoid, whose analytical expression on the xy plane is known in parametric form"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0001777_951293-Figure1'3-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0001777_951293-Figure1'3-1.png",
+        "caption": "Figure 1'3. An Example of Dynamic Hysteresis of the Dry Friction Clutch",
+        "texts": [],
+        "surrounding_texts": [
+            "Assuming that one can record the relationship between friction forces and displacement, energy dissipated Ed can be measured as an area of the dynamic hysteresis loop (refer to Fig. 9). Of course, the value of energy Ed may be a function of many parameters such as the coefficient of friction, frequency and amplitude of motion, etc.\nStatic and Dynamic Hysteresis; Level of Participation Factor\nFrom an engineering point of view, especially if analysis of clutches and drivetrains is considered, it may be preferable to describe the dry friction damping, which is a highly nonlinear and dynamic phenomenon, in terms of static hysteresis. Let us assume that one can measure energy dissipated by the system and, for instance in a static condition, energy dissipated by a system is Edst. This system can dissipate energy of Ed in a dynamic condition and Ed is usually less than Ed,,. The ratio of dynamic and static energy losses,\ndefines a level of participation factor for the dynamic hysteresis. Factor K expresses actual energy dispersion as a fraction of static energy losses per cycle. The static hysteresis is a reference point in this particular approach. Static hysteresis H of a dry friction clutch is defined across the automotive industry as a double friction force,\nin the clutch damper. Using this, energy dissipated by the system in dynamic conditions can be expressed as\nEd = ~KIXIH ...( lo)\nand in any static case as,\nEdst = 21XIH (K = 1). ...( 1 1)\nAccording to the definition of K, the dynamic hysteresis can be expressed in terms of the static friction as,\nAn Engineering Model of Hysteretic Damping Difficulties in performing any reliable calculation of the clutch damper lie not only in the fact that friction is a nonlinear phenomenon but also in some serious singularities introduced by the dry friction forces. In the friction clutch for instance, a normal load which works well under one set of excitation conditions may cause the friction contact to lock up and does not dissipate any energy under other conditions.\nDry Friction Clutch Model According to the weak damping model introduced in Fig. 1 and 4a, a physical model of the dry friction clutch can be understood as a system of two bodies D and H (see Fig. 11) that represents the disc or discs lumped together with the clutch flywheel ring and engine flywheel (body D), and the hub or hubs of the driven disc lumped together with the transmission input shaft (body H). The spring constant k is the damper torsional stiffness and S is the excitation torque. Forces kx and F represent the elastic torque of the clutch\ndamper and the dry friction torque, respectively. The total force resisting the relative motion of bodies D and H is\nRF = - kxT F ...( 13)\nand exemplifies the reaction torque exerted by the input shaft on the engine flywheel. Let us consider that the clutch damper has a constant stiffness and experiences dry friction damping determined by different but constant (independent of the relative speed) static and dynamic coefficieiits of friction. If the damper relative movement is slow (static), one can present the resisting force as a function of displacement as shown in Fig. 12 (compare to Fig. 2). Now, let us assume that the above system osculates about a certain equilibrium position E (see Fig. 13) with certain considerable frequency, and that the amplitude of its motion in a steady state is 1x1. During the motion, the torquedisplacement relationship will be a closed curve as shown in\nFig. 13 (compare Fig. 9). One can expect that the size and shape of the hysteresis loop will depend on many parameters of the driveline such as: the amplitude and frequency of oscillations, location of the equilibrium position, friction characteristics, inertias of bodies D and H, stiffness of the damper, etc.\nTo identify and investigate significant parameters and reject secondary (negligible) modeling factors, many hypothetical relationships between potentially sensitive parameters must be examined in an experimental way.",
+            "Measurement of Dynamic Hysteresis Experimental determination of system damping in dry friction clutches has been based on measurements pertinent to the steady-state dynamic behavior of a certain dry friction system especially designed for this purpose. Particularly, damping characteristics of the system have been determined from the dynamic hysteresis curves using measurements and computations of energy dissipated for various operating conditions of the testing device. The test setup for the dynamic hysteresis measurement is presented in Fig. 14. Testing was performed on a testing specimen that was to simulate the dry friction clutch damper. The device consisted of a coil spring arrangement allowing to control the torsional stiffness of the test specimen and a friction pack which allowed to introduce various levels of dry friction. A great deal of attention was paid to control the specimen friction level and eliminate or minimize friction other than that designed by the friction pack. The spinning rotary shaker was used to rotate the specimen with a certain angular speed Q and load it with harmonic, torsional impulses having frequency o. The frequency o of the fluctuating torque was independent of rotating speed Q. The controller generated a single frequency, steady-state torsional input that was under displacement control. The torque fluctuations resulted from the displacement and torsional stiffness of the specimen. The torque was reacted into machine and the displacement feedback was continually monitored by the controller. The torque F transmitted through the specimen was the actual resisting torque RF (refer to equation (13)) consisting of both the elastic and friction components. The specimen torque F and the relative displacement x of the testing device were measured and recorded in each experiment as functions of time. Both Ithe unfiltered and filtered output signals were processed to obtain dynamic hysteresis loops of the friction device. The filtered hysteresis loops represent the pure harmonic response of the specimen to the single harmonic excitation of the same frequency that was generated by the shaker. In the present paper, all the energy dissipation investigations were based on these signals. It is important to understand that the unfiltered signals consist of some internal disturbances introduced by the shaker, controller, and measuring devices themselves. Those undesirable harmonics are small in comparison to the basic signal corresponding to the excitation frequency and can be neglected in the fmal energy balance. Of course, this type of idealization can be and usually is a source of some quantitative errors in the final mathematical model.\nTest Matrix To the knowledge of the authors of this paper, the present effort is the first, full scale effort to model dynamic hysteresis in clutches. Hence, various hypothetical situations had to be considered and a relatively large test matrix had to be used. The most important parameters investigated experimentally included: levels of static friction pN in the\nfriction pack, rates of the torsional stiffness k of the specimen, rotational speeds R, amplitudes of torsional displacement 1x1, and frequencies w of the torsional excitation.\nResults of Hysteresis Testing\nResults obtained from the first sequence of tests are to be interpreted as preliminary. Because the test matrix had to be large, test repeats for each particular parameter had to be quite limited and the experiments were performed only for specific values of parameters. For instance, static friction torques were chosen for the experiments from the medium and high range of those used in the industry. Because of this, the experimental results must not be extrapolated to low static hysteresis cases. Moreover, one should remember that the results presented here are valid only for the specific test specimen that was used in this experimentation. Authors believe that damping characteristics obtained in the present project can change, if applied to clutches, and they will probably vary from one clutch damper to another depending on particular designs. The damping ability of the test specimen is determined by the amount of energy dissipated per cycle of the simple harmonic motion. The area of the dynamic hysteresis loop which is equivalent to energy dissipated per cycle was measured during hysteresis testing. Fig. 15 shows an example of tests performed for a sequence of increasing frequencies of excitation when the other parameters of the system are fixed. Note the change in the orientation of the dynamic loops (clockwise rotation) with frequency. This may be attributed to decreasing phase angle between the torque F and displacement x. Based on the empirical results, energy dissipated per cycle and per unit of amplitude was described as a function of the amplitude of relative motion X and the circular frequency of excitation w as,",
+            "48 dB/Octave < - Unfiltered Hysteresis\nSignal Processor H Unfiltered K p i h i h\n489"
+        ]
+    },
+    {
+        "image_filename": "designv11_60_0000088_j.ifacol.2015.06.369-Figure1-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0000088_j.ifacol.2015.06.369-Figure1-1.png",
+        "caption": "Figure 1. The 3-RPR planar parallel robot",
+        "texts": [
+            " One ordered by Grey code, which holds the traversal path stands in the surface of a hypercube; the other ordered by B(2,6) De Bruijn sequence, which has the shortest path through all terminals. In this work, we use the Grey sequence in the optimization algorithm. Notice that we also use l instead of u here. The optimization process is shown in Figure 5, and we have the results max 4.717l  and min 2.989l  . INCOM 2015 May 11-13, 2015. Ottawa, Canada Rui Zeng et al. / IFAC-PapersOnLine 48-3 (2015) 1930\u20131935 1935 L A x y al O a b c a bc bl cl B C D EF l Figure 1. The 3-RPR planar parallel robot The absolute difference of Cauchy index between two points in the iteration is never bigger than 2. And the optimization result is surrounded by non-degenerate singularity surfaces. The relationship between the iteration and the planning results is presented by Figure 3 also, in which branch optimum method is shown clearly to deduce the maximum of max minl l . The iteration step numbers in Figure 3 are not namely equal to the computing time as the time loss of each step isn\u2019t presented here"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0002601_004051759106100901-Figure1-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0002601_004051759106100901-Figure1-1.png",
+        "caption": "FIGURE 1. Schematic diagram of Eltex ATHS air-jet texturing machine: DZ = drawing zone, TZ = texturing zone, MSZ = mechanical stabilizing zone, HSZ = heat stabilizing zone, and WZ = winding zone.",
+        "texts": [
+            " The raw material characteristics for both yam groups are given in Tabk I. ,\u2019 TABLE 1. Raw material data for polyester filament yarns. \u00b0 Yarns shown are fully drawn. Sd = semiduU finish, Bt = bhght finish. Textured yarns were produced using a HemaJet with T 100 core at 300 m/min speed on Eltex AT/HS ma- at UNIV OF CALIFORNIA SANTA CRUZ on January 4, 2015trj.sagepub.comDownloaded from chine. The yam path on this machine and the points for measuring delivery and stabilizing tensions are shown in Figure 1. Air pressure of 9 kg/cm2, mechanical stretch of 196, and winding underfeed of 0.7 were used, as were overfeeds of 25% for group I and 33.3% for group Il yams. Wet textured yams were produced with the same parameters as dry textured yarns, except that yams were prewetted before entering the jet with a HemaWet wetting head using 0.51 water per hour at a pressure of 2 kg/cm2. Figure 2 shows the spin finish application technique we used to apply spin finishes to the feeder yarns. The feeder yam passed through guides and a guide pulley, which positioned the yam on the lick roll partially im- mersed in the spin finish solution in a trough, from where the yam went to the winding unit"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0000617_s1064230716030126-Figure3-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0000617_s1064230716030126-Figure3-1.png",
+        "caption": "Fig. 3. Kinematic diagram of the manipulator mechanism.",
+        "texts": [
+            " In accordance with this, a vector of the generalized coordinates , i = I, \u2026, III is introduced for each mechanism, where index i indicates the unit number, which includes the mechanism. Index i = M corresponds to the real space manipulator. Let us consider the structure of vectors in more detail. The current position of the mechanism, which is a system of interconnected bodies-links is defined by the joint coordinates , j = 1, 2, \u2026, ni, i = I, \u2026, III.1 They are the angles of rotation and/or linear displacement of one ith link with respect to the previous (adjacent) (j \u2013 1)th link. These coordinates can be combined in the ni-dimensional vector of joint coordinates . Figure 3 shows a kinematic diagram of one of the mechanisms, i.e., a manipulator mechanism for moving the full-scale model of the space manipulator gripper. The joint coordinates are changed by actuators, the current status of which is determined by the ni-dimensional vector of the coordinates of actuators di. Its components are the angles of rotation of the motor shafts. The relationship between these vectors is represented by the following relationships:2 , (2.1) , (2.2) where is the ni-dimensional vector of the rotation angles of the actuator\u2019s motor shafts reduced to joint angles and linear displacements, is the ni-dimensional vector of the elastic deformations of the gears, and is the ni-dimensional continuously differentiable invertible vector function of reduction transformations",
+            " The deformation of all nII links is determined by the coordinates of 6knII elastic elements that approximate these links and are combined into the vector l = . It is necessary to note that the arm (Unit I) and the manipulator mechanism (Unit III) are equipped with wrist force-torque sensors. The supporting structures of these sensors are connected with the joystick at the end of the arm held by the operator\u2019s hand and with the working tool mechanism, respectively. The elastic deformations of the sensor\u2019s supporting structure with respect to its base part combined with the last link of the arm and the mechanism are wrist coor- dinates (Fig. 3). They form six-component vectors and of the arm and manipulator mechanism. Thus, the current state of the above-mentioned mechanisms is represented by the following vectors: ( , , , , ) with the dimension of for the arm. ( , , , , ) with the dimension of for the space manipulator. ( , , , , ) with the dimension of for the manipulator mechanism that moves the full-scale model of the gripper. However, since the ni-dimensional vectors , , , are related by two vector equations (2.1) and (2.2), in order to describe the current state of the ith mechanism, it is sufficient to use two of them, such as gi and \u03c4i"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0000333_0047239515626704-Figure2-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0000333_0047239515626704-Figure2-1.png",
+        "caption": "Figure 2. Individual connection.",
+        "texts": [
+            " The Virtual Steel Sculpture The interactive virtual steel structure is an accurate 3D replica of the physical steel structure located next to the Civil Engineering building on the Purdue University campus in West Lafayette, IN. The sculpture was modeled to scale and can be observed from any point of view. Figure 1 shows the full model of the sculpture. The interface allows for 360 rotation of the view, zooming in and out, 434 Journal of Educational Technology Systems 44(4) panning, and tracking. In addition, students can click on the individual connections to observe them more closely (Figure 2). Every connection is identified by a number; by clicking on the number, students can access additional information such as limit states and real world examples (Figure 3). The interactive steel structure is deliverable via web or CD-ROM on standard personal computers (PCs and Macs). The platform for the project is based on the highest end in 3D interactive animation. We used Autodesk Maya software to model and texture the structure and to animate its functionality. Interactivity with the 3D components was programmed in C# using Unity 3D (2013) game development platform"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0000856_s0025654416040014-Figure1-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0000856_s0025654416040014-Figure1-1.png",
+        "caption": "Fig. 1.",
+        "texts": [
+            " Considering these integrals as new variables, we can reduce the equations of motion of the rigid body to a form which is convenient for the application of the integral manifold method [5]. This type of equations of motion of a rigid body is not new. For example, similar equations were used in the asymptotic analysis of equations of motion of a rigid body in the atmosphere [6, 7]. The rigid body orientation with respect to the cable direction is determined by the following three Euler angles: the precession angle \u03c8, the nutation angle \u03b1, and the intrinsic rotation angle \u03d5 (Fig. 1). The integral manifold method is used in the nonlinear case, which, in contrast to [8], permits analyzing the rigid body motion without the assumption that the nutation angle is small. Of course, it is assumed in this case that the cable does not touch the surface of the rigid body, which imposes certain constraints on the nutation angle depending on the rigid body geometry. For example, for a nearly spherical rigid body (or a rigid body with a spherical tip), the nutation angle does not exceed \u03c0/2"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0002948_jae-2002-486-Figure1-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0002948_jae-2002-486-Figure1-1.png",
+        "caption": "Fig. 1. The concept of the thin steel plates transported system with magnetic levitation technology.",
+        "texts": [
+            " \u2217Corresponding author: Susumu Torii, E-mail: torii@ee.musashi-tech.ac.jp; Fax: +81 3 5707 2212. 1383-5416/01/02/$8.00  2001/2002 \u2013 IOS Press. All rights reserved In case of the steel plate transporting system with magnetic levitation technology, the object of the transportation is very thin. It is difficult to support the weight of the thin steel plate with one electromagnet in magnetic levitation system because of the effect of magnetic saturation. Thus, plural electromagnets support the thin steel plate in the magnetic levitation system. The Fig. 1 shows the concept of the thin steel plates transported system with magnetic levitation technology. When the magnetic levitation control system support the thin steel plate by the electromagnets, vibration on the plate occurs on action point because of the magnetic levitation technology is unstable as the levitate direction. When the thin steel plate is levitated by the plural electromagnets, there are interactions of supporting forces as the action points. It is necessary to consider the dynamic vibration"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0002002_s0094-5765(99)00125-3-Figure2-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0002002_s0094-5765(99)00125-3-Figure2-1.png",
+        "caption": "Fig. 2. Kinematic descriptions of an arbitrary link.",
+        "texts": [
+            " Each truss boom is treated as a separate body or link, and its \u00afexibility is modelled using the \u00aenite element method (FEM), such that only axial deformation of its members is considered (truss members can take only axial load). The prismatic actuators are modelled as separate links, consisting of a cylinder component and an extending piston rod, which is typical of the basic geometry of a hydraulic or servo cylinder actuator. The actuator members are considered as rigid elements. The spatial position and orientation of an arbitrary link i with respect to an inertial reference system Xo, Yo, Zo is illustrated in Fig. 2. A body-\u00aexed frame Xi, Yi, Zi is de\u00aened on each individual link and situated at its revolute base, such that axis Zi corresponds to the axis of rotation. Angle yi is used to represent the rotation about the positive Zi-axis. The prismatic actuation of a piston-rod is de\u00aened to occur along the rod's Xi-axis. Vector pi represents the position vector of the origin of link frame i relative to that of the inertial frame. The inertial position vector of any point on link i is denoted by si, given by si pi ri, 1 where ri is the position vector of the point relative to the link frame origin"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0003339_s0022-460x(87)81301-9-Figure3-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0003339_s0022-460x(87)81301-9-Figure3-1.png",
+        "caption": "Figure 3. System composed of a disc and two crank-and-rocker mechanisms.",
+        "texts": [
+            " If the Fourier expansion of the function '1'CtP) is composed of harmonics having only odd tP multipliers, then the Fourier expansion of the coefficients in equation (9) contains only harmonics having even multipliers of cPo Consequently, the resonances associated with oi] (2wa.) = II m and wi(wa , +wa) = II m exist only for even m. For the linear equation (9) one may find the regions where resonance vibrations may start. The amplitude of the resonance vibrations can be determined only from the set of non-linear equations (5). The system considered, shown in Figure 3, is composed of a disc and two crank-and rocker mechanisms. The external torque MAl = Cl(n -dtPJdt) is applied only to the disc AI; M A 2 = M A ) = 0 and M B>= M B> = O. For the mechanism shown in Figure 3 the angle of oscillation of the rocker is given by ,--::----;;:----;; sin 1JI' = (0102 - c 3Jo~+ o~ - oDIC o~+ CD, (24) where O2=== 2( L2- t; COS lfJ)I L\" 0 3 = 2(L1+ Lc sin lfJ)1L; The matrices A, B, C, S, K, D and H in equation (9) have the forms A = diag (AI, A 2 , A 3), B= diag (0, B2, B3 ) , C= diag (C I , 0, 0), [ S, -82 -~'l K~[~ 0 -~'lS= -:2 82+83 K3 -83 83 -K3 K3 [ D, -D2 -~'l H~[~ 0 -~'lD= -~2 D 2+D3 H3 (25) -D3 D3 -H3 H 3 Detailed computations were carried out for Lei L, =0\u00b72, LI L, =1, L 21L, = 1, Ld L, = "
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0002728_112515.112577-Figure13-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0002728_112515.112577-Figure13-1.png",
+        "caption": "Fig. 13 Algebraic constraints forfi[sz contact. ,,",
+        "texts": [
+            " (5)Fim2: Constraints forfirw arerepeatedly given for two CVS, Suppose CFJ and CF2 represent CFS at two CVS. Let tll\u2019,lEl\u2019 and ACI are transformations for CFI and tn\u2019, tm\u2019and Ac2 are for CF2. If the direction of the X axis of CF2 in the world is in the negative direction of the X axis of CFI, m\u2018 and tm\u2019 are changed to M * tn\u2019 and M * tE2 \u2018 . ACI = A[&,O,&, &I, C5yl,6ZI] and Ac2 = A [&,0,42,&,&2,&2] where dil = h, & = &, 61 = &2, and & = - &. If the angle between the Y axes of CFJ and CF2 is q, the following constraints are additionally given (see also Fig. 13): e= ~,d=21r-r\u2019lc0s6 (28) where lx is signed distance from the origin of CF2 to the origin of CFI along the X axis of CF1. (6)Fim: Ac = A[&O,d,&,O,O] (see Fig. 14). 5.2 Inequality Constraints When CVS are touching the extended feature, other vertices of the internal feature must be locate above the extended feature to avoid the interference (see Fig. 3). This condition is interpreted as inequalities. Suppose a vertex V of the internal feature and FI = origin({V ) ). FE is determined in the same manner for the equation constraints"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0003547_0954406001523074-Figure2-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0003547_0954406001523074-Figure2-1.png",
+        "caption": "Fig. 2 Pitch outlines in the form of \u00aerst-order ellipses, with the base outlines arrived at by meeting the",
+        "texts": [
+            " The terms `involute', `base outline' and others will from here on be applied to loci generated from noncircular pitch pro\u00aeles. Under these circumstances, these terms will not meet all the constraints associated with their use with circular gears. Four combinations of constraints will be settled upon, although others could have been selected. The resulting loci will be given the analogous names that are applied with circular gears, with appropriate quali\u00aecations where the context does not make their special nature clear. Figure 2 shows four identical pairs of \u00aerst-order ellipses rolling on each other. The LAs are arrived at by four di erent means which, in turn, generate four different base outlines. In Fig. 1, for circular gears, neither the angle between the line of action and the tangent at the pitch point (PAT) nor that between the line of action and the line of centres (PAC) changes as the gears rotate. In non-circular gears that is not the case, and only one of these angles can be \u00aexed. It was decided to \u00aex the \u00aerst and then the second angle, PAT or PAC, in turn and to examine the outcome",
+            " It was therefore decided to examine separately these two terminating conditions: one where the line of action begins and ends where it is normal (-N) to radii from each gear centre, and the second where it is tangent to some envelope (-E). The tangency position was determined to be where two successive lines of action intersect. This is based on the principle that, aside from discontinuities, two in\u00aenitely close tangents to a planar curve will intersect on that curve. These end conditions will give rise to two di erent base outlines. In summary, \u00aexing the angle and the end-points of the line of action, each by two means, leads to four di erent `base outlines'. These are shown in Fig. 2 for pitch outlines that are simple ellipses: PAC-E, PAT-E, PACN and PAT-N. The points on the \u00aerst-order ellipses shown in this paper were arrived at directly by calculations from equations of simple ellipses. All higher-order ellipses could also have been calculated but were arrived at by a process of iteration. In this method a known outline was used to generate numerically its `rolling mate'. This method was developed in order to cope with a broad range of outlines. In the \u00aegures shown here, the initial pitch outline is drawn on the left and is called the driver or driving pitch outline",
+            " This is the conventional means of de\u00aening the pressure angle for circular and non-circular gears and is equivalent to the action of rolling a straight rack around the pitch outline. Figures 2b and d di er only in the method used in arriving at the end-points of the line of action, which has been mentioned above but will be discussed more fully below. Figures 3a and b show qualitatively the direction and location of a PAT line of action in a frame \u00aexed with respect to the gear centres. The angle of the conjugate line, CA, measured from the LC is given by CA TA\u00ff PAT The pressure angle PAT and the tangent angle TA are indicated in Fig. 2b. Since these are non-circular pitch outlines, the common tangent at the pitch point will oscillate about the vertical as the pitch point moves to and fro between the gear centres. As a consequence of this, a line that makes a \u00aexed angle to this tangent will also oscillate as it moves with the pitch point. For the single-order ellipses in Fig. 2, the largest di erence in the direction of the conjugate line will be at the mid-point of the line of centres. In the half-cycle shown in Fig. 3a, at the mid-point the conjugate line will be at the maximum angle: CAmax TAmax \u00ff PAT In the next half of the cycle (Fig. 3b), again at the midpoint, the conjugate line angle CA will be minimum: CAmin TAmin \u00ff PAT C02197 \u00df IMechE 2000 Proc Instn Mech Engrs Vol 214 Part C at OhioLink on November 7, 2014pic.sagepub.comDownloaded from If CAmin reaches 0 , which is depicted for the two teeth shown in Fig",
+            " The complementary pairs for rotation in the opposite direction are, for the cases used here, mirror images about the horizontal axis of symmetry of the pitch pro\u00aeles. For circular gears there is no di erence between these interpretations of the length, location and direction of the line of action. Put into more practical terms, there is no di erence in the outcome if a generating rack is set at some angle to the tangent patch, or at the complementary angle to the line of centres. However, there clearly is a di erence for non-circular gears. The PAT-E base outline shown in Fig. 2b permits useful involutes to be unrolled, as was done by Baxter. Note that the base outlines indeed appear to be confocal ellipses as reported by Olsson [1]. However, note also that the generation of involutes by unrolling a tangential cord to higher-order ellipses, as in Fig. 4b, leads to seemingly impossible situations. It may be seen from this \u00aegure that not only do the base outlines move in and out of the pitch outlines but they exhibit discontinuities. To understand the reason for these discontinuities, it is necessary, for example, to examine the curvature of the second-order ellipses shown in Fig"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0001146_jctn.2015.4303-Figure8-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0001146_jctn.2015.4303-Figure8-1.png",
+        "caption": "Fig. 8. Z thermal deformation on upper surface at 20 r/min.",
+        "texts": [],
+        "surrounding_texts": [
+            "AND INITIAL CONDITIONS The 3D model of h drostatic bearing table is shown in Figure 1. This paper built the finite element model of worktable by using the software named ANSYS Workbench. Because the rotary worktable has periodicities, it just needs to mesh 1/24 of one single cycle. As shown in Figure 2. The initial conditions and boundary conditions are necessary to solve specific problems. The center of the worktable which is studied by this paper is touching radial cylindrical roller bearing, so it is applied by Fixed Support code to the fixed boundary conditions in inner surface of the circumference. The distribution of corresponding J. Comput. Theor. Nanosci. 12, 3917\u20133921, 2015 3919 Delivered by Publishing Technology to: Chinese University of Hong Kong IP: 117.253.218.180 On: Fri, 26 Feb 2016 00:22:32 Copyright: American Scientific Publishers R E S E A R C H A R T IC L E Fig. 9. Z thermal deformation on upper surface at 80 r/min. temperature field must be applied to the simplified model to calculate thermal deformation field, and its specific command is Thermal Condition code. As same as temperature field, it needs to use Commands code to apply symmetric boundary conditions to both sides of the model. The specific settings are shown in Figure 3."
+        ]
+    },
+    {
+        "image_filename": "designv11_60_0002074_1.1359772-Figure2-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0002074_1.1359772-Figure2-1.png",
+        "caption": "Fig. 2 Test rig for the identification of dynamic coefficients",
+        "texts": [
+            " It turns out that the accuracy of the damping coefficients is APRIL 2001, Vol. 123 \u00d5 383 001 by ASME 16 Terms of Use: http://www.asme.org/about-asme/terms-of-use Downloaded F improved considerably when measurements are carried out at backward whirling frequencies, too. The dynamic coefficients in this paper are identified by a least square fit using the changes in dynamic behavior at minimum three forward and three backward whirling frequencies. The test apparatus employed in numerous identification runs is displayed in Fig. 2. The diameter of the rotor carried in hydrodynamic bearings is 23 mm. The spacing between the bearing and, accordingly, the whirling frequency varies from 770 to 890 mm or 26 to 35 Hz, respectively. The test-seals are located by pairs symmetrically to the inflow region in the midspan position of the shaft. The prerotation of the flow entering into the seals is generated with a guide vane ring. High amplitudes of the unstable rotor are avoided with catcher bearings. As the magnetic bearing is not placed at the same axial position as the test-seal, a conversion procedure is used ~@12#"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0002558_jahs.36.57-Figure2-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0002558_jahs.36.57-Figure2-1.png",
+        "caption": "Fig. 2 Definition of rariahles aud coordinate axes for inertial and body framcs.",
+        "texts": [
+            " The tethers are modeled as massless rigid links and all the attachment points arc assunicd to be frictionless pin joints. These assumptions are identical to thosc o f Ref. 6 and lead to seven degrees of freedom for the system. From a controller design point of view, i t is useful to express thc equations o f motion for this sytem i n terms o f the following nine variables instead of just seven: p , . 2,. >I,. 2 1 , y,.. z,. en. Q,. +?. A definition of thcsc variables along with the coordinate systems uscd for the inertial and body rcfcrcncc frames are included in Fig. 2. The total force equilibrium of the systeni along the z and ?. inertial directions leads to the following two cquations: MI;, + M,i, + (M, + ML)& - M,H(&, sin E, + E; cos F,,) = - (MI + M2 + MD + ML)R NONLINEAR STATE FEEDBACK 59 Since the tether attachrncnt points at the sprcadcr bar arc assumcd to bc frictionless pin joints. thc morncnts about thcsc points should vanish. This results in thc followin_r two cqu;~lions. (Z, + Z,,, - M 1 - 1 = - M , , - J,. - 11, sill L . (z, - ZL - A , cos 6, + - sln E, - H cos E,) = 0 (3 ) 2 (Z"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0001457_2015-01-2190-Figure1-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0001457_2015-01-2190-Figure1-1.png",
+        "caption": "Figure 1. Driveline units of test vehicle a) Engine and transmission b) Front sub frame c) Rear sub frame with RDM unit",
+        "texts": [
+            " Both the front and rear sub frames are attached to the body through 4 elastomeric mounts each. The power transfer unit, drive shaft and independent rear differential assembly are installed to meet the functional requirements of AWD. The vehicle is equipped with McPherson strut type suspension. The drive shaft (or propeller shaft) connects power transfer unit and rear drive module (RDM). The RDM is attached to rear cradle which in turn connected to body. The driveline units of test vehicle are illustrated in Figure 1 OPAX approach differentiates from the existing methods in the identification of the operational loads. It uses parametric model characterization of the operational forces as a function of measured path inputs (mount accelerations) and stiffness of the mounts. The parametric model is built from the data pertaining to operational path accelerations across mounts, few indicators, target response signal and frequency response functions from the passive side locations to targets and indicators. The outline of OPAX method [1] is illustrated in Figure 2 The description for each notation in Figure 2 is given below"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0000837_red-uas.2015.7441004-Figure1-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0000837_red-uas.2015.7441004-Figure1-1.png",
+        "caption": "Fig. 1. Pure pitching motion",
+        "texts": [
+            " With the previous considerations, we obtain the model by applying the Newton\u2019s laws of motion. The dynamic model in order to control the altitude of the MAV is given by [11]: \u03b8\u0307 = q (1) q\u0307 = Mqq +M\u03b4e\u03b4e (2) h\u0307 = V sin(\u03b8) (3) where V is the magnitude of the airplane speed, \u03b8 denotes the pitch angle. q is the pitch angular rate with respect to the y-axis of the aircraft body, h is the airplane altitude and \u03b4e represents the elevator deviation [11]. In aerodynamics Mq and M\u03b4e are the stability derivatives implicit in the pitch motion. We can see these variables in the in Figure 1. The lateral dynamic generates the roll motion and, at the same time, induces a yaw motion (and vice versa), then a natural coupling exists between the rotations about the axes of roll and yaw [2]. In our case, to solve it, we have considered that there is a decoupling of yaw and roll movements [12]. Thus, each movement can be controlled independently. Generally, the effects of the engine thrust are also ignored [2]. In the Figure 2, the yaw angle is represented, which can be described with the following equations: \u03c8\u0307 = r (4) r\u0307 = Nrr +N\u03b4r\u03b4r (5) where \u03c8 represents the angle of yaw and r denotes the yaw rate with respect to the centre of gravity of the airplane"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0003947_rtd2002-1642-Figure2-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0003947_rtd2002-1642-Figure2-1.png",
+        "caption": "Figure 2: Friction surface on the axle-box",
+        "texts": [],
+        "surrounding_texts": [
+            "The representation of the friction forces following the classical Coulomb\u2019s law. Often this formulation lead to numerical simulation problem, due to the discontinuity introduced by the friction force behaviour. To avoid this problem the friction force has been modelled with a continuos function of the relative velocity between the friction surface described by the following equation: 2 1       \u22c5 \u22c5 + \u22c5 = \u00b5 \u03c7 \u03c7 N V V F f (1) \u00b5 \u03c7 \u22c5\u00b1= \u22c5= \u00b1\u221e\u2192 \u2192 NF VF fV fV lim lim 0 Where Ff is the friction force, V is the relative velocity. The \u03c7 parameters represent the angle between the velocity axis and the friction force curve around the origin, \u00b5 is the kinetic friction coefficient, N is the force normal to the friction surfaces. In the case of the Y25 bogie the friction surfaces allow the relative motion on the plane which include the vertical and the lateral direction (see fig. 3), so it is necessary to extend the friction force formulation to two degrees of freedom. The two components of the vertical and lateral velocity are z& and y& ; the absolute value of the relative velocity is: 22 yzV && += (2) The components of the velocity are ( )\u03b8cos\u22c5= Vz& and ( )\u03b8sinVy \u22c5=& , while the components of the friction force in the z and in the y direction are: ( ) ( ) 22, 11 cos cos       \u22c5 \u22c5 + \u22c5 =       \u22c5 \u22c5 + \u22c5\u22c5 =\u22c5= \u00b5 \u03c7 \u03c7 \u00b5 \u03c7 \u03b8\u03c7 \u03b8 N V z N V V FF fZf & (3) ( ) ( ) 22, 11       \u22c5 \u22c5 + \u22c5 =       \u22c5 \u22c5 + \u22c5\u22c5 =\u22c5= \u00b5 \u03c7 \u03c7 \u00b5 \u03c7 \u03b8\u03c7 \u03b8 N V y N V sinV sinFF fZf & (4)"
+        ]
+    },
+    {
+        "image_filename": "designv11_60_0001948_5326.760575-Figure9-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0001948_5326.760575-Figure9-1.png",
+        "caption": "Fig. 9. Link assignment of the reclaimer.",
+        "texts": [
+            " The candidates of the feasible borders are saved for later use and the procedure starts from the candidates when the borders get lost. Fig. 8 shows the extracted contour map that can be used to determine a landing point. III. INVERSE KINEMATICS OF A RECLAIMER The autonomous reclaimer is servoed in the joint-variable space, whereas the points on the pile are expressed in the world coordinate system. In order to control the position of the bucket to reach the points, the inverse kinematics solution should be solved. The disk-like end-effector has an action to produce another degree of freedom when it lands on the pile. Fig. 9 shows the assignment of link frames for the reclaimer. A Cartesian coordinate system XiYiZi can be established for each link at its joint axis with the reference coordinate system X0Y0Z0: and are the fixed angles between the boom and the rotating disk by the mechanical structure. helps the reclaimed ore fall onto the conveyor belt on the boom. Let (xd; yd; zd) be a point on the pile. We would like to find the corresponding joint angles, 2 and 3; and translation distance d1 of the reclaimer so that the bucket can be positioned as desired"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0000471_icuas.2016.7502516-Figure3-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0000471_icuas.2016.7502516-Figure3-1.png",
+        "caption": "Figure 3. Photo of propulsion unit on a test stand (zoom on AXI 2814/12 GOLD LINE BLDC motor wtih GWS-HD9050x3-SW 9x5\" propeller)",
+        "texts": [
+            " MODEL OF PROPULSION UNIT From a variety of BLDC motors and propellers (in various configurations and sizes) currently used in multirotor flying robots created in the Institute of Control and Information Engineering of the Poznan University of Technology, for further work on the FOPID type speed controller synthesis one selected AXI 2814/12 GOLD LINE BLDC motor from Model Motors company with three-bladed propeller GWS-HD9050x3-SW 9x5\" due to its dynamics properties, i.e. providing a relatively large thrust (nominally up to 2 kg) at relatively low rotational speeds and low power consumption. This set was mounted on a specially designed and built test bench (Fig.3) described in more detail in [15]. That article presents a description and results of acquisition works by the use of dedicated software DYNO Terminal. From conducted tests it is known that the maximum thrust (19.09 N) of the tested propulsion unit is achieved at PWM=77 %, RPM=8893 r/min and the current is equal to I=15.551 A, but the maximum speed RPM=9039 r/min is achieved at PWM=75 %, and I=14.419 A. Useful thrust starts from minimum value 3.924 N (RPM=4492.5 r/min). The rotational speed corresponding to the gain of 1 in below models is equal to the value of 9550"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0002252_s0890-6955(98)00006-6-Figure15-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0002252_s0890-6955(98)00006-6-Figure15-1.png",
+        "caption": "Fig. 15. The constitution of modified FHH.",
+        "texts": [
+            " dB = S1S2\u00b7cosae(coturef 2 tanae) (44) VB2 = PP9(coturef1 2 tanaS2) (45) VB3 = CC1(coturef2 2 tanae) (46) whereuref1 is the side relief angle on the main cutting edge (Fig. 13)uref2 is the side relief angle on the front cutting edge (Fig. 14) A modified cutting force model included the effects of size, shape and tool edge wear, is presented in this study and it based on the force model in Fig. 4. Besides the (FH)Umin force, the plowing force FP, due to the effects of tool edge specification, and wear force FW, from the effects of flank wear, are considered to predict the horizontal cutting force (Fig. 15) [14]. That is: (FH)M = FHH(FH)Umin + FP + FW (47) Lf1 = d/cosCs + (l1 + l2)tanCs 2 h1 2 h2 + (2R1uR1 + PC + 2R2uR2 + 2R3uR3)/cosaS2 + (50) [fCM 2 (l1 + l2 + l3 + l4)]/cosaS2\u00b7cos(Ce 2 Cs)] Lf2 = 2R1uR1/cosaS2 + d/cosCs + (l1 + l2)tanCs 2 h1 2 h2 (51) Lf3 = 2(R2uR2 + R3uR3)/cosaS2 + [fCM 2 l1 + l2 + l3 + l4)]/[cosas2cos(Ce 2 Cs)] (52) where Lf1, Lf2 and Lf3 are the contact lengths between the cutting edge and workpiece, as shown in Fig. 5; HB is the Brinell hardness of workpiece; r is the roundness of tip between the main cutting edge and flank; R is the tool nose radius and dB is the tool worn depth (Fig",
+            " Once the shear area (A) and friction area (Q) are obtained, the shear energy (Us) and the friction energy (Uf) can be calculated according to Eqs (30) and (31). By using the values of Us and Uf, the theoretical principle component of cutting force, (FH)Umin , was obtained from Eqs (32) and (33). The theoretical vertical cutting force (FV) and the theoretical transverse cutting force (FT) were obtained from Eqs (34)\u2013(38). The flank wear force (FW) and plowing force (FP) must be taken into account when it is not zero. The modified cutting forces along the three axes, FHH, FVV and FTT, shown in Fig. 15, were obtained from Eqs (47)\u2013(60). The values of theoretical, modified theoretical and experimental results for each of FHH, FVV and FTT, were plotted in Fig. 20(a,b,c) and Fig. 21(a,b,c). The computational flow chart is illustrated in Fig. 18. The results shown in these figures imply that: 1. Fig. 20(a,b,c) indicates that among the experimental, modified and theoretical cutting forces, there is a good agreement between the experimental values and modified results, for the wear tool geometrical configuration, l1 = 0"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0000457_978-81-322-2671-0_78-Figure2-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0000457_978-81-322-2671-0_78-Figure2-1.png",
+        "caption": "Fig. 2 Experimental quadcopter model",
+        "texts": [
+            " Two motors M1and M3 will rotate in clockwise direction, while the other two motors M2 and M4 rotate in anticlockwise direction. Two motors adjacent to each other are always in the opposite direction of rotation. Thrust produced by motors should be twice that of the total weight of the quadcopter. If the thrust generated by the motors is too little, the quadcopter does not take OFF. However, if the thrust is more than the design level, the quadcoptor might become too nimble and hard to control. BLDC electric motor used to produce thrust along with the propeller is shown in Fig. 2. DJI 2212 low weight BLDC motor produces 920 rpm/V; a peak power of 370 W is used in the modeling [8]. A propeller is a type of fan that transmits power by converting rotational motion into thrust. Four propellers of span 10 in., with fixed pitch angle of 4.5\u00b0 are used in modeling [9]. Diagonal propeller pairs spin in the same direction and also have opposite tilting combination producing lifting thrust. Power Bank LiPo battery is used as a power source for quadcopter due to high power to weight ratio [10]"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0003648_app.1989.070380903-Figure1-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0003648_app.1989.070380903-Figure1-1.png",
+        "caption": "Fig. 1, Schematic drawing of bioelectrode.",
+        "texts": [
+            " Different power levels were used to generate GLUCOSE OXIDASE MEMBRANE 1593 the plasma which was used to treat the PP film. The treated films were taken out under N, atmosphere, and then washed with deionized water. The washed films were dipped in 25% glutaraldehyde for 12 h, washed again with deionized water, and soaked in GOD solution. The obtained films were used for the preparation of biosensors. Preparation of Biosensors The biosensors were prepared by combining the enzyme immobilized membranes with the electrode (Fig. 1). RESULTS AND DISCUSSIONS Preparation of Immobilized Glucose Oxidase Membrane via the Plasma-Initiated Polymerization Method The effect of plasma power on the activity of immobilized glucose oxidase is shown in Figure 2. The maximum activity was obtained when the plasma power was 40 W. When the plasma power was higher than 40 W, the activity of immobilized glucose oxidase decreased as the power increased. This might be due to the destruction of the glucose oxidase by the high energy particles. When the plasma power was 20 W, no activity of immobilized glucose oxidase was detected"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0003685_itsc.1997.660569-Figure9-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0003685_itsc.1997.660569-Figure9-1.png",
+        "caption": "Figure 9: Bond giraph for reverse gear",
+        "texts": [
+            " In the torque phase, the dynamic constraints on the third gear range are satisfied with the 1-2 band clutch pressure being not zero. The torque phase of the 3-2 shift is bond I i T I I r wCif RGi Rsr 1 Wsr graphed in Figure 10. K 1 Wt Tt+ 1 1 h T t - 7 OF-TF- 1 Figure 10: Bond graph for torque phase of 3-2 shift In the reverse gear, the turbine torque is supplied to the i.s.g. by the first clutch, and the i.c.g. is held to ground by the reverse band, and the r,s.g. free-wheels. This word model is bond graphed in Figure 9. 3.2.2 Speed phase. The speed phase model of the 3-2 shift is identical to the 2-3 shift model. In Figure 10, if the third clutch pressure decreases and the turbine is disconnected to the i.s.g., the shift model becomes as Figure 4. 3.3 Fourth-to-third shift model In the 4-3 shift, the i.s.g. rotates with being reduced. As mentioned in the 3-4 shift, the third clutch transmits power through the one-way sprag. Therefore, Figure 7 is the bond graph when the fourth clutch pressure is not zero and the speed of the turbine is equal to that of the i"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0002728_112515.112577-Figure6-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0002728_112515.112577-Figure6-1.png",
+        "caption": "Fig. 6 Algebraic treatment of position errors.",
+        "texts": [
+            " Accordingly, we can detertrtinefww. We then determine the positions of the axes of FPE based on the axes of FNPE. Since FP.E and FNPE are in the same part, the following constraint equation can be derived (9)W *fPE *fNPE-l = W * tP where tp is a transformatiort which transforms the axes of FNPE into the axes of FPE with respect to the frame of FNPE. rP is determined according to the positional error type and FD. Here we give an example of tP. Example: tP for position errors of the axis of a pin (see Fig. 6). When FNPE is properly positioned by referring to the datum, tpfor position errors concerning YE and ZE directions of the datum ix tP = rot(X,9 * (Z + A [0,uj,4,0,0,0]) * ro@,-@ (10) where rot(X, 9 rotates the axes of FNPE to orient its Y and Z axes in parallel to YE and ZE directions. Small displacements by uj and d place its origin on the origin of FPE. Finally, r&(X,-@ orients its axes to the axes of FPE. If position tolerances Ty and TZ are given in YE and ZE directions, the following inequalities must be satisfied"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0000688_j.asr.2016.08.002-Figure2-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0000688_j.asr.2016.08.002-Figure2-1.png",
+        "caption": "Fig. 2. Illustration of the CONSCAN movement.",
+        "texts": [
+            " The work is divided as follows: the beam deviation estimation technique is described in Section 2, the control law is presented in Section 3, the simulations are presented in Section 4, and the conclusions are presented in Section 5. The theory about the CONSCAN presented here is based on Gawronski and Creparo (2002). The CONSCAN consists in the addition of harmonic movements in both axes, azimuth and elevation, making the antenna to perform a circular pattern while follows the spacecraft, as represented in Fig. 2. This movement is circular, with radius r and angular velocity x. In Fig. 3, the frame of reference is presented. The origin represents the original antenna path, sk the spacecraft position, s\u0302k the estimated spacecraft position, ak the antenna position during the scan, and ek is the difference vector ystem for satellite tracking based on Kalman filtering and model pre016/j.asr.2016.08.002 between the spacecraft and the antenna positions. The sub index k refers for the k-th sample performed at the time tk \u00bc k T s"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0003042_iros.1993.583851-Figure14-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0003042_iros.1993.583851-Figure14-1.png",
+        "caption": "Fig. 14: a result of sliding toward the fringe of the table",
+        "texts": [],
+        "surrounding_texts": [
+            "We have implemented the system that plans motions to escape from situations where there are no possible grasp using sliding operations. Figs 12 through 14 show the results obtained applying our method. Fig.12 and Fig.13 illustrate the same problems but actions are different,. Fig.12 shows a result of combination of actions lb-2a-3a, and Fig.13 shows one of la-2a-3a. The problem in Fig.l2(a) is that the object which is the smaller of the 2 rectangular parallelepipeds cannot be grasped because of an obstacle. Fig.l2(b) shows the motion this system has planned, which is to slide the obstacle in order to make the object graspable. Fig.l2(c) shows the st,ep where the manipulator can grasp it after sliding. Fig.l3(a) illustrates the same problem as Fig.l2(a). Fig.l3(b) shows the motion which slides the object in order to make it graspable. Fig.l3(c) shows the step where the manipulator can grasp it after sliding. Fig.l4(a) illustrates the problem where a flat and wide rectangular pahllelepiped object cannot be grasped. Fig.l4(b) shows the motions which is to slide the object toward the fringe of the table in order to make it graspable. Fig.l4(c) shows that the manipulator grasps it after sliding it, at the fringe of the table."
+        ]
+    },
+    {
+        "image_filename": "designv11_60_0001235_b978-0-08-099425-3.00005-4-Figure5.82-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0001235_b978-0-08-099425-3.00005-4-Figure5.82-1.png",
+        "caption": "FIGURE 5.82",
+        "texts": [
+            " As with any model based on a physical discretisation, the model refinement or number of cross-sectional elements must be such that the width of the \u2018slices\u2019 is sufficiently small to deal with obstacles that are narrow compared with the overall Actual tyre Tyre model Discretisation of tyre profile for durability analysis. width of the tyre. The road model is based on the finite element representation described in Section 5.6.3. The algorithm developed carries out initial iterations to identify road elements that are subject to potential contact, at the current integration time step, before evaluating the position of each tyre element slice with each of the candidate road elements. An example of this is shown schematically in Figure 5.82 where one tyre cross-sectional element is seen to intersect a step defined by a number of triangular road elements. For each of the discrete elements used to model the tyre cross-section, the interaction with the road surface elements produces a line projection of the intersection on the tyre element. From this it is possible to compute the area and hence volume related to the penetration of tyre cross-sectional element by the road, for example by summing the three components shown in Figure 5.83"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0001049_cgncc.2016.7828850-Figure1-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0001049_cgncc.2016.7828850-Figure1-1.png",
+        "caption": "Figure 1. Four rotor hover system",
+        "texts": [
+            " Problem statement, adaptive SMC design, dynamics approximation using RBFN and stability analysis of the system is given in Section III. Simulation results are described in Section IV. Finally the conclusion of the proposed work is presented in Section V. The 3-DOF four rotor hover system is an electro-mechanical experimental system made by Googol technology [22], used to implement and test flight controllers for VTOL vehicles. The control algorithms developed for the system can easily use for other configurations of MUAV. Fig. 1 shows the laboratory based experimental hover system. The four-rotor hover system platform consists of four propellers connected to a gimbal ring named as front, back, left and right propellers. The gimbal is mounted on a pivot joint that permits the vehicle to move about roll, pitch and yaw axes. The torque generated by front, left and right propeller motors allows the hover system to move about pitch axis, the left and right propeller motors causes the roll movement of the system, while the back propeller motor moves the system about its yaw axis"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0001427_978-3-319-06596-0_36-Figure4-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0001427_978-3-319-06596-0_36-Figure4-1.png",
+        "caption": "Fig. 4. Example distribution of strain in bone-fixator system: a) Taylor-type fixator, b) tibia bone",
+        "texts": [
+            " Based on the obtained results it was found that the biggest displacement values were observed in the bone fracture along Y axis and averaged 0.99 mm - Fig.3. These values didn\u2018t exceed the permissible value of the fracture gap about 1 mm, which determines the correct bone union. It was found additionally that maximum strain and stress for the maximum force F=1500N were observed for connecting elements telescopic strut with holder constrained in fixator full ring (frame) and equals \u03b5max = 0.42%, \u03c3max = 838MPa respectively - Table 1, Fig.4, 5. The values of stress and strain in other parts were much smaller and equals \u03b5max = 0.21%, \u03c3max = 416MPa for wire, \u03b5max = 0.10%, \u03c3max = 200MPa for half pin and \u03b5max = 0.49%, \u03c3max = 91MPa for bone respectively - Table 1. Obtained in the analysis of stresses don\u2018t exceed the yield strength of the metal biomaterial (Rp0,2 = 690MPa) for loads of the range F = 100-1400 N - Fig.6. It assures fixator elements strain in elastic range. However, the load force F > 1400N following border crossing point of Rp0,2 frame material which in turn can cause damage"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0002982_iros.1994.407427-Figure1-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0002982_iros.1994.407427-Figure1-1.png",
+        "caption": "Fig. 1. Relationship between gradient vectors and the null s p a c e o f J a t W ) . (a)zh=O.(b)zh+zgipi=O.",
+        "texts": [],
+        "surrounding_texts": [
+            "The basic motion task is to trace a circle trajectory, which is represented as (27).\nX( t ) = -l.0c0~(2a t ) + 3.0 (27)\ny ( 2 ) = -1.0sin(2n;t)\nLet an additional task be to maximize the manipulability measure9. H(0)=- , for avoiding singularities. The initial configuration is [-1 .521,1.951,1 .353Ir which corresponds to [2.0.0.0]r in task space. Let the other additional task be an obstacle avoidance. Generally, the minimum distance between workspace obstacles and a manipulator can be expressed as a function of 8. For simplicity, suppose that a point of interest on the manipulator is the third joint. Then. it is required that the third joint avoid a workspace obstacle. Let the obstacle be a circle with radius of fl and center of [4.3,-3.0IT. We need to solve the following optimization problem:\nnraninize H ( e ) = m (28)\nsubject to x - f(0) = 0\nThe results are depicted in Figs.4 and 5. For comparison, the results for the case that does not consider obstacle avoidance are also depicted. In Fig.4, the manipulator configurations and the forbidden region , the interior of the circles , are shown. It is observed that the manipulator successfully performs additional tasks as well as the basic motion task. Getting a solution 'with disregarding the inequality constraint is blocked by G at t =0.36. which means that the third joint of the manipulator contacts with the boundary of the forbidden region at this time and the solution of the next time step must be obtained along that boundary ,i.e., the higher priority 'must be given to the obstacle avoidance task. The change of the Lagrange multiplier is depicted in Fig.S(b). The Lagrange multiplier becomes zero at I =0.54, meaning z h=O at that configuration. At this point, it is possible to get a solution of the next time step that satisfies zh=O, without being blocked by the inequality constraint. This means that the active inequality constraint can be relaxed, so a solution can be obtained with disregarding G , In other words, the higher priority is given to singularity avoidance, at this time step. The proposed method maximizes H (e) with the inequality constraint being satisfied, i.e., the obstacle is avoided. In Fig.S(c). the manipulability measure for both cases are depicted.\nIn the second simulation, the basic task is same as before and the manipulability is again chosen to be maximized for the task of singularity avoidance. The obtained solution trajectory performing only these two tasks is the same one that is obtained by the Extended Jacobian method, and is shown in Figd(a) as a curve. Suppose that the requirements for another additional tasks are\nexprtsstd by the following inquality constraints:\n(29) G ,(e) = -(e2 - 1.72 -(e3 - 1.312 + 0.22 s o G2(e) = -(e2- 1.64)~ -(e3- 1.17)~ + 0.1~ I o\nThe boundary of each inequality constraint is also shown in Fig.qa). The region enclosed by these two inequality constraints is forbidden for the manipulator. Note that DOR is not 'sufficient' , i.e.. DOR is smaller than the number of the additional tasks. Then the redundancy resolution problem can be stated as follows:\n(30) maximize ~ ( e ) = G i E \" j subject to x - fie) = 0\n-(e2 - 1.7)~ -(e3- 1.3)~ + 0.2~ s o -(e2- 1.64)~ -(e3- 1.17)~ + 0 .1~ I o\nThe solution of (30) is shown in Figs.6(b)-(d). In Fig.6(b). the solution trajectory is plotted in & 4 3 space. The upper right end point corresponds to the task space point of [ZO,O.O]r, which is the desired position for the end-effector at t =O and t = 1.0. At t = O S , the end-effector must be located at [4.0.O.OIT and the lower left end point corresponds to that position In Figs.6(b) and (c), we can find the fact that the proposed method has a cyclic property. Figsd(c) and (d) show the time history of the solution. For comparison, the solution 10 the problem without considering inequality constraints is also plotted in Fig.6(c). The solutions are recursively obtained by using (20) until it is blocked by the inequality constraint G1 at f =0.09. From this h e , solutions are obtained along the boundary of GI = 0 by using (23). The change of the corresponding Lagrange multiplier pI is shown in Fig.qd). At about t=0.25, a blocking inequality constraint G2 is encountered. At that time instance, the proposed method relaxes the inequality constraint G I , therefore making pI be zero , and modifies the value of p2 at the corresponding configuration. Then, it gets solutions along the boundary of G2=0 until t =0.33, at which p2 becomes zero so that the corresponding configuration becomes another switching point. At this point. zh=O and the method again relaxes the inequality constraint G 2 and thereafter gets solutions that maintain zh=O until any blocking inequality constraint is encountered.\n4. CONCLUSION In this paper, we showed that a redundancy resolution problem with multlple criteria could be transformed into a local equality and inequality constrained optimization problem, and proposed a method to solve it at velocity level. In the proposed scheme, equality constraints are imposed for the basic motion task. One additional task is performed by optimizing an objective function while the other additional tasks are performed by satisfying a set of inequality constraints. Since the proposed method uses the differential kinematic relationship and the \" ices involved are always square, the computation can be efficiently done. The method is efficient especially when the number of additional tasks",
+            "are larger than DOR. It also gives a way to systematically assign the priorities between the additional tasks by using the Lngrange multipliers. In addition, the method has a cyclic property which is mcial in cyclic tasks. Some limitation8 are also examined along with these benefits of the proposed method.\nRE\"as\nD. E. Whitney, \"Resolved motion rate control of manipulators and human prosthesis.\" IEEE Trans. on Man-Muchine Systt?ms, vol. 10, no. 2. pp. 47-53. 1969.\nD. N. Nenchev. \"Redundancy resolution through local 0ptimization:a review,\" J. qfRobotic Sptems, vol. 6, no. 6. pp. 769-798, 1989.\nJ. Baillieul, \"Kinematic programming alternatives for redundant manipulators,\" in Proc. IEEE Int. Conf. on Robotics and Automan'on, pp. 722-728, 1985.\nJ. Baillieul, \"A constraint oriented approach to inverse problems for kinematically redundant manipulators,\" in Proc. IEEE Int. CO@. on Robotics and Automation, pp.\nP. H. Chang. \"A closed-fon solution for inverse kinematics of robot manipulators with redundancy.\" IEEE J. of Robotics and Automation, vol. 3, no. 5. pp. 393-403, 1987.\nH. Seraji, \"Configuration control of redundant manipulatomtheory and implementation,\" IEEE J. of Robotics and Automation, vol. 5, no. 4, pp. 412-490, 1989.\nH. Seraji and R. Colbaugh. \"Improved configuration control for redundant robots,\" J. of Robotic Systems, vol. 7, no. 6, pp. 898-928, 1990.\nC. R. Carignan, 'Trajectory optimization for kinematically redundant arms,\" J . of Robotic Systems, vol. 8, no. 2, pp.\nT. Yoshikawa, \"Manipulability and redundancy control of robotic mechanisms.\" in Proc. IEEE Int. Conf. on Robotics and Aufomation, pp. 1004-1009, 1985.\nD. Luenberger, Optimization by vector space methods, Wiley & Sons, 1969.\n1827-1833, 1987.\n221-248, 1991.\nFig. 2. Adjustment of the Lagrange multiplier pi at 0(k) whcn getting B(k+l) is blocked by the inequality constraint Gj 10.",
+            "1, = 3.0\nI , = 2.5 I3 = 2.0"
+        ]
+    },
+    {
+        "image_filename": "designv11_60_0003173_s0141-6359(02)00117-4-Figure17-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0003173_s0141-6359(02)00117-4-Figure17-1.png",
+        "caption": "Fig. 17. Experimental setup for measuring torsional stiffness.",
+        "texts": [
+            " Sample results are shows in Figs. 15 and 16. Fig. 15 is a plot of displacement magnitude in an aluminum model of the perpendicular adjustment moment carrying variation; displacement is expressed in units of milli-inches. Fig. 16 shows strain energy for the three-dimensional printed perpendicular adjustment non-moment carrying variation. These figures show good strain and stress distribution indicating a robust design. Physical models were tested to validate analytical and finite element models of torsional stiffness. Fig. 17 shows a test fixture that was constructed to permit the testing of three-dimensional printed couplings, as well as an aluminum model fabricated specifically for the tests. Except where noted, all structural components were aluminum, and all fasteners and bearings were steel. The base of the test fixture was a 25 mm (1\u2032\u2032 nominal) thick plate, resting on 4 legs. To this plate were bolted two bearing blocks and a coupling mounting block. The bearing blocks were mounted approximately 13 cm (5\u2032\u2032) apart, and supported a shaft with 25"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0002229_jsvi.1996.0017-Figure1-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0002229_jsvi.1996.0017-Figure1-1.png",
+        "caption": "Figure 1. Bead on a rotating wire.",
+        "texts": [
+            " 7 1996 Academic Press Limited Many applications in engineering and physics lead to the model ordinary differential equation x\u0308=p(x)x\u0307+f(x)+G cos nt, (1) where p(x) and f(x) are 2p-periodic functions of x. These include the rotating pendulum, governors and phase-locked loops in circuits (see references [1\u20133]). We shall examine here the mechanical analogy which can be interpreted as either the rotating pendulum or the motion of a bead on a rotating circular rigid wire of large inertia. Without damping or forcing, the base example is a simple illustration of bifurcation in a nonlinear system (see reference [4], p. 26). From Figure 1, we can see that, in the unforced case, the equation of motion for the angle u is ma2u =ma(V2a cos u\u2212g) sin u+map(u)u , (2) where a is the radius of the wire, and m is the mass of the bead in the rotating wire analogy. It will be assumed that p(u) is p-periodic, and contains both damping and self-exciting terms of the form p(u)=\u2212k+b(1\u2212cos 2u). (3) 231 0022\u2013460X/96/020231+18 $12.00/0 7 1996 Academic Press Limited Other periodic functions are possible: the important requirement is that p(0) and p(p) are both negative"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0002018_s0094-114x(97)00044-x-Figure1-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0002018_s0094-114x(97)00044-x-Figure1-1.png",
+        "caption": "Fig. 1. Parallel manipulator of the generalized Stewart platform type.",
+        "texts": [
+            " The problem is formulated in the domain of complex numbers, the general optimization methods are used for the solution.' The presented method is very easily applicable on the manipulators of arbitrary structure, it has practically unlimited potential of parallelization. \u00a9 1997 Elsevier Science Ltd 1. INTRODUCTION The generalized Stewart platform is a fully parallel mechanism with six degrees of freedom. It is used in flight and automotive simulators, robotic end-effectors, and other applications requiring spatial mechanisms with high structural stiffness. The geometry of the Stewart platform is shown in Fig. 1. A great attention has been paid to the forward displacement analysis of this type of manipulator in the recent years. The kinematic problem can be stated as follows: given the lengths of the 6 variable limbs, find all the possible positions of the movable platform (end-effector of the manipulator). For some simplified types of the parallel manipulator, it is possible to determine the number of solutions of the forward displacement analysis (direct kinematic problem) by means of geometric theorems [1]",
+            " The complications were probably caused by the very strong non-linearity of the objective function arising from the products of complex goniometric functions containing exponential members (sin(cPRE + iCplM) = sin(q~RE)COsh(q~XM) + i COS(~PRE)Sinh(tpIM)). The effort was therefore concentrated to formulate the objective function on the basis of cartesian coordinates of points and their mutual distances. Eighteen optimization variables were chosen, particularly real and imaginary components of cartesian coordinates of points B~, B2 and B3 (see Fig. 1). The objective function was created in such a way that its zero values (global minima) correspond to the situation, when the kinematic constraint equations were satisfied. The position of a rigid body is uniquely given by the position of its three non-collinear points (Bt, B2, B3). Then the positions of points B4, Bs, B6, have to be uniquely expressed by means of position of triangle Bj B2B3. If this condition was not satisfied, our objective function would be valid for more different platforms simultaneously. The following relationships hold between the points of basic triangle [BIBzl 2 - ~ = 0,1B~B312 - ~ = 0,1B2B31' - ~ = 0. (1) The unique position of the auxiliary point B with respect to the triangle BI, B2, B3 is given by (see Fig. 3 and Fig. 1) It = B~ + (B2 - B,) x (B3 - Bt)/x/d, d2 sin/3,. (2) Forward displacement analysis of the generalized Stewart platform 247 The division by the constant ~/dld2 sin(fit) is introduced because of consistence of physical dimensions. The derivation of the linear dependencies for the positions of points//4, B5 and//6 comes from the procedure used for the description of multibody systems by means of so-called natural coordinates [7]. We shall use the following notation: B, B:, j = 1 , . . . , 6--position vectors of centers of spherical couplings and of auxiliary point in the system x, y, z ~0, B~o, j = 1 ",
+            " If the number of global extremes, and thus the maximum possible number of configurations of the given manipulator type, is to be found reliably, it is necessary to perform the whole above described calculation for more different manipulators of different dimensions. If the same number of global extremes is calculated repeatedly (10 times) for the series of manipulators of given type with randomly generated dimensions, then the maximum number of configurations of this type of manipulator can be considered to be reliably determined. The set of manipulators of randomly generated dimensions and states of drives was investigated. We have used the following strategy for the random generation of manipulators. The points A~ and Bi (see Fig. 1) were randomly generated in the 3D cube (x ~ ( - 10, 10); y E ( - 10, 10); z ~ ( - 10, 10)). The dimensions of manipulator described in Section 2 were computed from the random values of the cartesian coordinates of these points. The subsequent global optimization search passes through an 18th dimensional cube in the space of optimization variables. Several different regions of optimization variables were sought through for each of these investigated manipulators. In particular we used (( - 15, 15) x ( - 15, 15) x "
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0000860_978-3-319-44735-3_1-Figure1.8-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0000860_978-3-319-44735-3_1-Figure1.8-1.png",
+        "caption": "Fig. 1.8 With high force, the hexacopter goes up",
+        "texts": [],
+        "surrounding_texts": [
+            "The proposed controller implements a closed loop that comprises the three layers. Data produced as output in one layer is passed as input to the next layer. Fig. 1.9 With low force, the hexacopter goes down The proposed multi-layer fuzzy controller is based on [12] and is depicted in Fig. 1.10. The Control box is composed by a pre-processing phase (first layer), a set of fuzzy controllers (second layer), and post-processing phase (third layer). As one can observe, after the post-processing phase, the control outputs are applied onto the plant by means of the hexacopter rotors that actuate on the hexacopter movement and stabilization. The sensors perceive the changes on the plant controlled variables, and hence, provide the feedback to the controller. The controller, in turn, compares these input valueswith the reference values established as setpoints thereby closing the control loop [8]. The pre-processing phase (first layer) is responsible for acquiring data from the input sensors, process the input movement commands, as well as calculate the controlled data used as input to the fuzzy controllers in second layer. Before the multilayer controller starts its execution, there is an initialization phase that is performed within the first layer. The target position is set as the current position, so that the hexacopter does not move before receiving any command. Gyroscope and accelerometer sensors are calibrated and the GPS sensor is initialized by gathering at least four satellites. During the execution phase, the first layer is responsible to calculate the input variables to the fuzzy controllers: (i) the angular and linear distance (delta error) for X, Y, and Z axes between the current hexacopter position and the target position; (ii) the rotation and translation movement matrices to translate 3 axes movement into the speed related to the ground (i.e. X and Y axis). In addition, it is responsible to convert the input movement commands into setpoints for X, Y and Z positions. Movements commands are composed of three values representing the positive or negative movement along X, Y and Z axes related to the current positions, i.e. a command indicates a relative position. Thus, when a new command is received, the first layer will convert it to a absolute position. Then, when the control system is executing, this layer uses the GPS coordinates to determine the error in the distance from the hexacopter to the target position. These calculated errors in position are the inputs to the fuzzy controllers (Euler X, Euler Y and Euler Z errors). The second layer contains five fuzzy controllers, which act on issues regarding the hexacopter movement, namely hovering stabilization, vertical and horizontal movement and heading. As mentioned, these controllers take as input the data produced in the first layer and generate output for the third layer. The generated outputs represent the actuation on the six rotors for performing pitch, roll, yawmoves for all maneuvers necessary to reach the target position. The fuzzy controllers are discussed in details in the next section. The post-processing phase (third layer) is responsible for coordinating the fuzzy controllers outputs. As mentioned, in order to perform a proper maneuver, the proposed multi-layer controller establishes a priority on movements needed to complete a maneuver. When a new command is received, i.e. a new target point is set, the hexacopter must firstly reach the target altitude. Then, the hexacopter must turn until its front aims the target position. Finally, the hexacopter moves horizontally towards the target position. This layers also performs a threshold limits control by means of output values saturation, in order to keep the hexacopter stability while flying or hovering."
+        ]
+    },
+    {
+        "image_filename": "designv11_60_0003708_6.1989-3529-Figure1-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0003708_6.1989-3529-Figure1-1.png",
+        "caption": "Fig. 1. Sketch of the polar robot.",
+        "texts": [
+            " This paper studies one of the fundamental maneuvers of a polar coordinate robot and develops a full understanding of the switching structure of the optimal control. The class of problems considered are ones in which the initial and final robot states could be connected by simply retracting or extending the prismatic joint. For simplicity, robot motion is restricted to the horizontal plane so that only the motor for the prismatic joint, and the motor for rotation about the vertical axis are used (see Fig. 1). Results by the direct method [2,3] show similar behavior when rotation in the vertical plane is also allowed. A general property of the time-optimal maneuver is that at all times there is at least one actuator that is acting at full strength (bang-bang control). In our case the radial motor has the largest burden in the control task, and i t always operates at its force limit. During periods when other motors can aid this one, for example by generating centrifugal force when this is helpful, the other motors will also be at the saturation limit, but at other times these motors may be on a singular arc",
+            " Because of the prechosen switching structure, and the proximity of some of the switching points in these problems, the most appropriate approach is a variable mesh technique [ll] treating the switching points/final time and the switching conditions as parameters and interior point conditions, respectively. For the iterative solution of our MPBVP we use the code OPCON [8,12], a GGN method based on a multiple shooting algorithm. A stable QR solver has been implemented to solve the linearized constrained least squares problems. Robot Motion in the Horizontal Plane In this section we introduce the polar robot shown in Fig. 1. We restrict ourselves to motion in the horizontal plane in order to make the analysis easier, but i t will be seen that the problem is still complex. The results obtained are directly applicable to 3 degree of freedom cylindrical coordinate robots executing planar maneuvers. On the other hand, they can be thought of as motion of a spherical polar coordinate robot restricted to motion in the horizontal plane. The way in which the results would generalize when the restriction is eliminated is indicated by the direct method results in [3,4], and the numerical methods developed here can also be used to handle this more complicated problem",
+            " Its \"naturai coordinates\" are the radius T E [-l,I] (measuring the distance from the axis to the bar's center of mass) and the angle 0 E R (measuring the rotation about the vertical axis). The bar is modelled as a (homogeneous) rigid body of mass mB and length 21, the hand with load are modelled as a point mass m,. As Lagrangian we have the kinetic energy where mLB = mL+mB, ao(r) = ~ o + m , r ~ + m , ( r + l ) ~ . Here I. represents the inertia of the sleeve of the prismatic joint indicated in black in Fig. 1. The dynamic equations with actuator force F and actuator torque M are where s ( r ) = r+m,l/m,, is the distance of the loaded bar's center of mass (CM) from the axis. If s(r) = 0, the bar is in its equilibrium position, and one can show this position minimizes the moment of inertia ao(r) seen by the torque actuator. For the control problem we choose \"CM coordinates\" and These definitions generate the state and adjoint ODE system in the relatively simple form and the Hamiltonian becomes Singular Control for the Polar Robot The control u1 is singular if the switching function and all its derivatives vanish"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0001900_156855395x00076-Figure4-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0001900_156855395x00076-Figure4-1.png",
+        "caption": "Figure 4. Walking on the slope.",
+        "texts": [
+            " It is noted that three of the six legs are always put on the ground - it is possible to walk stably compared with the biped walking robot and the quadruped walking robot. Specific features of this robot are summarized as follows: (1) It can move in any directions and rotate. (2) The movement of the ball thread does not interfere with each other. (3) It can move on an uneven surface. (4) It can control the inclination of the upper surface. The length of each leg can be extended up to 130 mm; therefore, the maximum step that this robot can go over is 130 mm. The maximum inclination angle of the upper surface is 11\u00b0, as shown in Fig. 4. This robot can walk at speed of 30 cm/min on the plane. D ow nl oa de d by [ U ni ve rs ity L ib ra ry U tr ec ht ] at 0 2: 10 1 7 M ar ch 2 01 5 3. RELATIVE POSITION CONTROL EXPERIMENTS OF THE FRAMES This robot has a parallel link mechanism and two frames which can slide relatively by changing the length of three linear actuators based on the parameters (x, y, B). The movable ranges of three parameters (x, y, 0) are limited by three ball threads whose extendable lengths are limited from 0 to 250 mm"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0003787_iros.1996.570634-Figure2-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0003787_iros.1996.570634-Figure2-1.png",
+        "caption": "Figure 2: Model of Free Link",
+        "texts": [
+            " In Section 5, experimental results show that the manipulator can reach the desired configurations. 2 Modeling of Free Link A manipulator with three degrees of freedom in a horizontal plane is considered (Fig.1). The third joint is passive and is not equipped with an actuator or a brake. The first and second joints, which are active, control the position of the third joint in the 2-D plane. The third joint is a revolute joint around a vertical axis. The dynamics of the free link can be modeled as shown in Fig.2, where C g is the base coordinate frame fixed in the horizontal plane, C L is the coordinate frame k e d to the link (The origin is at the joint 0 and the x-axis coincides with OG, where G is the center of gravity of the link.), [z, U] is the position of the joint 0 in the base frame C g , and 6' is the angle b e tween the base frame C g and the link frame C L . The generalized coordinates which represent the configuration of the manipulator are [z, y, 6'1. The equations of motion with respect t o the link are: fz = mx - madsin 0 - m d 2 cos 0 r = -max sin 0 + m a y cos 6 + ( I + ma2)d where, m is the mass of the link, I is the moment of inertia of the link around G, a is the distance IOGI, between the joint and the center of gravity, [fz,f,] is the translational force at the joint 0, and T is the torque around the joint 0",
+            " Next, a composite trajectory from the initial configuration to the desired configuration is composed of the trajectory segments. Due to the drift term in the state equation, the p a s sive joint generally continues moving even when the active joints stop moving. Therefore, the trajectory segments should be special trajectories along which the active and passive joints simultaneously start and stop. Dynamic singularity, where the variation of one coordinate does not influence the other coordinates, is utilized for this purpose. In Fig.2, no angular acceleration occurs at the link when the joint is accelerated from rest in the direction of the z-axis of the link frame, CL. This is also true when the joint is decelerated from translational m e tion in the direction of the x-axis of C L , and is caused by a dynamic singularity where the translational acceleration in this direction is decoupled from the angular acceleration. This effect enables translation of the link without rot at ion. Next, rotation of the link is discussed. Suppose that the joint is accelerated in the direction of the y-axis of C L from rest"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0003694_00207170010023160-Figure2-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0003694_00207170010023160-Figure2-1.png",
+        "caption": "Figure 2. Single-valued non-linear I/O characteristics.",
+        "texts": [],
+        "surrounding_texts": [
+            "Suppose the non-linearity n\u2026u\u2020 in \u00ae gure 1 possesses the I/O properties listed in the previous section and the signal at its input is u\u2026t\u2020 \u02c6 A sin\u2026!t\u2020. Then, its output can be expressed by the Fourier series expansion y\u2026t\u2020 \u02c6 X1 k\u02c61 k:odd Ak\u2026A\u2020 sin\u2026k!t \u2021 \u00bfk\u2026A\u2020\u2020 \u20263\u2020 Retaining only the fundamental harmonic in (3), the DF is de\u00ae ned as N\u2026A\u2020 \u02c6 phasor representation of A1\u2026A\u2020 sin\u2026!t \u2021 \u00bf1\u2026A\u2020\u2020 phasor representation of A sin\u2026!t\u2020 \u02c6 A1\u2026A\u2020 A e j\u00bf1\u2026A\u2020 \u20264\u2020 In other words, the DF is the (real or complex) ratio of the \u00ae rst harmonic of y\u2026t\u2020 to the driving sinusoid u\u2026t\u2020. The error induced by this approximation is given by \u00a2N\u2026A\u2020 \u02c6 1 A X1 k\u02c63 k:odd Ak\u2026A\u2020e j\u00bfk\u2026A\u2020 \u20265\u2020 Note that, when the non-linear element is single valued, no phase shift in u\u2026t\u2020 occurs, i.e. \u00bfk\u2026A\u2020 \u02c6 0 for all harmonics. In order to obtain the closed-form version of (4) for a particular non-linearity, it is necessary to express it in terms of y\u2026t\u2020. Multiplying both sides of (3) by either sin\u2026!t\u2020 or j cos\u2026!t\u2020, and integrating, yields A1\u2026A\u2020 cos\u2026\u00bf1\u2026A\u2020\u2020 \u02c6 1 \u00ba \u20262\u00ba 0 y\u2026t\u2020 sin\u2026!t\u2020d\u2026!t\u2020 jA1\u2026A\u2020 sin\u2026\u00bf1\u2026A\u2020\u2020 \u02c6 j \u00ba \u20262\u00ba 0 y\u2026t\u2020 cos\u2026!t\u2020d\u2026!t\u2020 Adding the above two equations, and dividing the result by A, gives A1\u2026A\u2020 A e j\u00bf1\u2026A\u2020 \u02c6 j \u00baA \u20262\u00ba 0 y\u2026t\u2020e\u00a1j!td\u2026!t\u2020 Comparing this with (4), it follows that N\u2026A\u2020 \u02c6 j \u00baA \u20262\u00ba 0 y\u2026t\u2020e\u00a1j!td\u2026!t\u2020 \u20266\u2020 A similar equation can also be derived for \u00a2N\u2026A\u2020. Multiplying both sides of (3) this time by either sin \u2026k!t\u2020 or j cos\u2026k!t\u2020, and performing the same algebraic manipulations as before, results in Ak\u2026A\u2020 A e j\u00bfk\u2026A\u2020 \u02c6 j \u00baA \u20262\u00ba 0 y\u2026t\u2020e\u00a1jk!td\u2026!t\u2020 Substituting this into (5) gives \u00a2N\u2026A\u2020 \u02c6 j \u00baA X1 k\u02c63 k:odd \u20262\u00ba 0 y\u2026t\u2020e\u00a1jk!td\u2026!t\u2020 \u00bb \u00bc \u20267\u2020 Thus, it is clear that, in order to determine the behaviour of \u00a2N\u2026A\u2020 over a certain amplitude range, the addition of in\u00ae nitely many terms is required. Yet, when the nonlinearity is continuous at the origin and it has a soft I/O characteristic, it is possible to obtain an anlytical expression of the point to which this series converges by making use of symbolic computation tools. As a typical example, consider the single-valued nonlinearity shown in \u00ae gure 2(a). For A > q2, the DF of this element can be easily calculated from (6) as N\u2026A\u2020 \u02c6 q4 \u00a1 q3 f \u2026q1 \u00a1 q4\u2020 f \u2026q2; A\u2020 \u20268\u2020 where f \u2026qi; A\u2020 \u02c6 \u20262=\u00ba\u2020\u2026 i \u2021 \u2026qi=A\u2020\u00aei\u2020 with  i \u02c6 sin\u00a11\u2026qi=A\u2020 and \u00aei \u02c6 \u20261 \u00a1 \u2026qi=A\u20201=2 . The closed form of \u00a2N\u2026A\u2020, on the other hand, can be obtained from (7) as \u00a2N\u2026A\u2020 \u02c6 2 \u00ba X1 k\u02c63 k:odd \u00a1q3 sin\u2026\u2026k \u00a1 1\u2020 1\u2020 k\u2026k \u00a1 1\u2020 \u2021 sin\u2026\u2026k \u2021 1\u2020 1\u2020 k\u2026k \u2021 1\u2020 \u00b3 \u00b4\u00bb \u2021 \u2026q3 \u00a1 q4\u2020 sin\u2026\u2026k \u00a1 1\u2020 2\u2020 k\u2026k \u00a1 1\u2020 \u2021 sin\u2026\u2026k \u2021 1\u2020 2\u2020 k\u2026k \u2021 1\u2020 \u00b3 \u00b4\u00bc \u20269\u2020 D ow nl oa de d by [ T he A ga K ha n U ni ve rs ity ] at 2 3: 07 2 5 Fe br ua ry 2 01 5 By the aid of symbolic computation (Wolfram 1993), the limit of this in\u00ae nite series can be determined to be \u00a2N\u2026A\u2020 \u02c6 2 \u00baA \u2026\u00a1q1q3\u2026tanh\u00a11\u2026\u00ae1\u2020 \u00a1 \u00ae1\u2020 \u2021 q2\u2026q3 \u00a1 q4\u2020\u2026tanh\u00a11\u2026\u00ae2\u2020 \u00a1 \u00ae2\u2020\u2020 \u202610\u2020 The non-linear element of \u00ae gure 2(a) has a rather general I/O characteristic from which several other non-linearities can be obtained as special cases. One such case is the dead zone element depicted in \u00ae gure 2(b). The corresponding N\u2026A\u2020 and \u00a2N\u2026A\u2020 follow directly from (8) and (10) by setting q3 \u02c6 q4. These ideas are illustrated in \u00ae gure 4 which shows the variation of (10), (8) and N\u00a2\u2026A\u2020 \u02c6 N\u2026A\u2020 \u2021 \u00a2N\u2026A\u2020 with q1 \u02c6 0:9, q2 \u02c6 4:1, q3 \u02c6 2:7 and q4 \u02c6 0:6. Note that, when A < 4:1, the non-linearity behaves like a dead zone. Moreover, as A ! 1, the di\u0152erence between N\u2026A\u2020 and N\u00a2\u2026A\u2020 vanishes since \u00a2N\u2026A\u2020 ! 0. Quite frequently, non-linearities encountered in practical control systems have I/O characteristics which are not one-to-one relationships. An example is the relay element depicted in \u00ae gure 3(a). It has both a hysteresis loop and a dead zone. The N\u2026A\u2020 and \u00a2N\u2026A\u2020 of this element are slightly more complicated in that they are complex due to the overlap region. Using (6) and (7) they can be shown to be N\u2026A\u2020 \u02c6 2q3 \u00baA2 \u2026A\u2026\u00ae1 \u2021 \u00ae2\u2020 \u00a1 j\u2026q2 \u00a1 q1\u2020\u2020 \u202611\u2020 \u00a2N\u2026A\u2020 \u02c6 2q3 \u00baA X1 k\u02c63 k:odd 1 k \u2026e\u00a1jk 1 \u2021 e\u00a1jk 2\u2020 \u202612\u2020 where the  i and the \u00aei are as before, and A > q2. The point to which (12) converges can be shown to be \u00a2N\u2026A\u2020 \u02c6 2q3 \u00baA \u2026tanh\u00a11\u2026e\u00a1j 1\u2020 \u00a1 e\u00a1j 1 \u2021 tanh\u00a11\u2026e\u00a1j 2 \u2020 \u00a1 e\u00a1j 2\u2020 \u202613\u2020 Figure 5 is the equivalent of \u00ae gure 4 for this particular non-linearity with q1 \u02c6 1, q2 \u02c6 2 and q3 \u02c6 3 over the A range \u20302; 100\u0160. As with (10), (13) also goes to zero for su ciently large A leading to N\u00a2\u2026A\u2020 ! N\u2026A\u2020. A special case of this element is the rectangular hysteresis non-linearity (\u00ae gure 3(b)) which is obtained by stretching q1 up to \u00a1q2 and \u00a1q1 up to q2. The corresponding N\u2026A\u2020 and \u00a2N\u2026A\u2020 follow easily from (11) and (13). { In this paper, an operator is said to be quasi-linear if it involves, in some way, the DF of a non-linearity. Figure 3. Multi-valued non-linear I/O characteristics. D ow nl oa de d by [ T he A ga K ha n U ni ve rs ity ] at 2 3: 07 2 5 Fe br ua ry 2 01 5"
+        ]
+    },
+    {
+        "image_filename": "designv11_60_0002728_112515.112577-Figure11-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0002728_112515.112577-Figure11-1.png",
+        "caption": "Fig. 11 Algebraic constraints for agaimtz contact.",
+        "texts": [
+            " Therefore, we can apply differential approximations to replace pr andpE with p =pI\u2019 * (Z + AI) (25) p,z =pE\u2019 * (Z + AE) (26) where pI\u2019 and pE\u2019 represent the nominal positions of the parts. AI and AE are differential motion matrices for pt\u2019 and pE\u2019. Consequently, the following equation is derived based on (22) - (26): tE\u2019 *fE * pE\u2019 * AE -tI\u2019 *fI*pI\u2019 *AI =Ac*tl\u2019*f[*pl\u2019 (27) Ac is defined according to the contact type: (l)Agairm: Ac = AIO,dY,&&,&,&] (see Fig. 10). (2)Agaimtz: Ac = A[0,4,&&,0,&] (see Fig. 11). (5)Fim2: Constraints forfirw arerepeatedly given for two CVS, Suppose CFJ and CF2 represent CFS at two CVS. Let tll\u2019,lEl\u2019 and ACI are transformations for CFI and tn\u2019, tm\u2019and Ac2 are for CF2. If the direction of the X axis of CF2 in the world is in the negative direction of the X axis of CFI, m\u2018 and tm\u2019 are changed to M * tn\u2019 and M * tE2 \u2018 . ACI = A[&,O,&, &I, C5yl,6ZI] and Ac2 = A [&,0,42,&,&2,&2] where dil = h, & = &, 61 = &2, and & = - &. If the angle between the Y axes of CFJ and CF2 is q, the following constraints are additionally given (see also Fig"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0002670_0301-679x(90)90059-x-Figure5-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0002670_0301-679x(90)90059-x-Figure5-1.png",
+        "caption": "Fig 5 Geometry o f inclined tilted ring configuration",
+        "texts": [
+            "07 RII The resulting load obtained from equation (27) was 1.442 \u00d7 10 5. Tilted ring configuration For this configuration, an analytical method was introduced so as to compare with the present numerical approach. In order to solve Reynolds' equation, the short bearing approximation was made. As the name implies, the term OP/OO can be omitted in Reynolds' equation. With this assumption, the one-dimensional lubrication equation can be written as (rh:' t iJP, = 6\"qmr (28) Or Or I aO The variables which are used in the analytical model are shown in Fig 5. General film thickness variation is written as h h~. + h.t (Z/ro) cos 0 (29) and considering the central clearance at r~. h~. = h. [ 1 + t O'c/r.) cos 01 (311) the general film thickness can be obtained as tt = h. 1 + t cosO (31) , r ( ) , TRIBO/OGY INTERNATIONAL where tl- is the central clearance, Z the radial increment from r~., r . the outer radius of the ring, t the tilt as defined earlier and he the central clearance. Since the bearing configuration considered is short (short bearing approximation) the pressure distribution being parabolic along the radial ring width can be written with respect to central pressure P~,: P - P~ [1 - (Z /b ) e] (32) where b is the half-width of the ring and Z can be varied from - b to b"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0001235_b978-0-08-099425-3.00005-4-Figure5.33-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0001235_b978-0-08-099425-3.00005-4-Figure5.33-1.png",
+        "caption": "FIGURE 5.33",
+        "texts": [
+            " As the tyre rolls the tread moving through the contact patch is constrained by the road to move along a straight line, the net reaction of these forces being the camber thrust. Figure 5.32 shows a typical plot of lateral force Fy with camber angle g for increasing tyre load with the slip angle set to zero. From the plot it can be seen that the camber stiffness Cg is the gradient of the curve measured at zero camber angle at a given tyre load. In order to understand why a cambered tyre rolling at zero slip angle produces an aligning moment, it is useful to consider the effect of the shape of the contact patch. Consider the situation shown in Figure 5.33 where the wheel and tyre are rolling at a camber angle g with the slip angle equal to zero. The lower part of Figure 5.33 is a plan view on the tyre contact patch. The three points A, B and C, shown in Figure 5.33, are initially in line across the centre of the contact patch. If the tyre rolls so that point B moves to B0 at the rear of the contact patch and the rubber in the centre line is not subjected to any longitudinal stress. Generation of lateral force due to camber angle. Plotting lateral versus camber angle. Generation of self-aligning moment due to camber angle. Due to the camber the tyre will corner and point A on the inside of the tyre will roll at a smaller radius of bend to point C on the outside of the tyre",
+            " If it is presumed that the stiffness of the tyre restricts this and the points remain in line across the rear of the contact patch (A0, B0 and C0) then a longitudinal compressive stress acts on the inner A side and a tensile stress acts on the outer C side at the front of the contact patch. A similar effect occurs at the rear of the contact patch but the increased pressure in the front of the footprint of a rolling tyre means a net torque into the turn is developed when the effects are summed along the contact patch Plotting aligning moment versus camber angle e note the reversed sign compared to slip angle aligning moment. length, as shown in Figure 5.33. The geometric aspect is less obvious in motorcycle tyres but the mechanism is essentially similar. A typical plot of aligning moment with camber angle, for a given tyre load and zero angle, is shown in Figure 5.34. From the plot it can be seen that the aligning moment camber stiffness is the gradient of the curve measured at zero camber angle at a given tyre load. The lateral forces due to camber angle tend to be small when compared with those resulting from slip angle for a typical car tyre. In the linear range it would not be untypical to generate as much as 20 times the amount of lateral force per degree of slip angle compared with that generated per degree of camber angle"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0002235_0924-0136(95)02159-0-Figure2-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0002235_0924-0136(95)02159-0-Figure2-1.png",
+        "caption": "Fig. 2. A calculation scheme for the determination of the relative motion between the tool and the part.",
+        "texts": [
+            " 0924-0136/96/$15.00 \u00a9 1996 Elsevier Science S.A. All rights reserved SSDI 0924-0136(95)02159-J out also, by the introduced coordinate systems OoxoYoZo, associated with the roller, and 01x~y~z, associated with the half-finished material, towards the fixed coordinate systems Oxyz and O'x'y'z', which are translated one towards the other along the axis Ox by a distance A, referred to as the centre-line distance. At any moment in time, the tool (2) and the processed (1) surfaces have a common contact point M (Fig. 2), in which the requirement for contact between two mutually rolling surfaces is fulfilled [3,4]: VonNoM = 0 (1) where l/ou and Nou are the vector of their relative speed and the vector of their common normal, represented in the coordinate system OoxoYoZo. The vector Vou is determined by differentiation from the vector equation of the relative movement of a point of the half-finished material, defined in the coordinate system Otxtytzt, towards its conjugate point from the roller defined in the OoxoYoZo system. This equation, using Fig. 2, can be defined as mtRoM = Ro,~ + St + m3m2RtM (2) where RO M m_ XOM YOM 20M and R t ,v = Xl M YI M 7. I M are the radius vectors of point M, expressed with its coordinates, respectively, in the coordinate systems OoxoYoZo and O~ x~y~ z~; Ix i Re: =;oOi is the radius vector of the coordinate system O'~x'~y'tz't center O] in the system OoxoYoZo; 0 I S, = - s t sin/~ -- s t c o s i t is the vector of the axial displacement of the center O~, defined in the system Oxyz; i!11 -- cos o sin o !l COS (0 o I% o m2 -- cos tpt sin ~ot 0 - sin ~o~ cos ~pi 0 0 and m 3 ---= li 0 0 cos/t sin/~ -s in/z cos/t are the matrices of the transitions of the coordinates from the mobile coordinate system OoxoYoZo and OlX~ytz~, towards the fixed coordinate system Oxyz, indicating their rotations at angles ~Po, ~Pl and/z"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0000772_aim.2016.7576964-Figure6-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0000772_aim.2016.7576964-Figure6-1.png",
+        "caption": "Figure 6. Experimental setup to measure \ud835\uded5\ud835\udc33\u2032 imparted by the array of ASMs: angular guide, mounted on the micromanipulation system, to tilt the IPM by \ud835\udec9\ud835\udc33=30 0 or 60 0 (a similar guide was fabricated to allow inclinations of 750 and 900). The IPM\u2019s centre is placed at x=y=0, z=33 mm and tilted by \ud835\udec9\ud835\udc33=60 0 .",
+        "texts": [
+            " A torque gauge (HTG2-40 supplied by IMADA) with its respective torque sensor held the IPM in place to measure the transmitted torque \u03c4z\u2032. The IPM was connected to the torque sensor via a plastic connector that was prototyped using a 3D printer. Additionally, we fabricated plastic angular guides that allowed us to incline the IPM with the angles of \u03b8z=300, 600, 750 and 900. The angular guides and the probe tip of the Gauss meter can be moved along the X, Y and Z axes. These displacements are controlled by a micromanipulation system constructed of XYZ stages as shown in Fig. 6. In the first set of experiments, we measured B along the z axis by fixing the external magnetic system (i.e., \ud835\udec9\ud835\udc04\ud835\udc0f\ud835\udc0c=0 at all times) and only moving the tip of the probe with increments of 3 mm from -36 mm to 36 mm along the z axis (x=y=0 at all times). Under these conditions, the cylindrical components of B can be expressed in Cartesian components as Bx=Br, By=\ud835\udc01\ud835\udec9 and Bz=Bz. Therefore, we measured Bx, By and Bz along the z axis, although only Bx is shown in Fig. 7 because By and Bz varied between -3 [mT] and 3 [mT]"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0000012_ilt-03-2015-0031-Figure1-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0000012_ilt-03-2015-0031-Figure1-1.png",
+        "caption": "Figure 1 A method testing oil-film temperature of fluid film bearings which are faced by Babbitt metal or PTFE layer",
+        "texts": [
+            " An increase in PTFE layer thickness can lead decreases in maximum pressure (Fillon and Glavatskih, 2008). In addition, the similar conclusions were also summarized in Glavatskih and Fillon (2006). However, none of these works is conducted on large-scale thrust bearing. Operation monitoring is important in the maintenance of bearings all the way, especially large-scale bearings. But for large-scale pad faced by PTFE, the soft material has good thermal insulation so that designing temperature monitoring is a very head-breaking issue. The method in Figure 1 (Glavatskih, 2004) is that used thermocouples are embodied by the bearing backing material, and then continuous oil flows around the thermocouples are ensured by two orthogonal holes. But this measure not only affects the integrity of pad surface and strength and the abrasion-resistance of the PTFE layers, but also may result in silting-up of the holes easily from abrasive dust and then lead to an inaccuracy of temperature test. As for numerical study on thrust bearings, many works (Abdel-Latif, 1988; Wang et al"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0002990_901764-Figure6-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0002990_901764-Figure6-1.png",
+        "caption": "Figure 6. Positions of contacton the pinion tooth and on the mating gear tooth",
+        "texts": [
+            " where bh is half the width of the surface of contact, p- is maximum pressure, q is the load per unit length of surface of contact, and p l and p2 are the radii of the cylinders, E l and E2 are Young's moduli and v l and v2 are Poisson's ratios. In this study, p l and p2 are taken as radii of curvature at the pitch point of the mating teeth. Let the load transmitted by the pair of gears be and the length of contact at a time be LC. Then the tooth force per unit length q is h / L c . As derived by Weber and Banaschek (21) under a force per unit length q, the displacement Wh(x) under the Hertzian pressure is where h l and h2, as shown in Figure 6, are the lengths on the lines of application from the contact point to the center of the teeth. The flexibility due to contact Kc is where L is the length of contact of a pair of teeth. With the deflections found above and the loads applied on the contact line, the stiffness can easily be evaluated by the definition of equivalent stiffness as where Py is the load transmitted by a particular contact line and Wo is the deflection on the contact line. According to Eqs. 10 and 16, And when we rewrite Eq"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0003787_iros.1996.570634-Figure3-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0003787_iros.1996.570634-Figure3-1.png",
+        "caption": "Figure 3: Circular Motion of Joint",
+        "texts": [
+            " The circular motion of the link without translation is possible using this method. The above predictions are verified below more exactly. The acceleration [a, 0] in the z-direction of C L is given to the joint 0. The acceleration [j;,$] in Cg is: (4) x = acose y = as in0 Substituting (4) into ( 3 ) results in A 0 = 0, so 8 = 0. Therefore, angular acceleration of the link does not occur when the joint is accelerated or decelerated in the direction of OG. Next, assume that the joint moves along a circle of radius r, centered at a point [zo,yy~] in C g (Fig.3). The position of the joint, [z, y], is: z = 20 +rcos@(t) i y = yo +rs in@(t ) where @(t) denotes the angle which represents the position of the joint on the circle. The translational acceleration of the joint, [5, G ] , is: { (5) (6) 2 = -r4sin 4 - r$ cos 4 = r;dcos4 - r@ sin4 Substituting (6) into ( 3 ) leads to: r cos(8 - 4)$ + r sin(O - +)qi2 + ~8 = o (7) When 4 = B + 7r and r = A, Eq.(7) holds true and the constraint ( 3 ) is satisfied regardless of the velocity and acceleration of 4. The position of the center of rotation in the link frame C L is [A, 0] and is indicated as P"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0000639_s00170-016-9204-1-Figure1-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0000639_s00170-016-9204-1-Figure1-1.png",
+        "caption": "Fig. 1 Components of nozzle used in experiments",
+        "texts": [
+            " To design a head for gas free delivery, initially, the delivery nozzle models using gas as a source to move the particles to the melt pool were studied along with the inherited shortcomings as far as the deposition efficiency of the process is concerned. After that, the designed nozzle models available in the literature were considered in which the vibratory heads were used to move the particles. From these different models, the relationship of powder flow and vibration of nozzle exits was studied. The nozzle assembly and its components are shown in Fig. 1. The designed nozzle consisted of two portions, its main housing that is connected with the laser machine and its inner powder-containing hopper that is having no rigid contact with the nozzle housing. The arrangement is envisaged to provide a constant rate of powder delivery during the deposition process with an ability to control the flow by the start, stop and change in discharge quantity. To achieve a desired steady state flow rate of metallic powder through the designed nozzle, the vibrating head was used and the powder flow downward due to vibration and gravity force"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0002309_s004490050289-Figure1-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0002309_s004490050289-Figure1-1.png",
+        "caption": "Fig. 1. Schematic diagram of the anaerobic hybrid filter. (1) Wastewater storage vessel; (2) feed pump; (3) inlet; (4) immobilising support; (5) water jacket; (6) biogas outlet; (7) effluent outlet; (8) gas-meter; (9) effluent storage vessel; (10) sampling ports; (11) sludge discharge",
+        "texts": [
+            "1 The original anaerobic digester to be transformed into hybrid filter was a stirred tank reactor, working at 25 \u00b0C. After anaerobic digestion, the effluent goes to an aerated pond followed by a settling pond. Immobilisation of biomass in this plant, that has been necessary to face high organic loads and composition variability, has been simulated on laboratory scale by setting-up two anaerobic hybrid filters, consisting of 1.00 m-high PVC cylinders, with diameter of 0.20 m and working volume of 25 litres (Fig. 1). The filling material used as biomass support was constituted by plastic rings with length of about 2\u20143 cm and diameter of 2.5 cm. The temperature in the columns was regulated at the selected values by recirculating heated water in external jackets. The reactor was fed by means of a peristaltic pump, Watson Marlow, type 503 U/RL, which was regulated according to the selected organic load to be tested. The reactor was inoculated by adding fresh anaerobic sludges coming from the secondary digester of a municipal sludge treatment plant up to 10% v v~1"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0001996_s0736-5845(98)00026-x-Figure8-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0001996_s0736-5845(98)00026-x-Figure8-1.png",
+        "caption": "Fig. 8. Local configuration planning.",
+        "texts": [
+            " This gradient projection technique uses an arbitrary vector +h\"Lh/Lh. The function h (h) represents a desired performance criterion to be minimized: h (h)\"Eh(s j~1 )!hE2. (14) The desired criterion in Eq. (14) is to keep the configuration h in the optimal configuration space of h '-0\",j~1 and h '-0\",j planned by the global method: h(s j~1 )\"h '-0\",j~1 # s j~1 s j~1,j (h '-0\",j !h '-0\",j~1 ). (15) The desired configuration h(s j~1 ) depends on the path length s j~1 since P '-0\",j~1 and the path length s j~1,j between P '-0\",j~1 and P '-0\",j (see Fig. 8). By integrating the dh of Eq. (12) we receive the optimal configuration h n along the path dx between h '-0\",j~1 and h '-0\",j . Finally, the deflection of the elastic-rigid modeled robot has to be compensated. Therefore, the elastic beams of the robot were considered as Bernoulli beams [17, 18] whose deformation is described by the Ritz\u2014 Raleigh Approach [19]. An efficient approach, which is exactly enough for offline-programming, gives [20]. Herein the elastic deformation is described by a characteristic polynomial"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0003274_3.48216-Figure6-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0003274_3.48216-Figure6-1.png",
+        "caption": "Fig. 6 Influence of steep turn on a) the Dutch-roll mode shape and b) short-period mode shape: tilt-rotor aircraft, helicopter mode, 60 knots.",
+        "texts": [
+            " The dynamic characteristics of high-g turns in an airplane mode at 240 knots of this rotorcraft were found to be similar, to a large measure, to the 120-knot configuration. The helicopter-mode configuration at 60 knots, however, proved to be the most complex situation of the three configurations studied. As is evident from Figs. 5 and 6, both the eigenvalues and eigenvectors (particularly the latter) undergo considerable changes as load factor increases. In a steep turn at 60 knots ( 7 = \u2014 2 0 deg, nT=\\.5) in the helicopter mode, the contamination of the longitudinal motion in the Dutch-roll mode (see Fig. 6a) is even more dramatic than that in the conversion mode at 120 knots, as previously described (see Fig. 4). The change in pitch attitude is more than twice as large as that in roll attitude, and the changes in airspeed and angle of attack are substantial. Further, the change in roll attitude now lags more than 90 deg behind the change in sideslip instead of leading it, as in the l-g flight case (which, also, is normally the case for fixed-wing aircraft in l-g flight).7 The contamination of the lateral-directional motion in the longitudinal shortperiod mode (see Fig. 6b) in the steep turn is also much more severe than that in the conversion mode at 120 knots (discussed previously). The severe coupling in the longitudinal and lateraldirectional motions in steep turns, as revealed by the mode D ow nl oa de d by U N IV E R SI T Y O F M IC H IG A N o n Fe br ua ry 1 9, 2 01 5 | h ttp :// ar c. ai aa .o rg | D O I: 1 0. 25 14 /3 .4 82 16 16 R. T. N. CHEN J. AIRCRAFT a) b) \u2014 \u2014 -20 RIGHT TURNS 1 .5 2 0 1 1 .5 2 NORMAL LOAD FACTOR, ny, g ir 30 cc < 20 D -10 0 1-20 7 = 0\u00b0 RIGHT TURNS Fig"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0002128_ma946435m-Figure8-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0002128_ma946435m-Figure8-1.png",
+        "caption": "Figure 8. Projection of the disclination loops on the shear rate gradient axis-vorticity axis plane: (x1) gradient direction; (x2) vorticity axis; (x3) flow direction.",
+        "texts": [
+            " So, we must assume that the nonhomogeneous distribution of scattered light in the vertical streak is due to a 3D diffraction phenomenon. To study this, we will consider the simplest case: N disclination loops of equal lengths and similar internal core structure will be considered as assembled into one set. The upper and lower horizontal parts of the loops will be regarded as rods which are located in different planes parallel to the screen in this set (Figure 1). The projection of this set onto the (x1, x2) plane formed by the incident beam and vorticity axis is shown in Figure 8. The distances l and h between the nearest points of the rods along x2 and x3, and the angle R are independent random values obeying Gaussian distributions: s ) -i sin \u03b8 2 + j cos \u03b8 2 sin \u00b5 + k cos \u03b8 2 cos \u00b5 (5.6) O ) j cos \u03c8 - k sin \u03c8 (5.7) (MO) ) 1 2 E0\u03b4{sin 2\u03c8 sin2 \u00e2 cos2 \u03c9 + cos 2\u03c8 sin 2\u00e2 cos \u03c9 - sin 2\u03c8 cos2 \u00e2} (5.8) f ) 1 2 E0\u03b4D 2LCj0(12qDs1)j0(12qDs2)A3(pL,\u00e2;\u03c8) (5.9) A3(pL\u03b4;\u03c8) ) sin 2\u03c8 sin2 \u00e2J2 + cos 2\u03c8 sin 2\u00e2J1 - sin 2\u03c8 cos2 \u00e2J0 (5.10) J0 ) j0(ab) J1 ) 1 2 {j0[a(b + 1)] + j0[a(b - 1)]} J2 ) 1 4 {2j0(ab) + j0[a(b + 2)] + j0[a(b - 2)]} a ) 1 2 \u03c0pL and b ) qs3 \u03c0p (5",
+            " We will assume that there is no disorder in the orientation of optical axes of the rods in the set. In this case the light scattering intensity from the set under consideration with 2N horizontal rods can be written in the following general form: where ri - rj is a vector connecting the centers of the ith and jth rods and the angular brackets, \u2329...\u232al,h,R, represent the averaging over the random values of l, h, and R. By performing the summation in eq 5.19 for the arrangement of the rods shown in Figure 8, one can show that the following expression for the sum holds: where H ) D + h is the distance between the rods in the two depth levels of the set and l1, l2, ..., ln are the distances between different neighboring rods. If we assume that there is no correlation between the position of the rods in the (x2, x3) plane, we can average over the values ln separately. Using eqs 5.14, 5.19, and 5.20, we obtain the light scattering intensity per rod volume in the vertical streak (\u00b5 ) 90\u00b0) from the set of rods: The sum in eq 5",
+            " The situation for HPC is different since there is a helical winding up of the director around the core axis, changing with the shear rate. Acknowledgment. This work has been supported by the Direction des Recherches, Etudes et Techniques du Ministe\u0300re de la De\u0301fense, and by the Commission of the European Community (Brite Euram program). Notation D diameter of the rod (theory) d unit vector of the optical axis e unit vector parallel to the transmission direction of the polarizer which is in the (x2, x3) plane h inter-rod distance (Figure 8) HV crossed polarization, with the H polarization direction defined as being along the flow direction i, j, and k the unit vectors along the x1, x2, and x3 axes of the laboratory coordinate system L length of the rod l inter-rod distance (Figure 8) M induced dipole moment N number of loops in one set (2N represents the number of rods in the set) O unit vector defining the direction of the electric vector passing through the analyzer p number of rotations per unit of length s unit vector (eq 4.5) x1, x2, and x3 gradient direction, vorticity axis, and flow direction R angle defined in Figure 9 Ri polarizabilities of the rod \u00e2 tilt angle around the rod axis \u03b8 (or q) polar scattering angle (or scattering vector q with q ) (4\u03c0/\u03bb) sin(\u03b8/2)) \u00b5 azimuthal scattering angle \u03c8 direction of the optical axis of the polarizer according to the flow direction \u03b3\u0306 shear rate \u03c9 rotation angle around the core axis \u03c3x dispersion of the parameter x References and Notes (1) Doi, M"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0002252_s0890-6955(98)00006-6-Figure6-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0002252_s0890-6955(98)00006-6-Figure6-1.png",
+        "caption": "Fig. 6. The condition of wear for nose radius tool tip.",
+        "texts": [
+            " The area of shear plane A and the friction plane Q for case (2) can be obtained as follows: For convenience of calculation, the shear plane must be projected in the plane perpendicular to the speed of cut, which makes the calculations and analysis much easier and saves the time required for calculations. Defining the chip flow angle in this perpendicular section as hc9, the relation between hc9, and hc on the tool face is presented in Appendix Eq. A(6). According to Eq. A(6), the shear plane area can be varied by changing hc9 in small increments. Geometrical wear of the cutting tool on the tool face is shown in Fig. 5 and the type of wear on the tool edge is defined by a curve of radius R1, a straight line and a curve of radii R2, R3 (R2 = R3). The side view of the worn tool is shown in Fig. 6, from which the flank wear and wear land can be readily realized. The measurable values (i.e. l1, l2, l3, l4, h1, h2, h3, h4 and h5) can be obtained by amplifying the profile projector screen by drawing the circumference of the tool edge before the cut, and can draw it after wear has occurred. The geometrical lengths, radius of curve and curve angles of the worn tool are shown in Fig. 5, and can be obtained by the following equations: EP = (l21 + h2 1)1/2; PC = (l2 2 + h2 2)1/2; CN = (l23 + h2 3)1/2 and MN = (h2 5 + l24)1/2 (10) uR1 = tan21(l1/h1); uR2 = tan21(h3/l3); uR3 = tan21(h5/l4) and uPC = tan21(h2/l2) (11) R1 = EP/(2sinuR1); R2 = CN/(2sinuR2) and R3 = MN/(2sinuR3); (R2 = R3) (12) Due to the variations in Cs and the different kinds of geometrical circumferences of tool edge, several different cutting conditions occur when the feed is varied"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0002537_a:1007917003845-Figure5-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0002537_a:1007917003845-Figure5-1.png",
+        "caption": "Figure 5. Tested tasks.",
+        "texts": [
+            " The distance between the two robots is 0.9 m, and the length and the mass of the object JINT1410.tex; 18/02/1998; 16:12; v.7; p.16 are assumed to be 0.1 m and 5 kg, respectively. The world coordinate system {X,Y,Z} coincides with the coordinate system of robot 1, which is located at the base of robot 1 in M.K.S. unit. Kinematic and dynamic parameters of a PUMA robot are so well-known in [18] and [19] that the formulations of both optimization methods are obtained immediately. We choose three tasks for test, as shown in Figure 5. Two PUMA robots are to move a common single object along the boundary line of rectangles in point-to-point linear motion in Tasks 1 and 2, and the circumference of a circle in uniform velocity motion in Task 3. Each line motion has a velocity profile with zero starting and ending velocity. The ordered coordinates of the points on the corner of the rectangular tasks are (0, 0.4, 0.96) \u2192 (0, 0.5, 0.96) \u2192 (0, 0.5, 0.86) \u2192 (0, 0.4, 0.86) \u2192 (0, 0.4, 0.96) in task 1, (\u20130.05, 0.4, 0.86) \u2192 (0.05, 0.4, 0"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0003219_s0092-8240(87)80026-5-Figure3-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0003219_s0092-8240(87)80026-5-Figure3-1.png",
+        "caption": "Figure 3. The dependency of the eigenvalues 2~ (l = 0, 1, 2, 3, 4) on the steady state radii R 0 .",
+        "texts": [
+            " As special solutions we found spherically symmetric steady states which are related to the nutrient supply z as shown in Fig. 2. We performed a linear stability analysis around these spherically symmetric steady states with respect to disturbances of the form r = R o + e Y~t (0, rp)e at, (2) where y m are spherical harmonics. We obtained a series ofeigenvalues 2 which do not depend on m. The relation between the leading eigenvalues 2~ for each l = 0, 1, 2, 3, 4 and the steady state radii R o is shown in Fig. 3. Concerning the spherical mode l = 0, the branch R o (Fig. 2) is unstable and the branch R~ is stable. The translational mode l = 1 is always marginally stable. With respect to higher modes there exist successive R* (R* < R?+ a) such that the mode 1 is stable on R o < R* and unstable on R o > R~'. In any case l = 2 is the dominant unstable mode, which can be interpreted as the onset of a division process. Altogether, the analysis results in the following: (i) for low nutrient supply (z < Zo), the protocell shrinks and dies",
+            " (16) x \u00b0 ~ x \u00b0 + ~ 2 Equation (16) is substituted into equation (14) and after the linearization one finds the eigenvalue 2 in the following form, Ko2(l+ 2)(l-- 1) 21 = 2(1 + h(tco)- C 1 - KoI(t\u00a2o) ) { - ( y I ' +g ' ) -exg l ( l+ l)h'}, (17) where the prime means the derivative d/dK o . One can see immediately that 2~ = 1 = 0, which corresponds to a translational invariance of the protocell model. All other modes are calculated numerically and an example is given in Fig. 8; again a good correspondence to Fig. 3 is obtained. 5. A Preliminary Numerical Result. A numerical analysis can be performed in the following way. The time evolution of the curvature is given by the geometrical relation, = - ( K2 + A s ) Vn. (18) Equation (18) together with equation (14) enables us to calculate the time evolution of the surface. Here it should be noted that the arclength changes also by the geometrical relation, equation (11). The numerical procedure we have used is the same as in the work of Brower et al. (1983), therefore we report here only the numerical results"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0003738_850439-Figure1-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0003738_850439-Figure1-1.png",
+        "caption": "Fig 1 - Cross-head Petter AVl ~<1ith",
+        "texts": [
+            " In a conventional engine this is difficult, due to the orbital motion of the connecting rod. To sinplify this problem an additional cylinder block was nounted on top of 850439 the existing one and the engine operated as a cross-head. The piston rod then had pure reciprocating motion and ~<1as thus relatively simple to seal. The seal used was a PTFE reciprocating seal manufactured by Fluoro Carbon IJSA. A cross-sectional dra<iJing of the top half of the engine, shm-ling the seal in position, is shotm in Fig 1. sliding seal The raethod of supplying oil to the liner Has designed to be as close to the original splash lubrication systera of the AVl as possible. A snall reservoir of oil was contained above the seal and) as the piston descended, the turbulence created splashed this oil onto the liner wall. Tests were carried out to ensure that sufficient oil \\4aS supplied by this r.lethod. The oil in this reservoir t<1as continually circulated from a 500 ml capacity sunp that contained the test lubricant. An additional container, and a system of inlet control taps, allowed different test oils to be blended in when required"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0000810_1754337115626636-Figure1-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0000810_1754337115626636-Figure1-1.png",
+        "caption": "Figure 1. (a) Octahedral sphere and (b) structured quadrilateral mesh.",
+        "texts": [
+            " The parameters measured from the impact test include the contact time, longitudinal deformation and the coefficient of restitution (COR). To the authors\u2019 knowledge, none of the published FE models of soccer ball were validated against reaction force data. Since the developed ball model will be used to simulate soccer heading, a reaction-force-validated soccer ball model is imperative. The soccer ball model was developed using a hollow spherical shell with a diameter of 220mm. The surface of the shell was partitioned to form a spherical octahedron and meshed with quadrilateral elements as shown in Figure 1. Shell elements are preferred to the threedimensional continuum elements since they are not only more efficient for thin-walled structure modelling, but also computationally inexpensive and superior in predicting bending behaviour. In addition, the composite shell elements allow multiple material definitions for each layer.3 The composite shell elements in our model include two layers: inner bladder (0.2mm thick) and outer panel (2.2mm thick). The stitching seam was not modelled since it has no significant effect on the simulation results"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0000837_red-uas.2015.7441004-Figure3-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0000837_red-uas.2015.7441004-Figure3-1.png",
+        "caption": "Fig. 3. Pure rolling motion",
+        "texts": [
+            " In the Figure 2, the yaw angle is represented, which can be described with the following equations: \u03c8\u0307 = r (4) r\u0307 = Nrr +N\u03b4r\u03b4r (5) where \u03c8 represents the angle of yaw and r denotes the yaw rate with respect to the centre of gravity of the airplane. \u03b4r is the rudder deflection. Nr and N\u03b4r are the stability derivatives for yaw motion. The following equations describe the dynamics for the roll angle: \u03c6\u0307 = p (6) p\u0307 = Lpp+ L\u03b4a\u03b4a (7) where p denotes the roll rate and \u03c6 describes the roll angle, it is observed that \u03b4a represents the deviation of the ailerons. Lp and L\u03b4a represents the stability derivatives of the roll motion [11]. In the Figure 3 are shown the variables of the roll motion. In this section have been described the controllers that have been designed in order to control the fixed-wing MAV. The backstepping control for altitude motion has been designed considering the equations (1) and (3),For the design of the altitude controllers the linear longitudinal velocity is considered to be constant, in the equation (2) \u03b4e represents the elevator deviation, and also is the control input. The altitude error has been defined as e\u0303h = hd \u2212 h, and it denotes the difference between the desired altitude hd with respect to the current altitude h, where h was obtained by integrating (3)"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0001847_pime_proc_1994_208_361_02-Figure2-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0001847_pime_proc_1994_208_361_02-Figure2-1.png",
+        "caption": "Fig. 2 Types of irregularities of a journal",
+        "texts": [
+            " The present analysis reveals that, at the critical Sommer eld number, thefiiction is very low compared to the very high value for an ideal journal. The present analysis also reveals that the Action and the attitude angle are lower for any Sommerfeld number and that the load-currying capacity is lower for any eccentricity ratio than that for an ideal journal. NOTATION radial clearance = (bearing bush diameter - journal diameter D)/2, where D is as defined below clearance ratio diameter of the circle circumscribing the peaks of the undulations of the journal (Fig. 2a) eccentricity oil-film thickness hlc length of bearing number of undulations on journal speed of journal oil-film pressure in the clearance space dimensionless oil-film pressure in the clearance space specific bearing load = WILD radius of the circle circumscribing the peak of the undulations of the journal Sommerfeld number critical Sommerfeld number (Sommerfeld number at E = 1) bearing wall thickness peripheral velocity of journal resultant oil-film force, that is loadcarrying capacity resultant oil-film force component along the line of centres resultant oil-film force component perpendicular to the line of centres Cartesian coordinates dimensionless circumferential and axial coordinates barrel/bellmouth error circularity error barrel/bellmouth error ratio circularity error ratio eccentricity ratio permeability parameter absolute viscosity of oil coefficient of friction The M S was received on I July 1993 d was accepted for publication on 8 December 1993",
+            " Since all machined journals have some geometric irregularities, a study is necessary to know their role on the performance of bearings. In the present analysis geometric irregularities of a journal, such as circumferential undulations and barrkl/bellmouth shapes, are taken into account. Similar studies for solid journal bearings have already been reported (9-1 1). 2 ANALYSIS Figure 1 shows the porous bearing geometry with the coordinate system and the notations used in the analysis with an ideal journal. Figure 2 shows various types of geometric irregularities of the journal considered in the analysis. It is assumed that the irregularities vary sinusoidally. In commercial porous bearings, the permeability increases gradually from the ends to the middle and in the present analysis the variation is assumed to be sinusoidal, as per the curves of reference (12), and the variable permeability @,,, = @ (0.3 + 1.4 sin(zz/l)}. Proc Imtn Mcch Eagrs Vol208 at CORNELL UNIV on September 23, 2016pij.sagepub.comDownloaded from 2"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0002281_1.2831582-Figure2-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0002281_1.2831582-Figure2-1.png",
+        "caption": "Fig. 2 Bertram and Eshel's model",
+        "texts": [
+            " For a web passing around an unnipped winding roll, Knox and Sweeney (1971) suggested that the air layer thiclcness may be approximated with the foil bearing equation, R, \\ T J (1) where u is the winding speed or the average velocity of two adjacent surfaces. The effects of a nip roller were reported by Bertram and Eshel (1980) and by Eshel et al. (1984). They solved the Reynolds equation for the case of an incompressible lubricant. d (,-i dp\\ ,^ dh \u2014 / j ' \u2014 = \\2pu \u2014 dx \\ dx I dx with two boundary conditions. Ps(-\u00b0\u00b0) = 0 and pj(0) = T/R, (2) (3) where x = 0 indicates the center of the nip as shown in Fig. 2, and T/Ri is the pressure under the outer layer of the winding roll. By solving Eq. (2) with the above boundary conditions, the air film thiclcness at the center of the nip {h\u201e) is obtained to be f* = J p,dx = 4puR 4 ha T ^2Rh\u201e 37r R, (4) where F* indicates the nip loading obtained by integrating the air pressure in the region of -\u00ab> <.;\u00ab:< 0. The second term on the right-hand side of Eq. (4), which represents the effect of tension, is nearly negligible compared to the other term for most practical web handling operating conditions"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0003547_0954406001523074-Figure6-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0003547_0954406001523074-Figure6-1.png",
+        "caption": "Fig. 6 Displacement or velocity diagram for the conditions at",
+        "texts": [
+            " Only PAC-N and PAT-N con\u00aegurations were used. It was found that neither extending the conjugate line beyond the normal condition -N nor stopping short of it was any use. Tracing a point beyond this condition generates the beginning of the involute of the opposite hand, that is, for the opposite direction of rotation. This oppositehanded involute begins where the \u00aerst involute reaches the N base outline. Likewise, if the conjugate line does not reach the N normal point, the full length of the involute is simply not described. Figure 6 is a displacement or velocity diagram for the conditions at the pitch point on non-circular gears. This diagram was inspired by that of Juvinall and Marsheck (reference [16], Fig. 15.26, p. 580). The current pitch point PP0 and the following PP1 are shown with their related instantaneous conjugate lines. Two points c0 and c1 are shown on these conjugate lines travelling from one base outline to the other, with the appropriate variable velocity. Here, c0 is superimposed on the current pitch point, and c1 will be superimposed at the next location of the gears. The initial points for the construction of Fig. 6 are available because as detailed in Section 4.1, the process began by generating a \u00aele of points, on the pitch outlines, which is the sequence of pitch points. The tangents and conjuugate lines associated with each pitch point can be readily constructed for the two con\u00aegurations PAC-N and PAT-N. First the displacements of the two points c0 and c1 are examined as they move between two succeeding pitch points in a coordinate system \u00aexed with respect to the gear centres. Vector t1\u00b1PP0 is the displacement of the current pitch point along the instantaneous tangent"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0001076_indicon.2016.7838972-Figure3-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0001076_indicon.2016.7838972-Figure3-1.png",
+        "caption": "Fig. 3. Various defective bearings used for the experimental study.",
+        "texts": [
+            " The experimental study has been performed on a three phase four-pole induction motor ratings 3.0 Hp, 415 V, 4.5 Amps, 1440 rpm. The rotor of the motor is mounted on two ball bearings with the numbers 6205 on shaft-side and 6204 on fan-side. The three phase current signals were recorded using Yokogawa WT3000 Power Network Analyzer. These current signals were sampled at 5.0 kHz i.e., 100 samples per cycle. Fig. 2 shows the picture of experimental set up for recording stator current signals of induction motor. The bearings with various defects used for the experimental study are shown in Fig. 3. Initially the motor was run with healthy bearings on both the sides in order to obtain reference features which would be utilized for detection and further diagnosis. Later, the bearings with various defects were mounted on shaft-side (Fig. 3(a)) keeping healthy bearing on the fan-side. In order to study the effect of defective bearing position on fan-side (Fig. 3(b)), the bearing of same defects are used on the fanside keeping healthy bearings at the shaft-side. The recorded signals are analyzed with the multi-resolution analysis based S-transform to obtain complex s-matrix S(f, t). The maximum magnitude Amax = max(| S(f, t) |) and maximum phase \u0398max = max( S(f, t)) vectors were obtained from this s-matrix. Various statistical properties of these vectors such as maximum, minimum, mean, mode, standard deviation and variance were analyzed in the viewpoint of fault detection and diagnosis"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0000644_tjj-2015-0018-Figure1-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0000644_tjj-2015-0018-Figure1-1.png",
+        "caption": "Figure 1: Turbine blade module: (a) schematic of turbine blade arrangement, (b) LP turbine blade showing fir-tree root and tip shroud and (c) LP turbine rotor assembly.",
+        "texts": [
+            " Turbine module of a twin spool turbojet engine comprises HP NGV (high-pressure nozzle guide vane), HP turbine, *Corresponding author: S. K. Panigrahi, Mechanical Engineering Department, Defence Institute of Advanced Technology (Deemed University), Pune, India, E-mail: panigrahi.sk@gmail.com Benudhar Sahoo: E-mail: bsahoo543@gmail.com, R. K. Satpathy: E-mail: rksatpathy@gmail.com, Regional Centre for Military Airworthiness, Koraput, India Unauthenticated Download Date | 6/12/16 11:42 AM LP NGV (low-pressure nozzle guide vane) and LP turbine, a sketch of which is placed in Figure 1(a). Pressure of gas reduces after HP turbine due to expansion and, therefore, LP turbine blades are made comparatively lengthier to accommodate higher specific volume. Creep and fatigue are the life-limiting damage mechanisms for turbine blade. Therefore, SRT (stress rupture test) and vibration fatigue tests are being carried out as a part of qualification test for turbine blades [5]. The LP turbine blade under study shown in Figure 1(b) is a solid blade with tip shroud and z-lock configuration having fir tree root for fixing with turbine disc. High aspect ratio of the blade causes higher vibration unless supported properly. Additional support in the form of tip shroud is provided to limit the amplitude of vibration, thereby improving its reliability during operation [1]. The blade rotates at 8,500 rpm having a tip diameter of about 1 m operating at 900\u00b0C and develops a maximum centrifugal stress of 402 MPa. It is made of a typical wrought nickel base superalloy and provided with an oxidation-resistant aluminizing coating"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0000154_icelmach.2014.6960384-Figure5-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0000154_icelmach.2014.6960384-Figure5-1.png",
+        "caption": "Fig. 5. Mechanical model for the MBD simulation. The model consists of two bodies, a flexible and unbalanced rotor (blue) and a fixed rigid stator (grey). Furthermore, two bearings (Joint1, Joint2) are modeled. The permanent magnet synchronous machine (PMSM) is modeled with the proposed FE-based machine model.",
+        "texts": [
+            " Furthermore, the output of both models can be directly compared because both models get the same input. In the third stage, the output quantities , , , , A B C A i i i \u03a8 , , , , B C r T F F\u03d5\u03a8 \u03a8 of both models are compared. ABC ABC r T F F\u03d5i \u03a8 can be carried out. V. VALIDATION SETUP A PMSM with surface mounted permanent magnets and a nominal power of 15kW is used for the validation. The machine has six magnetic pole pairs and a nominal speed of 5000rpm. The mechanical model used for the validation is shown in Fig. 5. It consists of an unbalanced rotor, a stator, the FE-based model of a PMSM and two bearings. Additionally, an inverter, a battery and a drive controller are used to model a simple electric power train. The influence of the additional magnetic force to the displacement for the used mechanical model is shown in Fig. 6. Considering the additional magnetic drag leads to an approximately 40% higher displacement compared with a simulation neglecting the magnetic force caused by rotor eccentricity. In this section the comparisons between the reference FEM model and the FE based model are presented"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0001606_1.1697892-Figure4-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0001606_1.1697892-Figure4-1.png",
+        "caption": "FIG. 4. Geometry for pattern computation.",
+        "texts": [
+            " If the power received is PR , then the power per unit area P is (14) The isotropic value of power, Po, may be ob tained by finding the total energy received by integration of the re~eiving pattern over all 873 [This article is copyrighted as indicated in the article. Reuse of AIP content is subject to the terms at: http://scitation.aip.org/termsconditions. Downloaded to ] IP: 130.113.86.233 On: Wed, 03 Dec 2014 15:51:04 angles, and redistributing this energy uni formly. The geometry of the situation is shown in Fig. 4. Assuming that the pattern is inde pendent of cp, the value of Po is 1f\" P o=- (PR /A)sin8d8. 2 0 (15) Calculation of gain values proceeds by graphical evaluation of this integral from the receiving patterns. Experimentally, power levels are re ferred to the maximum value for the receiving pattern, PRM\u2022 Hence the expression used for computation takes the form 1 G (16) N ! L (PR .! P RM)sin8ifl8 i i=l where the summation extends from 8 = 0 to 8=1r. The effective area A may be obtained further from computed values of G by use of Eq"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0000902_ilt-07-2015-0097-Figure1-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0000902_ilt-07-2015-0097-Figure1-1.png",
+        "caption": "Figure 1 Geometric configuration of a constant flow valve-compensated non-circular hole-entry hybrid journal bearing",
+        "texts": [
+            " It is observed from this work that proper selection of micropolar parameters and capillary design parameter significantly affects the performance characteristics of circular and the non-circular bearing configuration. The above literature review based on the non-circular bearing configuration and micropolar lubrication reveals that the performance of a constant flow valve-compensated non-circular bearing has not been analyzed till date. Therefore, in the present work, efforts have been made to fill this research gap. The geometric configuration of a constant flow valvecompensated non-circular (i.e. two-lobe) hole-entry hybrid journal bearing is shown in Figure 1. There are two rows of holes with 12 holes per row, and it is assumed that the journal Non-circular hole-entry hybrid journal bearing Rajneesh Kumar and Suresh Verma Industrial Lubrication and Tribology Volume 68 \u00b7 Number 6 \u00b7 2016 \u00b7 737\u2013751 will rotate with uniform angular velocity about its equilibrium position. The modified Reynolds equation governing the laminar flow of iso-viscous, incompressible micropolar lubricant in the convergent space of a journal bearing system with usual assumptions is given by: x h3 12 p x y h3 12 p y jRj 2 h x h t (1) Where: 1 12l2 h2 6Nl h coth Nh 2l and N 2 1/2 ; l 4 1/2 is named as the micropolar function, whereas is the Newtonian viscosity coefficient, is the spin viscosity, is material coefficient, h is fluid film thickness and p is the micropolar fluid-film pressure"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0003666_aim.2001.936770-Figure1-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0003666_aim.2001.936770-Figure1-1.png",
+        "caption": "Fig. 1. Left: Parameters and frames for the pin-hole perspective view model; Right: Global set-up and frames",
+        "texts": [
+            "00 0 2001 IEEE 810 On the other hand, many researchers reported vision based feedforward control. However, in contrast to this work, they mostly use off-line generated models e.g. [SI or partially known paths or workpieces e.g. [lo]. [l] already dealt with the problem of keeping the continuous curved contour in the camera field of view in addition to maintaining and even improving the force controlled contact. The present work focuses on the used egdelcorner detection algorithm and control while taking a corner. C. Notation Figure 1 (right) gives the global experimental set-up defining the tool (or task), the camera and the robot end effector frames. They are denoted by preceding subscripts t , cam and ee, respectively. Figure 1 (left) shows the pin-hole perspective view model. f is the focal length of the camera lens. c a m ~ C is the distance along the .,,t-direction between the camera frame, located\u2019 a t the optical center of the lens, and the (contour in the) object plane. Following notations are used: H o is a frame of object 0. Superscripts C , OC and COG\u2018 refer to contour, offset contour and corrected offset contour respectively; :H is a homogeneous transformation from frame b in frame a. x, y and z are distances in [mm] except when used with subscript p , which denotes a parameter expressed in [pixels]"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0002923_978-94-011-3514-6-Figure4.11-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0002923_978-94-011-3514-6-Figure4.11-1.png",
+        "caption": "Fig. 4.11 Shaft of machine supporting an unbalanced disk",
+        "texts": [
+            "969 for the resonance frequency + 0.62 % (over estimate) \u00a3(1110) 1II0a - 1II0e 1II0e 120 - 120.45 120.45 - 0.37 % (under estimate) The range of application of the approximate method is quite large. In effect : \u00a3(1) , + 5 % 0.75 0.14 2.9 % In practice, measurement of the damping factor by means of a test in the free state (examples 3.7.1 and 3.7.2) is usually preferable to measurement in the steady state described above. The shaft of a machine supports a disk at its centre of mass m1 , having a mass unbalance of m2 (figure 4.11). Determine the ranges of speeds of rotation for which the maximum bending stress, from vertical motion, stays below a value 00 specified, by taking into account the stress due to the weight of the flywheel, but neglecting the mass of the shaft. Let us use the following numerical values g 200 kg 0.2 kg 9.81 m/s2 L 0.8 m R 0.25 m r = 0.03 m E 00 2.1.10 11 Pa 100 MPa - 76 - ml X2 + k Xz mz R wZ cos wt The result (4.10) gives the amplitude displacement of the steady state mZ R w2 k 11 - aZI a = w/wo being the relative angular frequency"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0003914_s1474-6670(17)54637-6-Figure3-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0003914_s1474-6670(17)54637-6-Figure3-1.png",
+        "caption": "Fig. 3. Two different arm configu rations in a horizontal plane",
+        "texts": [
+            " (a) Salisbury's approach K \" A~ \\ I \\ V \" (b) DCC Formulation For simplicity, we ignore compliance compo nents of moment, and focus our analysis on the relationship between end-effector com pliance (or stiffness) and joint compliance (or stiffness) within the same plane. Let us introduce the next notation : Cl1~Cx' C12~CXY' C2l~CYX' and C22~Cy ' Since n=2 holds for the arm in a horizontal plane, Cx ' Cxy ' and Cy can be freely varied wi th m=3, which means that DCC arm is materialized by setting three variable compliance elements in the arm . To realize this mechanically, two approaches are conceivable, as shown in Fig. 3 . and Eq . (8) is rewritten for both approaches. (i) Serial arm (ii) Parallel arm Cq1 Cq2 Cq3 ....... (14) 368 M. Kaneko et al. k 2 2 x x 2 xl k q1 1 k =---------- x 2Y2 x 1Y1 xy k (det I J~ I) 2 2 2 k y Y2 Y1 q3 2 2 x 4 x3 k q3 1 + ---------- x 4Y4 x 3Y3 (detIJ~I)2 y2 4 y2 3 kq4 ... . ... (15) where (xi' y i) expresses the joint posi tion in a left task-coordina te system, and J a' J b are arm Jacobian matrix and right arm Jacobian matrix, respectively. The formulation of stiffness a parallel arm for simplicity is focus sed on of notation",
+            " This condition is given by: (i) Serial arm (16) (ii) Parallel arm X2Y2kq1+x1Y1kq2+r(x3Y3kq4+x4Y4kq3)=0 (17) where r=(detIJ~I)2/(detIJ~I)2. If assuming all joint compliances keep positive values, that is Cqi>O (or kqi>O), at least one of should have a different sign from the others . This signifies that the arm attitude plays an important role for decoupling of compliance matrix. Simulation (i) Serial Arm Figure 4 (a) gives the selectable compliance map CX/C y computed from Eq. (14) for the arm attitude such as shown in Fig. 3 (a), where the decoupling presupposed. The condition shaded area C =0 xy in Fig . is 4 (a) represents a range that provides physically stable solutions (Cqi>O), while all the other areas provides at least one Cqi with negative value. With negative Cqi , serial arm can not maintain its arm attitude. C IC can be y x seleted exceptionally large or small in a=20\u00b7, 50\u00b7, and 80\u00b7, and these attitudes correspond to the condition that the point 0i lies on either the x or the y axis. Let us for example take up the case where the point 01 lies on the x axis",
+            " Compliance Cq1 existing at point 01 will give no influence whatever on C XY and Cx' but will merely influence on Cy ' Accordingly, it is easily inferred that by selecting a large Cq1 ' Cy may be selected relatively large against Cx Direct Compliance Control of \\Ianipulator .\\rllls 369 (E) Parallel Arm The parallel arm is assumed to be provided with variable compliance elements at joints 1 through 3, and it is assumed that two adjacent arms connected to joint 4 are coupled together with a pin joint having no active compliance element. Fig. 4(b) gives an example of decoupling selectable k /k x y condition k =0 for xy map the with arm attitude such as shown in Fig. 3(b). This simulation is also performed presupposing positive stiffness values for the three joints with variable compliance elements. The exceptionally large values in Fig. 4 (b) of k /k are due to the same reason as that x y given earlier for the serial arm. Experimental Verification Dee scheme will be verified experimentally by using one finger of robot hand with capability of adjusting its joint compliance. This hand shown in Fig. 5 has three joints per finger and its motion is limited in a horizontal plane"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0000077_icoin.2016.7427086-Figure13-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0000077_icoin.2016.7427086-Figure13-1.png",
+        "caption": "Fig. 13. The path taken by UAV in scenario 3 with increasing obstacle velocity. (a) Using To Goal search (b) Using Maximum Velocity search",
+        "texts": [
+            " 12(b), the UAV first moves towards the goal at t1 then starts to move towards the left from t2 until t25 to avoid collision with the obstacle. At t26, it starts aiming back to the goal. It reaches the goal at t45 or after 4.5 seconds. In this scenario, using TG search gives shorter path to the goal by 1 second. In this scenario, the obstacle\u2019s position is the same with previous scenario t0 = (6, 1) but this time the obstacle is increasing its velocity. The obstacle has initial velocity of (vx = \u20133 m/s, vy= 4 m/s) and obstacle is changing its velocity by (vx = \u20130.1 m/s, vy= 0.1 m/s). In TG search, as shown in Fig. 13(a), the UAV moves with maximum speed at t1 then stops at t2. It starts to move again with varying speed starting from t7 until t10 then maintains to move at maximum speed from t11 until it reaches the goal at t32 or after 3.2 seconds. Using MV search, as shown in Fig. 13(b), the UAV first moves towards the goal at t1 then starts to move towards the left from t2 until t11. It moves towards the right at t12 then moves directly towards the goal from t19 .It reaches the goal at t34 or after 3.4 seconds which is slower than TG search by 0.2 seconds. In this scenario, the obstacle\u2019s position is at t0 = (5, 11) with constant velocity (vx = \u20132 m/s, vy= \u20131 m/s). Using TG search, as shown in Fig. 14(a), the UAV maintains maximum speed until t7. At t8, it stops moving and starts again to move at slow speed at t24 then starts to accelerate after the obstacle has passed"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0000617_s1064230716030126-Figure1-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0000617_s1064230716030126-Figure1-1.png",
+        "caption": "Fig. 1. Space robot remote control circuit.",
+        "texts": [
+            " It involves the use of a plausible model of the space robot and its environment at the ground control center [6] with the simulation of conditions of gravity in the robot\u2019s operating area (in particular, microgravity conditions) [7]. In this model environment, using their intellectual capacity people should ensure the accomplishment of the required actions by controlling the space robot model in the master-slave mode using the arm with six degrees of freedom, which can reflect the strength of the interaction of the robot model with its environment\u2019s model (Fig. 1). The arm motion trajectory in this control mode without any delay will be repeated in time and space by the model of the working tool with the accuracy determined by the capabilities of the robot model\u2019s control system. This trajectory should be programmed for a real space robot and transferred to space to its control system along with the information on the magnitude and direction of the reflected interaction force vector in the form of the respective signals, which are functions of time. The space robot control system should provide the execution of the program obtained from these signals",
+            " 4 2016 REMOTE CONTROL OF SPACE ROBOTS 637 Since the above-mentioned additional information is the result of the operation of the robot\u2019s sensor system, further, we will call it a sensor image. The correction signal is a function of the mismatch between the model and the real sensor images. It vanishes in the case of a zero mismatch between them. Images of a set of feature points belonging to objects of the model of the robot\u2019s environment are the simplest example of a sensor image. For example, images of the vertices of polyhedrons (Fig. 1). They are selected by a special recognizing program on the image of the environment acquired by the TV-cameras located on the robot\u2019s working tool model. Images of similar points of the real environment are generated at the time of the execution of the programmed trajectory by the robot using TV cameras mounted on both the real and model robot\u2019s working tools. Therefore, in the case of an ideal execution, images of these points should match the images of the model points. However, in reality, because of the possible inaccuracy of the environment model, they do not match"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0003827_robot.2000.846348-Figure2-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0003827_robot.2000.846348-Figure2-1.png",
+        "caption": "Fig. 2: Model of the system",
+        "texts": [
+            " We further define the neighborhood equilibrium which means that the system can be shifted to another equilibrium state close to the original one even when the current equilibrium is momentarily broken. We show that the neighborhood equilibrium can be achieved only when the rolling based redundancy is not zero. We evaluate the robustness of the equilibrium state by utilizing the rotating angle, where the system loses the neighborhood equilibrium. We show several numerical examples to confirm our idea. 2 Modeling Fig2 shows the grasp of m objects by n fingers, where finger j contacts with object i, and additionally object i has a common contact point with object 1. CR, C B ~ (i = 1,. . . , m) and C F j k ( j = 1,. . . , n, k = l,...,cj> denote the coordinate systems fixed at the base, at the center of gravity of the object i and at the finger link including the kth contact of finger j, respectively. Let p B i E R3 and R B ~ E R3x3 be the position vector and the rotation matrix of C B ~ , and P F j k E R3 ER, respectively"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0003340_0010-4361(89)90337-6-Figure4-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0003340_0010-4361(89)90337-6-Figure4-1.png",
+        "caption": "Fig. 4 Geometry and position of the specimen for the MFWtest",
+        "texts": [
+            " The initial notches of the specimens were also made in these two directions by driving a razor blade into the material. The geometry of the specimen is shown in Fig. 3. The direction of the crack perpendicular to the MFD is referred to as the transverse (T) direction, that parallel to the MFD as the longitudinal (L) direction. 2. Instrumented modular falling weight test Modular falling weight tests were performed on an instrumented modular falling weight device at room temperature (the same data collecting unit as in the NII test was used). The rectangular plates (Fig. 4), which were 50 x 50 x 3.3 mm in size, were impacted by a metallic striker (tip radius 20 mm). The load to the maximum peak (Fp) and the energy to punch a hole through the thickness of the plates (Eto Q were again measured (impact velocity = 6.26 m s- ' ) . 224 COMPOSITES. MAY 1989 RESULTS AND DISCUSSION Notched Izod impact test 1. Nil-test data The principle load-displacement curves obtained from the Nil tests with the pure LCP resin and its glass- and carbon-fibre reinforced composites are shown in Figs 5a-f"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0003124_70.88068-Figure8-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0003124_70.88068-Figure8-1.png",
+        "caption": "Fig. 8. Wrench representation of a line contact on a polyhedral surface.",
+        "texts": [
+            " Line Contact In general, line contact between the object and the obstacle environment can be expected along the surface of the quadratic representation of the object, along the edge of the polyhedra representation, or on the surface of polyhedra located at the edges of the obstacle environment (Fig. 7). In line-contact configurations, the contact axis can be treated as the axis of the wrench acting between the objects in contact. In general, this wrench can be defined with respect to its coordinate frame as shown in Fig. 8 [9] The above contact configuration can be expected in the case shown in Fig. 7(c). The Plucker line coordinate values of the wrench axis with respect to compliance frame can be written as al = wI sin 8 + w2 cos 8 a 3 = w l c o s 8 - w 2 s i n 8 a4= w4 sin 8- w3d cos 8 Us= W,5 + d W 2 a6 = w4 cos 8 + wjd sin 8. IEEE TRANSACTIONS ON ROBOTICS AND AUTOMATION, VOL. 5 , NO. 4 , AUGUST 1989 537 Pr icker Line the Wrench A x i s 2 c Coordinotes o f I \u2018A2 From the above we can obtain the following relationship between the location and orientation ( d , 8) of the wrench axis (contact line) in terms of the measured values (Fig. 8): (14) The equation of a plane perpendicular to the line can be written as - d = ( u 4 / u 2 ) cos 8 - ( a 6 / a 2 ) sin 8. XI+ Y m + Z n = d (15) where (X, Y, and Z ) are coordinates of any point on the plane, ( I , m, n) are the direction cosines of the plane normal, at this point, and d is the distance between this normal and the origin of the compliance frame. Therefore, the equation of this plane becomes X c o s O + Z s i n O = d (16) or The above plane contains pencil of lines which pass through point ( - 04/a2,0, a6/(12)"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0001860_0094-114x(95)00022-q-Figure1-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0001860_0094-114x(95)00022-q-Figure1-1.png",
+        "caption": "Fig. 1. A machine for blanking sheet metal and its coupling graph; an equivalent over-constrained coupling network and its coupling graph.",
+        "texts": [
+            " Such overclosure is to be found in the epicyclic gear train shown in Fig. 3 and used as an example in this paper. However, it is not the actions attributable to over-constraint that concern this paper but those attributable to active couplings. Before the general procedure is explained a simple machine is used to illustrate the concept of circuit action and to show how an active coupling can be replaced by an equivalent passive coupling to create an over-constrained coupling network. 5. ACTIONS IN MACHINES Figure 1 shows a slider-crank mechanism for use in a machine driven by an electric motor to punch or blank parts from sheet metal. The interaction between punch and sheet metal therefore provides the second action coupling, a sink for mechanical power. This example is simple enough for intuition to identify the circuit actions that are a consequence of the active couplings present in the machine. Figure 1 also shows the coupling graph of the machine with three numbered circuits. Clearly the circuit action in circuit 2 is the action transmitted by couplings b, c. The action in circuit 2 attributable to active couplings is a force through the centre line of the connecting rod. The connecting rod and its bearings could transmit other actions because of the over-constraint inherent in this planar mechanism but those actions do not concern us except to state that the unit actions attributable to over-constraint could be found by using all six action screw coordinates in the reciprocity equation [5]",
+            " The Oldham coupling is an indirect coupling comprising two prismatic couplings in series that permits translational velocities u, v but not rotation t. For an active coupling transmitting force, as in the punch/metal interaction, the replacement passive coupling is a binary link with two R couplings. Circuit actions attributable to active couplings 1005 Consider intuitively now the task of closing the circuits of the equivalent over-constrained network. The active coupling of the electric motor, represented by a dashed line in the coupling graph of the machine shown in Fig. 1 is replaced by two prismatic couplings e and f i n series with an intermediate member 5 as in an Oldham coupling. Bearing a completes circuit !. Before the closure of other circuits, circuit 1 is simply stiff with the crank angle determinate; the circuit is neither mobile nor over-constrained. The closure of Circuit 2 does not alter this. However to close circuit 3 the binary link replacing the second active coupling of the machine needs to be fitted. This can only be done if the points of attachment are precisely spaced and exactly equal to the distance between the bearings of that binary link"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0002220_s0263574797000544-Figure3-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0002220_s0263574797000544-Figure3-1.png",
+        "caption": "Fig . 3 . Geometric constraints .",
+        "texts": [
+            " The errors u X d ( t 1 1) 2 X ( t 1 1) u and u X ~ d ( t 1 1) 2 X ~ ( t 1 1) u are calculated . We assume that convergence is ensured when these two errors are less than 10 2 6 during several integration steps . The constraints have been decomposed into \u2018\u2018geometric\u2019\u2019 and \u2018\u2018dynamic\u2019\u2019 constraints . 3 . 2 . 1 Geometric constraints . First of all , an obvious geometric constraint is due to the fact that the feet touch the ground all over the simulation . This induces a maximal vertical value for the trunk center of gravity position which must not be exceeded during simulation (see Figure 3) . Thus : z G ( t ) # z G max (10) In fact , this means that the real trunk trajectory along the vertical axis must asymptotically converge on the desired trunk vertical trajectory . Moreover , the angular perturbation cannot be imposed at random but must also satisfy this constraint . Since it is not possible to produce ankle torques without feet , another geometric constraint is related to the feet lengths . In other words , the feet must be long enough so that the ground reaction application point (also called Zero Moment Point) is within the feet support area . The following constraints are therefore taken into account : where X p and Y p are , respectively , the antero-posterior and lateral positions of the Zero Moment Point with respect to the ankles axis . L h and L t are , respectively , the distance from heel to ankle axis and from angle axis to toes (see Figure 3) . 3 . 2 . 2 Dynamic constraints . In the method described above , the legs and the arms seem to play the same part . However , in the erect stance the legs have a supporting role that the arms have not . Moreover , the arms motion can be considered as a perturbation to the trunk equilibrium . 16\u201318 In this way , as the arms motions are known , the forces / moments imposed by the trunk to the arms to insure these motions are also known . So they must be imposed as equality constraints in the optimization problem "
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0000409_j.jmatprotec.2016.05.028-Figure5-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0000409_j.jmatprotec.2016.05.028-Figure5-1.png",
+        "caption": "Figure 5: Cutting and placement method of flax and polypropylene sheets to manufacture doubly-curved composites",
+        "texts": [
+            " tH (50min) tC (25min)tI (40min) 0 Vacuum Pressure (mbar) Temperature (\u00b0C) 1000 Time (minutes) 30 500 180 25 VAOC 60 90 120 0 120 8/27 The material preparation for singly-curved composites was quite straightforward because 200 mm by 300 mm rectangular layers of flax fabric and polypropylene sheets were cut and placed inside the singly-curved mould before the composites were manufactured. However, it was challenging to cut and place the fibre reinforcement and polypropylene sheets in the doublycurved mould prior to manufacturing. Figure 5 shows how the flax fibres and the polypropylene sheets were cut and how they were placed inside the mould to manufacture doubly-curved composites. Each fibre layer was placed on the previous one and the gap was rotated until the gap of the lower layer was completely covered by the mass of the upper layer. The gap could have created a mechanically weaker section in the composite but this rotation in the placement method was intended to make the stress concentration uniform around the composite. The rotation also helped to stabilize the mass of fibres throughout the composite so that its density was as homogeneous as possible",
+            " Given these dimensions, each circular layer of flax and PP had to have a circumference of 44 cm and an area of 308 cm 2 . Since it was a conversion from a 2D shape to a 3D shape, it was not possible for a circular layer to take the shape of a hemisphere. Hence, a segment was cut from each circular layer so that it becomes a hemisphere when placed on the mould. The angle \u201c\u019f\u201d measured in radians and the radius \u201cr\u201d of the segment were calculated using Equation 1 where the circumference of the hemisphere was equated to the length of the circular layer (Figure 5). ( ) (1) At an angle of 60\u00b0 the radius of the circular layer was calculated to be 8.4 cm, and these dimensions were used to cut the fibre and the PP layers as shown in Figure 3. Once the layers were cut as shown in Figure 5, they were placed in the doubly-curved mould like a filter paper is placed in a funnel. Due to the doubly-curved mould, the layers pressed themselves with a force enough to vacuum bag the mould. Once the 3S3C manufacturing cycle was developed, singly-curved and doubly-curved composites were manufactured using flax and PP layers. Composites were manufactured at three distinct thickness values for each shape and flax reinforcement. Table 2 presents the plan to manufacture these composites using coarse twill, unidirectional and fine twill flax reinforcements, respectively"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0000871_ultsym.2016.7728402-Figure1-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0000871_ultsym.2016.7728402-Figure1-1.png",
+        "caption": "Fig. 1: Schematics of the US beam reflection pattern at needle interface, which creates shadowing below the needle.",
+        "texts": [
+            " The main contributions of this study are: (1) fast detection of a shadowed region that corresponds to the needle in the volume and (2) automated selection of the optimal scan plane from the 3D US volume for in-plane visualization of the needle even for very large insertion angles. 978-1-4673-9897-8/16/$31.00 \u00a92016 IEEE 2016 IEEE International Ultrasonics Symposium Proceedings During acquisition of US images, the transmitted beams passing through the tissue around the needle are selectively attenuated (see Fig. 1). This is due to the fact that first, at the needle interface a great portion of the waves is reflected back to the probe. Second, depending on the effective beam width, waves passing around the needle are weak. And third, in cases of large insertion angles, needle reflects the majority of US beams away from the probe and backscattered beams remain undetected by the transducer [9]. Therefore, a dark-colored representation of tissues below the needle occurs, which is called a shadowed region. As discussed in the previous section, the shadow of the needle is a 3D surface with a defined thickness"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0001197_978-3-658-12701-5-Figure6.8-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0001197_978-3-658-12701-5-Figure6.8-1.png",
+        "caption": "Figure 6.8: Planar robot \u2014 Singularity avoidance, feasible configuration",
+        "texts": [
+            "7 on the next page, \ud835\udefe = arctan (\ud835\udc593 sin (\ud835\udc5e3) , \ud835\udc592 + \ud835\udc593 cos (\ud835\udc5e3)) \ud835\udf05 = arctan (\ud835\udc66E \u2212 \ud835\udc591 sin (\ud835\udc5e1) , \ud835\udc65E \u2212 \ud835\udc591 cos (\ud835\udc5e1)) , one can find an expression for the remaining joint angle \ud835\udc5e2, \ud835\udc5e2 = \ud835\udf05 \u2212 \ud835\udc5e1 \u2212 \ud835\udefe. In each of the inverse kinematics approaches presented in Sections 6.1.2 to 6.1.4, one joint angle is treated as redundant, i.e. assumed to be known, as it is determined separately in terms of the presented separation method. The following example will show that the redundant joint angle \ud835\udc5er may not assume arbitrary values for a given end-effector position but that an admissible interval can be determined under certain conditions. In Figure 6.8 the third joint is treated as redundant, i.e. \ud835\udc5er = \ud835\udc5e3. As described in Section 6.1.4, links 2 and 3 can be reduced to a single, virtual link with length \ud835\udc4e. From Figure 6.9 on the next page one can see that it is geomet- rically not feasible to find an inverse kinematics solution for the suggested choice of \ud835\udc5e3, since no connection point for the first link and the virtual link can be found. Obviously, the feasibility of the inverse kinematics problem is bounded in the pose where the virtual link and the remaining link 1 assume a singular configuration. From the configuration depicted in Figure 6.8 on the previous page it can be seen that such a singularity will occur if the length of the virtual link, \ud835\udc4e = \ud835\udc4e (\ud835\udc5e3), is decreased to the difference between the distance of the end-effector position from the origin, \ud835\udc59, and the length of link 1, \ud835\udc591, i.e. \ud835\udc4esing = \ud835\udc59 \u2212 \ud835\udc591. From rearranging (6.3) on page 59, \ud835\udc5e3,sing = \u00b1 arccos ( \ud835\udc4e2 sing \u2212 \ud835\udc5922 \u2212 \ud835\udc5923 2 \ud835\udc592 \ud835\udc593 ) , bounds for \ud835\udc5e3 can be obtained at which a singularity of the virtual link and link 1 will occur. According to the procedure for obtaining bounds for \ud835\udc5e3 above, bounds for \ud835\udc5e2 can be found for the inverse kinematics case \ud835\udc5er = \ud835\udc5e2 as described in Section 6"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0002669_jsco.1998.0200-Figure4-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0002669_jsco.1998.0200-Figure4-1.png",
+        "caption": "Figure 4.",
+        "texts": [
+            " Kova\u0301cs and Hommel (1990)) that its solution could be greatly simplified if the determining equation could be solved by a sequence of lower-degree equations. Example 1.3. The inverse kinematics problem of the robot ROMIN (see GonzalezLopez and Recio (1993)) can be solved by many different methods, but it is specifically interesting since it is one of the few examples in which a new \u201clazy evaluation method\u201d for solving systems of equations, the dynamic evaluation procedure (Duval, 1990), has been used. Given a position (a, b, c) of the tip point P and the length of the links m,n (see Figure 4), the algebraic kinematic equations of the ROMIN are: \u2212s1(mc2 + nc3) = a, c1(mc2 + nc3) = b, ms2 + ns3 = c, plus the trigonometric identities: s2 1 + c21 = 1, s2 2 + c22 = 1, s2 3 + c23 = 1. After a triangulation, the fourth degree equation determining the angle \u03b82 is f(s2, c2) = (\u22124m2a2 \u2212 4m2c2 \u2212 4m2b2)c22 + (4mcn2 \u2212 4mc3 \u2212 4 cmb2 \u2212 4m3c \u22124 cma2)s2 \u2212 2n2c2 + n4 + c4 + b4 + a4 \u2212 2n2a2 + 2 c2a2 + 6m2c2 \u22122m2n2 +m4 + 2 b2a2 + 2m2b2 + 2m2a2 \u2212 2n2b2 + 2 c2b2. Now, this equation, can be rewritten as f(s2, c2) = g(h(s2, c2)) + q(s2 2 + c22 \u2212 1), where g(x) = n4 + b4 + a4 \u2212 2n2c2 + 2 c2a2 + 2 b2a2 + 2m2c2 \u2212 2m2n2 \u2212 2n2a2 + 2 c2b2 +c4 \u2212 2n2b2 \u2212 2m2b2 \u2212 2m2a2 +m4 +(4mcn2 \u2212 4mc3 \u2212 4 cmb2 \u2212 4m3c\u2212 4 cma2)x +(4m2a2 + 4m2b2 + 4m2c2)x2, h(s2, c2) = s2 and q = 4(b2 + a2 + c2)m2"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0001090_icpeices.2016.7853155-Figure1-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0001090_icpeices.2016.7853155-Figure1-1.png",
+        "caption": "Fig. 1: The Force Inspection on the InPd System",
+        "texts": [
+            " MATHEMATICAL MODELlNG OF REAL TIME SYSTEM A pendulum rod mounted on a movable cart platform is jointly called a cart pendulum system. The physical system in which the designed controller is applied possesses a linear cart motion on a predefined track and a single pendulum rod is attached on the cart platform. Thus the combined system is termed as a linear 1 stage InPd system. In this section a simplified linearized mathematical model is derived from the actual non linear equation describing the dynamic feature of a cart-inverted pendulum system. In fig. 1 the cart mass is denoted as-M, m-the mass of pendulum rod, I is the distance of the rod from the revolution axis to the centroid, e - the deviation angle of the rod from vertical rising location. I-the moment of inertia (MOl) of pendulum bar about its centroid position. F is the input force on the cart, F gbeing the resultant of horizontal and vertical component of the disturbing force on the pendulum rod. P - the vertical component of the reaction force at the pivotal point and the horizontal component is N"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0000984_j.ifacol.2016.10.148-Figure3-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0000984_j.ifacol.2016.10.148-Figure3-1.png",
+        "caption": "Fig. 3. Flying formation of UAVs for the concentration gradient measurements.",
+        "texts": [
+            " The computational grid for the estimator was adapted dynamically in order to achieve greater computational efficiency. In this work, this approach is implemented using a rigid formation of UAVs whose utility is twofold. First, the concentration gradients appearing in the desired inertial velocities of the leader UAV, and thus in its guidance, are computed via the use of the point concentration measurements from the \u201cfollower\u201d-UAVs that form a cross-shape formation around the \u201cleader\u201d UAV as shown in Figure 3. Secondly, the estimator model is modified to incorporate the additional concentration measurements provided by the follower UAVs. The remainder of this paper is organized as follows: Section 2 provides the mathematical model for the gaseous release process (plant) and the estimator modified to account for the additional plant information from the follower UAVs, as well as the UAV dynamical model and guidance scheme. Section 3 summarizes the numerical implementation procedure of the proposed approach",
+            " The desired inertial velocities for the UAV ere chosen in terms of the concentration estimation error and the error gradien s through the appropriate choice of a Lyapunov functional. The comput tional grid for the estimator was adapted dynamically in order to achiev greater computational efficiency. In this w k, this pproach is imple ented using a rigid formation f UAV whose utility is twofold. First, the concentration gradient appearing in the desired inertial velocities of the leader UAV, and thus in its guidance, are computed via the use of the point concentration measure en s from \u201cfoll wer\u201d-UAVs that form cross-sh pe formation around the \u201cleader\u201d UAV as shown in Figure 3. Secondly, the estimator model is modified to incorpo ate th ad itional con cen ration measurements provided by the follower UAVs. The remainder of this paper is organized as follows: Section 2 provides the mathematical model for the gaseous release process (plant) and the estimator modified to account for the addi ional plant information from the foll wer UAVs, as well as the UAV dynamical model and guidance scheme. S ction 3 summarizes the numerical implementation proc dure of the propos d approach",
+            " The computational grid for the estimator was adapted dynamically in order to achieve greater computational efficiency. In this work, this approach is implemented using a rigid formation of UAVs whose utility is twofold. First, the concentration gradients appearing in the desired inertial velocities of the leader UAV, and thus in its guidance, are computed via the use of the point concentration measurements from the \u201cfollower\u201d-UAVs that form a cross-shape formation around the \u201cleader\u201d UAV as shown in Figure 3. Secondly, the estimator model is modified to incorporate the additional concentration measurements provided by the follower UAVs. The remainder of this paper is organized as follows: Section 2 provides the mathematical model for the gaseous release process (plant) and the estimator modified to account for the additional plant information from the follower UAVs, as well as the UAV dynamical model and guidance scheme. Section 3 summarizes the numerical implementation procedure of the proposed approach",
+            " The computational grid for the estimator was adapted dynamically in order to achieve greater computational efficiency. In this work, this approach is implemented using a rigid formation of UAVs whose utility is twofold. First, the concentration gradients appearing in the desired inertial velocities of the leader UAV, and thus in its guidance, are computed via the use of the point concentration measurements from the \u201cfollower\u201d-UAVs that form a cross-shape formation around the \u201cleader\u201d UAV as shown in Figure 3. Secondly, the estimator model is modified to incorporate the additional concentration measurements provided by the follower UAVs. The remainder of this paper is organized as follows: Section 2 provides the mathematical model for the gaseous release process (plant) and the estimator modified to account for the additional plant information from the follower UAVs, as well as the UAV dynamical model and guidance scheme. Section 3 summarizes the numerical implementation procedure of the proposed approach",
+            " The computational grid for the estimator was adapted dynamically in order to achieve greater computational efficiency. In this work, this approach is implemented using a rigid formation of UAVs whose utility is twofold. First, the concentration gradients appearing in the desired inertial velocities of the leader UAV, and thus in its guidance, are computed via the use of the point concentration measurements from the \u201cfollower\u201d-UAVs that form a cross-shape formation around the \u201cleader\u201d UAV as shown in Figure 3. Secondly, the estimator model is modified to incorporate the additional concentration measurements provided by the follower UAVs. The remainder of this paper is organized as follows: Section 2 provides the mathematical model for the gaseous release process (plant) and the estimator modified to account for the additional plant information from the follower UAVs, as well as the UAV dynamical model and guidance scheme. Section 3 summarizes the numerical implementation procedure of the proposed approach",
+            " The desired inertial velocities for the UAV ere chosen in terms of the concentration estimation error and the error gradien s through the appropriate choice of a Lyapunov functional. The comput tional grid for the estimator was adapted dynamically in order to achiev greater computational efficiency. In this w k, this pproach is imple ented using a rigid formation f UAV whose utility is twofold. First, the concentration gradient appearing in the desired inertial velocities of the leader UAV, and thus in its guidance, are computed via the use of the point concentration measure en s from \u201cfoll wer\u201d-UAVs that form cross-sh pe formation around the \u201cleader\u201d UAV as shown in Figure 3. Secondly, the estimator model is modified to incorpo ate th ad itional con cen ration measurements provided by the follower UAVs. The remainder of this paper is organized as follows: Section 2 provides the mathematical model for the gaseous release process (plant) and the estimator modified to account for the addi ional plant information from the foll wer UAVs, as well as the UAV dynamical model and guidance scheme. S ction 3 summarizes the numerical implementation proc dure of the propos d approach",
+            " The computational grid for the estimator was adapted dynamically in order to achieve greater computational efficiency. In this work, this approach is implemented using a rigid formation of UAVs whose utility is twofold. First, the concentration gradients appearing in the desired inertial velocities of the leader UAV, and thus in its guidance, are computed via the use of the point concentration measurements from the \u201cfollower\u201d-UAVs that form a cross-shape formation around the \u201cleader\u201d UAV as shown in Figure 3. Secondly, the estimator model is modified to incorporate the additional concentration measurements provided by the follower UAVs. The remainder of this paper is organized as follows: Section 2 provides the mathematical model for the gaseous release process (plant) and the estimator modified to account for the additional plant information from the follower UAVs, as well as the UAV dynamical model and guidance scheme. Section 3 summarizes the numerical implementation procedure of the proposed approach",
+            " The computational grid for the estimator was adapted dynamically in order to achieve greater computational efficiency. In this work, this approach is implemented using a rigid formation of UAVs whose utility is twofold. First, the concentration gradients appearing in the desired inertial velocities of the leader UAV, and thus in its guidance, are computed via the use of the point concentration measurements from the \u201cfollower\u201d-UAVs that form a cross-shape formation around the \u201cleader\u201d UAV as shown in Figure 3. Secondly, the estimator model is modified to incorporate the additional concentration measurements provided by the follower UAVs. The remainder of this paper is organized as follows: Section 2 provides the mathematical model for the gaseous release process (plant) and the estimator modified to account for the additional plant information from the follower UAVs, as well as the UAV dynamical model and guidance scheme. Section 3 summarizes the numerical implementation procedure of the proposed approach",
+            " It takes the form of a Luenberger observer supplemented with an output injection term, which is specified by the difference between a \u201ctrue\u201d (measured) concentration and a state estimate at the current sensor locations, multiplied by the filter gain matrix \u0393 and the dual C\u2217 of the output operator C evaluated at the current sensor location \u02d9\u0302x = Ax\u0302+ C\u2217\u0393C(x\u2212 x\u0302), (5) where x\u0302(t) = \u3008c\u0302\u3009(t, \u00b7, \u00b7, \u00b7) denotes the estimated concentration state. Unlike the earlier effort Egorova et al. (in print, 2016), the filter gain \u0393 is now a positive definite matrix \u0393 =   \u039311 0 \u00b7 \u00b7 \u00b7 ... . . . 0 0 \u00b7 \u00b7 \u00b7 \u0393mm   (6) where m is the number of sensors. Here m = 7 which accounts for the leader UAV (numbered as 4) and the six follower UAVs in the 3D rigid formation (cross-shape) of Figure 3. Please note that a full gain matrix \u0393 would require information exchange between all UAVs whereas a diagonal gain matrix \u0393 significantly reduces the communication requirements at the possible expense of estimator performance. Figure 2 shows the free-body diagram for a UAV climbing at a flight path angle \u03b3 and a bank angle \u03c6. The derivation of the UAV equations of motion are based on the pointmass model of a fixed-wing aircraft Beard and McLain (2012); Zhao and Tsiotras (2010); Menon et al. (2012)",
+            " The forces L(CL) and D(CL) are the lift and drag forces, respectively, and are given by L(CL) = 1 2 \u03c1V 2 a SCL, D(CL) = 1 2 \u03c1V 2 a SCD(CL), (8) where Va is the airspeed, \u03c1 is air density, S is the planform area of the UAV wing, CL and CD are the lift and drag coefficients which are related via CD = CDp + C2 L \u03c0eAR , (9) where CDp is the parasitic drag caused by the shear stress of air moving over the wing, b is the wing span, AR b2 S is the wing aspect ratio, and e is the Oswald efficiency factor Beard and McLain (2012). The thrust T , the bank angle \u03c6, and the lift coefficient CL are the control inputs to the UAV model (7). The flying formation is assumed rigid throughout the whole estimation procedure, i.e., the \u201cfollower\u201d-UAVs maintain a constant distance from the \u201cleader\u201d-UAV at all times. Figure 3 depicts the top view of the formation of UAVs. The guidance scheme for the \u201cleader\u201d-UAV is coupled to the estimator (5) performance through the state-estimation error e(t) = x(t) \u2212 x\u0302(t). Combining Eq. (2) and Eq. (5) yields the governing equation for the state estimation error e\u0307 = (A\u2212 C\u2217\u0393C)e+ S(t,\u0398c) = Acle+ S(t,\u0398c). (10) The Lyapunov redesign method is used to derive the control inputs for the \u201cleader\u201d-UAV, according to which the flying formation is guided towards areas of larger estimation error",
+            " The forces L(CL) and D(CL) are the lift and drag forces, respectively, and are given by L(CL) = 1 2 \u03c1V 2 a SCL, D(CL) = 1 2 \u03c1V 2 a SCD(CL), (8) where Va is the airspeed, \u03c1 is air density, S is the planform area of the UAV wing, CL and CD are the lift and drag coefficients which are related via CD = CDp + C2 L \u03c0eAR , (9) where CDp is the parasitic drag caused by the shear stress of air moving over the wing, b is the wing span, AR b2 S is the wing aspect ratio, and e is the Oswald efficiency factor Beard and McLain (2012). The thrust T , the bank angle \u03c6, and the lift coefficient CL are the control inputs to the UAV model (7). The flying formation is assumed rigid throughout the whole estimation procedure, i.e., the \u201cfollower\u201d-UAVs maintain a constant distance from the \u201cleader\u201d-UAV at all times. Figure 3 depicts the top view of the formation of UAVs. The guidance scheme for the \u201cleader\u201d-UAV is coupled to the estimator (5) performance through the state-estimation error e(t) = x(t) \u2212 x\u0302(t). Combining Eq. (2) and Eq. (5) yields the governing equation for the state estimation error e\u0307 = (A\u2212 C\u2217\u0393C)e+ S(t,\u0398c) = Acle+ S(t,\u0398c). (10) The Lyapunov redesign method is used to derive the control inputs for the \u201cleader\u201d-UAV, according to which the flying formation is guided towards areas of larger estimation error. For this purpose, the choice of the Lyapunov functional follows Demetriou (2010) E = \u2212\u3008e,Acle\u3009, (11) where \u3008\u00b7, \u00b7\u3009 denotes the L2(\u2126) inner product. The time derivative of the Lyapunov function along (10) is E\u0307 = \u22122|Acle|2+ m\u2211 i=1 \u0393ii(\u03b5i\u03b5XiX\u0307i+\u03b5i\u03b5Yi Y\u0307i+\u03b5i\u03b5ZiZ\u0307i), (12) where \u03b5i, \u03b5Xi , \u03b5Yi , and \u03b5Zi , i = 1, . . . ,m are the concentration estimation error and the error gradients at each sensor location. The summation above is performed over the total Fig. 3. Flying formation of UAVs for the concentration gradient measurements. number of sensors m. Stability arguments require E\u0307 \u2264 0. Cancelling the indefinite terms in the expression yields the desired inertial velocities for each of the UAVs X\u0307d i = \u2212kX\u0393ii\u03b5i\u03b5Xi Y\u0307 d i = \u2212kY \u0393ii\u03b5i\u03b5Yi for i = 1, ...,m, Z\u0307d i = \u2212kZ\u0393ii\u03b5i\u03b5Zi (13) where kX , kY , and kZ are user-defined positive guidance gains. On the other hand, the condition for rigid formation of the UAVs requires that X\u0307d 1 = . . . = X\u0307d m, Y\u0307 d 1 = ",
+            " (17) For the feasibility of the proposed control algorithm, the control inputs should stay within certain ranges 0 \u2264 T \u2264 Tmax, |\u03c6| \u2264 \u03c6max, 0 \u2264 CL \u2264 CLmax. (18) Tmax is chosen according to the engine specifications for a particular UAV. The maximum attainable lift coefficient CLmax should be smaller than the lift coefficient at aircraft stall. The choice of the maximum available bank angle \u03c6max is dictated by the maximum load factor (n L Mg ): 1 \u03c6max \u2264 nmax. (19) Recapitulating, the six follower UAVs provide their concentration measurements to the estimator (5) and ensure that they maintain a rigid formation as depicted in Figure 3. The leader UAV also provides its concentration measurement to the estimator (5) and to generate the output error \u03b5i = \u03b54 in the guidance (13). Using the fact that all UAVs maintain a rigid formation, then the Lyapunov-based guidance, as described by the desired velocities (13), is used as an input to the UAV equations of motion to produce the UAV control signals (17) while enforcing the control constraints (18), (19). Since the guidance (13) requires the knowledge of spatial gradients, the measurements of the six follower UAVs are also used to compute them via (14)"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0000544_9781119268628-Figure7.1-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0000544_9781119268628-Figure7.1-1.png",
+        "caption": "Figure 7.1. Experimental setup (adapted from [KIT 09])",
+        "texts": [
+            " The two persons in each pair should be acquainted with each other very well and know each other\u2019s driving attitude and knowledge about driving, so they are expected to provide information necessary for safe and enjoyable driving to the driver when their partner is driving along an unfamiliar route. Fast Externally-paced Navigation 97 Each pair will participate in three sessions with different purposes. Each session consists of a set of two on-road drives with driving times ranging from 30 to 90 min, followed by an interview. The purpose of the interview is to determine the information needed for safe and enjoyable driving. Each drive is videotaped using two digital video cameras as shown in Figure 7.1: one for the outside view (camera B), and the other for the driver and navigator view (camera A). that his/her partner is completely familiar with. The partner thus serves as a human navigator who is expected to provide information that will help the driver to make the journey safe and enjoyable. In the second drive in the first session, the roles are reversed, that is the person who drives in the first drive serves as the navigator, and vice versa. The navigator is expected to provide information necessary for safe and enjoyable driving to the partner, i"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0000758_978-981-10-0471-1_34-Figure3-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0000758_978-981-10-0471-1_34-Figure3-1.png",
+        "caption": "Fig. 3 (a) CAD prototype and (b) final assembly of AE sensor and magnetic mounting",
+        "texts": [
+            " This is appropriate for the application with 3D printer which has rough surface and vibration during the working process. Therefore, vacuum silicone grease was chosen as a sensor coupling in this study. In order to use grease as sensor coupling, clamp was required. However, the dimension of the 3D printer motor block caused commercial mountings to be not so applicable. Therefore, AE sensor mounting was design using CAD software to perfectly match the dimension of the motor block. The rendering of CAD file of the mounting is shown in Fig. 3a and can be downloaded at [12]. Then, the mounting was fabricated using 3D printer which gave freedom in designing and also low-cost product. After that magnets and a bolt were assembled to the printed object to complete the fabrication as shown in Fig. 3b. The screw bolt was utilized in the design to help secure the sensor and benefits in removing trapped air between surfaces. In addition, AE sensor is attach to motor block using double-sided tape as a simple connecting agent and collect data for normal working condition to test the effectiveness of magnetic mounting and vacuum grease as preconditioned experiment. This is to use as a reference data to compare with tested data in the main experiment. An overview of the test is shown in block diagram in Fig"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0003685_itsc.1997.660569-Figure2-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0003685_itsc.1997.660569-Figure2-1.png",
+        "caption": "Figure 2: Bond graph for a compound planetary gear",
+        "texts": [
+            " Note that the bond graph method is well suited for modeling a power system such as an automatic transmission, because the method is based on generalized power variables. 2. IN-GEAR AND UPSHIFT MODELS 2.1 Compound planetary gear model A compound planetary gear being modeled in this paper consists of two planetary gears, as shown in Figure 1. This is one of the possibl automatic transmission configurations, and other configurations are also possible. The bond graph of the compound planetary gear in Figure 1 is shown in Figure 2. From Figure 2, the following relationship is derived [7], [8], [9 ] : The clutch control schemes for achieving various gear positions are given in Table 1. 0-7803-4269-0/97/$10.00 0 1998 IEEE 759 First gear 1-2 shift Fisrt clutch Second clutch Third clutch Fourth clutch 1-2 band Reverse band ON OFF OFF OFF ON OFF ON h OFF OFF ON OFF 2-1 shift 1 ON .U- OFF OFF I ON I OFF Secons gear 2-3 shift ON ON OFF OFF ON OFF U ON h OFF .II OFF 3-2 shift Third gear 3-4 shift 4-3 shift Fourth gear Reverse gear 0 ON ii OFF l i OFF OFF ON ON OFF OFF OFF OFF ON ON fi OFF OFF OFF ON ON U OFF OFF OFF ON ON ON OFF OFF ON OFF OFF OFF OFF ON 2"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0001235_b978-0-08-099425-3.00005-4-Figure5.2-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0001235_b978-0-08-099425-3.00005-4-Figure5.2-1.png",
+        "caption": "FIGURE 5.2",
+        "texts": [
+            " As a final introductory comment it should be noted that what is referred to as a \u2018tyre\u2019 model is actually a tyre-road interaction model. Placing the same tyre on a different surface will modify its behaviour substantially, as anyone who has ever driven on ice will have noticed. The most versatile tyre-model architectures allow the tyre-specific model to encounter different surfaces during the course of a single manoeuvre, but require proportionally more data to support this functionality. To assist with the description of the forces and moments generated by a tyre, an axis frame shown in simplified form in Figure 5.2 has been defined by the SAE (1976). SAE tyre axis system. In this frame the X-axis is the intersection of the wheel plane and the road plane with the positive direction taken for the wheel moving forward. The Z-axis is perpendicular to the road plane with a positive direction assumed to be acting downwards. The Y-axis is in the road plane and its direction dictated by the use of a right-handed orthogonal axis frame. The angles a and g represent the slip angle and camber angle respectively. The SAE frame will be used throughout this text unless stated"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0002305_1.2833509-Figure2-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0002305_1.2833509-Figure2-1.png",
+        "caption": "Fig. 2",
+        "texts": [
+            " Altogether there were 6 different shaft/pole piece combina tions which were tested, and one configuration in which only the magnetic field strength in the gap was changed, for a total of 7 configurations. These are listed in Table 1. (The magnetic field was changed by varying the amount of permanent magnet material in the magnetic circuit.) There were two general classes of seal geometry which were tested, one with a tapered stage on the shaft (configurations A - F ) and one with a tapered stage on the pole (configuration G). These are shown in Fig. 2, Contributed by the Tribology Division of THE AMERICAN SOCIETY OF MECHANI CAL ENGINEERS and presented at the ASME/STLE Joint Tribology Conference, San Francisco, Calif., October 13-17,1996. Manuscript received by the Tribology Division February 21, 1996: revised manuscript received June 28, 1996. Paper No. 96-Trib-31. Associate Technical Editor: R, F. Salant. The spindle speed was monitored by a tachometer output signal on the motor controller (the speeds were independently verified using a high speed strobe light)"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0000287_indicon.2015.7443715-Figure1-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0000287_indicon.2015.7443715-Figure1-1.png",
+        "caption": "Fig. 1. Engagement Geometry - Proportional Navigation.",
+        "texts": [
+            " Section II presents the dynamics of proportional navigation guidance. Section III presents the design of sliding mode con- 978-1-4673-6540-6/15/$31.00 c\u20dd 2015 IEEE troller while section IV presents the design of the disturbance observer and its integration of the disturbance observer with sliding mode controller. It also presents a simple stability analysis based on Lyapunov function method. Section V presents the simulation results. Section VI draws the conclusions. Consider the two dimensional engagement geometry shown in Fig. 1. We assume that (A1) Target and missile are point masses. (A2) Target and missile velocities are constant. The co-ordinate system used is that of inertial co-ordinates fixed to the surface of a flat-Earth model. Thus the components of acceleration and velocities along the two axes or directions can be integrated without having to worry about the additional terms due to the Coriolis effect. With angles as defined in Fig. 1 and using polar coordinates, the following definitions can be summarized. (a) \ud835\udc5f\ud835\udc61 and \ud835\udc5f\ud835\udc5a are the target and missile ranges and \ud835\udc5f is the relative range between the target and the missile. (b) \ud835\udf06 is the angle by which the line-of-sight (LOS) has rotated from the initial value. (c) \ud835\udefd and \ud835\udefe are the target and missile flight path angles. (d) \ud835\udc49\ud835\udc47 and \ud835\udc49\ud835\udc40 are the target and missile linear velocities. (e) \ud835\udc5b\ud835\udc47 and \ud835\udc5b\ud835\udc36 are the target and missile lateral accelera- tions. The dynamics of nonlinear proportional navigation guidance which are important for the control law formulation and are derived in [16] are reproduced below: "
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0000622_b978-0-12-800782-2.00003-8-Figure3.32-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0000622_b978-0-12-800782-2.00003-8-Figure3.32-1.png",
+        "caption": "Figure 3.32 Graphical phasor representation of superposition of torques as given in Fig. 3.31.",
+        "texts": [
+            " 240 Power Quality in Power Systems and Electrical Machines The relation between s1 (fundamental slip) and sh (1) (reflected harmonic slip) is s 1\u00f0 \u00de h \u00bc h\u03c9s1 \u03c9m \u03c9s1 \u00bc\u03c9s1 \u03c9m \u03c9s1 + h 1\u00f0 \u00de\u03c9s1 \u03c9s1 \u00bc 1 \u03c9m \u03c9s1 + h 1\u00bc h 1 s1\u00f0 \u00de: Therefore, the reflected harmonic slip, in terms of fundamental slip, is s 1\u00f0 \u00de h \u00bc s1 + h 1\u00f0 \u00de: (3-31a) For the backward rotating 5th harmonic one gets s 1\u00f0 \u00de 5 \u00bc s1 + 6 thus s1\u00bc s 1\u00f0 \u00de 5 + 6; and for the forward rotating 7th harmonic s 1\u00f0 \u00de 7 \u00bc s1 + 6 thus s1\u00bc s 1\u00f0 \u00de 7 6: Note that s 1\u00f0 \u00de 5 \u00bc 5\u03c9s1 +\u03c9m \u03c9s1 : In general, the reflected harmonic slip sh (1) for the forward and the backward rotating harmonics is a linear function of the fundamental slip s1: 241Modeling and Analysis of Induction Machines s 1\u00f0 \u00de h \u00bc s1 + h 1\u00f0 \u00de f or forward rotating field s 1\u00f0 \u00de h \u00bc s1 + h+ 1\u00f0 \u00de f or backward rotating field: (3-31b) Superposition of fundamental (h\u00bc1), fifth harmonic (h\u00bc5), and seventh harmonic (h\u00bc7) torqueTe\u00bcTe1+Te5+Te7 is illustrated in Fig. 3.31. At\u03c9m\u00bc\u03c9mrated the total electrical torque Te is identical to the load torque TL or Te1+Te5+Te7\u00bcTL, where Te1 is a motoring torque,Te7 is a motoring torque, andTe5 is a braking torque. Figure 3.32 shows the graphical phasor representation of superposition of torques as given in Fig. 3.31. 242 Power Quality in Power Systems and Electrical Machines In this section, the relation between the reflected harmonic slip sh (1) and the harmonic slip sh is determined as s 1\u00f0 \u00de h \u00bc f sh\u00f0 \u00de: The harmonic slip (without addressing the direction of rotation of the harmonic field) is sh\u00bc h\u03c9s1 \u03c9m h\u03c9s1 \u00bc h 1\u00f0 \u00de\u03c9s1 h\u03c9s1 + \u03c9s1 \u03c9m h\u03c9s1 with s1\u00bc\u03c9s1 \u03c9m \u03c9s1 : It follows that sh\u00bc h 1\u00f0 \u00de\u03c9s1 h\u03c9s1 + s1 h : With s 1\u00f0 \u00de h \u00bc s1 + h 1\u00f0 \u00de, one obtains for forward rotating harmonics s1\u00bc s 1\u00f0 \u00de h h 1\u00f0 \u00de or sh\u00bc h 1\u00f0 \u00de\u03c9s1 h\u03c9s1 + s 1\u00f0 \u00de h h 1\u00f0 \u00de h \u00bc h 1\u00f0 \u00de h + s 1\u00f0 \u00de h h h 1\u00f0 \u00de h \u00bc s 1\u00f0 \u00de h h : Similar analysis can be used for backward-rotating harmonics"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0002837_ip-b.1991.0017-Figure6-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0002837_ip-b.1991.0017-Figure6-1.png",
+        "caption": "Fig. 6 a Equivalent circuit 0 < p < 4 2 b Equivalent circuit ~ / 2 < p Q II c Phasor diagram 0 < p < s/2 d Phasor diagram si2 < p < I[",
+        "texts": [
+            "02 T 3.3 Comparison between measured and calculated results Fig. 2 shows the comparison of the measured and calculated performance of the experimental motor. The calculated result shows, like the measurements, that the relation between stator current and overexcitation ratio, V,/V,, is a V curve. 4 Discussion of results The V curve of a hysteresis motor has been clarified both by experimental result and by calculations. This phenomenon is discussed below. 4.1 Interpretation by equivalent circuit Fig. 6a shows a simplified equivalent circuit of the hysteresis motor obtained from the equivalent circuit in Fig. 4 by omitting the stator leakage impedance and stator iron-loss resistance. When the hysteresis angle, p, is greater than 4 2 , the simplified equivalent circuit becomes the same 3s that in Fig. 6b, since the equivalent rotor impedance Z, becomes capacitive, as can be seen from eqn. 3. The phasor diagrams are shown in Fig. 6c and d. Fig. 6d shows that the stator current i can lead the stator voltage when the hysteresis angle, p, increases. Fig. 7 shows the relation between the power factor of the simplified equivalent circuit shown in Fig. 6 and the magnetic state (p and p) of rotor hysteresis ring. The solid line is the locus of unity power factor. Regions to the right of the solid line correspond to a leading power factor and regions to the left correspond to a lagging power factor. 140 Simplified equivalent circuits and phasor diagrams The locus of unity power factor is given from eqns. 2 and 3 by Fig. 7 shows that the power factor of a hysteresis motor can be leading when p is greater than 4 2 and p is small. When the stator voltage of the hysteresis motor is reduced from V, to V, without changing the load torque, the B-H relation around the rotor changes from loop L, to loop L, in Fig. 3a. Replacing the loop L, by an equivalent elliptical loop, it is found that the hysteresis angle p of this elliptical loop is greater than 4 2 . The equivalent rotor impedance, 2, , becomes capacitive, and the stator current can lead the stator voltage. 4.2 Comparison with a synchronous motor If the simplified equivalent circuit of a hysteresis motor shown in Fig. 6a is changed into the circuit shown in Fig. 8, which takes the same form as that of the conventional synchronous motor, then V = jx, i + E , (7) V = jxgZ, (8) v = i,i, (9) i = i, + i, (10) E , = -jx, i, (1 1) E , = -jx, V/Z, (12) The voltage E , corresponds to the excitation voltage of a conventional synchronous motor which is basically proportional to the field current. Consideration of the mag- From Fig. 6a Substituting eqns. 8,9 and 10 into eqn. 7 gives IEE PROCEEDINGS-B, Vol. 138, NO. 3, M A Y 1991 nitude of Eo gives W O 1 l/IZ,l= 1/P (13) Eqn. 13 shows that the inverse of permeability of rotor hysteresis ring corresponds to the excitation voltage of the conventional synchronous motor. Fig. 9 shows the relation between permeability, p, and overexcitation ratio, Vo/V,, of the experimental hysteresis motor at a constant voltage V , = 0.24 pu, and with a constant load torque TL = 0.32 pu. As Vo/V, increases in the range where Vo/V, is greater than two, p decreases (or, l/p increases), as can be seen in this Figure"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0002619_1.2829168-Figure4-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0002619_1.2829168-Figure4-1.png",
+        "caption": "Fig. 4 Dual-lead worm",
+        "texts": [
+            "245 mm) are determined based on the multiplication of two sides axial modules with the worm gear teeth number (i.e., 72 teeth). Then, the pitch diameters of worm (i.e., 54.356 mm and 49.755 mm) are calculated by two times of worm gear set center distance (i.e., 133 mm) subtracts the above-mentioned pitch diameters of worm gear. By applying the developed gear set mathematical model and computer graph ics techniques, the three-dimensional computer graphs of duallead worm and worm gear surfaces are shown in Fig. 4 and Fig. 5, respectively. Since two different modules are used in the dual-lead worm, the thicknesses of worm threads are different. Thicknesses of the dual-lead worm have been calculated, mea sured at the mean pitch radius, for the worm rotation angles of 0 deg, 60 deg, 120 deg, 180 deg, 240 deg and 300 deg, respec tively. Table 2 shows the worm thickness variations measured at its nominal pitch radius of 25 mm. 5 Undercutting Regular surfaces are designed for gear drives to operate smoothly during gear meshing"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0000878_978-3-319-47614-8-Figure2.16-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0000878_978-3-319-47614-8-Figure2.16-1.png",
+        "caption": "Fig. 2.16 Transformation giving principal directions and principal stresses",
+        "texts": [
+            "18a) is r0 \u00bc ArrT r011 r012 r021 r022 \u00bc a11 a12 a21 a22 r11 r12 r21 r22 a11 a21 a12 a22 \u00f02:12\u00de Substituting, the coordinate transformation above gives the plane stress components in X\u2032Y\u2032 coordinates to be rx0 \u00bc rx \u00fe ry 2 \u00fe rx ry 2 cos 2h\u00fe sxy sin 2h ry0 \u00bc rx \u00fe ry 2 rx ry 2 cos 2h sxy sin 2h sx0y0 \u00bc rx ry 2 sin 2h\u00fe sxy cos 2h \u00f02:13\u00de Normal stresses \u03c3x\u2032 and \u03c3y\u2032 and shear stress \u03c4x\u2032y\u2032 vary smoothly with respect to the rotation angle \u03b8, in accordance with the coordinate transformation Eq. (2.13). We see that there exist a couple of particular angles where the stresses take on special values, Fig. 2.16. First, there exists an angle \u03b8p where the shear stress \u03c4x\u2032y\u2032 becomes zero. That angle is found by setting \u03c4x\u2032y\u2032 zero in the shear transformation equation of (2.13) and solving for \u03b8 (set equal to \u03b8p). The result is, rx ry 2 sin 2h\u00fe sxy cos 2h \u00bc 0 tan 2hp \u00bc 2sxy rx ry \u00f02:14\u00de \u03b8p defines the principal directions where the only stresses are normal stresses. These stresses are called principal stresses see Fig. 2.16b and are found from the original stresses by substituting \u03b8 = \u03b8p from (2.14) in (2.13) r1;2 \u00bc rx \u00fe ry 2 ffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi rx ry 2 2 \u00fe s2xy r \u00f02:15\u00de We will learn later the relation between stress and strain determined from tension tests. These tests are all uniaxial in nature. The material properties are therefore given by tension tests with tension in one direction; the real structures are however subjected to two dimensional or even three dimensional cases of stresses"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0003898_oceans.1998.724374-Figure3-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0003898_oceans.1998.724374-Figure3-1.png",
+        "caption": "Figure 3 : The simulation test",
+        "texts": [
+            " The force sensor is modelled as an elastic transmission in the motor which command the articulation between the arm and the vehicle. Consequently, we only measure the torque produced by the arm on the platform. We know that the arm produces a torque and a resulting force on the platform. For the moment we consider only the action of the torque on the platform. It is the reason why we do not command the linear thrusters of the platform (X and Y) during the simulation test. IV - RESULTS The simulation test The test consists in a succession of simulation with the same goal (Cf. Figure 3), with and without force control loop and with the maximum joint velocities. The maximum joint velocities are varying between 10% and 200% of the maximum indicated in table 3. The linear thrusters of the vehicle are not controlled. We compute a performance criteria Ip (Cf. Equation l), which is the sum of the scared errors of the three last articulations of the robot (without considering the linear positioning of the platform) during each simulation. 5 N We obtain the results shown in figure 4. and the maximum speed allowed to stay in a laminar regime of IP 200 150 100 50 O I With force compensation loop ' / I Without force compensation loop I I /-; 50% -- 100% ,-"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0003772_j.1540-8159.1983.tb04395.x-Figure4-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0003772_j.1540-8159.1983.tb04395.x-Figure4-1.png",
+        "caption": "Figure 4. Struc(ure of a fiber optic pH sensor: (aj pH probe; fbj improved design. (Courtesy 0/ /. /. Peterson.'^)",
+        "texts": [
+            " Optical Sensors Many dyes exist which absorb or transmit light according to eitber ionic concentration or gas partial pressure in solution; others will alter the fluorescent intensity according to gas, ion, or molecular concentration. Optical fibers can be used to introduce ligbt to these dyes that are in communication with the body. The measurement of this absorption or fluorescence property is tben transmitted out through another fiber and can be analyzed by other electronic devices or instrumentation either attached to the subject or implanted. Fiber optic pH, pO2. pCO2, lactate and alcohol measuring devices have been reported. A sample structure is shown in Figure 4, Results of pH, pO2, pCO2, alcohol, and lactate have been reported in tbe literature,'^''* With the integrated optics now developing rapidly, the light emitting and light detecting devices can be fabricated on tbe same semiconductor chip together witb the signal analyzing circuitry. It is, therefore, conceivable that an implantable package can be developed that generates light at different wavelengths and trans- permeabte envetope seal piMtic optical fiber / seal indicator dye filling 1 mm Ctmstructum of pH {yrohe"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0000709_012002-Figure1-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0000709_012002-Figure1-1.png",
+        "caption": "Figure 1. Conceptual diagrams of precesional transmissions that operates efficiently in the multiplication regime.",
+        "texts": [
+            "he majority of precessional planetary transmissions diagrams developed previously operate efficiently in reducer\u2019s regime [1]. Depending on the structural diagram, precessional transmissions fall into two main types \u2013 K-H-V and 2K-H, from which a wide range of constructive solutions with wide kinematical and functional options that operate in multiplier regime. The kinematical diagram of the precessional transmission K-H-V (figure 1) comprises five basic elements: planet career H, satellite gear g, two central wheels b with the same number of teeth, controlling mechanism W and the body (frame). The roller rim of the satellite gear g gears internally with the sun wheels b, and their teeth generators cross in a point, so-called the centre of precession. The satellite gear g is mounted on the planet (wheel) career H, designed in the form of a sloped crank, which axis forms some angle with the central wheel axis . Content from this work may be used under the terms of the Creative Commons Attribution 3.0 licence. Any further distribution of this work must maintain attribution to the author(s) and the title of the work, journal citation and DOI. Published under licence by IOP Publishing Ltd 1 Revolving, the sloped crank H transmits sphero-spatial motion to the satellite wheel regarding the ball hinge installed in the centre of precession. For the transmission with the controlling mechanism designed as clutch coupling (figure1), the gear ratio (gear reduction rate) varies in the limits: g bg HV b z cos z i ; z     g bg HV b z cos z i , z cos      (1) reaching the extreme values of 4 times for each revolution of the crank H. If necessary, this shortcoming can be eliminated using as a controlling mechanism the constant Cardan joint (Hooke\u2019s joint), the ball synchronous couplings, etc. This kinematical diagram of the precessional transmission ensures a range of gear ratios i = 8...60, but in the multiplication regime it operates efficiently only for the range of gear ratios i = 8\u202625"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0001148_apmc.2016.7931338-Figure5-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0001148_apmc.2016.7931338-Figure5-1.png",
+        "caption": "Figure 5. Matching pyramid",
+        "texts": [
+            " Proceedings of the Asia-Pacific Microwave Conference 2016 978-1-5090-1592-4/16/$31.00 \u00a92016 IEEE For maintaining low profile of antenna system it\u2019s necessary to choose the minimum distance between mirrors equal to \u03bb/2. The criterion for such choice is maximum aperture efficiency, not maximum gain because this element is planned to be used as an array element. The resonator CAD model is shown on fig. 4. Figure 4. Fabry Perot cavity CAD model The open waveguide with matching dielectric pyramid is used as feeder (fig. 5). It\u2019s known that matching is the weak point of FBC so that additional auxiliary elements are required for this purpose. In our case the dielectric pyramid made from PTFE (\u03b5 = 2.1) was introduced to allow broadband matching (Fig. 6) At decreasing w1 from 13 mm to 12.5 mm the frequency band widens and gain decreases. This means that electrical distance between the reflectors remains the same in wider range of frequencies (the ascendant sector of reflection phase in fig. 2). However, with further decreasing of w1 the reflection coefficient of PRS significantly decreases (fig"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0000811_978-981-10-2374-3_5-Figure5.1-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0000811_978-981-10-2374-3_5-Figure5.1-1.png",
+        "caption": "Fig. 5.1 The model of the quadrotor",
+        "texts": [
+            " In Sect. 5.4, the proposed sliding mode control scheme is applied to a quadrotor for numerical simulation. Section5.5 presents some conclusions. In this section, we formulate the dynamicmodel fromNewton-Euler equations. First, we derive the nonlinear model of the quadrotor including rotational and translational dynamics. Then, we present some properties associated with the rotational dynamics to facilitate the stability analysis in the later section. The rotational dynamics of the quadrotor shown in Fig. 5.1 is: \u03c4B = IB \u03c9\u0307B + \u03c9B \u00d7 IB\u03c9B + 4\u2211 i=1 B RPi \u03c4Pi (5.1) where: \u03c4Pi = IPi \u03c9\u0307Pi + \u03c9Pi \u00d7 IPi\u03c9Pi \u2212 \u03c4di (5.2) \u03c4di = [ 0 0 \u2212 km\u03c92 i sign(\u03c9i ) ]T (5.3) We define: sign(x) = \u2223\u2223\u2223\u2223\u2223 1 if x > 0 \u22121 if x < 0 \u03c91 < 0, \u03c93 < 0 (clockwise along the positive ZPi axis) \u03c92 > 0, \u03c94 > 0 (counter-clockwise along the positive ZPi axis) The angular velocity in the propeller frame is \u03c9Pi = B RPi \u03c9B + [0 0 \u03c9i ]T (5.4) and the toque in body frame is: \u03c4B = 4\u2211 i=1 BOPi \u00d7B RPi TPi (5.5) and the force in the propeller frame is: TPi = [ 0 0 k f \u03c9 2 i ]T (5"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0000153_imece2015-50907-Figure1-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0000153_imece2015-50907-Figure1-1.png",
+        "caption": "FIGURE 1. THE MANIPULATOR WITH A MOBILE PLATFORM.",
+        "texts": [
+            " In this paper, we aim to propose a method to describe the workable location space of the manipulator\u2019s base which simplifies the problem of finding the proper stop position. In general, kinematics modeling of the industrial manipulator is based on D-H method presented by Denavit, J. and Hartenberg, R. S. [8] in 1955. The base frame is set on the fixed base of the manipulator, and the workspace of the end effector can be obtained by homogeneous transformations from the base to the end effector. However, it is not satisfactory to do so when an appropriate stop position is required for the painting manipulator on a mobile platform as shown in Fig. 1. Since the end effector should always maintain normal direction to the surface according to the requirements of painting process, it is difficult for the calculation of the workspace and to optimize the stop position with the original point set on the base. To make a better determination of the stop position, we propose a creative model to set the base frame at the center wrist point O and a new configuration is formulated. The scheme is mainly implemented on ABB IRB 5500 in this paper. For its special non-spherical wrist, the original point of the base frame is actually at the virtual wrist point which is the projection of the intersection point of axis 5-1 and 5-2 onto axis 4 when the joint angles of axis 4 and 5 are both 0, as shown in Fig"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0003291_bf00261843-Figure2-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0003291_bf00261843-Figure2-1.png",
+        "caption": "Fig. 2. Hemocoelic muscles in the wing of Clione limacina form a three-dimensional lattice. Type A cells (stippled) appear in transverse (tr), longitudinal (lg), and dorsoventral (dr) muscles and are believed to be involved in wing deflation. Type B cells appear only in transverse muscles and are believed to produce wing withdrawal. Inset is a dorsal-lateral view of Clione limacina. The dotted line represents the plane of section for the larger drawing",
+        "texts": [
+            " The dorsoventral muscles consist of numerous (about 1000) individual stellate cells with several processes that extend to the dorsal and ventral wing epithelium. They are dis tr ibuted throughout the wing hemocoel (Fig. l). The longitudinal muscles always lie dorsal and close to the transverse muscle bundles (Figs. 1, 3). The dorsoventral muscle cells are typically located at the intersections of transverse and longitudinal muscles (Fig. 1). Thus, these three muscle groups form a well-organized, latticelike structure in the wing hemocoel (Fig. 2). All the wing muscles act against a fluid-filled hemocoelic cavity. Dur ing wing retraction, the hemocoelic fluid is forced into the head and body cavities. There are two smooth muscle cell types in the retraction musculature (Figs. 2-4). The first (type A) is found in transverse, longitudinal, and dorsoventral muscle bundles. Type B is found only in transverse bundles. Al though always present in each transverse muscle bundle, type A cells never number more than three and are always located at the dorsal edge of the bundle, close to type A cells of the longitudinal bundles (Fig. 4). Thus, it is type A cells that form a three-dimensional lattice structure (Fig. 2). Type B cells, which are the major components of the transverse bundles, run in only one direction. Overall, type A cells have a smaller diameter than type B cells (Table 1). The following is an ul trastructural compar ison between the type A and type B muscle cells based on Nicaise and Amsel lem's (1983) classification system. F o r convenience, organelles in type A and type B cells are termed type A and type B organelles, respectively. Fig. 3. Electron micrograph of a transverse section of Clione wing, showing a cross section of longitudinal muscles (Ig), or type A cells (A), and longitudinal section of a transverse bundle (tr), which consists mainly of type B cells (B)"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0002874_881783-Figure5-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0002874_881783-Figure5-1.png",
+        "caption": "Fig 5. Response lag to steering torque (Running pattern II",
+        "texts": [
+            " While we are actually operating a motor cycle on a public road, we experience or en counter a variety of situations - we may be changing the lane, passing other traffic and coming back on the original lane J or we may be averting a person or motorcycle that has sud denly jumped into our way. In consideration of such situations these two courses were set up. On these courses the Japanese highway lane widths were used for a lane passing width and an obstacle averting width. Running Patterns (1) In Running Pattern I, the motorcycle ran in 881783 3 1\":.lr the ,-lep<?ndence of the response l::tg on the motorcycle speed, Fig. 5 (Running pattern I) and Fig. 6 (Running pattern II) show the relation of the lag in the test motorcycle response to the motorcycle speed. The fOllowing describes the equivalent characteristics which are represented by the similar dependency of all four test motorcycles on the motorcycle speed. (l) A lag in the lateral acceleration response -\u00b7of the frame in Running pattern I. (2) A lag in the yaw rate response in Running pattern II. (3) A lag in the roll rate response in Running pattern II. (4) A lag in the lateral acceleration response of the frame in Running pattern II"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0000477_j.ifacol.2015.09.726-Figure2-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0000477_j.ifacol.2015.09.726-Figure2-1.png",
+        "caption": "Fig. 2. 2-D Mesh cross-section of the alternator",
+        "texts": [
+            " The aim of this model is to provide the same global and local temporal waveforms as the real machine in the healthy operating condition. Thus, it can be used to simulate faulty cases (see Iamamura et al. (2010); Fiser et al. (2010)). As the studied electromagnetic phenomenon is invariant along the rotation axis z, a 2D extruded mesh representing a machine cross-section is used. Furthermore, in order to reach accurate results and avoid numerical oscillations, special care is taken to elaborate the mesh which finally contains 22444 prismatic elements and 22638 nodes. (see Fig. 2) All simulations are then carried out using the A \u2212 \u03d5 formulation. It is a magnetodynamic formulation based on the magnetic vector potential A and the electrical potential \u03d5 in order to account for the currents induced into the rotor bars. Thus, to limit time calculation, a low excitation current magnitude is used in order to avoid the effect of the magnetic material non-linearity. The developed model is used to simulate different no load healthy operating points at the rated speed and then to calculate different variables such as the electromotive forces (emf) at the tree phase stator windings and a magnetic flux of the airgap probe"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0001658_ie50389a020-Figure6-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0001658_ie50389a020-Figure6-1.png",
+        "caption": "FIGURE 6. APPARATUS FOR IN- on tires B and C, only DOOR DRUM TESTS",
+        "texts": [
+            " The results so far obtained fail to show any significant difference between the resistivity measured by cured-on electrodes and that measured with similar electrodes on surfaces adequately treated with Aquadag and with the use of moderate electrode pressures. During the minter i t was impossible to make road tests on tires. The tests were then continued on an indoor, 6-f00tI steel-drum, tire testing machine. The tire was mounted on an insulated wheel, and a brush contact with the rim was the collector for the electrostatic voltmeter. Figure 6 is a diagram of the apparatus. The drum was run a t a peripheral one tire out of each set of four used for the outdoor tests was available for the indoor tests. Consequently, the discrepancy between the indoor and the outdoor tests in these cases may have been due to an abnormal variation in the tires themselves. TABLE 111. Tire A B C D E F G W I J K L Potentials, % of COMPARISOS O F ROAD TESTS WITH I N D O O R DRUM TESTS control Of car (roadtests) 39 39 44 56 59 64 81 87 80 82 109 73 Ofrim(drumtests) 41 73 76 68 56 60 88 87 79 74 86 70 The problem of shock hazard to a passenger leaving or entering a motor vehicle depends upon the potential retained on the vehicle after it stops"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0000494_fst15-249-Figure7-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0000494_fst15-249-Figure7-1.png",
+        "caption": "Fig. 7. (a) NIF Cepheus 2015 physics package support structure and cone flange, (b) detail of physics package support (CAD model), and (c) detail of cone flange (CAD model).",
+        "texts": [
+            " 6b uses only the 3-D printed material, which makes assembly easier and makes the target more rigid. Complex structures can also be designed to avoid any interference with beams and diagnostic views. The target is fielded using a boresight with two overlapping features instead of a rotation fiber for alignment, which increases alignment accuracy at LLE. Additive manufacturing is also used to make substructures in targets that are impossible or expensive to machine. Simpler parts, such as spacers, are also printed to conserve machinist time. The structures shown in Fig. 7 include structures for accurate positioning of filter packs, diagnostic windows, and mirrors. The physics package holder was custom designed with multiple surfaces to hold each of the multiple layers in the assembly at different positions. IV. CONCLUSIONS Additive manufacturing has had a huge impact, improving the way targets are assembled while making them more robust and at the same time decreasing fabrication costs. FUSION SCIENCE AND TECHNOLOGY \u00b7 VOLUME 70 \u00b7 AUGUST/SEPTEMBER 2016 Many different applications have been developed from simple jigs and fixtures to complex structures built into the final target assembly"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0000451_j.ifacol.2015.09.326-Figure8-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0000451_j.ifacol.2015.09.326-Figure8-1.png",
+        "caption": "Fig. 8. Phase portrait of the system (14) for the case of k = 0.5, \u00b5 = 0.5.",
+        "texts": [
+            " The characteristic equation of system (14) linearized in the neighborhood of stationary solution v1 \u2217, w1 \u2217 has the form (\u03bb+ \u00b5)2 = 0; consequently, this point is a stable diacritical node. This stationary solution describes the vertical downward motion with a constant speed in the plane (x, y). The eigenvalues of system (14) linearized in the neighborhood of stationary solution v2 \u2217, w2 \u2217 satisfy the equation \u03bb2 = 2\u00b52(k + \u221a 1 + k2) 2k + \u221a 1 + k2 , therefore \u03bb2 > 0 and the point v2 \u2217, w2 \u2217 is saddle-type point. Solution v2 \u2217, w2 \u2217 corresponds to a sloping motion of the particle in the (x, y) plane with a constant speed. Phase portrait of the (14) is presented in Fig. 8. The point A is the diacritical node type and the point B is the saddle type. Phase portrait allows to analyze the qualitative features of the solution of the boundary-value problem. Initial value, belonging to the circle v2(0) + w2(0) = V0 2, should be chosen between separatrix directed to the point B and straight line w = \u2212kv, otherwise the second condition in (15) could not be achieved. Note, that under consideration are phase trajectories, belonging to \u2126 for any value of \u03c4 , where u = u+(v, w)",
+            " The eigenvalues of system (14) linearized in the neighborhood of stationary solution v2 \u2217, w2 \u2217 satisfy the equation \u03bb2 = 2\u00b52(k + \u221a 1 + k2) 2k + \u221a 1 + k2 , Fig. 6. Domains of constant sign for u+(v, w) when k \u2265 1\u221a 3 . Fig. 7. Domains of constant sign for u\u2212(v, w) when k \u2265 1\u221a 3 . therefore \u03bb2 > 0 and the point v2 \u2217, w2 \u2217 is saddle-type point. Solution v2 \u2217, w2 \u2217 corresponds to a sloping motion of the particle in the (x, y) plane with a constant speed. Phase portrait of the (14) is presented in Fig. 8. The point A is the diacritical node type and the point B is the saddle type. Phase portrait allows to analyze the qualitative features of the solution of the boundary-value problem. Initial value, belonging to the circle v2(0) + w2(0) = V0 2, should be chosen between separatrix directed to the point B and straight line w = \u2212kv, otherwise the second condition in (15) could not be achieved. Note, that under consideration are phase trajectories, belonging to \u2126 for any value of \u03c4 , where u = u+(v, w)"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0001149_1.5063125-Figure4-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0001149_1.5063125-Figure4-1.png",
+        "caption": "Figure 4: Nominal dimensions of the test geometries as defined for the comparison of SLM and LMD",
+        "texts": [
+            " In the beginning of the deposition tests the focus of the powder jet is adjusted to the surface of the substrate. The SLM tests are carried out on a commercially available SLM machine. The SLM machine is in a standard set-up. As beam source an Ytterbium fibre laser is used. The wavelength of the laser radiation is 1.070 \u03bcm. The beam diameter in the work plane is approximately 0.1 mm. The processing parameters are shown in Table 3. In order to evaluate both technologies, three test geometries are defined. These geometries are shown in Figure 4. An overview of the complete scope of experiments is given in table 4. During the SLM tests the samples are manufactured with a nominal wall thickness of 1.3 mm. For the LMD tests a beam diameter of 1.3 mm is chosen. Each test geometry is manufactured four times. In sum 24 test geometries are built. All test geometries are built on mild steel (CK 45) substrates as shown in Figure 5. To investigate the form tolerances a coordinate measuring machine is used. The form and position tolerances are investigated referring to the international norm DIN EN ISO 1101"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0002102_00032719808001846-Figure1-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0002102_00032719808001846-Figure1-1.png",
+        "caption": "FIG. 1. Schematic diagram of the screen-printed enzyme electrode body , 1, PVC matrix: 2. Carbon electrode; 3, Enzyme and mediator (Potassium fei-ricyanide) layer; 4, Ag strip; 5 , Contacts. The insulation layer is not given here in order to show the Ag strips.",
+        "texts": [
+            "5mm thick) was obtained from BASF, Japan. Silver ink (Electrodag 477 SS) and carbon ink D ow nl oa de d by [ Se lc uk U ni ve rs ite si ] at 1 9: 18 0 5 Fe br ua ry 2 01 5 (Electrodag 423 SS) were from Acheson Colloids company (Michigan, USA). Chronoamperometry was performed on an EG & G Potentiostat/Galvanostat Model 273A (Princeton Applied Research, USA) connected to a Gateway-2000 computer (South Dakota, USA) to record the voltammograms. A schematic representation of the electrochemical printed sensor is shown in Fig. 1. The sensor consisted of a PVC substrate ( OSmm thick) upon which was printed three layers. All layers were printed through polyester screens. Forty five electrodes were printed onto each piece of PVC substrate. Conducting tracks, consisting of ink containing silver was applied to the pieces and allowed to dry at 45\u00b0C for 24 hours. Carbon ink was applied to each track and also allowed to dry at least 48 hours. An insulation shroud of vinyl ink was then applied, leaving terminals and active surfaces exposed"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0003681_robot.1991.131809-Figure3-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0003681_robot.1991.131809-Figure3-1.png",
+        "caption": "Figure 3: Description of Foot Placement Algorithm",
+        "texts": [
+            "2 The Foot Placement Algorithm (FPA) In contrast to the 1-DOF model, the 2-DOF model requires a foot placement algorithm to actively control and balance its forward motion. The FPA does not enter explicitly in the equations of motion because the leg is assumed massless, but it determines where the robot foot should be located a t touchdown. This determines the initial conditions for the ensuing phases. A complete description of the FPA may be found in [1,5]. Given a desired forward velocity, i d e s i r e d , the algorithm places the foot a distance xeztension in front of the body to regulate the actual forward velocity (Figure 3). xe,tens~on is computed as, where i: is the actual forward velocity, nj. is a \"gain\" and Tjtance estimates the duration of the stance phase. This expression was developed for passive hopping without thrust and the factor of f in the first term may be replaced by another gain, &time, to approximate for the effects of thrust. Note that the algorithm is largely kinematic in nature. 3. The Poincari. Return Map for the 2 DOF Model We are interested in analyzing the global dynamic behavior of the 2-DOF hopping systems and the effect of system and control parameter variations on their behavior"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0002008_025-Figure1-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0002008_025-Figure1-1.png",
+        "caption": "Figure 1. Schematic representation of contact configuration.",
+        "texts": [
+            " Closure of the equation set is facilitated by the equation set h = h0 + x2 2R\u2032 + v(x) and R\u2032 = R1R2 R1 + R2 . This parabolic approximation is a good fit over comparatively narrow nip widths and for the purpose of the present study is very appropriate. The nominal film thickness for a rigid contact (h0) can assume a negative value and under this circumstance represents the mechanical engagement of the rollers. The distortion v(x) is derived from a solution of the elasticity equations in the elastomeric layer. The calculation domain and boundary conditions are shown in figure 1. Inlet and outlet pressures are zero and the latter also satisfies the zero gradient condition by setting negative pressures to zero as they occur in the solution back substitution. To perform the analysis also requires the specification of a contact width. The extent of the region upstream of the nip dictates whether the junction is either starved or flooded, where a large nip width is a reflection of a flooded condition. The extent of the region downstream from the nip is determined from the calculation and the satisfaction of the rupture condition. Thus the initial nip width was chosen to represent either starved or flooded conditions and was also estimated from the experimental investigation and a separate numerical study on a dry contact. This latter work was based on the progressive engagement of stationary rollers where the engaging surfaces were modelled using contact elements. In the elastomer domain shown in figure 1, circumferential straining of the elastomer was allowed and perfect bonding to the steel core was assumed. Material parameters for the rubber were chosen based on a separate investigation where the static elastic modulus was found to be 2 MPa and a Poisson\u2019s ratio of 0.45 was adopted, the rubber thickness was 8 mm and is typical for rubber covered rollers as used in the printing process. The solution method employed iteration between the hydrodynamic and structural models where it was necessary to apply relaxation to obtain convergence"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0002229_jsvi.1996.0017-Figure14-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0002229_jsvi.1996.0017-Figure14-1.png",
+        "caption": "Figure 14. Poincare\u0301 map for b=bc=0\u00b78842 with a=ac , d=0\u00b71, v=1, G=1.",
+        "texts": [
+            " This is where we expect to find the \u2018\u2018most\u2019\u2019 chaotic regime which exists for this system. The motion of the bead is a chaotic mixture of the \u2018\u2018bounded\u2019\u2019 oscillations around the two homoclinic fixed points, in a manner similar to that motion observed in the standard Duffing oscillator [11], and full loop oscillations similar to those found in the standard pendulum oscillator. The bead \u2018\u2018hops\u2019\u2019 chaotically between the two distinct chaotic states. (11b) 0\u00b76390QbQ0\u00b78842. The behaviour is as in (11a). (12) b=bc=0\u00b78842 (Figure 14). This is one of the special cases as predicted in section 3, where for the critical parameter values there co-exist two distinct, non-interactive, regions of transversal manifold intersections, namely, all heteroclinic manifolds, and all homoclinic manifolds ultimately each with its own attracting set. At this critical parameter value we see two non-interacting examples of chaotic motion: (i) true Duffing-type chaos with the homoclinic horseshoes centred around the homoclinic saddle point, (ii) true pendulum-type chaos with the heteroclinic horseshoes present around both heteroclinic saddle points"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0001619_004051755002000402-Figure5-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0001619_004051755002000402-Figure5-1.png",
+        "caption": "FIG. 5. Schematic drawing of the waterdrop during the first part of the impact.",
+        "texts": [
+            " In order to calculate this pressure the following assumptions are made: (1) the foundation is assumed to be a tensionless membrane, thus influencing the course of impact only by its mass; (2) the membrane is hydrophobic, and thus the vertical component of the surface tension is small and can be neglected; (3) the drop does not penetrate into the foundation. (The surface tension and the state of motion within the drop determine the shape of the drop. An exact calcula- tion of this shape is extremely complicated. This investigation is founded on the shape determined from the photographs. In this way the action of the surface tension is taken into account.) Figure 5 is a schematic drawing of the impact of a waterdrop, made on the basis of these assumptions; further discussion is given after the calculations. Acceleration of the Deflection of the Foundation.According to the above assumptions the membrane is deformed only below the surface of contact between the drop and the membrane. The accelera- tion of the deflection is given by equation (1), derived from the second law of motion: at EMORY UNIV on April 19, 2015trj.sagepub.comDownloaded from 218 in which so that we obtain Centey of Gravity of the Deformed Drop",
+            " In order to describe these relationships we consider the foundation as an isotropic, semi-infinite, elastic medium, affected by a static pressure uniformly distributed over a circular surface with radius p. By applying the theory of Boussinesq [2], we derive the following expression for the maximum deflection : where E = modulus of elasticity of the medium and v = Poisson\u2019s constant. The other symbols have the same geometrical meaning as before, but they are not the same functions of the time. According to equation (4) _ gt2/2 can be neglected, and hence From Figure 5: Assuming equation (15) to be valid also under dynamic conditions and substituting in equation (15) for p and w, we obtain If in equation (5a) we neglect g, 6/3 in comparison with 1, and 6 in comparison with 2, the following expression is obtained: and from equation (3), neglecting the mass of the drop, we obtain Furthermore, P = p~p2; hence, Substituting expression (16) for p, we get Equation (17) can be written where The differential equation (17) cannot be solved exactly, and it is necessary to use successive approximations"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0000577_978-3-319-33714-2_5-Figure1-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0000577_978-3-319-33714-2_5-Figure1-1.png",
+        "caption": "Fig. 1 PM with cylindrical joints",
+        "texts": [
+            " Utenov Al - Farabi Kazakh National University, Almaty, Kazakhstan e-mail: umu57@mail.ru \u00a9 CISM International Centre for Mechanical Sciences 2016 V. Parenti-Castelli and W. Schiehlen (eds.), ROMANSY 21 - Robot Design, Dynamics and Control, CISM International Centre for Mechanical Sciences 569, DOI 10.1007/978-3-319-33714-2_5 39 Following the above-mentioned trends in the development of PM, we proposed a novel structure of six-DOF three-limbed PM with cylindrical joints (PM 3CCC) (Baigunchekov et al. 2009), as shown in Fig. 1. This PM is formed by connection of a mobile platform 3 with a base 0 by three spatial dyads ABC, DEF and GHI of type CCC (C\u2014cylindrical joint). Each of spatial dyads of type CCC do not impose restrictions on motion of the mobile platform, and six-DOF of the mobile platform are remained. Each cylindrical joint has two-DOF: one rotation and one translation. In the considered PM the joints A, F and I are active joints, and the joints B, C, D, E, G and H are passive joints. Six variable parameters s7, \u03b87, s8, \u03b88, s9, \u03b89 of active joints A, F and I are the generalized coordinates"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0002923_978-94-011-3514-6-Figure10.2-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0002923_978-94-011-3514-6-Figure10.2-1.png",
+        "caption": "Fig. 10.2 Linear elastic system subject to n forces Qi",
+        "texts": [
+            " One confirms that the matrices [Ml, [Cl and [Kl are symmetrical and that they clearly have the values found from Newton's equations (8.4). On the other hand, one notes that the structure of [Cl is the same as that of [Kl This similarity is not general but never-the-Iess frequent. Conversely, the structure of intrinsically different. 10.3.2 Potential energy of a linear elastic system [Ml is One considers a linear elastic system, subject to n generalized forces Q\" ... , Qn . The point of application Ai of the force Qi is displaced by Ai (figure 10.2). I,et us designate by xi the component of AiAi along Qi' and by bi the component in the orthogonal plane. The deformed configuration of the system, from the initial confiquration (Qi=O) is defined by the set of displacements xi ' bi and one can write ( 10.12) - 179 - One knows that the reciprocal stiffnesses are equal (Maxwell-Betti theorem) The system being linear, the potentiel energy of deformation is equal to the half-sum of the products between the forces and the displacements in their directions (Clapeyron's equation) 1 n V = - 1: Qi Xi 2 i By replacing the Qi with their values (10"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0001034_iecon.2016.7793095-Figure1-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0001034_iecon.2016.7793095-Figure1-1.png",
+        "caption": "Fig. 1. Geometric sketch of the multi axes hybrid kinematic",
+        "texts": [
+            " [25], the resulting non-linear optimization task is often declared as follows. min u(k)\u2208Rn Jqual(x0, u(k)) (3) subject to umin(k) \u2264 u(k) \u2264 umax(k) (4) \u2206umin(k) \u2264 \u2206u(k) \u2264 \u2206umax(k) (5) y min (k) \u2264 y(k) \u2264 y max (k) (6) Here (4), (5) and (6) are possible formulations of the constraints. More complex formulations, e.g. of system state constraints, are also possible. In general a hybrid kinematic consists of a parallel kinematic with additional serial axes. The hybrid kinematic considered in this paper is shown in Fig. 1. It consists of a tripod with three linear moving axes q1, q2, q3 and an additional linear axis qshaft, which is mounted at the Tool Center Point (TCP) of the tripod. In the literature this resulting hybrid kinematic is also called a tricept [27] or a Hybrid Kinematic Machine (HKM). Enlarging the workspace, the additional linear axis creates the first redundancy, as movements in z-direction can be achieved by either moving all three parallel linear axes, the additional axis or all four together",
+            " Since the machines coordinate system is fixed to the HKM, the HKM\u2019s position is given in this coordinates. WTCP = WxTCPW yTCP W zTCP  = WxOB W yOB W zOB + MxTCPMyTCP MzTCP  (7) The HKM\u2019s position depends on the three parallel axes as well as on the additional serial component. The positions of the three parallel axes as a vector MKiPi in machine coordinates, where i = 1, 2, 3 denotes the respective axis, can be described as a linear combination of vectors between the respective spherical joint Pi and the universal joint Ki (8) (see Fig. 1). MKiPi =M KiBi +M BiOB +M OBOP +M OPPi (8) The vector elements in (8) can be calculated by inserting the geometric associations (9), (10), (11). For an uniquely solvable equation system the vector MOBOP between the moved platform of the HKM OP and the center of the joint for moving the HKM OB has to be definite. Of course \u03c8 or \u03b8 could also be defined to gain a unique solvable solution, but this would lead to a massive restriction of the TCP\u2019s movements. MxKiPi = \u2212 cos \u03b3i \u00b7R\u2212 sinM\u03b8 \u00b7 \u2223\u2223MOBOP \u2223\u2223+ r \u00b7 cos \u03b3i \u00b7 cosM\u03b8 (9) MyKiPi = \u2212 sin \u03b31 \u00b7R\u2212 cosM\u03b8 \u00b7 sinM\u03c8 \u00b7 \u2223\u2223MOBOP \u2223\u2223+ r \u00b7 (cos \u03b3i \u00b7 sinM\u03c8 \u00b7 sinM\u03b8 + cosM\u03c8 \u00b7 sin \u03b31) (10) MzKiPi = cosM\u03c8 \u00b7 cosM\u03b8 \u00b7 \u2223\u2223MOBOP \u2223\u2223\u2212 zK \u2212 zB\u2212 r \u00b7 (cos \u03b3i \u00b7 cosM\u03c8 \u00b7 sinM\u03b8 \u2212 sinM\u03c8 \u00b7 sin \u03b3i) (11) Parameter R defines the distance between the three universal joints Ki as a radius of a circle with its origin on the zaxis, r being the distance between the three spherical joints Pi (circle with origin in OP )"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0002566_88-gt-92-Figure2-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0002566_88-gt-92-Figure2-1.png",
+        "caption": "Fig. 2 : The meridional profile of the most sensitive impeller with a specific speed of n, =0.83",
+        "texts": [
+            " Therefore it was a logical approach to seek a correlation involving the most important consequence of tip clearance, overall efficiency deterioration and specific speed. The efficiency loss for each clearance was obtained by substracting from a hypothetical zero clearance on = na=o- nA, which was in turn found by extrapolation from the efficiency/clearance plot (1) . A sensitivity factor y was defined as the ratio of efficiency loss to clearance ratio, y = An /A . Fig. 1 and 2 show the meridional profile of the two impellers with opposite extreme sensitivities. The impeller with specific speed n s = 0.83, fig. 2, was the most sensitive of the tested series; especially at small clearances it was highly sensitive. On the contrary the impeller with specific speed n s = 0.57 showed the least sensitivity; especially at small clearances it was insensitive. This sensitivity trend is supported by Senoo et al. (18) , who investigated two centrifugal blower impellers of specific speed n = 0.43 and 0.58. The impeller with n s 0.58 was also relatively insensitive. From the present investigation, it is not clear why a machine of smaller specific speed is less sensitive, one would expect the opposite, i"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0000441_iccsnt.2015.7491028-Figure4-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0000441_iccsnt.2015.7491028-Figure4-1.png",
+        "caption": "Fig. 4. Coverage Diagram of Base Station 43792",
+        "texts": [
+            " The longitude and latitude of the base station are 112.99061 and 28.22061, and its real location is Office Building. The antenna height is 35m, with direction angles of 30, 120, 240, 30, 120, 240, 30, 120 and 240, respectively. The total downtilt is 14 (the mechanical downtilt is 4, the electrical downtilt is 10 and the pre-set downtilt is 0), the transmitting power is 13 and the antenna gain is 18 (horizontal half power: 65; vertical half power: 8). First, calculate the coverage of the base station according to the basic information. In Figure 4, according to the principle of geometry )/tan( RHac  L H  ) 2 tan(   S H  ) 2 tan(   where is downtilt,  is the angle of vertical half power(65 degree), H is height and R is coverage distance. It can be calculated that the proximal distance S is 107.7m, remote distance L is 198.5m and center distance R is 140.4m. In order to analyze the correlation between signal strength and distance to base station better, we approached base station 4380 gradually from 43972 in this experiment and checked the signal strength through 3G module at an interval of 2000ms"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0001070_iccas.2016.7832415-Figure1-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0001070_iccas.2016.7832415-Figure1-1.png",
+        "caption": "Fig. 1. Two-link rigid robot manipulator.",
+        "texts": [
+            " Since the output derivatives may not be practical to be measured in all situations, they need to be estimated from an available output. Normally, this is not a problem for many applications, since advanced observer or estimator solutions can be easily found in the literature. Consider the physical model of a two-link rigid robot manipulator, with each joint equipped with a motor for providing input torque, an encoder for measuring joint positon, and a tachometer for measuring joint velocity, as shown in Fig.1. Using the Euler-Lagrangian equations well known in classical dynamics, we can obtain the dynamic equations of the robot as [25] (24) where and are the two joint angles, and are the joint torques, and are the lumped uncertainties acting on the system, and ; ; ; ; ; ; ; ; ; where and are the link masses in 1 kg, and are the link lengths in 1 m, and is the gravitational acceleration. Then, the above matrix elements represents physical properties of the system, i.e. Coriolis, centripetal and gravitational effects"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0003865_ijeee.39.1.6-Figure1-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0003865_ijeee.39.1.6-Figure1-1.png",
+        "caption": "Fig. 1 Equivalent circuit for a d.c. motor-generator group with a tachometer.",
+        "texts": [
+            " The latter is generally the goal of a first course in control while the former, although relying on a strong background, can only be achieved with the help of a good control laboratory. It is well accepted that a good control laboratory must not only illustrate the concepts introduced in the theoretical course but have to be realistic as well.2 Motivated by these facts, a control laboratory3 for the Electrical Engineering Course of the Federal University of Rio de Janeiro (UFRJ) has been developed and successfully used. The plant consists of a d.c. motor-generator group represented in Fig. 1, where va(t) is the input voltage, vt(t) is the voltage at the tachometer terminals and ig(t) denotes the current supplied by the generator when a load is connected at UNIVERSITE LAVAL on July 13, 2015ije.sagepub.comDownloaded from to its terminals. The problem here is to control the shaft velocity in the presence of plant uncertainties and external disturbance signals. A d.c. motor-generator group has been chosen for the plant and the shaft velocity as the control variable because they allow the following concepts to be illustrated: 1 Modeling; 2 Linearity, identification and measurement noise; 3 The effects of parameter identification errors and external disturbance in an open-loop control system; 4 The benefits of feedback; 5 The need for dynamic compensation",
+            " Finally, in the fifth section, a very simple electronic network will be introduced for the physical implementation of the controller derived in the previous section. Since this laboratory is meant for a first course in control, analog devices are deployed. However, it is important to note that the same controller could be easily implemented by using a digital controller if this laboratory takes place after a course in discrete-time control systems. A mathematical model for the d.c. motor-generator group depicted in Fig. 1 can be obtained simply by considering the equivalent circuit of the armature-controlled d.c. International Journal of Electrical Engineering Education 39/1 at UNIVERSITE LAVAL on July 13, 2015ije.sagepub.comDownloaded from motor of Fig. 2, where va(t) and ia(t) denote, respectively, the input voltage and the current in the armature circuit, w(t) is the shaft velocity and J and f are, respectively, the load inertia and the bearing friction. It is not difficult to show that:5 (1) where te = La/Ra, tm = J/f, Km is the torque constant, Ke is the counter electro-motor force and td(t) denotes the disturbance torque which, in this case, appears when a load is connected to the generator terminals. However, since La/Ra << 1 then for low and intermediate frequencies tes + 1 \u00aa 1 and, therefore, a simpler model for this system is given by: (2) where In order to obtain a complete model for the system, it remains to take into account the effects of the disturbance current (ig(t)) and the tachometer as well. The former can be accounted for, with the help of Fig. 1, by remembering that td(t) = ig(t) while the latter is usually modeled as a constant gain system, i.e. vt(t) = Ktw(t). Finally, defining Kg = Kd, the transfer function which relates va(t) and ig(t) to vt(t) can be written as follows: (3) leading to the block diagram of Fig. 3. V s K K s V s K K s I st a t a g t g( ) = + ( ) - + ( ) t t1 1 , K K K R R f K K JR R f K Kd a a e m a a g m= +( ) = +( )and t . K K R f K Ka m a e m= +( ), W s K s V s K s T sa a d d( ) = + ( ) - + ( ) t t1 1 W s K R f s s K K R f V s s f s s K K R f T s m a e m e m a a e e m e m a d ( ) = ( ) +( ) +( ) + ( ) ( ) - +( ) +( ) +( ) + ( ) ( ) t t t t t 1 1 1 1 1 International Journal of Electrical Engineering Education 39/1 at UNIVERSITE LAVAL on July 13, 2015ije"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0000895_carpi.2016.7745620-Figure3-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0000895_carpi.2016.7745620-Figure3-1.png",
+        "caption": "Fig. 3 Calculate the PTZ offset angle by tangent formula",
+        "texts": [
+            " Sort each pair of feature points\u2019 position in the two images and then the pixel error offset of real-time image and template image can be calculated by: 1 ( ) ( ( ))temp cap n H offset pix mean C C V          (1) where, ( )offset pix is the result of pixel error offset, horizontal (H) and vertical (V) direction. tempC , the position of feature point in template image; capC , the position of the other one feature point in real-time captured image. The calculation method of converting image pixel deviation to PTZ angle offset is as following described. (1) PTZ Angle Offset Calculation Based on tangent proportion as distance and focal length, calculate PTZ angle offset with image pixel deviation as figure 3 shown. In figure 3, the \u201coffset\u201d means real distance what the image pixel deviation in the camera projection of the CCD or IR device. It can be calculated by: ( ) H H offset offse pix resolution V V              (2) Where, offset is the distance of image pixel in the imaging device, horizontal (H) and vertical (V) direction. resolution is one pixels width in the camera imaging device. In figure 3, Due to the target position T point relative to the camera's distance is far greater than the focal length of the distance. T point during the acquisition of the images in the template image and the corresponding T\u2019 point can be approximate treated as translation. t and t\u2019 points respectively present the position of T and T\u2019 points in the projection of camera imaging device. The image pixel deviation between template image and real-time image can be presented by the displacement of t\u2019 point relative to t point"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0000897_978-3-319-50472-8_2-Figure3-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0000897_978-3-319-50472-8_2-Figure3-1.png",
+        "caption": "Fig. 3. Case 3.2.",
+        "texts": [
+            " the number of nodes occupied by other robots on each side. The robot may do so simply by counting the number of 1s and 0s in the strings. \u2013 Step 3: - This is the most important step since in this step the robot makes a decision on its movement. The cases are listed as follows: - \u2022 Case 3.1: - If the occupancy rate is nil on both sides then the robot does not make any movement (Fig. 2). \u2022 Case 3.2: - If the occupancy rate is equal on both sides then the robot makes one hop movement to the side with closer neighboring occupied nodes (Fig. 3(a)) or any of the sides if there is a tie (Fig. 3(b)). \u2022 Case 3.3: - If the occupancy rate is more on the counter-clockwise direction but the clockwise string is nil then the robot does not make any movement (Fig. 4(a)) else it makes one hop movement to the counterclockwise direction (Fig. 4(b)). \u2022 Case 3.4: - If the occupancy rate is more on the clockwise direction then the robot makes one hop movement to the clockwise direction (Fig. 5). 2-Node Problem: The main concern which may act as a thorn on the path of gathering is the 2-node problem, where two nodes in the ring are occupied by the robots"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0000195_iccas.2014.6987737-Figure1-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0000195_iccas.2014.6987737-Figure1-1.png",
+        "caption": "Fig. 1 Mechanical configuration of a quadrocopter with body fixed and inertial frame.",
+        "texts": [
+            " In presence of a col lision risk the collision avoidance layer adapts the refer ence velocity based on the repulsive force approach, thus, an evasive manoeuvre is performed by the UAY. More over, it has to be investigated in simulations if and how the presence of obstacles affects the convergence of the PEG solution in opposition to the solution in an obstacle free environment facing static and moving obstacles. 3. SYSTEM DESCRIPTION 3.1 Dynamical Model The detailed derivation of the dynamic model used in this work can be found in [10]. For modeling the quad rotor dynamics the mechanical configuration depicted in figure 1 was assumed. The body fixed frame and the in ertial frame are denoted by eB and eI, respectively. The UAV is defined as a point mass. To derive the equations of motions, the following notations are necessary. pI = (x, y, z)T is the position vector of the quad-rotors' center of gravity in the inertial frame, pB = (x B, Y B, Z B) T is the position vector of the quad-rotors' center of gravity in the body fixed frame, v = (u, v, w ) T are the linear velocities in body fixed frame, w = (p, q, r ) T are the an gular rates for roll, pitch, and yaw in body fixed frame, and 8 = (cp, e, 1/)) T is the vector of Euler angles"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0001000_1.4035601-Figure3-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0001000_1.4035601-Figure3-1.png",
+        "caption": "Fig. 3 Inclined grooves",
+        "texts": [
+            " There are strengths and weaknesses for using either of these orientations. However, the focus of these investigations is on a radially mounted seal. Advantages for this orientation include being more applicable to turbine split line designs, having to cope with less movement perpendicular to the flow direction, and being easier to scale at various diameters. There are many possible groove shapes, several of which originate from bearing design. These include wedges (flat and tilted) (Fig. 1), circumferential pockets and Rayleigh steps (Fig. 2), inclined grooves (Fig. 3), and herringbone grooves (Fig. 4). Wedges are reported by Dhagat et al. [32] and the tilted variety is recorded by Galimutti et al. [33]. Pockets are also shown by Dhagat et al. [32] and Rayleigh steps are described by Cheng and Wilcock [34]. Herringbone grooves are assessed by several authors although there is no general agreement on whether the center land region should be included or not. The center land is omitted in Dhagat et al. [32] and Liu et al. [35] but included in Cheng and Wilcock [34] and Proctor and Delgado [29]"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0000050_iemdc.2015.7409124-Figure6-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0000050_iemdc.2015.7409124-Figure6-1.png",
+        "caption": "Fig. 6. Magnetic field distribution in BLDC motor as obtained from the 2 nominal load; (b) radial component of no-load.",
+        "texts": [
+            " until a threshold ld) at low control manufacturers to performed on a onstantly monitor lements). If the unit increases the e rotor speed and in a cooling fan oothly. Bearings t high speeds, and ancy of a cooling ns. Three types of ) sleeve bearings, rings. d two-phase PM The disassembled ing and fan blades case fan motor 1.10 W 1500 rpm 12 V DC 0.091 A 4 29.4 mm 5 mm 35 mm 31.6 mm 9.2 mm isotropic barium ity Br = 0.4 T and PWM solid state from Hall sensors The cross section of the distribution is shown in Fig. 6 the magnetic flux density dis plotted in Fig. 6b. The speed is controlled by t that the PWM 1 signal is 0% rotating at the minimum speed, The motor can start up at 3 range of 3.5 to 14 V. The pea 1 PWM is a modulation of the dut switch between two fixed values, such the signal never changes in the proces rhythm of \"on time\" and \"off time\". F a fan \u2013 such a CPU or casing fan \u2013 r maximum speed r for PC fan: (a) inner stator with the investigated two-phase fan PM D FEM analysis: (a) flux lines at the air gap magnetic flux density at motor and magnetic flux a"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0001256_jahs.60.042007-Figure1-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0001256_jahs.60.042007-Figure1-1.png",
+        "caption": "Fig. 1. Gearbox with a face-gear drive: (a) schematic diagram and (b) structural dynamics model.",
+        "texts": [],
+        "surrounding_texts": [
+            "A(t) state-space system matrix ag tooth addendum bg tooth dedendum c1, c2 damping coefficients of input/output shafts cm, km average mesh damping/stiffness cr contact ratio cp circular pitch D Rayleigh dissipation function d tooth face width E gear elastic modulus gi(t) deformation of the ith mesh stiffness h thickness of face-gear body plate Jm, J1, J2, JL mass moment of inertia of input side, pinion, face gear, and load k1, k2 torsional stiffness of input/output shafts kt tooth stiffness L1, L2 inner/outer limiting value for face-gear tooth Lsp boundary curve of fillet surface M, C(t), K(t) system mass, damping, stiffness matrices M2s rotation matrix between shaper and face-gear coordinates md gear module m12 gear ratio Ns, N1, N2 number of shaper, pinion, and face-gear teeth Nv1, Nv2 number of teeth for formative spur gears ns unit normal to shaper tooth surface Pd diametral pitch \u2217Corresponding author: e-mail: mpeng1@vols.utk.edu. Presented at the American Helicopter Society 68th Annual Forum, Fort Worth, TX, May 1\u20133, 2012. Manuscript received August 2014; accepted December 2014. q(t) generalized coordinates vector R1, R2 inner/outer limiting radii rs, r2 position vectors in shaper/face-gear coordinates rb1 pinion base circle radius rds shaper dedendum radius rk(t), \u03b8k(t) radial/circumferential positions of contact point rvai, rvbi addendum/base circle radii of formative spur gears T total system kinetic energy T0 mesh period t time (us, \u03b8s) Gaussian coordinates of the involute shaper surface V(s2) s relative velocity of shaper to face gear V total system strain energy z\u2217 s limiting value along the z axis of shaper surface \u03b1PA pressure angle \u03b1vai addendum pressure angle of formative spur gears \u03b3 shaft angle of face-gear drive \u03b3s apex semiangle of shaper pitch cone \u03b31, \u03b32 apex semiangle of pinion/face-gear pitch cone \u03b8r initial angle \u03bbi ith Floquet multiplier \u03bc mesh stiffness correction factor \u03be small real number \u03c1 material density (T0,0) Floquet transition matrix \u03c6s generalized parameter of motion \u03d5m, \u03d51, \u03d52, \u03d5L total rotation angles of input side, pinion, face gear, and load \u03c81, \u03c82, \u03c8L elastic deviation angles of pinion, face gear, and load m constant input speed pn natural speed 1, 2 pinion/face-gear rotation speed \u03c9 mesh frequency DOI: 10.4050/JAHS.60.042007 C\u00a9 2015 The American Helicopter Society042007-1 Introduction Recently, face-gear drive technology has been successfully applied to helicopter main transmissions. However, torsional vibration and parametric instability due to the unique face-gear meshing kinematics are still in need of thorough investigation to improve dynamic performance and reduce noise. Over the past two decades, a number of scholars have made significant contributions to the development of face-gear drives, including: face-gear generation and meshing simulation in Ref. 1, finite element analyses for tooth contact and bending stress optimizations in Refs. 2\u20134, experimental investigations on failure modes and durability tests in Refs. 5\u20139, performance evaluations of tooth profile modifications in Refs. 10\u201312, and dynamic stability studies on transverse vibrations in Ref. 13. Moreover, the instantaneous load sharing on all teeth in mesh was explored for a face gear theoretically and experimentally in Ref. 14. Two simple analytical formulae were presented to predict the radius limits of face-gear tooth based on statistical methods by Ref. 15. Finally, some prototype face-gear drives for helicopter transmission were designed and validated in Refs. 16 and 17. The research on torsional stability of face-gear drives is scarce, even though many results exist for spur (Refs. 18\u201320) and bevel gear systems (Refs. 21 and 22). High-frequency tooth mesh loads on spur gear teeth with pitch errors were analyzed by employing an equivalent springmass dynamics model in Refs. 23 and 24. The concept of variable mesh stiffness was proposed for analyzing dynamic loads by considering tooth deformations, profile modifications, transmission errors, and so on, in Ref. 25. The effects of shaft and bearing compliance were also included in a cylindrical involute gear model to estimate vibrations resulting from tooth profile errors in Ref. 26. The individual and combined torsional mesh stiffnesses were investigated based on finite element methods in Ref. 27. It has been shown that the fluctuation of tooth mesh stiffness due to a nonunity contact-ratio can excite dynamic instability, which was observed and modeled for spur gears in Ref. 19. Another significant source of instability in face-gear drives arises from the unique kinematic variation in the position of the out-of-plane mesh loads normal to the face-gear body. According to previous research in Refs. 1, 12, and 13, the meshing tooth pairs of a face-gear/pinion drive act in so-called line contact mode when the pinion is an identical copy of the generating shaper. The contact line represents the locus of the effective mesh loads at any given instant. During each meshing cycle, the contact line progresses across the tooth surface, which introduces a periodically time-varying stiffness term into the equations of motion leading to a potential source of parametric instability in face-gear drives. The schematic and corresponding structural dynamics model of a typical face-gear drive are shown in Figs. 1(a) and 1(b), respectively. The objectives of this paper are to (a) establish a structural dynamics model for torsional vibration of a single-stage face-gear drive and (b) analyze the system stability behavior. A parametric study involving gear system rotational inertias and shaft torsional stiffness values is conducted, and recommendations for reduced vibration designs are made. Since the system is periodically time varying, Floquet theory is employed to solve a Mathieu\u2013Hill type equation to evaluate the system stability. Also, a perturbation analysis based on Hsu\u2019s method (Ref. 28) is performed to obtain approximate closed-form expressions for the stability boundaries for design purposes. Face-Gear Geometry and Meshing Kinematics In this paper, the face gear and input pinion are cut for line contact which occurs when the pinion gear and the shaper have identical profiles. The cross section of a sample face-gear drive is shown in Fig. 2. According to the mathematical model developed by Litvin et al. in Refs. 1 and 29, contact lines, which generate the face-gear tooth surface, are derived from the so-called equation of meshing\u23a7\u23aa\u23aa\u23a8 \u23aa\u23aa\u23a9 f (us, \u03b8s, \u03c6s) = ns \u2022 V(s2) s = 0 V(s2) s = [ v(s2) xs v(s2) ys v(s2) zs ] T r2(us, \u03b8s, \u03c6s) = M2s(\u03c6s)rs(us, \u03b8s) (1) where ns is the unit normal vector to the shaper tooth surface and V(s2) s is the relative sliding velocity between the shaper and face gear at a point of contact in shaper coordinates. Furthermore, M2s is the rotation matrix, 042007-2 which transforms shaper coordinates into face-gear coordinates, rs and r2 are position vectors in the shaper and face-gear coordinates, respectively, (us ,\u03b8s) are the Gaussian coordinates of the involute shaper surface, and \u03c6s is the generalized parameter of motion. To avoid undercutting, the position vector for the inner limiting end (section A in Fig. 2) of facegear tooth is solved via the following equations (Ref. 29):\u23a7\u23aa\u23a8 \u23aa\u23a9 rs = rs(us, \u03b8s) = [ xs ys zs ]T f (us, \u03b8s, \u03c6s) = 0 2 1 + 2 2 + 2 3 = 0 (2) with 1 = \u2223\u2223\u2223\u2223\u2223\u2223\u2223\u2223\u2223\u2223\u2223\u2223\u2223\u2223 \u2202xs \u2202us \u2202xs \u2202\u03b8s v(s2) xs \u2202ys \u2202us \u2202ys \u2202\u03b8s v(s2) ys \u2202f \u2202us \u2202f \u2202\u03b8s \u2202f \u2202\u03c6s d\u03c6s dt \u2223\u2223\u2223\u2223\u2223\u2223\u2223\u2223\u2223\u2223\u2223\u2223\u2223\u2223 = 0 (3a) 2 = \u2223\u2223\u2223\u2223\u2223\u2223\u2223\u2223\u2223\u2223\u2223\u2223\u2223\u2223 \u2202xs \u2202us \u2202xs \u2202\u03b8s v(s2) xs \u2202zs \u2202us \u2202zs \u2202\u03b8s v(s2) zs \u2202f \u2202us \u2202f \u2202\u03b8s \u2202f \u2202\u03c6s d\u03c6s dt \u2223\u2223\u2223\u2223\u2223\u2223\u2223\u2223\u2223\u2223\u2223\u2223\u2223\u2223 = 0 (3b) 3 = \u2223\u2223\u2223\u2223\u2223\u2223\u2223\u2223\u2223\u2223\u2223\u2223\u2223\u2223 \u2202ys \u2202us \u2202ys \u2202\u03b8s v(s2) ys \u2202zs \u2202us \u2202zs \u2202\u03b8s v(s2) zs \u2202f \u2202us \u2202f \u2202\u03b8s \u2202f \u2202\u03c6s d\u03c6s dt \u2223\u2223\u2223\u2223\u2223\u2223\u2223\u2223\u2223\u2223\u2223\u2223\u2223\u2223 = 0 (3c) Furthermore, the inner limiting value of the face-gear tooth, L1, is obtained as (Ref. 29) L1 = z\u2217 s \u2212 rds tan \u03b3 and \u03b3 \u2264 90\u25e6 (4) Here z\u2217 s represents the limiting value along the z axis of the shaper surface (zs axis in Fig. 2), solved by the nonundercutting condition described in Eqs. (2) and (3), rds is the shaper dedendum radius, and \u03b3 is the intersecting shaft angle of the face-gear drive. The outer limiting end (section B in Fig. 2) is determined by the condition of the face-gear tooth top land becoming pointed, and the outer limiting value, L2, is given by (Ref. 29) L2 = Ns 2Pd ( 1 tan \u03b3s \u2212 1 tan \u03b3 ) + ag tan \u03b3 + Ns 2Pd tan \u03b3s ( cos \u03b1PA \u2212 cos \u03b1 cos \u03b1 ) (5) where Ns is number of shaper teeth, Pd is diametral pitch, \u03b3s is apex semiangle of shaper pitch cone, ag is tooth addendum, \u03b1PA is pressure angle, and \u03b1 is solved by (Ref. 29) \u03b1 \u2212 (Ns \u2212 2) sin \u03b1 Ns cos \u03b1PA = \u03c0 2Ns \u2212 (tan \u03b1PA \u2212 \u03b1PA) (6) The corresponding inner limiting radius, R1, and outer limiting radius, R2, for the face-gear tooth are scaled in the plane of the face-gear body disk, respectively, as R1 = L1 sin \u03b3, R2 = L2 sin \u03b3 (7)"
+        ]
+    },
+    {
+        "image_filename": "designv11_60_0002154_jsvi.1996.0870-Figure2-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0002154_jsvi.1996.0870-Figure2-1.png",
+        "caption": "Figure 2. Squeeze film and deflections of tube/support.",
+        "texts": [
+            " When the tube vibrates at large amplitude or large initial eccentricity in its support hole, both phenomena may occur physically. Therefore both squeeze and solid contact forces have to be calculated simultaneously. When the normal velocity of the tube vn is positive, a squeeze film occurs between the surfaces of the tube and the support. As the tube approaches very near the support, the normal squeeze force becomes quite large and tends to deform the tube and the support walls. The deflections increase the minimum film thickness as shown conceptually in Figure 2. At the same time, the deflections also influence the actual velocity and acceleration of the changes in the minimum squeeze film thickness. By not considering surface deflections, the minimum film thickness is h=Cr \u2212 e, where Cr is the radial clearance between tube and support, and e is the instantaneous radial displacement of the tube centre. With the elastic tube and support deflections, the minimum film thickness becomes hx =Cr \u2212 e+Xt +Xs =Cr \u2212 e+Xw, (5) where Xt and Xs are the deflections of tube and support, respectively; Xw =Xt +Xs =combined deflection"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0002904_s1474-6670(17)52428-3-Figure1-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0002904_s1474-6670(17)52428-3-Figure1-1.png",
+        "caption": "Fig. 1. Definition of Links Frames.",
+        "texts": [],
+        "surrounding_texts": [
+            "Copyright \u00a9 IFAC Identification and System Parameter Estimation, Budapest. Hungary 1991\nIDENTIFICA TION AND ADAPTIVE CONTROL IN MANUFACTURING\nAND ROBOTICS\nCALCULATION OF THE IDENTIFIABLE PARAMETERS FOR ROBOTS CALIBRATION\nW. Khalil and M. Gautier\nLaboratoire d'Automatiquc de Names URA CNRS 823, ENS,M. 1 Rue de la Noe,\n44072 Names Cedex 03, France\n~. This paper presents a numerical method to determine the identifiable parameters used in geometric robot calibration. The determination of these parameters facilitates the identification and calibration process, The calculation is based on QR decomposition, it proceeds in two steps: - at first the number of identifiable parameters is determined, - then a set of base parameters is determined by eliminating some standard parameters, The algorithm is easy to implement using available software package. An application for a 6 degree of freedom robot is given.\nKeywords. Calibration, geometric parameters, identifiability, identification, model reduction, robots.\nINTRODUCTION\nThe absolute errors of robots are both non geometric and geometric (Whitney 86, Roth 87). The former may be due to friction, compliance, gear transmission and backlash, The later may result from imprecise manufacturing of the robot links and joints, or from the deviation of the encoder offsets. It may be also due to the poor estimation of the parameters defining the location (position and orientation) of the robot with respect to the fixed reference frame, and of the poor estimation of the parameters defining the end effector with respect to the terminal link, This paper proposes a numerical method to determine the identifiable geometric parameters needed in robots calibration. A general identification model is firstly presented, then an algorithm to determine the identifiable geometric parameters based on QR decomposition is given,\nDEFINITION OF THE GEOMETRIC PARAMETERS\nWe define: frame -I : a fixed reference frame, frame 0 : a frame fixed with respect to the base of the robot, frame n : a frame fixed with respect to the terminal link of the robot, frame n+ I: a frame fixed with the tool or the end effector.\nThe end effector frame location, position and orientation, can be calculated with respect to the reference frame by the direct geometric model given as :\n(I)\nwhere: -ITo is the transformation matrix defining the base of the robot with respect to the fixed frame, <Tr n is the transformation matrix of the robot, it defines a frame\nfixed with the terminal link with respect to the robot base frame. nT n+ I is the matrix defining the end effector frame with respect to frame n.\nOwing to the poor knowledge of the values of the geometric parameters, defining the previous matrices, there will be a difference between the real location of the end effector frame and its location as calculated using the direct geometric model. Three sets of geometric parameters are thus taken into account: the robot parameters, the parameters defining the base of the robot with respect to a fixed reference frame, and the parameters defining the end effector with respect to the terminal link.\nDefinition of the Robot Geometric Parameters\nWe consider open loop robots consisting of n joints and n+1 links. Link 0 is the base and link n is the terminal link, The coordinate frame j is assigned fixed with respect to link j. The definition of the link frames will be carried out by the modified Denavit and Hanenberg notation (Khalil 86, Dombre 88). The Zj axis is along the axis of joint j , the Xj axis is along the common perpendicular of Zj and Zj+I ' Frame j is defined with respect to frame j-I by the matrix j-I T j which is function of the four\nparameters (aj, dj, Sj, rj), Fig. I, such that:\nj-I Tj = Rot(x, aj}Trans(x, dj}Rot (z, Sj) Trans(z, rj}\nwhere:\nj-IA' J\no\no j-lpj 1\nCajCSj -SUj\nSajCSj Caj\no 0\n(2)\nj-I Aj is a 3x3 matrix defining the orientation of frame j with\nrespect to frame j-I, j-I Pj defines the position of the origin of frame j with respect to\nframe j-I.",
+            "The joint variable j will be given as :\nqj = crj 6j + (Jjfj (3)\nwhere: (Jj = 0 for j rotational, (Jj = I for j translational, and\ncrj = (1- (Jj)'\nWe define also :\nqj = (Jj 6j + crjfj (4)\nThe robot base frame, frame 0, is defined such that Zo is along\nzl,and xOis along XI when ql = 0, this leads to get uI= 0,\nd I = 0 and Cl I = O. Frame n will be defined such that Xn is\nalong Xn-I when qn = 0, thus Cln = O.\nDefinition of the End Effector Parameters\nIn general the end effector frame, frame n+ I, is defined arbitrarily by the user. Therefore its location with respect to frame n will need 6 parameters. An arbitrary frame j can be defined with respect to a frame j-I by the 6 parameters\n(Yj,bj,uj,dj,6jofj) shown in Fig. 2 (Khalil 86, Stone 86).\nThus we can define the end effector frame with respect to frame n as :\nConcatenating this matrix with the matrix n-I T n we get:\nn-I Tn+1 = n-ITn nTn+1 n-IT n+ I = Rot(x, un)Trans(x, dn)Rot (z, 6n)\nTrans(z,rn)Rot(z,YE)Trans(z,bE)Rot(X,UE)\nTrans(x,dE)Rot(z,6E)Trans(z,fE) which gives :\nn-I T n+1 = Rot(x,un)Trans(x,dn)Rot(z,6n')Trans(z,rn ')\nRot(x,<Xn+ I )Trans(x,dn+ I )Rot(z,6n+ I )Trans(z,r n+ I) (6)\nThus the parameters YE, and bE cannot be identified separately,\ntheir errors will be calibrated through the parameters 6n and rn respectively. Frame n+ I will be represented by 4 independent parameters similar to those defining the link frames such that\n(<Xn+ I = UE, dn+ I = dE, 6n+ I = 6E, rn+ I = fE). Remark: From this study we conclude that the concatenation of successive frames defined by six parameters, as given in Eq.5 , can be reduced to the representation of four parameters . So Denavit and Hartenberg notation is not an under determined parameterization as it may seem to be (Ziegert 88).\nDefinition of the Robot Base Frame\nThe world fixed frame, frame - I, can be defined arbitrarily by the user. So frame 0 will be defined with respect to frame -I by six parameters. The matrix -I T 0 can be defined by the\nparameters : (Yz, bz, u z, dz, 6z, rz)\u00b7 Since uI = 0, d l = 0, the transfomlation -IT I can be given as:\n-IT I = -IToD-r1 -I T I = Rot(z,YzlTrans(z,bz)Rot(x,uz)Trans(x,dz)\nRot(z,6 z)Trans(z,rz) Rot (z,6 1) Trans(z,rl )\nwhich can be written as :\n-IT 0 D-r I = Rot(x,O)Trans(x,O)Rot(z,yz)Trans(z,bz)\nRot(x,ul')Trans(x,dl')Rot (z,6],) Trans(z,rl ') (7)\nFrom Eq.7 we conclude that this change of variables on frame 1 parameters, will enable us to represent the fixed frame using similar parameters defining the link frames, such that :\nUo = 0, do= 0, 60 = Yz' rO = bz (8)\nAs a conclusion the calculation of -ITn+1 can be obtained as\nfunction of 4(n+2)- 2 geometric parameters.\nTHE IDENTIFICATION MODEL\nThe end effector frame location can be calculated by the direct geometric model given as :\n(9)\nOwing to the errors on the geometric parameters defining the previous matrices, there will be a difference between the real location of the end effector frame and its location as calculated using the direct geometric model. The problem of calibration is to adjust the values of the geometric parameters such that this error is minimum . As -I T n+ I is nonlinear funclion of the geometric parameters , this is a nonlinear problem. Using a first order Taylor development of the parameters errors, a linear differential model can be used (Wu 84, Sugimoto 84, Payannet 86). If the errors are too large an iterative procedure, based on the differential model also, can be applied using Newton-Gauss procedure or any nonlinear optimization method. The differential vectors defining the deviation of the end effector frame from the nominal value due to the differential error in the geometric parameters can be given by (Dombre 88):\n(10)",
+            "where: D is the (3x I) translational differential error vector of the end effector frame,\n/) is the (3x I) rotational differential error vector of the end effector frame,\nLlX defines the (kxl) vector of the errors of the standard\ngeometric parameters, - J is (6xk) matrix, with k = 4(n+2)-2\nTo identify LlX, equation (ID) will be applied for sufficient number of configurations ql, . .. ,qffi, the corresponding Di and\n/)i, for i=I, ... ,m, will be measured and the resulting system of equations will be solved to get the least squares errors solution.\nDefinition of Standard Components of LlX\nFrom section 2, the calculation of \u00b71 T n+ 1 needs 4(n+2)-2\nparameters. So the standard elements of LlX are the differential\nerrors of the corresponding parameters (Lluj' Lldj ' Ll9j' Llrj>. But the representation of equation (4), on which this modeling is based, is singular when the Z axes of two successive frames are parallel. Therefore if Zj.1 is parallel to Zj an additional\ndifferential parameter Ll~j must be added (Sugimoto 84, Payannet 86, Wu 88) , this new parameter will represent a rotation around the axis Yj.l. In this case the error Mj will not be calculated because it is in the same direction of Llrj.} , thus the maximum number of parameters for each frame will remain equal to 4(n+2)-2.\nCalculation of The Columns of.I\nAssuming the matrix defining frame j with respect to the fixed frame as :\n(11 )\nThe calculation of the columns of the J matrix can be calculated as follows (Dombre 88, Khalil 89) :\n- Since the parameter Uj represents a rotation around Xj_l, then\nthe corresponding column is given as :\n. [ sj-IxLj.l,n+1 ] Ju'=\nJ Sj_1 (12)\nWhere: x denotes the vector product, Li,n+ I is the (3x I) vector between the origin of frame i and the\norigin of frame n+ I.\n- Since the parameter dj represents a translational movement\nalong Xj.} , then:\n(13)\nwhere 0(3x I) is the (3x I) zero vector. Similarly we can deduce that:\n. [ ajxLj,n+ I ] J9j = a'\nJ\njrj = [ O(:~I) ] . [ nj-IxLj.l,n+1 ]\n,J~ ' = J \"' J.I\n(14)\nAll the vectors of equations (12 , .. . ,14) are referred to the measuring fixed frame.\nIDENTIFIABLE (BASE) GEOMETRIC PARAMETERS\nUsing equation (ID) we find that all the parameters of 6X cannot be identified separately : - if the elements of a column of J is equal to zero, for all q, then the corresponding parameter cannot be identified. This\nparameter must be eliminated from LlX, and the corresponding column in J will be eliminated too. Let the number of the non zero columns be c. - if the matrix J contains only b independent columns, such that the other (c-b) columns can be obtained as linear combinations of the b independent columns. Then the solution of the system of equation (ID), using sufficient number of configurations ql, ... ,qffi, is not unique. We choose to calculate the base solution which reduces the num ber of parameters to be identified and simplifies the calibration process. Let us rewrite the system (10) as :\nwhere: Ji represents b independent columns of J, h represents (c-b) dependent columns of J,\n(15)\nLl X 1 is the vector of the geometric parameter errors corresponding to the columns of J),\nLl X 2 is the vector of the geometric parameter errors corresponding to the columns of h ,\nSince the columns of h can be written as linear combinations of the columns of Ji, then we can write:\n(16)\nwith B is ( bx(c-b\u00bb matrix with constant elements.\nUsing (16) ,equation (ID) can be written as :\n(17)\n(18)\nIn the identification process, equation (17) will be used instead\nof (10) . The solution will give directly LlXb which is called the base parameters vector. The matrix B is not needed in the identification process. The study of the minimum parameters can be carried out by studying the dependence of the symbolic expressions of the columns of J (KhaliI91).\nNumerical Determination of the Base Parameters\nNumerically the study of the base parameters is equivalent to study the space span by the columns of (rxc) matrix W with r~ c, calculated from J at m random configurations ql, ... ,qffi such that:\nw.[ (19)\nFrom a linear algebra point of view, the classification of the independent columns appears to be a rank deficiency problem."
+        ]
+    },
+    {
+        "image_filename": "designv11_60_0001996_s0736-5845(98)00026-x-Figure33-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0001996_s0736-5845(98)00026-x-Figure33-1.png",
+        "caption": "Fig. 33. Curve segment length S i,i`1 between two knots P i and P i`1 .",
+        "texts": [
+            "3) Because the integration formula in Eq. (A.1) could not explicitly been calculated we used the Simpsons rule [22, 23] for the numerical integration of Eq. (A.4) with *s j \" *u 6 (s@(u j~1 )#sN @ (u j )#s@(u j )) (A.4) and sN @(u j )\"4 s@Auj! *u 2 B . (A.5) Hereby the interval of the curve parameter u is split up into K pieces (A.6): *u\" u i`1 !u i K , u i \"0 and u i`1 \"1. (A.6) The segment length S i,i`1 of the spline curve is the sum of the K pieces\u2019 curve length, calculated in Eq. (A.4), as Fig. 33 shows S i,i`1 \" K + j/1 *s j (A.7) Next step is planning the path parameter s (t) for each curve segment i. Knowing the velocity profile v(t) and having the curve segment length S i,i`1 we can use the characteristic formula to plan the parameter s i (t#*t) along a specified path with Eq. (A.8): s i (t#*t)\"s i (t)#v i (t)*t#1 2 a i (t)*t2. (A.8) where \u00b9 i (t(\u00b9 i`1 and s i \"s i~1 #S i~1,i as Fig. 34 shows. The necessity of transforming the path parameter s (t) into the parameter u (t) of the spline curve is given by the polynomial of the curve (see [11, 12])"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0000865_iciea.2016.7603944-Figure1-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0000865_iciea.2016.7603944-Figure1-1.png",
+        "caption": "Fig. 1. The scheme of eight-rotor aircraft",
+        "texts": [
+            " Eventually, good path tracking performances validate the effectiveness and robustness of the designed controller. The rest of the paper is organized as follows: in Section 2, a dynamic nonlinear model for eight-rotor aircraft is described. Based on the nonlinear model, controllers on account of LADRC and PD strategies are designed for the control system in Section 3.In Section 4, numerical simulations are carried out to show the good behavior of the closed-loop system. Conclusions are drawn in Section 5. II. DYNAMIC MODEL OF EIGHT-ROTOR AIRCRAFT The aircraft is configured as Fig.1. Eight rotors driven by eight brushless DC motors are installed at the end of the connecting rod vertically compose the power units of the aircraft. Movements of the aircraft can be realized through changing the speed of each rotor. The aircraft is a typical unactuated system moves with 6 DOF but only 4 DOF can be controlled independently with four control inputs, which are the total thrust F, produced by four rotors and the torques \u03c4 obtained by varying the rotor speed. Four rotors (1,4,5,8) rotate counterclockwise, while the other four (2,3,6,7) rotate clockwise",
+            "00 c\u00a92016 IEEE 2 2 1 1 2 2 2 2 2 2 1 1 1 1 1 1 1 1 3 3 2 2 1 1 1 1 4 4 2 2 1 1 1 1 5 5 2 2 2 2 2 2 2 2 2 2 6 6 2 2 7 7 2 2 8 8 0 0 0 0 0 0 0 0 U F k k k k k k k k k l k l k l k l R k l k l k l k l k k k k k k k k \u03c6 \u03b8 \u03c8 \u03c4 \u03c4 \u03c4 \u03a9\u2212 \u03a9 \u03a9 \u03a9 \u03a9 \u03a9 \u03a9 \u2212 \u2212 \u03a9 \u03a9 = = \u2212 \u2212 \u03a9 \u03a9 \u2212 \u2212 \u2212 \u2212 \u03a9 \u03a9 \u03a9 \u03a9 \u03a9 \u03a9 (1) Where , ,\u03c6 \u03b8 \u03c8\u03c4 \u03c4 \u03c4 are roll torque, pitch torque and yaw torque respectively, i\u03a9 is the rotor i speed, UR\u03a9\u2212 denotes the transformation matrix from control inputs to the rotor speed, thrust factor 1k and drag factor 2k are positive coefficients, l expresses the distance between rotor and center of the aircraft. Let us consider two frames: earth-fixed reference frame G and body-fixed reference frame B, as depicted in Fig.1.The absolute position of the center gravity and orientation described by Euler angle of the aircraft are expressed in G frame by [ ], , zTP x y= [ ], ,T\u03b7 \u03c6 \u03b8 \u03c8= . The Euler angles are roll angle ( )/ 2 / 2\u03c0 \u03c6 \u03c0\u2212 < < , pitch angle ( )/ 2 / 2\u03c0 \u03b8 \u03c0\u2212 < < , and yaw angle ( )\u03c0 \u03c8 \u03c0\u2212 < < respectively. Vector [ ], ,T u v w\u03c5 = and [ ], ,T p q r\u03d1 = are the translational velocity and angle velocity in B frame. The orientation of the aircraft from G to B is given by the rotation matrix R: cos cos cos sin sin sin cos cos sin cos sin sin sin cos sin sin sin cos cos sin sin cos cos sin sin cos sin cos cos R \u03c8 \u03b8 \u03c8 \u03b8 \u03c6 \u03c8 \u03c6 \u03c8 \u03b8 \u03c6 \u03c8 \u03c6 \u03c8 \u03b8 \u03c8 \u03b8 \u03c6 \u03c8 \u03c6 \u03c8 \u03b8 \u03c6 \u03c8 \u03c6 \u03b8 \u03b8 \u03c6 \u03b8 \u03c6 \u2212 + = + \u2212 \u2212 (2) Since the aircraft are regarded as a rigid body with 6 DOF in a three-dimensional space, whose dynamic model can be derived by Newton-Euler equation"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0003611_ac000427o-Figure1-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0003611_ac000427o-Figure1-1.png",
+        "caption": "Figure 1. Schematic drawing of the differential pCO2 microelectrode employing carbonic anhydrase in the hydrogel layer.",
+        "texts": [
+            "24 The addition of enzyme carbonic anhydrase (CA) to an internal bicarbonate solution in electrochemical and optical pCO2 sensing systems has also been proposed to enhance the response time, since this biocatalyst accelerates the rate of the CO2 hydration reaction.25-28 In our work, we combined two advanced technologies in an attempt to obtain a planar pCO2 microsensing device with faster preconditioning and response characteristics for dissolved CO2 measurement in physiological samples: one is a differential sensing arrangement to facilitate the microfabrication of potentiometric pCO2 electrodes, and the other is the use of CA to shorten total measurement time (Figure 1). The pH-sensitive polymeric membranes adapted for use in constructing a differential pCO2 sensor system in this work function as both a gas-permeable membrane and an internal pH sensing element. In the differential configuration, the pCO2 electrode is made with an unbuffered recipient layer including CA; hence the pH changes are promoted and detected. The reference electrode, on the other hand, employs a strongly buffered hydrogel layer; therefore, diffused CO2 cannot change the pH in the recipient layer"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0000188_0954406214560420-Figure3-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0000188_0954406214560420-Figure3-1.png",
+        "caption": "Figure 3. Diagram of a limb of 6-URS mechanism.",
+        "texts": [
+            " The iterative search method is an important numerical method for the FK problem. But the failure may occur when the iterative search method is directly used to solve the FK problems of a class of mechanisms, whose workspace is restricted. Generally, the workspace restriction is caused by the extremely displacement singularity in the limbs and the coupled motion of the mechanism. at Univ Politecnica Madrid on January 14, 2015pic.sagepub.comDownloaded from Workspace restricted by extremely displacement singularity in the limb As shown in Figure 3, assuming lbc 4 lac, the constraint equation of point bi is bi aij j4lac \u00fe lbc \u00f08\u00de The workspace of point bi is a sphere, whose center is point ai. There is an extremely displacement singularity in the limb. Figure 3 shows that when mobile platform moves from point p to point p0, the middle hinge point ci has no solution. The workspace of the mechanism is restricted by the extremely displacement singularity in its limbs. This situation makes the iterative search method for the FK problem failure. Generally, existed methods to conquer this situation are referred to the damping coefficient to reduce the chances of workspace excess,22 but it is powerless when the target poses are at the edge of the workspace. So, it is necessary to expand the workspace of the limb"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0000936_amm.859.15-Figure3-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0000936_amm.859.15-Figure3-1.png",
+        "caption": "Fig. 3 Part model design.",
+        "texts": [
+            " Additionally, high insulated material was installed to the liquefier to maintain the heat temperature. All these factors was investigated to determine their impacts on the extrusion process [6\u20139]. The features of the new nozzle are depictured in Fig. 2. Notes: Fig. 2 a: screw slot, b: outlet nozzle, c: heating element slot, d: cylindrical liquefier installed with insulator, e: thermocouple slot Sample Fabrication. Sample was designed using Autodesk Inventor Software (Autodesk, USA) and converted into standard triangular language (STL) format as shown in Fig. 3. The dimension of the part model is 132 mm x 86 mm x 9 mm. The part model has been designed based on previous research [10,11]. The first model was printed using the original nozzle (nozzle 1) and the second model was printed using a newly developed nozzle (nozzle 2). The parameters for the printing process were fixed for both nozzles. Printing temperature was set at 220\u00b0 with printing speed of 60mm/s for outer layer and infill. The air gap was maintained at 0 with layer thickness of 0.3 mm and setting for raster angle was set at 45\u00b0"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0000012_ilt-03-2015-0031-Figure3-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0000012_ilt-03-2015-0031-Figure3-1.png",
+        "caption": "Figure 3 Schematic diagram of film shape for thrust bearing with pivoting-linearly pad",
+        "texts": [
+            " The generalized Reynolds equation for a Newtonian compressible laminar fluid was established in cylindrical coordinate system: 1 r e h3 p r e h3r p r 6 r vs ( h) vsr \u00b7 r \u00b7 ( h) r vsr \u00b7 h (1) where / e and are defined by the following equations, respectively: l e 12 e e/ e 2 e e e (2) In equation (2), p e, e, e and are defined as: Thermal elastohydrodynamic lubrication analysis Shun Wang, Qingchang Tan and Zunquan Kou Industrial Lubrication and Tribology Volume 68 \u00b7 Number 1 \u00b7 2016 \u00b7 67\u201375 D ow nl oa de d by R ye rs on U ni ve rs ity A t 1 1: 26 0 2 N ov em be r 20 16 ( PT ) e 0 h dz/h e 0 h 0 z dz dz/h2 e 0 h 0 z z dz dz/h3 e h/ 0 h dz e h2/ 0 h zdz (3) In equation (1), the slip velocity, vs, is given by the express (Jin, 2005): vs 1.98478 10 4 r h 1.4137 104 0.78806 0 r h 1.4137 104 r h 1.4137 104 (4) Oil film temperature is simulated according to the energy equation: cp h3 12 r2 p h T h3 12 p r T r kT h r2 2T 2 h r r r T r T z h3 12 r2 p 2 r2 h \u00b7 4 2 h3 12 p r 2 (5) Apart from these, thermal field of pad may be determined by solid heat transfer equation: 2Tp r2 2 2Tp r2 Tp r r 2Tp zp 2 0 (6) As shown in Figure 3, oil film thickness is expressed by the following equation when the pad is pivoted linearly: h h0 \u00b7 rsin( 0 ) ve r, vt(r, ) s(r, ) (7) Here, ve r, and vt r, are elastic and thermal distortions of composite pad surface, respectively, while s r, represents pad surface shape, including tapers of leading and trailing edge. Elastic deformation of pad surface is formulated by: ve r, 2 E0 A p(r , )r dr d (rsin r sin )2 (rcos cos )2 (8) where E0 is effective elasticity modulus of bearing and is calculated by: 1 Epad 1 Esteel base Dsteel base Dsteel base DPTFE layer 1 EPTFE layer DPTFE layer Dsteel base DPTFE layer E0 2 (1 pad 2 )/Epad (1 collar 2 )/Ecollar (9) Discretize equation (8): D ow nl oa de d by R ye rs on U ni ve rs ity A t 1 1: 26 0 2 N ov em be r 20 16 ( PT ) ve ri, j 2 E0 k 0 M-1 l 0 N-1 K(ri rk, j l)p(rk, l) (10) In equation (10), K ri rk, j l , also tagged as Kk, l i, j , is influence coefficient"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0003339_s0022-460x(87)81301-9-Figure2-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0003339_s0022-460x(87)81301-9-Figure2-1.png",
+        "caption": "Figure 2. Torsional buckling of the system composed of two crank-and-rocker mechanisms.",
+        "texts": [
+            " MECHANISM OF EXCITING VIBRATIONS Many factors having effects on the behaviour of the system can be found directly from the set of equations (5) and (9). Neglecting the time derivatives and MAn in equation (5), one obtains a set of transcendental equations. Assume that the set has the solution 4'n-1 = (Pn = (P,,+l= (Po, where (Po is the root of the equation d1Jl'(4J)/d4J=O. If besides (Po the set of transcendental equations has additional solutions 4Jn- l:;t; 4J\", (P,,:;t; (Pn+l, then torsional buckling of the system may take place. As is shown in Figure 2, under the action of critical torques M B n the cranks move in directions opposite to one another. Since the rocker shaft is subjected to little or no distortion it does not contribute much to the stiffness of the mechanism. Assume that the crank-shaft is operating at a certain speed wand the maximum distortion I(Pn - (pn+11 occurs when the rocker is close to one of the extreme positions \u00abP \"'\" (Po). From Figure 2 one may see that in this case the moments A\" d24J\"/dt2 and B\"(d1Jl'\"/d4J,,)(d24J,,/dt2) may contribute to the dynamic torsional buckling of the system; the moment B,,(d21J1'n/d4J~)(d(P,,/dt)2reduces the vibration amplitude and thereby it stiffens the system. On the other hand, it is possible to show that the same moment may induce relative angular motion of the system elements. Substituting TORSIONAL VIBRATION OF SHAFTS tPn-1(t) = tP,,(t) = <P,,+I(t) 225 (15) into the set of non-linear equations (5) and using relations (14), one finds that all equations in the set become identical",
+            " With the lapse of time the amplitude of vibration decreases. For a certain small amplitude the angular velocity increases to the value corresponding to the actual value of n. It was observed that the resonance vibration was initiated by the function being in phase with the periodic solution of the linear homogeneous equation (9). On the boundary c (Figure 8(a\u00bb the jump in the amplitude is smaller then on the boundary s (Figure 8(b\u00bb. This is due to the stiffening effect of B; (d2 1/'nl d4>~)(dcPnl dt)2 (Figure 2). When this effect vanishes, the second jump up in the amplitude takes place (Figure 8(a\u00bb. The phenomenon of the sudden decrease of the average angular velocity w to the value close to zero is initiated by dynamic torsional buckling of the system (Figure 2). Figure 9 illustrates the behaviour of the system in the combination resonance for a == 10, a21a=0\u00b7558823529 (~=-0'2), w~,/w~=0'25, S2=S3, C t/ (wt2L: A n)=O.002, H3/(wt2B3) = 0'05, e] wg == \u00b10'0002, where Wf2 = (W~, + w~,)/2. For increasing n (Figure 9(a\u00bb resonance occurs for the value of wl(2w~,) corresponding to the lower boundary of the combination resonance (Figure 4(a), a21ii = O'558823529). For decreasing O (Figure 9(b\u00bb the average angular velocity w rapidly decreases from the value corresponding to (0) (b) 0 \u00b78 oe 0 \u00b77 0 '7 0 \u00b76 0\u00b76 0 \u00b75 0 \u00b75 *-;; 13 0 "
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0000814_s1068366616050056-Figure2-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0000814_s1068366616050056-Figure2-1.png",
+        "caption": "Fig. 2. The scheme of the device for tribological tests in the distributed contact conditions.",
+        "texts": [
+            " The L-AN 68 machine oil with a f low rate of 30 drops/min was used as the lubricant. The volume sample wear was determined according to the formula where D0 is the roller diameter, l is the sample width, and b is the width of the wear groove. 2 0 0 0 2arcsin sin 2arcsin , 8V D l b bI D D \u23a7 \u239b \u239e \u23a1 \u239b \u239e\u23a4\u23ab= \u2212\u23a8 \u23ac\u239c \u239f \u239c \u239f\u23a2 \u23a5 \u23a9 \u239d \u23a0 \u23a3 \u239d \u23a0\u23a6\u23ad The relative volume wear was calculated as or where \u03c4 is the friction time and L is the friction path. Tribological tests with the distributed contact were performed using a special device made on the basis of a vertical drilling machine (Fig. 2). Counterbody 6 was fastened in the machine chuck, a working part of which is fulfilled in the form of a tube made of VK6 o V V II = \u03c4 o ,V V II L = 456 JOURNAL OF FRICTION AND WEAR Vol. 37 No. 5 2016 FELDSHTEIN et al. hard alloy with the wall thickness of 1.5 mm and average diameter of 16 mm. Chuck rotation with a frequency of 2300 rpm provided the linear velocity of 115 m/min. Sample 7 was a disc with a central hole. The sample was fastened in bath 1 intended to collect the lubricant with the help of double-end bolt 3 and nut 2"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0003947_rtd2002-1642-Figure8-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0003947_rtd2002-1642-Figure8-1.png",
+        "caption": "Figure 8: Slant between the pivots.",
+        "texts": [
+            " For this reason, vehicles with very high torsional stiffness have serious trouble to cross these kinds of defects, cause the vertical load on some wheels may decrease with derailment risk. In our work we have considered three different events which may cause the wheel unloading and which usually occur together during curving. The first event is caused by a slant between the two rail extended to the entire side of a wagon, and that lead to a superelevation of only one side (left/right) of a bogie respect to the second bogie. We indicate this events as slant between the pivots (fig. 8). The second event is a slant between the rails with a shorter extension, which is applied to a single bogie lifting the first wheel and pulling down the second of only one side of the bogie. This second event is shown in fig. 9 and is indicated as bogie slant. The third effect we have considered is the one due to the lateral acceleration non-compensated by the cant (anc) which, for freight vehicles, can reach a value of 0.6 m/s2 during curving. The effect of anc is an unloading of all the wheels of the inner side of the curve and a loading of the wheel of the outer side"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0000269_9781782421955.540-Figure5-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0000269_9781782421955.540-Figure5-1.png",
+        "caption": "Fig. 5 \u2013 Generation of the base circle B as envelope of the normal lines N.",
+        "texts": [
+            " 4 BASE CIRCLE OF INVOLUTE CIRCULAR GEARS The base circle of circular gears is the evolute of each involute tooth profile and can be determined as the envelope of the family of all straight lines, which are normal to the rack-cutter profile \u03b7 and passing through the instant center of rotation P0. In fact, when the normal line N to both profiles at the contact point P passes through P0, they are conjugate, since the relative sliding velocity is tangent to them at point P. As introduced in (6) and extended here, this family of normal lines N is generated by the pure-rolling motion of the auxiliary centrode \u03b5 that coincides with the pitch line of the rack-cutter, whose motion parameter is the clockwise angle \u03b1, as shown in Fig. 5. The position vector N\u03c0 of a generic point Q of the normal line N with respect to the moving frame O2X2Y2, whose X2-axis coincides with the auxiliary centrode \u03b5 or pitch line of the rack-cutter, is given by the homogeneous coordinates below: 2 cos , sin , 1 T pr r r\u03b1 \u03c6 \u03c6\u23a1 \u23a4= +\u23a1 \u23a4\u23a3 \u23a6 \u23a3 \u23a6N\u03c0 (2) where rp, r, \u03b1 and \u03c6 are the pitch radius of the pitch circle P, the oriented segment from P0 to Q, the clockwise angle that gives the position of P0 with respect to the Y1-axis, and the pressure angle of the rack-cutter profile \u03b7, respectively",
+            " The foregoing procedure proves that the base circle for circular gears and, in general, those of non-circular gears, as shown in (6), can be obtained as the envelope of the family of normal lines N to the rack-cutter profile \u03b7, but this approach is quite complex. Instead, the proposed approach makes use of the Aronhold theorem and the return circle, which eases the formulation and its application. In fact, the same result of Eq. (10) could be obtained by simply tracing the return circle R of diameter \u0394 equal to the pitch radius rp of the gear to generate, as shown in Fig. 5. Thus, the intersection of N with R gives the point T of the base circle B by repeating the procedure for different positions of the pitch line of the rack-cutter, one has the base circle. 5 BASE CURVES OF INVOLUTE NON-CIRCULAR GEARS The approach proposed above for circular gears can be extended to the case of non-circular gears by giving their base curves as sets of points coming from the intersection of the return circle R with the line N normal to the rack-cutter profile \u03b7, which is defined by the pressure angle \u03c6 "
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0001235_b978-0-08-099425-3.00005-4-Figure5.8-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0001235_b978-0-08-099425-3.00005-4-Figure5.8-1.png",
+        "caption": "FIGURE 5.8",
+        "texts": [
+            " Evenwith the simplified laws of friction as described above, a block ofmaterial standing on a surface can be imagined to be subject to frictional forces at the interface of the surface and block.When the surface is level then the block and the surface are pressing on each other with equal and opposite normal forces due to theweight of the block and its reaction. With no other forces acting, there are no frictional forces present, although the capacity for them to arise is present. The frictional forces themselves only arise in reaction to an applied external force e an inclination of the surface or an externally applied force parallel to the plane of the surface (Figure 5.8). The frictional forces calculated by the \u2018laws\u2019 of friction are in fact the maximum capacity for reactive forces to be generated, and when the applied forces are lower than this level then the frictional forces rise and fall to preserve equilibrium. Thus the frictional force in the stationary block is time-varying if the applied force is also time-varying. This concept becomes particularly important for modelling clutches, as discussed in Chapter 8. Frictional forces for a block on an inclined plane"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0001487_ire-i.1956.5007018-Figure1-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0001487_ire-i.1956.5007018-Figure1-1.png",
+        "caption": "Fig. 1",
+        "texts": [
+            " INTRCDUCTION Gray T1) Code, the accuracy is limited by the inside diameter. With sliding contacts it is The art of converting shaft position into possible reliably to resolve a motion of .005\". coded electrical signals has expanded rapidly With a diameter of 3\", the inside circumference dulring the last five years. A number of dif- is 3T. or roughly 10\", and the best practical ferent techniques have been developed; however, angular resolution is 1/2000 of 360' or about this discussion will be confined to encoders similar to Figure 1, and will apply particularly to encoders using sliding contacts. However, the sine of .18' is .0033 (1/300). Therefore, the best practical resoluIn most applications to date these encoders tion of the sine function of the angle is 1/300. have been constructed to produce a coded value Using photocells7Ti is-possible reliably to proportional to the angular displacement of the resolve a motion of .0005\". Therefore, with the input shaft. It has become apparent, however, same size encoder, a photoelectric device can that it is possible to produce coded values equal resolve angles to 1/20,000 of 360' or can resolve to the sine or cosine function of angular shaft the sine function to 1/3000"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0001489_jrproc.1954.274732-Figure2-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0001489_jrproc.1954.274732-Figure2-1.png",
+        "caption": "Fig. 2-The short and open circuit impedances are obtained by joining P to the zero points and points at infinity on the two reactance scales.",
+        "texts": [
+            " 1 represent corresponding reactances. Although the data specifying the transducer have been presented in the form of electrical measurements of a certain kind, the transformation diagram may also be set up from any other specification of the transducer such as the equivalent-T impedances, or short and open circuit impedances; and conversely, beginning from the diagram, any other specification may be deduced. It is sufficient to note the simple constructions which give the open and short-circuit impedances Z1, Zo2, Z,1, Zs2 (Fig. 2). TRANSFORMATION OF COMPLEX IMPEDANCES It remains to be shown how complex impedances may also be transformed by the diagram, and for this purpose we shall need the following two results from the theory of functions of a complex variable. (1) The transformation w =f(z), where w is an analytic function of the complex variable z, is conformal; i.e. angles are preserved. (2) The conformal transformation represented by the bilinear relation a + AZ z' + 8z transforms circles into circles. The constants a, /, y, 6, may be complex"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0000077_icoin.2016.7427086-Figure12-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0000077_icoin.2016.7427086-Figure12-1.png",
+        "caption": "Fig. 12. The path taken by UAV in scenario 2 with constant obstacle velocity. (a) Using To Goal search (b) Using Maximum Velocity search",
+        "texts": [
+            " At t17 it moves at maximum speed until it reaches goal at t33 or 3.3 seconds. Using MV search with same velocity and initial position, as shown in Fig. 11(b), UAV maintains its maximum speed throughout the entire path. It moves to the left at t3 then it starts moving to the right at t13. It reaches goal at t38 or at 3.8 seconds which is slower than TG search by 0.5 seconds. For this scenario, the obstacle\u2019s position at t0 = (6, 1). The obstacle is moving with constant velocity (vx = \u20133 m/s, vy= 4 m/s). Using TG search, as shown in Fig. 12(a), the UAV moves with maximum speed at t1 then stops at t2. It starts again to move at slow speed at t7 and moves at changing speed until t21. From t22, it moves at maximum speed until it then reaches the goal at t35 or after 3.5 seconds. Using MV search, as shown in Fig. 12(b), the UAV first moves towards the goal at t1 then starts to move towards the left from t2 until t25 to avoid collision with the obstacle. At t26, it starts aiming back to the goal. It reaches the goal at t45 or after 4.5 seconds. In this scenario, using TG search gives shorter path to the goal by 1 second. In this scenario, the obstacle\u2019s position is the same with previous scenario t0 = (6, 1) but this time the obstacle is increasing its velocity. The obstacle has initial velocity of (vx = \u20133 m/s, vy= 4 m/s) and obstacle is changing its velocity by (vx = \u20130"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0000772_aim.2016.7576964-Figure1-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0000772_aim.2016.7576964-Figure1-1.png",
+        "caption": "Figure 1. The main components of the proposed drug delivery system for WCE. A: ring-shaped external magnetic system, B: drug release module, C: the capsule robot, D: complementary modules within the capsule (anchoring mechanism, active locomotion system and localization and orientation detection module), E: patient bed, F: clinician, G: joystick, H: Human Capsule Interface. Point P represents the origin of the general coordinate system XYZ, \ud835\udec9\ud835\udc04\ud835\udc0f\ud835\udc0c is taken with respect to the x axis, and \ud835\uded7 is taken with respect to the z axis.",
+        "texts": [
+            " To the best of our knowledge, this is the first time that such analysis is carried out for a drug delivery system in capsule robots. The remainder of this paper is organized as follows. Section II provides the details of the overall system under study. Section III presents the theoretical methods for the analysis of the torque transmitted to a tilted IPM. Section IV provides the verification of the theoretical results with experimental results for flux density, magnetic torque and rotational movement of the IPM and the crank to release drug. Finally, discussions of the results and future work are presented in Section V. Figure 1 shows a ring-shaped external magnetic system made of EPMs, the position (XYZ) and orientation (\u03b8EPM, \u03c6) of which can be controlled from a joystick. A rotating magnetic field around the patient bed is generated when the EPMs are physically rotated about the z axis. In this work, we have used an external magnetic system made of an array of 24 arc-shaped permanent magnets (ASMs) that we have optimized to generate a theoretical value of 500 mT at the centre of the system (x=y=z=0) as shown in Fig. 2",
+            ", N50) is considered to be placed in a prototype of a capsule robot. The capsule robot is to operate within the cylindrical region of interest defined by the maximum radial operating distance r1 and the length of \u2206z=60 mm. Since the drug release mechanism is actuated by a magnetic torque, the operating distance r1 can be increased to meet the requirements of a more realistic medical application (i.e., r1>120 mm) while the dimensions of the external magnetic system are scaled up at the same time [13, 17]. A magnetic torque \u03c4z\u2032, as shown in Fig. 1, will be imparted to the IPM embedded in the capsule robot as its magnetization vector m interacts with the rotating magnetic field created by the ASMs. The transmitted torque \u03c4z\u2032 is then converted into a piston force F by means of a slider-crank mechanism that is physically connected to the IPM. In this way, a piston will push drug out of a reservoir when the ASMs are rotated around the patient\u2019s body as shown in Fig. 1. Within the capsule robot, we define a coordinate system X\u2032Y\u2032Z\u2032, the origin of which coincides with the IPM\u2019s centre. When the IPM\u2019s centre coincides with the centre of ASMs and the axes of the two coordinate systems are aligned, it is possible to approximately transmit a magnetic peak torque \u03c4z\u2032 of 3.5 [mNm] on a 3.1 mm cubic IPM by using only 4 ASMs that generate approximately 114 [mT] at the IPM\u2019s centre [17]. However, in a real application, the IPM can be off the centre and/or tilted as it will move along with the capsule robot",
+            " According to these results, peak torques from 2 to 4 [mNm] were obtained when the IPM was tilted by an angle \u03b8z between 750 and 600, respectively. Consequently, the IPM and the crank of the drug release module can be rotated even if the IPM is tilted by these angles. However, the peak torque of \u03c4z\u2032 continued decreasing if the IPM was further inclined (i.e., for \u03b8z>750), reaching 0 [mNm] at \u03b8z=900. Therefore, the IPM and the crank stall for any angle \u03b8EPM if the IPM has the specific orientation determined by \u03b8IPM=900 (or 2700) and \u03b8z=900. At these values, the inclination \u03c6 (see Fig. 1) and perhaps the position of the external magnetic system would need to be adjusted by the clinician to activate the drug release mechanism. Depending on the need to generate different drug profiles (i.e., changes in the number of doses or changes in release rates), the clinicians may be able to follow different real-time control strategies for the capsule robot by moving the external magnetic system to tailor therapeutic treatments to individuals\u2019 needs. As shown in Fig. 9, we have inserted the cubic IPM into its IPM case (A) which is connected to the crank of the slider-crank mechanism"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0000609_b20057-129-Figure1-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0000609_b20057-129-Figure1-1.png",
+        "caption": "Figure 1. Geometry of the tank support saddle including a segment of the deck.",
+        "texts": [
+            "01 Geometric model of the tank support saddle is shown at Figure 1. In order to build a cost effective structure, the tank saddle includes a polymer layer (Grudzi\u0144ski 2014) used as an insulation. The order and thickness of layers is shown at detail view B (Fig. 1). The overall dimensions of the tank support saddle, not including the deck, are 0.4 m \u00d7 1.4 m \u00d7 4.6 m. Two-dimensional models are often used in order to decrease the calculation time. Such a simplification is allowable if one of the overall dimensions of the concerned body is an order of magnitude smaller than the other dimensions, or if two opposite border surfaces are insulated. In other cases, a distortion in analysis results occur. A difference between 3D and 2D, evaluation is shown at Figures 2 and 3 on the example of a LNG tank support saddle"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0000837_red-uas.2015.7441004-Figure2-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0000837_red-uas.2015.7441004-Figure2-1.png",
+        "caption": "Fig. 2. Pure yawing motion",
+        "texts": [
+            " In aerodynamics Mq and M\u03b4e are the stability derivatives implicit in the pitch motion. We can see these variables in the in Figure 1. The lateral dynamic generates the roll motion and, at the same time, induces a yaw motion (and vice versa), then a natural coupling exists between the rotations about the axes of roll and yaw [2]. In our case, to solve it, we have considered that there is a decoupling of yaw and roll movements [12]. Thus, each movement can be controlled independently. Generally, the effects of the engine thrust are also ignored [2]. In the Figure 2, the yaw angle is represented, which can be described with the following equations: \u03c8\u0307 = r (4) r\u0307 = Nrr +N\u03b4r\u03b4r (5) where \u03c8 represents the angle of yaw and r denotes the yaw rate with respect to the centre of gravity of the airplane. \u03b4r is the rudder deflection. Nr and N\u03b4r are the stability derivatives for yaw motion. The following equations describe the dynamics for the roll angle: \u03c6\u0307 = p (6) p\u0307 = Lpp+ L\u03b4a\u03b4a (7) where p denotes the roll rate and \u03c6 describes the roll angle, it is observed that \u03b4a represents the deviation of the ailerons"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0002020_s003660050027-Figure1-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0002020_s003660050027-Figure1-1.png",
+        "caption": "Fig. 1. Optimal shape design of a rotor blade support.",
+        "texts": [
+            " In these cases, the question of optimisation was not dealt with until damage had occurred in-service: the author\u2019s industrial partners realised (often too late) that their designing left quite a bit to be desired. They would then call for the author\u2019s help in using optimisation programs to supply them with an improved \u2018shape\u2019. These shapes were reached despite the technological limitations being very severe at this stage; so severe, in fact, that engineers were powerless to resolve the problem. Innumerable problems such as this were dealt with. Figure 1 exemplifies this well. Some results in investigating the optimal shape of a helicopter gear box are presented here. The part is located just below the rotor blades. This axisymmetrical structure is long enough to be considered as being clamped at its base. The design variables and constraints were given by the manufacturer. The objective is the minimisation of the maximal tangential stress along the exterior boundary. Figures 1 and 2 show the initial and final boundaries and the tangential stress distribution along the exterior boundary"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0000878_978-3-319-47614-8-Figure2.27-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0000878_978-3-319-47614-8-Figure2.27-1.png",
+        "caption": "Fig. 2.27 Robert Hooke\u2019s apparatus to determine elasticity of a spring",
+        "texts": [
+            " We have to bring this material property into the scene to find a deformation under load. Robert Hooke, a contemporary of Newton in 1660 discovered the law of elasticity which bears his name and which describes the linear variation of tension with extension in an elastic spring. We find from tests that stresses and strains are related amongst themselves through the material characteristics. That\u2019s how different materials find applications under different conditions. Robert Hooke used experimental apparatus like that shown in Fig. 2.27 to measure the extension of a spring as a function of the imposed load and thereby formulated his famous law; \u201cAs the extension, so is the force.\u201d In modern parlance, this is \u03c3 = E\u03b5; Stress is proportional to Strain with the constant of proportionality given by the Young\u2019s modulus, E. Uniaxial tension test is commonly used to determine the stress-strain behavior of isotropic materials (materials having identical properties in any direction). A typical modern tension testing machine is shown in Fig"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0003190_1.1424598-Figure1-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0003190_1.1424598-Figure1-1.png",
+        "caption": "Fig. 1. Free-body diagram of a spool of mass m and moment of inertia I about its center C. A constant tension force T has a line of action making an angle with respect to the forward horizontal direction (which is to the right in this sketch). The directions of the static frictional force fs and of the linear acceleration of the center of mass a follow from the equations of motion and depend on , as discussed in the text.",
+        "texts": [
+            " In addition to the familiar angle c in the forward direction at which the spool cannot roll without slipping, there is a second characteristic angle m in the backward direction at which the spool always rolls without slipping. (Throughout this article, \u201cforward\u201d refers to the direction in which the string would unwind off the bottom of the spool.) In standard homework problems and class demos involving pulled cylinders, m is not investigated. The critical pulling angle c at which the spool slips in place without rolling occurs when the line of action of the pulling force T is directed through O, as illustrated in Fig. 1. All other forces (the normal force N, the spool\u2019s weight mg, and the frictional force f ) also have lines of action through O; thus, there is no net torque about O and the spool does not rotate about this point. Let the inner radius of the spool (about which the string is wrapped) be R1 and its outer radius (with which it contacts the floor) be R2. From the geometry it follows that c = R1/R2. For example, if R1 = 0.75 R2 then c = 41 . Now consider pulling at a steeper angle,3 > c. A free-body diagram for the case in which the spool rolls without slipping (so that Acceleration of a Pulled Spool 482 THE PHYSICS TEACHER \u25c6 Vol. 39, November 2001 the friction is static, fs ) is sketched in Fig. 1. The directions of the friction and acceleration vectors are correct as drawn provided that c < < m , where m is some maximum angle whose value will be calculated below. Applying Newton\u2019s second law horizontally gives fs \u2013 T cos = ma, (1) from which we see indeed that the frictional force must point in the same direction as the horizontal linear acceleration a. The rotational analog of Newton\u2019s second law counterclockwise about point C is = TR1 \u2013 fsR2 = I = Ia/R2, (2) where is the angular acceleration of the spool and I is its moment of inertia about C",
+            " 39, November 2001 where use was made of Eqs. (1) and (3). Substituting this into Eq. (3) gives an expression for the maximum acceleration if the spool is to not slip, amax = . (7) This is plotted in Fig. 2 for the case of s = 0.2 and, as above, = 1 and R1/R2 = 0.75. As expected from the preceding discussion, amax = 0 at = c. Thus, if the tension is smaller than the value given by Eq. (6), the spool will not move at this angle.5 Repeating the analysis for a shallow pulling angle, < c, we get the same free-body diagram as in Fig. 1, except that now the acceleration points to the right rather than to the left. The friction must be the source of the clockwise torque about C, and must therefore continue to point to the left in the diagram. Hence the acceleration and its maximum value are given by the negatives of Eqs. (3) and (7), respectively. Once again a must be positive, because < c \u27af cos > R1/R2. Nevertheless I have plotted it as being negative on the graph in Fig. 2 to show that the maximum acceleration varies smoothly as crosses through c",
+            ", one for which I > mR1R2) to be: (i) c = cos-1[(R \u2013 /R2)2] with 0 < c < 90 at which the spool rotationally slips in place without rolling regardless of how large the coefficient of static friction is; and (ii) m = cos-1 [-(R \u2013 /R)2] with 90 < m < 180 at which the spool rolls without slipping regardless of how small the coefficient of static friction is. In the Classroom These results can be applied to an introductory physics course in two ways. On the one hand, a three-part homework problem can be constructed. The students are given the freebody diagram in Fig. 1. In the first step, the students are asked to write down the horizontal component and rotational analog of Newton\u2019s second law using only the symbols in the freebody diagram. In the second part, the students are to solve these two equations to find the special angle m at which the frictional force falls to zero. Finally, they should be told to observe from their result that this angle is unphysical unless I is larger than some minimum value and then asked what this limiting value is in terms of the mass and radii of the spool"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0001235_b978-0-08-099425-3.00005-4-Figure5.14-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0001235_b978-0-08-099425-3.00005-4-Figure5.14-1.png",
+        "caption": "FIGURE 5.14",
+        "texts": [
+            " While it can be asserted with some certainty that the integrated vertical forces throughout the contact patch must sum to the applied load, it is an error to assert that the contact patch dimensions can be simply predicted by, for example, load, width and inflation pressure. It can be generally observed for conventional tyres that vertical stiffness is broadly linear until rim contact occurs (op het Veld, 2006) and (Reimpell and Sponagel, 1988). A moment\u2019s consideration suggests that the growth of contact patch area is nonlinear with vertical stiffness and therefore the resulting average contact pressure cannot scale linearly. In Figure 5.14, it can be seen that the contact patch length can be found using Pythagoras, as given in Eqn (5.24) \u00f05:24\u00de Presuming the width of the contact patch is constant, for a linear stiffness we can say the average contact pressure P must follow a form as follows \u00f05:25\u00de Contact patch length. where k is the tyre vertical stiffness and W is the tyre width. Considering the tyre as an isothermal volume, we might estimate the change in nominal pressure PDz from initial pressure P0 by calculating a change in volume compared to initial volume V0 \u00f05:26\u00de \u00f05:27\u00de \u00f05:28\u00de \u00f05:29\u00de where rw is the radius of the wheel rim"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0000197_icma.2014.6885793-Figure1-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0000197_icma.2014.6885793-Figure1-1.png",
+        "caption": "Fig. 1. The pursuit evader model",
+        "texts": [
+            " In Section VII, the simulation results are presented. Finally, conclusions and future work are given in Section VIII. The application that we will use for this study is the pursuitevasion differential game [20]. In the pursuit-evasion game there are one or several pursuers that attempt to capture one or several evaders in minimal time while the evaders try to escape or to maximize the capturing time [20]. Hence, this problem can be considered as an optimization problem with conflict objectives [18]. Fig. 1 shows the pursuit-evasion model with its parameters. The dynamic equations that describe the motions of the pursuer and the evader robots in this game are [11], [21] x\u0307i =Vi cos \u03b8i y\u0307i =Vi sin \u03b8i \u03b8\u0307i = Vi Ri tan ui (1) where i is \u201de\u201d for the evader and is \u201dp\u201d for the pursuer. Also, (xi,yi), Vi, \u03b8i, Ri, and ui refer to the position, the velocity, the orientation, the turning radius, and the steering angle respectively. The steering angle is bounded and is given by \u2212uimax \u2264 ui \u2264 uimax , where uimax is the maximum steering angle"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0002669_jsco.1998.0200-Figure2-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0002669_jsco.1998.0200-Figure2-1.png",
+        "caption": "Figure 2.",
+        "texts": [
+            " This issue, known as the inverse kinematics problem, amounts to solving a polynomial system where the unknowns are the sines and cosines {si = sin(\u03b8i), ci = cos(\u03b8i), i = 1, . . . , 6} of the six joint angles {\u03b8i, i = 1, . . . , 6}. For general robots the solution of such systems is quite involved, as noted in the next example. But for robots of particular geometry the solution can be easier to achieve. For instance, if the robot is constructed so that the last three joint axes intersect at one point, the corresponding system essentially simplifies, since the robot has a sort of wrist (this is represented in Figure 2 by a point where the three joints coincide) that takes care of the tip\u2019s orientation. Thus, instead of six unknown angles we are reduced to finding the first three (to position the wrist). Craig\u2019s (1989) well-known book contains a detailed exposition of this particular case. There it is shown that the system solution for the third joint angle can be expressed as (r \u2212 k3)2 4a2 1 + (z \u2212 k4)2 sin(\u03b11)2 = k2 1 + k2 2 where k1, k2, k3, k4 are linear functions of s3 and c3, a1, \u03b11 are parameters describing the robot\u2019s geometry (such as the length of the links or the relative angles between two consecutive joints) and z, r are some input data for the tip position"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0000185_20140824-6-za-1003.00239-Figure3-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0000185_20140824-6-za-1003.00239-Figure3-1.png",
+        "caption": "Fig. 3. Test setup for the hydraulically-actuated fin loading",
+        "texts": [
+            " Considering the conditions given above, the control law is obtained for the sliding mode control system with the varying sliding surface as follows:      xsv ddd cKK xxxssgndddx u 111112212     (41) In the second control algorithm with a varying sliding surface, parameters \u03bb and \u03b2 are specified by means of the fuzzy logic-based triangular membership functions unlike the preceding linear approach. In this second case, the control law formulated in equation (41) is used, too (\u00d6zakal\u0131n, 2010). In order for the control systems whose mathematical models are developed as explained above to be realized, the test setup consisting of a hydraulic actuator used as the actuator, speed reducer, and torquemeter is given in Fig. 3. system (\u00d6zakal\u0131n, 2010). Having performed the system identification works using the test setup shown in Fig. 2, the natural frequency and damping ratio values of the FLS whose dynamic behavior is described as in equation (22) are calculated to be 7.071 Hz and 0.349, respectively. The numerical values of the system parameters are determined as submitted in Table 1 so as to be used in computer simulations. Furthermore, the bandwidth (c) and damping ratio (c) parameters of the control system are considered to be 62",
+            " In the real-time computer simulations and tests, the electrical current capacity of the driver of the CAS, sampling frequency of the control system, and duration of the computer simulations are chosen as 8 A, 2000 Hz, and 1 s, respectively (\u00d6zakal\u0131n, 2010). Here, since the expected settling time value as per the designated bandwidth quantity becomes smaller than 200 ms, the simulation duration is selected to be 1 s in order to increase the resolution of the results. In the end of the computer simulations carried out in the MATLAB SIMULINK environment and the tests conducted on the setup seen in Fig. 3., the settling time, average input (control) voltage of the servovalve, maximum overshoot, and steady state error values are presented in Table 2 and Table 3 versus the step input with the amplitude of 10 as the abbreviation SMC stand for the sliding mode control regarding all the PID-type control, sliding mode control with a constant sliding surface, sliding mode control with a linearly-varying sliding surface, and sliding mode control with a fuzzy logic-based varying sliding surface approaches such that they include all kinds of switching functions, namely the signum, hyperbolic, and fuzzy logicbased switching functions"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0000878_978-3-319-47614-8-Figure2.19-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0000878_978-3-319-47614-8-Figure2.19-1.png",
+        "caption": "Fig. 2.19 Mohr\u2019s circle stress transformation",
+        "texts": [
+            " 52 2 Continuous Solid Note that the circle Eq. (2.19) is rx0 rAvg 2 \u00fe s2x0y0 \u00bc R2 \u00f02:19a\u00de The two principal stresses and the maximum shear stress are shown on Mohr\u2019s circle. Recall that the normal stresses equal the principal stresses when the stress element is aligned with the principal directions, and the shear stress equals the maximum shear stress when the stress element is rotated 45\u00b0 away from the principal directions. Also remember from Eq. (2.18), the angle on Mohr\u2019s circle is 2\u03b8, see Fig. 2.19. As the stress element is rotated away from the principal (or maximum shear) directions, the normal and shear stress components will always lie on Mohr\u2019s Circle. The angle between the current axes (X and Y) and the principal axes is defined as \u03b8p, and is equal to one half of the angle between the line Lxy and the \u03c3axis as shown in Fig. 2.20; the procedure to obtain stresses on another plane is illustrated in Fig. 2.20 with the following steps. 1. Locate the two points (\u03c3x, \u03c4xy) and (\u03c3y, \u2212\u03c4xy). 2"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0002537_a:1007917003845-Figure4-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0002537_a:1007917003845-Figure4-1.png",
+        "caption": "Figure 4. Top view and side view of two PUMA robot system.",
+        "texts": [
+            " If the resultant force/moment is applied to the object via each robot as sx = Do{\u03d5\u0308d +K1(\u03d5\u0307d \u2212 \u03d5\u0307) +K2(\u03d5d \u2212 \u03d5)}+ ho(\u03d5\u0307, \u03d5) (25) where \u03d5d, \u03d5\u0307d, and \u03d5\u0308d means the desired values, then, combined with (24), the motion is governed by the following equation (\u03d5\u0308\u2212 \u03d5\u0308d) +K1(\u03d5\u0307\u2212 \u03d5\u0307d) +K2(\u03d5\u2212 \u03d5d) = 0 (26) where K1 and K2 are [6 \u00d7 6] gain matrices which guarantee asymptotic stability. The resultant force/moment (25) is used in the force distribution part of Figure 3 to obtain optimal force/moment ix of each robot by the proposed Dual method and GIQP method. In the following subsections, the optimal force distribution problem is solved by using sx of (25) in the equality constraints with the assumption that \u03d5 \u2261 \u03d5d for convinience. The two cooperating PUMA robot system is illustrated in Figure 4. The distance between the two robots is 0.9 m, and the length and the mass of the object JINT1410.tex; 18/02/1998; 16:12; v.7; p.16 are assumed to be 0.1 m and 5 kg, respectively. The world coordinate system {X,Y,Z} coincides with the coordinate system of robot 1, which is located at the base of robot 1 in M.K.S. unit. Kinematic and dynamic parameters of a PUMA robot are so well-known in [18] and [19] that the formulations of both optimization methods are obtained immediately. We choose three tasks for test, as shown in Figure 5"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0000077_icoin.2016.7427086-Figure14-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0000077_icoin.2016.7427086-Figure14-1.png",
+        "caption": "Fig. 14. The path taken by UAV in scenario 5 with decreasing obstacle velocity. (a) Using To Goal search (b) Using Maximum Velocity search",
+        "texts": [
+            " It starts to move again with varying speed starting from t7 until t10 then maintains to move at maximum speed from t11 until it reaches the goal at t32 or after 3.2 seconds. Using MV search, as shown in Fig. 13(b), the UAV first moves towards the goal at t1 then starts to move towards the left from t2 until t11. It moves towards the right at t12 then moves directly towards the goal from t19 .It reaches the goal at t34 or after 3.4 seconds which is slower than TG search by 0.2 seconds. In this scenario, the obstacle\u2019s position is at t0 = (5, 11) with constant velocity (vx = \u20132 m/s, vy= \u20131 m/s). Using TG search, as shown in Fig. 14(a), the UAV maintains maximum speed until t7. At t8, it stops moving and starts again to move at slow speed at t24 then starts to accelerate after the obstacle has passed. It then reaches the goal at t50 or after 5 seconds. Using MV search, as shown in Fig. 14(b), the UAV moves towards the right at t8 then it starts aiming back to the goal at t24. It reaches goal at t33 or at 3.3 seconds which is faster than TG search. In this scenario, the obstacle\u2019s position is the same with previous scenario t0 = (5, 11) but the velocity of the obstacle is decreasing. The obstacle\u2019s initial velocity is (vx = \u20136 m/s, vy= \u2013 3 m/s) and obstacle is changing its velocity by (vx = 0.2 m/s, vy= 0.1 m/s). In TG search, as shown in Fig. 15(a), UAV moves at maximum speed then stops at t6 to avoid the approaching obstacle"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0001235_b978-0-08-099425-3.00005-4-Figure5.60-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0001235_b978-0-08-099425-3.00005-4-Figure5.60-1.png",
+        "caption": "FIGURE 5.60",
+        "texts": [
+            " For the purposes of this document it is assumed the road is flat and only one point of contact occurs. Tyre model geometry. Inclined tyre geometry. The camber angle g between the wheel plane and the surface of the road is calculated using \u03b3 = \u03c0/2 \u2212 \u03b8 \u00f05:50\u00de where \u03b8 = cos-1 ({Ur} \u2022 {Us}) \u00f05:51\u00de The vertical penetration of the tyre dz at point P is given by \u03b4z = R2 - |CP| \u00f05:52\u00de In order to calculate the tyre forces and moment it is also necessary to determine the velocities occurring in the tyre. In Figure 5.60 the SAE coordinate system is located at the contact point P. This is established by the three unit vectors Tyre geometry and kinematics. {Xsae}1, {Ysae}1 and {Zsae}1. Note that referring back to Chapter 2 the subscript 1 indicates that the components of a vector are resolved parallel to reference frame 1, which in this case is the Ground Reference Frame (GRF). Using the triangle law of vector addition it is possible to locate the contact point P relative to the fixed GRF O1 {RP}1 = {RW}1 + {RPW}1 \u00f05:53\u00de At this stage it should be said that the vector {RPW}1 represents the loaded radius and not the effective rolling radius of the tyre"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0002359_pesc.1998.703387-Figure4-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0002359_pesc.1998.703387-Figure4-1.png",
+        "caption": "Fig. 4. Optimal current in field weakening region 11.",
+        "texts": [
+            " Field Weakening Region I In field weakening region I, maximum torque point can be determined as the intersection of the voltage limit circle (19) and current limit circle (4) as shown in Fig. 3. D. Field Weakening Region II If the speed of motor is getting increased, the radius of voltage limit circle is getting reduced. In this case, the following function is weld to decide optimal currents. The contact point of the volltage limit circle (19) with this function (23) gives d q ament references (24) and (25) as shown in Fig. 4. E. Discussion The conventional field weakening scheme can produce maximum torque at steady state. However, the proposed field weakening can also guarantee the maximum torque operation both at steady state and transient state. At steady state, the equation A:r = Lmi; is valid. The substitution of &,=Lmi; into (21) anti (24), and rearranging of the resulting equation produces the exactly same d-axis current of conventional approach in (9) and (14). So, the proposed field weakening scheme is same as the conventional one in steady state"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0001562_t-aiee.1941.5058220-Figure4-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0001562_t-aiee.1941.5058220-Figure4-1.png",
+        "caption": "Figure 4. Manual reset protector with bi- Figure 5. Manual reset prometallic element tector with \"solder-pot\"",
+        "texts": [],
+        "surrounding_texts": [
+            "TI L t t I n.ftories as \"Inherent Overheating Protec-he Inherent Overheating Protection OF tive Devices.\" Motor protectors have many modifica-\ntions. When applied on certain types ofSingle-Phase Motors service they are of the automatic reset type and as the name implies such a pro-\nC. P. POTTER tector stops the motor when it reaches a FELLOW AIEE dangerous temperature and automatically\nstarts the motor when its temperature decreases to a safe value. For other types of\nTHE problem of controlling and restrict- velopment and are widely used today in service, protectors are of the manual reset Ting the temperatures of motor wind- motor starting switches. As the name type which must be reset by an attendant,\nings, is of vital importance to the motor indicates, these devices are actuated by who presumably inspects the installation manufacturer, the motor user, and the the heating of a resistor, the temperature to find out what caused the protector to general public. The motor manufacturer of which depends upon both the ambient open, before resetting it. Protectors is anxious to have winding temperatures temperature and the motor current. If which are used in motors 1 hp and smaller restricted to a safe value, so that his prod- the thermal characteristics of the resistor usually break the line current, thus elimiucts may have customer acceptance. The are like those of the motor, and if the re- nating the necessity of providing the ormotor user is interested in winding tem- lays have the same ambient temperature dinary overcurrent protection required by peratures because they affect motor life as the motor, thermal relays installed in the Underwriters' Laboratories. Protecand maintenance cost. The general pub- starting switches should be entirely ade- tors which are used in motors larger than lic is not conscious of any interest in the quate for the protection of motor wind- 1 hp do not ordinarily break the line curmatter, but is certainly affected in all ings. The first of the above conditions is rent, but are connected in the pilot circuit cases where excessive temperatures result of the magnetic switch which starts and in fire hazards. CONTACT WELDED stops the motor.\nIn the early days of the industry, mo- Motor protectors of the automatic reset tors did not have adequate protection type contain some sort of a thermostatic against overheating, because suitable pro- device such as a bimetallic disk or strip, tective equipment was not available. which may be mounted adjacent to a Motor circuits were fused, but the fuses BUIT-IN LOW EXPANSIVE heater coil through which the motor curhad to be large enough to accommodate rent flows. Figure 1 shows a cross-section the motor starting current and were, HIH EXPANSIVE of a device employing a heater coil, manutherefore, too large to properly restrict the / SURFACE factured by the Spencer Thermostat running current or protect the windings. SCREW e e Company. In other types the heater coil However, the lack of proper protection is omitted and the bimetallic strip acts as was partially offset by the special atten- MOULDED PHENOLIC a resistor. A protector of the latter type tion which was given motors by their HOUSING manufactured by Cutler-Hammer, Inc., owners and attendants. Figure 1. Cross-section of protector with is shown in figure 2. In both cases the biThe shortcomings of fuses led to the de- heater coil metallic elements carry contacts which\nvelopment of magnetic type relays pro- are normally closed and which are convided with dash-pots, which retard the only approximated and the second one nected in the motor line. These protecoperation of the relays to such a point cannot always be met, therefore, starting tors are mounted inside the motor and that they can be set to operate close to switches provided with thermal relays when the temperature inside the motor the full load current of the motor and still must also be used with discrimination, reaches the maximum safe value, the bihave enough time delay to allow the motor metal snaps to its hot position, the conto start. These relays were incorporated Inherent Overheating Protective tacts open, and the motor is disconnected in motor starting switches and in most Devices from the line. When the temperature of cases provide adequate motor protection, the motor decreases sufficiently, the bibut they have one undesirable character- Motor overheating is usually due to metal snaps back to its normal position istic. Their operation depends entirely overloading, buit may also be caused by and the motor is automatically started. on motor current, while the temperature too frequent startiig, by foreign material Automatic reset protectors are used on of motor windings depends upon the room becoming lodged in the motor openings, motors which drive stokers, fans, blowers, temperature as well as the motor current. by installing the motor in a non-venti- pumps, small refrigerating outfits, etc. A magnetic type relay which will protect lated compartment, or in a number of Automatic reset protectors are built in a motor located in a room of normal tem- similar ways, and such overheating can- many styles, some of which are illustrated perature may therefore not protect the not be prevented by any device which is in figure 3. same motor when it is operated in a high mounted outside of the motor. Complete Motor protectors of the manual reset ambient temperature. protection of motor windings can only be type may employ a bimetallic element or Thermal type relays were the next de- obtained by using thermal devices which they may have a thermal device of the\nPaper 41-154, recommended by the AIEE com- are mounted inside the motor enclosure mitee on electrical machinery, and nresented at the and which make use of resistors or \"heater AIEE South West District meeting St. Louis, Mo.,whctemorcuen October 8-10, 1941. Manuscript submitted July coils through which the motor current 30,1941; made available for preprinting August 28, passes. These devices are ordinarily\nc- P PorrERiswiththewanerElectriccorpo known as \"motor protectors\" and are tion, St. Louis, Mo. designated by the Underwriters' Labora- Figure 2. Protector without heater coil\nNOVEMBER 1941, VOL. 60 Potter-Overheating Protection of Motors TRANSACTIONS 993",
+            "are used principally on motors which operate oil burners and other devices where it is desirable to inivestigate the\n994 TRANSACTIONS Potter-Overheating Protection of Motors ELECTRICAL ENGINEERING",
+            "___ ___ ___ ___ ___ U~~~~~~~PM PER CENT FLsC 4o PERCENT OF FULL LOAD0 0 > PER C N E N\n2OOI20Hol<80 X cX 16 1t700 w w mUJ_~~~~~~~__0 CURRENLi 6'0 '~~~~~~~~~~~~ CURRENT INGMEATR~.. __120 PER CENT OF 1300\nW T USYNCHRONOUSESPEED\nH LU Q.~ ~ ~ ~ ~ ~ ~~~~~.L\nFiure 6. Winding temperature versus time 0.\nol ~ poetrNomlvlde28 degee :miet 0 1/\n0 30 60 90 120 150 180 40 45 50 55 60 65 70 MINUTES FREQUENCYFigue6.Wining empeatue vesus ime 60 70 80 90 100 110 120 Figure 10. Motor characteristics versus fre-PE-R CENT OF NORMAL VOLTAGE\ntmxmmlaspr db\nNormal VOltage and frequency. 'A.-horse- Figure 8. Motor characteristics versus voltage qunyptmxiuroadspemiteob power repulsion-start induction motor with at maximum loads permitted by protectorprtco protector Normal voltage. 28 degrees ambient. 1I3Curve A: At mdximum current permitted Normal frequency. 28 degrees ambient. horsepower repulsion-start induction motor\nwihot,ntrrptonbyprto 1/3-horsepower repulsion-start induction mo-\nis usually much less than the ultimate time and protector limited temperature qut apid wt te volte. t willge temperature which would be reached of the 1/3 hp repulsion start induction seen that if this motor isdvriving a comwithout a protector. motor at normal voltage and frequency pressor oripumptan is drably ovCurve A in figure 6 shows the relation with the rotor stalled. This is an ab- loaded, it will be un ableythel\nbetween time and winding temperature normal condition, but is one which might at the lower voltages, because tbe protecfor the maximum current which the pro- exist in case of a mechanical defect in the ato lwi r andtdisconnect the mroto tector will allow the motor to carry con- motor or the machine which it drives. from the supply circuit tinuously when it is operating at normal The protector trips and resets many Locked tests were also made at various voltage and frequency. In this case the times before the winding temperature voltages and the results are plotted in protector does not trip but any additional reaches a constant value, which, however, figure 9. In this case the locked current load would cause it to do so. The maxi- is considerably less than for the running varies at about the same rate as the mum winding temperature is 94\u00b0C, which condition and is well within safe limits. voltage, but the protector limited temis much lower than could be allowed. Tests were made with the motor operat- peratures change very little as the voltage Curve B is plotted for a current 10% ing at voltages above and below normal, varies. This difference in protector greater than the maximum current which to determine the effect of voltage on the operation between running and locked the protector will carry continuously and winding temperatures permitted by the condition is due to the fact that when the curve C is plotted for a current 20% protector, and the results are shown in motor is locked, the protector is essengreater than the maximum current which figure 8. It is interesting to note that tially a constant temperature device and the protector will carry continuously. In under these conditions the protector is its operation depends almost entirely on both these curves the protector trips and essentially a constant current device and the temperatul-e of the heater coil which resets several times before the winding allows the same current to flow at voltages is adjacent to the bi-metal. Since the temperature reaches a constant value, from 70 to 120% of normal voltage. The heater coil temperature and the winding\nprotector limited temperatures at 70% temperature both depend on time and o00 _ _ _ _ _ voltage are slightly less, and at 120% m\nX ~~~~motor current, it logically follows that\n100 '5X9 10 loo 10fs.I 80 I__\n80 t | 1< 1 M|W#M WIND)NG U_\nTEMPERATURE 8o-JU8 ID EPERATURE 8cocX 60< t1 S j V ~~~~~~80 0 TP08J V80 8 -j ;:4 ui 0<W60>6 t 1 660_ 70 80 9 12 0 60 30 60 90 120 150 PER CENT OF NORMAL 0 AMPERES\ncr 40 Z~~~~~~~~~~~~~~~~~~~ a:\ncrAO~ ~ INTE Fiur 9. Loke chrctrstc zessvl- F 1 okdcaatrsisvru\nLu 404< H40 4-0.Figur_ 7. Winding temperature versus time age frequency\nRoto loced LUmlvldeddfeuny omlfeuny 1dgre min. Nrd ot.3 ere min.13 /horeoe____io-sdr inucio mo- l/-hrspoe __usinstrinucio mo_ospwr_plin- nuto oo 20rHit prteto to ihpotco2Ch rtco\nNOVEMBER M1NUTE Figur 9. LockedhrhatnPracteritics vesu volt-r FigureION119okdchrce9tisvru"
+        ]
+    },
+    {
+        "image_filename": "designv11_60_0002266_21.214794-Figure3-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0002266_21.214794-Figure3-1.png",
+        "caption": "Fig. 3). Lemma 3: For a cylinder C symmetric about z-axis, a finite line segment lies outside C if, when projected to the z-y plane, its normal through the origin is greater than the radius of C (see Fig. 4).",
+        "texts": [
+            " In the case of Lemma 3, projections of the end-points to the I-y plane are defined by We can arrive at the solution similarly: Using the 2-D equivalent of (33, we can obtain the normal from the line through the axis of the cylinder, p,,(X). Note that p(X) or pzu(X) is a point on the given line segment only if 0 I X 5 1. On the other hand, if x > 1 or X < 0 in (39) or (41), the finite line segment does not have a normal that passes through the origin. In that case one only needs to test if one of the two end points of the line segment lies outside the sphere for Lemma 2 or outside the cylinder for Lemma 3. Fig. 3 illustrates the two cases of X < 0 and X > 1, repsectively. B. Feasibility Evaluation for PUMA 560 Referring to Fig. 1, the workspace of PUMA 560 is characterized by three 3-D surfaces, a cylinder and two spheres. For a trajectory to be feasible, it must lie outside the cylinder of radius d3 and the sphere of radius R2, and it must lie inside the sphere of radius RI. First of all, in order for p to lie inside SI, Lemma 1 must be satisfied, Le., it is necessary that lP l l IR1 and IP2lIR1. (42) Secondly in order for it to be outside S2, first use (39) to find the corresponding for the given p1 and p2"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0001948_5326.760575-Figure7-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0001948_5326.760575-Figure7-1.png",
+        "caption": "Fig. 7. Slice for the contour line.",
+        "texts": [
+            " Therefore, the occluded regions are interpolated with gradually decreasing gray value in the radial direction and are also merged to bear the final height map for extracting contour lines. Fig. 6 shows the results of interpolation and merging of the height map. In the final image processing step, the contour lines for the landing height of the buckets are extracted. Sliced images corresponding to the contour lines are obtained by thresholding the height map from the prior image processing step. Since the back side of the sliced image consists of false data by image processings and has no relation with determination of the landing point, it is cut out from the image. Fig. 7 illustrates the procedure for contour line extraction. The highest point in the height map and the center line of the horizontal scanning angle are obtained, respectively. Then, the split line which passes the highest point and perpendicularly crosses the center line is obtained. The rear region of the split line in the sliced image are excluded from consideration. A contour line is extracted from the edge points in the front side of the split line in the sliced image by using the edge following technique [7]"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0000018_2016-01-1468-Figure3-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0000018_2016-01-1468-Figure3-1.png",
+        "caption": "Figure 3. Transition in Striation Direction as a result of Braking (not to scale).",
+        "texts": [
+            " 121 In order to analyze striations in practice, an understanding of the sensitivity of the equations and how error propagates through the analysis would be useful. In this section, the equation for longitudinal slip (Equation 1) will be studied. Application of the brakes has the effect of changing the angle of the striation marks [1,2,3,4]. Specifically, the marks change from a direction perpendicular to the tire heading with no braking, to a direction parallel to the wheel hub velocity (parallel to the tire mark) with full braking. The change in striation angle with braking is depicted in Figure 3. No braking, partial braking and full braking scenarios are depicted from left to right. In all scenarios the tire is at the same slip angle, \u03b1. The tires are moving from bottom to top on the page, as indicated by the blue arrow. The value \u03ba is also introduced in Figure 3 which represents the angle between perpendicular to the tire and the striation direction. In image on the left of the figure, no braking is occurring. The striations (shown in bold lines) are perpendicular to the tire heading. Mathematically, this case can be described as \u03b1 + \u03b8 = 90 and \u03ba = 0. Partial braking is depicted in the middle image. The brakes have changed the direction of the striations an angle \u03ba from perpendicular to the tire. In the case of partial braking, \u03b1 + \u03b8 + \u03ba = 90. On the right, the brakes are applied fully, locking the wheel. With full braking, the striations are parallel to the velocity direction of the wheel hub, which is parallel to the tire mark. Under the full braking condition, \u03b8 = 0 and \u03b1 + \u03ba = 90. The tires in Figure 3 all have the same slip angle. Now consider the effect of changing the slip angle. Under full braking, the angle \u03ba will be largest when the tire slip angle is small, and vice versa. In other words, when the slip angle is small, there is a relatively large angular difference in \u03ba between no braking and full braking. When the slip angle is large, there is a relatively small difference in \u03ba between no braking and full braking. In Figure 4, full brake scenarios for slip angles of 10 and 80 degrees are depicted on the left and right, respectively"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0000301_12.2222201-Figure1-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0000301_12.2222201-Figure1-1.png",
+        "caption": "Figure 1. A single degree of freedom rotating arm is driven by antagonistically acting FAM muscle bundles. Working fluid is supplied to the actuators by an ideal constant source pressure Psys , creating mechanical output work by the robotic arm.",
+        "texts": [
+            " The FAM force-strain curve has been quantified by various analytical and empirical techniques. Equation (1) uses FAM geometry to relate actuator contraction force to applied pressure and actuator strain4. Characterization parameters include initial radius r0, initial braid angle \u03b10, applied pressure Pa, initial length l0, and actual length l. Parameters a, b, and strain \u03b5 are used to simplify the representation of the modeling equations.  2 2 0 (1 )ideal aF Pr a b   \u00a0 (1) 0 2 2 00 0 3 1 , , tan( ) sin( ) l l a l b       \u00a0\u00a0 (2) The rotating robotic limb in Figure 1 of length Larm, mass marm, and tip mass Mext is actuated by a FAM antagonistic pair. This is a useful plant for demonstration because the nonlinear gravitational torque loading of the arm couples with the decreasing force-strain behavior of a FAM. The antagonistic actuators are FAM bundles, applying torque on a pulley of mass mpulley and radius rpulley. The equation of motion for the system is given in (3), where \u03c4ext is the external torque, I is the rotational mass moment of inertia of the system, and c is the system damping due to FAM braid friction and friction losses",
+            " Two control schemes are proposed, with each handling post-transition activation levels differently. The overall feedback controller diagram in Figure 2 is divided into sections representing the plant/actuator hardware, pressure dynamics with control valves, and controller. Before addressing the switching control scheme, a controller must be created that is applicable to a single motor unit. A simplified PID and model-based control algorithm is developed16 to control applied motor unit pressure. As shown in Figure 1, position error E is input to the PID controller, outputting a representative corrective force control signal Uc in (12). KP, KI, and KD represent the proportional, integral, and derivative control gains, respectively. 0 t c P I D dE U K E K Ed K dt    (12) Known system gravitational forces Ug are calculated in (13) using the actual measured system position \u03b8act, and added to the corrective PID control signal to yield a representative desired total force, Ftot. cos( cM os() ) 2 arm arm act ext arm act g pulley L m g gL U r     (13) tot c gF U U  (14) Desired applied pressure is then calculated with (15) by using the representative total force, the desired rotational position \u03b8des of the robot arm, and an inversion of the modified FAM model"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0003063_1.428360-Figure5-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0003063_1.428360-Figure5-1.png",
+        "caption": "FIG. 5. Rotating shaft finite element.",
+        "texts": [
+            " 1, the triad XYZ is a global coordinates system with its origin at the geometrical center of the shaft left bearing, where the X axis coincides with the shaft bearing\u2019s center line in the nonworking ~zero speed! position of the system. The orientation of the de- 852awzi M. A. El-Saeidy: Rotating machinery dynamics simulation /content/terms. Download to IP: 129.101.79.200 On: Fri, 29 Aug 2014 09:52:47 Redistr flected rotor element in space ~Fig. 3! is monitored using Euler angles ~Fig. 4!. The elastic rotating shaft is discretized using a C0 four-node isoparametric beam finite element ~Fig. 5! with four degrees of freedom ~DOF! per node: two translational motions plus two total rotations. The C0 beam finite elements were first used in rotor dynamics in Ref. 15, using a weak formulation, where the authors demonstrated the excellent performance of the C0 elements compared to their counterparts of C1, reported in the literature. The author14 has presented a rotating shaft C0 finite-element model using the finite-element displacement method taking into account translational motions, rotary inertia, shear deformations, gyroscopic moments, and mass unbalance forces",
+            " x\u0304 y\u0304 z\u0304 is an auxiliary, moving frame system that initially coincides with XYZ. Euler angles are: ~1! rotation c about the X axis results in Y coinciding with y\u0304; ~2! rotation u about y\u0304 results in the moving frame coinciding with ay\u0304z% ; and ~3! spin f about the a axis results in the moving frame coinciding with frame abc. The angular velocity vector is v\u03045v\u0304aia 1v\u0304bib1v\u0304cic . Its components, expressed in the body coordinate system, are14 @v\u0304a v\u0304b v\u0304c# T5@c\u0307 cos u1f\u0307 c\u0307 sin u sin f 1 u\u0307 cos f c\u0307 sin u cos f2 u\u0307 sin f#T. ~1! Figure 5~a! shows a typical four-node beam element of length 2lk e where e and k refer to an element and its number. A generic point located in a general position within the kth element has two translational motions v , w along b, c axes, respectively, plus two total rotations ub , uc around b, c. These local deformations are related to their global counterparts dk e5@y z uy uz# T by the transformation matrix Q\u0304s , such that hk e5Q\u0304sdk e , hk e5@v w ub uc# T, Q\u0304s5FQs 02 02 Qs G , Qs5F cos Vst sin Vst 2sin Vst cos Vst G . 0\u03042 is a 232 null matrix. The shaft kth finite element has 853 J. Acoust. Soc. Am., Vol. 107, No. 2, February 2000 F ibution subject to ASA license or copyright; see http://acousticalsociety.org 16-DOF @Fig. 5~b!# and its 1631 global deformation vector, qk e , and the local counterpart, uk e , are qk e5@y j z j uy j uz j# T, uk e5@v j w j ub j uc j# T5A\u0304sqk e , j51,2,3,4. In transformation matrix, A\u0304s5diag@Q\u0304s Q\u0304s Q\u0304s Q\u0304s#; each of its off-diagonal entries is a 434 null matrix. The global deformation vector, dk e , is related to the element nodal point variables vector, qk e , such that dk e5Nqk e , N5FN j 0 0 0 0 N j 0 0 0 0 N j 0 0 0 0 N j G , j51,2,3,4. ~2! N is 4316 matrix of the element C0 shape functions where N1 , N2 , N3 , N451/16@2(12r)(123r)(113r), 9(12r) 3(123r)(11r),9(12r)(113r)(11r),2(123r)(113r) 3(11r)# ",
+            " M k e is the inertia matrix and G\u0304k e is the matrix due to gyroscopic effect, both defined in the XYZ frame. det(Jk e) is a determinant of the Jacobian matrix, Jk e . ~G\u0304k e ,M k e!5E 21 1 ~N !TS F 0\u03042 0\u03042 0\u03042 rk eVsIpk e I\u03032 GN , F rk eAk eI2 0\u03042 0\u03042 rk eIdk e I2 GN D det~Jk e!dr . I\u030325F 0 1 21 0G , I2 is a 232 identity matrix, and Ak e is the element cross sectional area. In Ref. 14, the element strain tensor ek e 5@]u/]a (]u/]b1]v/]a) (]u/]c1]w/]a)#T, where a, b, and c are the element spatial coordinates @Fig. 5~a!# and u 5u(a ,b ,c)5(cub2buc) is the axial displacement field of a point located in a general position in the cross section, from the centroidal axis, and stress vector sk e 5diag@Ek e GI k eK\u0304k e GI k eK\u0304k e#T, where (Ek e ,GI k e) are the moduli of elasticity and shear, respectively, and K\u0304k e is the cross sectional shear factor, along with the assumption of symmetrical cross section, were used to obtain the element potential energy Pk e51/2~qk e!TKk eqk e , Kk e5E 21 1 ~Bk e!TDk eBk e det~Jk e"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0001454_20140313-3-in-3024.00185-Figure2-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0001454_20140313-3-in-3024.00185-Figure2-1.png",
+        "caption": "Fig. 2. Showing the human arm and the reference angles",
+        "texts": [],
+        "surrounding_texts": [
+            "We considered several methods to calculate joint angles. The first option was to use Inverse Kinematic Algorithm. Since the manipulator is redundant, the algorithms provided analytical solutions to the equations. [1] gives the DenavitHartenberg (D-H) parameters for the human arm and also describes an optimized Inverse Kinematics Algorithm for finding out the joint angles and also suggests an optimization for the same. We computed the transformation matrices and also implemented this algorithm but found it too slow. [2] uses a Jacobian control method for redundant systems which implements a fast closed loop algorithm. The main reason that these algorithms are much slower than ours is that they assumes knowledge of only the end effector position and orientation, whereas with the Kinect Sensor, we also have the 3D global coordinates of each of the joint position, which helps us in easily finding out the joint angles. In addition to being slow, these algorithms employ a fairly complex algorithm for an otherwise simple problem. We researched some other papers to look into other techniques and know more about Inverse Kinematics and the manipulator {[3], [4], [5], [6]}. Our own algorithm derives inspiration from [7]. In this, the authors have tried to simplify the Inverse Kinematics problem for various cases when one of the joint positions is known. Now with our Kinect Sensor, we know all the joint positions, using analytical 3D geometry, we have computed all the joint angles. [8] provides all the information on Amtec PowerCube manipulator dynamics and control."
+        ]
+    },
+    {
+        "image_filename": "designv11_60_0000962_tmag.2016.2636208-Figure4-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0000962_tmag.2016.2636208-Figure4-1.png",
+        "caption": "Fig. 4. Electromechanical converter. (a) Linear. (b) Cylindrical. (c) Spherical.",
+        "texts": [
+            " (29) The material (anisotropy) force and torque component are equal to FeM = 1 2 \u222b V (\u03bdxy \u2212 \u03bdyx)Bx \u2202 By \u2202y dV (30) TeM = 1 2 \u222b V (\u03bd12 \u2212 \u03bd21)B1 \u2202 B2 \u2202x2 dV (31) and they do not vanish for regions where asymmetrical magnetic anisotropy appears, i.e., \u03bd12 = \u03bd21. Under the conditions specified in theorems, it can be written as Fe = FeL + FeM (32) Te = TeL + TeM. (33) The first and second theorems are satisfied if anisotropy (material) either force FeM or torque TeM vanishes, respectively. It results from equality M = 0, which is guaranteed under conditions (3) or (12), respectively. The first and second theorem presentations are carried out based on linear motor forces analysis (Fig. 4 and 5). Linear motor dimensions and parameters of work are a = 0.02 m (conductive layer width), l = 1 m (rotor length), g = 0.01 m (gap width), \u03b81 = 4870 A [magnetomotive force (mmf) first harmonic magnitude], f1 = 5 Hz, Y = 1 m (pair-pole length), \u03b3 = 30 \u00b7 106 S/m (carriage conductivity), \u03bdx x = 0.4\u00b7\u03bd0 (cross-layer axis reluctivity), \u03bdyy = 0.4\u00b7\u03bd0 (move direction axis reluctivity), and different reluctivities \u03bdxy , \u03bdyx (see Tables V and VI). Fig. 6(a) and (b) confirms [see Table V for the cases a) and b)] that if the condition (3) is satisfied, the first theorem thesis is fulfilled"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0003658_esej:20000302-Figure5-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0003658_esej:20000302-Figure5-1.png",
+        "caption": "Fig. 5 The windings for the translational motor (TLIM coils) and for the rotational motor (TRIM coils)",
+        "texts": [
+            " Alnscn projerl 111 tlic Alasca project a device for placing 1C:s on a priiitcd circuit board has bccn devclopcd. I t s l i o u l d replace older, difilciilt to control pnciumatic cquip~iient by an electric servo system that can rot:itc and translate simultaneously. accul-ately arid Cut. A design team of a mechanical and clrctrical cnginccr wgs Eomicd to design tlie niotoi- and its control (both students fiaoi the 'Mcchatronic Ilcsigncr' postgraduate course). An induction type of iiiotor W:IS developed witti two sets ofwitidings, one to realise the rotation ami another one to realise tlic translation (Fig. 5). To achieve the required accuracy, air bearings were used. T h i s could only be achieved by using contactless sciisors to iiicasure the two motions. The inductive sensor fur measuring the translation w x iiiore or less stnndaiil, although care lind to be taken due to the magiictic fields of the motor. A contactless rotational s e i i s w was developed that could accurately nicasurc tlic rotation even wlieii the actuator peiforiiis trauslatiotial motions. The setisor coiisists of a coiiihiiiatioii of four LEDs at the stator, a slicet of polarising material at tlie rotor/traiislator and four photodiodes at the stator covered with sheets of polarising material whose directions of polarisation arc spread over 120\u00b0 (Fig"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0002210_10402009608983594-Figure1-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0002210_10402009608983594-Figure1-1.png",
+        "caption": "Fig. 1-Hemispherical porous bearing configuration.",
+        "texts": [
+            " The oil film pressure is solved and, then, applied to predict the load capacity and the thickness-time relation. A comparison bctwccn the results derived by the BM, DM and SFM is made to show the influence of the viscous shear effects on the sqitcczc film performance. As the value of the permeability parameter approaches zero, the bearing characteristics predicted using the BM approach that of a solid bearing, thus providing a support to the present theoretical study. ANALYSIS 'The physical configuration of a hemispherical porous bearing is shown as in Fig. 1. Assuming that there is n o variation of pressure across the fluid film, n o external forces are acting on the film, the fluid inertia is small compared to the viscous shear, an incompressible Newtonian lubricant having constant properties, the porous bearing material is homogeneous and isotropic, the flow of lubricant in the bearing is laminar, and the pressure distribution within the porous medium is a linear function in the radial direction, then the reduced Navier-Stokes equations and the BM equations are given in Cartesian coordinates by the following: where )3"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0001916_1.555403-Figure1-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0001916_1.555403-Figure1-1.png",
+        "caption": "Fig. 1 Motion of contact lens during blinking",
+        "texts": [
+            " In many regards a contact lens\u2019 behavior is akin to that of a slider bearing: the lens itself represents the slider, the eye plays the role of the stationary pad, and the tear film is the lubricant. The motion of the contact lens on the cornea and its proper functionality is greatly affected by the blinking action. Hence, it is crucial to investigate the influence of the blinking action in designing soft contact lenses. The interaction of the eyelid with a contact lens during the event of blinking may be described as follows ~cf. Hayashi @1#, Mandell @2#!. First the upper lid descends until it comes into contact with the upper edge of the contact lens and pushes it downward ~Fig. 1(a1)!. The lens tilts ~Fig. 1(a2)! and, then, slides over the eye surface and into the lower corneal position until its lower edge slides into the lower limbus about 2 mm and it comes to rest ~Fig. 1~b!!. At the very beginning of the sliding motion ~when the lens is on the central zone of the eye!, due to a close match of curvatures of the lens and the cornea, the lens moves parallel to the eye. With the tear fluid acting as the lubricant between the lens and the eye surface, the lens behaves akin to a slider bearing. The next stage of the blinking process involves the upper lid covering the lens ~Fig. 1~c!!. After about 0.05 seconds from the completion of the closure, the upper lid starts to open ~Fig. 1~d!!. The eye ball, then, undergoes a rotational motion\u2014what is commonly referred to as the Bell\u2019s phenomenon\u2014whereby the upper lid pull the lens upward, thus positioning the lens at a slightly higher location than its resting spot. When the upper lid sweeps around the center of the lens, the lens is squeezed downward ~Fig. 1~e!!, and centered ~Fig. 1~f!! by the centering force caused by meniscus deformation. During this interval, the contact lens behavior is akin to that of a squeeze-film bearing. Research performed by Hayashi @1# represents a very comprehensive study of the motion of the hard contact lens. He modeled a hard lens as a parallel slider bearing with a peripheral curve. It was assumed that the slider is a rigid surface and infinitely wide perpendicular to the direction of motion. An analytical derivation for pressure distribution was presented showing a skewsymmetric profile for the pressure and the velocity profiles in the lubricating film",
+            "1 Pressure Distribution Before Tilting. Figure 3 shows the dimensionless pressure distribution with various aspect ratios for l50.2 mm3/N before tilting takes place. A notable trend is the occurrence of positive pressure build up in the leading edge and a negative pressure in the trailing edge for all aspect ratios L. In reality, region R3 and a part of region R2 extending from region R3 experiences P>0 and the rest of the part of region R2 and region R1 experiences P<0 shortly after the upper lid meets the lens in Fig. 1(a1). It is shown that as L increases ~in the range of 3 to 9! the absolute value of the peak pressures decrease at the junctions of the region R1 and R2 and at the junction of the region R2 and R3 . Each pressure distribution crosses P50 at X50.4. It is also interesting to note that the pressure distribution in the peripheral regions R1 and R3 is concave for large values of L~L>7!. This trend\u2014which is not typically seen in hydrodynamic bearings\u2014occurs because of the relatively sharp inclination angle a at the peripheral edges"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0000524_978-4-431-55013-6-Figure8.2-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0000524_978-4-431-55013-6-Figure8.2-1.png",
+        "caption": "Fig. 8.2 Bisection method to find the stable manifold. The two initial points are given. When each point is mapped (shown as 0, 1, and 2), the midpoint and the initial point 2 approach each other. The initial point 1 and the midpoint are more proximately located on either side of the stable manifold than the two initial points",
+        "texts": [
+            " To apply the external force control to stabilize the unstable quasi-periodic orbit, it is necessary to determine the orbit itself in advance. If we find the stable manifold, we would be able to determine the unstable quasi-periodic orbit because the orbit is stable on it. In general, however, it is difficult to determine the stable manifold analytically. Fortunately, a method that numerically estimates the orbit on the stable manifold has been proposed based on a bisection method [6]. We consider two initial points on either side of the stable manifold and their midpoint (Fig. 8.2).When each point is mapped by the system equation, themidpoint and the initial point 2 approach each other, whereas the initial point 1 is separated from them. Therefore, by observing the dynamics, we can identify the side on which the midpoint is located relative to the stable manifold. We replace the initial points with themidpoint and the initial point 1,which aremore proximately located on either side of the stable manifold than the two initial points. These two points approximate a point on the stable manifold with arbitrary precision by iterating this process"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0003173_s0141-6359(02)00117-4-Figure20-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0003173_s0141-6359(02)00117-4-Figure20-1.png",
+        "caption": "Fig. 20. Radial coupling concept with self-locking axially tapered adjustment compliance.",
+        "texts": [
+            " Analytical models of the coupling designs agreed with experimental results for stiffness to within 20%. Our finite element models were within 15% of measured values. Thus, the models and design metrics are useful for designing the couplings. Further work is planned within this family of couplings. We will pursue a more accurate bending model and also plan to investigate the possibility of a radial version of this new type of coupling. The concept of the radial version of this coupling is as shown in Fig. 20. In addition it may be possible to create these types of features on silicon wafer, so that the accuracy in assembly of MEMS devices can be increased. This work started as a team project for the course 2.75\u2014 Precision Machine Design, taught by Prof. Alex Slocum at MIT. Heather Dunn was part of the team during the course project and participated in the initial concept development. We appreciate her valuable initial contribution to this project. We would like to thank Dr. Zafar Shaikh and the Ford Motor Company for the three-dimensional printed parts"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0001256_jahs.60.042007-Figure5-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0001256_jahs.60.042007-Figure5-1.png",
+        "caption": "Fig. 5. Tregold\u2019s approximation.",
+        "texts": [
+            " 4, the circumferential position of contact point over one mesh cycle, \u03b8k(t), is simply expressed as \u03b8k(t) = { 2t + \u03b8r 0 \u2264 t \u2264 crT0 no contact crT0 < t \u2264 T0ceil(cr) (9) where the angle, \u03b8r , is some initial phase value. Without loss of generality, \u03b8r is set as zero for simplicity. In this paper, the contact ratio, cr, is calculated via Tregold\u2019s approximation (Ref. 31) to avoid extensive numerical computations. The idea is to imagine a formative spur gear instead of the face gear in mesh with the spur pinion (Refs. 1 and 29). This then transforms the face-gear drive into a formative spur gear drive. This method is a common practice used in bevel gearing and is illustrated in Fig. 5. The number of teeth in the formative spur gear pair, Nv1 and Nv2, are obtained as (Ref. 31) Nv1 = N1/ cos \u03b31, Nv2 = N2/ cos \u03b32 (10) where the apex semiangles of pinion and face-gear pitch cones, \u03b31 and \u03b32, are (Ref. 29) \u03b31 = arc cot ( m12 + cos \u03b3 sin \u03b3 ) , \u03b32 = arc cot ( m21 + cos \u03b3 sin \u03b3 ) (11) Here N1 and N2 are the number of pinion and face-gear teeth, respectively, and m12 = N2/N1 and m21 = 1/m12 are gear ratios. The contact ratio of the formative spur gear pair is used to approximate the face-gear drive contact ratio"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0002671_bf01282292-Figure3-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0002671_bf01282292-Figure3-1.png",
+        "caption": "Fig. 3 a, b. Swimming trajectory of P. minimum, a Plots for middle points of anterior and posterior end of swimming P. minimum. The cell began to swim smoothly (A), stopped to change the direction (B), resumed swimming (C), and stopped to change the direction again (D). Time durations from point A to B, B to C, and C to D were 1696 ms, 416 ms and 2064 ms, respectively, b Sketched swimming path of P. minimum. Pp pitch of helical swimming trajectory; Rp radius of helical swimming trajectory; Fpath angle of the cellular swimming trajectory against the axis; O the pitch angle of the cell against the axis of the trajectory",
+        "texts": [
+            " 2 a and b, as they were not clearly visible. The cellular length along the antero posterior axis, the width in the suture plane, and the height perpendicular to the suture plane are given in Table 1 with the equivalent spherical diameter (ESD), which is a geometrical average of the length, width, and height of the cell, ~ , and corresponds to the diameter of a sphere whose volume is equivalent to that of P. milffmum cell. While cells of P. minimum swam uniformly, they changed their swimming direction from time to time. Figure 3 a shows a typical swimming trajectory of P. minimum, in which the middle points of the anterior and posterior ends of the swimming cell were plotted at 16 ms time intervals and connected sequentially. The cell swam smoothly (A-B in Fig. 3 a), then changed its direction (B-C) and finally resumed swimming (C-D). Time durations from A to B, B to C, and C to D were t696 ms, 416 ms, and 2064 ms, respectively. The smooth swimming path (A-B and C-D in Fig. 3 a) was a right-handed helix. The entangled line from point B to C shows that the cell swam more slowly in turning motion than in helical swimming. Figure 3 b shows a sketch of a cell swimming toward the right with its anterior end forward, its antero-posterior axis along the helical trajectory, and one of its valval sides faced toward the axis of helical trajectory. The pitch, radius, path angle, F = tan -1 ( 2 7 ~ R p / P p ) , o f t h e helical trajectory and pitch angle (Fig. 3 b), i.e., the angle between the cellular anteroposterior axis and the axis of the helical trajectory, are shown in Table 1. Because the cellular antero-posterior axis was inclined along the swimming trajectory, the pitch angle of the cell was not significantly different from the path angle (Table 1). The cell was swimming along the cellular antero-posterior axis and rotating around the axis in a right-handed direction to the cellular swimming direction (Fig. 2 a). Revolution around the axis of the trajectory almost synchronized with that of the cellular rotation so that the same side of the cell faced the axis of the trajectory",
+            " Here, we termed this pattern of the transverse flagellum \"bi- form\" helix, since the helix is not uniform but alternates two half pitches, and described the \"bi- form\" transverse flagellum by p, which is the ratio of pout to the pitch of a whole period of helix (Pout + pin) . When the helix is uniform, p is 0.5, p decreases to zero when the outer parts are parallel to the antero-posterior axis of the cell, and increases to one when the inner part are parallel to the cell antero-posterior axis. The radius, frequency, and wave number of the helical wave, Pout, Pin, and p (Fig. 2 b) are shown in Table 1. p was 0.33 + 0.05 (n = 7) with a range from 0.27 to 0.41. Cells swimming forward stopped and changed their swimming direction occasionally (Fig. 3 a). Figure 5 a and b shows sequential images of a typical turning motion of P. minimum. Figure 5 a shows a cell stopping to change the swimming direction and Fig. 5 b shows the cell resuming its swimming. The time interval between Fig. 5 a xvii and b i was 184 ms. During this time interval the cell seemed to continue the same motion as in Fig. 5 a xiii-xvii . The sequential motion is described below. 1. Turning was preceded by a specific wave (Fig. 5 a iii-xi). As the wave propagated, the longitudinal flagellum was thrust forward gradually and started pointing forward and the direction of the cellular antero-posterior axis was directed upward (Fig",
+            " 5 a vi-xvii). 2. The longitudinal flagellum beat with a planar undulating wave (Fig. 5 a xiii-xvii) from the base to the tip, and the cell moved backward. The plane in which the longitudinal-flagellum beat was perpendicular to the valval sutural plane, similarly to that in forward swimming (Figs. 2 and 4). The transverse flagellum still remained quiescent. The backward swimming motion continued for some time. This was 184 ms between Fig. 5 a xvii and b i. The backward motion corresponded to B-C in Fig. 3 a. 3. The transverse flagellum resumed beating with the ordinary helical wave (Fig. 5 b i-iv) and the cell began to swim forward (Fig. 5 b vi). The longitudinal flagellum gradually directed backward in undulating motion (Fig. 5 b iii-xii). When the cell resumed swimming, no specific wave thrusting the longitudinal flagellum backward was observed (Fig. 5 b vi-xii), contrary to the beginning of turning motion (Fig. 5 a iii-xi). D i s c u s s i o n In the present study, we obtained continuous still images of flagella in vivo (Fig"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0002650_pola.1991.080291005-Figure2-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0002650_pola.1991.080291005-Figure2-1.png",
+        "caption": "Figure 2. cycle, (b ) cooling cycle. DSC thermograms for PMS-c: (a ) heating",
+        "texts": [
+            " Observations of this polymer on a polarizing microscope revealed the formation of a black-and-white Schlieren-type texture at 62\u00b0C on slow cooling from the isotropic phase, but this texture was not unequivocal evidence for the presence of a liquid crystalline phase. Therefore, the mesophase formation ability of this side chain was evaluated by attaching it to a polymethylsiloxane, PMS, backbone, polymer PMS-c. Polymer PMS-c exhibited a T, at 126\"C, and an isotropization transition, Ti, at 142\u00b0C as shown in Table IV and Figure 2. On cooling from the isotropic state, a bitonnet texture appeared at 151\u00b0C and a fan-type texture formed at 140\u00b0C. This type of textural behavior is typical of the smectic A phase.17 In most cases, the enthalpy of isotropization, AHi, of an SCLCP generally falls in the range of 0.1-0.9 kJ/ mol for the nematic phase and 1.6-7.5 kJ/mol for Table IV. Phase Transitions of SCLCPs Transition Temperature\" Polymers (\"0 Texture AH: (kJ/mol) PI-a PI-b PI-c PMS-c PI-d PI-e PII-e PIII-e PMS-e PI-f PII-f PMS-f G 19 I G 22 I G-10 K 73 I G-2 K 126 S 142 I G-1 K 94 (N96)d (K85 N 90 I) G-12 K 106 (N 104)d (K96 N 101 I)f G-14 K 116 N 120 I G-13 K 120 N 127 I G-12 K 152 (S1151)d Sz179 I G-12 K198 K2("
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0001076_indicon.2016.7838972-Figure1-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0001076_indicon.2016.7838972-Figure1-1.png",
+        "caption": "Fig. 1. Parts of a ball bearing",
+        "texts": [
+            " INTRODUCTION Induction motors are workhorse of the industry due to its wide range of applications. They are generally exposed to harsh industrial conditions which on prolonged exposure induce various defects in the motor. Studies have reported that majority of these defects in induction motors are caused due to damage in rolling bearings which shares approximately 40% in total faults [1]. Bearing faults are caused due to improper lubrication, installation, contamination of lubricant. The typical ball bearing is shown in Fig. 1 on which defects may occur in its outer race, inner race, cage and balls. Thus bearing faults are named as: 1. Outer Race Fault, 2. Inner Race Fault, 3. Cage Fault, and 4. Ball Defect. These faults may result in excessive vibrations in the machine which leads to system function incorrectly. These damages also impose heavy repairing, maintenance and replacement charges. Therefore, it is important to monitor, analyze and diagnose incipient bearing faults of the induction motors for its better operation"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0001948_5326.760575-Figure10-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0001948_5326.760575-Figure10-1.png",
+        "caption": "Fig. 10. Bucket on the estimated plane.",
+        "texts": [
+            " Therefore, r is another variable to be solved to determine d1; 2; and 3: With the link assignment, the kinematic equations are obtained as follows: xd = sin 3(l3 cos sin r + sin (d6 cos cos r)) + cos 3( l3 cos cos r d6 sin + a4) + l2 sin 3 + h2 (2) yd = sin 2(cos 3(l3 cos sin r + sin (d6 cos cos r)) sin 3( l3 cos cos r d6 sin + a4) + l2 cos 3) + cos 2(cos (d6 cos l3 sin cos r) l3 sin sin r) (3) zd = cos 2(cos 3(l3 cos sin r + sin (d6 cos l3 sin cos r)) sin 3( l3 cos cos r d6 sin + a4) + l2 cos 3) + sin 2(cos (d6 cos l3 sin cos r) l3 sin sin r) + d1: (4) r gives redundancy in the kinematics of the reclaimer as indicated in the above equation. Therefore, the number of the equation is fewer than that of the variables to be determined. The circles in Fig. 10 are made to be the trajectories of the rotating buckets, which means that the way that the buckets can contact with the landing point exist infinitely. But the dotted circles shows cases where the buckets are deeply embedded in the pile and they already collided with pile. Only the solid circle is desirable and r is uniquely determined in this case. It gives us a constraint equation. It is derived from the fact that the circle should meet the surface of the pile at the one point. In general, the piles have irregular curved surfaces and it is very hard to extract the mathematical model of the surface"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0003668_isie.2001.931828-Figure4-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0003668_isie.2001.931828-Figure4-1.png",
+        "caption": "Fig. 4 The robots leave the unit-center in the environment with obstacles",
+        "texts": [
+            " But the previous work has several demerits. For example, in Fig. 3, both obstacles produce repulsive forces(Fomt 1 , F ~ s r z ) , respectively, which make & t i and 60brtz . These directions are combined with attractive direction 0 - b ~ . Therefore, the robot can not pass through between the obstacles. Such a mistake is due to no obstacle gap. Therefore, in this case, the strategy can let a robot leave in any direction to dynamic environment which moving obstacles. According to previous works, in situation shown as Fig. 4, the robot has no choice to deviate from the center of formation, because the previous work proposed the behavior control method that are only used in limited cases and may be fall into a local minimum or maximum. To cope with the uncertain environment and achieve the real-time performance, fuzzy logic and neural network controller have been implemented as a useful means combing fuzzy logic and neural network. 2.2. Fuzzy and Neural Network Neural networks and fuzzy systems represent two distinct methodologies that can deal with uncertainty"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0002796_l90-034-Figure10-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0002796_l90-034-Figure10-1.png",
+        "caption": "FIG. 10. The eigenvalue parabola for the conservative problem.",
+        "texts": [
+            " His next paper on the subject was Leipholz (1972c), one of only two papers that he published in the Journal of Applied Mechanics. Here he analysed a generic problem which is either of divergence type (must diverge) or may diverge, using Galerkin's method and two coordinate functions. He showed that if it diverges, it will do so at a higher load than the corresponding conservative problem. The argument is very simple. With two Galerkin functions, the eigenvalue problem for the conservative system has the form where q is the load parameter. The parabola f(X) is shown in Fig. 10; it has two positive real roots, Xy and g. As q increases, the parabola and its two roots move to the left, and at the critical q = q,, Xy becomes zero. The system diverges. The corresponding equation for the follower-force system is found to be Now there are two cases. The system may diverge or flutter. It will flutter, i.e., X I and X2 will be equal, if k2 = aT14 - a2. Thus, if k2 < a114 - a2 for all q, then the system can only diverge. Figure 11 shows that X I will become zero for a higher q than that for which Xy becomes zero, thus Leipholz applied this analysis to the divergence-type systems of Fig"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0001594_bf00411970-Figure2-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0001594_bf00411970-Figure2-1.png",
+        "caption": "Fig. 2. One of t he coils a f t e r t i l t i n g i t s \" p l a n e \" a b o u t t h e l ine A A t h r o u g h t h e cen t re .",
+        "texts": [
+            " It will further be clear that, owing to the absence of forces, the centres of the successive coils are not shifted with respect to each other. They all coincide witch the centre of the initial spiral, and the change in position of the \"plane\" of each coil only consists of a tilting about a straight line through the centre. After this tilting phenomenon has taken place, such a coil situated originally in the xy-plane will approximately lie in a plane intersecting the xy-plane under the (small) angle 9, as indicated in fig. 2. The line of intersection AA of this plane with the xy-plane passes through the centre 0 and includes the angle ~v with the y-axis. Hence it follows that the angle between the normal ON to the \"plane\" of the coil and the z-axis is also equal to 9, while the plane NOB passing through this normal and the z-axis includes the angle ~v with the x-axis. In consequence of the load which at the outer end of the spring wire is given by the moment vectors M~, My and M z along the axes of x, y, z the following moments must, as a first approximation, be transmitted by the coil about the axis AA: M A = M x sin ~ - - My cos ~v, ] about the ax i sBB: M B = M s cos g - / My sin ~v-+-M~0, } (1) about the axis ON: M N = M~"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0000786_ijbic.2016.078664-Figure2-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0000786_ijbic.2016.078664-Figure2-1.png",
+        "caption": "Figure 2 Missile statement in pitch plane",
+        "texts": [
+            " The outputs of the guidance produce proper commands and autopilot applies a computed control signal to enforce the system to track desired commands. The actuator force the physical control input to track the commanded control input. The missile dynamics respond to the control inputs. The aim of the flight control system is to force the missile dynamics to perform a desired mission. The missile dynamics are basic equations of motion. Assuming that missile motions are vertically constrained, the equations of motion that govern the missile dynamics can be achieved more easily. Figure 2 shows the missile flying in the vertical axis. The angle between the inertial reference axis and the velocity vector of the missile is named the flight-path angle (\u03b3). The angle of attack (AOA) \u03b1 is the angle from the velocity vector to the missile centreline. The angle from the inertial reference to the missile centreline is named the pitch angle (\u03b8). AZ is the acceleration in the normal direction. where \u03b8 is the angular acceleration and J is the moment of inertia. M(\u03b1, \u03b4) is the moment applied to the missile body"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0002678_1.2829318-Figure8-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0002678_1.2829318-Figure8-1.png",
+        "caption": "Fig. 8 Flank numbering",
+        "texts": [
+            " Each arc is developed in space with a screw motion and forms one of the flanks of the screw (see Fig. 4). On the contrary, we can consider that the base of each flank is the corresponding arc. The vertexes of the profiles of the central screw define the helical lines. The part of the cylindrical external side of the central screw that comes into contact with the satellite screws, and the segments of hehcal line, that are also common to the satellite screws determine the seal curves on the satellite screws (see Fig. 8). As well, the seal curves on the central screw can be determined by considering the vertexes of the satellite screws. The pump case and the parts of the screw surfaces between the exhaust zone and the seal curves define the chambers (see Fig. 5) in which the pressure can be assumed constant and equal to the value of delivery pressure. The parts of screw helical surfaces that are walls of the chambers are called sectors of surface. Finally we define as domain the set of surface parameters corresponding to a sector of surface (see Fig",
+            " Once the geometry is given, force and moment can eventually be considered as functions 'F(ip) and Moiip) of the rotation tp. For the sake of brevity the procedure is illustrated on two flanks of one of the satellite screws and on the corresponding sectors of surface. Therefore the method can be generalized for the other flanks and for the central rotor. Calculation of the Loads For convenience when we consider the left satellite screw, we will number the flanks which compose the surface of the screw, starting from a reference position, as shown in Fig. 8. Figure 9 shows some images of the left satellite screw for increasing values of the rotation angle. Notice that, since the screw is double-threaded, the same configuration is repeated with period IT. Therefore, it is sufficient to consider what hap pens to flank 2 of Fig. 8 in the interval [0, TT], to know that it corresponds to what happens to flank 7 in the interval [TT, ITT] . The shaded zones in Fig. 9 represent the sectors of surface of the relative flanks. We will analyze what happens for the sectors of surface that are formed by flanks 2 and 3. We will begin by considering the surface which represents flank 2. It is necessary to begin from the parametric expression of the arc at the base of the same flank on plane xy, which is given by the following equation: Journal of Mechanical Design DECEMBER 1998, Vol",
+            " Now it is necessary to determine the sector of surface on the flank. Starting from the value <\u0302 = 0, it is possible to observe from Fig. 9 that the upper part of the sector of surface is bounded by the top of the screw towards the exhaust corresponding to the value of the parameter \u00ab = 27r; other two bounds are the extreme values of the flank corresponding to 1? = 0 and \u00a7 = 2 (7 - /S). The last of the bounds is the seal relative to the cylindrical part of the central screw. This segment belongs to the line r2 (see Fig. 8) that is represented by the following equation: r2: JT, = 0 (16) If ip varies, the bounds of the surface sector remain the same, even if the segment of the seal curve changes its position in space. In order to specify the segment it is necessary to put the components of vector R2, given by (15), equal to those of r2 (16). - ( 2 r - Te) &in {-d - u - if + P - y) - 2r = - r , , (2r - r j cos (1? - M - (/) + ,5 - 7) = 0 au = au. (17) The third equation of (17) is an identity, while the solutions of other two are respectively: ^p + p \u2014 y = \u00b1 Ik-n -d \u2014 u \u2014 ip + p \u2014 y = \u2014 \u00b1 kn",
+            " The solution of those minimization problems is too far from the aim of the present paper, a general approximate solution can however be suggested by considering the square function of the pitch in some compo nents of the moment, from which follows that a choice charac terized by short pitch and large radiuses is preferable. The calculation of the loads on the other flanks of the satellite screw, particularly those which have as base an arc of trochoid, presents many more difficulties. In that case, the seal curves belong to lines such as the Fa of Fig. 8 and the transcendental equations such as (17) do not have an immediate solution, in contrast to what has been shown previously for flank 2. If we consider for example flank 3 of Fig. 8, its parametric expression in homogeneous coordinates can be determined by following the same procedure of previous formulas (11 ) - (15 ) : than a pitch and the same dependence of force and moment on the angular pitch. The sum of the values calculated for the components of force and moment on all the sectors of surface considered, gives the values of the loads as function of the rotation angle ip, i.e., F((p) and Mo(ip). The integration operations made here, thanks to the suitable choice of the surface parametric forms and to the regularity of the surfaces, guarantee that the resulting vectors are those corresponding to the actual force and moment vector acting on the rotors in magnitude and direction"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0003868_6.2002-1460-Figure5-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0003868_6.2002-1460-Figure5-1.png",
+        "caption": "Fig. 5 An annular membrane",
+        "texts": [
+            " The four corners of the membrane are cut flat in the numerical analysis. In the finite element discretization, 264 constant strain triangular elements are used and the symmetry of boundary condition is also considered. The wrinkle pattern and distribution shown in Fig. 4 are in good agreement with the experimental result shown in Fig. 3. Example 3: An annular membrane An annular membrane of outer radius = 6 and inner radius r = 3 is subject to a pair of normal force and tangential force along the inner edge, see Fig. 5. A 192-element mesh of the membrane and predicted stress distribution (under = 0.01 and = 0.01) is shown in Fig.6. Also, the evolution of the wrinkle pattern of the membrane under a changing tangential load is presented in Figs. 7 (a) to (d). or i p \u03c4 p \u03c4 \u03c4 To better illustrate wrinkle patterns, here a wrinkle line drawing technique is developed. Here, a wrinkle line is defined as a trajectory of the first principal stress \u03c3 that continuously crosses a number of wrinkled elements. For elements of constant strains, the segment of the curve within each element is a straight line with its orientation in the direction of \u03c3 of the element"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0003035_dftvs.1994.630030-Figure2-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0003035_dftvs.1994.630030-Figure2-1.png",
+        "caption": "Figure 2: The 1; track model in (b) 3D.",
+        "texts": [
+            " 2: The 3D mesh architecture can be considered as layers (in the zy plane) of 2D meshes stacked in the z direction. A 3 x 3 x 3 mesh is shown in Figure 1. Each P E is connected to its six nearest neighbors. The six neighbors are in the North (N), South (S), East (E), West (W), Z+ and Z- directions. A 1; track, 3D mesh model 194 1063-6722/94 $4.00 0 1994 IEEE Reconfiguration Techniques 195 To augment a 3D mesh into the 3D I f track model, a six port switch is inserted between a PE and each of its six neighbors. The entire 3D array is covered in each of the six sides by a 2D n x n layer of spare PE\u2019s. Figure 2(a) shows a 2 x 2 x 2 mesh with the corresponding switches in the 1; track model. Three types of switches are shown depending on where they are located: an SV switch is located between two PE\u2019s that are N and S neighbors, an S H switch is located between two PE\u2019s that are E and W neighbors and an SZ switch is located between two PE\u2019s that are neighhors in the Z direction. In addition to being connected to the two PE\u2019s that a switch lies between, an SV switch is connected to four other SV switches. Namely, its four SV neighboring switches in the yz plane. Similarly, an SZ switch is connected to 4 other SZ switches and an SH switch is connected to 4 other SH switches. Given an n x n x n 3 0 mesh, six 2D n x n meshes of spares are added such that one is added to each side (face) of the 3D mesh. In this model which is shown in Figure 2(b), 6n2 spares are added to the n3 non-spare PE\u2019s. 196 International Workshop on Defect and Fault Tolerance in VLSI Systems The connection in a switch of port i to port j will be denoted by C<i, j> where C is a connection operator, { i , j } E {N, S, E, W, Z+, Z-} and i + j . Obviously the operator C i s symmetric, that is C<i,j> + C<j, i>. The state of a switch will be denoted by the set of ports that are connected together. For example the state in which N and W are connected and E and S are connected will be denoted by {C<N, W>,C<E,S>} "
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0000380_1464419316631862-Figure5-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0000380_1464419316631862-Figure5-1.png",
+        "caption": "Figure 5. RX 130L robotic manipulator in the reference configuration q1 \u00bc q2 \u00bc q3 \u00bc 0.",
+        "texts": [
+            " At the switching points, the total energy jumps whereas it is preserved otherwise (Figure 4(b)). Notice that the geometric locking constraint is exactly satisfied because the transition conditions (32) yields _q2 \u00bc 0, which is exactly integrated to q2 \u00bc q2 t1\u00f0 \u00de, for t5t1, and analogously for q3. Comparing Figures 2(a) and 3(a) the difference of the obtained joint trajectories is apparent. Joint locking in a 6 DOF serial robot The second example is an industrial 6 DOF serial robot Sta\u0308ubli RX130L (Figure 5). All geometric and inertia parameter were identified and used in the dynamic model. The serial robot can be split into a wrist (the last three revolute joints) and the serial chain, consisting of the first three revolute joints, which achieves the positioning of the wrist. Since wrist locking has no significant effect on the overall motion, a simplified scenario is considered where joint 1 is locked at t\u00bc 0.05 s, then joint 2 is locked at t\u00bc 0.1 s, and finally joint 3 is locked at t \u00bc 0:15 s. In the following figures the motor angles and rates are shown that are related to the joint angles and rates by a gear ration of 100"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0000602_s11771-015-2876-0-Figure11-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0000602_s11771-015-2876-0-Figure11-1.png",
+        "caption": "Fig. 11 Three types of bearing cage: (a) Cylindrical pocket cage; (b) Spherical pocket cage I; (c) Spherical pocket cage II",
+        "texts": [
+            " It is suggested that for the oil\u2212air lubrication method, or any air as transporting material lubrication method, the lubrication device can be mounted at either A or B side at low rotating speed. However, when the shaft rotates at high speed, it is better to mount the lubrication device at the B side in order to diminish the barricading effect from the air vortex and to increase the lubrication efficiency. 4.4.1 Structure of bearing cage To investigate the influence of internal structure parameters on the air phase flow pattern, three kinds of bearing cages are used in the first set of model. Figure 11 shows the specific structure of the three types of bearing. They are cylindrical pockets cage, spherical pockets cage I, and spherical pockets cage II. The difference between the two types of spherical cage is that the type II has been trimmed on the cage ring. Figure 12 shows the temperature contour on the central radial plane for different cage types bearing at shaft rotating speed of 3\u00d7104 r/min. It could be seen that when the structure of the cage changes, the temperature distribution changes correspondingly"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0001948_5326.760575-Figure13-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0001948_5326.760575-Figure13-1.png",
+        "caption": "Fig. 13. Graphic display of the yard.",
+        "texts": [
+            " pi becomes an element of the candidate set for the landing points S: \u2022 Obtain i which is an angle between the landing point and slew limit. o) ELSE \u2022 the point pi is discarded. It can not be used as a candidate point for landing on the pile. 3) ENDFOR i 4) Obtain max = maxf ijpi 2 Sg for the pi: 5) The point pi for the max is chosen as the optimal landing point. It has the widest slew angle, which is the same operation result as for the well-experienced worker. The algorithm for extracting the optimal landing point is summarized as follows. The obtained optimal landing point that satisfies Fig. 13 shows the graphical display for the pile and reclaimer. The scanned pile data from the range finder are reconstructed for the workers to be able to see the shape of the pile by the 3-D graphics. The reclaimer is drawn by animation techniques and the display of its on-line configuration is performed using the position data of the sensors in the joints. The contour line illustrates the reclaiming height of the pile and the optimal landing point and slew limit are extracted from the above algorithm"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0002033_(sici)1097-4563(199710)14:10<729::aid-rob3>3.0.co;2-w-Figure9-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0002033_(sici)1097-4563(199710)14:10<729::aid-rob3>3.0.co;2-w-Figure9-1.png",
+        "caption": "Figure 9. Mobile manipulator. J 5FJrp o Jhp Jhm G . (16)",
+        "texts": [
+            " Now, denoting the error of Ur for the desired trajectory as Ure , and assuming Ur is small, we obtain the error in the joint coordinate space as follows,6. PUSHING OPERATION WITH A MOBILE MANIPULATOR qe Q J*Ure , (17)An approach to realizing the pushing operation with a mobile manipulator, that is, a mobile robot with a q\u0307e 5 J*Ur\u0307e (18) manipulator, is discussed in this section. We assume the mobile manipulator to be a mobile robot with wheretwo independent driving wheels on which a threedimensional manipulator with n degree-of-freedom (DOF) is mounted (see Figure 9). J* 5F J21 rp o 2J1 hmJhpJ21 rp J1 hm G (19) We adopt a proportional and differential feedback control rule for the arm tip of the manipulator. As for the mobile robot part, we use a feedback control method proposed in reference 14. In this method the mobile robot is pulled by the reference point fixed in front of the mobile robot. Using this method, although the orientation of the mobile robot cannot be controlled, the consideration of the nonholonomic constraint can be removed. A unified control method for the mobile manipulator combining these two control rules is decribed below. As in Figure 9, we denote the position of the reference point expressed in the universal frame SU as Upr , the position of the arm tip as Uph , the angular displacement vector of the right and left wheels as qp 5 [qr , ql]T, and the joint vector of the manipulator Figure 10. Reference point of mobile manipulator. and p1 is the pseudo inverse matrix of p. Based on the arm tip on the XOYO plane when the manipulability15 is maximum (see Figure 10). Then the desiredthis error information, we give the control rule for the mobile manipulator as follows: trajectory of the reference point is obtained from that of the object"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0002923_978-94-011-3514-6-Figure13.5-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0002923_978-94-011-3514-6-Figure13.5-1.png",
+        "caption": "Fig. 13.5 Trajectories of the mass corresponding to the two real natural modes of the system of figure 13.1 without damping",
+        "texts": [
+            "2 are written o and the initial conditions { xj{O) X2 (0) Xo - 1.883 Xo The equations of motion are then Xl (0) X2 (0) cos (2:1.27 t) X2/ XO - 1.883 cos (23.27 t) The trajectory is a segment of the straight line making an angle 92 with the axis Xl Arc tg(-1.881) - 62.03 \u2022 o o The theoretical considerations of paragraph 13.3 have shown that the orthogonality of the modes imp] ies, in this particular case, the (13.42 ) (13.4:: ) - 279 - perpendicularity of the modal straight lines. This i.s co\"fitliled by the numerical results above, which are shown in figure 13.5. In this chapter, we have s1-udied the \\\"'havi ,)lIr \"f it coplanar system consisting of a point lOaf,S attached to any number of springs and viscous resistances (di screte 1 inear elements). Such\" system has two degrees of freedom, referred t.o by x 1 and xl . The 1\",\u00b7\";,,113 of t.his study can be summarized and commented upon as follows. When the system is dissipative, the two natural mndes, known QS complex modes, correspond to the trajectori<.';:; of the mass whi,;h are two elliptical spirals"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0003645_i2002-10009-1-Figure5-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0003645_i2002-10009-1-Figure5-1.png",
+        "caption": "Fig. 5. Schematic drawing of the splitting of the central point defect and of the radius of disclination in a TGBA Robinson spherulite.",
+        "texts": [
+            " Therefore the defects classes C\u03bb (respectively C\u03c7) carry each of them only one type of singularity, viz. a singularity of the \u03c7 field (respectively the \u03bb field). The class C\u03c4 carries \u03c7 and \u03bb singularities. We therefore advance the idea that the k = 1 terminations are linked by a short C\u03bb line segment, (size \u223c= p), and a long C\u03bb loop segment. The global geometry of the distorted \u03c7 field opposes the presence of a C\u03c7). The \u03bb-director is uniformly oriented inside the loop (perpendicular to the smectic layers, Fig. 5) which is slightly elongated in the direction of this director, in order to facilitate a \u03bb non-singular geometry in the core, i.e. favoring a wedge C\u03bb line vs. a twist C\u03bb line. The extremities of the loop are therefore locations of stronger singularities, and are preferred locations for the short segment. These singularities are often observed. It has been observed [9] that in thick free-standing films prepared in square holes (thickness of the film \u223c= 5\u00b5m; square edge \u223c= 42 \u00b5m), the TGBA phase grows from the homeotropic SmA phase, when the temperature increases, in the shape of a regular texture with the \u03c7-axis in the plane of the film, almost everywhere perpendicular to the nearest edge of the limiting square, Figure 6a,b"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0000622_b978-0-12-800782-2.00003-8-Figure3.11-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0000622_b978-0-12-800782-2.00003-8-Figure3.11-1.png",
+        "caption": "Figure 3.11 Field for the determination of the unsaturated stator leakage flux of a 3.4 MW, two-pole, three-phase induction motor. One flux tube contains a flux per unit length of 0.005 Wb/m [19].",
+        "texts": [],
+        "surrounding_texts": [
+            "Numerical approaches such as the finite-difference and finite-element methods [16] enable engineers to compute no-load and full-load magnetic fields and those associated with short-circuit and starting conditions, as well as fields for the calculation of stator and rotor inductances/reactances. Figures 3.8 and 3.9 represent the no-load fields of four- and six-pole induction machines [17,18]. Figure 3.10a\u2013e illustrates radial forces generated as a function of the rotor position. Such forces cause audible noise and vibrations. The calculation of radial and tangential magnetic forces is discussed in Chapter 4 (Section 4.2.14), where the concept of the \u201cMaxwell stress\u201d is employed. Figures 3.11 to 3.13 represent unsaturated stator and rotor leakage fields and the associated field during starting of a two-pole induction motor. Figures 3.14 and 3.15 represent saturated stator and rotor leakage fields, respectively, and Fig. 3.16 depicts the associated field during starting of a two-pole induction machine. The starting current and starting torque as a function of the terminal voltage are shown in Fig. 3.17 [19]. This plot illustrates how saturation influences the starting of an induction motor. Note that the linear (hand) calculation results in lower starting current and torque than the numerical solution. Any rotating machine design is based on iterations. No closed form solution exists because of the nonlinearities (e.g., iron-core saturation) involved. In Fig. 3.13 the field for the first approximation, where saturation is neglected and a linear B\u2013H characteristic is assumed, permits us to calculate stator and rotor currents for which the starting field can be computed under saturated conditions assuming a nonlinear (B\u2013H) characteristic as depicted in Fig. 3.16. For the reluctivity distribution caused by the saturated short-circuit field the stator (Fig. 3.14) and rotor (Fig. 3.15) leakage reactances can be recomputed, leading to the second approximation as indicated in Fig. 3.16. In practice a few iterations are sufficient to achieve a satisfactory solution for the starting torque as a function of the applied voltage as illustrated in Fig. 3.17. It is well known that during starting saturation occurs only in the stator and rotor teeth and this is the reason why Figs. 3.13 and 3.16 are similar. 220 Power Quality in Power Systems and Electrical Machines F2 = 3500 N/m F1 = 3500 N/m F 2 = 3 11 0 N /m F 1 = 3 15 0 N /m F 2 = 1250 N/m F1 = 4340 N/m F 2 = 1770 N /m F 1 = 5160 N /m F 2 = 1400 N /m F 1 = 4110 N /m F 2 = 1 30 0 N/m F 1 = 5 45 0 N/m 1 \u2013B \u2013B 2 3 A A 4 5 \u2013C \u2013C \u2013A \u2013A C C B B 6 7 8 9 10 11 12 (a) 1 \u2013B \u2013B 2 3 A A 4 5 \u2013C \u2013C \u2013A \u2013A C C B B 6 7 8 9 10 N S 11 12 (b) f f 1 \u2013B \u2013B 2 3 A A 4 5 \u2013C \u2013C \u2013A \u2013A C C B B 6 7 8 9 10 N S 11 12 (c) Figure 3.10 Flux distribution and radial stator core forces at no load and rated voltage for (a) rotor position #1, (b) rotor position #2, (c) rotor position #3, Continued 221Modeling and Analysis of Induction Machines 222 Power Quality in Power Systems and Electrical Machines 223Modeling and Analysis of Induction Machines Figure 3.16 Field distribution (second approximation) during starting with rated voltage of a 3.4 MW, two-pole, three-phase inductionmotor. One flux tube contains a fluxper unit lengthof 0.005 Wb/m [19]. 0.1 0.2 0.3 0.4 0.5 0.6 0.7 0.80 0 1.00.9 2 4 6 8 10 12 0 0.4 0.8 1.2 1.6 2.0 2.4 2.8 3.2 I [kA ] ph START T [kNm ] TOTAL START I HAND I NUMERIC MEASURED CURRENT MEASURED TORQUE V ph [pu ] Vnom . ph TNUMERIC THAND Figure 3.17 Starting currents and torques as a function of terminal voltage for a 3.4 MW, three-phase, induction motor [19]. 225Modeling and Analysis of Induction Machines"
+        ]
+    },
+    {
+        "image_filename": "designv11_60_0001781_s0167-8922(08)70009-9-Figure6-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0001781_s0167-8922(08)70009-9-Figure6-1.png",
+        "caption": "Figure 6 - Short Bearing Mobility Map for a Cavitating Film Showing",
+        "texts": [
+            " If the shaft is at a given point (A) as shown its movement over a time period (dt) will be given by the 'Squeeze' Component in the direction of the squeeze path. (see Figure 5). Clearly this will be scaled. (d) Now the angular velocities can be taken into account. movement due to these will be a component (see equation 12(b)), The shaft directed tangentially. Here a sign convention is necessary and anticlockwise rotations have usually been taken as positive. This component can be vectorially added to the squeeze component to give the resultant motion, as shown in Figure 6. Strictly speaking since the value of the average mobility number of the resultant will be different from that used in calculating the squeeze step, an iteration to some acceptable tolerance will be required. However, for small time steps when the procedure is carried out on the computer this is not necessary. (e) The procedure described above has enabled the movement over a small time period to be predicted. The procedure can now be repeated starting from the new position. Normally the average values of quantities such as (P, Rb and a,) are used over a step"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0000582_978-3-319-27247-4_36-Figure2-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0000582_978-3-319-27247-4_36-Figure2-1.png",
+        "caption": "Fig. 2 3D simulation of Pendubot system with solid-works",
+        "texts": [
+            " Then, (10) and (11) can be reduced to, \u00f0a1 \u00fe a2 \u00fe 2a3 cos h2\u00de\u20ach1 \u00fe\u00f0a2 \u00fe a3 cos h2\u00de\u20ach2 a3 sin\u00f0h2\u00de _h22 2a3 sin\u00f0h2\u00de _h1 _h2 \u00fe a4g cos h1 \u00fe a5g cos\u00f0h1 \u00fe h2\u00de \u00bc s1 \u00f012\u00de a2\u20ach2 \u00fe\u00f0a2 \u00fe a3 cos h2\u00de\u20ach1 a3 sin\u00f0h2\u00de _h21 \u00fe a5g cos\u00f0h1 \u00fe h2\u00de \u00bc 0 \u00f013\u00de In which a1 \u00bc m1l 2 c1 \u00fem2l 2 1 \u00fe I1; a2 \u00bc m2l 2 c2 \u00fe I2; a3 \u00bc m2l1lc2 a4 \u00bc m2lc1 \u00fem2l1; a5 \u00bc m2lc2 In matrices form, the equations of motion of the two link Pendubot can be described as follows: D\u00f0h\u00de\u20ach\u00feC\u00f0h; _h\u00de _h + G \u00bc s \u00f014\u00de with s \u00bc s1 0 \" # D \u00bc a1 \u00fe a2 \u00fe 2a3 cos h2 a2 \u00fe a3 cos h2 a2 \u00fe a3 cos h2 a2 C \u00bc a3 sin\u00f0h2\u00de _h2 a3 sin\u00f0h2\u00de _h2 a3 sin\u00f0h2\u00de _h1 a3 sin\u00f0h2\u00de _h1 0 G \u00bc a4gcosh1 + a5gcos(h1 \u00fe h2\u00de a5gcos(h1 \u00fe h2\u00de \" # Figures 2 and 3 illustrate the 3D simulation of Pendubot with Solid-works in unstable balancing position (Fig. 2) and in stable balancing position (Fig. 3). The sliding mode controller is derived using the state dynamics described by equation, d11\u20ach1 \u00fe h1 \u00fe \u20acu1 \u00bc s1 \u00f015\u00de In which d11 \u00bc d11 d12 d22 ; h11 \u00bc h1 d12 d22 h2 ; u1 u1 d12 d22 u2 Consider a state space form of Pendubot system as ~x1 \u00bc h1 and _h1 \u00bc _x1 \u00bc ~x2: We have, _~x1 \u00bc ~x2 and _~x2 \u00bc f1(~x) + b1s1 with f1( x) = ( x11 _h1 \u00fe c12 _h2 \u00fe u1\u00de d11 ; b1 \u00bc 1 d11 ; d11 \u00bc d11 d12d21 d22 u1 \u00bc u1 d12 d22 u2 ; c11 c11 d12d21 d22 ; c12 \u00bc c12 Forwardly, consider a sliding surface as follows: S \u00bc k h1 hd1 \u00fe _h1 _hd1 \u00bc k~x1 \u00fe~x2 \u00f016\u00de The goal is to choose the scalar \u03bb value such that the system restricted on the surface (16) is of the stable characteristics"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0001976_s002490050241-Figure3-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0001976_s002490050241-Figure3-1.png",
+        "caption": "Fig. 3 Calculated current-voltage loops of model system (Fig. 1) upon single saw-tooth command voltages as in Fig. 2 with di erent slopes dV/dt: A dV/dt = 0.1 V s)1 (same as in Fig. 2); B dV/ dt = 1.0 V s)1; B1 as B but with Cm = 0; C dV/dt = 10 V s)1",
+        "texts": [
+            " This I(t) tracing exhibits non-linear features which contain all the properties of the model system. In such diagrams, the capacitive current IC = Cm dV/dt can readily be identi\u00aeed. Although it can hardly be seen in the I(t) tracing in Fig. 2, its discontinuous change by 2|dV/dt|Cm = 0.4 nA becomes obvious when the slope changes its sign at V1 and at V2. IC is also plotted in Fig. 2 in an expanded I scale. The V(t) and I(t) data from Fig. 2 are converted to the current-voltage diagram, I(V), in Fig. 3A. Knowing the general shape of the saw-tooth voltage-protocol and |dV/dt| ( 0.1 V s)1 in this case), Fig. 3A provides in principle the complete information from Fig. 2 which can be used to identify the four free parameters (Cm, g0, k0a, and k0i ) of the model system in Fig. 1. In contrast to Fig. 3B and C, the particular |dV/dt| in Fig. 3A was too small to reveal the \u00aerst parameter, Cm, by the capacitive currents. The second parameter g0 = g(V > 0)/cK,l, can be determined by g(V > 0), i.e. the asymptotic straight I/V line (not drawn) through the origin which is approached by the left portion of the curve. The remaining parameters k0a and k0i can be determined by the right part of the graph, e.g. by the two currents at zero voltage, I+,V = 0 and I),V = 0 for rising and falling V, respectively. If the voltage sensitivity coe cients da and di in the more general form of Eq. (2), ka = k0aexp(dau) and ki = k0i exp(diu), were not \u00aexed (at )0.5 and 0.5, respectively) but free as well, the information on the right hand side of the curve (e.g. the slopes in I+,V = 0 and I),V = 0 could be used to determine these two parameters also. Figure 3B and C shows the responses of the same system when |dV/dt| of the recording protocol was increased to 1 and 10 V s)1, respectively. In these cases the discontinuous changes of the capacitive currents become evident at the left and right ends of the loops. Correspondingly, the left side of the plot is split into two branches. Here the two intersections of the curves with the V-axis do not indicate, of course, the precise equilibrium voltage of the system (here EK). The right hand side of these tracings steepens up dramatically with |dV/ dt| (notice di erent I scales in Fig. 3B and C), owing to the relatively slow and therefore incomplete reequilibration between a and i after fast and short V changes over the same amplitude. Figure 3B1 shows Fig. 3B without the capacitive up and down shifts [\u00aerst term on right side of Eq. (4) ignored]. This tracing demonstrates that with a saw-tooth clamp the capacitance-corrected tracings of a V-gated conductance has the general shape of an 8, and that the internal intersection of the 8 is located at the equilibrium voltage of the conductance. This characteristic feature is evident because the second term in Eq. (4) [with Eq. (1)] is always zero at the equilibrium voltage (IGHK = 0), irrespective of the conductance (g0) and the activity (pa) of the pathway",
+            " Only closer inspection shows a slight kink around V0 (which will be focused on below with tracing 4E), and a slight widening of the loop in the negative voltage range indicating a K inward recti\u00aeer by rules 3 and 4. Compared with the other examples in Fig. 4, the tracing of Fig. 4A is displayed with the most expanded current scale; it demonstrates that with the applied jdV =dtj 1 V s)1 in all the samples, the capacitive currents IC = CmdV/dt can be ignored. This current would be most obvious as a discontinuous jump of 2IC at the two extreme voltages where dV/dt changed its sign abruptly (see Fig. 3B and C). The simple curvature of the current loop in Fig. 4B is reminiscent of the intrinsic outward recti\u00aecation of IGHK for K [Eq. (1)] where the limiting slopes for outward and inward conductance correspond to cK,c and cK,l, respectively. The ratio of these slopes in Fig. 4B, however, is much smaller than cK,c:cK,l 40 in C. wailesii (Gradmann and Boyd 1999). One probable reason for this discrepancy is an additional `leak' pathway (probably a Cl) conductance) which (1) adds, on a percent basis, more to the small K inward conductance than to the large K outward conductance, and (2) shifts VI = 0 more positive than EK",
+            " As in tracing 4E, the occurrence of such a kink is frequently coupled with a double-8 shape. It is assumed that these kinks re\u00afect steep gating of a Cl) conductance. Thus far, all examples here show monotonic currentvoltage loops. The following three examples (Fig. 4F, G, and H), however, show portions with negative slope. In principle, the occurrence of negative slope conductances in such IV loops does not necessarily point to a complex gating system. This is evident from the portions with dI/ dV < 0 in the theoretical examples in Fig. 3A, B, and D, which arise from a very simple gating system. In Fig. 4F, negative slope conductance is evident in the far negative voltage range. In this example, this feature appears to be a steady-state property, because it is rather independent of the sign of dV/dt. Other features of this loop are known already from previous examples: widening in positive voltage range (Fig. 4D), kink, and double-8 shape (see Fig. 4E). The portion of the current-voltage loop in Fig. 4G with negative slope conductance is also in the negative voltage range, however, without coincidence for rising and for falling voltage (as in Fig",
+            " 4C (=5A); D steady-state IV data (Gradmann and Boyd 1999) recorded from a di erent cell with a step protocol tative feature is a reversible, hyperpolarization-induced increase of the K conductance (rule 4). The smaller jdV =dtj, the more time the system spends at elevated negative voltages, and the K conductance has more time to increase towards the maximum, g0. At about )200 mV, g0 seems to be reached nearly in Fig. 5C and completely in Fig. 5D, when the extrapolated slope (DI/ DV, not illustrated) passes the origin [see Eq. (1) and discussion of Fig. 3A]. The tendency to saturate can already be seen at the end of the negative going branch in tracing B (dV/dt = 0.25 V s)1) but not in the fastest tracing A (dV/dt = 1.0 V s)1). Characteristic for a V-gated inward recti\u00aeer is also the \u00aending that within the negative going branch the ``critical'' voltage of the conductance increase (the voltage at which the curvature of downward bending is strongest) occurs already at smaller hyperpolarizations when jdV =dtj is small. In order to analyze the processes in Fig"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0003768_mfi.1996.572183-Figure6-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0003768_mfi.1996.572183-Figure6-1.png",
+        "caption": "Fig. 6 Coordinate systems for adaptive walking",
+        "texts": [
+            " As for lower-limb trajectories, the trajectory of the lower side of foot parts is set with constant leveling. A flow chart of this walking control methcd is shown in Fig. 5. A processing methcd for the information on the landing sm?gce acquired and the control methods of lowr-limb trajectories, which are vital factors in the systeq are described below. 5.1 Processing method of information on landing surfaces All kinds of information acquired by WAF-3 for detecting a Iandmg surfBce is sent to the computer for walking control in real time, and arranged into the following information Fig. 6 shows each system of m r h a t e s set for the (1) Information on the grounding of the lower-foot plate: The grounding is judged with ON/OFF ofthe microswitches installed in the four comers ofthe lower side ofthe lower-foot plate. Basically, the leg wasjudged grounded when all microswitches in the four comers turned on. (2) Information on the relative angle and the relative distance of landing surfaces and upper-foot plates: With respectto ohwing the relative position to the landing surfice, it is p i b l e to determine a plane based on the measurement of the positions of three points"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0000991_tmag.2016.2645541-Figure2-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0000991_tmag.2016.2645541-Figure2-1.png",
+        "caption": "Fig. 2. Estimation of the error angle with torque commands at (a) 45\u00b0 and (b)",
+        "texts": [
+            "org/publications_standards/publications/rights/index.html for more information. pole position is along with dmotor. Reference commands are given along the dcontrol-qcontrol axes. The error angle between the two reference frames is denoted by \u03b8err. To find the actual initial pole position, \u03b8err should be estimated. If the real pole position, \u03b8r is same as the estimated pole position, \u03b8est, the error angle, \u03b8err will be equal to zero. Let us consider, a torque command is applied along the qcontrol+45\u00ba axis as shown in Fig. 2(a), then , the real torque along qmotor can be expressed as follows,      45cos45cos * 45 * 4545 errestre TTT  (1) Now as shown in Fig. 2(b), if the control axes are shifted by 90\u00ba and another torque command is applied to the qcontrol45\u00ba axis, the resultant real torque, along qmotor axis can be defined as,       45cos45cos * 45 * 4545 errestre TTT  (2) The position information supplied to the current controller is expressed as \u03b8ref. It includes \u03b8fb, a feedback relative position, obtained using incremental encoder; \u03b8shift, the shift angle command to decide whether the command signal is applied at 45\u00ba or at -45 \u00ba; and \u03b8cmp, the compensated IPP obtained from IPP estimator",
+            " But this leads to errors in the IPP estimation under load condition, because, in real system, the effect of the load is present. Therefore while finding \u03b8err, if the load torque term is ignored, then the estimated signal keeps on fluctuating [18] and does not converge to the real IPP under load condition. As a result, a different method is required which can give accurate estimation of the IPP both under no load and loaded condition. Again, it is seen from (9) and (10), the reference frame method needs a reference command, to be applied along qcontrol axis. The idea is illustrated in Fig. 2. Therefore, it is important to select an appropriate command signal. Selection of the reference command signal for the accurate IPP estimation is very important. A reference command must have the following characteristics, (i) When the reference command is applied, it must not result in the movement of the IPP and cause error. (ii) If there is a movement due to the command signal, then the movement must be compensated to make the resultant movement to zero. To achieve the above mentioned conditions, a test command pattern signal is selected as shown in Fig"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0002879_6.1998-3285-Figure9-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0002879_6.1998-3285-Figure9-1.png",
+        "caption": "Figure 9 Original Seal/Rotor pressure distribution viewed in the axial direction for 0.016\" axial clearance",
+        "texts": [
+            " 8, reveals a very limited region in which the flow from the air bearing holes is influential. This flow pattern surrounding the orifice is consistent with that observed on the seal after testing. Note that over the entire arc, there is mixing of the flows from the dam and air bearing regions of the seal, and the strong radial component originating at the dam hinders the formation of a hydrostatic film at the air bearing. The radial flow from the dam slows upon entering the hydrostatic bearing region, resulting in increased pressure at the bearing face as shown in Fig. 9. This increased pressure inhibits the seal's ability to close. Figure 8 Original Seal/Rotor flow field viewed in the axial direction for 0.016\" axial clearance The analytical results indicate that in its original configuration, the force balance generated at a 0.016\" axial clearance will not allow the seal to continue moving closer to the rotor. Flow from the air dam is intended to exit the seal via the seal vent slots. However, the introduction of flow from the dam into the air bearing region of the seal raises the pressure at the air bearing face, limiting seal closure"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0001996_s0736-5845(98)00026-x-Figure35-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0001996_s0736-5845(98)00026-x-Figure35-1.png",
+        "caption": "Fig. 35. Interpolating the actual curve point P(u i (t#*t)).",
+        "texts": [
+            "8) to a curve parameter planning u (t#*t) we approximate u (t#*t) with Taylor\u2019s equation around the interpolation time *t: u i (t#*t)\"u i #*t ) uR i #A *t2 2 B ) u( i#2. (A.11) Terminating Eq. (A.11) after the quadratic term should be enough and we get Eq. (A.12) by putting Eqs. (A.9)\u2014 (A.11) together u i (t#*t)\"u i # *t 2s@(u i ) (*u v #*t )*u a ), (A.12) where *u v \"2 ) v i (t) (A.13) and *u a \"a i (t)#sA(u i ) A v i (t) s@ i (u i )B 2 . (A.14) Interpolating the end-effector path P (t#*t) in Fig. 35 with the cubic polynomial and the curve parameter u(t#*t) yields Eq. (A.15), P (u i (t#*t))\"A x(u i (t#*t)) 2 t (u i (t#*t))B . (A.15) The inverse kinematics for the redundant robot at the desired end-effector position P (t#*t) will be, as usual, calculated at velocity level. Using kinematic transformation [24, 25] we get the difference dP(u i (t)) between the desired end-effector position P(t#*t) and the actual position P (t) as a function F : dP(u i (t))\"F(P(u i (t)), P (u i (t#*t))). (A.16) The solution of dh at velocity level with the general formulation of the generalized inverse of the Jacobian matrix [3]: dh\"J`dP#(E"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0003340_0010-4361(89)90337-6-Figure25-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0003340_0010-4361(89)90337-6-Figure25-1.png",
+        "caption": "Fig. 25 Schematic picture of interlaminar shear failure and tensile failure of a specimen in MFW test",
+        "texts": [
+            " This is the same as has already been found in the NII test. The failure behaviour of a LCe matrix plate seems to be similar to that of the continuous fibre reinforced conventional composite. The highly oriented liquid crystal domains act like continuous fibres and the interfaces between the domains are potential lines of weakness. The multiaxial stress system that imposes shearing forces along these lines of weakness may therefore initiate shear failure. The tensile failure, however, occurs in the plane perpendicular to the MFD (see Fig. 25). Because of the cross-laminating caused by fibres delamination of composites becomes difficult, both the tensile strength in plane B and the shearing strength in plane A are more or less improved. But, in the direction parallel to the ~vo, the relative movement between fibres and matrix becomes easier if the fibre-matrix interface is poor, and tensile strength in this direction is lowered. The degree of aniostropy will be reduced ~, 9 Furthermore the addition of brittle fibres will result in brittle mechanical behaviour; the displacement of the composites to total failure of the specimen becomes smaller"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0002678_1.2829318-Figure7-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0002678_1.2829318-Figure7-1.png",
+        "caption": "Fig. 7 Infinitesimal surface element",
+        "texts": [
+            " From a purely analytical point of view, once the domain of the surface parameters T and the transformation 3' are defined, the sector of surface S is also defined. The transformation ^ is here represented by three equations for the coordinates x, y, z in the surface parameters \u00a7 and u. Rotor Loading Due to Pressure In the quasi-static case, which we have considered as a hy pothesis, two types of loads due to the pressure: force and moment, are present (Adams and Soedel, 1995). We will con sider a general position of the rotor and an infinitesimal element of area dS (see Fig. 7) , in the neighbourhood of a general point (2(x, y, z) on the surface S which has the equation P = P{u, \u2022& ). The expressions of the force d\u00a5 on the infinitesimal element dS and the moment dMo, as regards the origin O of the refer ence system, are respectively: d\u00a5 = pndS, dUo = p(Q - O) X ndS, from which, if the considered surface S is regular: F = p I ndS =p I n | P\u201e X Ptf I rfr Js JT X P. (1) (2) X P,, \\PuXPJdT Mo = p I (Q- 0)X ndS = \u0302 J/\" X P^dT, (3) = P j (Q- O) xn\\P\u201exP,\\dT JT = P ! (Q-O)X (P\u201e X P,)dT"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0002879_6.1998-3285-Figure1-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0002879_6.1998-3285-Figure1-1.png",
+        "caption": "Figure 1 Cross section of the enhanced 36\" diameter aspirating face seal",
+        "texts": [
+            " As such, they have a potentially significant performance advantage over conventional labyrinth seals, particularly at large diameters. In addition, aspirating seals are inherently not prone to wear, owing to their non-contacting nature, so their performance is not expected to degrade over time. A 36\" diameter aspirating seal, for application to the GE90 low pressure aft outer seal location, was designed and fabricated by the Stein Seal Company2'3. A cross-section of the seal design, enhanced by the presence of a flow deflector on the rotor face, is shown in Fig. 1, and major seal components are listed in Table 1. A test plan has been Copyright\u00a9 1998 by the American Institute of Aeronautics and Astronautics, Inc. All rights reserved. 1 American Institute of Aeronautics and Astronautics D ow nl oa de d by U N IV E R SI T Y O F FL O R ID A o n Se pt em be r 19 , 2 01 6 | h ttp :// ar c. ai aa established to evaluate seal performance under a variety of conditions that the seal would be subjected to in the GE90 aircraft engine application. The tests are being executed on a full scale rotary test rig, and analyses have been performed using 3-D Computational Fluid Dynamics (CFD) in order to validate test results"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0002252_s0890-6955(98)00006-6-Figure10-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0002252_s0890-6955(98)00006-6-Figure10-1.png",
+        "caption": "Fig. 10. The shear plane area A and projected area Q on the tool face when f . fCM.",
+        "texts": [
+            " And second, the line PC is intersected by curve , or curve EP is intersected by curve . In the first case, wherein the straight line PC is intersected by curve , feedrate (f) can be divided into another two cases, depending on the cutting condition, as: 1. when condition f . fc occurs 2. when condition f # fc occurs The present work only concerns the case (i), i.e. f . fc. future work is intended to extend the present analysis to case (ii). The real cutting conditions are shown in Figs 8 and 9. For simplification of the calculations, the simplified model is shown in Fig. 10, in which the actual feed in Figs 8 and 9 have been modified. The length of WW9 can be found from Fig. 8 as: WW9 = ll/sin(uPC + Ce 2 Cs) (17) ll = (l1 + l2 + l3 + l4)sin(Ce 2 Cs) + (h1 + h2)cos(Ce 2 Cs) + R2cos2uR2 2 R3cos2uR3\u00b7cos(Ce 2 Cs) 2 fsinCe 2 R1\u00b7[1 + sin(Ce 2 Cs) 2 2uR1] (18) The modified feed fCM is calculated as: f CM = fcosCs + R1(1 2 cos2uR1) + WW9cosuPC (19) If WW9 , 0 i.e. the condition as shown in Fig. 9, the intersection angle uSS can be calculated as: uSS = Ce 2 Cs + sin21(mm/R1) (20) where mm = (l1 + l2 + l3 + l4)sin(Ce 2 Cs) + (h1 + h2)cos(Ce 2 Cs) + (R2cos2uR2 2 R3cos2uR3)\u00b7cos(Ce 2 Cs) 2 fsinCe 2 R1sin(Ce 2 Cs) (21) Thus yielding modified feed as: f CM = fcosCs + R1(1 2 cos2uss) + WW9cosuPC (22) After the modified feed fCM is obtained, the shear plane area and the projected area on the tool face can be calculated from Figs 4 and 10, as: A = A1 + A2 + A3 + A4 + A5 + As (23) where A1 = 0"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0003689_robot.2001.932951-Figure2-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0003689_robot.2001.932951-Figure2-1.png",
+        "caption": "Fig. 2. Base-ward normal accumulation property of the proposed coordinate transformation",
+        "texts": [
+            " U RE:MARK I : The theorem establishes that the coordinate axes consisting of { 'TIS,} are the geometrical one corresponding to the tangential and normal coordinate axes a t each contact Ckz. Note that these two coordinate axes are sufficient for expressing the other coordinate axes due to joint torques and frictional forces of the upper contact points. Fortunately, the base-ward normal accumulation property states that they affects only along the normal direction at the contact Cki. To grasp the meaning of the above theorem, consider a three degrees-of-freedom planar planar k-th finger; shown in Fig. 2. For the current configuration, the numerical values of I I T k and n<k are equal to I I T l and ll,, given in Appendix B. 221 2 3.2 Grasp feasibility inequality Using the geometrically decomposed contact force (10) and (11) we can formulate the grasp feasibility inequality using the friction constraint (9). By choosing the coordinate system the normal force constraint corresponds to 3=r+L 3=2+1 Then the friction cone constraint reduces to which is called the decomposed friction constraint. Noting that the above two inequalities are of linear inequality form, they can be simplified to the following linear inequality S ( 4 , PIE I N P ) R ( 4 1 PI7 (15) where S, A, R are all diagonal matrices"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0000622_b978-0-12-800782-2.00003-8-Figure3.13-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0000622_b978-0-12-800782-2.00003-8-Figure3.13-1.png",
+        "caption": "Figure 3.13 Field distribution (first approximation) at starting with rated voltage of a 3.4 MW, twopole, three-phase induction motor. One flux tube contains a flux per unit length of 0.005 Wb/m [19].",
+        "texts": [
+            "16 depicts the associated field during starting of a two-pole induction machine. The starting current and starting torque as a function of the terminal voltage are shown in Fig. 3.17 [19]. This plot illustrates how saturation influences the starting of an induction motor. Note that the linear (hand) calculation results in lower starting current and torque than the numerical solution. Any rotating machine design is based on iterations. No closed form solution exists because of the nonlinearities (e.g., iron-core saturation) involved. In Fig. 3.13 the field for the first approximation, where saturation is neglected and a linear B\u2013H characteristic is assumed, permits us to calculate stator and rotor currents for which the starting field can be computed under saturated conditions assuming a nonlinear (B\u2013H) characteristic as depicted in Fig. 3.16. For the reluctivity distribution caused by the saturated short-circuit field the stator (Fig. 3.14) and rotor (Fig. 3.15) leakage reactances can be recomputed, leading to the second approximation as indicated in Fig"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0003013_a:1008115522778-Figure2-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0003013_a:1008115522778-Figure2-1.png",
+        "caption": "Figure 2. Simplified free body diagram for the deformation of a flexible part.",
+        "texts": [
+            " Flexible parts can be buckled or bent by the axial or lateral forces during misalignment compensation. The relationship between part deformation and reaction forces depends on the part shape, the contact conditions between a part and its mating hole, the grasping condition in the end effector of a robot, and whether or not other external forces act on, etc. In flexible peg-in-hole tasks, it can be assumed that a part is deformed only by the axial or lateral forces. Accordingly, it is necessary to investigate the force-deformation relations under the load conditions as shown in Figure 2. Because the behaviors of most flexible parts can be approximated to elastic within the range in use mainly, this paper deals with only the elastic deformation. Parts deformation in flexible parts assembly is usually large [13, 14], though it is elastic. The deformation equation can be obtained from the relations between the curvature and the deformation [15, 16]. When the inclination angle of a part is large, namely, when it is large deformation, the differential equation that is related with the bending momentM and the lateral deformation v along the xv direction is obtained by d2v/dx2 v (1+ (dv/dxv)2)3/2 = M EI , (1) where EI is the flexural rigidity. Using Equation (1), the differential equation related to the deformation under the load conditions as shown in Figure 2 is obtained by d2 dx2 v ( EI (d2v/dx2 v ) (1+ (dv/dxv)2)3/2 ) + dv dxv ( Fv dv dxv ) = q, (2) where q is a uniformly distributed load intensity per unit length, which is equivalent to a lateral concentrated load Fh. And Fv is a vertical axial load. The elastic deformation curve to be obtained from Equation (2) is called an elastica, its mathematical solution can be obtained by the trial and error method. However, its solving process and solution are very complex and time-consuming. Accordingly, it is not practical to apply its mathematical solution to actual assembly tasks",
+            " To obtain the robotic error-corrective motion corresponding to the measured assembly errors is to obtain the relations between the five components of the assembly errors and the five components of the error-corrective motion. Because the behavior of a flexible part during error compensation is controlled by the force-deformation relations, the relations between the errors and the corrective motion can be obtained by using Equation (2). If it is assumed that a flexible part is deformed only by a lateral force Fh and a vertical force Fv as shown in Figure 2, and that its behavior is elastic, then the deformation v is just the function of the forces Fh and Fv , the flexural rigidity EI , and the boundary conditions Cbc at both ends of the part. It is expressed by v = fv(Fh, Fv,EI,Cbc). (3) From Equation (3), the relation between the errors and the corrective motion is expressed by  mre mr\u03c6h mr\u03c6v m\u03b8x m\u03b8y  = fn(  em \u03c6h \u03c6v \u03b8v \u03b8h v(Fh, Fv,EI,Cbc) ). (4) In order to obtain the relationship between misalignment and its corrective motion which is available for all tasks in flexible parts assembly, all parameters in Equation (4) have to be considered"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0003851_s0167-8922(08)70928-3-Figure9-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0003851_s0167-8922(08)70928-3-Figure9-1.png",
+        "caption": "Fig. 9 : Isotherms in t h e mid-plane bear ing f o r a n eccen t r i c i ty r a t i o of 0.8 and rotat ional speed of 2000 rpm",
+        "texts": [
+            " In these conditions t h e ambient tempera ture is assumed to be constant and equal to 45\u00b0C as shown in Table 1. Figures 9 and 10 produce t h e isothermal c h a r t in t h e journal bearing at a relative eccentr ic i ty of 0.8, a rotational speed of 2000 rpm and for a convection hea t t ransfer coeff ic ient of t h e bush of 80 W/m \"C. 2 40 Dif fe ren t s ca l e s a r e used f o r t h e sha f t , t h e bush and for t h e fi lm which appea r s to b e of cons t an t thickness because of t h e var iables used. Figure 9 which gives t h e isothermal c h a r t in t h e mid-plane shows t h a t t h e maximum t e m p e r a t u r e is s i tua t ed in t h e oil fi lm near t h e bear ing edge downstream t h e minimum film thickness zone. The h e a t f lux l ines which a r e perpendicular to t h e isothermal l ines are deduced f r o m fig. 9 and a r e given in fig. 10. This f igu re shows t h a t t w o d i f f e ren t zones ex i s t in t h e bearing. In t h e first one, which corresponds to t h e coldest p a r t of t h e fi lm t h e s h a f t and t h e bush give h e a t to t h e fluid. In t h e second one, which corresponds to t h e ho t t e s t p a r t of t h e fi lm t h e fluid gives h e a t to t h e bush and to t h e sha f t . T h e s h a f t and t h e bush c a r r y ove r t h e h e a t f r o m t h e ho t t e s t zone of t h e fi lm to t h e co ldes t one"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0003773_robot.1997.619353-Figure1-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0003773_robot.1997.619353-Figure1-1.png",
+        "caption": "Fig. 1 Choice Coordinate systems for kinematic constraints",
+        "texts": [
+            ", x4, x,) describing the configuration of the mobile robot is 3 + N,, + ( N j + Consider that the origin of the choice of coordinate system OeJ, e E {f, c , oc, od}, j E { 1, . . . , N e } , the choice of coordinate frame of the j t h wheel of type e with the origin on the wheel axle above the center point of wheel-ground contact for fixed, centered, and omnidirectional wheels and at the pivot of the arm of off-centered wheels is given by (1, Q j in polar coordinates in the choice of coordinate system M and the radial line cy is the x-axis of O,,(see Fig. 1). Similarly, the origin of the choice of slip coordznatefmmes M e , (or t,he origin of M e , j is (cl, p - i) in the choice of coordinate system U,, and the radial line /r - is the x-axis of the frame Me,. Let ;( be the angle that the direction of complementary rolling 4 of an omnidirectional wheel makes with the direction of &, the axis about which the wheel is driven. The three scalar components of the constraints imposed by the wheels are as follows: Nc + No, + N o d ) + Nod + Nc = 3 + Nj + 2 ( N c + N o c + N o d ) "
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0000544_9781119268628-Figure8.1-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0000544_9781119268628-Figure8.1-1.png",
+        "caption": "Figure 8.1. t-Translation invariant principle",
+        "texts": [
+            " Designing for Future Needs 111 This situation is depicted by the following equation: DA(t) = DB(t\u2212 \u03c4) for \u2212 \u03c3 \u2264 t \u2264 \u03c4 In other words, A is behind B with the amount of time, \u03c4 , in terms of the change of desire state. In this scheme, defining future needs of A during the period spanning from the present, t = 0, to the future time, t = \u03c4 , reduces to revealing the history of desire state of B from the past time, t = \u2212(\u03c3 + \u03c4), to the present, t = 0. It is suggested that this scheme should be called the \u201ct-translation invariant principle\u201d for making inaccessible future needs of a person accessible to those who want to understand and use the person\u2019s future needs at the present time. Figure 8.1 depicts it schematically. The issue of predicting people\u2019s future needs has been dealt with in the field of ergonomics under the name of prospective ergonomics, which is about the conception and design of future systems, products or services [ROB 09]. However, as described, the problem of defining a person\u2019s future needs should be reduced to the problem of describing the other person\u2019s behavioral selection history, who is experiencing a certain amount of time ahead of the person in question; this is done by understanding a person\u2019s behavioral selection process in the past, which is, in large part, governed by memory processes, and therefore, it is the issue of the field of cognitive sciences"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0001783_aim.1999.803252-Figure3-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0001783_aim.1999.803252-Figure3-1.png",
+        "caption": "Fig. 3 Experimental Testbed for Impedance Control",
+        "texts": [
+            " This induces that the value of the steady state torque is the stiffness multiplied by the steady state deformation 68 , and fingertip behaves like a programmable spring. 5. Experimental Results To verify the proposed control approach, a DSP(Digita1 Signal Processor)-based real-time control system has been built. The kernel of the system is a commercially available processor board with a TMS320C40. Using the Cartesian impedance control scheme developed in section 4, an experimental testbed has been built as shown in Fig. 3. According to a desired trajectory, the fingertip should move along X-axis in the finger base coordinate system from point 'a' to point 'b' with a constant linear speed at 5 0 m d s . However, because there is an obstacle between these two points the fingertip can not reach point 'b' physically and an impact appears when the fingertip makes contact with the obstacle. The entire experimental results are shown in Fig.4. From Fig.4(a) the fingertip can not continue to move along the dot-lined predefined trajectory in X-axis at about 0"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0000153_imece2015-50907-Figure9-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0000153_imece2015-50907-Figure9-1.png",
+        "caption": "FIGURE 9. THE INTERSECTION CURVES BETWEEN OUTER BOUNDARY SURFACES.",
+        "texts": [
+            " The positions of those corner points are C1 = [0,0,0]T C2 = [1980,0,0]T C3 = [1980,1750,0]T C4 = [0,1750,0]T (18) corresponding to the outer boundary surfaces E1, E2, E3, E4, which are shown in Fig. 8. We firstly calculate the min distance between E1, E2, E3, E4 two by two and the results are all zero which means that the four outer boundary surfaces intersect with each other. After the start points for the continuation method are solved out by the Newton-Raphson method, we obtain the complete intersection curves utilizing the PITCON code package [11], as shown in Fig. 9. The upper limit of the bounding boxes is 1300 mm. So we choose the point from the center intersection area which has the intermediate z coordinate enveloped by all intersection curves as 5 Copyright \u00a9 2015 by ASME Downloaded From: http://proceedings.asmedigitalcollection.asme.org/pdfaccess.ashx?url=/data/conferences/asmep/86880/ on 01/31/2017 Terms of Use: http://www.asme.org/about-asme/terms-of-use the optional stop position which is far away from the inner and outer barrier boundary. Indeed the coordinates of the point are [775,875,1899]T "
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0002055_ip-epa:19960203-Figure10-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0002055_ip-epa:19960203-Figure10-1.png",
+        "caption": "Fig. 10 current 'Torque' and 'jZux' components of rotor flux-linkage due to stator",
+        "texts": [],
+        "surrounding_texts": [
+            "With the stator current given by i s = I , sinwt, or in phasor presentation = ISLO (12) Two important observations picked up later are: - the resultant rotor flux linkage Y', is always in quadrature with the rotor current I, - the area of the flux linkage triangle is a maximum at the frequency w = UT, i.e. when @ = 7d4. 2.5 Frequency response: constant resultant secondary flux linkage The coupled circuit model is used later to derive torque, so we need to recognise that in practice it will be desirable to operate the motor so that its magnetic circuit is fully utilised. In the coupled-circuit model this corresponds to keeping the resultant secondary (rotor) flux linkage at rated value at all frequencies. It is convenient to normalise the currents by defining two current references. For the primary, I,, is the current to produce rated resultant secondary flux linkage under DC conditions, i.e. with the secondary open circuited. Similarly, I,, is the current in the secondary to produce rated secondary flux linkage if the primary was open circuited. The reference currents are thus Hence from eqns. 13 and 14, with constant resultant rotor flux linkage, the normalised currents are given by L d - ; I S , L W T Iro (17) The most important point to note is that with Y, constant the secondary current is directly proportional to frequency. 2.6 Sudden transition in secondary current Now examine how to control the primary current to make the secondary current change instantaneously from a steady state at one amplitude and frequency to a steady state at another amplitude and frequency, while the amplitude of the resultant secondary flux linkage is maintained constant. We see later that this is an essential requirement for instantaneous torque control. Suppose that for t < 0, the secondary current is Under steady-state conditions the corresponding primary current and resultant secondary flux linkage are given from eqns. 2 and 3 as i,l = Irl sin(w1t + a ) (18) (19) RITi = ~ cos(w1t + a) w1 It is assumed that the amplitude and frequency of the secondary current are to be changed at t = 0 so that for t > 0 the secondary current is Zr2 = I72 sin(w2t + P ) (22) The steady-state solutions for the primary current and resultant secondary flux linkage are where (25) w2Lr R z , ~ = Je ; 42 = tan-' - From eqns. 20 and 24 note that for YJ, to remain constant the secondary currents must be proportional to the frequency, i.e.: (26) 4 2 - w2 Ir1 W I Since there must be no discontinuity in the flux linkage at t = 0, it follows from eqns. 20 and 24 that p = a. Hence, the necessary conditions for obtaining an abrupt transition of the amplitude of the rotor current from I,, to I,, are: (a) The supply frequency must be abruptly changed in the ratio of the secondary currents, as indicated by eqn. 26 (b) The amplitudes of the primary current must be related by the equation IEE Proc-Elect,. Power Appl., Vol. 143, No. I , January 1996 61 (e) The phase of the primary current must change abruptly by an angle (A@) given by 2.7 Experimental verification These findings have been verified on a three-phase slipring motor with the rotor locked in the position of alignment between stator and rotor phases. The rotor current waveforms shown in Figs. 6 6 relate to a sudden change in the stator frequency from 0.8 to 4Hz. In Fig. 4, there is no change in the amplitude or phase of the stator current, and consequently the rotor current displays its characteristically long transient. In Fig. 5, the amplitude of the stator current is changed as indicated by eqn. 27, but there is no change in phase, so again the rotor current displays a long transient. But in Fig. 6, the phase is also abruptly changed (as given by eqn. 28) and the rotor current jumps immediately into its new steady state (with five times its previous amplitude and frequency) with no evidence of any transient. Fig.5 I 1 f 1 I 14 A)div] Rotor current response: Frequency and umplitude control 2.9 Summary It has been shown that if the primary current is abruptly changed with the correct amplitude, frequency and phase, the secondary current jumps immediately to its new steady state while the resultant secondary flux linkage remains unchanged. The ability to suddenly change the secondary current from one steady state to another is the key to vector control of torque, as will now be shown. I 1 sb %- A Fig.7 Two-phase model for cage induction motor 3 current-fed operation 3.7 Motor model The simplified model (Fig. 7) draws on two well-established methods for easing the treatment of the threephase motor with a cage rotor. First, the three stator windings are replaced by a two-phase equivalent, consisting of two identical sinusoidally distributed windings in electrical space quadrature. And secondly, the cage rotor is replaced by a two-phase wound rotor, comprising two identical sinusoidally distributed windings in space quadrature. It is well known that this model correctly represents the real motor, except in relation to second-order effects such as space harmonics which are not of interest here. The rotor has no saliency, so all four self inductances are constant and may be written as Locked-rotor analysis under balanced stator Ls, = Lsb Ls; LT, = LTb = L, (29) The mutual inductances between the two stator windings and the two rotor windings are zero because both sets are in space quadrature. The mutual inductances between stator and rotor windings vary sinusoidally with rotor position, and with the conventions indicated in Fig. 7, they are Msa ,a = +M cos e, Msb ,, = +M cos The voltage equations for the rotor windings are Fig. 6 Rotor current response: Frequency, amplitude and phase control 62 IEE Proc.-Electr. Power Appl., Vol. 143, No. I , January 1996 impressing in the stator a balanced set of sinusoidal currents given by is, = I , cos(w,t); or in phasor form E= I, LO i,b = I , sin(w,t); or in phasor form z= I , L - - 2 7r (32) Solution of eqn. 31 yields the corresponding steadystate rotor currents as 3.2 Steady-state torque/frequency relationship The torque is given by (35) Applying this to the model in Fig. 7 and using the mutual inductance eqn. 30 gives the torque as T = -M(Z,,i,, + i s b i r b ) sin@, + M(is&, - is=&) cos@, (36) In practice, the steady-state locked-rotor torque is independent of the rotor position, and this finding can be verified by substituting the expressions for stator and rotor currents given by eqns. 32 and 33 in eqn. 36. We therefore choose to study the behaviour when 0, = 0, i.e. with the rotor and stator windings directly aligned, without loss of generality. The torque expression then simplifies to The solution for the currents also becomes much easier with the stator and rotor windings aligned, because the pairs of statodrotor windings are then decoupled from each other, and all the results obtained in Section 2 can be applied directly. T = M ( i s b i r a - i s a i r b ) (37) With balanced two-phase currents given by is, = I , cos(w,t); z s b = Is sin(w,t) (38) the torque becomes sin(w,t - 4) ws MIS +I, cos(w,t) 7 cos(w,t - d ) ] It was shown in Section 2 that it is necessary to maintain constant resultant rotor flux linkage if the rotor IEE Proc-Electr. Power Appl., Vol. 143. No. I , January 1996 current is to make a sudden jump from one steady state to another. The variation of I, and I, to ensure constant rotor flux linkage is given by eqn. 17, and under these conditions the torque expression (eqn. 39) becomes As the resultant rotor flux linkage (YJ is kept constant, the torque expression can be written in normalised form as 9 2 - = w,r; where To = 2 = (MIso ) ITo (41) T T O Lr The normalised torque To represents the product of the resultant rotor open-circuit flux linkage (rated value) and the rotor current which would produce the same flux linkage if the stator was open circuited. Eqn. 41 shows that the variation of torque with normalised frequency (0 ,~) at constant resultant rotor flux linkage is a simple linear function of frequency. Note that this is an exact expression which is valid for all frequencies, and should not be confused with the approximately linear torque/frequency relationship for voltage-fed operation, which is only valid for small values of slip frequency. The similarity between eqn. 40 and that for torque in a DC machine makes clear why, when the induction motor is operated so that the resultant rotor flux linkage is kept constant it has the same steady-state torque characteristics as a DC machine with armature current control. (It is not obvious at this point that the dynamic torque can be controlled in the same way as for an armature-controlled DC machine, but this is shown subsequently.) 4 To strengthen our physical insight we now move from a circuit viewpoint and develop an alternative picture of behaviour under locked-rotor conditions by utilising the well established concept of space phasors. Space phasors can provide an accurate picture of the instantaneous spatial relationships between distributed quantities which vary sinusoidally in space inside the machine, such as the ampere conductor distribution of a sinusoidally distributed winding, or a sinusoidal flux density wave. With care they can also be used to represent integrated quantities such as flux linkage, but if the winding in question and the associated flux density are not sinusoidally distributed in space we resort to the notion of an equivalent sine-distributed winding with the same flux linkage as the actual winding. 4. I Ampere-conductor distribution and MMF space phasors When balanced sinusoidally distributed windings are fed with balanced sinusoidal currents the resultant amperexonductor distributions (and MMFs) of stator and rotor consists of waves of constant amplitude which rotate at synchronous speed with respect to the stator. The steady-state relationship between the associated rotor and stator MMF space phasors will now be explored. At an angular position 0, in the air gap with respect to the axis of the stator winding sa (Fig. 7), the resultant stator MMF is F, = Nsi,,cos6,+Nsis~ sin@, = N,I, cos(w,t-0,) (42) T = q,Ir (40) Space phasors under locked rotor conditions 63 or in space-phasor notation where N, is the effective number of turns per stator winding. The corresponding resultant rotor MMF, at an angular position a, (a, = 0,-8,) in the air gap with respect to the rotor phase winding ra (Fig. 7), is F, = Nrzra cos a, + N r i r b sin a, - (43) F - N I & l ( w a t - @ s ) s - s s 7r (44) = NrIr cos w,t - 8, - 4 - -) ( 2 or in space-phasor notation - (45) F , = N I &3 (w,t-& -$- 5) where Nr is the effective number of turns per rotor winding. It is evident from eqns. 43 and 45 that the spatial angle between the space phasors of stator and rotor MMFs is $+7c/2, and it is defined only by the frequency and rotor time constant. Notice that the rotor MMF, and consequently the rotor ampere-conductor distribution, do not depend on the rotor position. 4.2 Space phasors for flux linkages Using the space-phasor notation, the resultant rotor flux linkage is given by In this equation the term MZs represents the space phasor of the rotor open-circuit flux linkage when the stator windings carry balanced sinusoidal currents. The magnitude and spatial position of this phasor depend only on the stator currents, and it rotates at a speed determined by the supply frequency, i.e. it is in phase with the stator MMF space phasor (Fs) and is given by - xPT = L,r+MMI, (46) - (47) Fs M C = MIse3(Wst--83); where I - - - N, - The space phasor LrZr represents the self flux linkage produced in the rotor windings owing to their induced currents I, as given by eqn. 34. This phasor is co-linear with the rotor MMF space phasor (Fr) and is - L T = L I $ ( W \" t - e s - + - % ) ; where z= N, Fr (48) r r T T The space phasor Y, is the resultant rotor flux linkage and is obtained from eqn. 46 as where - (49) Q/ r - q r e 3 ( W g t - - 8 . - # 9 It is important to notice from eqns. 48 and 49 that, as shown in Fig. 8, the resultant rotor flux-linkage space phasor is in space quadrature with the space phasor of rotor MMF. 4.3 Torque from rotor flux-linkage space phasors The rotor flux-linkage space-phasor diagram (Fig. 8) is identical to the time-phasor diagram for each phase which is represented by Fig. 3. The area of the triangle is 1/2Y&J,., and referring to eqn. 40 we make the important observation that the torque is proportional to the area of the flux-linkage triangle. In addition, we make some illuminating observations about the mechanism of steady-state torque production by reflecting on the physical significance of the space phasors. Fig.9 Resultant rotor flux-linkage and ampere-conductor distributions Under steady-state conditions at any frequency the rotor ampere-conductor distribution is seen to remain in space phase with the resultant rotor flux linkage Yr (Fig. 9). But Yr is notionally attributable to the resultant radial flux density wave at the rotor windings, and it follows that since torque can be expressed in terms of the integration of the rotor '3IZ' product over the periphery, the torque is simply proportional to the product of Y, and I,. Hence when the machine is operated so that the rotor flux linkage is kept constant, the rotor current wave is in the optimum position as far as torque production is concerned, and torque is directly proportional to rotor current (and slip), just as in a DC machine. This confirms eqn. 40 derived previously. This picture contrasts sharply with that which arises from the conventional voltage-fed analysis, where torque is pictured in terms of the interaction between an air-gap flux density wave and an induced rotor ampere-conductor wave, which, even at low slips, are never quite in space-phase. 4.4 Flux and torque components of stator current By resolving the stator flux-linkage phasor MIs into its IEE Proc -Electr Power Appl, Val 143, No 1, January 1996 64 components perpendicular and parallel to the resultant rotor flux-linkage phasor Yr (Fig. lo), the significance of the terms \u2018flux component\u2019 and \u2018torque component\u2019 of stator current (which are used widely in the literature of vector control) can be seen clearly. The torque component MI,, (which is directly proportional to the rotor current) can be regarded as being responsible for nullifying the demagnetising effect of the rotor (torque producing) current, leaving the flux component MIsy to set up the resultant rotor flux linkage. 5 Dynamic (vector) torque control 5.1 Locked rotor We have already shown how the secondary (rotor) current in one pair of coupled coils can be caused to jump from one steady-state condition to another by appropriate control of the stator currents. We have also shown in Section 3.2 that the locked-rotor torque is directly proportional to the rotor current if the resultant rotor flux linkage is kept constant. Hence, to cause a step change in the locked-rotor torque, one simply has to change the amplitude, frequency and phase of stator currents simultaneously in the manner indicated in Section 2.6 so as to cause the required jump in the rotor current. For example, let the motor be in the steady state at the frequency mSl and producing torque Tl and let the corresponding rotor flux-linkage phasor diagram be represented by triangle OABl in Fig. 11. Suppose we now want the torque to suddenly increase to a new value T2 which is twice the old torque. To double the torque (while keeping the resultant rotor-flux linkage constant), we need to double the rotor current (eqn. 40), so from Section 2.6 we see that we must simultaneously make the following abrupt changes: (U) double the stator frequency, i.e. increase it in the ratio of rotor currents (Ir2/Irl) (b) increase the amplitude of the stator currents in the ratio of rotor impedances (Zr2/Zrl) corresponding to new (oS2) and old (ql) stator frequency, and (e) increase the phase of the stator currents from corresponding to frequency (mSI) to @2 corresponding to the new frequency (as2). IEE Proc -Electr Power Appl , Vol 143, No. I , January 1996 den increase in the amplitude and speed of the stator MMF wave, together with an instantaneous forward jump in its position. 5.2 Constant speed In this Section we apply essentially the same ideas as for locked rotor to the running condition. We again regard the stator space phasor (MIs) as the independent variable and the remaining two space phasors (&Ir and Yr) are again derived from it. But whereas with the rotor stationary all three variables are at supply frequency (as) and all three space phasors rotate at synchronous speed (us), when the rotor is rotating at a constant angular speed (or) the relative velocity of the stator space phasor as seen by the rotor is cos,,, = os - or, i.e. the slip velocity. As far as the rotor is concerned, all three flux linkages vary at the slip frequency. Hence all the relationships governing rotor behaviour developed so far (in particular the torque) can be obtained simply by replacing the supply frequency by the slip frequency. (This contrasts sharply with the voltage-fed machine analysis, where it is not possible to follow such a simple line of reasoning because the stator MMF is not independent of the rotor speed.) Establishing the conditions for step change of torque under running conditions now becomes easy. For example, if the motor is running with a slip frequency of msr at torque T, and a step change to a new torque T = y Tis required, we apply the results from Section 2.6 to deduce that the following simultaneous changes must be made in the stator currents: (a) The frequency must be changed so that the slip frequency changes to yoSr, i.e. I iJ,,=y (51) u s , (b) The amplitude must be changed in the ratio of rotor impedances corresponding to frequencies yo, and a,,, i.e. (e) The phase must be changed by an amount equal to the difference between the rotor power factor angles corresponding to frequencies yo,, and os,, i.e. Of course, these changes will only yield a step increase of torque from one steady state to another if the rotor speed remains constant. If the rotor accelerates, it will clearly be necessary to alter the stator frequency and amplitude to maintain the torque constant, as the slip frequency will be changing. The currents and torque for an induction motor in which the stator currents are changed, are obtained by using a conventional d-q-axis machine model, are shown in Figs. 12-14. These results relate to a threephase four-pole cage induction motor with L, = L, = 61.lmH, M = 59mH and R, = 0.4Q rotating at speed of 300rev/min. The perfect step increase in torque confirms the validity of the space-phasor approach followed in this work. 65 20'01 5 0- 4.0- 3 0- 2.0 I I 10.0 4 0 -1 0.0 -20.0 -0.5 0.5 time, s Time variations of stator currents when all conditions for step Fig.12 change of torque are fulfilled 10.0 a o -1 0.0 -0 5 0.5 time, s Time variations of rotor currents when all conditions for step Fig.13 change of torque are fulfilled 6.01 1 I 5.3 Scalar torque control It has been stressed in the discussion that three conditions (magnitude, frequency and phase) must be satisfied by the stator currents to achieve ideal step changes in torque. This is equivalent to controlling the magnitude, speed and instantaneous position of the stator MMF wave. It is the inclusion of instantaneous position in the trio of conditions which gives rise to the term 'vector' control, and which differentiates it from scalar control. For the sake of interest Figs. 15-17 show the effect of making the same sudden changes in amplitude and 66 frequency as in Figs. 12-14, but without the change in phase. Comparison of the torque responses in Figs. 14 and 17 clearly indicates the superiority of vector over scalar control. * O ' O l 10.0 4 0 - 10.0 -20.0 -0.2 0.8 time, s Time variations of stator currents when step changes in amplitude Fig. 15 and frequency (but not phase) of stator currents are imposed 6 Experimental verification Three separate PWM current controllers were built to provide independent control over the stator currents of a standard three-phase, 3.0kW, 240/415V, 11.216.5A, 1420revlmin cage induction motor. Digitally stored sinewaves were used so that it was possible to make IEE Proc.-Electr Power A p p l , Vol 143, No I , January 1996 almost instantaneous changes to the magnitude, frequency and phase of the three current reference signals, and the power output stage was provided with a 500V DC link so that the loop-gain of each current controller was high and very rapid changes could be obtained in the stator currents. This is verified by Figs. 20 and 21, which show that the risetimes of the step changes in the stator currents are negligible. 6. I Steady state The rotor time constant was obtained from a lockedrotor, variable-frequency test with a constant stator current of 3A (RMS) per phase, the torque being measured with a torque transducer. The peak torque occurs when the normalised frequency is unity (0.71Hz), so the rotor time constant is 225ms. A series of steadystate tests with constant resultant rotor flux linkage were carried out under locked rotor and running conditions. The aim was to show that torque depends only on slip and rotor time constant, and hence to verify that under current-fed conditions the absolute speed of the rotor is unimportant. To avoid complications due to saturation, the resultant rotor flux linkage was limited to 75% of rated value, tests being conducted at speeds of 750 and 1500rev/min and with the rotor locked. The speed was held constant with a synchronous machine coupled to the motor shaft, and the stator current at each frequency was adjusted according to eqn. 17 to keep the resultant rotor flux-linkage constant. The results of these experiments are shown in Fig. 18, from which it is clear that, as predicted, the torque depends only on the slip, and not on the absolute speed of the rotor. (The minor variations which occur are probably due to the change of rotor resistance with temperature.) Normalised stator current and torque against slip frequency under Accordingly, all the tests were carried out with the rotor at rest (but not locked), and with torque applied by means of weights as shown in Fig. 19. One weight is attached to the motor shaft via a stiff rope, and applies a constant load torque. (It was important to minimise system inertia, so a large weight near to the axis was desirable.) The other weight is suspended by an electromagnet from the end of a horizontal arm fixed to the motor shaft. The current in the magnet is set so that this weight is only just suspended, so that when the current in the magnet is switched off, the weight immediately falls away and there is a sudden step reduction in the total load torque. The test procedure involves adjusting the motor current so that the motor torque is exactly equal to the total load torque, and then simultaneously de-energising the electromagnet and switching the amplitude, frequency and phase of the stator currents to the values calculated to balance the new (lower) torque. Ideally, the result of this experiment should be no movement of the rotor, because the torques before and after switching should exactly balance the load torques. To detect any acceleration (which will be proportional to net torque), an accelerometer [9] was fitted to the shaft. Typical results which correspond to sudden reduction in the load torque from 7.4\" to 2\" are shown in Figs. 20 and 21. These plots show the stator current in one phase, and the resultant torque (Tr) as derived from the accelerometer signal. 6.2 Open-loop torque control The ultimate test of the theoretical findings is to show that the motor produces sudden step changes from one steady-state torque to another. Measurement of torque under dynamic conditions is difficult, mainly because of the relatively low bandwidth of most torque transducers, but here we were able to take advantage of the fact that torque has already been shown to be independent of speed and therefore we can investigate dynamic torque control with the motor stationary. IEE Proc.-Electr. Power Appl., Vol. 143, No. 1. January 1996 Fig.20 Time variations of resultant rotor torque (acceleration) and stator current when only amplitude and frequency of stator current are suddenly changed i: 2A/div.; T: 2.4Nddiv. In Fig. 20 the amplitude and frequency of the stator currents are switched to the correct new values (to balance the new lower load torque after switching), but there is no sudden change in the phase. This is an example of scalar control and the motor torque therefore oscillates before settling down to its new steadystate value because there is no instantaneous change in the position of the stator MMF space phasor. 61 In contrast, in Fig. 21 the amplitude, frequency and phase of the stator current are all changed instantaneously to their new steady-state values, so that a transient-free step of the motor torque occurs at the instant that the load torque suddenly changes. This is confirmed by the fact that the rotor stays at rest, i.e. the acceleration signal remains at zero. This ideal (vector control) condition involves an instantaneous jump in the position of the stator MMF space phasor. Fig.21 to? current when amplitude, frequency and phase are all suddenly changed Time variations of resultant rotor torque (acceleration) and staz 2AId1v1 T 2 4\u201cidlv The way in which a cage induction motor can be made to produce sudden step changes in steady-state torque has been explored in ways which are intended to appeal to machines engineers who find the complexities of vector control schemes daunting. By considering the problem of open-loop torque control, the conditions for obtaining step changes in torque have been quantified in terms of two parameters. In essence, the stator MMF wave must be made to change its magnitude, slip and position instantaneously whenever a sudden change of torque is called for, thereby causing the rotor current to jump to a new steady-state value. Simple analytic formulas have been obtained for the changes which must be made to the stator currents to produce given step changes in torque. These results are not in themselves new, but have been presented in a way which should make understanding of the underlying mechanisms clear to those who find complex mathematics difficult. References LEONHARD, W.: \u2018Control of electric drives\u2019 (Springer-Verlag, 1985) YAMAMURA, S.: \u2018Spiral vector theory of AC circuits and machines\u2019 (Clarendon Press - Oxford, 1992) HO, E.Y.Y., and SEN, P.C.: \u2018Decoupling control of induction motor drives\u2019, IEEE Trans., 1988, IE-35, (a), pp. 253-262 ENSLIN, N.C., and VAN DER MERVE, F.S.: \u2018Improving the open-loop torque step response of induction motors\u2019, ZEE Pvoc. B, November 1987, 134, (6), pp. 317-323 HUGHES, A., CDRDA, J., and ANDRADE, D.A.: \u2018An inside look at cage motors with vector control\u2019,Proceedings of IEE conference on Electrical machines and drives, September 1993, Oxford, pp. 258-264 ANDRADE, D.A.: \u2018Dynamic control of inverter-fed cage induction motors\u2019. PhD thesis, Electronic & Electrical Engineering Dept., University of Leeds, April 1994 DIANA, G., and HARLEY, R.G.: \u2018An aid for teaching field oriented control applied to induction machines\u2019, ZEEE Trans., 1990, PWRs-4, (3), pp. 1258-1261 JONES, C.V.: \u2018The unified theory of electrical machines\u2019 (Butterworths, London, 1967) STEPHENSON, J.M.: \u2018Frequency-response analysis and parameter measurement of a DC-excited drag-cup tachogenerator\u2019, Prac. IEE, 1970, 117, (12), pp. 2301-2305 68 IEE Proc.-Electr. Power Appl., Vol. 143, No 1, January 1996"
+        ]
+    },
+    {
+        "image_filename": "designv11_60_0000050_iemdc.2015.7409124-Figure1-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0000050_iemdc.2015.7409124-Figure1-1.png",
+        "caption": "Fig. 1. PC fan: 1 \u2013 PM BLDC motor, 2 \u2013 housing, rear surface, 5 \u2013 Venturi set up, 6 \u2013 struts of a spide 8 \u2013 cylindrical exterior of the housing, 9 \u2013 leads [5].",
+        "texts": [
+            " The impeller blades are surrounded by a ring studded with 12 magnets, which are acted upon by four coils that're located at the corners of the housing of the fan. The tips of the blades can also be made of a hard magnetic material and magnetized in radial direction [9]. II. CONSTRUCTION OF PM BLDC FAN MOTORS The cost effective two-phase brushless motors for computer fans have a salient pole inner stator and ringshaped outer PM rotor. The outer PM rotor is integrated with the fan blades facilitating air flow [1 \u2013 4]. The housing is mechanically connected with the inner stator of the motor with the aid of a spider structure (Fig. 1). The details of construction of a PC fan motor are shown in Fig 2. A Hall sensor detects the polarity of PMs and via solid state devices switches the DC voltage from one stator coil to another. The speed of the fan motor is controlled by adjusting either the DC voltage or pulse width in lowfrequency PWM [10]. In spite the PM BLDC motor has four dead spots per revolution, it has god self-starting capability. Since the rotor rests between the poles of PMs at zero-current state Analysis of Steady-State and Transient Performance of Two-Phase PM Motors for Computer Fans J",
+            " S s of steady-state shows the current ng, Fig 11 shows locked rotor, and e at no load. The nvestigated motor ue; (b) input current 7 to 9) have been er input voltages, age slightly drops Fig. 7a, 8b) 13.2 0.32 10.5 0.22 8.55 0.15 le) with the load 12 and 10 V DC 4 V, 0.086 W for Fig. 10. Voltage and current waveform under maximum load of 0.36 Ncm. Te t power versus torque; (b) DC at constant DC voltage. s at starting: (a) at no-load; (b) st results. At starting, the steady-state voltage achieved after about 3.5 ms at no load (Fi 2.5 ms at maximum load 0.36 Ncm (Fig. 1 The total power losses in the moto controller are high (Fig. 9) and exceed m the maximum output power (Fig. 8a). IV. COMPARISON OF CALCULATIONS WIT Figs 12 and 13 show the comparis performance characteristics with the calculations of characteristics have b analytical approach. The efficiency includ and solid state controller losses. As expe and size of PM BLDC motors, the effic and does not exceed 10% at 14 V, 6.5% a 10 V (see also Fig. 9). ltage and current e and current curves and current is g"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0001946_0273-1177(92)90261-u-Figure2-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0001946_0273-1177(92)90261-u-Figure2-1.png",
+        "caption": "Fig. 2. Competition of phototaxis and physical torque:",
+        "texts": [
+            " Thus gravity, or shear, could influence the long-time average of metabolic activity of single swimming algal cells exposed to directional illumination. The effect is necessarily joint and dynamic, between gravity and/or shear and a sensory orienting stimulus. Gravity alone would not produce an effect. These considerations suggest an experiment for measuring sensory stimulus\u2014 response behavior in terms of an equivalent physical aligning torque. For example, shear or gravitational torque could be used to measure the \u201cstrength\u201d of phototaxis. Figure 2a shows a Chlamydomonas \u2014 like cell oriented at an angle 0 with respect to the z-axis. Light of intensity I shines along the minus x\u2014axis. A torque N, exerted by gravity and/or shear, is applied to the cell. Figure 2(b) shows an imagined plot of the functional dependence of M(9) vs. I, for fixed 0. In region A, weak light, the cell attempts to swim toward plus x, toward the light source. In region B, high light intensity, the cell tends to swim away from the light. In region C, very high intensity, the cell\u2019s response weakens. From plots like Fig. 2b, it should be possible to quantify taxis in physical units. GRAVITY REDUCES ENERGY OF TAXIS Gravitational orientation may affect the efficiency of taxis by an additional mechanism, one which results in a reduction of energy consumption. When cell orientation by gravity dominates fluctuations, upward swimming will be fairly smooth and direct, as in Fig. 3a. A weak transverse stimulus, due to illumination, for example, will deflect the trajectory, as in Fig. 3b. When a cell has travelled a distance L in the direction of the stimulus, its total travel distance will be approximately L/sina, where the angle a measures the orientational bias due to attraction and if there were no other stimulus, the cell\u2019s trajectory would appear randomly contorted as in Fig"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0000007_eleco.2015.7394494-Figure1-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0000007_eleco.2015.7394494-Figure1-1.png",
+        "caption": "Figure 1. 3-DofHelicopter System",
+        "texts": [],
+        "surrounding_texts": [
+            "1. Introduction\nAircraft control system has improved a lot with the recent technological advances in the computer science and improvements in the control theory. There are main difficulties in designing a controller for the aircrafts: parametric uncertainty and high nonlinearity of the aircraft are one of them. Fuzzy control is often used in order to overcome these kind of difficulties due to its nonlinear nature. It can provide an effective solution to the control ofplants which are complex and uncertain. Moreover, LQR controller can be used in order to control a system which has parametric uncertainty and nonlinearity.\nThere are many control methods in order to control 3-Dof helicopter. A fuzzy controller is designed in order to control the elevation axis [2]. Fuzzy logic controller and LQR controller is used in together in order to contro13-Dofhelicopter [1]. 3-Dof helicopter system is controlled by using a Pill controller, a fuzzy controller and LQR controller separately [4].\nIn this paper, fuzzy controller and LQR controller are compared. First of all, the linear quadratic regulator controller designed by Quanser firm is investigated and 3-Dofhelicopter is controlled by using the LQR controller. Then, three separate fuzzy Pill controllers have been employed so as to control three axes of the system. Finally, robustness of LQR controller and fuzzy controllers to model uncertainties are tested and the results are compared.\n2. 3-Dof Helicopter\nThe 3-DOF Helicopter system is shown in Figure l. The elevation, pitch and travel axes of helicopter are shown as E, 9 and ep respectively. The positions of three axes are measured by three encoders. The output of the system are positions of 3 axes and the inputs of system are voltages which are applied to front and back motors. The motors attach two propellers and drive them, so motor force is generated. The positions of 3-Dof helicopter is changed by moving the propellers. There is a counterweight on the helicopter in order to test parameter uncertainty. The nonlinear equation of elevation, pitch and travel axes are given below [5].\n2.1. Mathematical Model of Elevation axis\nFree body diagram of elevation axis can be shown in Figure 2. Direction of elevation axis is like to gravitational axis. Mathematical model of elevation axis is given in Figure 2. Je is moment of inertia of the system about elevation axis, La is distance from the middle point to the base point, Vf and Vb are the voltage applied to the front and back motor. The mathematical equation of elevation axis is given in equation (1).\nL leE = -(Mf +Mb)B~cos(e - Oa)cos Ua\nLc . +McB --Ii: cos(e + Oc) - 'YJee (1)cos Uc\n+KmLa(Vf + Vb)cos ()",
+            "velocity. Also, output of the fuzzy controllers are voltages which are applied to the motors.\n2.2. Mathematical Model of Pitch axis\nThe Ke, Kde, a and ~ are scaling factors of fuzzy controller. Takagi-Sugeno fuzzy model is used in order to control the system. A second order low pass filter is used before the derivative term in order to attenuate high frequencies.\n(4)\nwn 2 15791.37\nH(s) = S2 + 2(wns + wn2 = ...,s2=-+----=2-\",2-\",6....,.1-=-95=-s-+----=1\"::\"5=-79=-l:-.3=7\nIn table 1, the rule table of fuzzy controllers is given [6]. Defuzzification method is chosen as ''weighted average\".\nInput membership functions of fuzzy controller is determined as triangular shapes and output membership function of fuzzy controller is determined as singleton. (In Figure 7, NH: negative high, NL: negative low, Z: zero, PL: positive low and PH: positive high.)\nFree body diagram of pitch axis is shown in Figure 3. The pitch model symbol is (}. I (J is the moment of inertia of the system about pitch axis. Mathematical equation of the pitch axis is given in equation (2).\n.. Lh1(J(} = -MtO--l'-cos\u00ab(} - 8h ) cos Uh\nLh+MbO-l'-cos\u00ab(} + 8h ) (2)cos Uh\n-TJ(J8 +KmLh(Vt - Vb)\n2.2. Mathematical Model of Travel axis\n(3)\nFree body diagram of travel axis is shown in Figure 4. The elevation model symbol is cp. l<t> is the moment of inertia of the system about travel axis. Mathematical equation of travel axis is given below in equation (3).\nl<t>CP = -TJ<t>cP - KmLaCVt +Vb)sin (}\n3. Design of a Fuzzy PID Controller for 3-Dof Helicopter\nFuzzy Pill control mechanism obtained in order to control the system is shown in Figure 5 Three separate fuzzy Pill controllers are used in order to control elevation, pitch and travel axes of the helicopter. As it can be seen in Figure 6, inputs of Pill controllers are determined as position error and angular\nThe input universe of discourse for position error is [-25, +25], hence scaling factor of position error input (Ke) is 0.04[7]. Scaling factor of angular velocity input (Kde) is determined as 0.07 by doing some test on the simulation. Scaling factor of pitch and travel axes fuzzy Pill controllers for position error input (Ke) are found as 0.04 and 0.05 by doing experiment on the real system. Also, scaling factor of pitch and travel axes fuzzy Pill controllers for angular velocity input (Kdc) are found as 0.07 and 0.02 by doing experiment on the real system. The output scaling factors fuzzy Pill controllers are determined as a=20 and ~=15 for elevation axis, a=1O and ~=l for pitch axis and a=12 and ~=6 for travel axis."
+        ]
+    },
+    {
+        "image_filename": "designv11_60_0001235_b978-0-08-099425-3.00005-4-Figure5.22-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0001235_b978-0-08-099425-3.00005-4-Figure5.22-1.png",
+        "caption": "FIGURE 5.22",
+        "texts": [
+            "9 it is clear for a block of rubber, or tread material, there is more force required at any given displacement during the loading phase than the unloading phase. As tread material moves through the contact patch it will be loaded until it reaches the midpoint of the contact patch and unloaded as it moves to the rear of the contact patch. This and the additional losses due to hysteresis in the side walls leads to a pressure distribution that is not symmetrical as shown for the stationary tyre in Figure 5.13 and has a greater pressure distribution in the front half of the contact patch as shown Figure 5.22. The pressure distribution implies that the resultant tyre load Fz acts through the centre of pressure, a distance dx forward of the wheel centre. For equilibrium, a couple exists that must oppose the tyre load and its reaction acting down through the wheel centre. The couple that reacts the wheel load couple results from the rolling resistance force FRx acting longitudinally in the negative XSAE axis and reacted at the wheel centre where l z Rx R x\u03b4F F = \u00f05:36\u00de The rolling resistance may also be referenced by a rolling resistance coefficient, this being the rolling resistance force FRx divided by the tyre load Fz. By definition therefore the rolling resistance moment My is Fz dx and the rolling resistance moment coefficient is dx. Rigorous adherence to the sign convention associated with the tyre reference frame is essential when implementing these formulations in a tyre model. In Figure 5.22, to assist understanding, Fz is represented as the vertical force acting on the tyre rather than the negative normal force computed in the ZSAE direction. Generation of rolling resistance in a free rolling tyre. The rolling resistance force is very small in comparison with other forces acting at the contact patch, a rolling resistance of the order of 1% of vehicle weight being typical for a car tyre. This and the fact that the rolling resistance force may vary by up to 30% of the average value during one revolution (Phillips, 2000) make accurate measurement difficult"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0000494_fst15-249-Figure2-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0000494_fst15-249-Figure2-1.png",
+        "caption": "Fig. 2. Coating fixtures: (a) Be disk coating, (b) gold cone for Cepheus, and (c) sombrero tray for Cepheus sombreros.",
+        "texts": [
+            " By using AM to produce the assembly fixture for the neutron pinhole assembly, complex ribs could be designed to make the FUSION SCIENCE AND TECHNOLOGY \u00b7 VOLUME 70 \u00b7 AUGUST/SEPTEMBER 2016 structure rigid. Assembly features to hold the assembly at the right height could be incorporated with no increase in manufacturing cost. We also use AM to produce coating, assembly, and radiography fixtures. Coatings are often needed on beryllium substrates; the coating fixtures are designed to trap the parts to prevent beryllium contamination of the source while masking the portion of the substrate that is required to be coating free. Figure 2a shows a coating mask for multiple beryllium targets. The cone for Cepheus (Fig. 2b) is rapid prototyped and requires coating with gold and then Parylene. To avoid adding extra weight, the inside of the cone must not be coated. A custom coating fixture provides masking in addition to mounting features for the coating chamber. Cepheus also requires a large shield of aluminum that is pressed into a dimpled pattern and then Parylene coated. A small fixture was made to hold several shields in a position so they are coated on all sides while sitting in the chamber (Fig. 2c). Rapid-prototyped stalk structures are designed with boresight features (Fig. 3a), which are used during alignment to overlap toward a specific direction in the chamber. Great care must be taken during assembly to make sure the support structure is aligned with the axis of the pin, due to clearances between the inner diameter of the structure and the outer diameter of the glass pin. Rapidprototyped fixtures (Fig. 3b) were made to hold the alignment boresight of the new structure at the correct angle with respect to the stalk while it is being glued in place"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0003258_bf02508012-Figure1-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0003258_bf02508012-Figure1-1.png",
+        "caption": "Fig. 1. - The simply connected domain D1 of integrated conductivity ~1 bounded by the closed contour C.",
+        "texts": [
+            " - We shall begin by giving integral relations connecting the limiting values of the electric potential and the current function. The two functions V,~(x,y) and z~l U,~(x,y) are harmonic and satisfy the Cauchy-Riemann conditions in the simply connected domain Din. This implies that V l ( x , y ) + i z ~ l U l ( X , y ) is an analytic function of the complex variable = x + iy in D~(i = ~rL--1). From the Cauchy integral theorem we can set where 0 = x(s') + iy(s') is a variable point of integration on the contour C. From fig. 1 we have (15) (1 - ~ = rexp[iO], (*) One of the referees indicated an alternative derivation of the integral equation (22), which is slightly shorter than our derivation. His analysis is based on the double layer representation technique (see, for instance, ref. (5)). (5) R. COURANT and D. HILBERT: Methods of Mathematical Physics, Vol. 2 (Interscience Publishers, New York, London, 1962), p. 260. 528 M . S . ABOU-DINA and A. A. ASHOUR where r = 14 - P [ is the distance between the field point (x, y) inside D1, and the variable boundary point (x(s'), y(s')) on the contour C, and 0 = arg(0 - /)) ",
+            " (2), (3), (4) and (5), the boundary conditions for the modifying current system can be obtained as (20) U~(s)=U~(s), (21) 5\" 1 (Y~ (8) - Y~ (8 ) ) = (o\" 1 - fro) E(x(8) c o s ~ + y(s) sin a). Using (18), (19), (20) and (21), the function G(s) defined by (9) on C is found to satisfy the following integral equation: (22) G(s) + qo -- ~1 C~ G(s ' )~ logrds ' 2(z0- zOi(x(s) - - - cos ~ + y(s) sin a). ~r(zo + ~1) ~o + ~1 A GENERAL METHOD FOR EVALUATING THE CURRENT SYSTEM ETC. 529 Using (13) and applying the Cauchy-Riemann conditions to the function log(rexp[iO]) (fig. 1), differentiation of (22) with respect to s is found to give 0.0--0-1 d c~G(s,)~Ods,_2(0-o_~.~_l)i d=_(x(s)cosa+y(s)sina)\" (23) F(s) + 7:(0-0 + 0-1--------) ds 0-0 -{- 0\"1 aS Performing the integration in (23) by parts yields (24) F(s) 0\"0 - - 0\"1 (~ 8 7:~o ~--~1) ~ F(s') ~ logr ds' = 2 ( 0 . 0 - 0-1) 0-0 -{- 0-1 I ~ss (x(s) cos a + y(s) sin a). Equation (24) is a linear Freedholm integral equation of the second kind for the function F(s). The kernel of this integral equation is also given as (25) y ( s ) - y ( s ' ) 8 1 o g r = 8 0 = -~s arCtg x(s) _ x(s, ) "
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0001020_s12221-016-6478-8-Figure1-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0001020_s12221-016-6478-8-Figure1-1.png",
+        "caption": "Figure 1. Schematics of the hybrid structure textile (HST).",
+        "texts": [
+            " Copper fibers have been known to be possibly treated without any significant problems due to mechanical strength larger*Corresponding author: jykim@ssu.ac.kr than silver fibers. Quite fast condition, 9,000 rpm, could be applied for incorporating Z twist between 560D PU fiber and copper fiber. Figure 3(c) shows in a graphical manner the details of the fibers prepared by the method above. PET fibers (45 fila. and 150 den.) were used for covering those fibers followed by S twist at 9,000 rpm. The final HST has been shown in Figure 1. Various types of samples using HST have been prepared in order to measure data transmission characteristics in Figure 2 and Figure 3. SEM (CX-100S, Semicoxem Co.) and optical microscopy (OM) have been employed in order to measure diameter and twist numbers for a variety of samples. Since the crosssectional area of copper fibers has a great effect on their electrical resistance, copper fiber diameter and outer coating thickness have been separately measured and analyzed for further use. OM has been mainly used in order to examine the twist features of copper fibers such as twist number, twist angle and contact area"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0003685_itsc.1997.660569-Figure11-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0003685_itsc.1997.660569-Figure11-1.png",
+        "caption": "Figure 11 : Bond graph for the i.s.g. free-wheeling",
+        "texts": [],
+        "surrounding_texts": [
+            "2.6 Fourth gear model\nWhen the i.s.g. speed goes to zero in the 3-4 shift, the transmission is in the fourth gear. The bond graph model is depicted in Fig. 8. In the fourth gear, the r.s.g. free-wheels times faster than the turbine, and the r.c.g. rotates times faster than the turbine. Hence, the fourth gear range is called overdrive mode [ 101. Note that the fourth gear range is distinguished from the second gear range whether the r.s.g. free-wheels and the i.s.g. is held stationary or the r.s.g. is held stationary and the i.s.g. free-wheels.\n3. DONWSHIFT MIODELS 3.1 Two-to-one shift model The shifting down from second to first gear begins with the second clutch pressure off-going. The bond graph model of the 2-1 shift is identical to that of the 1-2 shift in [8] and [9]. The equations of motion as well as the kinematics and dynamic constraints are also identical.\n3.2 Third-to-second shift model 3.2.1 Torque phase. With the first clutch pressure on-coming, the third clutch pressure off-going, and the 1-2 band clutch pressure on-coming, the torque phase of the 3-2 shift is initiated. In the torque phase, the dynamic constraints on the third gear range are satisfied with the 1-2 band clutch pressure being not zero. The torque phase of the 3-2 shift is bond I i\nT I\nI r wCif RGi Rsr 1 Wsr graphed in Figure 10. K 1 Wt\nTt+ 1 1 h T t - 7 OF-TF- 1\nFigure 10: Bond graph for torque phase of 3-2 shift\nIn the reverse gear, the turbine torque is supplied to the i.s.g. by the first clutch, and the i.c.g. is held to ground by the reverse band, and the r,s.g. free-wheels. This word model is bond graphed in Figure 9.\n3.2.2 Speed phase. The speed phase model of the 3-2 shift is identical to the 2-3 shift model. In Figure 10, if the third clutch pressure decreases and the turbine is disconnected to the i.s.g., the shift model becomes as Figure 4.",
+            "3.3 Fourth-to-third shift model In the 4-3 shift, the i.s.g. rotates with being reduced. As mentioned in the 3-4 shift, the third clutch transmits power through the one-way sprag. Therefore, Figure 7 is the bond graph when the fourth clutch pressure is not zero and the speed of the turbine is equal to that of the i.s.g. If the fourth clutch pressure is not zero, and the speed of the i.s.g. is slower than that of the turbine, the bond graph for the 4-3 shift is in Figure 1 1. In Figure 1 1, it can be seen that both the r.s.g. and the r.c.g. are free-wheeling.\n4. SIMULATION RESULTS The developed model is implemented as a module in a powertrain simulation tool called, \u201c AUTOTOOL\u201d. For detailed simualtion reseults, see [ 1 11.\n5. CONCLUSION In this paper, a mathematical model of an automatic transmission was developed using the bond graph method. The dynamic and kinematic constraints, as well as the governing differential equations of motions for the 1-4 in-gear ranges, upshifts, and downshifts were developed. The addition of proper dynamic and kinematic constraints allows the correct determination of in-gear ranges and shift conditions. The shift command influences only the pressures of each clutch, and the proper range of the automatic transmission is determined by the dynamic and kinematic constraints internally in a simulation. This allows dynamically-correct simulations of automatic transmissions. The developed model can be used in the design of a hydraulic pressure controller,\nintelligent cruise controller, and vehicle platooning controller.\nReferences [l] L. M. Fisher, \u201cShifting auto gears by computer\u201d,\n[2] K. Glitzenstein and J. K. Hedrick, \u201cAdaptive Control of Automotive Transmissions\u201d, Proc. of American Control Con$, vol. 2, pp. 1849-1855, June 1990.\n[3] R. L. Anderson and R. L. Bierley, \u201cMeasuring Automatic Transmission Shift Performance\u201d, SAE paper 654465,1965.\nThe New York Times, Wednesday, March 7,1990.\n[4] A. K. Tucgu, K. V. Hebbale, A. A. Alexandridis, and A. M. Karmel, \u201cModeling and Simulation of the Powertrain Dynamics of Vehicles Equipped with Automatic Transmission\u201d, Proc. of Symp. on Simulation of Ground Vehicles and Transportation Systems, ASME Winter Annual Meeting, Anaheim, Dec. 1986.\n[SI A. M. Karmel, \u201cA Methodology for Modeling the Dynamics of the Mechanical Paths of Automotive Drivetrains with Automatic Step-Transmissions\u201d, Proc. of American Control Con$, vol. 1, pp. 279- 284, June 1986.\n[6] D. Hrovat, W. E. Tobler, and M. C. Tsangarides, \u201cBond Graph Modeling of Dominant Dynamics of Automotive Powertrains\u201d, presented at the ASME Winter Annual Meeting, New Orleans, Dec. 1984.\n[7] J. Runde, \u201cModeling and Control of an Automatic Transmission\u201d, S.M. thesis, Dept. of Mech. Eng., M.I.T., Jan. 1986.\n[8] D. Cho and J. K. Hedrick, \u201cAutomotive Powertrain Modeling for Control\u201d, ASME J. Dyn. Sys. Meas. Contr., vol. 11 1, no. 4, pp. 568-576, Dec. 1989.\n[9] D. Cho, \u201cNonlinear Control Methods for Automotive Powertrain Systems\u201d, Ph. D. thesis, Dept. of Mech. Eng., M.I.T., Dec. 1987.\n440-T4,2nd Ed., 1983. [ 101 GM Hydra-matic, Principles of Operation: THM\n1113 J. Kim, B. Y. Jeong, D. Cho, and H. Kim, \u201cAuTOTOOL, a PC-based Object-oriented Automotive Powertrain Simulation Tool\u201d, IEEE Conference on Intelligent Transportation Systems, Boston, MA, Nov. 1997."
+        ]
+    },
+    {
+        "image_filename": "designv11_60_0000913_j.ifacol.2016.10.638-Figure1-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0000913_j.ifacol.2016.10.638-Figure1-1.png",
+        "caption": "Fig. 1. Schematic diagram of lane change maneuver",
+        "texts": [
+            " Kyoto, Japan Copyright \u00a9 2016 IFAC 497 vergence is mostly guaranteed with a sufficiently narrow interval determined with an approximate solution to the equation. In this paper, minimization of the total vehicle force for the obstacle avoidance problem is reduced to finding the solution of one equation in one unknown. In order to estimate the initial guess for the root solving method, Chebyshev and least squares function approximations are performed. The dimensionless final time is obtained with high precision using Brent\u2019s method, and this leads to the determination of optimal control that is as precise as required. 2. PROBLEM FORMULATION Figure 1 shows a vehicle that moves on a straight road with initial vehicle longitudinal velocity vx0, and initial lateral velocity vy0, at a given position. The vehicle performs a lane change in order to avoid an obstacle blocking its forward path. The longitudinal distance to the obstacle is denoted by xf . The lateral distance at final time tf , is given as yf . For simplicity, the vehicle is treated as a particle with mass m. In this problem, total vehicle force Ft, is to be minimized for the lane change maneuver"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0001124_icrom.2016.7886769-Figure1-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0001124_icrom.2016.7886769-Figure1-1.png",
+        "caption": "Fig. 1: Kinematics of E-puck mobile robot.",
+        "texts": [
+            " Moreover, after simulation, the proposed algorithm is implemented practically on E-puck mobile robot and the results are demonstrated. In this section, fundamental components of the proposed system such as the governing equations on the kinematics of the mobile robot and constituted elements of the controller are discussed. The E-puck mobile robot actuators are able to receive velocity commands. Also, precise model of system is one of 978-1-5090-3222-8/16/$31.00 \u00a92016 IEEE 386 the prerequisite of RHC. As shown in Fig. 1, robots location in the main coordinate system can be expressed as bellow: X = [ xe ye \u03b8e ]T (1) where \u03b8e is the angle between robot\u2019s main axis (parallel with the wheels) and the horizontal axis. Time derivative of robot\u2019s location is: X\u0307 = [ x\u0307e y\u0307e \u03b8\u0307e ]T (2) Let\u2019s consider linear and angular velocities as control inputs and robot\u2019s velocity along the X and Y axis as outputs. The control equation can be expressed as: x\u0307e y\u0307e \u03b8\u0307e  =  (vr + vl) cos \u03b8e 2 (vr + vl) sin \u03b8e 2 (vr \u2212 vl) l  (3) where vr and vl are linear velocities of right and left wheels respectively, and l is the distance between them"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0002299_s0043-1648(96)07481-9-Figure1-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0002299_s0043-1648(96)07481-9-Figure1-1.png",
+        "caption": "Fig. 1. Mechanical setup of test rig.",
+        "texts": [
+            " [7] performed experiments on journal bearings, determining the temperature distribution on the bearing bush. Mitsui, Hori and Tanaka [8] performed experiments mapping the temperature on both the journal and the bush, whereas Andrisano [9] mapped the temperature distribution on the journal, as a function of circumferential position. In the present study, a test rig is designed for testing journal bearings with cooling and heating capabilities of the journal. The experimental results are compared with those obtained from a THD lubrication model proposed by the authors. 2. Test rig Fig. 1 shows the mechanical equipment of the test rig. The test bearing bush (1) floats on the journal (2) which is supported by preloaded ball bearings (3) (4). The test bush is a 3608 single bore bearing, made by rolling a thin sheet of aluminium onto a steel backing. The journal (5) is made from carbon steel. Load is applied downward to the bearing, by means of dead weights (6), applied through a linking mechanism (7) Journal: WEA (Wear) Article: 7481 (8) (9) (10). All seven wheels in the load mechanism are equipped with ball bearings, in order to reduce friction"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0001953_1.2833876-Figure5-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0001953_1.2833876-Figure5-1.png",
+        "caption": "Fig. 5 Qualitative explanation of stability and instability of a bouncing vibration",
+        "texts": [
+            " With a steady bouncing vibration, the contact point should be at the vicinity of the top or bottom of the waviness because the momentum transfer from the disk to the slider must be as small as the dissipation of the suspension damping. Other contact points may be possible only when they are a symmetrical pair of positive and negative veloc ity points as in the case of Fig. 2(c) . We can explain that the periodic vibration containing a top contact point is stable, whereas one that only contains bottom contact points is unstable by looking at Figs. 5(a) and (b). In Fig. 5(a) , the solid line shows the synchronous bouncing vibra tion whose collision point is at the top of the waviness. If the Downloaded From: http://tribology.asmedigitalcollection.asme.org/ on 02/20/2018 Terms of Use: http://www.asme.org/about-asme/terms-of-use Waviness frequency/;], kHz Fig. 3 Maximum bounce height versus waviness frequency Ith collision point deviates from the ith top by positive angle A6i, as shown with the dotted line, the bounce velocity de creases due to the negative velocity of disk surface at the contact point",
+            " If the deviation AS,+i is negative, as shown with the dashed line, then the ;' + 1th bounce velocity increases due to the positive surface velocity at the contact point and then the deviation A6',+2 becomes positive. Therefore, the deviated bouncing vibration converges to the solid line, because the condition of | A6i+i/A9i \\ < 1 holds. When AS, is negative, the same result can be easily obtained. However, provided that the synchronous steady vibration comes into contact with only the bottom of the waviness as shown in Fig. 5(b), the positive deviation AS, at the ith colh- Journal of Tribology sion from the ith bottom will results in a larger i + 1th positive deviation AS^+i, because of the increase in bounce velocity due to the positive velocity at the contact point. When the deviation AS, is negative, on the other hand, the i + 1th deviation AS,+i will increase further in a negative direction, as seen in Fig. 5(b). Thus I ASi+i/AS,| > 1 and then a synchronous vibration with only the bottom contact point is unstable. When a collision takes place at a bottom point between two top contact points, as in the case of a 1/2 period ratio vibration shown in Fig. 2(b), the bottom point, to some extent, will add a destabilizing effect to the bouncing vibration. However, a stable bouncing vibration will still be possible because the addi tion or subtraction of velocity at the bottom point is compen sated by the subtraction or addition of larger velocity at the top point"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0001235_b978-0-08-099425-3.00005-4-Figure5.45-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0001235_b978-0-08-099425-3.00005-4-Figure5.45-1.png",
+        "caption": "FIGURE 5.45",
+        "texts": [],
+        "surrounding_texts": [
+            "In order to obtain the data needed for the tyre modelling required for simulation, a series of tests may be carried out using tyre test facilities, typical examples being the machines that are illustrated in Figures 5.42 and 5.43. The following is typical of tests performed (Blundell, 2000a) to obtain the tyre data that supports the baseline vehicle used throughout this text. The measurements of forces and moments were taken using the SAE coordinate system for the following configurations: 1. Varying the vertical load in the tyre 200, 400, 600, 800 kg. 2. For each increment of vertical load the camber angle is varied from 10 to 10 with measurements taken at 2 intervals. During this test the slip angle is fixed at 0 . 3. For each increment of vertical load the slip angle is varied from 10 to 10 with measurements taken at 2 intervals. During this test the camber angle is fixed at 0 . 4. For each increment of vertical load the slip and camber angle are fixed at zero degrees and the tyre is gradually braked from the free rolling state to a fully locked skidding tyre. Measurements were taken at increments in slip ratio of 0.1. The test programme outlined here can be considered a starting point in the process of obtaining tyre data to support a simulation exercise. In practice obtaining all the data required to describe the full range of tyre behaviour discussed in the preceding sections will be extremely time consuming and expensive. The test programme described here does not, for example, consider effects such as varying the speed of the test machine, changes in tyre pressure or wear, changes in road texture and surface contamination by water or ice. The testing is also steady state and does not consider the transient state during transition from one orientation to another. High Speed Dynamics Machine for tyre testing formerly at Dunlop Tyres Ltd. Most importantly the tests do not consider the complete range of combinations that can occur in the tyre. The longitudinal force testing described is limited by only considering the generation of braking force. To obtain a complete map of tyre behaviour it would also, for example, be necessary to test not only for variations in slip angle at zero degrees of camber angle but to repeat the slip angle variations at selected camber angles. For comprehensive slip behaviour it would be necessary at each slip angle to brake or drive the tyre from a free rolling state to one that approaches the friction limit, hence deriving the \u2018friction circle\u2019 for the tyre. Extending a tyre test programme in this way may be necessary to generate a full set of parameters for a sophisticated tyre model but will significantly add to the cost of testing. Obtaining data requires the tyre to be set up at each load, angle or slip ratio and running in steady state conditions before the required forces and moments can be measured. By way of example the basic test programme described here required measurements to be taken for the tyre in 132 configurations. Extending this, using the same pattern of increments and adding driving force, to consider combinations of slip angle with camber or slip ratio would extend the testing to 1452 configurations. In practice this could be reduced by judicious selection of test configurations but it should be noted the tests would still be for a tyre at constant pressure and constant speed on a given test surface. Examples of test results for a wider range of tyres and settings can be obtained by general reference to the tyre-specific Flat Bed Tyre Test machine. (Courtesy of Calspan.) publications quoted in this chapter and in particular to the textbook by Pacejka (2012). For the tyre tests described here the following is typical of the series of plots that would be produced in order to assess the force and moment characteristics. The results are presented in the following Figures 5.44e5.53 where a carpet plot format is used for the lateral force and aligning moment results: 1. Lateral force Fy with slip angle a 2. Aligning moment Mz with slip angle a 3. Lateral force Fy with aligning moment Mz (Gough Plot) 4. Cornering stiffness with load 5. Aligning stiffness with load 6. Lateral force Fy with camber angle g Lateral force Fy with slip angle a. (Courtesy of Dunlop Tyres Ltd.) 7. Aligning moment Mz with camber angle g 8. Camber stiffness with load 9. Aligning camber stiffness with load 10. Braking force with slip ratio Aligning moment Mz with slip angle a. (Courtesy of Dunlop Tyres Ltd.) Lateral force Fy with aligning moment Mz (Gough Plot). (Courtesy of Dunlop Tyres Ltd.) Before continuing with the treatment of tyre modelling, readers should note the findings (van Oosten et al., 1999) of the TYDEX Workgroup. In this study a comparison of tyre cornering stiffness for a tyre tested on a range of comparable tyre test machines gave differences between minimum and maximum measured values of up to 46%. Given the complexities of the tyre models that are described in the following section the starting point should be a set of measured data that can be used with confidence to form the basis of a tyre model. Cornering stiffness with load. (Courtesy of Dunlop Tyres Ltd.)"
+        ]
+    },
+    {
+        "image_filename": "designv11_60_0000200_chicc.2014.6896402-Figure1-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0000200_chicc.2014.6896402-Figure1-1.png",
+        "caption": "Fig. 1: A three-DOF planar space manipulator",
+        "texts": [
+            " The boundedness of \u0394x\u0308 shows that \u0394x\u0307 is uniformly continuous, and using Barbalat\u2019s Lemma [28, p. 122], we obtain \u0394x\u0307 \u2192 0 as t \u2192 \u221e. Remark 2. Here we assume that the estimate of the inertia matrix of the spacecraft H\u0302b is positive definite and the estimate of the extended Jacobian G\u0302 is of full row rank, which can possibly be guaranteed by the parameter projection algorithm [30]. In this section we present simulation results for the proposed adaptive control law via a three-DOF planar space manipulator (Fig. 1). In this simulation, both the spacecraft attitude regulation and the end-effector trajectory tracking are required. Without loss of generality, the desired value of the spacecraft attitude is set to be zero. In this simulation, the desired trajectory of the end- effector is given by xd = [ 3.7 + 0.3 cos(\u03c0t) 0.2 + 0.3 sin(\u03c0t) ] . The initial state of the 3-DOF space manipulator is as follows: the position of the center of mass of the spacecraft is RC0 = [ 0 0 ]T , the initial configuration of the FFSM is q(0) = [ 0 \u03c0/3 \u2212 2\u03c0/3 \u03c0/3 ]T , and the initial position of the FFSM end-effector is x0 = [ 3"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0000175_icamechs.2014.6911582-Figure1-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0000175_icamechs.2014.6911582-Figure1-1.png",
+        "caption": "Fig. 1. Relation between the inertial system and the quadcopter coordinate system.",
+        "texts": [
+            " However, in conventional MPC, because the prediction input to a reference trajectory is computed definitely, the probable element of a prediction horizon cannot be taken into consideration. Thus, in this paper, PF-MPC (particle filtermodel predictive control) [5], [6] that considers the abovementioned uncertainly is introduced for the predictive control of a quadcopter. This method improves the stability of the quadcopter system and prevents the generation of an excessive control input. II. DYNAMICS OF QUADCOPTER The relation between the inertial system \u03a3O and the quadcopter coordinate system \u03a3r is shown in Fig. 1. Here, the upper part of the z axis in \u03a3r is positive. The posture of \u03a3r relative to \u03a3O is represented by rotation transformation as follows: RO r (\u03b8z, \u03b8y, \u03b8x) = Rz(\u03b8z)Ry(\u03b8y)Rx(\u03b8x) (1) where \u03b8y , \u03b8r, and \u03b8p represent the yaw, roll, and pitch angles, respectively. Moreover, Rz(\u03b8z), Ry(\u03b8y), and Rx(\u03b8x) represent the rotation transformation around the z axis, y axis, and x axis, respectively. The shift vector from the origin of \u03a3O to the origin of \u03a3r is represented as z(t) [x(t) y(t) z(t)]T . The quadcopter has four propellers, the angular velocity of the i-th propeller is expressed as si(t), and the perpendicular upward thrust can be expressed as follows: fi(t) = bs2i (t) (i = 1, 2, 3, 4) (2) where b is the thrust constant"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0001035_iecon.2016.7793572-FigureI-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0001035_iecon.2016.7793572-FigureI-1.png",
+        "caption": "Fig. I. Flux density distribution of TPMSM with the intertum short fault in the upper part (red s:.juare) and shows less flux than other parts, given by FEA simulation",
+        "texts": [],
+        "surrounding_texts": [
+            "Rf Resistance that models insulation degradation 11 Ratio of shorted-turns to total turns in a faulty winding Vf Voltage drop of shorted turns if Circulating current in shorted turns l/JPM Flux linkages of permanent magnets (PMs) e Electrical angular position Rs Resistance of stator coils L Self-inductance of stator coils M Mutual-inductance of stator coils Lsm Static component of L Lsi Leakage component of L Vabc Stator voltage vector iabc Stator current vector N Number of windings per phase Lf Fault inductance component y Coupling factor of windings Va,h a-phase voltage in healthy condition Lan Nth winding in phase a I. INTRODUCTION Permanent magnet synchronous machine (PMSM) is widely used in many applications due to its many advantages such as high efficiency and high power to volume ratio. These properties make it an attractive candidate for many industrial applications, for example, electric and hybrid vehicles [1]. As a consequence of its popularity, many researchers have studied fault diagnosis of PMSM recently to maintain its high performance and efficiency during the past few years [2]. The fault diagnosis of the PMSM has emerged as a point of concern at the same time since several faults may degrade overall system's performance and even result in catastrophic damages. Numerous types of faults can occur in PMSM such as bearing faults [3], electrically short/open circuit faults in stator windings [4], eccentricity faults [5], and demagnetization fault [6]. Among the various faults, the intertum short fault is one of the most popular electrical faults [7]. The interturn short fault is a short circuit between the turns of a stator winding caused by insulation degradation [8]. The degraded insulation is usually modeled as a fault resistance (Rf) and a fault tum ratio (11). Such insulation degradation induces a circulating current and produces excessive heat. The thermal stress caused by excessive heat may lead to a more severe insulation degradation, which means a severe intertum short fault. It also creates a reverse magnetic flux (Fig. 1) which results in lower torque than healthy condition and may lead to a demagnetization fault at last. Therefore, it is necessary to detect the interturn short fault more accurately in early stage. There have been many researches to diagnose the interturn short fault including motor current signature analysis (MCSA) [9]-[12], model-based analysis [l3]-[14], and artificial intelligence [15]. MCSA is a conventional and the most often used method to diagnose the fault based on the extraction of signal features from stator winding currents. Many researchers have studied the interturn short with MCSA such as fast Fourier transform [9], short time Fourier transform [10], and wavelet transform [11]. These approaches do not need extra sensors to apply with, so they can be easily applied to various types of faults and motors [12]. However, the frequency analysis cannot diagnose the degrees of failure. They also have a weakness in distinguishing different faults which have similar harmonic component in specific frequency areas. There are other approaches to diagnose the intertum short fault. For example, a zero sequence method which suggested a fault index not affected by the rotor speed [l3] and model-based approach with Kalman Filter [14]. But they dealt with either fully-shorted condition or very severe fault condition, which may already result machine failure or system breakdown. A neuro-fuzzy approach was also tried to diagnose the fault [15], but it usually demands a lot of data to train and diagnose the fault. 978-1-5090-3474-1/16/$31.00 \u00a92016 IEEE 1513 This paper addresses a fault diagnosis based on model based analysis with a fault index based on a rearranged fault model. The intertum short fault model is already analyzed by Romeral et al [16] and Bon-Gwan Gu et al [8]. We analyzed and simplified those fault models and rearranged the model so that it can be applied to diagnose the intertum short fault with simple form. The least square (LS) method is used for calculating the fault index. We verified the diagnosis method with Finite Element Analysis (FEA) tool, Maxwell. Our proposed method shows satisfactory result of fault diagnosis with various fault cases. It worked well with weak faults, which mean lower degradation of insulation (bigger Rf) and less fault tum ratio than conventional approaches. This paper is organized as follows. In section II, the intertum short fault model is described. We also introduce the rearranged fault model. In section III, we suggest the fault diagnosis scheme with the fault index based on the rearranged fault model. In section IV, the simulation environment with FEA tool Maxwell is introduced. We also give simulation result with the proposed fault diagnosis method. Finally, we present our conclusion in section V. II. INTERTURN SHORT FAULT MODEL OF PMSM A. Series Connected PMSMwith the lnterturn Short Fault An electrical model of wye connected PMSM in the stationary (abc) frame is given as [Vabcl = [Rs][iabcl + :t ([Ls][iabc] + [1/JPM,abel) (1) where [Vabcl = [Va Vb Vc]t, [iabcl = [ia ib ic]t, [1/JPM,abc] = [RS 0 01 1/JPM[COSe cos(e - 21f) cos(e + 21fw, [Rs] = 0 Rs 0 , 3 3 0 0 R [L , ] \ufffd [\ufffd \ufffd 71 ' If the intertum short fault occurs in phase a of the PMSM, another closed loop circuit is added in whole PMSM circuit system (Fig. 2) and fault terms are added to (1) as following equation [16]. [Vabcl = [Rs][iabc] + :J[Ls][iabc] + [1/JPM,abcD - !!:. [RsJ [Ai] if -!!:...:!.. ([L M MFif) (2) N N dt where [Ai] = [10 OF. (2) shows an influence of fault current of intertum short on phase voltages Va' Vb' and Vc. The work of Romeral et al. described equation which shows relationship between fault voltages and fault current at the short part as follows t d Vf = Rfif = I1Rs(ia - if) + 11 [Lf] dt [is,abel -112 L..:!.. if + 11'!!.-(1/JPMcose) dt dt (3) However, they did not consider multi-pole structure of the PMSM, which usually has series or parallel windings. Therefore, we adopted a representation of the relationship of fault part from [8] since they considered series windings and analyzed fault phenomenon more accurately. B. Rearranged lnterturn Short Fault Model Based on the work of Bon-Gwan Gu et at. [8], we modified the representation of the fault equation as in [17]. ( 1 1) 1 Lf = 1 - -- L + (1 - -)L I 1-y N sm N s The first row of (2) can be rewritten as v: - V h = -!!:.R if -!!:''':!''(Lif) a a, N s N dt (4) (5) (6) where Va,h is a-phase voltage without the intertum short fault, in healthy condition, same as the first row of (1). Assume that the coupling factor of winding is ignorable, y \"\" 0, we can rearrange (4) and get the following equation through a series of calculation. (7) The left side of (7) is equal to the voltage of one winding (Lan' Fig. 2) which have the intertum short fault. The rearranged model fits well with the fault condition (Fig. 3). III. DIAGNOSIS SCHEME OF THE INTERTURN SHORT FAULT A. Ideal Fault Indices for the Interturn Short Fault There are two major parameters affect the severity of the intertum short fault. One is the degraded insulation Rf, which directly indicates the fault severity since smaller Rf induces bigger fault current in short circuit. The other is a fault tum ratio 11 and it also affects the characteristics of the fault current. These values are independent of stator currents, rotational speed and other operation conditions. Therefore, these two parameters are the ideal indices to determine the intertum short fault severity. Unfortunately, the intertum short fault model is not sufficient to determine those two parameters uniquely. Usually we cannot measure the fault voltage Vf or the fault current if directly in real applications, and we can only estimate these parameters through (2). B. Proposed Fault Indexfor the Interturn Short Fault Because it is not possible to estimate the ideal fault indices separately, we suggest a new fault index to diagnose the intertum short fault. The proposed fault index makes us possible to distinguish the severity of the fault. The equation (7) can be rewritten as . ((Rf 1 ))-1 N-l If= -;+(l-l1)iVRs (Va-iVVa,h) (8) The term in the right side of (8), c: + (1-11 ) \ufffd Rs) can be represented as fault impedance, Rfault impedance since it shows a relationship between fault current and fault winding's voltage. We can substitute fault current if with (8) in the first row of (2) as following equation v: -V h = -!!:.-R if -!!:.-\ufffd (Lif) a a, N s N dt 1 J1 d -- (RsVa fw + -(LVa fw)) (9) N Rfault impedance ' dt ' N-l where Va fW = Va --Va h' which has a same value of the , N ' faulty winding's voltage. Here, we declare a fault index by /1 = ((R\ufffd + (2:. _ 1) 2:. Rs)) -l since its value is R fault impedance /1 /1 N determined by only the fault conditions, Rf and 11 unless Rs and N do not change. The term -\ufffd Rsif -\ufffd :J Lif ) indicates the residual component in voltage due to the intertum short fault. The introduced fault index is zero without the intertum short fault and it becomes bigger according to the increasing severity of the fault, in other words, smaller Rf or higher 11. C. Diagnosis of the Interturn Short Fault with fault index The intertum short fault is diagnosed using the residual component of the voltages. The phase with the fault among the three phases (abc) is determined by comparing the fault index from the voltage residual component of each phase. The biggest one has the intertum short fault. The severity of the fault is directly got from the value of the fault index. The fault model of the PMSM with intertum short in phase a is described as (2). If the fault occurs in phase b or c, it changes as following equations. [Vabc] = [Rs][iabc] + :t ([Ls] [iabc] + [1/JPM,abcD - \ufffd[Rsl[Az]if -\ufffd\ufffd([M L M]tif) (10) N N dt [Vabcl = [Rs][iabcl + :t ([Ls] [iabcl + [1/JPM,abcD - \ufffd[Rs][A3]if -\ufffd\ufffd([M M L]tif) (11) N N dt [Az] = [01 O]t and [A3] = [00 l]t. We can represent the rearranged fault model with each case and get following equations from (8) and (9) Vb -Vb,h = -\ufffdRsif -\ufffd :t (Lif) /1 1 d -----(RSVb fW +-(LVb fw)) (12) N Rfault lmpedance ' dt ' v: -V = -!!:.-R if -!!:.-\ufffd(Lif) c c,h N s N dt /1 1 d = -- (RsVc fw +-(LVc fw)) N R fault impedance ' dt ' (13) . N-l N-l . WIth Vb,fw = Vb -iV Vb,h and Vc,fw = It;; -iV Vc,h' Vb,h IS b-phase voltage without the intertum short fault, in healthy condition, same as the second row of (1) and Vc,h is same as the third row of (1). We can easily get the fault index from (9), (12), and (13) with the application of least square (LS) method. As we assumed the intertum short fault with only one phase and the fault phase usually have bigger residual component than other phases, comparing the fault index from those three equations will show which phase has the fault. The fault index indicates the severity of the intertum short fault. Proposed diagnosis method could figure out the degree of the intertum short fault successfully with FEA simulation result and we will represent it in the next section. IV. STMULATTON RESULTS IPMSM drive system was built in FEA tool, Ansoft Maxwell for PMSM implementation and Simplorer for inverter implementation and control the PMSM (Fig. 4). The PMSM was operated with Maximum Torque per Ampere (MTPA). Specification of the IPMSM used in the simulation is listed in Table I and structure of the motor is given in Fig. 1. The motor maintained 3000 rpm during the simulation with varying torque from 1 to 4.5 Nm (Fig. 5). Sampling frequency of the simulation was 50 kHz. Various fault conditions and healthy PMSM have been run with the Maxwell and the Simplorer. We gave a variation of fault resistance and fault tum ratio to verify the proposed algorithm (Table II). With our notation of f.1=20, as there are 6 series windings, it has a value of 3.3% short turn ratio in whole windings of one phase. The result of the fault diagnosis is given in Table III. The proposed method was applied to the simulation results given by Maxwell, and the fault indices were estimated by applying the LS method to (9), (12), and (13). The fault index of phase a is the biggest among the values of those three in all cases except fault case 1, which is a normal case. From the result, we can conclude that the proposed algorithm can detect the phase with the interturn short fault well. The severity of the fault is determined by the value of the fault index. The ideal value of the fault index is calculated from the condition of the interturn short fault, Rt and f.1. Comparing the fault index given by the proposed method with an ideal value of the fault index shows satisfactory result with fault case 3\ufffd6. With the proper threshold, we can determine whether the PMSM has a fault or not. In these cases, we set the threshold as 0.05 to distinguish the healthy condition (a green line, Fig. 6). The threshold is determined according to the intertum short fault in our simulation environment because the weakest fault has a value of the fault index as 0.099 (Table III). However, the estimated result of the fault case 2 was not close to ideal value of the fault index and it results in misdiagnosis (Fig. 6). The ideal fault index is over the threshold and it has to be determined as having fault, but estimated value is lower than the threshold. The proposed scheme has to be updated to deal with much weak fault cases. V. CONCLUSION In this paper, a PMSM intertum short fault diagnosis method has been proposed. With the rearranged fault model, a fault index was introduced to identify the severity of the intertum short fault. Applying the LS method with the rearranged fault model showed accurate estimation result in most cases. We could figure out which phase has a fault and the severity of the fault at the same time. FEA simulation results demonstrate that the proposed method works well."
+        ]
+    },
+    {
+        "image_filename": "designv11_60_0003166_0168-874x(90)90028-d-Figure5-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0003166_0168-874x(90)90028-d-Figure5-1.png",
+        "caption": "Fig. 5. Finite element models used in present study.",
+        "texts": [
+            " Linear interpolation functions were used for approximating each of the stress resultants and strain components, and quadratic Lagrangian interpolation functions are used for approximating each of the generalized displacements. The integrals in the governing equations were evaluated using a two-point Gauss-Legendre numerical quadrature formula. Because of the symmetry of the shell meridian and loading, only one half of the tire meridian was analyzed. The finite element models used are shown in Fig. 5. To assess the accuracy of the shell model of the tire, the deformations produced by uniform inflation pressure of Po = 2.2063 x 10 6 Pa, acting normal to the inner surface, were calculated the outer surface. 226 A.K. Noor et aL / Analysis of aircraft tires (nonorthotropic) stiffness coefficients. Ero = 8 \u00d7 10 6 Pa. A.K. Noor et al. / Analysis of aircraft tires ~ = 0 = h i 227 using the geometrically nonlinear shell theory. Twelve finite elements were used in modeling half the cross section (a total of 384 strain parameters, 384 stress-resultant parameters, and 243 nonzero generalized displacements--see Fig. 5(a)). Comparison was made with the experimental data obtained on the shuttle nose-gear tire (see Fig. 2). The results are summarized in Figs. 6 to 8. Close agreement between the predicted deformations and experimental results is demonstrated in Fig. 6. Figures 7 and 8 show the meridional variations of the generalized displacements, stress resultants and strain energy densities. As can be seen from Fig. 8, for the case of inflation pressure, the transverse shear strain energy density is considerably smaller than the extensional/bending energy density",
+            " The normal loading in pascals is given by the following equations, which model experimental data obtained at NASA Langley on the shuttle tire: , o - - Y'~p, cosnO, - 0 . 2 < ~ < 0.2, P = n = l O, I~1 >0 .2 , (24) where p, = (2po/n 'n) sin nil, and po and fl are functions of ~ as shown in Fig. 9. Because of the symmetry of the shell meridian and loading, only one-half of the meridian is analyzed using 37 elements (a total of 1184 stress-resultant parameters, 1184 strain parameters, and 743 nonzero displacement degrees of f reedom--see Fig. 5(b)). The boundary conditions at the centerline are taken to be the symmetric or antisymmetric conditions. Typical results are presented in Figs. 10 and 11 and in Tables 3 and 4. The foregoing procedure was applied to this problem, and 10 global approximation vectors were evaluated at n o = 5 and used to generate the tire response in the range n = 1 to 10. Accuracy of the generalized displacements obtained by the foregoing strategy with 8, 10 and 15 global approximation vectors is indicated in Figs"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0000392_11663_2015_5001-Figure1-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0000392_11663_2015_5001-Figure1-1.png",
+        "caption": "Fig. 1 (A) Schematic representation of a tBLM formed on a mixed self-assembled monolayer of tether (cholesterol) and spacer molecules (6-mercaptohexanol). (B) Cole\u2013Cole plots for cbo3 before and after formation of a tBLM measured at 0 V vs. SCE. Adapted with permission from [35]. Copyright (2006) American Chemical Society",
+        "texts": [
+            " In case of tBLMs, the space between the electrode and the lower leaflet of lipid bilayer was used for the incorporation of a hydrophilic spacer. The spacer in turn avoided the direct interaction of proteins with the electrode surface and their subsequent denaturation. The tBLM approach was successfully used to immobilise cytochrome bo3 (cbo3), a ubiquinol oxidase from Escherichia coli [35]. The tBLM was formed on gold surface functionalised with cholesterol tethers which inserted itself into the lower leaflet of the membrane (Fig. 1A). 6-Mercaptohexanol was used as the hydrophilic spacer. The planar membrane architecture was formed by selfassembly of proteoliposomes, and its structure was characterised electrochemically by EIS. Normally, the double-layer capacitance of an ideal phospholipid bilayer should be around 0.5 \u03bcF cm 2. A drastic increase in the capacitance is generally caused by some disorder or defect in the bilayer. Incorporation of a protein also increases the capacitance values, as a result of the higher dielectric constant and disorder in the bilayer. As shown in Fig. 1B, the double-layer capacitance of the tBLMs on gold electrode was around 0.7\u20130.8 \u03bcF cm 2 which was only slightly larger than that of the ideal and was almost the same both in the absence (figure not shown) and presence of cbo3. These results showed that the inclusion of cbo3 had almost no effect on the double-layer capacitance of the tethered membrane and did not induce large defects in the tethered bilayer. The functionality of tBLMimmobilised cbo3 was investigated by CV and was confirmed by catalytic reduction of O2"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0003827_robot.2000.846348-Figure4-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0003827_robot.2000.846348-Figure4-1.png",
+        "caption": "Fig. 4: Rotation of the hand-object system",
+        "texts": [
+            " As we increase the rotational angle, we may keep a neighborhood equilibrium. However, for a particular rotational angle, the system finally results in the state where no more equilibrium is guaranteed. Since z direction of the base coordinate system coincides with the direction of gravity, the rotation around the z axis brings no effect on the neighborhood equilibrium. Therefore, we rotate the hand-object system for a[rad] around a unit vector k lying on the x - y plane of the base coordinate system as shown in Fig.4. Due to this rotation, the potential energy of the objects is rewritten as follows: m i=l k:V, +C, kxkzV, - k,S, kxk V, - k,S, kvkzV, i- kxS , kxkyV, + k,S, kyV, I + C, sincr, k = [kx ky kZlT = [cos$ sin$ 0IT. Using this potential energy, eqs.(l4) and 15 can be computed. Then, we obtain the maximum \\~Q!I which satisfies the neighborhood equilibrium as amax( k ) . Since crmax(k) depends on k , we can consider the robustness of the grasp by examining the minimum of amaX(k) with respect to the direction of k "
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0002229_jsvi.1996.0017-Figure18-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0002229_jsvi.1996.0017-Figure18-1.png",
+        "caption": "Figure 18. Poincare\u0301 map for b=bc=0\u00b78842 with a=ac=0\u00b77338, d=0\u00b71, v=4\u00b73, G=50.",
+        "texts": [
+            " However, what relations (15) and (16) do show us is that for a given v the trapping region for the homoclinic manifolds remains extremely narrow and consequently stable which suggests that they could withstand a far greater forcing amplitude than G=20 before losing stability, and moreover, the heteroclinic manifolds show a much larger deviation from their mean and will with forcing amplitudes only slightly higher (around G=50 derived from numerical investigations) than G=20, lose their stability and begin to interact with the homoclinic manifolds. If we increase the forcing amplitude further to G=50, say, and correspondingly increase the forcing frequency to v=4\u00b73, according to the behaviour v0ln G implied by (15) and (16), then the Poincare\u0301 map shows some of the features displayed in Figure 18. Here the homoclinic manifolds have collapsed into two narrow bands in the Poincare\u0301 section which show signs of long period behaviour. A typical quasi-periodic solution associated with the right-hand homoclinic section plotted over 150 periods after a transient of 20 000 periods is shown in Figure 19: the outer boundary of the solution is clearly visible. The heteroclinic manifolds, on the other hand, continue to give rise to a stable attractor which does not interact with any homoclinic manifolds"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0000743_gt2016-56592-Figure17-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0000743_gt2016-56592-Figure17-1.png",
+        "caption": "Figure 17 COMPARISON OF MERIDIONAL VELOCITY CONTOURS AT SECTION IV ( / 1.00mx S  , FIGURE 7)",
+        "texts": [
+            " The flow diffusion in the optimized impeller is much smoother than in the baseline impeller, indicating an improvement of flow field in the former. 9 Copyright \u00a9 2016 by ASME Downloaded From: http://proceedings.asmedigitalcollection.asme.org/pdfaccess.ashx?url=/data/conferences/asmep/89477/ on 02/05/2017 Terms of Use: http://www.asme.org/about-asme/terms-of-use The comparison of meridional velocity contours for the de- sign operating condition at impeller exit (section IV, / 1.00mx S  in Figure 7) is shown in Figure 17. The jet-wake pattern can be seen near the SS in the both initial and optimal impellers, but the wake region of the optimal impeller is significantly reduced from that of the baseline. This is a further indication of flow field improvement after the optimization and the mixing loss in the downstream diffuser will be reduced. The comparison of entropy contours (reference point 94450Pa, 303oK) and streamlines (white curves) at sections I ( / 0.25mx S  ), II ( / 0.50mx S  ), III ( / 0.75mx S  ) and IV ( / 1"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0000114_0954406216640806-Figure3-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0000114_0954406216640806-Figure3-1.png",
+        "caption": "Figure 3. The force distribution along the longitudinal profile of a wrinkle: (a) the curved profile is simplified into a straight line, (b) the two instantaneous profiles FV and F0V0, and (c) F overlaps with F0 in the two instantaneous profiles.",
+        "texts": [
+            " Then, the expression for the relationship between the hydraulic pressure and main forming parameters is established based on the energy conservation law. Finally, the analytic formula that a useful wrinkle needs to meet is derived using the relationship expression. In this paper, the prediction approach is named the mechanics-based prediction method, or MPM. A curved longitudinal profile of a wrinkle shown in Figure 2 is selected to be analysed here. The longitudinal profile of the deformed tube is simplified into a straight yield line FV,12 as shown in Figure 3. In the next moment of time, FV moves to F0V0, remaining a straight line. For the sake of a convenient analysis, the following assumptions are made with the present method: 1. The tubular material is homogeneous, isotropic, and incompressible, and meets the von Mises yield criterion. 2. The elastic deformation of the tubular material is ignored, and the external work is completely converted to the plastic strain energy and work done against friction. 3. The length of the simplified profile (straight line FV) of the wrinkle remains unchanged in the process of deformation. 4. The peak F overlaps F0 in the process of deformation, as shown in Figure 3(c). In Figure 3(c), the bending strain energy Wb when the yield line FV moves to F0V0 can be expressed as13 Wb \u00bc 2 Mp\u00f0hF hV l sin \u00de \u00f012\u00de where the plastic modulus Mp per unit length and the length l of the yield line are, respectively Mp \u00bc et 2 2 ffiffiffi 3 p \u00f013\u00de l \u00bc ffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi hF hV\u00f0 \u00de 2 \u00fel2y q \u00f014\u00de Under an external force, the plastic strain energy Wc in the hoop direction can be expressed as Wc \u00bc Z 2 et0 \"ej jVd \u00f015\u00de where the equivalent strain \"e and the volume V of the merging region between the peak and valley of a wrinkle are, respectively \"e \u00bc d\" d \u00bc d d \u00f0hF hV lsin \u00de hV \u00f016\u00de V \u00bc \u00f0h2F \u00fe h2V \u00fe hFhV\u00dely 3 \u00f017\u00de Because the yield line FV is in the free bulging zone without friction with the dies, the work caused by friction when the yield line FV moves to F0V0 is Wf& 0",
+            "comDownloaded from Combining equations (12), (15), (18), and (21), one can write P \u00bc lcos ly 2 Fz \u00fe et 2ffiffiffi 3 p \u00f0hF hV l sin \u00de \u00fe 2 et \u00f0hF hV l sin \u00de \u00f0h2F \u00fe h2V \u00fe hFhV\u00dely ( ) 3hV 8>>>>>>< >>>>>>: 9>>>>>>= >>>>>>; d0 lg S 2 \u00fe h2V lcos ly 2 3 ly 2h2F\u00fe3hFl sin \u00fel2 sin2 2 hFhV h2V ! 8>>>>< >>>>: 9>>>>= >>>>; \u00f022\u00de Let m \u00bc lcos ly 2 n \u00bc et 2ffiffiffi 3 p \u00f0hF hV l sin \u00de \u00fe 2 et \u00f0hF hV l sin \u00de\u00f0h2F \u00fe h2V \u00fe hFhV\u00dely 3hV q \u00bc d0\u00f0lg S 2\u00de \u00fe h 2 V lcos ly 2 3 ly 2h2F\u00fe3hFl sin \u00fel2 sin2 2 hFhV h2V ! 8>>< >>: 9>>= >>; Then, equations (22) can be rewritten as P \u00bc mFz \u00fe n q \u00f023\u00de As shown in Figure 3(c), when \u00bc 0, the wrinkle is just flattened (or the contour between the peak and valley of the wrinkle varies from oblique to horizontal, and likely convex later), so the critical hydraulic pressure Pc can be obtained as Pc \u00bc mcFzc \u00fe nc qc \u00f0 \u00bc 0\u00de \u00f024\u00de If the hydraulic pressure P increases high enough in the subsequent process to flatten the wrinkle, the wrinkle becomes a useful one. Consequently, the prediction formula for useful wrinkles can be expressed as P4Pc \u00f025\u00de Experiments in THF with axial feeding To validate the two analytic formulas for predicting the wrinkle type, which were derived in previous section, experiments in the THF under pulsating hydraulic pressure with axial feeding were carried out on a self-developed THF experimental system",
+            "comDownloaded from d0 (mm) initial outer diameter of a tubular blank (Figure 1) d\"e ( ) equivalent strain increment (equation (5)) d\"t( ), d\"y( ), d\" ( ) strain increments at valley V of the wrinkle in the thickness, axial and hoop directions, respectively (equation (5)) f (c/mm) fluctuation intervals of the hydraulic pressure per axial feeding (Figure 8) Fd (KN) axial force on the cross section at the wrinkle valleys (Figure 2) Ff (KN) friction force between tube and locating ring (Figure 2) FP, FP\u2019 (KN) generated forces on the pusher and the inner wall of deformed tube by the hydraulic pressure, respectively (Figure 2) Fy (KN) push force from the pushers (Figure 1) Fyc (KN) critical value of push force Fz from the pushers (Figure 1, equation (24)) hc(mm) critical geometry parameter for geometry-based prediction method (GPM) in this paper (equation (11)) hV, hF (mm) bulging height at the wrinkle valley and peak, respectively (Figure 2) l (mm) length of the straight yield line FV of the wrinkle (Figure 3) l0 (mm) Initial length of a tubular blank (Figure 1) lb (mm) bulging zone length of a tubular blank (Figure 1) lg (mm) locating zone length of a tubular blank (Figure 1) ly (mm) Projected length of the straight yield line FV of the wrinkle on y-axis (Figure 3) Mp (MPa/mm) plastic modulus per unit length (equation (12)) P (MPa) hydraulic pressure in tube (Figure 1) Pc (MPa) critical hydraulic pressure to get a wrinkle a useful one (equation (24)) S (mm) axial feeding distance on the end of the tube (Figure 1) t (mm) wall thickness at the wrinkle valleys (equation (2)) t0 (mm) initial wall thickness of a tubular blank (Figure 1) T total bulging time (Table 3) Ti (s) sampling time in total bulging time (Table 3) V (mm3) volume of the merging region between the peak and valley of a wrinkle (equation (17)) Wb, Wc, We (KJ) the bending strain energy of the yield line AB, plastic strain energy in hoop deformation, external work (equations (12), (15), and (18)) Wp, WFy (KJ) external work by hydraulic pressure P and external work WFy by axial force Fy, respectively, (equations (19) and (20)) Dh (mm) displacement of the valley of the wrinkle along z-axis in each time (Figure 3) P (MPa) fluctuation amplitude of the hydraulic pressure per axial feeding (Figure 8) V (mm3) volume difference when the yield line moves from FV to F\u2019V\u2019 (equation (19)) \"e( ) equivalent strain (equation (16)) ( ) Friction coefficient between the tube and the locating rings, 0.125 (equation (1)) e (MPa) Equivalent stress (equation (5)) z, (MPa) Axial and hoop stresses at the wrinkle valleys, respectively. (equations (2) and (4)) ( ) Angle between yield line FV of the wrinkle and the y-axis (Figure 3) at UNIV OF CINCINNATI on May 23, 2016pic.sagepub.comDownloaded from"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0000739_978-1-4471-4976-7_91-1-Figure1-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0000739_978-1-4471-4976-7_91-1-Figure1-1.png",
+        "caption": "Fig. 1 The illustration of the symbols",
+        "texts": [
+            " In the following formulation, the recursive method is applied to a manipulator with n rigid links. Coordinate systems are established by following Denavit-Hartenberg\u2019s convention (D-H convention) (Denavit and Hartenberg 1955). In the inverse dynamics problems, the positions, velocities, and accelerations (yi, _yi, \u20acyi for a revolute joint and di, _di, \u20acdi for a prismatic joint) of the joints are known. The outward recursion propa gates velocities, and accelerations of each link from the base to the end link. Referring to Fig. 1, the angular velocity of a revolute joint, propagated from link i 1 to link i, is given by vi \u00bc vi 1 \u00fe y : izi (1) or in local components Riv 0 i \u00bc Ri 1v 0 i \u00fe y : Riz 0 i (2) whereRi stands for the orientation matrix of link i andvi ' is the angular velocity of link iwith respect to the ith frame (local coordinate system). Premultiplying both sides of Eq. 2 with Ri T and simplifying yield v0 i \u00bc Ri i 1v 0 i \u00fe y : iz 0 i (3) whereRi 1 i is the rotation matrix relating link i 1 and link i. z ' i is the unit vector [0,0,1] T parallel to the zi axis",
+            " With the velocities and accelerations found, the inertial force and torque applied to link i at its center of mass can be calculated by applying the Newton-Euler equations: f 0Ci \u00bc mi _v 0 Ci (13) n0 Ci \u00bc Ii _v 0 i \u00fev0 i Iiv 0 i (14) where mi is the mass of the link and Ii denotes the inertia tensor of link i about its center of mass, calculated in a frame identical to the link\u2019s coordinate system except the origin. In the equations above, all items are expressed in the local coordinate systems. The forces and torques acting on the links are calculated through the outward recursion. The joint torques that generate the forces and torques will be calculated through the inward recursion. Referring to Fig. 1, the inward recursion propagates the joint forces and torques from the end link to the base as f 0i \u00bc f 0Ci \u00fe Ri i\u00fe1f 0 i\u00fe1 (15) n0 i \u00bc n0 Ci \u00fe Ri i\u00fe1n 0 i\u00fe1 \u00fe p0 Ci f 0Ci \u00fe pii\u00fe1 Ri i\u00fe1f 0 i\u00fe1 (16) where fi ' and ni ' are the force and torque exerted on link i with respect to the ith frame. The actuating torque on a revolute joint can be calculated by ti \u00bc n 0T i z 0 i (17) For prismatic joints, the driving force is f i \u00bc f 0T i z 0 i (18) In the forward dynamics problem, the joint accelerations are to be calculated with known joint torques"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0001217_978-3-319-08072-7_38-Figure2-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0001217_978-3-319-08072-7_38-Figure2-1.png",
+        "caption": "Fig. 2 Schematic drawing of an example SLIP model, including trunk segment. VPP is the \u2018virtual pivot point\u2019, a virtual point in which the GRF vectors intersect [3].",
+        "texts": [
+            " For these experiments advanced testing environments are required, for example tilting, rotating, or fast accelerating walking surfaces, or additional devices to perturb humans walking on a treadmill or over ground. The main focus is on ankle- and foot placement strategies, but also the stiffness adaptation of the leg is being considered. C) As human modelling environment we have selected to build on the SLIP (spring-loaded inverted pendulum) modelling approach, starting with a highly simplified 2D model and from there adding the elements needed to adequately describe the experimental results (a more complex SLIP model example shown in figure 2.) D) To develop the \u2018sense of balance\u2019 we are studying the approaches used in bipedal robot control and in biomechanics, and translating them for use in BALANCE. We classified methods in \u201cInstantaneous\u201d and \u201cRetrospective\u201d, where the first can be used in online control to detect the onset of a fall, and the second to evaluate performance over a number of strides. Most of the existing methods only are valid during a specified continuous task, or at least require information about the actual environment and task carried out"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0002869_s1474-6670(17)40029-2-Figure1-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0002869_s1474-6670(17)40029-2-Figure1-1.png",
+        "caption": "Fig. 1. General N -trailer system. The leading vehicle may be cart- or car-like.",
+        "texts": [
+            " T- I 1JHTr zr )), each of its components is a sum of polynomials in ;;2, . . . , Zn of degree larger or equal to one. Therefore, we may decompose this vector field as a sum of b~ -homogeneous v.f. with degrees not smaller than m - M = -n + 2 (since n ~ 3) . From Definition 2, C2 must be a sum of vector fields homogeneous of degree larger than -q. This condition is satisfied if -1 > -q and -n+2 > -q, that is if q > max{l, n - 2} = n - 2, as was to be shown. \u2022 Let us consider the general N-trailer system with off-axle hitching, as shown on Fig. 1. In order to derive a kinematic model, two assumptions are made: (i) the vehicles' wheels roll on a plane without slipping; (ii) the leading vehicle (the tractor) is a car-like vehicle equipped with a front steering wheel. When the leading vehicle is cart-like, the extension of the present results is straightforward. The notation for various physical parameters and angles is detailed on Figure 1. The vehicles are numbered starting with the one farthest from the tractor. The relative orientations of the ve hicles with respect to each other are given by { 0i h ~ i ~ N +1, while \u00b0 N + 2 is the orientation of the car's front wheel. Besides these angles, the posi tion and orientation of one of the vehicles, say ve hicle M, must be determined to completely char acterize the system's configuration in the plane. To this purpose, one may consider a given planar curve C, whose known curvature K is a function of the distance 8 measured along C from some point Co E C, and consider a set of Frenet coordinates (8,y,(3) as shown on Figure 2"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0000670_s11071-016-3035-3-Figure1-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0000670_s11071-016-3035-3-Figure1-1.png",
+        "caption": "Fig. 1 A simplified model of the vehicle and its front axle",
+        "texts": [
+            "eywords Brachistochronic motion \u00b7 Nonholonomic system \u00b7 Wheeled vehicle \u00b7 Optimal control R. Radulovic\u0301 \u00b7 A. Obradovic\u0301 \u00b7 Z. Mitrovic\u0301 Faculty of Mechanical Engineering, University of Belgrade, Kraljice Marije 16, Belgrade 35 11120, Serbia S. \u0160alinic\u0301 (B) Faculty of Mechanical and Civil Engineering in Kraljevo, University of Kragujevac, Dositejeva 19, Kraljevo 36000, Serbia e-mail: salinic.s@ptt.rs The subject of this research paper is a wheeled vehicle shown in Fig. 1. The motion of the vehicle is observed with respect to the fixed reference frame O\u03be\u03b7\u03b6 whose coordinate plane O\u03be\u03b7 coincides with the horizontal plane of the vehicle motion. The moving coordinate frame Axyz is rigidly attached to the vehicle body, so that the coordinate plane Axy coincides with the plane O\u03be\u03b7 where point A represents the mass center of the front vehicle axle. The unit vectors of the axes x , y, and z are \u2212\u2192 i , \u2212\u2192 j , and \u2212\u2192 k , respectively. The axis Ax passes through the mass center C of the vehicle body, and it is normal to the rear vehicle axle",
+            " (5) on the axes of coordinate frame Axyz yields: M [ V\u0307 \u2212 ( l2 + M2 M l1 ) \u03c92 ] = F1 \u2212 RA sin \u03b8, (6) M [ \u03c9V + ( l2 + M2 M l1 ) \u03c9\u0307 ] = RA cos \u03b8 + RB, (7) 0 = N1 + N2 \u2212 Mg, (8) J \u2217\u03c9\u0307 + J2\u03b8\u0308 + M ( l2 + M2 M l1 ) \u03c9V = RAl cos \u03b8, (9) 0 = M1gl2 + M2gl \u2212 N2l, (10) where \u03c9 = \u03d5\u0307 is the vehicle body angular velocity, M = M1 + M2, J = J1 + J2, J \u2217 = M1l22 + M2l2 + J is the moment of inertia of the vehicle about the axis B\u03b6 , N1 and N2, respectively, are normal reactions of the horizontal plane on the rear and front axles, g is the gravity acceleration, and cos \u03b8 = V/ \u221a V 2 + l2\u03c92 (see Fig. 1). Further, the differential equation of the front axle rotation about the axis Az reads: J2 ( \u03c9\u0307 + \u03b8\u0308 ) = L1. (11) Now, based on Eqs. (6)\u2013(11) it is possible to determine the reactions of nonholonomic constraints, as well as the driving force and the turning torque required to realize motion as follows: RA (t) = 1 l \u221a V 2 + l2\u03c92 V [ J \u2217\u03c9\u0307 + (Ml2 + M2l1) \u03c9V + J2\u03b8\u0308 ] , (12) RB (t) = 1 l [ M1l1\u03c9V + (M1l1l2 \u2212 J ) \u03c9\u0307 \u2212 J2\u03b8\u0308 ] , (13) F1 (t) = MV\u0307 + \u03c9 V ( J \u2217\u03c9\u0307 + J2\u03b8\u0308 ) , (14) L1 (t) = J2 ( \u03c9\u0307 + \u03b8\u0308 ) , (15) as well as the reactions N1 and N2: N1 = M1l1g l , (16) N2 = M1gl2 l + M2g"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0003348_0898150021000039301-Figure1-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0003348_0898150021000039301-Figure1-1.png",
+        "caption": "FIGURE 1 Schematic illustration of experimental setup.",
+        "texts": [
+            " In the present study, Stellite 6, one of the popular grades of Stellite, was used for the development of graded overlay on AISI 304 stainless steel substrate. Stellite 6 is a derivative of Co\u2013Cr\u2013W alloys, originally developed by Haynes [10]. An indigenously developed 2.5 kW continuous wave CO2 laser [11] was used for the development of graded overlay of Stellite 6 on AISI 304 stainless steel substrate. The experimental set-up consisted of a 2.5 kW laser system, integrated with a beam delivery system and a computer-controlled workstation, as shown in Fig. 1. A beam delivery system, consisting of a plane mirror and a 500mm focal length concave mirror, was employed to focus the laser beam. The cladding process involved scanning the surface of the substrate with a 3mm wide defocused laser beam and simultaneous injection of powder into the molten pool on the surface of the substrate. Overlapping tracks of about 3mm width and with 50% overlap were made to hardface larger surface area. Powder of Stellite 6 and AISI 316 stainless steel (size\u00bc 45\u201375 mm) were employed for the formation of graded overlay"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0001235_b978-0-08-099425-3.00005-4-Figure5.90-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0001235_b978-0-08-099425-3.00005-4-Figure5.90-1.png",
+        "caption": "FIGURE 5.90",
+        "texts": [
+            " One DOF is associated with the spin motion of the tyre, that is dependent on the longitudinal forces generated and the slip ratio. The other DOF is the height of the wheel centre above the road, that is controlled by the applied force representing the wheel load. Computer graphics for the tyre rig model. The tyre test rig model has been used to read the tyre model data files used in a study (Blundell, 2000a) to plot tyre force and moment graphs. The graphics of the tyre rig model are shown in Figure 5.90. The results obtained from a series of tyre tests (Blundell, 2000a) have been used to set up the data needed for the various modelling approaches described here. In summary the following procedure was followed: 1. For the Interpolation method the measured numerical values were reformatted directly into the SPLINE statements within an MSC.ADAMS data file as shown in Table 5.8. For each spline shown in Table 5.8 the X values correspond to either the slip or camber angle and are measured in degrees"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0000050_iemdc.2015.7409124-Figure5-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0000050_iemdc.2015.7409124-Figure5-1.png",
+        "caption": "Fig. 5. Investigated PM BLDC moto PCB; (b) ring-shaped PM.",
+        "texts": [
+            " Measurement sensors c temperatures (such as on cooling e temperature is too high, then the control operating voltage for the fan and hence th air flow. Bearings are a critical component because bearings make the fan rotate sm reduce friction, allow the fan to operate a are partly responsible for the life expect fan in a computer and the noise level of fa bearings can be used in a cooling fan: (a (b) ball bearings, and(c) fluid dynamic bea The specifications of the investigate brushless motor are given in Table 1. motor is shown in Fig. 5. The stator hous have been removed. TABLE I Specifications of investigated computer Rated input power No-load speed Rated voltage Rated current Number of poles Stator core outer diameter Axial length of stator stack PM Outer diameter (OD) PM inner diameter (ID) Axial length of PM The ring-shaped PM is made of an ferrite with remanent magnetic flux dens coercivity Hc = 260 kA/m. The built-in converter receives position information placed in the q-axis of the stator (Fig. 3). trolling computer nected to a 4-pin er) pin is used to control pin is an h requires a pullke linear voltage oportional to the upply voltage; the sed on the control erating at 25 kHz, peed"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0000672_mecatronics.2016.7547109-Figure6-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0000672_mecatronics.2016.7547109-Figure6-1.png",
+        "caption": "Fig. 6. Developed filling module within a practical course",
+        "texts": [
+            " Sprint: Realization of storing, separating and filling \u2022 4. Sprint: Testing of storing, separating and filling \u2022 5. Sprint: Implementation of virtual commissioning \u2022 6. Sprint: Design, realization and testing of transporting and positioning \u2022 7. Sprint: Physical installation and integration \u2022 8. Sprint: Global review and performance check From a component-based point of view, the developed module by the representative team is composed by a metallic hopper, containing the bulky filling material that can be refilled manually throughout the process (see Fig. 6). Underneath the hopper, a separator detaches a certain amount of filling material as a result of the ground area and the velocity of the slide feed moving the separator. 023 The drive of the filling machine is controlled by a PLC, being connected with sensors and actors via input-output clips. Whenever a packet is recognized at the entry and the module is free of other packets, the conveyor is activated in order to transport a new packet to the filling station. Once the packet is located under the hopper, the conveyor will be stopped and the filling process is initiated and ended automatically"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0002762_s0094-5765(98)00040-x-Figure6-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0002762_s0094-5765(98)00040-x-Figure6-1.png",
+        "caption": "Fig. 6. The manipulator model and the commanded trajectory.",
+        "texts": [
+            " To verify e ectiveness of this control strategy, several di erent cases were simulated numerically. Here three typical sets of results are presented. Deformation of the links is expressed by the \u00aerst mode in the inverse kinematics, inverse dynamics, and direct dynamics. 6.2.1. Case 1. Consider a simple example of the manipulator with one \u00afexible slewing link and one \u00afexible deployable link. The manipulator is \u00aexed to a large platform, for example the Canadarm attached to the Shuttle, and the translation of the manipulator is absent. Figure 6 shows the manipulator model and desired trajectory of the end-e ector (dotted line). The mass, length, and sti ness of the two links are m = 1.0 kg, l = 10 m, and EI = 3193 Nm2, respectively. The maneuver requires the end-e ector to move 10 m in the y direction in 10 s, and the required velocity of the end-e ector in the y direction is taken as a fourth order polynomial function in time. The solid lines in Figs 7\u00b110 show the controlled response of the \u00afexible manipulator, while the dashed\u00b1dotted lines present time histories of the desired velocity, acceleration, the desired deployed length, and slew angle"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0002199_hlca.19920750107-Figure2-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0002199_hlca.19920750107-Figure2-1.png",
+        "caption": "Fig. 2. Cyclic uoltammetry a J O / r l i e ~n6iinif rnodek 6c18 und b) of 7b in D.WF/ (Bu,N) C/O, (0.1 M). I I = 0.05 V/s at a glassy carbon electrode; a ) 1 : 5 mixture of 6c (0.25 mM) and 8 (1.25 mM); b ) 7b (ca. 0.25 mM).",
+        "texts": [
+            "+ subunits in 7b which might be expected from the behavior of oligomeric viologens [2]. The viologen subunits in 7b behave as multiple non-interucting redox sites (an important condition for the 'battery effect' to be discussed in the following chapter) as the observed E,,, = (Epc + EPJ2 involving the 5 viologen subunits in 7b is essentially identical with that of a single redox site in 8 [3] (Table 1) . To simulate the redox behavior of 7b (Co\"-V;+) as close as possible, its electrochemistry was compared with that of 1 : 5 mixtures of the subunits models 6c (Co\") and 8 (SV++) (Fig. 2). The Co\"'/Co\" redox couple shows the intensity expected for a slow redox system at low concentration in the 1:5 mixture. Under the same conditions, no Co\"'/Co\" 6c lc 8 7b +0.38\") +0.39\") +0.31\") -0.49b) -0.52b) -0.43b) -0.Xlb) -0.40b) ca. -0.5') -0.829 \") From spectroelectrochemistry. b, FromCV. ') Estimated from spectroelectrochemistry reduction and oxidation waves are observed for 7b. Two reasons may by responsible for this behavior. Either the thermodynamic Co\"'/Co\" reduction potential is shifted in negative d"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0001081_icarcv.2016.7838776-Figure14-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0001081_icarcv.2016.7838776-Figure14-1.png",
+        "caption": "Figure 14. Vibration model for RS during movement",
+        "texts": [
+            " 12 and 13 show the relationship between the stick and keyboard and the angle of the stick-holding robot hand \u03d5, respectively. First, the robot hand engaged in z-axis rotation by \u03d5 [\u00b0]; then, the key-to-key movement Y [mm] results were obtained. As \u03d5 increases, the RS vibration in the y-axis movement (zaxis) direction is suppressed (increased), because the external force causing the movement is divided. Thus, Y sin\u03d5, which causes the z-axis direction vibration, increases in the 0 \u03d5 90 [\u00b0] range. We considered this vibration model and attempted to control it. Fig. 14 shows the RS coordinate system. The wooden-ball vibration was modeled as a transient vibration with displacement y0. The quarter sine << \u2264\u2264 = )4,0(.0 )40(.sin 0 tTt TttY y m m\u03c9 (4) was used to express the motion equations (a) 40 mTt \u2264\u2264 , .0)cos(sin 1 2 1 2 1 2 1 2 =+\u22c5\u22c5\u2212++ \u03d5\u03b8\u03c9\u03c9\u03b8\u03b8\u03b8 mmmm tAmlklclml (5) (b) tTt h << 4,0 , .02 2 2 2 2 2 =++ mmm klclml \u03b8\u03b8\u03b8 (6) where is the input angular frequency [rad/s], Tm is the input power cycle [s] and m1, m2 are the RS swing angles [\u00b0]. To travel to B6 from C7, the robot hand moves in the y-axis direction only"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0002002_s0094-5765(99)00125-3-Figure1-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0002002_s0094-5765(99)00125-3-Figure1-1.png",
+        "caption": "Fig. 1. Schematic of a manipulating truss.",
+        "texts": [
+            " All rights reserved Printed in Great Britain 0094-5765/00 $ - see front matter 717 PII: S0094-5765(99)00125-3 {Paper IAF-96.A7.04 presented at the 47th International Astronautical Congress, 7\u00b111 October 1996, Beijing, China {Corresponding author. Tel.: +1-514-398-6288; fax: +1- 514-398-7365; e-mail: misra@mecheng.lan.mcgill.ca vibration control schemes are required. The objective of this paper is to present a dynamical model and use it to devise vibration control schemes. The system considered is a manipulating truss structure (Fig. 1) such as a truss crane. It is modelled as a collection of substructures or links consisting of: . statically determinate or indeterminate truss booms; . prismatic actuators that act as length varying elements to produce the required geometry change of the overall structure; . possibly a remote manipulator as the end e ector of the truss crane. Each truss boom is treated as a separate body or link, and its \u00afexibility is modelled using the \u00aenite element method (FEM), such that only axial deformation of its members is considered (truss members can take only axial load)"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0002879_6.1998-3285-Figure11-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0002879_6.1998-3285-Figure11-1.png",
+        "caption": "Figure 11 Enhanced Seal/Rotor pressure distribution viewed in the axial direction for 0.016\" axial clearance",
+        "texts": [
+            " This wall, or flow deflector, effectively prevents radial flow leaving the air dam region from entering the air bearing region. Instead, flow from the air dam is forced to exit the seal via the vent slots, and is prevented from influencing the flow from the orifice holes in the air bearing. Fig. 10 shows the flow field for the enhanced seal/rotor configuration, again for a clearance of 0.016\" and a pressure differential of 7.1 psid. Note that the flows from the air dam and air bearing regions are completely isolated, with all of the air dam flow exhausting the seal before reaching the air bearing. More importantly, Fig. 11 illustrates the pressure distribution for the case where the flow deflector is present, and reveals reduced pressures acting on the air bearing face in comparison to the original configuration. Integrating the pressure forces for the two configurations shows that the net pressure acting on the air bearing face is increased by >10% when the air dam flow is allowed to influence the air bearing region. Approx. deflector location Figure JO Enhanced Seal/Rotor flow field viewed in the axial direction for 0"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0000656_ccdc.2016.7532158-Figure1-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0000656_ccdc.2016.7532158-Figure1-1.png",
+        "caption": "Figure 1: Coordinate systems for ship sailing in restricted waters",
+        "texts": [
+            " In order to design a controller for a ship sailing in restricted waters, it is necessary to establish the ship dynamic model firstly. This section gives detail knowledge about ship dynamic model considering the nonlinear force and moment induced by the river bank, according to the works by Sano et al. [6] and Zhang et al. [7]. It is assumed that, 1) ship surge speed is constant; 2) compared with surge speed, sway speed is small. As a matter of experience, such assumptions are proper, and will not affect results accuracy. Two right-handed coordinate systems as shown in Figure 1 are adopted, where the \ud835\udc42\ud835\udc65\ud835\udc66\ud835\udc67 is the ship-fixed coordinate system, with origin \ud835\udc42 at geometrical center of the ship, \ud835\udc42\ud835\udc65 points to the ship bow and \ud835\udc42\ud835\udc66 points to the starboard, \ud835\udc42\ud835\udc65\ud835\udc66 plane is parallel to the undisturbed free surface. \ud835\udc42\ud835\udc67 is vertical to the undisturbed free surface and points downwards. The \ud835\udc42\ud835\udc60\ud835\udc65\ud835\udc60\ud835\udc66\ud835\udc60\ud835\udc67\ud835\udc60 is the earth-fixed reference frame, with the origin on the channel centerline, and \ud835\udc42\ud835\udc60\ud835\udc65\ud835\udc60 is coincident to the river centerline,\ud835\udc42\ud835\udc60\ud835\udc66\ud835\udc60 is pointing to the right side, \ud835\udc42\ud835\udc60\ud835\udc65\ud835\udc60\ud835\udc66\ud835\udc60 plane is parallel to the undisturbed free surface"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0000962_tmag.2016.2636208-Figure8-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0000962_tmag.2016.2636208-Figure8-1.png",
+        "caption": "Fig. 8. Induction motor with solid rotor\u2014the view.",
+        "texts": [
+            " The coenergy method is applied only for checking the Maxwell force value. The cases c and d show that if condition (3) is not satisfied first, theorem cannot be applied, and in Fig. 6(c) and (d), both the forces are different. The material (anisotropy) force FeM = 0. First theorem is biconditional (equivalence). Examples for the second theorem are presented in Fig. 7 for the same data as for linear motor in case b), and are given in Table VI. For presenting theorems for electromagnetic torque, an induction motor with solid rotor is considered (Fig. 8). There is assumed: \u03b3 = 35 \u00b7106 S/m (rotor conductivity), a = 0.02 m (conductive rotor layer width), R = 0.05 m (rotor outer radius), l = 0.25 m (rotor length), g = 0.0005 m (the gap width), \u03b81 = 1504 A (magnetomotive force first harmonic magnitude), f1 = 50 Hz, p = 2 (pair pole number), \u03bdrr = \u03bd0/3 (radial reluctivity), \u03bd\u03b1\u03b1 = \u03bd0/2 (tangential reluctivity), and reluctivities \u03bdr\u03b1 , \u03bd\u03b1r (see Tables VII and VIII). If condition (3) is satisfied, the first theorem is true [see Table VII for cases a) and b)\u2014presented in Fig"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0003879_6.2001-4392-Figure5-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0003879_6.2001-4392-Figure5-1.png",
+        "caption": "Fig. 5 Rotating Model",
+        "texts": [
+            " - Tik jf\u2014\u2014\u2014\u2014^-j-2\u2014\u2014J-^ (3) (4) American Institute of Aeronautics and Astronautics D ow nl oa de d by K U N G L IG A T E K N IS K A H O G SK O L E N K T H o n Se pt em be r 11 , 2 01 5 | h ttp :// ar c. ai aa .o rg | D O I: 1 0. 25 14 /6 .2 00 1- 43 92 Armjk Fig. 4 Analytical Model Rotating Motion We consider the rotating motion; n satellites rotate on the c.m. of the system with the objective formation in same plane. We establish the coordinated control method using tether, arm, thruster and CMG. Rotating Model Figure 5 shows the rotating model in case of n = 3. The definition of rotating coordinate system {r} is the followings. The origin of {r} is the c.m. of the system. (5) r^ plane shows the rotating plane, and rz axis shows the rotating axis. rx axis shows the direction of the satellite 1 in the rotating plane from the c.m. of the system. The relationship between {r} and {i} is defined by the following Euler angles. where C is the rotating matrix; {m} is the medium coordinate system; and superscript d shows the objective (set) value"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0000582_978-3-319-27247-4_36-Figure3-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0000582_978-3-319-27247-4_36-Figure3-1.png",
+        "caption": "Fig. 3 Matlab/Simulink",
+        "texts": [
+            " Then, (10) and (11) can be reduced to, \u00f0a1 \u00fe a2 \u00fe 2a3 cos h2\u00de\u20ach1 \u00fe\u00f0a2 \u00fe a3 cos h2\u00de\u20ach2 a3 sin\u00f0h2\u00de _h22 2a3 sin\u00f0h2\u00de _h1 _h2 \u00fe a4g cos h1 \u00fe a5g cos\u00f0h1 \u00fe h2\u00de \u00bc s1 \u00f012\u00de a2\u20ach2 \u00fe\u00f0a2 \u00fe a3 cos h2\u00de\u20ach1 a3 sin\u00f0h2\u00de _h21 \u00fe a5g cos\u00f0h1 \u00fe h2\u00de \u00bc 0 \u00f013\u00de In which a1 \u00bc m1l 2 c1 \u00fem2l 2 1 \u00fe I1; a2 \u00bc m2l 2 c2 \u00fe I2; a3 \u00bc m2l1lc2 a4 \u00bc m2lc1 \u00fem2l1; a5 \u00bc m2lc2 In matrices form, the equations of motion of the two link Pendubot can be described as follows: D\u00f0h\u00de\u20ach\u00feC\u00f0h; _h\u00de _h + G \u00bc s \u00f014\u00de with s \u00bc s1 0 \" # D \u00bc a1 \u00fe a2 \u00fe 2a3 cos h2 a2 \u00fe a3 cos h2 a2 \u00fe a3 cos h2 a2 C \u00bc a3 sin\u00f0h2\u00de _h2 a3 sin\u00f0h2\u00de _h2 a3 sin\u00f0h2\u00de _h1 a3 sin\u00f0h2\u00de _h1 0 G \u00bc a4gcosh1 + a5gcos(h1 \u00fe h2\u00de a5gcos(h1 \u00fe h2\u00de \" # Figures 2 and 3 illustrate the 3D simulation of Pendubot with Solid-works in unstable balancing position (Fig. 2) and in stable balancing position (Fig. 3). The sliding mode controller is derived using the state dynamics described by equation, d11\u20ach1 \u00fe h1 \u00fe \u20acu1 \u00bc s1 \u00f015\u00de In which d11 \u00bc d11 d12 d22 ; h11 \u00bc h1 d12 d22 h2 ; u1 u1 d12 d22 u2 Consider a state space form of Pendubot system as ~x1 \u00bc h1 and _h1 \u00bc _x1 \u00bc ~x2: We have, _~x1 \u00bc ~x2 and _~x2 \u00bc f1(~x) + b1s1 with f1( x) = ( x11 _h1 \u00fe c12 _h2 \u00fe u1\u00de d11 ; b1 \u00bc 1 d11 ; d11 \u00bc d11 d12d21 d22 u1 \u00bc u1 d12 d22 u2 ; c11 c11 d12d21 d22 ; c12 \u00bc c12 Forwardly, consider a sliding surface as follows: S \u00bc k h1 hd1 \u00fe _h1 _hd1 \u00bc k~x1 \u00fe~x2 \u00f016\u00de The goal is to choose the scalar \u03bb value such that the system restricted on the surface (16) is of the stable characteristics"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0003768_mfi.1996.572183-Figure1-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0003768_mfi.1996.572183-Figure1-1.png",
+        "caption": "Fig. 1 Struc ture of WAF-3,0'-X'Y'Zt: detection coordinate.",
+        "texts": [
+            " From the foregoing considerations, the purpose of this study was to develop a foot mechanism which mounts a landing surface detection system, as an external sensor, which measures the gradient of the landing surface during dynamic walking as well as the relative position to the surface, to devise a walking control method that constitutes a closed loop control system which provides a lower-limb trajectory control against the landing surface on a real-time basis by utilizing the information of the landing surface, obtained by the above-mentioned foot mechanism, and to develop a life-size biped walking robot which can adapt to the real world in which the lower-limb trajectory deviation and the surface shape deviation can not be 0-7803-3700-X/96 $5.00 01996 IEEE 233 reduced to zero. 2 Required condition of the walking surface Required condition of the walking surface to be researched was set as follows. ( 1 ) Height and gradient of the path surtace are unknown factors. ( 2 ) All landing spikes (see Fig. 1) in the four comers of the biped walking robot's foot must be able to make contact with the walking surlke. ( 3 ) Extent of deform of the path surface aid that of movement that are expe- rienced when the machine model walks are limited to the ranges where there is no need to make a compensating movement that considers the dynarmG. It should be noted that the gradient of the lateral plane is provided to be horizontal (AccuracytO.5' ) and known in advance. This is because the present WL-I 2 Series has no active degree of freedom around the roll axis of the lower-limb re&g materials ofthe surt8ce, such that it does not deform at all when the machine mcdel walks and that sags approximately 2 mm when it is standing on a single le& satisfy the conditions",
+            " It is I& that floors which are not likely to deform, such as floors covered with a carpet or other flooring materials which may deform slightly (such as atatami mat and the like) satisfy such conditions. 3 Foot mechanism WAF-3 Like a m ' s palm finger pad and sole, although WAF-3 is flexible, it dces not transform any further when load is applied up to a point because of its nonlinear hardness property. WAF-3 is also equipped with sensors enough to detect the shapes and the gradient of the path surfaces during dynamic biped wallung. Fig. 1 presents a rough sketch of WAF-3 and tk detecziofi coo- 0'- XTZ' which was established in order to measure the landing path surface. Fig. 2 presents a photograph which shows the actual foot mechanism attached to a biped walking robot WL-12RVII. This mechanism consists of an upper foot plate which is directly fitted to the foot beneath the ankle actuator, a lower foot plate which makes direct contact with the path surface, several pieces of wire which connect the upper foot plate and the lower foot plate, and a sandwich structure which clamps an opencell-foam shock absorbing material between the upper foot plate and the lower foot plate. In reference to the movable range (MR) of the passive DOF of human lower limbs, a similar range was set for WAF-3. Specifically, on Cartesian coo~dinates as show in Fig. 1, the MRonthe X' axis was set at t5 .0 mm, the MR on the Z' axis was set at 5.5 mm, the MRaroundtJx X' axis was set at i 2.3\" ,and the MR m d theY axis was set at 31 I .7. The weight ofthis foot unit is 3.0 kg for one leg i.e., 1.8 kg heavierthan the conventional foot weight ofthe WL-I 2RV. The mechanisms which WAF-3 possesses are summarized below. (A) Landing path surfice detection mechanism (\u20ac3) Shock absorbing mechanism (C) Suppofl leg change stabilization mechanism 'Ihe following describes the individual mechanisms"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0002074_1.1359772-Figure3-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0002074_1.1359772-Figure3-1.png",
+        "caption": "Fig. 3 Arrangement of swirl brakes at the entrance of the seal",
+        "texts": [
+            " Leie and Thomas @13# and Benckert and Wachter @14# used swirl brakes mounted in the cavities or in front of the seal. An almost 384 \u00d5 Vol. 123, APRIL 2001 rom: http://gasturbinespower.asmedigitalcollection.asme.org/ on 01/27/20 total reduction of the lateral force is reported by the latter authors, when 8 swirl webs are placed at the entrance of the labyrinth. In the present paper, an arrangement with 4 ~4DB! and with 8 ~8DB! swirl brakes in front of the reference seal ~TOR29! is investigated ~see Fig. 3!. The best known use of swirl brakes was reported in the high-pressure-fuel turbopump of space shuttle ~@15#!. In this application a honeycomb-stator is inserted in the seal beside the swirl brake. In general, it is expected that honeycomb seals show favorable dynamic characteristics and, last but not least, smaller leakage rates. First investigations of Hawkins et al. @4# indicate that a 16-cavity TOR seal does not improve the stability significantly. Unfortunately the number of cavities, the honeycomb cell width and the pressure levels differ and therefore a direct comparison with the actual results is not possible",
+            " The cross-coupled stiffness shows the usual almost linear dependency on the preswirl. rom: http://gasturbinespower.asmedigitalcollection.asme.org/ on 01/27/20 The behavior of the direct damping is distinctive different. Even without entry swirl, the damping values differ from zero and do not depend on the preswirl. With an increased pressure difference, the leakage rate and the preswirl are higher, too. In a next step, swirl brakes are placed at the entrance of the seal. The swirl brakes are equally spaced on the circumference as illustrated in Fig. 3. Four swirl brakes suffice to reduce the crosscoupled stiffness by more than 50 percent ~see Fig. 7!. Eight brakes on the circumference allow an even further reduction. It is remarkable to see that the direct damping is increased simultaneously. In case of zero preswirl and small drag due to rotation, there is no difference between the measurement with and without brakes. The swirl brakes have no influence on the leakage. Now the swirl brakes are removed and the stator of the seal is replaced by a honeycomb stator"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0000942_j.renene.2016.11.054-Figure26-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0000942_j.renene.2016.11.054-Figure26-1.png",
+        "caption": "Fig. 26. (a) The mechanical fault simulator and (b) the experimental bearing.",
+        "texts": [
+            " Therefore, this example gives powerful evidence of the superiority of the proposed method for enhancing performance in the weak feature identification of generator bearing pitting damage. The experimental vibration data gathered from the experimental rig of crack damage simulation for generator bearing is also analyzed to further validate the effectiveness on structural damage identification of the proposed method. The rolling bearing running data were also acquired from a faulty generator of Spectra Quest Inc.\u2019s mechanical structure damage simulator shown in Fig. 26(a). The crack damages are generated by wire cutting machine on bearing outer-race, shown in Fig. 26(b). And the corresponding complete bearing damage simulation system also includes CoCo 80 data acquisition system and Spectra Quest Inc.\u2019s mechanical structure damage simulator. The test bearing supports the motor shaft. Condition vibration data was collected using accelerometers, which were mounted on to the housing, with the sampling frequency of 12800 Hz for output end bearing experiments. The experimental bearings with weak damages used in this test are the deep groove ball bearings with the type of NSK 6208 and the parameters of this type of bearing are displayed in Table 3"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0001051_phm.2016.7819861-Figure8-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0001051_phm.2016.7819861-Figure8-1.png",
+        "caption": "Figure 8. A rigid-flex coupling dynamic model for the planetary gearbox in ADAMS.",
+        "texts": [
+            " 1 1 1 1 1 0 , ( ) ( , ,1, ,0) ,e x x F dxK x x step x x d x C x x dt > =  \u2212 \u2212 \u2212 \u2264 (1) 216 4 9 3 REK RE= = 1 2 1 1 1 R R R = + (2) 2016 Prognostics and System Health Management Conference (PHM-Chengdu) where R1 and R2 represent the pitch circle radiuses of the two meshing gears; E-Young modulus; e-force exponent; dpenetration; C-damping; K-stiffness; x1-initial displacement; x-actual displacement [9]. The generated rigid-flex coupling dynamic model for the planetary gearbox in ADAM is shown in Fig. 8. The simulation parameters are shown in Table V. The HASTIFF solver and integrator SI2 are chosen to ensure the computational accuracy and avoid excessive impacts. A torque of 50N\u2022m is exerted on the carrier, at the same time a motion which is defined by a step function is exerted on the input shaft. The simulation time is 0.5s while the step number is 10000 to ensure iterative step is sufficiently small. Using these settings, the vibration response is simulated and analyzed. The meshing incentives between planet gears and the sun gear will be transmitted to the carrier which is an output component"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0000622_b978-0-12-800782-2.00003-8-Figure3.14-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0000622_b978-0-12-800782-2.00003-8-Figure3.14-1.png",
+        "caption": "Figure 3.14 Field for the determination of the saturated stator leakage flux of a 3.4 MW, two-pole, three-phase induction motor. One flux tube contains a flux per unit length of 0.15 Wb/m [19].",
+        "texts": [
+            " No closed form solution exists because of the nonlinearities (e.g., iron-core saturation) involved. In Fig. 3.13 the field for the first approximation, where saturation is neglected and a linear B\u2013H characteristic is assumed, permits us to calculate stator and rotor currents for which the starting field can be computed under saturated conditions assuming a nonlinear (B\u2013H) characteristic as depicted in Fig. 3.16. For the reluctivity distribution caused by the saturated short-circuit field the stator (Fig. 3.14) and rotor (Fig. 3.15) leakage reactances can be recomputed, leading to the second approximation as indicated in Fig. 3.16. In practice a few iterations are sufficient to achieve a satisfactory solution for the starting torque as a function of the applied voltage as illustrated in Fig. 3.17. It is well known that during starting saturation occurs only in the stator and rotor teeth and this is the reason why Figs. 3.13 and 3.16 are similar. 220 Power Quality in Power Systems and Electrical Machines F2 = 3500 N/m F1 = 3500 N/m F 2 = 3 11 0 N /m F 1 = 3 15 0 N /m F 2 = 1250 N/m F1 = 4340 N/m F 2 = 1770 N /m F 1 = 5160 N /m F 2 = 1400 N /m F 1 = 4110 N /m F 2 = 1 30 0 N/m F 1 = 5 45 0 N/m 1 \u2013B \u2013B 2 3 A A 4 5 \u2013C \u2013C \u2013A \u2013A C C B B 6 7 8 9 10 11 12 (a) 1 \u2013B \u2013B 2 3 A A 4 5 \u2013C \u2013C \u2013A \u2013A C C B B 6 7 8 9 10 N S 11 12 (b) f f 1 \u2013B \u2013B 2 3 A A 4 5 \u2013C \u2013C \u2013A \u2013A C C B B 6 7 8 9 10 N S 11 12 (c) Figure 3"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0003280_j.1460-2687.2001.00076.x-Figure3-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0003280_j.1460-2687.2001.00076.x-Figure3-1.png",
+        "caption": "Figure 3 Schematic drawing of propulsion model and parameters.",
+        "texts": [
+            " With the general design of the pitching machine established, the details of the propulsion system design remain to be worked out including; the air tank volume, the inner diameter of the air tank outlet pipe, the volume between the air tank and ball, and the barrel length and inner diameter. In addition to these design parameters, it was also important, for the ultimate design parameter choices, to understand the functional relationship between the initial air tank pressure and the ball exit velocity. Once this relationship is known, it will be possible to control the speed of the ball at release by controlling the pressure in the air tank. Therefore a dynamic model of the propulsion system (see Fig. 3) was developed. This model includes the volume in the air tank, air \u00afow through the opened release valve, the pressure build up in the volume behind the ball, and the ball acceleration down the barrel. This dynamic model is a set of differential equations that, given a set of initial conditions, were integrated in time using the software package MATLABTM (Hanselman & Little\u00aeeld 1995). Running multiple simulations with different design parameter settings led to an understanding of the system dynamics and determination of an acceptable set of design parameters. The parameters for the model are shown schematically in Fig. 3. Each may be varied independently to understand its effect on the resulting ball exit speed. The model begins with the air tank \u00aelled to a known initial absolute pressure, Pa. Initially the pressure Pb in chamber B and the portion of the barrel in front of the ball are considered to be at atmospheric pressure. Once the valve is opened (this is assumed to happen instantaneously) the air in the tank is free to \u00afow into chamber B. The difference between the rising pressure in chamber B and atmospheric pressure in front of the ball accelerates the ball"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0000050_iemdc.2015.7409124-Figure2-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0000050_iemdc.2015.7409124-Figure2-1.png",
+        "caption": "Fig. 2. Construction of a PM BLDC motor drive disassembled motor; (b) inner stator with four salie external rotor with 4-pole PM rotor.",
+        "texts": [
+            " The tips of the blades can also be made of a hard magnetic material and magnetized in radial direction [9]. II. CONSTRUCTION OF PM BLDC FAN MOTORS The cost effective two-phase brushless motors for computer fans have a salient pole inner stator and ringshaped outer PM rotor. The outer PM rotor is integrated with the fan blades facilitating air flow [1 \u2013 4]. The housing is mechanically connected with the inner stator of the motor with the aid of a spider structure (Fig. 1). The details of construction of a PC fan motor are shown in Fig 2. A Hall sensor detects the polarity of PMs and via solid state devices switches the DC voltage from one stator coil to another. The speed of the fan motor is controlled by adjusting either the DC voltage or pulse width in lowfrequency PWM [10]. In spite the PM BLDC motor has four dead spots per revolution, it has god self-starting capability. Since the rotor rests between the poles of PMs at zero-current state Analysis of Steady-State and Transient Performance of Two-Phase PM Motors for Computer Fans J",
+            " 3), and instantly rotates 45o when fi will not stop on one of its dead spot. Two phase stator winding consists of f around the stator pole cores. There are rst switched on, it 3 \u2013 front surface, 4 \u2013 r structure, 7 \u2013 blades, for computer fans: (a) nt poles; (c) PCB; (d) ition: (a) opposite coils in pairs. 1 \u2013 PM, 2 \u2013 econd phase, 5 \u2013 Hall our coils wrapped four coils in the inner stator (Fig. 3b), while different magnetic polarity. T pairs, either each one with its with its neighboring coils (Fig of the outer rotor, there are fo (Fig 2b, Fig. 3). Typically, a 12-V DC cool rotor-blade assembly containin stator. A Hall sensor detects th switches 12 V DC from one st 3). Varying the supplied DC v most fans. A 12-V DC fan 3.5\u20265.0 V DC voltage appl when increasing voltage is supp Typical electronic circuits PC fan motors are shown in F fans used in computers use s two to four pins. The first tw deliver power to the fan mo optional, depending on fan desi \u2022 ground; \u2022 power (+12 V); \u2022 sense: provides a tachomet actual speed of the fan as a p proportional to speed (with eac pulses sent through this pin; \u2022 control: provides a PWM si to adjust the rotation speed voltage delivered to the cooling two neighboring coils have he coils are connected in opposite coils (Fig"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0003645_i2002-10009-1-Figure1-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0003645_i2002-10009-1-Figure1-1.png",
+        "caption": "Fig. 1. TGBA phase, schematic (adapted from M. Kleman, O.D. Lavrentovich, Soft Matter Physics, an Introduction (Springer, 2002).",
+        "texts": [
+            " Renn and Lubensky [3] have suggested that the analog of the Abrikosov phase is a helical SmA characterized by a set of regularly spaced (repeat distance lb) twist grain boundaries (TGBs). These TGBs are made of equidistant screw dislocations that play the role of the lattice of vortex lines in the Abrikosov phase, and that separate SmA slabs rotated one with respect to the other by an angle \u03c9 \u2248 d/ld (more precisely [4] 2 sin \u03c9/2 = d/ld). Here d is the SmA repeat distance, and a e-mail: maurice.kleman@mines.org ld is the distance between two neighboring screw dislocations, see Figure 1. Hence the name of TGBA was proposed for such a phase. We have p = 2\u03c0/q = 2\u03c0lb/\u03c9, p being the pitch. The Ginzburg-Landau parameter \u03ba can be written [5] \u03ba2 = \u03bb2/\u03be\u22a5 = d \u221a 2/2\u03c0\u03be\u22a5, when expressed in these same parameters. We shall call in the sequel vortex lines those screw dislocations that build the TGBs. TGBA phases are documented in many materials [6]. By taking into account thermal fluctuations one expects, in type-II superconductors [7], to find a liquid phase [7, a)] and a glass phase [7, b)] of vortex lines; these phases have indeed been observed"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0000878_978-3-319-47614-8-Figure2.18-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0000878_978-3-319-47614-8-Figure2.18-1.png",
+        "caption": "Fig. 2.18 Mohr\u2019s circle principal and Max shear stresses",
+        "texts": [
+            "13) as rx0 rx \u00fe ry 2 \u00bc rx ry 2 cos 2h\u00fe sxy sin 2h sx0y0 \u00bc rx ry 2 sin 2h\u00fe sxy cos 2h \u00f02:18\u00de Squaring and adding and using cos2 2h\u00fe sin2 2h \u00bc 1, we get rx0 rx \u00fe ry 2 2 \u00fe s2x0y0 \u00bc rx ry 2 2 \u00fe s2xy \u00f02:19\u00de This is a circle plotted on a graph where the abscissa is the normal stress and the ordinate is the shear stress. This is easier to see if we interpret \u03c3x and \u03c3y as being the two principal stresses, and \u03c4xy as being the maximum shear stress. Then we can define the average stress, \u03c3avg, and radius R (which is just equal to the maximum shear stress), rAvg \u00bc rx \u00fe ry 2 R \u00bc ffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi rx ry 2 2 \u00fe s2xy r \u00f02:20\u00de The circle is centered at the average stress value, and has a radius R equal to the maximum shear stress, see Fig. 2.18. 52 2 Continuous Solid Note that the circle Eq. (2.19) is rx0 rAvg 2 \u00fe s2x0y0 \u00bc R2 \u00f02:19a\u00de The two principal stresses and the maximum shear stress are shown on Mohr\u2019s circle. Recall that the normal stresses equal the principal stresses when the stress element is aligned with the principal directions, and the shear stress equals the maximum shear stress when the stress element is rotated 45\u00b0 away from the principal directions. Also remember from Eq. (2.18), the angle on Mohr\u2019s circle is 2\u03b8, see Fig"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0002128_ma946435m-Figure1-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0002128_ma946435m-Figure1-1.png",
+        "caption": "Figure 1. Optical system and the reference coordinate frame: (C) upper cone; (PL) lower plate; (S) sample; (A) analyzer; (P) polarizer; (x1) shear rate gradient direction; (x2) vorticity direction; (x3) flow direction; (\u03b8) scattering angle; (\u00b5) azimuthal angle; (\u03c8) angle between polarizer direction and flow direction.",
+        "texts": [
+            " The sample thickness at the point where the light is incident was about 400 \u00b5m. The direction of the scattered radiation is defined by \u03b8 (or q) and \u00b5 which are the scattering angle (or scattering vector q with q ) (4\u03c0/\u03bb) sin(\u03b8/2)) and the azimuthal angle, respectively. The direction of the optical axis of the polarizer with respect to the flow direction is defined by the angle \u03c8. The optical axis of the analyzer is either parallel or perpendicular to the optical axis of the polarizer. The gradient, vorticity, and flow directions are named x1, x2, and x3 (see Figure 1), respectively. SALS patterns are observed in a plane (x2-x3 plane) perpendicular to the incident beam. They are recorded at a rate of 25 frames/s with a video recorder and digitized into a 512\u00d7 512 pixel matrix with 256 gray levels. These gray levels give semiquantitative information about the scattered intensity. All the experiments were performed at room temperature (21-22 \u00b0C) between crossed polarizers. A more complete description of the experiments and the results will be published later. The evolution of the intensity of the vertical streak (the intensity at the azimuthal angle \u00b5 ) 90\u00b0) was measured for different values of the angle \u03c8",
+            " This can also give rise to diffraction effects. So, we must assume that the nonhomogeneous distribution of scattered light in the vertical streak is due to a 3D diffraction phenomenon. To study this, we will consider the simplest case: N disclination loops of equal lengths and similar internal core structure will be considered as assembled into one set. The upper and lower horizontal parts of the loops will be regarded as rods which are located in different planes parallel to the screen in this set (Figure 1). The projection of this set onto the (x1, x2) plane formed by the incident beam and vorticity axis is shown in Figure 8. The distances l and h between the nearest points of the rods along x2 and x3, and the angle R are independent random values obeying Gaussian distributions: s ) -i sin \u03b8 2 + j cos \u03b8 2 sin \u00b5 + k cos \u03b8 2 cos \u00b5 (5.6) O ) j cos \u03c8 - k sin \u03c8 (5.7) (MO) ) 1 2 E0\u03b4{sin 2\u03c8 sin2 \u00e2 cos2 \u03c9 + cos 2\u03c8 sin 2\u00e2 cos \u03c9 - sin 2\u03c8 cos2 \u00e2} (5.8) f ) 1 2 E0\u03b4D 2LCj0(12qDs1)j0(12qDs2)A3(pL,\u00e2;\u03c8) (5.9) A3(pL\u03b4;\u03c8) ) sin 2\u03c8 sin2 \u00e2J2 + cos 2\u03c8 sin 2\u00e2J1 - sin 2\u03c8 cos2 \u00e2J0 (5"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0000099_indiancc.2016.7441157-Figure2-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0000099_indiancc.2016.7441157-Figure2-1.png",
+        "caption": "Figure 2: Angle Notations of Leader and Follower 1 Formation",
+        "texts": [
+            " vi is the forward velocity of the aircraft. The normal accelerations ani and tangent acceleration ati are defined as ani = Hn i (s)a n ci (5) ati = Ht i (s)a t ci (6) where anci and atci are the commanded accelerations as given by eqn 9 and eqn 10. The transfer function for the 20Hz low pass filters Hn i (s) and Ht i (s) are given in [6] Hn i (s) = 1 0.05s+1 (7) Ht i (s) = 1 0.05s+1 (8) The leaders path, follower 1\u2019s desired position with respect to the leader along with its various angles are shown in figure 2. The angle definitions are the same between leader and follower 2 also. Heading angle is defined as the angle between the aircraft forward x-axis and the horizon. \u03c81 and \u03c82 are the heading angles of the leader and the follower 1 respectively. l12 is the relative distance between the leader and follower 1. Similarly l13 is the relative distance between the leader and follower 2. \u03bb1 is the Line Of Sight (LOS) angle between the leader and the follower 1 and similarly \u03bb2 is the LOS angle between leader and follower 2"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0000185_20140824-6-za-1003.00239-Figure1-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0000185_20140824-6-za-1003.00239-Figure1-1.png",
+        "caption": "Fig. 1. View of the hydraulically-actuated fin loading system.",
+        "texts": [
+            " When these functions are designated, it is accounted to relinquish from the stability as minimum as possible in addition to diminishing the chattering effect. Having completed the computer simulations for all the control system algorithms established, the real-time experiments are conducted using the test setup developed. Thus, it is seen that the sliding mode control system with a varying sliding surface yields the most satisfactory results. The schematical representation of the considered FLS combined of four identical hydraulic actuation units is shown in Fig. 1 where p1 and p2 denote the pressure values of the hydraulic fluid in the inlet and outlet ports of the hydraulic cylinder, and pS and pR represent the pressure values of the 978-3-902823-62-5/2014 \u00a9 IFAC 10920 hydraulic fluid in the inlet of the fluid control valve and return line to the hydraulic fluid tank, respectively. The dynamics of the servovalve as the flow control valve can be expressed by the forthcoming first-order transfer function: 1  sT K I x v v   (1) where x, I, Kv, and Tv denote the displacement of the servovalve control spool, control current to the servovalve, gain, and time constant of the valve, respectively",
+            " Using equation (6), pL and its first time derivative are obtained in the following manner:   ppppL A/Fybymp    (7)   ppppL A/Fybymp    (8) As JTp represents the mass moment of inertia of the half portion of the torquemeter put between the fin connecting rod and transmission rod to measure the amount of the torque applied on the fin connecting rod around the rotation axis of the fin connecting rod, Fp and its first time derivative can be determined as follows:   Tpp JF  (9)   Tpp JF  (10) The relationship between the angular displacement of the fin and linear displacement of the piston can be established from Fig. 1 along with its successive time derivatives using the small angle assumption in the following manner:  pLy  (11)   pLy  (12)   pLy  (13)   pLy  (14) Substituting equations (9) through (14) into equations (7) and (8) give the forthcoming expressions:    pppppTpL A/LbLmJp    (15)    pppppTpL A/LbLmJp    (16) Inserting equations (12), (15), and (16) into equation (5) and arranging it, the differential equation describing the dynamic behaviour of the hydraulic FLS can be derived in the following manner (\u00d6zakal\u0131n, 2010 and \u00d6zakal\u0131n et al"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0001996_s0736-5845(98)00026-x-Figure34-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0001996_s0736-5845(98)00026-x-Figure34-1.png",
+        "caption": "Fig. 34. Parametrization of current path length s i (t#*t) within the curve segment.",
+        "texts": [
+            "6) The segment length S i,i`1 of the spline curve is the sum of the K pieces\u2019 curve length, calculated in Eq. (A.4), as Fig. 33 shows S i,i`1 \" K + j/1 *s j (A.7) Next step is planning the path parameter s (t) for each curve segment i. Knowing the velocity profile v(t) and having the curve segment length S i,i`1 we can use the characteristic formula to plan the parameter s i (t#*t) along a specified path with Eq. (A.8): s i (t#*t)\"s i (t)#v i (t)*t#1 2 a i (t)*t2. (A.8) where \u00b9 i (t(\u00b9 i`1 and s i \"s i~1 #S i~1,i as Fig. 34 shows. The necessity of transforming the path parameter s (t) into the parameter u (t) of the spline curve is given by the polynomial of the curve (see [11, 12]). First we write the path velocity v(t), v(t)\"A ds duB A du dtB\"s@(u) uR (A.9) and the path acceleration a(t), a (t)\" dv dt \"s@(u) ) u(#sA(u) ) uR 2 (A.10) as a derivation of the spline curve parameter u(t). Transforming the path parameter planning s (t#*t) in Eq. (A.8) to a curve parameter planning u (t#*t) we approximate u (t#*t) with Taylor\u2019s equation around the interpolation time *t: u i (t#*t)\"u i #*t ) uR i #A *t2 2 B ) u( i#2"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0002821_analsci.6.351-Figure1-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0002821_analsci.6.351-Figure1-1.png",
+        "caption": "Fig. 1 Electrochemical reactor (ECR). 1, glassy carbon rod; 2, working electrode (glassy carbon grain, ca. 300 mesh);",
+        "texts": [
+            " Experimental Chemicals and standard solutions VMA and HVA were purchased from Sigma Chemical Co. (St. Louis, MO, USA). All other chemicals were of reagent grade. Water was deionized and glass-distilled. Stock standard solutions were prepared by dissolving VMA and HVA, respectively, in 0.1 M acetic acid at concentration of 10 mM and stored at -30\u00b0C. The working standard solution was prepared daily by dilution of the stock solutions with 0.1 M acetic acid. Electrochemical reactor A glassy carbon electrode was used as the electrochemical reactor (ECR) (Fig. 1). The reactor consisted of a supporting electrode (platinum wire), a reference electrode (silver-silver chloride) and a working electrode 3, filter stopper; 4, saturated potassium chloride sol.; 5, reference electrode (silver-silver chloride); 6, silicone rubber; 7, supporting electrode (platinum wire); S, 0-ring; 9, Teflon block; 10, HPLC eluent. 352 ANALYTICAL SCIENCES JUNE 1990, VOL. 6 packed with glassy carbon grains (ca. 300 mesh, Tokai Electrode Mfg. Co., Ltd., Tokyo, Japan) by drytapping"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0001197_978-3-658-12701-5-Figure2.6-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0001197_978-3-658-12701-5-Figure2.6-1.png",
+        "caption": "Figure 2.6: NURBS example \u2014 modified curve",
+        "texts": [
+            " In general, a convex set of two points xu\ufffd and xu\ufffd is their connecting line, i.e. conv (xu\ufffd,xu\ufffd) = {xu\ufffd + \ud835\udf06 (xu\ufffd \u2212 xu\ufffd)} , 0 \u2264 \ud835\udf06 \u2264 1. In the two-dimensional example in Figure 2.1, the set \ud835\udc9c is a convex set of all points xu\ufffd. In Figure 2.2 this is not the case, \u212c is only the convex set of points x1 and x3, and x2 and x5 but not for any other combination of the depicted points. The convex hull of a set of points is the convex set of minimum size. Examples for the convex hull property can be found in Figure 2.5 and Figure 2.6 on page 13. \u2022 Local approximation: If a control point du\ufffd is moved, only the portion of p (\ud835\udc61) on the interval \ud835\udc61 \u2208 [\ud835\udc61u\ufffd, \ud835\udc61u\ufffd+u\ufffd+1) is affected. An example for the local approximation property can be found in Figure 2.6 on page 13. A computation method for NURBS that is particularly interesting for computer implementation is the recursion formula by de Boor, Cox and Mansfield. The simple case of this recursion is called a base function, which in this case is a constant that is either one or zero, i.e. \ud835\udc410 u\ufffd = \u23a7{ \u23a8{\u23a9 1 \ud835\udc61u\ufffd \u2264 \ud835\udc61 < \ud835\udc61u\ufffd+1 0 otherwise where \ud835\udc57 = 0, \u2026 , \ud835\udc5a\u22122. The general case that reduces the functions \ud835\udc41u\ufffd u\ufffd towards the simple case \ud835\udc51 = 0 is \ud835\udc41u\ufffd u\ufffd (\ud835\udc61) = \ud835\udc61 \u2212 \ud835\udc61u\ufffd \ud835\udc61u\ufffd+u\ufffd \u2212 \ud835\udc61u\ufffd \ud835\udc41u\ufffd\u22121 u\ufffd (\ud835\udc61) + \ud835\udc61u\ufffd+u\ufffd+1 \u2212 \ud835\udc61 \ud835\udc61u\ufffd+u\ufffd+1 \u2212 \ud835\udc61u\ufffd+1 \ud835\udc41u\ufffd\u22121 u\ufffd+1 (\ud835\udc61) where \ud835\udc51 is the local degree of the polynomial functions, \ud835\udc5b is the maximum degree of \ud835\udc41u\ufffd u\ufffd (\ud835\udc61) and \ud835\udc5a is the number of knots",
+            "11) yields \ud835\udc412 u\ufffd (\ud835\udc61) = \u23a7 { { { \u23a8 { { { \u23a9 u\ufffd\u2212u\ufffdu\ufffd u\ufffdu\ufffd+2\u2212u\ufffdu\ufffd u\ufffd\u2212u\ufffdu\ufffd u\ufffdu\ufffd+1\u2212u\ufffdu\ufffd \ud835\udc61u\ufffd \u2264 \ud835\udc61 < \ud835\udc61u\ufffd+1 u\ufffd\u2212u\ufffdu\ufffd u\ufffdu\ufffd+2\u2212u\ufffdu\ufffd u\ufffdu\ufffd+2\u2212u\ufffd u\ufffdu\ufffd+2\u2212u\ufffdu\ufffd+1 + u\ufffdu\ufffd+3\u2212u\ufffd u\ufffdu\ufffd+3\u2212u\ufffdu\ufffd+1 u\ufffd\u2212u\ufffdu\ufffd+1 u\ufffdu\ufffd+2\u2212u\ufffdu\ufffd+1 \ud835\udc61u\ufffd+1 \u2264 \ud835\udc61 < \ud835\udc61u\ufffd+2 u\ufffdu\ufffd+3\u2212u\ufffd u\ufffdu\ufffd+3\u2212u\ufffdu\ufffd+1 u\ufffdu\ufffd+3\u2212u\ufffd u\ufffdu\ufffd+3\u2212u\ufffdu\ufffd+2 \ud835\udc61u\ufffd+2 \u2264 \ud835\udc61 < \ud835\udc61u\ufffd+3 0 otherwise \u2200 \ud835\udc57 = 0 \u2026 4. Figure 2.4 on the next page shows the B-spline basis functions \ud835\udc41u\ufffd u\ufffd for \ud835\udc57 = 0 \u2026 \ud835\udc5a\u2212\ud835\udc5b\u22122 = 4 and \ud835\udc51 = 0 \u2026 \ud835\udc5b = 2. The weighted functions \ud835\udc45u\ufffd u\ufffd can be easily computed using (2.3) on page 6. Following (2.1), the NURBS function p (\ud835\udc61) can now be obtained and visualized, see Figure 2.5. To show the local approximation property from Section 2.1.1, one of the control points from the example is modified, i.e. d3 = \u239b\u239c \u239d 4.5 1.5 \u239e\u239f \u23a0 . The result is depicted in Figure 2.6 on the next page where the modified curve pmod (\ud835\udc61) only differs from the original curve p (\ud835\udc61) in the area between the adjoining control points d2 and d4. For the approximation of a series of \ud835\udc41 \u2265 \ud835\udc5a \u2212 \ud835\udc5b \u2212 1 points pu\ufffd at parameters \ud835\udf0fu\ufffd \u2208 [\ud835\udc4e, \ud835\udc4f] with weights \ud835\udc64u\ufffd where \ud835\udc58 \u2208 {0, 1, \u2026 , \ud835\udc5a \u2212 \ud835\udc5b \u2212 3, \ud835\udc5a \u2212 \ud835\udc5b \u2212 2} a NURBS curve \u0303p (\ud835\udc61) of degree \ud835\udc51 with \ud835\udc5a knots \ud835\udc61u\ufffd where \ud835\udc58 \u2208 {0, 1, \u2026 , \ud835\udc5a \u2212 2, \ud835\udc5a \u2212 1} is to be found such that \u0303p (\ud835\udc61 = \ud835\udf0fu\ufffd) \u2248 pu\ufffd, see [7]. The goal is to minimize the sum of error squares of the approximation \u0303p with respect to the given points pu\ufffd at the given parameters \ud835\udf0fu\ufffd, i"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0000082_dscc2015-9762-Figure1-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0000082_dscc2015-9762-Figure1-1.png",
+        "caption": "Figure 1. This knee extension neuroprosthesis uses electrical stimulation of the quadriceps muscles, V , to elicit a knee extension torque, \u03c4ke. The angle \u03b8 is the angle of the lower leg relative to the equilibrium position, \u03b8eq, which is located relative from vertical.",
+        "texts": [
+            " Because an NMPC method provides approximate optimal control it may reduce the amount of stimulation required to produce a desired motion, thus reducing the effects of FES-induced muscle fatigue. Potentially, the proposed NMPC method may benefit FES-based gait restoration devices by increasing walking durations. The dynamics of a leg extension musculoskeletal system can be described as J\u03b8\u0308+G\u2212 \u03c4p = \u03c4ke, (1) where \u03b8, \u03b8\u0307, \u03b8\u0308 \u2208 R are the angular position, velocity, and acceleration of the lower leg (shank and foot) relative to equilibrium as illustrated in Fig. 1, J is the moment of inertia of the lower leg, and G(\u03b8) = mglc sin(\u03b8+ \u03b8eq) is the gravitational torque. In the gravitational torque m is the mass of the lower leg, g is gravitational acceleration, lc is the distance from the knee joint to the center of mass, and \u03b8eq is the equilibrium position of the lower leg relative to vertical. The passive musculoskeletal torque of the knee joint, \u03c4p(\u03b8, \u03b8\u0307), is modeled as \u03c4p = d1(\u03c6\u2212\u03c60)+d2\u03c6\u0307+d3ed4\u03c6\u2212d5ed6\u03c6, where the anatomical knee joint angle, \u03c6 \u2208 R, is defined as \u03c6 = \u03c0 2 \u2212 \u03b8\u2212 \u03b8eq"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0002252_s0890-6955(98)00006-6-Figure2-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0002252_s0890-6955(98)00006-6-Figure2-1.png",
+        "caption": "Fig. 2. (a) Detailed model of the shear plane area A and the friction area Q of Fig. 1, (b) chip flow angle hc9.",
+        "texts": [
+            " The shear plane area A includes the area of (1) besides the cylindrical area of the tool nose [14,16]. 3. nose radius of the tool (R) is larger than the feedrate (f), R\u00de0 R . f, according to the depth of cutting, which can be subdivided into three parts: (a) d . R, (b) d = R and (c) d , R, as shown in reference [14]. In this section the case of small radius, that is case 2, R\u00de0 R , f, will be evaluated. The shear plane A and the projection area of the cutting cross-section Q, for the nose radius of the tool with a chamfered main cutting edge is shown in Fig. 2(a). The calculation of shear area A and friction area Q is based on the following Eqs (2)\u2013(9), [14,16]. R , {f\u00b7cosCs/[1 + sin(Cs 2 Ce)]}, shear area A. (Figs 1 and 2) A = A1 + A2 + A3 + As, (2) A1 = 0.25 \u00d7 [4a2n2 2 (a2 + n2 2 c2)2]1/2, (3) and As = (0.5\u00b7W2 e\u00b7cos2aS1\u00b7tanCs)/(cosab\u00b7sinfe) (6) where A1 is the area of triangle BCK, A2 is the area of cyclindial CC9EK, A3 is the area of trapezoid C9EFD9, and As can be seen from Fig. 2(a) to be a triangular shape of the secondary chip DD9YJ. The area of the projected cross-section Q is equal to Q1 + Q2 + Q3 (Figs 1 and 2), where Q1 = {f1[d 2 0.5\u00b7f1sinCscosCs 2 R(1 2 sinCs 2 cosCs)] 2 R2[1 2 p/4 + tan((Ce 2 Cs)/2) 2 0.5(Ce 2 Cs)] 2 0.5[f1cosCs 2 R(1 + tan((Ce 2 Cs)/2))]2\u00b7 (7) [tan(Ce 2 Cs)/cosaS2\u00b7cosab)]};(Q1 is the area of the BC9DL), Q2 = WecosaS1(d/cosCs 2 WecosaS1\u00b7tanCs)/cosab(Q2 is the area of the rectangle CC9DD9), (8) and Q3 = 0.5\u00b7W2 ecosaS1tanCs/cosab(Q3 is the area of triangle DD9Y) (9) Expressions for a, n, c, e, g, hc9, ds, f1 and P are given in Appendix A; and the angles F and hc9 are defined in Fig. 2(b). Abdelmoneim et al. [2] have suggested that the tool edge may in fact wear rapidly to form a cylindrical surface with a larger radius and an adjoining flat wear land. Takeyama and Murata [5] further showed that the mechanism of tool wear during turning can be classified into two basic types: the rate process and mechanical abrasion. According to Takeyama et al. [17], calculation of cutting force components could be one approach to confirm the wear behaviour of lathe tools during the cutting process",
+            " (A9) Coefficients of the tool with nose radius (R\u00de0,R , f) with wear are: a1 = {fCM 2 (l1 + l2) 2 R2sin2uR2 2 R3sin2uR3 2 R2sin(Ce + Cs) 2 2R2tan[Ce 2 Cs)/2]}(B1) {cos2ae + [tanaS2sinab + tan(Ce 2 Cs)]2cos2ab}1/2/(cosaS2cosab) n1 = [fCM 2 (l1 + l2) 2 R2sin(Ce 2 Cs) 2 R22uR2 2 R32uR3 + R2(Ce 2 Cs)]/(coshc9sinfe) (B2) c1 = (e2 1 + g2 1 2 2e1g1sinab)1/2 (B3) e1 = {fCM 2 (l1 + l2) 2 [R2sin2uR2 + R32uR3 2 R2sin(Ce + Cs) 2 2R2\u00b7tan(Ce 2 Cs/2)]} (B4) {tanhccosab 2 [tanas2sinab + tan(Ce 2 Cs)]cosaS2}/(cosaS2\u00b7cosab) g1 = [fCM 2 (l1 + l2) 2 R2sin(Ce 2 Cs) 2 R22uR2 2 R32uR3 + R2(Ce 2 Cs)]\u00b7 (B5) (cotfetanae)/coshc9 const1 = {cos2aS2 2 sin2fe[tanhccoshc9 2 cosaS2cosab\u00b7(cotfe + tanab)]2}1/2/ (B6) (sinfecosaS2cosab) F = p 2 + Ce + Cs (Fig. 2(a)). (B7) k1 = d/cosCs + (l1 + l2)tanCs 2 h1 2 h2 (B8) i1 = d/cosCs 2 h1 2 h2 + h3 + h4 + (l1 + l2)tanhc9 2 fCM[tanhc9 2 tan(Ce 2 Cs) + tanCs] (B9) 2 (l1 + l2 + l3 + l4)tan(Ce 2 Cs) j1 = fCM (B10) [1] H.T. Young, P. Mathew, P.L.B. Oxley, Allowing for nose radius effects in predicting the chip flow direction and cutting forces in bar turning, Proceedings of Institute of Mechanical Engineers (IMech E) 201 (C3) (1987) 213\u2013216. [2] M.E. Abdelmoneim, R.F. Scrutton, The tribology of cutting tools during finish machining, Wear 25 (1973) 45\u201353"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0003997_20020721-6-es-1901.00325-Figure1-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0003997_20020721-6-es-1901.00325-Figure1-1.png",
+        "caption": "Fig. 1. An admissible trajectory from x0 to xT .",
+        "texts": [
+            " (Kaya and Noakes, 1996)) to refer to such switching time computation problems. A segment of the trajectory x(t) corresponding to the time interval from tk\u22121 to tk represents a smooth arc. The dynamical system (1) can also be written as a sequence of initial value problems dx dt = fk(x), t \u2208 (tk\u22121, tk\u22121 + \u03bek), x(tk\u22121) = { x0, if k = 1, x(tk\u22121 \u2212 0), if k > 1, (2) where \u03bek is the time-duration of the k-th arc, or simply the k-th arc-time, given by \u03bek = { tk \u2212 tk\u22121, if k = 1, 2, . . .N, tf \u2212 tN , if k = N + 1. A sketch of a trajectory for an admissible control is shown in Figure 1. We will call such a trajectory an admissible trajectory. The STC problem is usually formulated in terms of arc-times as in (Kaya and Noakes, 1996). The segment of a trajectory x(t) corresponding to the interval [tk\u22121, tk] can be parametrised as x(tk\u22121 + \u03c4), where \u03c4 \u2208 [0, \u03bek] (k = 1, . . . , N + 1). We will use the notation x(\u03c4 ; \u03bek\u22121, . . . , \u03be1) \u2261 x(tk\u22121 + \u03c4) , (3) which explicitly shows that the behaviour of x in the k-th arc also depends on the previous arctimes. Note that for the first arc this notation simply becomes x(\u03c4) as there are no previous arcs"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0001146_jctn.2015.4303-Figure7-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0001146_jctn.2015.4303-Figure7-1.png",
+        "caption": "Fig. 7. Total thermal deformation field at 80 r/min.",
+        "texts": [],
+        "surrounding_texts": [
+            "AND INITIAL CONDITIONS The 3D model of h drostatic bearing table is shown in Figure 1. This paper built the finite element model of worktable by using the software named ANSYS Workbench. Because the rotary worktable has periodicities, it just needs to mesh 1/24 of one single cycle. As shown in Figure 2. The initial conditions and boundary conditions are necessary to solve specific problems. The center of the worktable which is studied by this paper is touching radial cylindrical roller bearing, so it is applied by Fixed Support code to the fixed boundary conditions in inner surface of the circumference. The distribution of corresponding J. Comput. Theor. Nanosci. 12, 3917\u20133921, 2015 3919 Delivered by Publishing Technology to: Chinese University of Hong Kong IP: 117.253.218.180 On: Fri, 26 Feb 2016 00:22:32 Copyright: American Scientific Publishers R E S E A R C H A R T IC L E Fig. 9. Z thermal deformation on upper surface at 80 r/min. temperature field must be applied to the simplified model to calculate thermal deformation field, and its specific command is Thermal Condition code. As same as temperature field, it needs to use Commands code to apply symmetric boundary conditions to both sides of the model. The specific settings are shown in Figure 3."
+        ]
+    },
+    {
+        "image_filename": "designv11_60_0003173_s0141-6359(02)00117-4-Figure14-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0003173_s0141-6359(02)00117-4-Figure14-1.png",
+        "caption": "Fig. 14. Self-locking taper effect.",
+        "texts": [
+            " This section describes the mechanics of frictional interlocking between the coupling fingers. The idea is extended to determine the angle of the sloping faces of the finger required for self-locking. During our experimental phase, our first design displayed relative slip between the fingers of the male and the female elements. As explained in Section 3.4, relative slip between fingers is detrimental to the torsional stiffness of the coupling. The design was analyzed and the reason for the slip was identified and corrected in the next design. Fig. 14 shows a block lying on a flat surface which is being acted upon by an external force F at an angle \u03b2 with the surface. If \u00b5s is the coefficient of static friction between the block and the surface, then, neglecting gravity, the block will remain stationary as long as the following conditions are valid: N = F sin(\u03b2) (19) \u00b5sN > F cos(\u03b2) (20) Therefore, a simple condition for self-locking is |tan(\u03b2)|\u00b5s > 1 (21) Given the tangential force at the mean radius of the coupling is the external force acting on the inclined surface, as shown in Fig. 14, let \u03c6 the angle made by the side face of the fingers with the radial line. Using the self-locking result of Eq. (21), we can determine an inequality relating \u03c6 and \u00b5s. \u03c6 = \u03c0 2 \u2212 \u03b2 (22) Therefore, \u2223\u2223\u2223tan (\u03c0 2 \u2212 \u03c6 )\u2223\u2223\u2223 \u00b5s > 1 (23) \u00b5s > |tan(\u03c6)| (24) For a given \u00b5s, the maximum value of the inclination angle is given by \u03c6 < tan\u22121(\u00b5s) (25) Finite element models were created and evaluated using ProENGINEERTM and ProMECHANICATM. Torsional loads were applied as point loads. Note that ProMECHANICATM automatically distributes a point load over a small circular area; this area is small relative to the size of the components, and does not significantly change our results"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0002412_942026-Figure5-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0002412_942026-Figure5-1.png",
+        "caption": "Figure 5. Schematic of the in-situ spectrometer",
+        "texts": [
+            " Location of optical access in cylinder of Caterpillar 1Y73 (Viton O-rings provide an oil seal to the liner and a water seal to the cooling jacket). The window surface was machined to a nominal cylindrical radius (67.5mm) corresponding to that of the liner bore and a final flush fitting was achieved by the use THE IN-SITU SPECTROMETER - The spectrometer provides the means of launching the interrogating IR beam through the liner window and of collecting the reflected return light for spectral analysis. A schematic of the design is provided in figure 5. ZnSe meniscus lenses of focal length 25mm served to relay an image of the source onto the moving piston and to collect the return light, all at a magnification of 1 : 1. Operation was maintained at f/l for high optical throughput by locating a lens immediately behind the liner window. The source itself was an electrically-heated (24W) ceramic 'globar' which approximates to a black body operating at 1500 K. Forward and return light signals were separated by means of a germanium 60140 beamsplitter, coated on one side for minimal reflection in the range 2850 cm-' to 700 cm-'"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0002258_1.1332396-Figure4-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0002258_1.1332396-Figure4-1.png",
+        "caption": "Fig. 4 Bearing geometry",
+        "texts": [
+            " Agreement of the dynamic modal method with the quasi-static model is illustrated by steady response to a steady applied load. With this initial condition, the same steady load is applied, released, and re-applied to produce a dynamic simulation where the externally applied load varies rapidly. In this case, the dynamic modal method results differ appreciably from results generated with the quasi-static modal method. Transactions of the ASME s of Use: http://www.asme.org/about-asme/terms-of-use Downloaded F 5.1 Physical and Finite Element Arrangement. The bearing geometry is shown in Fig. 4, and all other bearing and fluid properties are given in Tables 1 and 2. Next, the structural mesh and the bearing fluid film mesh are shown in Fig. 5. The structural mesh was created with the PATRAN 3 finite element modeler and utilizes HEX8 elements. Stiffness and mass matrices were created and reduced with the MSC/NASTRAN 67 finite element analysis computer program, as documented by MacNeal @18#. Finally, Fig. 6 shows the first seven axially symmetric mode shapes, where a total of 483115528 modes were computed"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0001996_s0736-5845(98)00026-x-Figure17-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0001996_s0736-5845(98)00026-x-Figure17-1.png",
+        "caption": "Fig. 17. Knots of the spline curve fitted to the engine\u2019s path.",
+        "texts": [],
+        "surrounding_texts": [
+            "The programming methods, path planning, configuration planning and motion planning will now be used for offline-programming the world\u2019s largest robot, the mobile aircraft cleaning manipulator SKYWASH SW33. The SW33\u2019s redundant kinematic consists of 5 main axes, 5 axes denoting the SW33\u2019s wrist and one axis denoting the angle of the SW33\u2019s rotating brush. The SW33\u2019s radius of action is 33 m. The first task to be programmed is moving the cleaning brush along a path on the hull of an aircraft. The path length on the front part of the aircraft\u2019s torso is 65.0 m. The 1350 data points representing the path are ordered in a relative distance of 0.05 m (see Fig. 14). The second task to be programmed is moving the cleaning brush along the aircraft\u2019s engine with a path length of 45.0 m. The 900 data points representing the engine\u2019s path are also arranged in a relative distance of 0.05 m (see Fig. 15). Fig. 15. Path of datapoints on the aircraft\u2019s engine. First a curve is fitted to the data points by a piecewise cubic polynomial used in the robot control. The tolerance of the fitted curve to the data points is 0.05 m in position and 0.5\u00b0 in orientation. Using the path planning method the number of knots on the torso\u2019s path is 27, the number of knots on the engine\u2019s path is 38, with respect to the required tolerance (see Figs. 16 and 17). The SW33\u2019s configuration will be calculated in knots having specified geometrical characteristics. The constraint for maximum pressure is 2.7]107 N/m2 and for minimum distance between SW33 and aircraft is 1.5 m each main axis and 0.5 m each wrist axis. Using the global configuration planning method we get the configurations in the six specified knots in the case of torso washing and in the 10 knots in the case of engine washing. The profit of putting manipulability and desired configuration into the optimization criteria show Figs. 18 and 19, because two globally planned configurations in order do not pass singularities. Energy optimal configurations in the rest 27!6\"21 and 38!10\"28 knots will be planned within the Fig. 16. Knots of the spline curve fitted to the torso\u2019s path. Fig. 20. Local configuration planning method used for torso washing. Fig. 21. Local configuration planning method used for engine washing. optimal configuration space obtained by the global method. Figs. 20 and 21 show how the method works between a pair of globally planned configurations. Controlling the collision distance between SW33 and aircraft in all knots results 1.72 m minimum distance of the main axes and 0.54 m minimum distance of the wrist axes. Finally, the time optimal velocity profile for the SW33\u2019s motion along the specified paths was calculated, where the following constraints are given. With respect to the constraints in Figs. 22\u201424 the SW33\u2019s time optimal velocity profile gained by the motion planning method leads to a complete motion time of 179 s along the torso path and 281 s along the engine path. Driving nearly always with maximum endeffector speed of 0.4 (m/s) at the torso path shows the effect of progressing manipulability and energy cost towards an optimum within the configuration planning method (see Fig. 25). The process speed of cleaning the engine\u2019s path is limited by 50% of the maximum speed, Fig. 26 shows that this speed is nearly reached along the whole path. Analyzing the dynamics of the flexible subsystem we get the joint torque trajectories q (t) of the wrist axes, with and without the average joint torque each, by Figs. 27 and 29. Figs. 28 and 30 show the FFT-analysis of the robot, washing the aircraft\u2019s torso and washing the aircraft\u2019s engine. In Figs. 27 and 29 1 (kNm) denotes 1000 (Nm), in Figs. 28 and 30 F(- )\"1 is equivalent to q\"3 (Nm). Knowing that the robot\u2019s natural frequencies are beginning at 2.0 Hz, Figs. 28 and 30 ensure that there is no Fig. 29. The robot\u2019s joint torque trajectory during washing the torso. vibration stimulation by the programmed motion, because the amplitude F (- ) at - 0 \"2.0 (1/s) is far below the influence of motion damping by joint friction."
+        ]
+    },
+    {
+        "image_filename": "designv11_60_0001235_b978-0-08-099425-3.00005-4-Figure5.21-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0001235_b978-0-08-099425-3.00005-4-Figure5.21-1.png",
+        "caption": "FIGURE 5.21",
+        "texts": [
+            " It is clear that the direction of slip changes several times as tread moves through the contact patch resulting in the distribution of longitudinal shear stress of the type shown at the bottom of Figure 5.20. The shear stress is plotted to be consistent with the SAE reference frame and is not symmetric; the net effect being to produce an overall force, the rolling resistance, acting in the negative XSAE direction. It should be noted that the two-dimensional model presented is not fully representative as components of lateral slip are also introduced in a free rolling tyre due to deformation of the side walls as shown in Figure 5.21. As the tyre carcass deforms in thevicinityof the contact patch thedeformationof the side walls creates additional inwards movement of the tread material (Moore, 1975). Generation of slip in a free rolling tyre. Lateral distortion of the contact patch for a free rolling tyre. This causes the contact patch to assume anhourglass shape creating an effect referred to as \u2018squirm\u2019 (Gillespie, 1992) as the tread material moves through the contact patch. Before moving on to consider the driven or braked tyre we will now consider the rolling resistance forces generated in a free rolling tyre"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0000962_tmag.2016.2636208-Figure5-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0000962_tmag.2016.2636208-Figure5-1.png",
+        "caption": "Fig. 5. Linear motor\u2014the view.",
+        "texts": [],
+        "surrounding_texts": [
+            "The first and second theorem presentations are carried out based on linear motor forces analysis (Fig. 4 and 5). Linear motor dimensions and parameters of work are a = 0.02 m (conductive layer width), l = 1 m (rotor length), g = 0.01 m (gap width), \u03b81 = 4870 A [magnetomotive force (mmf) first harmonic magnitude], f1 = 5 Hz, Y = 1 m (pair-pole length), \u03b3 = 30 \u00b7 106 S/m (carriage conductivity), \u03bdx x = 0.4\u00b7\u03bd0 (cross-layer axis reluctivity), \u03bdyy = 0.4\u00b7\u03bd0 (move direction axis reluctivity), and different reluctivities \u03bdxy , \u03bdyx (see Tables V and VI). Fig. 6(a) and (b) confirms [see Table V for the cases a) and b)] that if the condition (3) is satisfied, the first theorem thesis is fulfilled. Lorentz force (volume integral) equals to the Maxwell force (surface integral) and FeM = 0. The coenergy method is applied only for checking the Maxwell force value. The cases c and d show that if condition (3) is not satisfied first, theorem cannot be applied, and in Fig. 6(c) and (d), both the forces are different. The material (anisotropy) force FeM = 0. First theorem is biconditional (equivalence). Examples for the second theorem are presented in Fig. 7 for the same data as for linear motor in case b), and are given in Table VI."
+        ]
+    },
+    {
+        "image_filename": "designv11_60_0003681_robot.1991.131809-Figure1-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0003681_robot.1991.131809-Figure1-1.png",
+        "caption": "Figure 1: Description of 2-DOF Hopping Model",
+        "texts": [
+            " The 2 DOF Model The total system mass is modeled as a point mass with no rotational inertia. The leg actuator and energy storage mechanism is an air cylinder with restoring force proportional to the inverse of the piston displacement, like the 1-DOF model. The leg and cylinder combination are assumed to be massless. We will use a polar coordinate system when the robot is in ground contact and a Cartesian system for ballistic motion. Let T denote the leg length and e denote the \u201chip angle\u201d during stance (Figure 1). The motion of the robot may be decomposed into four distinct phases : thrust, decompression, flight and compression (Figure 2). i. Thrust Phase. At time t j the leg is a t its minimum radial length, ~j = ~ ( t j ) (i.e., +( t i ) = 0 ), the control valves are opened and a constant supply pressure CH2969-4/91/0000/1392$01 .OO 0 1991 IEEE 11. ... 111 is connected to the leg cylinder for a fixed time, 6 t . This applies a constant radial force, ?, to the mass. The equations of motion in this phase are: m(i: - r e 2 ) - "
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0001817_0141-0296(93)90032-y-Figure6-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0001817_0141-0296(93)90032-y-Figure6-1.png",
+        "caption": "Figure 6 Special guyed structures",
+        "texts": [
+            " On some projects, the shear and bending loads applied to the single mast of the structures in Figure4(e,g) were large enough to produce a single leg that weighed more than the two masts of a guyed V tower. The oft prescribed demand for torsional strength to resist the possible loads of broken wires or to resist the complex of loads that might be imposed by failure of an adjacent structure (the duty of failure containment) can create a problem as some of the guy systems with crossed guys, as in Figure 4 (e, g) and Figure 6(d), will permit a rotation that reduces the spacings of the resisting guys, the guy tensions and strains then increase with more rotation etc. until failure or snap through may occur. For this reason some designers have used auxiliary or longitudinal secondary crossarms that will ensure that under torque forces, the spacings of the acting guys will move apart and stabilize the problem as shown in Figure 4(h) and Figure 6(e). These extra arms do not add significantly to tower weights as they keep some of the heavier guy loads out of the upper part of the structure but they can and do make the tower lay out and assembly on the ground more difficult. The tower shown in Figure 4(a) was used in Finland as early as 1927, fabricated as a lattice work in steel. In the 1960s, a similar tower of aluminium was being used in North America for helicopter transport and erection on lines of remote or difficult access. Thousands were installed in Western Canada and Alaska in areas of permafrost or unstable ground conducive to frost heave",
+            " Improved procedures for the construction of the CRS with adoption of the single rope suspension for nonice areas and of the method of precutting guys and thus automatic alignment of the masts (see below) should lead to much greater use of this crossarmless type of tower for future lines. Some special guyed towers Some guyed structure types do not fit easily into the categories of guyed masts, guyed rigid frames nor of the guyed and pinned masts of Figures 4 and 5. These special types are shown in Figure 6. Several interesting attempts have been made to reduce the lengths of the masts of the guyed V tower by converting into forms of pinned or hinged Ys for the reason that three half-length masts should be and are much less costly than two full-length masts while retention of the pinned or hinged mast principle prevents the problems of the guyed rigid structures. The guyed V on a guyed post is shown in Figure 6(a) where the short centre mast requires four small guys to hold it in position. Three half-length masts replace two full-length masts at the expense of an extra four small guys. The saving in mast material is evidently greater than the cost of the extra guy system although the tower is now critically vulnerable to the failure of one of these small guys. This contravenes a principle that holds for almost all other guyed towers in that after stringing, they will remain upright and still be able to carry at least 50% of design loads with one guy removed",
+            " It is an interesting structural feature that the hinged connection to the lower vertical leg ensures that the loads in the two upper legs will always be equal and this reduces the uncertainties regarding the load distribution into the crossarm and guy system and ensures that each o f the mast sections works fully under all loads. Erection was made simple by attaching two of the inner guys (one from each side of the tower) to the base of the centre mast to form a temporary rigid Y, erecting and plumbing the tower with the four outer guys and then transferring the two inner guys from the base of the mast to their anchor positions. The guyed V and inverted delta or 'T ' ~0 is shown in Figure 6(c). The need to build compact lines and to control the magnetic fields beneath a line that some believe may have influences on health have focused attention on the inverted delta or T arrangement of the phases. One method of supporting a compact T is within a guyed V, a V of two tripods with a simple spreader strut to hold the mast tops apart. This tower holds great promise for magnetic field control, for building on narrow rights-ofway and for the benefits that compaction of the phases bring in reducing the reactance of long distance lines and increasing power transfer capability. So far, no one seems to have found a totally acceptable way of suspending the three phases, as there is an evident reluctance to suspend the lower phase from the other two with phase-to-phase insulators. However, it is but a matter of time before someone finds an acceptable suspension arrangement and the new guyed V with T becomes one of the standards for EHV lines. Two guyed single mast direct current (two pole) towers are shown in Figure 6 (d, e). Though but single masts, these towers are more appropriately classified as guyed rigid frames as the four guys do not converge on a single point. At first glance this is a simple and straightforward solution but there are difficult spatial problems of trying to maintain electrical clearances between guys and conductors and at the same time of trying to have the guys centric with the line of action of the resultant of the major loads and the centre line of the mast. If the guys extend beyond the mast and attach to the opposite crossarms, the tower under torque load will rotate and the lines of action of the resisting guys will come closer and guy tensions will increase, a problem mentioned above with regard to the guyed portal. It is common practice with this tower to use extra secondary arms to keep the guys apart although the ground assembly problems with these arms are causing some designers to look again at the need for the torque loading. The guyed X tower is shown in Figure 6(f). The problems of frost heave that are common to many northern lands have led to the development in Alaska of a special type of X tower that behaves as a rigid tower under vertical and transverse loads and is guyed longitudinally. Many hundreds of miles of these X towers are in place, some of lattice steel, of lattice aluminum and of tubular steel in some of the areas considered more environmentally sensitive. The usual footings are two single H piles on each of which is attached, by U bolts, the bracket on which the tower can pivot back and forth"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0000054_oceans.2015.7404427-Figure6-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0000054_oceans.2015.7404427-Figure6-1.png",
+        "caption": "Fig. 6: Lateral proje",
+        "texts": [
+            " Sever (6) ))0 , (7) | || || || | (8) (9) direction, and V generates the th two azimuth ns Per Minute les , rboard and port near relationship re 240 is the in (10)-(14). d. (10) (11) (12) (13) (14) arboard moment r of gravity and 3). For station e neglected. tion of the wind al models can be found in the literature [26][27][28] is based on [25]. For marine vehicles that have a f pressure is: 12 , Where 1.2 / is the apparent wind speed. Both apparen can be measured directly by th described in Section II.C. The wind forces and moments ca( )( )( ) , Where FwA and LwA are the fr windage areas (Fig. 5, Fig. 6). The by the computer with projected 3D apparent wind speed and angle of a be measured directly from the a section IIC. , , are empirical wind co and yaw direction respectively, and ( ) c ( ), ( )( ) c (2 ) In [25], it is suggested that 1.0, 0.7 , and c 0.2, 0. [29]. The one adopted here orward speed, the dynamic (15) air density, and is the t wind speed and direction e ultrasonic anemometer n be expressed as: (16) ontal and lateral projected y are calculated precisely model. and are ttack respectively, that can nemometer introduced in cted area"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0000187_icsens.2014.6985409-Figure2-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0000187_icsens.2014.6985409-Figure2-1.png",
+        "caption": "Fig. 2. Schemetic of half and full bridge circuit.",
+        "texts": [
+            " As the former research introduces, the strain on the diaphragm part shows a sinusoidal pattern which can be easily modeled than other locations [5]. We have two sets of sensor configurations, a full bridge and a half bridge as shown in Figs. 1 and 2. We use the rosette type strain gauge of Micro Measurement. In Fig. 1, the sensor components grouped with dashed line are connected with a full bridge circuit like previous researches. The sensor component highlighted with solid line is connected with a half bridge. The circuits used to each component are illustrated in Fig. 2. When the harmonic drive is in operation and an external force or torque is exerted on the harmonic drive, tangential and radial directional strains are generated. The models of strain which affect the sensing component can be modeled as (1). 1 1 1 2 2 2 t r t r \u03b5 \u03b5 \u03b5 \u03b5 \u03b5 \u03b5 = + = \u2212 + (1) The subscript 1 and 2 represents each strain gauges on a single sensing component. The subscript t and r means tangential and radial strain component. The reducing operation mainly affects to tangential strain and the radial strain affects to the other [5]"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0003013_a:1008115522778-Figure16-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0003013_a:1008115522778-Figure16-1.png",
+        "caption": "Figure 16. Experimental results for misalignment compensation: initial error = 1 mm.",
+        "texts": [
+            " It is checked every step of the corrective motion whether misalignment compensation is completed or not. Figure 15 shows the initial rough shapes of the part used in the experiments for misalignment compensation in cylindrical peg-in-hole tasks. The experiments to compensate for lateral misalignment only were performed in nine cases as shown in Figure 15(a)\u2013(c). And the experiments to compensate for inclination error and to insert a part into its mating hole were performed in one case as shown in Figure 15(d). Figure 16 shows the experimental results for lateral misalignment compensation when em = 1 mm and cr = \u22122, 1, 4 mm. The em denotes lateral misalignment, namely, the distance between a hole center and the center of the bottom of its respective mating part. And the cr denotes the center position of the upper surface of a part. When em is small, the error in estimating a hole center using the implemented visual sensing system becomes large because the visible part of the hole is a little [23]. Figures 16(a) and (b) show the large error between the actual prescribed value and the measured value by the sensing system inferred from the neural network",
+            " Another reason is that a initial fitted part into its mating hole does not escape easily from the hole due to its deformation although the corrective motion is going on. Figure 17 shows the experimental results when em = 4 mm and cr = 1, 4, 7 mm. The error between the actual prescribed value and the measured value of em, and the error between the inferred corrective motion and the actual motion until misalignment compensation is actually accomplished are not large compared with the case em = 1 mm in Figure 16. When cr = 1 mm, the lateral misalignment was compensated by one times of corrective motion inferred from the neural network. When cr = 4, 7 mm, the lateral misalignment was compensated by two times of corrective motion. Misalignment compensation was accomplished at the beginning of the second corrective motion, because the difference between the inferred motion and the actual motion until misalignment compensation is actually accomplished is not large in the first corrective motion m1. However, there is quite large difference in the second corrective motion m2. This is due to the same reason as em = 1 mm in Figure 16 because the bottom of the peg was moved near to the hole by m1. These successful compensation does not guarantee the success in more general tasks such as high speed or high precision assembly. However, the success rate of misalignment compensation in such tasks will be increased by considering more parameters in a neural net-based inference system. Figure 18 shows the experimental results when em = 7 mm and cr = 4, 7, 10 mm. Both the measurement error by the sensing system and the error in the inferred corrective motion are not large"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0003825_icsmc.1996.561326-Figure1-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0003825_icsmc.1996.561326-Figure1-1.png",
+        "caption": "Fig. 1. The excessive scheme of MGC.",
+        "texts": [
+            " MATHEMATICAL MODELS The model of manoeuvring SC takes into account: controlled orbital motion of SC\u2019s mass center around the Earth; spatial angular motion of a SC frame as a carring rigid bogy (RB); movements of flexible contructional units (solar array panels, communication antennae) about SC frame by means of two-gimbal gear drives; flexible-viscous fixation of the optical telescope\u2019s strnctnre on the SC frame; MGC in the form of minimumexcessive scheme of four GD\u2019s of the \u201d2-SPEED\u201c- type [9], represented on Fig. 1; MGC fixation on RB by means of a damping vibroabsorbing frame; environment external torques (gravity-gradient, aerodynamic, magnetic). The mathematical model of each GD describes the flexibility of gyrorotor\u2019s (GR\u2019s) ball bearings, the proper GR rotation dynamics with regard to it\u2019s static and dynamic disbalance; flexibility of gyroshell\u2019s (GS\u2019s) preloaded ball bearings; flexibility, dead band and kinematic defects in the gear; nonlinear dynamics of stepwise motor on the GD precession axis taking into account the dry friction torque as well as flexible GD fixation on the SC\u2019s frame"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0001774_1999-01-1767-Figure2-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0001774_1999-01-1767-Figure2-1.png",
+        "caption": "Figure 2. DU-Gates predicted transmission error",
+        "texts": [],
+        "surrounding_texts": [
+            "The overall simulation approach is shown in Figure 1. Three un-coupled models are used to calculate the excitation, dynamic shaft response and casing response as follows: \u2022 Tooth Contact Analysis: To predict the excitation from the gear mesh using an influence coefficient method \u2022 Dynamic Geared Shaft Model: To predict the dynamic response of the transmission internals using a time domain solution \u2022 Transmission casing models: To predict the vibration and radiated noise from the casing using FE and BE techniques A complete generic modelling approach was developed as described in the following sections. The approach was applied to the simulation of whine from an off-highway automotive single mesh transmission and was validated by measurements of casing vibration and radiated noise. SHAFT EXCITATION DUE TO GEAR MESHING - TOOTH CONTACT ANALYSIS \u2013 The process of establishing the excitation caused by the meshing gears begins with calculation of the quasi-static transmission operating loads in order to establish the range of gear mesh misalignments occuring in the gearbox assembly and estimate rolling element bearing stiffnesses for use in the subsequent dynamic shaft analysis. Using proprietary integrated transmission analysis software, the gears, shafts and bearings are modelled in three degrees of freedom. Inputs to the model include full shaft geometry, bearing specification and \u2018macro\u2019 gear geometry (module, pressure angle, facewidth, number of teeth, helix angle) for the mating gears. The flexibility of all the represented transmission components is modelled except for the gear bodies for which stiffness is assumed infinite. Quasi-static operating loads and deflections are calculated for the range of operating conditions of interest as well as estimates of the rolling element bearing stiffness which are derived from in-built catalogue data and special algorithms which calculate the stiffnesses of the bearings according to the loads and geometry of the gearbox. 3 The effects of casing deflections due to operating loads are included in the analysis by using a separate finite element model of the casing in conjunction with the above model. Predicted bearing reaction loads from the gear, shaft and bearing model are applied to the finite element model in order to calculate bearing housing deflections which are then fed back into the shaft analysis in order to calculate a new set of bearing loads and stiffnesses. If the casing is subject to external structural loading this too can be included as it may contribute to bearing house deflections and thus misalignment of the gears. This iterative process is repeated to obtain a convergent solution. Even though these two models are not coupled, convergence was achieved in three iterations for the transmission used in the validation exercise described later in this paper. For the next stage the DU-GATES tooth contact analysis software [Reference 4] was used to calculate the transmission error and mesh stiffness variation for input into the dynamic model. This program uses a finite element based code to initially calculate the direct and cross compliances (or influence coefficients) of all the contacting tooth flanks, in a loaded gear pair, due to bending and shear stresses. The detail tooth metrology data, of an average tooth, are also input into the program. These data include the net effect of both profile modifications (e.g. crowning and profile relief) and manufacturing errors. The tooth contact analysis program then solves iteratively for transmission error and mesh stiffness while including the effects of elastic compliances, tooth flank topology and non-linear (Hertzian) contact deflections. The solution is carried out at 64 quasi-static positions within a single meshing cycle and at a range of gear misalignments. A three dimensional map of transmission and mesh stiffness at each load condition is created, as shown in Figures 2 and 3, for input to the dynamic model as described in the following section. DYNAMIC BEARING LOAD - DYNAMIC GEARED SHAFT MODEL \u2013 An approach and software was developed by Ricardo to predict the three dimensional dynamic behaviour of the geared shaft system. Within the software the system is represented as a series of lumped masses, each with six degrees of freedom, connected by massless flexible elements. The gear and bearing inner race masses and inertias are included in the lumped masses. The compliances of the following elements are included: \u2022 The gear mesh stiffness, along the line of action, as calculated by the tooth contact analysis as a function of the gear pair misalignment and the phase of gear meshing (Figure 3) \u2022 The shaft flexibility calculated from the section properties of the shaft \u2022 The shaft support bearing compliance, in five degrees of freedom, as calculated by the quasistatic shaft and casing finite element analyses described in the previous section. Using excitation derived from the tooth contact analysis, the shaft vibratory response is calculated in the time domain in six degrees of freedom at each mass node of the model. At each time step, the nodal forces, displacements, velocities and accelerations from the previous step are used to calculate the tooth contact force, bearing reactions and centrifugal loads which are then applied to the relevant model node. The system equations of motion are then solved, using a variable step 'stiff' solver, to extract the new nodal forces, displacements, velocities and accelerations at this point in time. The key to the whole approach adopted for prediction of the system response is the method used to calculate the tooth contact force. The results of the tooth contact analysis detailed previously are input into the dynamic model in the form of maps of transmission error and mesh stiff- 4 ness across one tooth pass for a range of positive and negative gear misalignments. At each time step the transmission error and mesh stiffness corresponding to the rotational position of the gears is interpolated from the maps and used within the calculation of the dynamic component of the tooth force. The dynamic component is then added to the steady state force component resulting from the nominal torque to give a total force transmitted between the gears. The use of a time domain rather than a frequency domain solution allows the instantaneous vibratory response of the gear pitch points to be derived at each time step and used within the tooth force calculation in two ways. First, gear misalignment resulting from the shaft dynamic response at this time step can be calculated and used when interpolating for transmission error and mesh stiffness from the maps as described above. Second, the vibrational response along the line of action of the tooth force is used to modify the effect of the transmission error on its dynamic component. This coupling of the shaft dynamic response from one timestep into the calculation of the excitation force and thus the dynamic response in the next timestep is summarised in Figure 4. where : Ftc = total tooth contact force Fnt = tooth force due to nominal torque km = mesh stiffness dte = transmission error along the line of action dvib = vibratory displacement along the line of action ct = gear mesh damping vvib = vibration velocity along the line of action The effect of coupling shaft dynamic response is best illustrated by considering a simple example. A positive transmission error will lead to an increase in the tooth mesh force equal to the product of transmission error and mesh stiffness at each time step. However, if the vibratory displacement of the gears is such that the gear teeth move apart by the amount equal to the transmission error along the line of action, then the net dynamic component of force applied at the gear teeth will be zero. Conversely if the vibratory response is such that the gear teeth move together by an amount equal to the transmission error along the line of action then the net dynamic component of force applied to the gear teeth will be twice that calculated without considering dynamic effects. 5 Figures 5 to 8 illustrate the effect of coupling the vibratory response between timesteps in this way for a twin helical gear and shaft system. In figure 5 the transmission error and predicted pitch point vibrational response at 1050 rev/min are of a similar magnitude and almost in phase which results in a calculated peak dynamic tooth force of approximately 60% of that calculated when neglecting the instantaneous dynamic response, which is shown in Figure 6. At 1300 rev/min (Figure 7) the opposite situation occurs. Here the transmission error and vibrational response are out of phase resulting in a predicted peak dynamic tooth force of more than twice that which is calculated when the vibrational response between subsequent timesteps is not coupled (Figure 8). The significant effect of this approach on predicted sound power is shown in Figure 9. The figure shows that the coupling of shaft dynamic behaviour in the solution decreases the predicted sound power radiated by an automotive gear casing by up to 8 dB(A) at 1050 rev/min, but increases it by up to 10 dB(A) at 1300 rev/min, indicating that the patterns of noise across the transmission\u2019s operating speed range predicted by the coupled and uncoupled calculation methods would be significantly different. The effect of coupling the dynamic shaft response in this part of the simulation is further discussed in the later section dealing with validation of the whole method. CASING NOISE AND VIBRATION PREDICTION \u2013 The process used in the prediction of casing noise and vibration is shown in Figure 10. A restrained modal finite element analysis of the transmission casing is carried out in order to calculate the natural modes of the casing on its mounts. These data together with the mass, inertia and stiffness matrices from the model are then used to calculate the casing forced vibrational response using a modal contributions method. The resultant vibration data are then used as input for further calculations, to predict sound power using the Rayleigh method, and to predict sound pressure using the Rayleigh and Boundary element methods in LMS-SYSNOISE [Reference 8]."
+        ]
+    },
+    {
+        "image_filename": "designv11_60_0000673_s11740-016-0696-1-Figure1-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0000673_s11740-016-0696-1-Figure1-1.png",
+        "caption": "Fig. 1 Illustration of the hydrodynamic conventions and the MHP structure. Legend: U1 = velocity of the sheet metal strip in sliding direction (mm/s), W1 = velocity of the sheet metal strip transversal to the sliding direction (mm/s), U2;W2 = correspondent velocities of the tool (mm/s), h0 = fluid film thickness (mm), hp = MHP pocket depth (mm), 2rp = pocket length (mm), 2rp2 = pocket width (mm)",
+        "texts": [
+            " Thus, Wang et al. proposed a mass-conserving cavitation algorithm [36] and so did Xiong and Wang for steady-state lubrication [37]. Gherca et al. provided a thorough investigation about the texture geometry in parallel bearings [38]. The analysis of the examined lubricant pockets shows a high potential of pocket geometries to maximize the load capacity and the fluid mechanical properties. Due to the superimposed machine feed and MHP head movement, the geometry of MHP structures is elliptic, see Fig. 1. As of the state-of-the-art, elliptic lubricant pockets as manufactured by MHP have not yet been analytically covered by any research and a solution for the RE is unknown. Therefore, elliptic texture geometry will be investigated in this work. Moreover, most researchers base their work on numerical approaches to model sophisticated state-of-the-art cavitation, inertia, and mass-conserving effects. However, this work begins with an analytical solution of the Reynolds equation, thus, mass-conserving effects have to be neglected, but will be covered in the future. The validity of this approach will be discussed in the paper. Objective The objective of this contribution is to analytically derive a solution of the Reynolds equation to describe the longitudinal fluid pressure distribution, load capacity, and coefficient of friction of a MHPed surface structure. This implies the possibility to enable an efficient MHP process design achieving defined friction coefficients and fluid film thicknesses in deep drawing. A MHP structure is shown in Fig. 1. Note, that the MHP structure is a hollow lubricant pocket. For graphical reasons it is shown upside down. To achieve the objective, the surface structure geometry is simplified by an elliptical shape. Thus, this work neglects lateral effects. However, this work can still be used to determine the minimum fluid pressure occurring, as the presented approach calculates a lower bound and lateral effects will increase the fluid pressure. Approach The structure of this work is as follows. In Sect. 2 the pressure formula obtained from previous work is presented"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0002887_bf03546242-Figure1-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0002887_bf03546242-Figure1-1.png",
+        "caption": "FIG. 1. Osculating Orbit Plane.",
+        "texts": [
+            " Rather general equations of motion for the satellites are derived considering one of the Dynamics and Orbit Determination of Tethered Satellite Systems 179 satellites to be perturbing the other's motion. These equations are simplified by assuming planar motion. The \"apparent\" gravitational constant [11, 12] is derived for a nonlibrating two-body TSS. Then, a two-stage process for identifying and determining the motion of two-body TSS is applied to obtain some interesting example results. A two-body TSS is depicted in Fig. 1. We neglect the mass of the tether in this paper to emphasize the motions of the two satellites. In doing so, we note that the mass of the tether may in some cases be substantial compared to that of the satellites. These cases will need to be considered elsewhere. Basically, our system looks like a classical perturbed two-body system. We assume, however, that the Earth, E, is much more massive than either satellite. That is, m(f) ~ m or m p , where m and m p are the masses of the tethered body of principal interest and the perturbing body, respectively",
+            " Similarly, the equation of motion for the \"perturbing\" satellite is (2) where r p is the magnitude of r p and a p is the acceleration due to forces not modeled by f T and the dominant two-body gravitational term. If we denote by p, the vector from m to m p , then by using equations (1) and (2) we may write 00 JL + JL M fp=--r -r--- T+a -a, r3 p r3 mm p p p (3) where M = m + mr: Because the motion of m may be considered perturbed two-body motion, we will derive explicit equations for m in an osculating orbit. The plane of the orbit is defined by the angles nand i, the longitude of the ascending node and inclination, respectively (see Fig. 1). The angular velocity A of the E~TJ~ coordinate system is therefore such that the velocity of m lies in the ~TJ -plane. Hence, (4) That is, A2 = O. It then follow that the velocity of m may be expressed as (5) The time rate of change of the velocity of m is (6) By using Fig. 1, we may express the angular velocity components Aj, j = 1,3, in terms of the time rates of change of n, i, and (). The results are A 0 0 () diAI = -1\u00a3 SIn i sIn + - cos () dt ~ \u00a3'\\ 0 () di. 0 /\\2 = -1\u00a3 sm i cos - - sin () = dt and . . A3 = n cos i + () n = AI sin ()/ sin i and di dt = Alcos (J (7) (8) (9) (10) (11) Dynamics and Orbit Determination of Tethered Satellite Systems 181 We can use equations (2), (5), (6), (9), (10), and (11) to get the following six equations for the motion of m: VI = vilr - JLlr2 + UI \u00b7 fTlm + UI \u00b7 a (12) V2 = -vlv2lr + U2 "
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0003547_0954406001523074-Figure1-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0003547_0954406001523074-Figure1-1.png",
+        "caption": "Fig. 1 Variables and geometric constructions used in de\u00aening base circles and involutes on circular gears",
+        "texts": [
+            " For circular gears it would have been possible just as e ectively to \u00aex the direction of the line of action by de\u00aening the pressure angle to be the angle between the line of action and the line of centres. For circular gears this angle, which is complementary to PAT, will be referred to here as the pressure angle with respect to centres (PAC). If a point on the cord or line of action is traced as the gears rotate, on two rotating coordinate systems each \u00aexed in one of the gears, a pair of mating involutes are generated (Fig. 1b). These involutes, or gear pro\u00aeles, are parallel to each other and perpendicular to the conjugate line at their point of contact, as that point moves on the conjugate line from one of its ends to the other. Many popular texts, for example references [12] and [13], examine this mechanism in detail and show from several perspectives that involutes have the desirable geometric properties mentioned here. The line of action for a circular gear is tangent to the base circles at its ends. The ends of this line also meet the radii from each gear centre perpendicularly",
+            " Under these circumstances, these terms will not meet all the constraints associated with their use with circular gears. Four combinations of constraints will be settled upon, although others could have been selected. The resulting loci will be given the analogous names that are applied with circular gears, with appropriate quali\u00aecations where the context does not make their special nature clear. Figure 2 shows four identical pairs of \u00aerst-order ellipses rolling on each other. The LAs are arrived at by four di erent means which, in turn, generate four different base outlines. In Fig. 1, for circular gears, neither the angle between the line of action and the tangent at the pitch point (PAT) nor that between the line of action and the line of centres (PAC) changes as the gears rotate. In non-circular gears that is not the case, and only one of these angles can be \u00aexed. It was decided to \u00aex the \u00aerst and then the second angle, PAT or PAC, in turn and to examine the outcome. It was also necessary to decide how to specify the condition for the termination of the line of action in order to arrive at the base outline",
+            " In this method a known outline was used to generate numerically its `rolling mate'. This method was developed in order to cope with a broad range of outlines. In the \u00aegures shown here, the initial pitch outline is drawn on the left and is called the driver or driving pitch outline. The outline that rolls on it and that was arrived at by iteration is drawn on the right and is termed the driven pro\u00aele. Sheppard [2] and others have de\u00aened the conditions that have to be met by any rolling pair capable of continuous rotation. With reference to Fig. 1, these conditions may be said to be as follows: 1. The point of contact, or pitch point, between the pitch outlines must be on the line of centres. 2. The two pitch outlines must be tangent to each other at the pitch point. 3. The outlines must roll without slipping. 4. A complete number of cycles must exist on each outline. For a given driver pitch outline, the iterative method used here to arrive at the driven pro\u00aele begins with an estimated centre of rotation for the driven gear. The pitch pro\u00aele of the driver is then rotated through one cycle, during which the outline of the driven pro\u00aele is generated while meeting the above four constraints"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0001235_b978-0-08-099425-3.00005-4-Figure5.3-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0001235_b978-0-08-099425-3.00005-4-Figure5.3-1.png",
+        "caption": "FIGURE 5.3",
+        "texts": [
+            " The Z-axis is perpendicular to the road plane with a positive direction assumed to be acting downwards. The Y-axis is in the road plane and its direction dictated by the use of a right-handed orthogonal axis frame. The angles a and g represent the slip angle and camber angle respectively. The SAE frame will be used throughout this text unless stated. It should be noted that not all practitioners adhere rigidly to this frame in their publications and another frame, the ISO 8855 tyre frame, is gaining favour. This frame, ISO 8855 tyre axis system. shown simplified in Figure 5.3 has become a de facto standard for new tyre-model architectures. The ISO 8855 tyre axis system is one of three axis systems described in the Tyre Data Exchange protocol, TYDEX, with the accompanying note that while the ISO definition presumes a horizontal ground plane e that is to say normal to the gravity vectore TYDEX does not enforce this condition. It will be seen later that distortions in the tyre carcass will cause the contact patch to move away from the rigid wheel plane shown in Figures 5"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0002377_elan.1140040207-Figure15-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0002377_elan.1140040207-Figure15-1.png",
+        "caption": "FIGURE 15. Mass transport corrected Tafel analysis of the current-voltage curves shown in Figure 13.",
+        "texts": [
+            " Computed voltammograms for this case, using the same parameters as in the EC' mechanism, are shown in Figure 13 from which it can again be seen that a marked distortion of shapevis-his the classical EC' process can arise for certain values of a*())* and the relevant (Leveque limit) normalized rate constant: K* = IZ(\\+ ~~,,,,i~*{4h4xx,2d2/9V,'o)'!~ (29) Figure 14 shows the limiting current-flow rate data for three different flow rates. Again, the Leveque conditions are not quite realized, so that remarks analogous to those made for the EC'( pre-eqm) mechanism apply with respect to the experimental mechanistic analysis requiring the computation of the current-voltage curve for the pertinent electrode geometry and other parameters employed. Figure 15 shows the Tafel plots corresponding to the current-voltage curves of Figure 13. Similarities to the preequilibrium mechanism are evident. Both mechanisms involve the product of the electron transfer,&, in a reversible reaction that regenerates A. The other product, C or intermediate, in each case, is involved in a further irreversible reaction. Thus, the trends discernible from Figure 15 can be explained in terms corresponding to those used in the discussion of Figures 11 and 12. Again a transition between negative and positive trends in halfwave potential is predicted as the normalized rate constant R is increased. This is shown in Figure 3 6. For low-rate constants, the equilibrium (vii) inhibits the electron transfer (vi) producing a cathodic whereas for fast-rate constants the reaction takes on the broad characteristics of an EC' mechanism (finite Y) . However, the second-order nature of the following kinetics induces more curvature in the Tdfel plot than is observed for the classical EC' mechanism. Figure 15 shows that Tafel slopes as high as 109 mV/decade (at 298 K) can be observed for extreme combinations of a*&* and P. Subcase (21: k(z,ilzjY] >> k(-,,,jA]. Figures 17 and 18 show the calculated current-voltage curves and corresponding Tafel plots for this case, again using the parameters and geometry cited for the EC' mechanism. Calculations showed the results for this mechanism to lie, within experimental resolution. inside the Leveque limit and, K I\" 0.6. 0 . 8 , I .o, 0.2, 0 . 4 , ( c l s a a FIGURE 12"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0002678_1.2829318-Figure10-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0002678_1.2829318-Figure10-1.png",
+        "caption": "Fig. 10 (a) Sector of surface on flank 2; (b) domain 7* corresponding to 0 = 0; (c) domain T as function of 0",
+        "texts": [
+            " Finally domain T, whose image is the surface sector S, has the frontier FT composed of the curves u = 2ir, 'd = Q, d = 2{y \u2014 P) and H(M, \u00ab9, <\u0302 ) and is the domain on which to calculate the integrals in formulas ( 5 ) - ( 1 0 ) . H ( M , 1?, i/j): < u = 'd-ip + l3-y + 7r/2 for Q^ip^y-I3 + -Kl2 u = 'd-ip + p - y + 57r/2 for y - /3 + 7r/2 < (/3 < TT (20) Journal of Mechanical Design DECEMBER 1998, Vol. 120 / 585 Downloaded From: http://mechanicaldesign.asmedigitalcollection.asme.org/ on 09/02/2013 Terms of Use: http://asme.org/terms If we consider the values of ^p also, it is possible to obtain the domain in the space W' for the three parameters u, \u2022& and (/3 as shown in Fig. 10(c). The reconstruction of the domain from its frontier FT is allowed since its boundary (38 corre sponds to the frontier FT of T, since S is a regular simple surface (Amerio, 1987; Citrini, 1992). Now we can consider the calculation of the scalar components of force and moment by using formulas ( 5 ) - ( 1 0 ) , since both the integration domain T and the integrating functions are known. The respective ex pressions are: conformation of the flank considered (see Fig. 10). Generally this result is not the same for the flanks which have arcs of trochoids as base instead of circumference arcs as base. A gen eral result can be obtained by considering the force components given by (21) - (23), which are a linear function of the angular pitch a of the screw and are evidently a linear function of the pressure p. This consideration also applies to the other flanks considered, due to the structure of formulas ( 5 ) - ( 7 ) and to the analytical models of the flanks such as (15)"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0002006_s0003-2670(98)00705-3-Figure1-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0002006_s0003-2670(98)00705-3-Figure1-1.png",
+        "caption": "Fig. 1. The scheme of the proposed flow-injection analytical system: S/ST, sample/standard: C, carrier; PPP1, PPP2, computer controllable peristaltic pumps; T, mixing point; CF, chemifold; spectrophotometer working at 210 nm with 18 ml flow-through cell (Hellma); PC, personal computer with 12 bits D/A converter.",
+        "texts": [
+            " After the predetermined sampling time has elapsed, the carrier probe has moved back (the pump was again off) to the carrier container. Improved time-based injection systems were later described by Ruzicka et al. [1] and Fogg et al. [14]. In the present work, a new sample introducing system for FIA and continuous \u00afow analysers is proposed. Because of its construction, it is rather robust and \u00afexible, enabling sample injection, sample dilution and quanti\u00aecation based on an on-line created calibration curve or by standard addition method. The proposed \u00afow-injection analytical system (Fig. 1) was composed of the following elements: computer controllable peristaltic pumps (GILSON, MINIPULS3, France), selection valve (replumbed electrically driven injection valve, installed in ASIA-Ismatec, Switzerland), Te\u00afon tubings and connectors, 6 channel and 12 bit resolution D/A converter (CIO-DDA06/Jr, Computer Boards, USA), homemade software, PC programmed in GWBASIC and diode-array spectrophotometer (Hewlett-Packard, 8453, Germany) collecting data at 210 nm. MINIPULS3 peristaltic pump can be controlled manually or remotely by applying an external voltage",
+            "03 rpm, which results in practically linear (step-free) change of the rotation speed of individual pump. The possibility of external control of the rotation speed of the pumps was exploited for the injection of samples having predetermined exactly known analyte concentration pro\u00aele (linear or nonlinear). In all our experiments, the maximal rotation speed of individual pump was preset at 8 rpm, enabling maximal \u00afow rate in the individual channel of 0.8 ml/min using Tygon tubing with 0.55 mm i.d. With the con\u00aeguration schematically shown in Fig. 1, the exchange of the samples was rather complicated and time consuming. The problem was solved by introducing an appropriate commutator whose scheme is shown in Fig. 2. The commutator was composed of an electrically driven low-pressure injection valve installed in ASIA analyser (Ismatec, Switzerland). The injector was modi\u00aeed in such a way that its rotor was replaced by the rotor usually installed on the low-pressure six-port selection valve of the same type, and two neighbour holes on the Te\u00afon stator were connected through an engraved bore",
+            "s.d. determined from the peak areas. This means that the dispersion of the sample in the chemifold after injection point is completely reproducible. Therefore, the main reason for the differences in peak heights lies in the nonideal reproducibility of the injected sample volume rather than in the nonreproducible sample dispersion. In the similar way the in\u00afuence of an individual part of the chemifold used (dilution chamber, separation unit, etc.) can be measured very rapidly. It is obvious from Fig. 1 that because of the structure of the pro- posed FIA injection/dilution system, any change of a particular part of the chemifold between the injection point and detector should change sample dispersion, and consequently, peak shape in the same way as in the case when sample is injected using a classical injection valve. In the same analogy, peak carryover should depend on the sample injection frequency as is the case when sample is injected by a rotary or slider valve. Few additional bene\u00aets of the proposed FIA injection dilution system can be observed in Fig. 1. Optimisation of the used FIA system based on the determination of optimal sample volume (S1/2) can be determined in the easiest possible way only by changing the injection time, avoiding the time consuming change of the sample loops. Because of the possibility to produce a steady-state signal, using an FIA system based on the proposed injection/dilution system can be readily converted to a continuous \u00afow system. This option could be of some advantage when samples containing low concentrations of an analyte have to be analysed",
+            " Besides simple sample injection (as classical slider or rotary injectors), the proposed system also allows on-line sample dilution. This can be achieved simply by changing the operation procedure of the pumps propelling the sample and the carrier during the injection period. An appropriate dilution can be obtained if sample and carrier propelling pumps are not switched on and off, respectively, during the injection time, but the rotation speed of both the pumps are modi\u00aeed (increased and decreased) in such a way that the total \u00afow rate tot at the merging point T (Fig. 1) remains permanently constant. For example, if the rotation speed of sample and carrier propelling pumps will be increased for 50% and decreased for 50%, respectively, two times diluted sample is injected. An example of on-line sample dilution, actually the creation of a calibration curve, is presented in Fig. 5. As a sample a standard solution containing 5 mg/l of NO\u00ff3 was used. During the injection period, rotation speed of the pump propelling the sample and the rotation speed propelling the carrier was increased or decreased for 10%. All measurements with the exception of the signal obtained for undiluted sample were made in duplicate. The calibration curve created on the basis of the measurements presented in Fig. 5 was a straight linear curve having r2 of 0.9998. In the next series of the experiments the very unique possibility of the proposed injection/dilution system, i.e. the injection of the samples having exactly known concentration gradient at the merging point T (Fig. 1), was tested. Speci\u00aec gradients can be obtained by the appropriate change of the rotation speed of individual pumps. The only condition is again a permanently constant \u00afow rate tot at merging point T. In the case when both peristaltic pumps deliver sample and the carrier at the same \u00afow rate (the same predetermined maximal rotation speed, the same diameter of the applied Tygon tubings), constant \u00afow rate at merging point T can be achieved by controlling external voltage applied to the individual pump"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0002229_jsvi.1996.0017-Figure17-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0002229_jsvi.1996.0017-Figure17-1.png",
+        "caption": "Figure 17. Poincare\u0301 map for b=bc=0\u00b78842 with a=ac=0\u00b77338, d=0\u00b71, v=4, G=20.",
+        "texts": [
+            " A detailed analysis of the homoclinic-heteroclinic manifold intersections in the b-section in the parameter space for the critical value of a at v=1, G=1 has been discussed extensively in the previous section with particular interest in parameter domains in which the two families do not intersect. As an example of the robust non-interacting nature of the heteroclinic and homoclinic manifolds, as predicted in the critical parameter case, consider the bifurcation diagram shown in Figure 16, and the Poincare\u0301 map of Figure 17. Here the forcing amplitude has been increased to G=20, and the forcing frequency to v=4. When we consider the relations (15) and (16) with v=4, we find that G=20 is considerably less than both maximum permissible homoclinic and heteroclinic upper bounds. In fact they predict maximum values of the orders of Go110 000 and Ge1150, and clearly, as we would perhaps expect, these values are far in excess of the actual numerical values, which may be explained as a breakdown of the Melnikov method for such extreme perturbing parameters"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0001235_b978-0-08-099425-3.00005-4-Figure5.57-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0001235_b978-0-08-099425-3.00005-4-Figure5.57-1.png",
+        "caption": "FIGURE 5.57",
+        "texts": [
+            " For vehicle handling studies we are generally concerned with the manoeuvring of the vehicle on a flat road surface and elaborate contact formulations such as these are not necessary and a single point of contact is used to calculate aggregate slip states in the contact patch. This is the so-called \u2018point follower\u2019 formulation. Whichever type of contact formulation is used, the function of the tyre model is to establish the forces and moments occurring at the tyre to road contact patch and resolve these to the wheel centre and hence into the vehicle as indicated in Figure 5.57. For each tyre the tyre model will calculate the three orthogonal forces and the three orthogonal moments that result from the conditions arising at the tyre to road surface contact patch. These forces and moments are applied at each wheel centre and control the motion of the vehicle. In terms of modelling the vehicle is actually \u2018floating\u2019 along under the action of these forces at each corner. For a handling model the forces and moment at the tyre to road contact patch, which are usually calculated by the tyre model, are: 1"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0002796_l90-034-Figure3-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0002796_l90-034-Figure3-1.png",
+        "caption": "FIG. 3. Nicolai's problem.",
+        "texts": [
+            " Writing at the height of the British Empire, when Britain's steel mills and shipyards were flourishing, he analysed the stability of shafts under the action of compressive forces and axial torques. Figure 2 shows five possible end conditions for the shaft. In each case the load is applied in the vertical direction, and this is the direction of the torque vector. Greenhill considered cases I, 11, and V. Using the equilibrium method, Greenhill found the values 2.861, 2, and 2 respectively for the critical torque parameter when P = 0. Greenhill did not consider cases I11 and IV. The next step came when Nicolai (1928) considered a variant of case IV (Fig. 3), in which the twisting moment, M, is tangential (i.e., the torque vector is directed along the tangent to the deformed axis of the bar). He found that there is no equilibrium configuration of the bar other than the rectilinear one. This observation leads to two possible conclusions: either the bar is stable for all possible values of the moment, i.e., the critical torque parameter is k = a, or the equilibrium method, which seeks values of the moment for which there are nontrivial equilibrium configurations, is invalid"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0003736_aim.2001.936501-Figure1-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0003736_aim.2001.936501-Figure1-1.png",
+        "caption": "Fig. 1 . Geometrical situation of a minimum infinity-norm solution",
+        "texts": [
+            " This algorithm is based on the geometrical relationship between the hypercube which represents the same magnitude of the infinity-norm of x and the solution space of A z = y in n-dimensional space. In the geometrical viewpoint, the solution is determined when the boundary of the hypercube first touches the solution space by increasing volume of the hypercube. It means that in m-dimensional space, the end point of the vector y must lie on a boundary plane of the convex polyhedron mapped from the contacting point. Figure 1 shows the geometric situation of above explanation. More structured procedures of this algorithm are presented in the following. S-Step 1. convex polyhedron by using Let k = 1 and select an arbitrary vertex p p ) of the where z ( l ) is a vector representing an arbitrary vertex on the unit n-box in Rn. S-Step 2 . Determine the points p i for i = 1,. . . , n. which can be connected to the point p r ) as p i = Aqi, i = 1,. . . r * (11) qi = E , z ( ~ ) (12) (13) Ei = diag(1,. . . , 1, -1,1,. . . ,1} S-Step 3"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0000183_1754337114568825-Figure2-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0000183_1754337114568825-Figure2-1.png",
+        "caption": "Figure 2. Camera setup for the serve shot.",
+        "texts": [
+            " The shuttlecock moment of inertia in the pitch direction measured with the bifilar swing method had a mean of 2.933 1026 kgm2 and standard deviation of 7.423 1028 kgm2 after five runs. Three synchronized Phantom Miro 120S highspeed cameras were set up to observe the flight path of the shuttlecock at 500 frames per second (fps). The cameras were adjusted for optimal framing of the flight path. All four modes of play were studied \u2013 clear, net, smash and serve. An example of a two-camera setup for the serve shot is presented in Figure 2. The videos were then processed for the initial flight conditions and the resultant flight paths of the four modes of play. The flight path of the serve shot was applied as input data for parameter identification using the nonlinear grey-box model (idnlgrey) in MATLAB . Previous attempts of drag and lift measurements using only raw acceleration data obtained from position plots showed high inaccuracy due to sensitivity to experimental error. The usage of the greybox model for identification eliminates this issue"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0000878_978-3-319-47614-8-Figure6.4-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0000878_978-3-319-47614-8-Figure6.4-1.png",
+        "caption": "Fig. 6.4 Equilibrium diagram of the beam element",
+        "texts": [
+            " Let us consider a long slender structure, we called this beam. Instead of getting into the detailed deformation field at a point, let us take a global element dx of the beam at distance x and consider its equilibrium. We begin with the existence of bending moment Mx and shear force Vx at distance x to keep the element in equilibrium against the applied load q. Note the convention, on the right face the shear force and the applied load are positive downwards and bending moment is positive in counter clockwise direction. Figure 6.4 gives the equilibrium diagram of the beam element. The moment and force equilibrium equations are bS LE Equivalent length of a Cantilever column 6.2 Beam with Axial Load (Beam-Column) 161 Mx \u00fe Mx \u00fe dMx\u00f0 \u00de Vxdx\u00fe qdx dx 2 \u00bc 0 Vx Vx \u00fe dVx\u00f0 \u00de qdx \u00bc 0 \u00f06:18\u00de Neglecting second order terms and rearranging, we get dMx dx \u00bc Vx dVx dx \u00bc q \u00f06:19\u00de We notice that the bending moment and shear force are related between themselves, while the shear force is directly connected with the loading. Now consider the beam geometry after deformation from pure bending, see Fig"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0003913_s1474-6670(17)54642-x-Figure3-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0003913_s1474-6670(17)54642-x-Figure3-1.png",
+        "caption": "Fig. 3: Elements for disturbed grasp analysis",
+        "texts": [
+            "2) cos(a2)*sin(alo-a30) N2= N3*'------------------ cos(a3)*sin(a20-alo) and the consequent Ti Having once fixed three contact points, it is necessary, in order to determine the normal actions, to assign the values to three degrees of indetermination: two due to the choice made for H (a,b) c A and one for N3. In the following part of the paper it will be shown that the solutions obtained from (3.2) are not equivalent if evaluated with respect to their behaviour when a generalized disturbance is ap plied. DISTURBED GRASP Each solution determined by the previous analysis of autoequilibrated grasp needs to be evaluated; for every solution we suppose that point H and the N3 value are assigned . When a generalized disturbance is applied (see Fig.3), the consequent values of Tid can be com puted by three equilibrium conditions where Ni are known from the autoequilibrated grasp analysis. Searching the maximum value Q* of the generalized disturbance applicable to the object before rolling or slipping at the contacts means solving (4.1) for increasing values of Q and for ~ach angle 6 and verifing (4.2) Q* corresponds, then, to the first value of Q in which, for at least a 6, one of the Tid makes (4.2) false. Q* can be found for every autoequilibrated solution characterized by Hand N3; given Pi, and Pq, if H varies within A (a and b assume all the possible values compatible with the friction conditions), a set of corresponding Q* is determined"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0002678_1.2829318-Figure4-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0002678_1.2829318-Figure4-1.png",
+        "caption": "Fig. 4 Helical lines and flanl<s",
+        "texts": [
+            " It is also important to choose a reference value for the rotor rotation angle ip (see Fig. 2) and we choose the initial value in Fig. 3 as zero. Notice also that with the adopted convention, the rotation of the central screw is counter-clockwise and that of the satellite screws is clockwise. Journal of Mechanical Design Copyright \u00a9 1998 by ASME DECEMBER 1998, Vol. 120 / 581 Downloaded From: http://mechanicaldesign.asmedigitalcollection.asme.org/ on 09/02/2013 Terms of Use: http://asme.org/terms Fig. 1 Screw pump rotors The profiles are composed of different arcs of curves (see Fig. 4). These are in particular arcs of circumference, cycloid or epitrochoid, whose analytical expression on the xy plane is known in parametric form. For the definition of the seals, it is important to define the points that represent the profile vertexes (see Fig. 4). Starting from the profiles, the screw can be modelled in space (Mimmi and Pennacchi, 1997). From a geometric point of view, the latter are helicoids composed of helical surfaces. The model obtained is given by equations in parametric form. Each arc is developed in space with a screw motion and forms one of the flanks of the screw (see Fig. 4). On the contrary, we can consider that the base of each flank is the corresponding arc. The vertexes of the profiles of the central screw define the helical lines. The part of the cylindrical external side of the central screw that comes into contact with the satellite screws, and the segments of hehcal line, that are also common to the satellite screws determine the seal curves on the satellite screws (see Fig. 8). As well, the seal curves on the central screw can be determined by considering the vertexes of the satellite screws"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0001256_jahs.60.042007-Figure2-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0001256_jahs.60.042007-Figure2-1.png",
+        "caption": "Fig. 2. Face-gear drive: inner and outer limiting ends of a face-gear tooth.",
+        "texts": [
+            " Since the system is periodically time varying, Floquet theory is employed to solve a Mathieu\u2013Hill type equation to evaluate the system stability. Also, a perturbation analysis based on Hsu\u2019s method (Ref. 28) is performed to obtain approximate closed-form expressions for the stability boundaries for design purposes. Face-Gear Geometry and Meshing Kinematics In this paper, the face gear and input pinion are cut for line contact which occurs when the pinion gear and the shaper have identical profiles. The cross section of a sample face-gear drive is shown in Fig. 2. According to the mathematical model developed by Litvin et al. in Refs. 1 and 29, contact lines, which generate the face-gear tooth surface, are derived from the so-called equation of meshing\u23a7\u23aa\u23aa\u23a8 \u23aa\u23aa\u23a9 f (us, \u03b8s, \u03c6s) = ns \u2022 V(s2) s = 0 V(s2) s = [ v(s2) xs v(s2) ys v(s2) zs ] T r2(us, \u03b8s, \u03c6s) = M2s(\u03c6s)rs(us, \u03b8s) (1) where ns is the unit normal vector to the shaper tooth surface and V(s2) s is the relative sliding velocity between the shaper and face gear at a point of contact in shaper coordinates. Furthermore, M2s is the rotation matrix, 042007-2 which transforms shaper coordinates into face-gear coordinates, rs and r2 are position vectors in the shaper and face-gear coordinates, respectively, (us ,\u03b8s) are the Gaussian coordinates of the involute shaper surface, and \u03c6s is the generalized parameter of motion. To avoid undercutting, the position vector for the inner limiting end (section A in Fig. 2) of facegear tooth is solved via the following equations (Ref. 29):\u23a7\u23aa\u23a8 \u23aa\u23a9 rs = rs(us, \u03b8s) = [ xs ys zs ]T f (us, \u03b8s, \u03c6s) = 0 2 1 + 2 2 + 2 3 = 0 (2) with 1 = \u2223\u2223\u2223\u2223\u2223\u2223\u2223\u2223\u2223\u2223\u2223\u2223\u2223\u2223 \u2202xs \u2202us \u2202xs \u2202\u03b8s v(s2) xs \u2202ys \u2202us \u2202ys \u2202\u03b8s v(s2) ys \u2202f \u2202us \u2202f \u2202\u03b8s \u2202f \u2202\u03c6s d\u03c6s dt \u2223\u2223\u2223\u2223\u2223\u2223\u2223\u2223\u2223\u2223\u2223\u2223\u2223\u2223 = 0 (3a) 2 = \u2223\u2223\u2223\u2223\u2223\u2223\u2223\u2223\u2223\u2223\u2223\u2223\u2223\u2223 \u2202xs \u2202us \u2202xs \u2202\u03b8s v(s2) xs \u2202zs \u2202us \u2202zs \u2202\u03b8s v(s2) zs \u2202f \u2202us \u2202f \u2202\u03b8s \u2202f \u2202\u03c6s d\u03c6s dt \u2223\u2223\u2223\u2223\u2223\u2223\u2223\u2223\u2223\u2223\u2223\u2223\u2223\u2223 = 0 (3b) 3 = \u2223\u2223\u2223\u2223\u2223\u2223\u2223\u2223\u2223\u2223\u2223\u2223\u2223\u2223 \u2202ys \u2202us \u2202ys \u2202\u03b8s v(s2) ys \u2202zs \u2202us \u2202zs \u2202\u03b8s v(s2) zs \u2202f \u2202us \u2202f \u2202\u03b8s \u2202f \u2202\u03c6s d\u03c6s dt \u2223\u2223\u2223\u2223\u2223\u2223\u2223\u2223\u2223\u2223\u2223\u2223\u2223\u2223 = 0 (3c) Furthermore, the inner limiting value of the face-gear tooth, L1, is obtained as (Ref. 29) L1 = z\u2217 s \u2212 rds tan \u03b3 and \u03b3 \u2264 90\u25e6 (4) Here z\u2217 s represents the limiting value along the z axis of the shaper surface (zs axis in Fig. 2), solved by the nonundercutting condition described in Eqs. (2) and (3), rds is the shaper dedendum radius, and \u03b3 is the intersecting shaft angle of the face-gear drive. The outer limiting end (section B in Fig. 2) is determined by the condition of the face-gear tooth top land becoming pointed, and the outer limiting value, L2, is given by (Ref. 29) L2 = Ns 2Pd ( 1 tan \u03b3s \u2212 1 tan \u03b3 ) + ag tan \u03b3 + Ns 2Pd tan \u03b3s ( cos \u03b1PA \u2212 cos \u03b1 cos \u03b1 ) (5) where Ns is number of shaper teeth, Pd is diametral pitch, \u03b3s is apex semiangle of shaper pitch cone, ag is tooth addendum, \u03b1PA is pressure angle, and \u03b1 is solved by (Ref. 29) \u03b1 \u2212 (Ns \u2212 2) sin \u03b1 Ns cos \u03b1PA = \u03c0 2Ns \u2212 (tan \u03b1PA \u2212 \u03b1PA) (6) The corresponding inner limiting radius, R1, and outer limiting radius, R2, for the face-gear tooth are scaled in the plane of the face-gear body disk, respectively, as R1 = L1 sin \u03b3, R2 = L2 sin \u03b3 (7) Following the work in Ref"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0003736_aim.2001.936501-Figure5-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0003736_aim.2001.936501-Figure5-1.png",
+        "caption": "Fig. 5. The four-link planar robot executes a linear trajectory(shown as dashed arrow)",
+        "texts": [],
+        "surrounding_texts": [
+            "Table I compares the average number of floating point operation of the three algorithms. As shown in the table, our algorithm needs a little more floating point operation than Cadzow\u2019s algorithm does, but much less than Shim\u2019s. Note that determination of a minimum two-norm solution takes almost a half of the operation during execution of proposed algorithm.\nIv. AVOIDING DISCONTINUITY IN TRAJECTORY PLANNING\nIn some cases, uniqueness or continuity of minimum infinity-norm solutions can not be guaranteed. So one may ask when the nonunique and discontinuous solution can be occurred and how one can recognize this situation? And why we have to recognize this? In a continuous-time environment, discontinuity means the solution of a system can $mp at some time. This may also cause control input to a system to jump with large deviation and cause backlashes or jerks. More detail information about uniqueness and continuity of least infinity norm solutions is available in [7].\nA . Ian\u2019s Approach The keywords of this approach are \u2018rate mixing\u2019 and \u2018subspace angle\u2019. Rate mixing approach adopts a linear combination of a minimum two-norm solution and a minimum infinity-norm solution to make a new feasible solution for avoiding discontinuity. So the rate mixing has the following form: .\ne* = re(co) + (1 - o 5 T 5 1 (33) In this case, the subspace angle is used for determination\nof the coefficient r . The subspace angle is defined geometrically as the angle between two hyperplanes(subspaces SI and S2) embedded in a higher dimensional space. Because non-uniqueness can occur when solution space of a system touches more than one points of a hypercube, it is good to measure the angle between solution space and subspace of the hyperplane which include solution point. Note that the solution space of a system is parallel to the nullspace of the system. So we can replace the solution space with the nullspace in calculating subspace angle. As a matter of fact, Ian\u2019s approach doesn\u2019t calculate actual subspace angle. To find the minimum subspace angle over all possible faces and edges of the hypercube, Ian suggested to concatenate nullspace N and 5\u20182 into n x n matrix and check its determinant. A zero determinant corresponds to a zero subspace angle; larger absolute value determinants indicate larger subspace angles.\nSo the new discontinuity index has the following form:\nAnd the mixing ratio is defined as follows:\n(35) ,,. = 1 - e-a.d,in\nThis approach is well defined, easy to understand and ready to be implemented in computer codes. But the \u201cZero Subspace Angle\u201d condition is not complete(necessary but not sufficient) in deciding discontinuity. So sometimes the mixing factor r drops to zero, even in a case where discontinuity is not imminent. If someone has to know exact situation, this approach can fail to give exact information. So we propose a new approach that can handle this situation in next section.\nB. A New Approach\nFigure 4 shows a geometry of subspace angle if dimension of matrix A of equation (20) is 2 x 3. As shown in figure 4, the resultant dimension of solution space is 1. Ian\u2019s approach measures subspace angles between solution space and every faces of hypercube. But there is no need to measure all of them. In figure 4, the solution space has no intersection with the face F3, so the subspace angle between the solution space and F3 needs not be calculated in this case. Because of the above reason \u201cZero Subspace Angle\u201d condition is not said to be complete. Now there is one question remained. How can one select subspaces which have an intersection with solution space\u2018? In figure 4, we can find an answer, the solution space has intersection with face Fl and F2, but not with face F3. In this case, the first and second component of optimal solution z* has the same magnitude with ~ ~ z * ~ ~ ~ , while the third one does not. During the running of algorithm, one can easily determine indices of saturated components of z*. Indices of columns which compose matrix A2 used in our algorithm or Cadzow\u2019s algorithm are the indices of saturated elements of z*. So one can easily find the faces that have an intersection with the solution space, and determine the singular index using the corresponding faces.\nNow, we know how to set the matrix S2 and N I which are included in Ian\u2019s approach(equation (34)). The index",
+            "of row which is the element of N I must be selected from the indices of saturated elements of optimal solution 2\u2019.\nIt is a time consuming to calculate nullspace of matrix A in Ian\u2019s method. If the m x n matrix A has rank m, nullspace of matrix A can be uniquely specified. So we can replace N in equation (34) with A, and get the same result. But the indices of row which is the element of NI are not the same as that of Ian\u2019s method. Matrix NI in this case must include the matrix AI\u2019. Each column of A1 is the column of A whose index correspond to that of non-saturated element in optimal solution 2*.\nOne can easily find the matrix AI in the last of procedure of finding optimal solution. So distance from discontinuity can be determined in the procedure of finding optimal solution with a few additional calculations.\nBut this singular index has intrinsic problem of discontinuity. Though the contact point may be changed continuously, the hyperplane which includes the contact point can hop from one hyperplane to another. So we apply moving average method to smoothen the changes of the resultant singular indices.\nC. Example\nFig. 7. Mixing ratio using proposed method applied to the same example\nTo show the versatility of the proposed method, we derive a joint trajectory by solving inverse kinematic equation of a four-link planar redundant manipulator.\nThe trajectory planning based on Ian\u2019s method is shown in figure 6.\nAs shown in figure 6, there is no discontinuity of each synthesized joint angle. But singular index around 300 in time axis fails to alarm discontinuity of pure I,-norm solution. Singular index around this point falls to zero thought the discontinuity of pure minimum infinity-norm solution doesn\u2019t occur near that point.\nFigure 7 shows the result of proposed method to the same example. In this case, singular index doesn\u2019t fall to zero near 300. So, the proposed method can detect more accurate distance from the singular situation.\nV. CONCLUSION\nIn this paper, we propose a new algorithm for finding minimum infinity-norm solution. The proposed new alge rithm doesn\u2019t need any condition while Cadzow\u2019s algorithm needs satisfaction of Haar condition. And this algorithm needs less computational time than Shim\u2019s algorithm does.\nWe also described how to determine the distance from discontinuity in solution more exactly than preceding research work done by Ian. The proposed approach does not need any separate procedure in determining nonuniqueness of solutions. And we proved that the proposed method can be successfully used through numerical example.\nThe proposed method can be used for optimal trajectory planning for a redundant manipulator and monitoring the singularity of a manipulator.\nFuture studies may include the efficient method which overcomes the intrinsic discontinuity of singular index based on proposed method.\nREFERENCES Y. Nakamura, Advanced Robotics, Redundancy and Optimization, Addison-Wesley Publishing Company B. Siciliano, \u201cKinematc Control of Redundant Robot Manipulators : a Tutorial,\u201d Journal of Robotic Systems, vol. 6 , no. 3, pp.201-212, 1990 T. Yoshikawa, \u201cManipulability and Redundancy Control of Robotic Mechanisms,\u201d IEEE Int. Conference of Robotics and Automation 1985. pp.1004-1009. James A. Cahow, \u201cA finite algorithm for the minimum 1, solution to a system of consistent linear equations\u201d SIAM Journal of Nunierical Analysis, vol. 10, 110.4, Sept. 1973, pp. 607-617. I.C. Shim and Y.S. Yoon, \u201cStabilized minimum infinity-norm torque solution for redundant manipulators\u201d Roboticcl pp. 193- 205. Jihong Lee, \u201cA structured algorithm for the minimum 1, solutions and its application to robot velocity workspace analysis\u201d to appear in Robotica, 2000. I. Gravagne, \u201cMinimum effort techniques for inverse kinematics of redundant robot manipulators,\u201d Master Thesis, Dept. of Electical Eng., Clemson University, 1999."
+        ]
+    },
+    {
+        "image_filename": "designv11_60_0000936_amm.859.15-Figure2-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0000936_amm.859.15-Figure2-1.png",
+        "caption": "Fig. 2 Features of the new nozzle.",
+        "texts": [
+            " The design of the die angle in our study was based on a previous research analyzing nozzle die angle during the extrusion of PLA material [6]. With regards to heat transfer and heat distribution, the liquefier design in cylindrical shape was better compared to the square shape of original nozzle. Additionally, high insulated material was installed to the liquefier to maintain the heat temperature. All these factors was investigated to determine their impacts on the extrusion process [6\u20139]. The features of the new nozzle are depictured in Fig. 2. Notes: Fig. 2 a: screw slot, b: outlet nozzle, c: heating element slot, d: cylindrical liquefier installed with insulator, e: thermocouple slot Sample Fabrication. Sample was designed using Autodesk Inventor Software (Autodesk, USA) and converted into standard triangular language (STL) format as shown in Fig. 3. The dimension of the part model is 132 mm x 86 mm x 9 mm. The part model has been designed based on previous research [10,11]. The first model was printed using the original nozzle (nozzle 1) and the second model was printed using a newly developed nozzle (nozzle 2)"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0003908_s1474-6670(17)36910-0-Figure7-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0003908_s1474-6670(17)36910-0-Figure7-1.png",
+        "caption": "Fig. 7: con\\'eyor belts. speed u (t)",
+        "texts": [],
+        "surrounding_texts": [
+            "This o\\'er\\'iew has certainly forgotten a lot of int.eresting papers, Howe\\,er, we hope it has pro dded some enlightenment to t he matter. and recalled some useful references among t he huge number of results connected ,\\'ith LTDS. Acknowledgments: In 1993, seofra} twm:> (mainly ill FmncL but also in Europt, Russia, Japan, Nor'th and South America) initiated a ntfll'ork for coopfmtion on dflay systems, whOM lIWVf1l!ent is nOli' sllpported by thf fnnch CNRS (National Cent.er for Scientific Research ), in the fmm.fwork of tllt \"GdR Automatiql/.( ,. (Research Group in Automatic control), TIn author would like to ac knowledge CNRS for this S'Upp01't. as well as all collwg'Ufs related to this group fOT theiT frllif:ful dis cussi 0 lI,s,"
+        ]
+    },
+    {
+        "image_filename": "designv11_60_0000813_tmag.2016.2620182-Figure4-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0000813_tmag.2016.2620182-Figure4-1.png",
+        "caption": "Fig. 4. The air bearing pressure distribution.",
+        "texts": [
+            " This difference is caused by the interaction of the disk with the ramps because the ramp restricts the displacement of the disk. If the ramp is a rigid body, the displacement of the disk never exceeds the initial gap (180 \u03bcm). However, with a flexible ramp, the displacement of the disk exceeds the initial gap, as shown in Fig. 3(c). A transient shock simulation is used to obtain the disk curvature, vertical force, and pitch and roll moments between the slider and disk. The air bearing distribution calculated by the slider dynamic simulation is applied to the down head slider, as shown in Fig. 4. The air bearing pressure distribution is changed at each time step. B. Slider dynamics simulation The air bearing pressure and slider behaviors are calculated in the slider dynamics simulation. Their distributions are obtained by simultaneously solving the generalized Reynolds equation and equation of motion, as shown in Eqs. (1) and (2) [10].       3 3 6 12 p p ph Q ph Q x x y y U ph V ph ph x y t                                (1) where \u03bc is the viscosity of air, p is pressure, U and V are the relative velocities in the x and y directions, respectively, h is the gap between the slider and the disk, and Q is the flow factor"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0002073_s0032-9592(97)00044-7-Figure1-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0002073_s0032-9592(97)00044-7-Figure1-1.png",
+        "caption": "Fig. 1. Encapsulated pF-EIS-CAP sensor.",
+        "texts": [
+            " Introduction Since the investigations of Bergveld [1,2] potentiometric pH-sensitive field effect transistors (FETs) have been used as transducers for biosensors. These sensors are sensitive to the buffer capacity of the sample. This drawback was eliminated by Hintsche et al. [3,4], by the development of a fluoride-sensitive FET (pF-FET) consisting of a Si/SiO2/Si3N4/LaF3 gate. In preceeding papers we reported on the development of an inexpensive and simple EIS-chip consisting of Si/SiO2/Si3N4/LaF3-1ayers, which changes its capacitance depending of the fluoride concentration in the contacting liquid [5] (Fig. 1). This chip was used as a transducer for Bio-pF-EIS-CAP-sensors [6]. In the *To whom correspondence should be addressed. 175 present paper we report on the application of these biosensors, which were integrated into a flow injection analysis (FIA) system, for on-line monitoring of various bioprocesses. On-line analyser system Aseptic samples were taken by a tubular in situ sampling unit, with a hyrophilized polypropylene microfiltration membrane of 0.2/~m pore diameter (ABC, Pucheim), on-line. The samples were preconditioned (e"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0000080_s00170-016-8529-0-Figure4-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0000080_s00170-016-8529-0-Figure4-1.png",
+        "caption": "Fig. 4 Experimental setup: a 20 MN hydroforming press at Harbin Institute of Technology and b experimental die for the investigation of flattening behavior",
+        "texts": [
+            " If the relationship between internal pressure and axial feeding during wrinkling stage is reasonable, the preformed wrinkles could be flattened completely in the calibration stage and a defect-free specimen could be obtained, as shown in Fig. 3. It can be seen from Fig. 3 that the maximum diameter of the hydroformed sample is 85 mm with a length of 63 mm. The semi-conical angle between the transition zone and the expansion zone (maximum diameter zone) is 20\u00b0; thus, a total length of 123.5 mm (l/d\u22482) for the deformation area can be achieved. The experimental investigations into flattening behavior of wrinkles were carried out on the 20MNhydroforming press at Harbin Institute of Technology, as shown in Fig. 4a. The 20 MN press could control two intensifiers and four axial cylinders simultaneously. Moreover, Fig. 4b gives the experimental die used in this investigation, and more information about this experimental setup has been discussed in [17]. Figure 5 shows the preformed wrinkles of 5A02 tubes under different internal pressures of 0.8 ps, 1.0 ps, 1.2 ps, 1.4 ps, 1.6 ps, and 1.8 ps. Clearly, the wrinkles only occur at both ends of the tube\u2019s deformation zone when the internal pressure was lower as 0.8 ps and 1.0 ps, and the two wrinkles are distorted and non-axisymmetric. Moreover, three axisymmetric wrinkles were formed for the cases with higher internal pressure (1"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0000329_j.ifacol.2016.03.133-Figure1-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0000329_j.ifacol.2016.03.133-Figure1-1.png",
+        "caption": "Fig 1. Geometry of vehicle in pitch plane",
+        "texts": [
+            " LAUNCH VEHICLE MODELING The vehicle attitude dynamics is modelled taking into account aerodynamics, control actuator dynamics, vehicle bending, propellant sloshing, variation in centre of gravity (cg) and moment of inertia etc. as the time progresses. This leads to a time variant system. Using time slice approach a short period model is evolved [2] which can be assumed time invariant for a small duration, so that linear time invariant control system principles can be used. Further it is assumed to be decoupled in pitch/yaw/roll and planar analysis is carried out. Referring to [2],[3], the equations can be represented as Consider the geometry of vehicle in pitch plane represented in Fig 1.      000 Um zF U z (1)  yMyyI  (2) where z is perturbation velocity of vehicle,  is pitch angle,  zF is total force acting parallel to vehicle body axis, yyI is moment of inertia about pitch axis, 0m is reduced mass of vehicle and 0 V U is the forward velocity of vehicle. Considering the forces and moments acting on the launch vehicle due to engine inertia, aerodynamics, elasticity, slosh etc, effective force and moment equations may be represented as Force Equation                 .]0[ 0][ 2 0 2 1 i qrm i i rl i Uzrlclrlrm piUpim i lnCAU ii q T TcTzF T l T l T l        (3)                        i q i i c l r l r m r I i c l r l r m i q i U c l r mU c l r m z c l r l r mclr m r IU r l r m c lrlr m r I i piU ip lpim lnCAU ii q T T ii q T TcT c lyM T l T l T l T l T l           )()(00 ][] 2 [ 0 ][0 ][ 2 0 2 1 ][ (4) where A is reference area, nC is normal force coefficient, cl and rl are distance from origin of body axis and c"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0002456_s1474-6670(17)48465-5-Figure7-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0002456_s1474-6670(17)48465-5-Figure7-1.png",
+        "caption": "Figure 7: Pen Examplr Grasp<>d by 4 Fiugrrs.",
+        "texts": [
+            " lIIass I/Iy = 1 S , finger initial distances frolll C(,lltpr of IllaSS alld slidillg sp('('d parameters 11 = 3 cm, 12 = 2 cm, 13 = 2 cm, 14 = 3 cm, iX; = 1 cm sec-I, time intervals to = 0 sec, I1 = 1 sec, t2 = 2 sec, t3 = 3 sec, t4 = 4 sec. Figure 9 shows the reference trajectories for the wrench intensitiE's , whE're the following do not vary with time Cl 0.1 N C2 -0.025 :V C3 -0.025 N C6 0.04 :V C9 -0.04 N ('10 0.1 N Cll -0.025 N Cl2 0.025 N 4.3 Control Algorithm Simulation Let us again assume the skill example of thp pen grasped by 4 fingers (see Fig. 7) and rotate the pen around thr ~-axis. The control goal according to section 3.3 is thrIl to control the pen position to rest rd(t) = [00 O]T [tnd thc RPY-angles to rotate around the z-axis cpd( t) = [sin wt 0 of, where w is the angular velocity of rotation . Figure 10 shows that the initial angular position error decreases to zero asymptotically. In this paper we described an Intelligent Assisting System - IAS as a forthcoming research topic approaching the under standing of human manipulative skill"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0000153_imece2015-50907-Figure7-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0000153_imece2015-50907-Figure7-1.png",
+        "caption": "FIGURE 7. THE SIMULATION ENVIRONMENT FOR PLANE PAINTING.",
+        "texts": [
+            " In practical, we tend to select the point in the center area of the intersection set so that a relatively high margin of the joints can be guaranteed which is proved by the simulation experiment in the next section. We have established a virtual production environment to test our method. It contains a 3D mobile platform with a manipulator fixed on it. The objective painting area is a plane surface whose size is 1980\u00d71750 mm2 and the painting trajectory is marked as the red line with the local frame located at the upper left corner as shown in Fig. 7. Obviously the minimum intersection set of the BWLSs appears when center wrist points are at 4 corner points respectively. The positions of those corner points are C1 = [0,0,0]T C2 = [1980,0,0]T C3 = [1980,1750,0]T C4 = [0,1750,0]T (18) corresponding to the outer boundary surfaces E1, E2, E3, E4, which are shown in Fig. 8. We firstly calculate the min distance between E1, E2, E3, E4 two by two and the results are all zero which means that the four outer boundary surfaces intersect with each other"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0003560_nme.1620180606-Figure9-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0003560_nme.1620180606-Figure9-1.png",
+        "caption": "Figure 9. Cooling tower shell",
+        "texts": [
+            "063 5 +0*124 89 1 LINE NODE AND TRANSITIONAL SHELL ELEMENT Modulus of elasticity = 519,100 Poisson\u2019s ratio = 1/6 Shear factor = 5/6 The results of this analysis (Figure 8) are compared with those obtained using the SHORE-I11 for the analysis of program which is a well-established public domain computer shells of revolution. The results from the SHORE analysis are shown as solid lines while those obtained using the new approach are indicated by small circles. Excellent agreement between the two analyses is obtained. Cooling tower shell-dead load The column supported hyperbolic cooling tower (Figure 9a) is analysed for self-weight loading. The discretization for rotational elements is shown in the figure with dashed lines. The region marked 'Imperfection' in the figure is subdiscretized with a combination of general and transitional shell elements. Four even rows are used vertically, whereas 14 horizontal divisions are made at variable intervals (Figure 9b). The first four lines are at 7-5-degree increments and the last 10 divisions are made at 15-degree intervals. Again, only half of the circumference is used in the analysis. The material properties are as follows: Modulus of elasticity=519,100 Poisson\u2019s ratio = 1/6 Shear factor = 5/6 The results of this analysis are plotted along with those obtained using the SHORE program (Figure 10). The results from the SHORE analysis are shown by solid lines and those calculated using the new program are indicated by small circles. Only membrane stress resultants are plotted since stress couples are negligible all along the meridian except near the lower boundary. The two analyses agree very well. Cooling tower shell-wind load A static wind load is applied to the shell in Figure 9 with the discretization as well as the material properties taken as the same as previously described. The static wind pressure is assumed to be constant over the height of the shell and the horizontal distribution of the pressure is expressed in terms of the following six Fourier harmonic coefficients: Harmonic number Fourier coefficients 0 -0.003836 1 -0.009268 2 -0.021457 3 -0.017974 4 -0.003790 5 +O-003377 In Figure 11, the results of this analysis, shown by small circles, are compared with those obtained using SHORE-I11 program which are shown by solid lines"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0000822_s11015-016-0331-6-Figure2-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0000822_s11015-016-0331-6-Figure2-1.png",
+        "caption": "Fig. 2. Schematic of the deformation zone in the CRE process.",
+        "texts": [
+            " At the fi rst stage of processing, the loose material is usually compacted until being capable of holding its shape. The next stage is extrusion of a porous compact to produce the desired cross-section. The process is schematized in Fig. 1a. Billet 1 is fed to the pass formed by two rolls 2. On its way, the billet meets die 3. The rolls create friction stresses that push the billet through the die. Figure 1b shows the distortion of the billet within the deformation zone. To solve the problem, we will use the following parameters and boundary conditions. The geometry of the deforma- tion zone is shown in Fig. 2. The effective roll radii R1 = 53.5 mm and R2 = 40.5 mm. The pass width b = 15 mm, the minimum 1 PLM Ural Company Group, Ekaterinburg, Russia; e-mail: eaa@plm-ural.ru. 2 Ural Federal University, Ekaterinburg, Russia; e-mail: j.n.loginov@urfu.ru. 3 Siberian Federal University, Krasnoyarsk, Russia; e-mail: kafomd_1@mail.ru. DOI 10.1007/s11015-016-0331-6 gap between the rolls h1 = 7 mm, the die face height hd = 22 mm, the extruded rod diameter dpr = 9 mm, the parallel land length Lpl = 2 mm, the die face angle \u03b8 = 90\u00b0, the billet height h0 = 14 mm, and the billet width b0 =14 mm"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0000992_tee.22376-Figure3-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0000992_tee.22376-Figure3-1.png",
+        "caption": "Fig. 3. Core linear magnetization curve",
+        "texts": [
+            " Moreover, the harmonic leakage resistance increases with the decrease of the phase number because the rotating harmonic magnetomotive force (MMF) increases, while the end-winding leakage resistance decreases with the decrease of the phase number. Consequently, there is no exact multiple relationship between stator leakage reactance X \u20191 under symmetrical fault condition and normal condition. 2.2. Magnetizing reactance To simplify the analysis process, the core saturation is ignored. Suppose the magnetization curve is a straight line. As shown in Fig. 3, the abscissa is per unit of excitation MMF, and the ordinate is per unit of the induced electromotive force (EMF). Assuming that excitation current I m remains constant when the number of stator winding phase reduces from n to (n \u2212 n \u2032), the excitation MMF is F \u2032 m = n \u2212 n \u2032 n \u00d7 n \u221a 2 \u03c0p W Im = n \u2212 n \u2032 n Fm (8) where F m is the excitation MMF under the rated condition, W is the turn number of each phase series winding, and p is the number of pole pairs. According to Ohm\u2019s law of a magnetic circuit, the flux is proportional to the excitation MMF when magnetic permeability is constant"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0000577_978-3-319-33714-2_5-Figure2-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0000577_978-3-319-33714-2_5-Figure2-1.png",
+        "caption": "Fig. 2 Binary link jk of type CC",
+        "texts": [
+            " Transformation matrix Tjk between the coordinate systems UjVjWj and XkYkZk, attached to the ends of the binary link with the j-th and k-th joints, has a form Tjk = t11 t12 t13 t14 t21 t22 t23 t24 t31 t32 t33 t34 t41 t42 t43 t44 2 664 3 775, \u00f01\u00de where t11 = 1, t12 = t13 = t14 = 0, t21 = ajk \u22c5 cos \u03b3jk + bjk \u22c5 sin \u03b3jk \u22c5 sin \u03b1jk, t22 = cos \u03b3jk \u22c5 cos \u03b2jk \u2212 sin \u03b3jk \u22c5 cos \u03b1jk \u22c5 sin \u03b2jk, t23 = \u2212 cos \u03b3jk \u22c5 sin \u03b2jk \u2212 sin \u03b3jk \u22c5 cos \u03b1jk \u22c5 cos \u03b2jk, t24 = sin \u03b3jk \u22c5 sin \u03b1jk, t31 = ajk \u22c5 sin \u03b3jk \u2212 bjk \u22c5 cos \u03b3jk \u22c5 sin \u03b1jk, t32 = sin \u03b3jk \u22c5 cos \u03b2jk + cos \u03b3jk \u22c5 cos \u03b1jk \u22c5 sin \u03b2jk, t33 = cos \u03b3jk \u22c5 cos \u03b1jk \u22c5 cos \u03b2jk \u2212 sin \u03b3jk \u22c5 sin \u03b2jk, t34 = \u2212 cos \u03b3jk \u22c5 sin \u03b1jk, t41 = cjk + bjk \u22c5 cos \u03b1jk, t42 = sin \u03b1jk \u22c5 sin \u03b2jk, t43 = sin \u03b1jk \u22c5 cos \u03b2jk, t44 = cos \u03b1jk, ajk\u2014a distance from the Wj axis to the Zk axis measured along the direction of the common perpendicular tjk between the Wj and Zk axes; \u03b1jk\u2014an angle between positive directions of the Wj and Zk axes measured counterclockwise about positive direction of tjk; bjk\u2014a distance from direction of tjk to direction of the Xk axis measured along the positive direction of the Zk axis; \u03b2jk\u2014an angle between positive directions of tjk and Xk axis measured counterclockwise about the positive direction of the Zk axis; cjk\u2014a distance from direction of Uj axis to direction of tjk measured along the positive direction of theWj axis; \u03b3jk\u2014an angle between positive directions of the Uj axis and tjk measured counterclockwise about the positive direction of the Wj axis. In comparision with the Denavit\u2013Hartenberg transformation matrix, having four parameters, the transformation matrix (1) has six parameters fully characterizing the relative locations of the coordinate systems UjVjWj and XkYkZk, because a free rigid body in space has six generalized coordinates. A binary link jk of type CC is shown in Fig. 2. Axes of the coordinate systems UjVjWj and XkYkZk, attached to the ends of this binary link, are chosen as follows: the Wj and Zk axes are located along the axes of rotation and translation of the cylindrical joints j and k; the origins Oj and Ok of the coordinate systems UjVjWj and XkYkZk are located in points of intersection of the Wj and Zk axes with the common perpendicular tjk between these axes; the Uj and Xk axes are located along the common perpendicular tjk; the Vj and Yk axes are completed the right-hand Cartesian coordinate systems UjVjWj and XkYkZk"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0000070_s1068798x15070175-Figure1-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0000070_s1068798x15070175-Figure1-1.png",
+        "caption": "Fig. 1. A bearing assembly for a crankshaft: (1) crankcase; (2) crankshaft bearing; (3, 4) upper and lower inserts of slip bearing; (5) crankshaft; tp1, tp2, working gaps, respectively, above and below the axis O1\u2013O5; Sp1, Sp2, radial gaps, respectively, above and below the axis O1\u2013O5.",
+        "texts": [
+            " This approach to precision permits automated selection of the components in assemblies\u2014for example, the optimization of diamet ric gaps in coaxial frictional pairs within bearing assemblies for shafts. In the crankshaft bearings of the 8ChVN15/16 die sel engine, the optimal gaps are Si = 0.088\u20130.11 mm. By selection, we may ensure the required minimum (hmin) and optimal (hopt) thickness of the oil layer in the joint and hence the maximum wear margin Sw and performance KT of the component in bearing assembly for shafts (Fig. 1) [2]. The structure of the automated control system for bearing assembly is shown in Fig. 2. Automated selec tion is based on specification of the bearing tolerances and fit for the crankcase, slip bearing inserts, and crankshaft pins in the bearings. State Standard GOST 26346\u201382 is adopted as the reference here; the specifications employed are summarized in Fig. 1 and in Tables 1\u20133. We use the following notation in automation of assembly with error compensation: \u0394_sh2, \u0394_sh3, \u0394_sh4 denote the misalignment of the crankshaft bearings 2\u20134 relative to the common axis O1\u2013O5; \u0394_k2, \u0394_k3, \u0394_k4 denote the misalignment of the crank pins 2\u20134 of crankshaft 5 relative to the common axis O1\u2013O5. Here 2\u20134 denote components assembled Keywords: crankshaft, crankshaft bearings, inserts, misalignment, assembly technology, selection, tolerances, automation, quality DOI: 10.3103/S1068798X15070175 with bearings 2\u20134 and pins 2\u20134, which have centers O2, O3, and O4, respectively. In addition, tp1 and tp2 are the working gaps, respectively, above and below axes O1\u2013O5; Sp1 and Sp2 are the radial gaps, respectively, above and below axes O1\u2013O5 (Fig. 1). The monitoring results\u2014the actual dimensions and misalignment of the crankshaft bearings and the thickness fluctuation due to machining\u2014are entered in a database. The mutually compensating characteristics for the upper and lower inserts are automatically selected from the corresponding database. These characteris tics are the actual dimensions of the crankshaft bear ings and pins and the misalignment (noncoaxiality) in the cross section corresponding to minimum discrep ancy between the frictional pair",
+            " In the automated sys tem, the thickness values of the upper and lower inserts are entered in the same database (Table 2) and a spe cially developed system of tolerance and fits is employed. We assume that the common axis of the first and fifth crankshaft pins is aligned with the axis O1\u2013O5. In other words, this is the common axis for the crankshaft bearings and crankshaft pins. The working gaps determine the difference in actual dimensions of the crankshaft bearings and pins corresponding to bearing p. This finding is used in cal culating the insert thickness. For example, the differ ence between the actual dimension D_p2 of crank shaft bearing 2 (Fig. 1) and the actual dimension D_sh2 of the corresponding pin is regarded as the actual diametric gap in the given frictional pair, which must be within the optimal values of the radial gap and is equal to the sum of the thicknesses of the upper and lower inserts in the plane of greatest discrepancy between the pin and bearing. For convenience of assembly, the model of the bearing system is developed so that the centers of the bearings are displaced in the direction of their favor able position vectors, while the axes of the extreme and intermediate bearings are in the plane of closest approach of the frictional surfaces of coaxial frictional pairs"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0001777_951293-Figure9-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0001777_951293-Figure9-1.png",
+        "caption": "Figure 9. Dynamic Hysteresis Loop",
+        "texts": [
+            " 8 shows the situation when loading (OQ) and unloading (QO,) processes are rapid and the load is removed immediately after the loading phase is completed without any time for relaxation. In the triangle OQO,, there is no elastic delay but the curve 0,O represents the reverse elastic delay which is still present in the system. If the load is applied in the opposite direction immediately after the state 0, is reached, the process of deformation and relative motion are running along the path O,Y,T as shown in Fig. 9. Fig. 9 consists of a continuation of the process shown in Fig. 8. It can be noticed that to remove the displacement 002 in this fast (dynamic) process, a negative load OY, is necessary. Let us assume that, when at state T, the next phase of motion starts without any delay and that both unloading and loading processes are similar to those between Q and T. The representation of this phase is a curve TO,Y,Q. The closed line QO,Y,TO,Y,Q is called a dynamic hysteresis loop. ) Mechanical Pro~erties of Elastic Hvsteresis In the case of dynamic hysteresis such as in Fig. 9, if phases of loading and unloading are being repeated in similar ! way, each cycle of motion would be expected to be a similar loop. Of course in actual mechanical systems, even if load I cycles are repeatable, hysteresis loops need some time to reach the steady state and be identical in each cycle. The area of the hysteresis loop in Fig. 9 represents an energy loss during a single cycle of motion. Strictly speaking, all of this energy does not have to be necessarily lost. Some part is always dissipated as heat but the remaining part may remain in the system as potential energy of residual stresses due to various local slips, lock-ups, etc. This fact creates major difficulties in empirical (for instance, calorimetric) investigations of energy losses. During oscillations, the stress-strain and forcerelative movement states are not homogeneous across the mechanical system",
+            " (7) The main practical inconvenience of this simple analytical approach is that, even in the case of a constant dynamic coefficient of friction, the amplitude of motion 1x1 varies and is a function of parameters of the system and, in particular, its initial conditions. This difficult analytical problem can be resolved in an empirical way. Assuming that one can record the relationship between friction forces and displacement, energy dissipated Ed can be measured as an area of the dynamic hysteresis loop (refer to Fig. 9). Of course, the value of energy Ed may be a function of many parameters such as the coefficient of friction, frequency and amplitude of motion, etc. Static and Dynamic Hysteresis; Level of Participation Factor From an engineering point of view, especially if analysis of clutches and drivetrains is considered, it may be preferable to describe the dry friction damping, which is a highly nonlinear and dynamic phenomenon, in terms of static hysteresis. Let us assume that one can measure energy dissipated by the system and, for instance in a static condition, energy dissipated by a system is Edst",
+            " Let us consider that the clutch damper has a constant stiffness and experiences dry friction damping determined by different but constant (independent of the relative speed) static and dynamic coefficieiits of friction. If the damper relative movement is slow (static), one can present the resisting force as a function of displacement as shown in Fig. 12 (compare to Fig. 2). Now, let us assume that the above system osculates about a certain equilibrium position E (see Fig. 13) with certain considerable frequency, and that the amplitude of its motion in a steady state is 1x1. During the motion, the torquedisplacement relationship will be a closed curve as shown in Fig. 13 (compare Fig. 9). One can expect that the size and shape of the hysteresis loop will depend on many parameters of the driveline such as: the amplitude and frequency of oscillations, location of the equilibrium position, friction characteristics, inertias of bodies D and H, stiffness of the damper, etc. To identify and investigate significant parameters and reject secondary (negligible) modeling factors, many hypothetical relationships between potentially sensitive parameters must be examined in an experimental way"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0000761_978-981-10-2404-7_29-Figure1-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0000761_978-981-10-2404-7_29-Figure1-1.png",
+        "caption": "Fig. 1 Single-legged robot constrained by a vertical slider",
+        "texts": [
+            " In this paper, the hopping height of the single-legged hopping robot is controlled by an energy compensation algorithm to compensate the energy loss in the course of the phase of robot hopping. The robot system, which is based on the kinematic model and dynamic model of the single-legged hopping robot, takes into account the inelastic impact collision with the ground. The articulated single-legged robot is composed of three rigid bodies: the upper body (consisting of hip), the thigh, and the shank with mass and their inertia with respect to the corresponding centers of mass as shown in Fig. 1. The thigh is attached to the main body by the hip joint, whereas the shank is linked to the thigh by the knee joint. The hip and knee joint are able to rotate actively within 120\u00b0 in the sagittal plane. Considering the weight of some kinds of actuator such as hydraulic cylinder, the center of mass (CoM) of the thigh and shank link deviate from the centroid of the link. The single-legged robot\u2019s upper body is constrained on a vertical slider, so that it is free to slide vertically on the slider in order to perform a vertical jumping motion",
+            " Robot kinematic are described as a function of the vector, q \u00bc qb qr\u00f0 \u00deT \u00f01\u00de which includes the vector qb describing the unactuated floating base coordinate and the actuated joint coordinate qr. As the x axis is fixed by the vertical slider, the information of x axis is not required here. So that the vector qb has only the vertical height z. The vector qr has the element of hip and knee joint angle qr \u00bc qH qK\u00f0 \u00deT. qH and qK are hip and knee joint angle, respectively, the symbol of which is as shown in Fig. 1. In summary, we get the minimal coordinates q \u00bc z qH qK\u00f0 \u00deT \u00f02\u00de According to the Kinematics of the robot, we get the CoG(Center of Gravity) of thigh rmT and shank rmS, here the abscissa value of the floating base coordinate is regarded as 0. rmH \u00bc 0 z\u00f0 \u00deT \u00f03\u00de rmT \u00bc xmT zmT \u00bc lmT sin qH \u00fe e1\u00f0 \u00de z lmT cos qH \u00fe e1\u00f0 \u00de \u00f04\u00de rmS \u00bc xmS zmS \u00bc lT sin qH \u00fe lmS sin qH \u00fe qK \u00fe e2\u00f0 \u00de z lT cos qH lmS cos qH \u00fe qK \u00fe e2\u00f0 \u00de \u00f05\u00de the CoG of legged robot, rCoG \u00bc mHrmH \u00femTrmT \u00femSrmS mH \u00femT \u00femS \u00f06\u00de and the end-effector position of the legged robot, ree \u00bc xee zee \u00bc lT sin qH \u00fe lS sin qH \u00fe qK\u00f0 \u00de z lT cos qH lmS cos qH \u00fe qK\u00f0 \u00de \u00f07\u00de where mH, mT, mS denote CoM of hip, thigh, and shank",
+            " Furthermore, a high-level controller can be implemented as a state machine that simply changes virtual component connections or parameters at the state transitions (Pratt et al. 2001). The virtual model control validates the instant the legged robot touches down the ground and invalidates when the leg lifts off the ground. During the stance time, the spring virtually retracts and extends along with the fluctuation of the hip and energy is interchanging between the robot system and the spring as Fig. 1 shows. On compression section, kinetic energy of the robot system transforms to the spring and on thrusting, potential energy of spring transforms to the robot system. Using the kinematics of the legged robot, the force of the virtual spring can be transformed into the joint torque with the transpose of the Jacobian matrix. s \u00bc JTFHFv \u00f023\u00de JFH \u00bc @rmH @ree @qr \u00f024\u00de While both, the matrix JFH is the relative Jacobian of the robot, which is solely a function of joint coordinate. The virtual force Fv, which is simply calculated as a linear spring element between the end-effector of robot and hip joint, was decomposed into a vertical and a horizontal component"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0001020_s12221-016-6478-8-Figure6-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0001020_s12221-016-6478-8-Figure6-1.png",
+        "caption": "Figure 6. Images of (a) schematics of fixture, (b) types of fixtures, (c) different strands, and (d) connection I details.",
+        "texts": [
+            " Resonance frequency, S11, and S21 of different lengths have been measured and compared by using network analyzer (E5061B, Agilent Co.) in order to quantify the data transmission characteristics. Effect of ply numbers has been investigated by placing two yarns in parallel on the specially invented fixture. Those yarns have been further used in order to study the effect of twist numbers by incorporating varying twist levels. Teflon based PCB has been used after etching in order to enhance measurement reliability, and showed fixtures with two separate sections for signal and ground in Figure 5 and Figure 6. Figure 7 shows cross-sectional SEM images of urethane coated copper fibers under different magnifications. It has been revealed that pristine and urethane coated copper fibers have 70 \u03bcm and 80 \u03bcm, respectively. In spite of irregular boundary between copper fiber and urethane coated, average coating thickness has been determined to about 5 \u03bcm, and thus cross-sectional area (A) was calculated to be 3.846\u00d7 10-9m2 by equation (1). dCu (Diameter of Cu fiber) = 7 \u00d7 10-5 (m) (1) A (Area of Cu fiber cross section) = = 3"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0000868_fpmc2016-1767-Figure3-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0000868_fpmc2016-1767-Figure3-1.png",
+        "caption": "Figure 3: Total Fluid Domain Assembly",
+        "texts": [
+            " Solving the problem different passages between mesh generator and solid modeler was necessary because of the very complex mating (i.e. splined coupling), very thin clearances, and complex geometries. The filling and subtraction tasks in general are undertaken for limited subset of assembly after the fixing of geometry errors. At the end of this very time consuming phase all the partial volumes have to be assembled to obtain the various computational subdomain. The subdivision of the computational domain into subdomains is necessary to define different speed rotating domains, which communicate via interface surfaces. In figure 3 the complete fluid domain is colored according to Figure 2 color scheme. The Range 1 Operating Conditions in Terms of shaft speeds are summarized on table 1: the various shaft speeds vary with respect to output speed that is bound between a minimum and a maximum. For obligation of confidentiality all the shaft speeds were normalized with respect to a reference nominal speed. In a previous work [11] the same geometry was analyzed, in particular consistent results were found for stationary condition but for rotating shafts conditions many inconsistencies have arisen"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0002252_s0890-6955(98)00006-6-Figure7-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0002252_s0890-6955(98)00006-6-Figure7-1.png",
+        "caption": "Fig. 7. The critical feed fc, and the critical length lc of the worn tool.",
+        "texts": [
+            " 5, and can be obtained by the following equations: EP = (l21 + h2 1)1/2; PC = (l2 2 + h2 2)1/2; CN = (l23 + h2 3)1/2 and MN = (h2 5 + l24)1/2 (10) uR1 = tan21(l1/h1); uR2 = tan21(h3/l3); uR3 = tan21(h5/l4) and uPC = tan21(h2/l2) (11) R1 = EP/(2sinuR1); R2 = CN/(2sinuR2) and R3 = MN/(2sinuR3); (R2 = R3) (12) Due to the variations in Cs and the different kinds of geometrical circumferences of tool edge, several different cutting conditions occur when the feed is varied. In order to understand the cutting conditions, a criterion for determining the critical length and critical feed are developed and shown in Fig. 7. The critical feed fc and critical length lc are: lc = (l\u00b7sinCs 2 m\u00b7cosCs)/sin(uPC 2 Cs) (13) f c = (l 2 lccosuPC)/cosCs (14) where l = l1 + l2 + R2sin2uR2 + R3sin2uR3 2 R1(1 2 cos2uR1) (15) m = h1 + h2 + R3cos2uR3 2 R2cos2uR2 2 R1(sin2uR1) (16) Two conditions are developed for the cutting process. First the straight line PC is intersected by curve only. And second, the line PC is intersected by curve , or curve EP is intersected by curve . In the first case, wherein the straight line PC is intersected by curve , feedrate (f) can be divided into another two cases, depending on the cutting condition, as: 1"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0002229_jsvi.1996.0017-Figure9-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0002229_jsvi.1996.0017-Figure9-1.png",
+        "caption": "Figure 9. Poincare\u0301 map for b=4\u00b7565 with a=ac , d=0\u00b71, v=1, G=1.",
+        "texts": [
+            " The behaviour is as in (2a) above, but with Ws ho is \u2018\u2018outside\u2019\u2019 Wu he . (3a) 4\u00b7464QbQ5\u00b7590. This is a range of interactive transversal manifold intersections between Wu ho and Ws he . However, within this range we do not expect to find a chaotic regime, since the solution paths become unbounded as t:a. Physically, the self-exciting term containing b becomes large and dominates the damping effect. (3b) \u22124\u00b7008QbQ3\u00b70127. Here the solution curves are bounded but numerical evidence suggests that they approach stable limit cycles or fixed points as t:a. (4a) b=4\u00b7564 (Figure 9). There are interactive tangential manifold intersections. Ws he is \u2018\u2018outside\u2019\u2019 Wu ho . There are no other transversal intersections. This value marks the end of this particular region of intersections. (4b) b=\u22123\u00b70127. The behaviour is as for (4a) above, but Wu he is \u2018\u2018outside\u2019\u2019 Ws ho . (5a) 3\u00b70771QbQ4\u00b7565 (Figure 10). There are no transversal intersections of any kind. This is essentially a window free from intersections and, therefore, the possibility of any chaotic motion. These exist two unstable fixed points in the Poincare\u0301 map which correspond to two unstable periodic orbits in phase-space"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0002310_20.619563-Figure1-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0002310_20.619563-Figure1-1.png",
+        "caption": "Fig. 1 Positions of B-sensing coils.",
+        "texts": [
+            " The magnetic losses of the motor at the PWM drive are greater than those expected from the material losses and increase by a certain rate from those at the sine drive in spite of the complex distribution of magnetic induction in the motor core. 11. EXPERIMENTAL PROCEDURE We have employed a 1.5 kW 3-phase squirrel-cage induction motor which has 4-poles, 36-stator slots and 44- rotor slots. The stator core and the rotor core are made from non-oriented electrical iron sheets (ASTM 47F460, JIS 50A800, DIN VSOO-SOA). As shown in Fig. 1, B-sensing coils (10 turns) have been wound at the stator core (position S), 9 stator teeth in one pole (positions T1-T9), and the rotor core (position R) of the motor so that waveforms of induced voltage and magnetic induction in the motor driven by a sinusoidal power supply (f = 60 Hz) and a V/f = constant PWM inverter power supply have been observed (Fig. 2). I. INTRODUCTION Squirrel-cage induction motors have recently been driven by PWM inverter power supplies for variable speed and torque controls"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0000978_tac.2016.2645178-Figure4-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0000978_tac.2016.2645178-Figure4-1.png",
+        "caption": "Fig. 4: Quadrocopter model and control problem.",
+        "texts": [
+            " 1, which explicitly allows for non-zero off-diagonal blocks in the multiplier resulting from such a shift. In contrast, Lma. 2 prescribes zero off-diagonal blocks under all circumstances. Consequently, a symmetric ball that wrongly considers also part of the negative real axis for the range of singular values is considered, [6]. The comparison of the respective methods for specific interconnection matrices documented in Fig. 3 reveals that method (i) is consistently providing better performance than methods (ii) and (iii). Consider the quadrocopter model illustrated in Fig. 4(a), of which an LFT-LPV model given in (10) is derived in [2]. Control of rotation about x3,k is assumed to be separate, leading to the orientation vector q = [q1, q2] \u22a4, where each angle is limited to [\u221240, . . . , 40]\u25e6. The mass m is 0.64 kg, J1 = 0.004 20 kgm2, I2 = 0.008 15 kgm2 and g = 9.81 kgm/s2. The inertial and body coordinate systems are denoted x = [x1, x2, x3] \u22a4 and X = [X1, X2, X3] \u22a4, resp. The input u1 = \u22114 i=1 u\u0303i is the total force in X3 direction, whereas u2, u3 are the torques around axes X1, X2, resp. A single leader and five followers are considered. Agent 1 is a virtual leader with an integrator model H1(s) = 1/sI3 in each of the position coordinates. To simplify, x3,k \u2261 0,\u2200t > 0. The agents Hk, k = 2, 3, . . . , 6 are quadrocopter models whose individual coordinates are yk = [x1,k, x2,k, x3,k] \u22a4 and qk = [q1,k, q2,k] \u22a4. Fig. 4(b) shows the generalized plant with shaping filters WS = 5 s+0.05I3, and WKS = s+0.1 s+10000I3 used to consider tracking and to penalize the control input. The interconnection matrix is chosen as \u2113(t) = [ 0 0 \u2113G2G1 (t) \u2113G2G2 (t) ] . The formation control system\u2019s performance is compared between the approaches (i)\u2013(iii) as in Section IV-A. The position information of each agent is broadcasted, which implies that the agents have no knowledge about the number of recipients. In graph theoretical terms, this means that the interconnection matrix may be row-normalized, but not column-normalized"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0000580_s12555-015-0054-7-Figure10-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0000580_s12555-015-0054-7-Figure10-1.png",
+        "caption": "Fig. 10. Schematic of ship roll damping system due to fin stabilizer.",
+        "texts": [
+            " Moreover, only 1 FSE-FLS type approximation is required to compensate for sum of the uncertainties in this note, independent of the system order. That effectively avoids the conflict of each adaptive laws in the conventional burden. Therefore, the merit \u201cless computational burden\u201d is obvious that would facilitate its implementation in the practical engineering. Example 3: To further show the developed algorithm\u2019s effectiveness in the industrial application, we consider a typical problem in field of ocean engineering, i.e., ship roll damping control using fin stabilizers shown in Fig. 10. The research topic has been revitalized by the luxury yacht industry in recent years [39]. The mathematical model for the roll damping control design can be written as follows: \u03d5\u0307 = p, (Ix +Kp\u0307) p\u0307+(Kp +2r f K\u03b1U) p+K|p|p |p| p+K\u03d5 \u03d5 =\u22122U2K\u03b1 \u03b1 +Kw (t) , (49) where \u03d5 , p denote the ship roll angle, angular rate respectively. (Ix +Kp\u0307) is the moment of inertia and added moment of inertia, Kp,K|p|p denote damping coefficients, K\u03d5 = \u03c1g\u2207GMt is a restoring term coefficient due to gravity and buoyancy. Kw (t) is the moment acted on ship by wave and wind, and the term \u22122U2K\u03b1 \u03b1 describes the roll moment induced by the fins with respect to the forward speed U of the vessel and the mechanical angle \u03b1 of the fin"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0000482_ijmic.2016.077748-Figure1-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0000482_ijmic.2016.077748-Figure1-1.png",
+        "caption": "Figure 1 System schematic diagram and the visco-elastic model (see online version for colours)",
+        "texts": [
+            " An example is the robot developed by RIKEN-TRI collaboration centre, RIBA, which uses its whole arms in human transfer (Mukai et al., 2010; 2011). In such application, the goal is to manipulate a multi-link heavy object using whole arm manipulation. Multi-link object manipulation could be considered as an initial framework for the most sophisticated process of the human manipulation; suggested by Onishi et al. (2003). This study presents the dynamic modelling, static analysis, and non-prehensile manipulation control of a three-rigid-link object manipulated by two cooperative robot arms in a two dimensional space (see Figure 1). Whole arm manipulation lies under the taxonomy of non-prehensile manipulation; manipulation without firm grip where simple manipulators are utilised instead of complex and dexterous ones (Lynch et al., 1998; Maeda et al., 2004; Behave et al., 2015). Gravitational, centrifugal, and Coriolis forces are utilised as virtual motors to exhibit more DOFs of a manipulated object such as slipping, rolling, and free flight (Asano et al., 2003). For non-prehensile manipulation, both the object and the manipulator geometries are important",
+            " A dynamic model of the system has been derived, static analysis has been considered and a controller for lifting up the object in a plane has been designed. The object\u2019s behaviours were considered with and without friction contacts with the arms, where friction has significant effect throughout implementation in motion control (Merola et al., 2015). In this work, a three-rigid-link object to be manipulated by two robot arms in a two dimensional space is proposed. Two links are kept in contact with one arm, while the third link is in contact with the other arm, see Figure 1. This configuration is more realistic framework to a lifted human rather than the early considered one by Zyada et al. (2010, 2011). This work is divided into two parts; the first one presents the dynamic modelling of the system, and the equilibrium contact points exploration at different configurations of the object links. The effect of changing the direction of the static frictional forces at the multi contact points on the equilibrium points is discussed. The other part includes the design of a controller to perform the object manipulation according to a desired trajectory of the object state variables",
+            " The static analysis with frictionless contact between the object and the arms is given in Section 3. The effect of the static friction on the static analysis problem is also discussed in this section. The design of the controller is given in Section 4. Simulation results and discussion are given in Section 5. Conclusions and recommendations for future work are presented in Section 6. 2 System dynamics In this section, system description, a dynamic model of the system, kinematic constraints, and visco-elastic model at a contact are presented. 2.1 System description Figure 1 illustrates the schematic diagram of the system. The object is represented by three rigid links (link-1, link-2, and link-3) connected by two passive joints: the knee and the hip joints. The angular positions of the links are given by \u03b81, \u03b82, and \u03b83, respectively, see Figure 1. The position of the knee joint (x, y) is referred to the world coordinate frame \u03a3o. The position of the hip joint is related to that of the knee one in terms of the length of link-2, L2 and its angle \u03b82 by the equation: 2 2 2 2 cos sin h h x x L \u03b8 y y L \u03b8 = + = + (1) The position of either the knee or the hip joint along with the angular positions of the three links complete the definition of the position and orientation of the object relative to \u03a3o. The masses and mass moment of inertia of the links are denoted by m1 and J1, m2 and J2, m3 and J3 for link-1, link-2, and link-3, respectively",
+            " B(q, l) and u are defined by: 1 1 1 1 2 2 2 2 3 3 3 3 1 0 1 0 1 0 0 1 0 1 0 1 ( , ) sin cos 0 0 0 0 0 0 sin cos 0 0 0 0 0 0 sin cos B q l l \u03b8 l \u03b8 l \u03b8 l \u03b8 l \u03b8 l \u03b8 \u23a1 \u23a4 \u23a2 \u23a5 \u23a2 \u23a5 \u23a2 \u23a5= \u2212 \u23a2 \u23a5 \u2212\u23a2 \u23a5 \u23a2 \u23a5\u2212\u23a3 \u23a6 (4) [ ]1 1 2 2 3 31 2 3 T TT T T x y x y x yu u u u F F F F F F= =\u23a1 \u23a4\u23a3 \u23a6 (5) ( ) ( ) ( ) ( ) ( ) ( ) ( ) ( ) ( ) ( ) ( ) ( ) ( ) ( ) ( ) ( ) ( ) ( ) 1 2 3 1 1 1 2 2 3 2 2 3 3 3 1 2 3 1 1 1 2 2 3 2 2 3 3 3 2 1 1 1 1 1 1 1 1 1 2 2 2 2 3 2 2 2 2 3 2 2 2 2 3 3 2 3 2 322 3 3 3 3 3 0 sin sin sin 0 cos cos cos ( ) sin cos 0 0 sin cos 0 cos sin cos k k k k k m m m m L \u03b8 m L m L \u03b8 m L \u03b8 m m m m L \u03b8 m L m L \u03b8 m L \u03b8 M q m L \u03b8 m L \u03b8 J m L m L m L \u03b8 m L m L \u03b8 J m L m L m L L \u03b8 \u03b8 m L \u03b8 m L \u03b8 + + \u2212 \u2212 + \u2212 + + + = \u2212 + \u2212 + + + + \u2212 \u2212 ( ) ( ) ( ) ( ) ( ) ( ) ( ) ( ) ( ) ( ) ( ) ( ) ( ) ( ) 2 3 3 2 3 2 3 3 3 3 1 1 1 1 2 2 3 2 2 2 3 3 3 3 1 1 1 1 2 2 3 2 2 2 3 3 3 3 3 2 3 3 2 3 3 2 3 2 2 3 , 0 cos 0 0 cos cos cos 0 0 sin sin sin , 0 0 0 0 0 0 0 0 0 sin 0 0 0 sin 0 k k m L L \u03b8 \u03b8 J m L m L \u03b8 \u03b8 m L m L \u03b8 \u03b8 m L \u03b8 \u03b8 m L \u03b8 \u03b8 m L m L \u03b8 \u03b8 m L \u03b8 \u03b8 C q q m L L \u03b8 \u03b8 \u03b8 m L L \u03b8 \u03b8 \u03b8 \u23a1 \u23a4 \u23a2 \u23a5 \u23a2 \u23a5 \u23a2 \u23a5 \u23a2 \u23a5 \u23a2 \u23a5 \u23a2 \u23a5\u2212 +\u23a3 \u23a6 \u23a1 \u23a4\u2212 \u2212 + \u2212 \u23a2 \u23a5 \u2212 \u2212 + \u2212\u23a2 \u23a2= \u23a2 \u2212\u23a2 \u23a2 \u2212 \u2212\u23a3 \u23a6 ( ) ( ) ( ) ( )1 2 3 1 1 1 2 2 3 3 2 3 3 3 , ( ) 0 cos cos cos T kG q m m m g m L g \u03b8 m L m L g \u03b8 m L g \u03b8 \u23a5 \u23a5 \u23a5 \u23a5 \u23a5 \u23a1 \u23a4= + + +\u23a3 \u23a6 (6) with 11 1 1 1 11 1 1 22 2 2 2 22 2 2 33 3 3 3 33 3 3 sin cos cos sin sin cos cos sin sin cos cos sin xN yT xN yT xN yT F\u03b8 \u03b8 F u F\u03b8 \u03b8 F F\u03b8 \u03b8 F u F\u03b8 \u03b8 F F\u03b8 \u03b8 F u F\u03b8 \u03b8 F \u2212 \u23a1 \u23a4\u23a1 \u23a4 \u23a1 \u23a4 = = \u23a2 \u23a5\u23a2 \u23a5 \u23a2 \u23a5\u2212 \u2212\u23a3 \u23a6 \u23a3 \u23a6 \u23a3 \u23a6 \u2212 \u23a1 \u23a4\u23a1 \u23a4 \u23a1 \u23a4 = = \u23a2 \u23a5\u23a2 \u23a5 \u23a2 \u23a5 \u23a3 \u23a6 \u23a3 \u23a6 \u23a3 \u23a6 \u2212 \u23a1 \u23a4\u23a1 \u23a4 \u23a1 \u23a4 = = \u23a2 \u23a5\u23a2 \u23a5 \u23a2 \u23a5 \u23a3 \u23a6 \u23a3 \u23a6 \u23a3 \u23a6 (7) The distances l1, l2 and l3 the relative velocities v1, v2 and v3 are given as: ( ) ( ) ( ) ( ) ( ) ( ) 1 1 1 1 1 2 1 2 1 2 3 2 3 2 3 sin cos sin cos sin cos A A A A A h A h l y y \u03b8 x x \u03b8 l y y \u03b8 x x \u03b8 l y y \u03b8 x x \u03b8 \u23ab= \u2212 + \u2212 \u23aa = \u2212 + \u2212 \u23ac \u23aa= \u2212 + \u2212 \u23ad (8) ( ) ( ) ( ) ( ) ( ) ( ) 1 1 1 1 1 2 1 2 1 2 3 2 3 2 3 sin cos sin cos sin cos A A A A A h A h v y y \u03b8 x x \u03b8 v y y \u03b8 x x \u03b8 v y y \u03b8 x x \u03b8 \u23ab= \u2212 + \u2212 \u23aa = \u2212 + \u2212 \u23ac \u23aa= \u2212 + \u2212 \u23ad (9) 2.3 Contact model The contact between the arm and the object can be modelled using the visco-elastic model (Zyada et al., 2010), see Figure 1, as: , for 0 0, for 0 i i i i i iN i k \u03b4 d \u03b4 \u03b4 F \u03b4 \u23a7 + \u2265\u23aa= \u23a8 <\u23aa\u23a9 (10) where FiN is the normal component of the interaction force, \u03b4i is the radial deformation of the arm in the direction of the normal force, i\u03b4 is its rate of change, ki, di denote the spring constant, and damping coefficient respectively, i = 1, 2, 3. \u03b4i can be calculated in terms of q and (xAj, yAj): ( ) ( ) ( ) ( ) ( ) ( ) 1 1 1 1 1 2 1 2 1 2 3 2 3 2 3 sin cos sin cos sin cos A A A A A h A h \u03b4 r x x \u03b8 y y \u03b8 \u03b4 r x x \u03b8 y y \u03b8 \u03b4 r x x \u03b8 y y \u03b8 \u23ab= + \u2212 \u2212 \u2212 \u23aa = \u2212 \u2212 + \u2212 \u23ac \u23aa= \u2212 \u2212 + \u2212 \u23ad (11) 2"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0000345_s0025654416010064-Figure1-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0000345_s0025654416010064-Figure1-1.png",
+        "caption": "Fig. 1.",
+        "texts": [
+            " In contrast to existing papers dealing with contact modes of motion of a flexible rotor on a rigid stator [3, 5\u201310] and on an elastic support with a gap [2, 4, 11, 12], the present paper takes into account the internal friction in the rotor material, the inhomogeneity of the rotor rotation, and the technological misalignment of the rotor and stator axes. These dynamical factors a priori intensify the impact modes of the supercritical rotor especially in the case of a rigid stator. We assume that the shaft is horizontal, cantilever, viscoelastic in bending and torsion, moderately long and sufficiently rigid in compression-extension, clamped at one end, and connected with an unbalanced rigid disk inside a rigid support with a gap at the other end. The \u201cflexible rotor\u2013rigid stator\u201d system under study is outlined in Fig. 1. In such a system, the axial vibrations of the rotor and the angular displacements of its axis are negligibly small, and hence we can restrict our consideration of the problem of the system motion to the linear approximation in the framework of the plane model; namely, we can assume that all points of the disk only move in directions parallel to the plane Oxy. Simultaneously, in addition to the immovable (Cartesian) coordinate system Oxyz, it is convenient to use the coordinate system Ouvz rotating with the rotor, namely, with its unbalance vector a",
+            " By substituting these expressions into the Lagrange equation, we obtain the equations of motion of the unbalanced unloaded flexible rotor in Cartesian coordinates, mRx\u0308R + (bR + dR)x\u0307R + kRxR + bR\u03d5\u0307yR = mRa(\u03d5\u03072 cos \u03d5 + \u03d5\u0308 sin \u03d5), mRy\u0308R + (bR + dR)y\u0307R + kRyR \u2212 bR\u03d5\u0307xR = mRa(\u03d5\u03072 sin \u03d5 \u2212 \u03d5\u0308 cos \u03d5), IR\u03d5\u0308 + mRa(y\u0308R cos \u03d5 \u2212 x\u0308R sin \u03d5) + bR(x\u0307RyR \u2212 xRy\u0307R) + bR\u03d5\u0307r2 R = 0. Here IR = IG + mRa2, where IR is the axial moment of inertia of the disk and rR = \u221a x2 R + y2 R. In the presence of an external torque M0, a considerable moment of inertia I0 of the electrical drive, and the shaft viscoelastic behavior in torsion, it is necessary to supplement this equation with the corresponding force factors. If we fictitiously cut the shaft of this system (Fig. 1), then we can assume that, in addition to the torque M0, its left part is subjected to the moments \u2212k\u0303R(\u03d50 \u2212 \u03d5)of elasticity forces and \u2212b\u0303R(\u03d5\u03070 \u2212 \u03d5\u0307) of viscosity forces, where k\u0303R is the shaft rigidity in torsion, b\u0303R = k\u0303R\u03a8R/(2\u03c0\u03bb\u0303R) is the coefficient of internal friction, whose absorption value \u03a8R is assumed to be the same as in bending, and \u03bb\u0303R is the rotor torsional vibration frequency. The right part of the system is subjected to the same moments of the shaft viscoelastic resistance forces but with opposite signs"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0000786_ijbic.2016.078664-Figure3-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0000786_ijbic.2016.078664-Figure3-1.png",
+        "caption": "Figure 3 Missile kinematics in three dimensions",
+        "texts": [
+            " cos( ) z zA A\u03b3 V V = \u2248\u03b1 (3) where V is the magnitude of the velocity vector. The normal acceleration (Az) is evaluated as following equation. ( , )z z F \u03b4A m = \u03b1 (4) where m is the missile mass and Fz(\u03b1, \u03b4) is the forces applied to the missile. By substituting (3) and (4) into (1) and combining with (2), the result is two nonlinear differential equations expressed in (5). ( ( , ),, )zF M\u03b8 \u03b8 m \u03b4 V \u03b4 J = \u2212 =\u03b1 \u03b1\u03b1 (5) In reality, the missile motion equations are expressed in three dimensions. The missile kinematics in three dimensions are shown in Figure 3. Where, the definitions of these variables are tabulated in Table 1. The roll (p), pitch (q), and yaw (r) rate are the three components of the missile angular velocity vector in bodyfixed coordinates. Assuming that products of inertia are zero, the governing equations on the translational and rotational motion are expressed in (6) and (7), respectively. 2 2 2 2 cos tan sec tan tan tan cos sec tan tan tan tan zb xb yb xb F Fp q r mu mu F Fr p q mu mu \u239b \u239e= \u2212 + \u2212 \u2212\u239c \u239f \u239d \u23a0 \u239b \u239e= \u2212 + \u2212 +\u239c \u239f \u239d \u23a0 \u03b1 \u03b1 \u03b2 \u03b1 \u03b1 \u03b1 \u03b2 \u03b2 \u03b2 \u03b2 \u03b1 \u03b2 \u03b1 \u03b2 (6) z yxb x x yb x z y y z yzb z z I IMp qr I I M I Iq pr I I I IMr pq I I \u2212 = \u2212 \u2212= \u2212 \u2212 = \u2212 (7) The lateral components of the missile inertial translational acceleration are evaluated as follows"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0003560_nme.1620180606-Figure6-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0003560_nme.1620180606-Figure6-1.png",
+        "caption": "Figure 6. Geometry and displacement components of transitional shell elements",
+        "texts": [
+            " This peculiar behaviour may be understood if it is recalled that the purpose of the shape functions on the upper boundary is to vary the nodal value A s in the q direction, and that the nodel value itself is changing along the 6 direction. A similar approach may be used to generate line nodes and shape functions for a cubic as well as for a linear element. The details of these procedures are not presented here and only the final results are shown in Figure 5 . TRANSITIONAL SHELL ELEMENT Two typical transitional elements are shown in Figure 6 . The elements are similar as those introduced in References 3 and 7, except for the line nodes on the upper boundaries. The polar co-ordinates of any point in the element may be related to the curvilinear co-ordinates by where N, is defined in the preceding section and V3i is the normal vector at a node i whose matnitude is equal to the thickness of the shell at the point. Co-ordinate 8, of the moving node in equation (1 1) varies along the line as 1-5 l+t ern = - O l + - O l + l 2 2 where Om, 0, and 8,+, represent the 8 co-ordinates at the moving, left corner and right corner nodes on the line, respectively. The Cartesian co-ordinates at the point may be obtained as Displacements are expressed in terms of five components at each node (Figure 6). These are three components of translation (Ui, Vi and Wi) and two rotational components (ai and P i ) about the two principal directions of the shell surface. Since a displacement component along the line node depends on the corresponding displacement component of the rotational shell element that encompasses this line on the common boundary, the displacement function of the rotational element will be examined first. A displacement vector for the rotational element4\" has five components: u4, U o , u,, & and Po"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0000050_iemdc.2015.7409124-Figure3-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0000050_iemdc.2015.7409124-Figure3-1.png",
+        "caption": "Fig. 3. Stator coils, PM polarity and Hall sensor pos connected in pairs; (b) neighboring coils connected stator pole, 3 \u2013 coil of one phase, 4 \u2013 coil of the s sensor located between two stator poles.",
+        "texts": [
+            " A Hall sensor detects the polarity of PMs and via solid state devices switches the DC voltage from one stator coil to another. The speed of the fan motor is controlled by adjusting either the DC voltage or pulse width in lowfrequency PWM [10]. In spite the PM BLDC motor has four dead spots per revolution, it has god self-starting capability. Since the rotor rests between the poles of PMs at zero-current state Analysis of Steady-State and Transient Performance of Two-Phase PM Motors for Computer Fans J. F. Gieras, Fellow, IEEE, D. Chojnowski, P. Mikulski S (Fig. 3), and instantly rotates 45o when fi will not stop on one of its dead spot. Two phase stator winding consists of f around the stator pole cores. There are rst switched on, it 3 \u2013 front surface, 4 \u2013 r structure, 7 \u2013 blades, for computer fans: (a) nt poles; (c) PCB; (d) ition: (a) opposite coils in pairs. 1 \u2013 PM, 2 \u2013 econd phase, 5 \u2013 Hall our coils wrapped four coils in the inner stator (Fig. 3b), while different magnetic polarity. T pairs, either each one with its with its neighboring coils (Fig of the outer rotor, there are fo (Fig 2b, Fig. 3). Typically, a 12-V DC cool rotor-blade assembly containin stator. A Hall sensor detects th switches 12 V DC from one st 3). Varying the supplied DC v most fans. A 12-V DC fan 3.5\u20265.0 V DC voltage appl when increasing voltage is supp Typical electronic circuits PC fan motors are shown in F fans used in computers use s two to four pins. The first tw deliver power to the fan mo optional, depending on fan desi \u2022 ground; \u2022 power (+12 V); \u2022 sense: provides a tachomet actual speed of the fan as a p proportional to speed (with eac pulses sent through this pin; \u2022 control: provides a PWM si to adjust the rotation speed voltage delivered to the cooling two neighboring coils have he coils are connected in opposite coils (Fig. 3a), or . 3b). Around the perimeter ur PMs in N-S-N-S pattern ing fan motor consists of a g a 4-pole PM, and a 4-pole e rotating magnetic field and ator coil set to another (Fig. oltage can vary the speed of might start rotating with ied, and increase its speed lied. for feeding and controlling ig. 4. The common cooling tandardized connectors with o pins are always used to tor, while the rest can be gn and type: er signal that measures the ulse train, frequency being h fan rotation, there are two gnal, which gives the ability without changing the input fan",
+            " Bearings are a critical component because bearings make the fan rotate sm reduce friction, allow the fan to operate a are partly responsible for the life expect fan in a computer and the noise level of fa bearings can be used in a cooling fan: (a (b) ball bearings, and(c) fluid dynamic bea The specifications of the investigate brushless motor are given in Table 1. motor is shown in Fig. 5. The stator hous have been removed. TABLE I Specifications of investigated computer Rated input power No-load speed Rated voltage Rated current Number of poles Stator core outer diameter Axial length of stator stack PM Outer diameter (OD) PM inner diameter (ID) Axial length of PM The ring-shaped PM is made of an ferrite with remanent magnetic flux dens coercivity Hc = 260 kA/m. The built-in converter receives position information placed in the q-axis of the stator (Fig. 3). trolling computer nected to a 4-pin er) pin is used to control pin is an h requires a pullke linear voltage oportional to the upply voltage; the sed on the control erating at 25 kHz, peed. Typically, a and 100% of the 100% duty cycle. until a threshold ld) at low control manufacturers to performed on a onstantly monitor lements). If the unit increases the e rotor speed and in a cooling fan oothly. Bearings t high speeds, and ancy of a cooling ns. Three types of ) sleeve bearings, rings. d two-phase PM The disassembled ing and fan blades case fan motor 1"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0003666_aim.2001.936770-Figure7-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0003666_aim.2001.936770-Figure7-1.png",
+        "caption": "Fig. 7. Division of desired path in finite control states",
+        "texts": [
+            " A corner is identified if the difference in orientation exceeds a given threshold. The position of tvhe corner then easily follows from the intersection of two lines. The exact location of the corner, however, is updated afterwards, when the optical axis of the vision systems is again positioned over the contour (as it was before the corner occurred) in which case the contour measurement is more accurate. Finally, the corner position is offset by the tool radius. This offset corner is the starting point of the tool path around the corner, as shown in figure 7. 111. CONTROL STRUCTURE Figure 4 gives an overview of the control structure. An external Cartesian (velocity) space control using vision and force sensors is implemented around the low level servo controlled robot. The Cartesian control loop, keeps the contour in the camera field of view, while maintaining a constant normal contact force. [l] describes this double control in full detail. It is implemented by dividing the orthogonal control space in either force controlled, vision controlled, velocity or tracking directions",
+            " The desired angular feedforward velocity e e ~ f f is -ee% Rt - tFy/kt eeWff = with eeux the tangent velocity of the end effector (or tool), Rt the tool radius, tFy the normal contact force and kt the tool stiffness. Due to the compliance of the tool, the z-axes of end effector and tool frames do not coincide. At the corner the end effector makes a sharper turn than the tool. The radius of this turn is smaller than the tool radius by a distance ,FY/kt. This explains equation 11. After the corner is taken, the (tangent) velocity will gradually build up and the controller returns to the normal operation state. Figure 7 gives an example which matches the desired path to the different control states. Some control specifications are given. IV. EXPERIMENTS A . Experimental set-up The experimental set-up consists of a KUKA 361 robot, with a SCHUNK force sensor together with an eye-in-hand SONY CCD XC77 camera with 6.15 mm lens. The CCD camera consists of 756 by 581 square pixels with 0.011 mm pixel width, from which however only a non-interlaced subimage of 64 by 128 pixels in 256 grey levels is used. Instead of the commercial controller, our own software environment COMRADE [14] is used"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0000655_jfm.2016.579-Figure1-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0000655_jfm.2016.579-Figure1-1.png",
+        "caption": "FIGURE 1. (a) Circular cylinder under a uniform electric field parallel to an adjacent wall. As a result of the electro-osmotic flow, a stationary cylinder experiences a hydrodynamic force in the x-direction and torque in the z-direction. (b) Circular cylinder under a uniform current perpendicular to an adjacent reactive electrode. As a result of the electro-osmotic flow, a stationary cylinder experiences a hydrodynamic force in the y-direction.",
+        "texts": [
+            " Boston University, Mugar Memorial Library, on 02 Jan 2017 at 22:54:01, subject to the Cambridge Core terms of use, torques on such a cylinder have been obtained, the resulting rectilinear and angular velocities of a comparable freely suspended cylinder are readily determined using the available resistance coefficients pertaining to the cylinder\u2013wall geometry (Jeffrey & Onishi 1981). This decomposition into pure electro-osmosis and pure rigid-body motion is useful in isolating the different singular scaling associated with each component. The problem considered herein is described in figure 1(a). An infinite dielectric cylinder of circular cross-section (radius a) is held stationary within an electrolyte solution (dielectric permittivity \u03b5, viscosity \u00b5) with its axis parallel to an adjacent dielectric solid wall. The separation distance between the cylinder and the wall is a\u03b4. The cylinder\u2013liquid system is exposed to an otherwise uniform electric field, of magnitude E\u221e, which is applied perpendicular to the cylinder axis and parallel to the wall. We employ the standard thin-double-layer description, applicable when the diffuse-charge layers surrounding the cylinder and wall surface are thin compared with both a and a\u03b4",
+            " As the outer contribution to that force is at most O(1) we conclude that Fx = \u03b4 \u22121 [Fx0 + o(1)], (5.7) where the leading term is dominated by the inner contribution, namely Fx0 =\u2212 \u222b \u221e \u2212\u221e \u2202U0 \u2202Y \u2223\u2223\u2223\u2223 Y=H(X) dX. (5.8) Integration yields Fx0 = (1\u2212 \u03b3 )\u03c0, (5.9) in agreement with (3.1a). Similar arguments readily show that the hydrodynamic torque T is also O(\u03b4\u22121), T = \u03b4\u22121 [T0 + o(1)], (5.10) with T0 = Fx0, in agreement with (3.1b). We now consider the case where the cylinder is held in the vicinity of a reactive electrode, emitting a uniform current density j into liquid: see figure 1(b). Upon identifying E\u221e with j/\u03c3 (\u03c3 being the electric conductivity of the liquid) we may retain the preceding dimensionless notation. The electrostatic problem is again governed by Laplace\u2019s equation and the no-flux condition (2.1), but with the no-flux condition (2.2) replaced by the imposed-current condition, \u2202\u03d5 \u2202y =\u22121 at y= 0. (6.1) The far-field condition (2.3) is consistently replaced by \u03d5 \u223c\u2212y as |x|\u2192\u221e. (6.2) The flow-problem formulation remains valid provided the electro-osmotic condition (2"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0000510_htj.21229-Figure3-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0000510_htj.21229-Figure3-1.png",
+        "caption": "Fig. 3. Conceptual diagram of contact thermal resistance measurement.",
+        "texts": [
+            " The joint surface between the spindle and test bar is conical, while other surfaces are cylindrical. The test method for contact thermal resistance of a cylindrical surface is similar to a flat surface. For test methods, refer to Zhao and colleagues [14]. The radius of the conical joint surface is changing, so the heat transfer law on the surface is different. Therefore, compared with a cylindrical surface, the research on contact resistance of a conical surface is not mature. This paper designed the experiment system as shown in Fig. 3, to measure the contact thermal resistance of a conical surface. During the test, electric heating was configured in the outer circle of the spindle specimen, thermal insulation membrane outside the electric heating was added, as was coolant in the central hole of the test bar. By measuring the temperature field distribution of the cross section, we can get the contact position temperature change of both specimens, and thus get the contact thermal resistance of the cross section. Measuring the contact thermal resistance value of these three cross sections and taking the average value as the contact thermal resistance value of the conical joint surface"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0003547_0954406001523074-Figure4-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0003547_0954406001523074-Figure4-1.png",
+        "caption": "Fig. 4 Second-order ellipse pitch outlines driving third-order outlines. The base outlines shown here were arrived at by meeting the same requirements as those in Fig. 2 for \u00aerst-order ellipses, i.e., (a) PAC-E, (b) PAT-E, (c) PAC-N and (d) PAT-N",
+        "texts": [
+            " Put into more practical terms, there is no di erence in the outcome if a generating rack is set at some angle to the tangent patch, or at the complementary angle to the line of centres. However, there clearly is a di erence for non-circular gears. The PAT-E base outline shown in Fig. 2b permits useful involutes to be unrolled, as was done by Baxter. Note that the base outlines indeed appear to be confocal ellipses as reported by Olsson [1]. However, note also that the generation of involutes by unrolling a tangential cord to higher-order ellipses, as in Fig. 4b, leads to seemingly impossible situations. It may be seen from this \u00aegure that not only do the base outlines move in and out of the pitch outlines but they exhibit discontinuities. To understand the reason for these discontinuities, it is necessary, for example, to examine the curvature of the second-order ellipses shown in Fig. 4. Note that the curvature varies considerably around these outlines. Where the vertical axis cuts these ellipses, the outline is nearly straight. Gear teeth on a \u00afat surface constitute a rack, for which the base outline is in\u00aenitely far away. Now, where the horizontal axis cuts these second-order ellipses the centre of curvature can be judged to be between the outline and the centre of the ellipse. C02197 \u00df IMechE 2000 Proc Instn Mech Engrs Vol 214 Part C at OhioLink on November 7, 2014pic.sagepub",
+            " From the distribution of points, which may be calculated around these outlines, these discontinuities appear to be of the second order. There seems to be continuity in position, i.e. points can be calculated as close as desired to each other all around the base outlines, but there appears to be stepwise changes in gradient in the base outlines in Figs 4a and b. The methods used to generate involutes were based on two concepts. Both concepts were applied here to only one of the pitch outlines shown (the second-order ellipse in Fig. 4). The selection was made on the basis that this outline, while demonstrating signi\u00aecant di culties, might provide manageable challenges. A broader examination did not appear necessary at this stage. The \u00aerst concept applied was that of carrying out the numerical equivalent of tracing points on tangent cords as those cords are unrolled from the base lines. The results of the PAC-E outline (Fig. 4a) are not shown because its discontinuities gave rise to similar outcomes as the PAT-E outline (Fig. 4b). Figure 5 shows the results of tracing points on inextensible tangential cords, as they are unrolled from each of the remaining base outlines. The cords were necessarily on the convex side of the outline, at the tangent point. No reasonable results seemed to be produced by relaxing or reinterpreting the requirement that the cord be tangent at the point of contact. The involutes in Fig. 5a have discontinuities where the tangential cords generating them encounter discontinuities in the base outline",
+            " An algorithm that does was used for Fig. 7, shown later. By using a su ciently large pressure angle, it would have been possible to eliminate the excursions of the base outline. However, upward of an equivalent of a PA of 55 would have to be used to contain the base line in this example. Gear teeth with unrolled from a base outline, maintaining the cord tangent to the base line at the point of contact. The pitch and base outlines used are the same as those for the second-order ellipses shown in Fig. 4, except that the PAC-E base pro\u00aele (Fig. 4a) is not included because it would show similar features to PAT-E: (a) PAT-E, (b) PAC-N and (c) PAT-N Proc Instn Mech Engrs Vol 214 Part C C02197 \u00df IMechE 2000 at OhioLink on November 7, 2014pic.sagepub.comDownloaded from such pressure angles would create relatively large normal forces at their point of contact and large bearing forces, with associated large frictional losses. There is also the possibility that the teeth will ultimately not slide (at locations away from the pitch points) but wedge and lock the gears together",
+            " 7, the angle of rotation at which the pitch outlines are shown, the conjugate line, appears to meet both the tangent and normal conditions (E and N). This condition occurs twice per cycle, and Fig. 7 shows just one of them. Figure 8 is similar to Fig. 7, except that here a pressure angle PAT-N of 45 was used. That is, the line of action made a \u00aexed angle of 45 to the tangent at the down the conjugate line from one base outline to the other. The points are traced on rotating coordinate systems \u00aexed in each gear centre. The PAC-N line of action and base outline of the second-order ellipse as in Fig. 4c are used. (b) Enlargement of the region about the pitch point, showing that normal conjugate action is taking place between the opposite-handed involutes of action and base outlines of the same second-order ellipse shown above is used. (b) Enlargement of the region about the pitch point, showing that normal conjugate action is taking place between the oppositehanded involutes Proc Instn Mech Engrs Vol 214 Part C C02197 \u00df IMechE 2000 at OhioLink on November 7, 2014pic.sagepub.comDownloaded from contact point"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0000860_978-3-319-44735-3_1-Figure1.23-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0000860_978-3-319-44735-3_1-Figure1.23-1.png",
+        "caption": "Fig. 1.23 Input linguistic variables and their membership functions for: Yaw angle error",
+        "texts": [
+            " The yaw controller (or heading goal controller) is responsible for pointing the hexacopter front to the target position, keeping this position until it arrives at the destination. Figure1.22 shows the block diagram of this controller. To allow this controller to perform such a task, two input variables are needed. The yaw angle is similar to roll and pitch angles, and hence, are calculated based on the Euler approximation of Z-axis angle between the current angle position of the hexacopter front and the target position. Figure1.23 shows the linguistic variable and the membership function values of yaw angle. The distance to the goal is the second input data utilized. It indicates how far the hexacopter from the target position. The goal distance is calculated fromX- and Y-axis positions of the hexacopter and results in a polar coordinate indicating the angle and the distance to the target point. Goal distance linguist variable is depicted in Fig. 1.24. This fuzzy controller has 28 rules as shown in Table1.3. These rules define the value of the output variable omega yaw (OYaw), presented in Fig"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0000188_0954406214560420-Figure6-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0000188_0954406214560420-Figure6-1.png",
+        "caption": "Figure 6. Equivalent schemes of a PR limb.",
+        "texts": [
+            " The generalized coordinate corresponding to the target position is q0m \u00bc \u00f0\u20192 lap\u00de T, where lap is the length of the second link, a fixed value. Also, the virtual P joint can be added to the first link, just as Figure 5(b) shows. The generalized coordinates of the equivalent mechanism is q0 \u00bc \u00f0\u20191e le\u00de T. The target position of the equivalent mechanism is q0m \u00bc \u00f0\u20191 loa\u00de T, and \u20192 is used to keep the orientation of the second link with respect to the first link. Just including P joints is not sufficient to remove the extreme displacement singularity. A PR limb is shown in Figure 6. Although there is a P joint, its workspace is a ribbon area, whose width is two Figure 4. Coupled motion of a rigid body. at Univ Politecnica Madrid on January 14, 2015pic.sagepub.comDownloaded from times of the link\u2019s length. In Figure 6(a), the PR limb is converted to the RP limb. In Figure 6(b), the PR limb is transformed to the PP limb. The generalized coordinates of the equivalent PP limb is q0 \u00bc \u00f0d1e le\u00de T, and the target position is q0m \u00bc \u00f0d1 0\u00deT. Figure 7 shows the equivalence of an RRR and a PRR limb. The equivalent process is similar to the above ones. Through the above analysis, the idea for removing the extreme displacement singularity of the limb can be summarized as follows: (1) fixing the genuine joint inputs; (2) converting some fixed parameter into new joint variables, with their current values as the target joint values; and (3) applying NR or QN method to find the pose to match the target joint values"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0003827_robot.2000.846348-Figure1-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0003827_robot.2000.846348-Figure1-1.png",
+        "caption": "Fig. 1: A cylindrical and an ellipsoidal objects on table",
+        "texts": [
+            " Since an equilibrium grasp can impose weaker constraints on the objects than a force-closure grasp, we can expect to realize it with fingers less than the number required for the force closure grasp. Based on this consideration, we focus on an equilibrium grasp aparting from a force-closure grasp in this work. Definition of the equilibrium grasp is that the sum of all forces and moments acting on each object balance within the hand. Now, let us introduce an interesting behavior for both a cylindrical and an ellipsoidal objects. Suppose that both are placed on a table as shown in Fig.1. Also, suppose that the table is a bit inclined. For such a small inclination of the table, both objects will lose the initial equilibrium state and start rolling motions. Once the cylindrical object starts a rolling motion, it never stops and finally falls down from the edge of the table. Under the same situation, the ellipsoidal object will also start a rolling motion. However, it soon stops and results in another equilibrium state, while it may not stop for a larger inclination angle of the table"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0003722_s0921-8890(00)00130-5-Figure1-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0003722_s0921-8890(00)00130-5-Figure1-1.png",
+        "caption": "Fig. 1. (a) Spring Flamingo: the 7-link planar bipedal robot whose hip, knee and ankle rotate about parallel axes. The axes are perpendicular to the XZ plane. (b) A simulation model of Spring Flamingo is generated for the simulation study in this paper.",
+        "texts": [
+            " In recent years, the MIT Leg Laboratory proposed a control language called virtual model control (VMC) [22] for legged locomotion. In the VMC approach, virtual components (e.g., springs and dampers) are used to specify actuator forces. The actuators are driven so that the robot, to the extent possible, behaves as if the virtual components were real. Virtual components\u2019 endpoints may be attached between parts of the robot, as well as between the robot and the environment. In the control of a planar biped (Fig. 1) for steady dynamic walking, Pratt et al. [22] implemented an algorithm called \u201cTurkey Walking\u201d 1 using the VMC 1 The label \u201cTurkey Walking\u201d is used because the algorithm was first applied to the dynamic walking task of a planar biped called \u201cSpring Turkey\u201d. approach. In the algorithm, they used a virtual parallel spring-damper component in the vertical direction to control the height, a virtual rotational spring-damper component for the pitch angle control of the body, and a virtual damper for the horizontal velocity control of the biped (Fig",
+            " A simple model is chosen for RAC and the resulting control algorithm preserves the low computation of the Turkey Walking algorithm. All the implementations are carried out in a simulation environment using a simulated biped that is modeled after a physical biped. This paper demonstrates that the modified algorithm is robust in that it enables the simulated biped to adapt to load variations of up to 20% of the body mass even while it is walking blindly over rolling terrain. The control algorithm presented in this paper is applied to a simulated 6-d.o.f. 7-link planar biped (Fig. 1). The biped is constrained to move in the sagittal plane. The biped has two slim legs and a body. The legs (each weigh about 1 kg) are much lighter than the body (10 kg). Each leg consists of a hip joint, a knee joint and an ankle joint. All the joints are revolute pin joints with axes perpendicular to the sagittal plane. The hip and knee joints are actuated, whereas the ankle joint is limp. The heel and toe of each foot have discrete sensors to detect ground contact. The detailed parameters of this model can be found in Appendix A",
+            " Since one of the original motivations for using the VMC approach is its low computation, it is desirable to maintain such a feature when incorporating an adaptive behavior into the biped. The walking task in the sagittal plane can be subdivided into three subtasks as follows: 1. body height control; 2. body pitch control; 3. horizontal velocity control. Let p = [x, z, \u03b1]T be the generalized coordinates of the body frame Oxz with respect to an inertia frame OXZ attached to the hind leg\u2019s ankle (during the double support phase) or stance leg\u2019s ankle (during the single support phase) (see Fig. 1). The body height, body pitch, and horizontal velocity controls correspond to keeping z, \u03b1 and x\u0307 close to the desired value zd, \u03b1d and x\u0307d, respectively. In the Turkey Walking algorithm, the body height is controlled by a virtual spring-damper. The body pitch control is controlled by a virtual rotational spring-damper. The horizontal velocity control is con- 3 A singular configuration is one in which the upper and lower links of the legs are completely aligned with each other. trolled by a virtual damper"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0000472_s00170-016-9136-9-Figure1-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0000472_s00170-016-9136-9-Figure1-1.png",
+        "caption": "Fig. 1 The calculation model for tensile stresses and movements caused by these stresses at an arbitrary point located inside the contact area",
+        "texts": [
+            " 8, 410003 Saratov, Russia 3 Department of Computer Aided Manufacturing Technologies, Technical University of Ko\u0161ice, Faculty of Manufacturing Technologies with a seat in Presov, \u0160t\u00farova 31, 080 01 Pre\u0161ov, Slovak Republic The specific work of tensile stresses under the rolling of elastic bodies at the arbitrary pointM can be determined from the following formula: dAm \u00bc 1 2 d\u03c3m\u22c5d\u03b4m; \u00f01\u00de where: Am is the specific work of tensile stresses under the rolling of elastic bodies at the arbitrary point M. \u03c3m is the tensile stress at the point M, MPa. \u03b4m is the material deformation at the point M, caused by tensile strains, mm [23\u201326]. The calculation model is shown in Fig. 1. Under the influence of the external load, the spherical rolling body is pressed against the rolling raceway, which results in the elliptical contact area having a minor semiaxis a and a major semiaxis b occurring between the rolling body and the raceway [27\u201329]. Let us create the Cartesian reference system centred in the middle of the elliptical contact area. The axis OX is directed along the minor semiaxis of the contact area, whereas the axis OYis directed along the major semiaxis of the contact area"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0003681_robot.1991.131809-Figure2-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0003681_robot.1991.131809-Figure2-1.png",
+        "caption": "Figure 2: 2-DOF Hopping Robot Cycle",
+        "texts": [
+            " The leg actuator and energy storage mechanism is an air cylinder with restoring force proportional to the inverse of the piston displacement, like the 1-DOF model. The leg and cylinder combination are assumed to be massless. We will use a polar coordinate system when the robot is in ground contact and a Cartesian system for ballistic motion. Let T denote the leg length and e denote the \u201chip angle\u201d during stance (Figure 1). The motion of the robot may be decomposed into four distinct phases : thrust, decompression, flight and compression (Figure 2). i. Thrust Phase. At time t j the leg is a t its minimum radial length, ~j = ~ ( t j ) (i.e., +( t i ) = 0 ), the control valves are opened and a constant supply pressure CH2969-4/91/0000/1392$01 .OO 0 1991 IEEE 11. ... 111 is connected to the leg cylinder for a fixed time, 6 t . This applies a constant radial force, ?, to the mass. The equations of motion in this phase are: m(i: - r e 2 ) - ? + mg cos(@) = 0 for t j 5 t 5 t j + St m(rli + 2 4 - mgsin(8) = o Decompression Phase. At the end of the thrust phase the valves are closed and the cylinder decompresses"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0002923_978-94-011-3514-6-Figure13.1-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0002923_978-94-011-3514-6-Figure13.1-1.png",
+        "caption": "Fig. 13.1 Point mass attached to resistances",
+        "texts": [
+            " Nevertheless, when all the coordinates x i are lengths and their number is less than or equal to three, the space defined is easily interpretable; it is either a straight line, a plane or a euclidian space depending on whether respectively. n has the value one, two or three Given the objective of simplifying as much as possible the mathematical formulation and presentation of the modes, we have chosen a system consisting only of a point mass which moves in a plane and which is kept in its equilibrium position by r springs of stiffnesses Moreover, t viscous resistances with constants Cj act on the mass. Such a system has two degrees of freedom (figure 13.1). r springs and t viscous - 265 - The lines of action of the springs are specified by the angles ai, and those of the resistances by the angles Cj . So as to retain the linearity of the problem, we assume that the elastic and dissipative elements are sufficiently long in order that the angles ai and F:j can be considered as constants at the time of the displacement of the mass. The kinetic energy has the elementary form The mass matrix which follows from it is then simply [M] m tI,j Fig. 13.2 spring ki placement Extension of the tor the dis- 00\" of the mass X, X, (13",
+            " tOO + ,'(If lL 'I\" 1, 21 po wp which correspond to the two angles B~ and B~ respectively, which define the directions of the principal axes of the trajectory. These angles, which we shall call the pr inc ipal directions in what follow~, are d~terruined by their tangents X2 I x, t t' P ttl P The principal directions of an actual natural mode are not thpm selves orthogonal, as we shall see in the numetical example dealt with below; nor are they orthogonal to the principal di rections of the other natural mode pither. (13.25 ) (13.26 ) (13.27) - 272 - So as to allow the comparison of certain results, let us go back to the system of figure 13.1, but assume that it is without viscous resistances. The matrices [M] and [K] remain unchanged while. t.he damping matrix is exactly zero. The eigenvalues are purely imaginary and the eigenvectors are real; consequently Ap = 0 and tp = 0 In order to isolate a morle, ld us retain the initial conditions (13.15), that is to say X1 (0) Xo and o X2 (0) ~p Xo and o The two initial velocities being zero, it is a matter this time of a simple relea.se of the mass, from a specific point in the plane X1 X2 ",
+            "1035) e- 2 ,35l cos(22.62 t + 0.4208) The trajectory described by the mass is an elliptical spiral shown in figure 13 .1. (13.37 ) - 276 - for the second mode. 1\\2 .Llli IP2 = Arc tg 0.2592 rr/2 - 1.3116 W2 21. 93 _1_ 1.0346 cos IP2 x '(0) Xo Xl (0) 0 x2 (0) 0.2171 Xo X2 (0) - 3.688 Xo Th\" muddl directions have the values e' - 12.55 0 2 8; ~ 72.88 0 The equations governing the second isolated mode are as follows Fig. 13.4 Trajectory of the mass corresponding to the ~econd complex mode of the system of figure 13.1 0,5 - 277 - When the resistances Cj are zero, the system is reduced to [M] x + .. [K] x .. o The mass and stiffness matrices remain defined by the relations ( 13.2) and (13.6) respectively. With the numerical values chosen, one obtains the following eigenvalues r W1 22.18 52 W2 23.27 53 - j W1 - j 22.18 5, W2 23.27 Since t.he modes are real, the changes of bn~i s matrix order 2 n , is written 22.18 e- jrr / 2 23.27 e- jrr / 2 22.18 e jrr / 2 23.27 11.78 e- jrr / 2 43.82 e jrr / 2 11 .78 e jrr/2 41",
+            " As for the principle directions of the product [MJ-1[Cl they have the values 81 = 72,79\u00b0, 17,21 \u00b0 in the numerical example ChU;i\"Il. The existence of d d~mping not respecting the Caughey condition leads to two consequences; on one hand the rectilinear trajectories are transformed into elliptical spirals, and on th,' other hand the principal directions of HI'> \"<\">Ir\\plex modes are. situated betw~~n ~he principal directions of [Ml l[Kl and those of [Ml-1 [Cl . They be,:ome closer to the la Her ao the damping becomes more important (figure 13.6). Fig. 1J.6 Principal directions for the system of figure 13.1 complex modes [M] 1 [Kl [MI-1[Cl x, Therefore, th~ Caughey condition clearly expresses the congruence of the principal directions of [Ml-llKl with those of [Ml-1[C]. - 281 - CHAPTER 14 FORCED STATE OF THE GENERALIZED OSCILLATOR 14.1 Introduction For rnechanical systerns having rnany deyre<es of freedorn, it is illusory to expect to undertake a systernatic analyt.ical study of the responses to the different types of excitation - that is to say, of the forced state - as we have done for the elementary oscillator, because of the very great diversjty of envisageable circumstances"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0000364_intelse.2016.7475150-Figure3-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0000364_intelse.2016.7475150-Figure3-1.png",
+        "caption": "Fig. 3.",
+        "texts": [
+            " (8) 9 As the intercept course angle the difference of course angle and the desired course angle i.e., XE = X - XR ' we can have: V (. + . ) t an \u00a2cos(x -1j;) = 9 XE XR. (9) 9 Solving (9) for XE we have . gtan \u00a2 . XE = -V- cos(X -1j;) - XR , (10) 9 The above state equation using (5) takes the following form . _ 9 tan \u00a2 ( . -1 ( - W n sin X W e cos X) ) . XE - -V--cos sm + - X 9 Va Va R ' (11) here XR is the rate of change in reference course angle. This term is non zero for following circular arcs but for straight flight XR = O. Cross track error deviation y is the closet distance to path Fig. 3 and can be its rate of change can be derived using the relation (12) The closed loop dynamics from \u00a2reJ to \u00a2 of the autopilot control loop is approximated using a first order filter [21]: \u00a2 1 \u00a2reJ (13) TS + 1 and incorporated into guidance design. Equations (11), (12) and (l3) represent the overall dynamics for the outer loop guidance design problem with y, XE and \u00a2 as state variables and to avert the cross lateral error guidance loop will gen erate \u00a2reJ as the control signal. Previously we showed that coordinated turn condition can be described by X = ~ tan \u00a2 cos(X -1j;) (14) 9 It can also be expressed in terms of heading and the airspeed as in [21]"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0003093_s1385-8947(01)00305-9-Figure1-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0003093_s1385-8947(01)00305-9-Figure1-1.png",
+        "caption": "Fig. 1. Apparatus setup.",
+        "texts": [
+            " (7) and (8) CD drag coefficient d sphere diameter (m) g acceleration due to gravity (m/s2) m consistency index (N sn /m2) n flow index Re Reynolds number Rec critical Reynolds number Ret terminal Reynolds number Ut terminal velocity (m/s) V sphere volume (cm3) Greek letters \u03c1 liquid\u2013solid density difference (kg/m3) \u03c1l liquid density (kg/m3) \u03c1p particle density (kg/m3) The purpose of this study is to determine the hydrodynamics of rising solid spheres at high Reynolds up to super-critical conditions in non-Newtonian (pseudoplastic) liquids. It will be established if the constant drag coefficient of 0.95 determined by Dewsbury et al. [6] is valid for Ret > 8500. The main goal is to propose a drag curve of free rise of solid spheres in pseudoplastic liquids at very wide range of Reynolds numbers. Fig. 1 shows the apparatus used in this research. The square column was constructed with transparent acrylic and had a height of 120 cm and internal length and width of 73 cm. The large width of the column ensured there would be no wall effects affecting our results [8,11]. To determine the terminal velocity of the rising spheres, a high resolution CCD video camera (Hitachi, Japan, model KP-M 1U) along with a graphics software program (SigmaScan Pro 4.0, SPSS, USA) were used with the same technique described earlier by Dewsbury et al"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0002923_978-94-011-3514-6-Figure8.4-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0002923_978-94-011-3514-6-Figure8.4-1.png",
+        "caption": "Fig. 8.4 Natural modes of the system consisting of two equal masses, symmetrically placed on a massless beam, oscillating by bending in a plane",
+        "texts": [],
+        "surrounding_texts": [
+            "~2 has the values ~21 and ~22 respectively ~22 The system of equations (8.6) consequently has the following specific solutions for Xl All e jlU1t All e- jlU1t A12 e jlU2t A12 e- jlU2t for X2 A21 e jlU1t A2l e- jlUlt A22 e jw2t A22 e- jlU2t The general solutions are obtained by linear combinations of the specific solutions (8.10) (8.11 ) (8.12 ) - 142 - The constants C 1 , Dl, C2 and D2 are arbitrary; moreover, since the Aij are only defined to a scale factor, the solutions can be put in the form { Xl X2 or in the matrix form with .. x 1st mode 2nd mode The above equations include 4 constants of integration, Xl, X2 , 'p 1 , ~2 , which are functions of the initial conditions. They introduce the concept of natural modes, to which we will return in detail in the study of the generalized oscillator. A natural .ade is the motion of a syst_ about a natural angular frequency (or at a natural frequency which comes to the same thing). We have assumed that the system examined does not include resistances. The natural angular frequencies Wl and W2 are thus equivalent to the angular frequency WO of the conservative elementary oscillator studied in section 3.1. It would have been more consistent to have adopted the notation WOl and W02 instead of and W2. However, as no confusion is possible in this chapter, we prefer the shorter notation. A system with two degrees of freedom thus possesses two natural modes. By choosing the initial conditions in a way that X2 = 0 , the system oscillates according to the first mode only. It oscillates according to the second mode if Xl = 0 . In the general case, the two modes exist together, but do not have mutual influence, in other words there is no exchange of energy from (8.13) (8.14 ) (8.15) - 143 - one to the other. This important property, called orthogonality of the natural modes, will be established later. It expresses the linear independance of the natural vectors ~ and is translated here by the two relations .. ~T ~T [M] ~2 0 [K] ~2 0 1 being, after development { m1 + m2 ~ 21 ~22 0 k1 + k2 ~21 ~22 + k3(1-~21)(1-~22) 0 The role of the natural modes appears clearly when the system has a geometric symmetry (figure 8.2). Tn this case and the characteristic equation (8.9) becomes 2(k + k3) k2 + 2 k k3 p4 + p2 + 0 m m2 It has the solutions p2 k + k3 13 m m The natural angular frequencies are thus = k k + 2 k3 .. 2 UJ~ 1 m m The ratios ~21 and ~22 of relation (8.11) then take the valu~s +1 and - 1 In fact ~21 + 1 - 1 (8.16 ) (8.17 ) (8.18 ) (8.19 ) (8.20) (8.21 ) - 144 - The first mode corresponds to identical oscillations of the two masses X2 Xl COS(Wlt - 'Ill) whereas the second mode corresponds to the oscillations being 160\u00b7 out of phase { Xl = X2 = - X2 COS(W2 t - 'Il2) Three examples of symmetrical oscillators with two degrees of freedom are shown in figures 6.2 to 6.4. (0) system in static equilibrium (1) 1st mode identical displacements of the 2 masses X2 = Xl (2) 2nd mode X2 = - xl opposed displacement of the 2 masses (2) 2nd mode: X2 - Xl (6.22) (6.23) -145 - (1) 1st mode X2 x, (2) 2nd mode : X2 - X, One can make several comments concerning these examples The static configuration corresponding to a natural mode is called a mode shape of the system. This concept will be generalized in what follows. Thus, the first and second mode shapes are shown in each of the figures. For the reference system and for the double pendulum, the elastic coupling does not participate in the first mode because of the symmetry, whereas it does in the second mode. This statement cannot be applied directly to the last example of the bending of a beam. However, one should note that the flIst mode shape has a single curved portion whereas the second has two, separated by an inflexion point. For an equal value of the displacement (for example the normalized value Ix, I = 1 ), the potential energy of the first mode is less than that of the second mode. It consists of the energy of deformation in the three examples together with, in the case of the double pendulum, the potential energy of position of the masses. - 146 -"
+        ]
+    },
+    {
+        "image_filename": "designv11_60_0002762_s0094-5765(98)00040-x-Figure11-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0002762_s0094-5765(98)00040-x-Figure11-1.png",
+        "caption": "Fig. 11. Manipulator model and the desired trajectory.",
+        "texts": [
+            " The maneuver requires the end-e ector to move 10 m in the y direction in 10 s, and the required velocity of the end-e ector in the y direction is taken as a fourth order polynomial function in time. The solid lines in Figs 7\u00b110 show the controlled response of the \u00afexible manipulator, while the dashed\u00b1dotted lines present time histories of the desired velocity, acceleration, the desired deployed length, and slew angle. It is apparent that the controller is remarkably successful in following the speci\u00aeed trajectory, even in the presence of \u00afexibility. 6.2.2. Case 2. Consider the manipulator with four links in free-\u00afying condition (i.e. g = 0). Figure 11 shows the manipulator model and its initial con\u00aeguration. The desired trajectory of the end-e ector is speci\u00aeed, as before, to be a straight line in the y direction. The maneuver requires the end-e ector to move 0.5 m in 2 s, and the desired acceleration is de\u00aened by the sine function with the period of 2 s. The mass and sti ness of the link 4 are m4=0.3172 kg and E4I4=0.006567 Nm2, respectively. The mass and sti ness of the other links are expressed as follows: the mass ratios from link 1 to 4 are 100:20:2:1; and the ratios of the sti ness are 1000:800:40:1"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0002859_ip-epa:19982165-Figure6-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0002859_ip-epa:19982165-Figure6-1.png",
+        "caption": "Fig. 6 Wound ring stuck",
+        "texts": [
+            " For this purpose, rings were stamped out, with an outer diameter of 163\u201c and an inner diameter of 90\u201d. This gave an annulus thickness of 36mm, which corresponds to the \u2018back iron\u2019 of quite a large machine. These \u2018ring stacks\u2019 or simulated motor stators were assembled and wound with magnetising and B sensing conductors. In this form, the power loss of the ring could be examined accurately. Ring stacks were treated in a variety of ways; annealed, not annealed, welded as in a motor stack, hydraulically squeezed to various pressures etc. In all, some 500 ring stacks have been evaluated. Fig. 6 shows a typical ring. Fig. 7 shows a device for squeezing ten ring stacks at the same time. Copper inserts allow stress equalisation and makes seam welding easier. IEE Proc.-Electr. Power Appl., Vol. 145, No 5, September 1998 late \u2018perfect\u2019 insulation, and in the other case using bare steel. The differences are very small, see Fig. 10. copper inserts 3 welded seams I 10 stacks of 20 rings 0\u20195 I u I I U hydraulic jack Fig. 7 Hydraulic pressurising apparatus for welding the laminae 4 Results of ring stack tests The number of results reported here will be limited to show the main features encountered"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0000842_wcicss.2015.7420318-Figure1-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0000842_wcicss.2015.7420318-Figure1-1.png",
+        "caption": "Figure 1. Schematic diagram of two-link manipulator.",
+        "texts": [
+            " The purpose of the present paper is to develop a design method of the combined MPC-DOB systems where the total control input, which is composed of the DOB output and MPC output, can be constrained. To this end, we introduce a time-varying input constraints into design, and constraints on the control signals generated by the MPC controller are properly adjusted at every sampling instant so that the total control input may satisfy the given constraints. II. Experimental results The schematic diagram of the two-link planar manipulator used in the following experiments is shown in Fig. 1. Also, the specifications of the manipulator is shown in Table I. We suppose that the amplitude of the total control input should be bounded within plus or minus 9 V in the experiments. We also assume that the difference of the total control input should be bounded within plus or minus 10 V. The prediction horizon Np and the control horizon Nc are respectively specified as Np = 20 and Nc = 2. Fig. 2, Fig. 3 and Fig. 4 respectively show the experimental results by the proposed DOB-based MPC, by the conventional MPC and by the conventional DOB-based MPC"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0000323_j.ifacol.2016.03.043-Figure2-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0000323_j.ifacol.2016.03.043-Figure2-1.png",
+        "caption": "Fig. 2. Goal pose 1: \u03b81 = 30o, \u03b82 = 130o, \u03b83 = \u2212110o; goal pose 2: \u03b81 = 10o, \u03b82 = \u2212120o, \u03b83 = 100o.",
+        "texts": [
+            " For the present study, in all cases, the start manipulator pose is taken as the stowed pose \u03b81 = 170o, \u03b82 = \u2212160o, \u03b83 = 160o, although it is conceivable that depending on the final relative pose of the chaser and target, it would be useful to bring the manipulator out of the stowed pose to a better starting pose, for more effective grabbing. For the present study, we selected some of the grid points as the goal poses, after checking whether the target placed with grab point appropriately positioned, is not interfering with manipulator or with chaser. A couple of such goal poses are shown in Fig. 2. It should be noted that the target pose may not correspond to a pre-specified grid point exactly. We can handle this by generating grids based on start and goal poses, or use a fixed master grid to which start and goal poses can be appended as intermediate points. Each grid point is considered as a node of the graph. In the three dimensional grid, an inside grid point has a maximum of 6 neighbours when only rectangular moves are allowed, and has 26 neighbours, when diagonal moves are also allowed"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0002252_s0890-6955(98)00006-6-Figure3-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0002252_s0890-6955(98)00006-6-Figure3-1.png",
+        "caption": "Fig. 3. Model of the chamfered main cutting edge nose radius tool when wear occurs, f . R, (R\u00de0).",
+        "texts": [
+            " [17], calculation of cutting force components could be one approach to confirm the wear behaviour of lathe tools during the cutting process. According to the last section (Section 2.1), the calculation of shear area A and the projected area Q for nose radius tools with a chamfered main cutting edge, when wear has occurred, can also be divided into following three categories. 1. Sharpness of the tool (R = 0) with wear, as shown in reference [15]. 2. Wear nose radius of the tool (R) is smaller than the feedrate (f), (R\u00de0,R , f), as shown in Fig. 3, and the shear plane A includes the area A = A1 + A2 + A3 + A4 + A5 + As 3. Wear nose radius of the tool (R) is larger than feedrate (f), (R\u00de0 R . f) according to the depth of cutting, which can be subdivided into three parts (a) d . R, (b) d = R and (c) d , R. According to the above conditions, case (1) had been discussed in reference [15]. This paper will therefore focus on the case of small radius, i.e. case (2). However, an experiment has been performed to study the case of large nose radius cutting, i"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0001235_b978-0-08-099425-3.00005-4-Figure5.10-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0001235_b978-0-08-099425-3.00005-4-Figure5.10-1.png",
+        "caption": "FIGURE 5.10",
+        "texts": [
+            " All these models are essentially empirical, in that they fit some mathematical function to observed behaviour. For real tyres the friction generated between the tread rubber and the road surface is generated through two mechanisms, these being hysteresis and adhesion. As already discussed in Chapter 3, rubber displays substantial hysteresis. In order to understand the influence of hysteresis on tyre/road friction, consider a block of rubber subjected to an increasing and then a decreasing load as shown in Figure 5.10. As the rubber is loaded and unloaded it can be seen that, for a given displacement d, the force F is greater during the loading phase than the unloading phase. Hysteresis in rubber. Loading and unloading of a single tread block of tyre rubber in the contact patch. If we consider the situation where the same block of rubber is sliding over a nonsmooth surface, it can be seen from Figure 5.11 that an element of rubber in the contact patch will be subject to continuous compressive loading and unloading"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0000804_jahs.61.042006-Figure4-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0000804_jahs.61.042006-Figure4-1.png",
+        "caption": "Fig. 4. Contact centroid, for 1 < cr < 2.",
+        "texts": [
+            " Afterward, the second tooth pair continues in engagement alone during the interval from the time B to time C. The cycle continues in the similar pattern from the second pair to the third pair. Figure 3 also illustrates the phase differences between meshing pairs. Contact centroid To simplify the following structural dynamics model, an equivalent contact point is needed to combine all effects of multiple contact points when the contact ratio is greater than one. This equivalent contact point is called \u201ccontact centroid\u201d and illustrated in Fig. 4. The contact centroid position (rc, \u03b8c) is given in space fixed frame (r, \u03b8 ) as rc(t) = \u2211ceil(cr) i=1 [ri(t)li(t)]\u2211ceil(cr) i=1 [li(t)] (2) \u03b8c(t) = \u2211ceil(cr) i=1 [\u03b8i(t)li(t)]\u2211ceil(cr) i=1 [li(t)] 042006-3 where (ri , \u03b8i) is the contact point position of the ith meshing pair, li is the corresponding contact line length, cr is contact ratio, and ceil(\u2022) is the function that rounds a number upward toward its nearest integer. An assumption made here is the longer contact line bears more load, thus the contact line length is used as a weight parameter to calculate the contact centroid position"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0002814_bf01128176-Figure3-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0002814_bf01128176-Figure3-1.png",
+        "caption": "Fig. 3. Schematic drawing illustrating the alternating bright and aark stripes imaged in planes above and below the sample plane of symmetry for a nematic solution of PBT subjected to a magnetic field as viewed with incident light polarized along n o (see text).",
+        "texts": [
+            " A more complicated behavior has been reported for PBT solutions under similar conditionsJ 2~ In the latter case, incident light polarized along the original director produces alternating bright and dark bands parallel to H focused in planes a distance h above and below the center plane of the slab when the transmitted light is viewed without an analyzer (h may be smaller or greater than half the slab thickness); bright bands in the upper plane are above dark bands in the lower plane, and vice versa (see Fig. 3t. The patterns are not observed with light polarized along H. In this case, A is initially independent of the field strength, by contrast to the behavior for 2 discussed above, which varies in a complex way with H,/Hc. ~4~ The separation of the two focal planes depends on both H and the time in the field. Although A is about independent of the time in the field initially, after some time, the interval suddenly decreases by half. These observations are attributed to cylindrical lenses formed by tilting of the director out of the plane of the slab as the reorientation proceeds, with the axes of the lenses parallel to H and spaced at more or less common interval A/2 along the original director"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0001128_icrom.2016.7886810-Figure2-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0001128_icrom.2016.7886810-Figure2-1.png",
+        "caption": "Fig. 2. An omni-directional wheel with roller in contact with the ground.",
+        "texts": [
+            " The robot considered here has omni-directional wheels which have passive rollers attached to the rim of the wheel hub. The rollers rotate about their axis which in turn rotates with the wheel hub. Also, as the rollers\u2019 axes are at an angle to the respective wheel axes, this can provide motion in the transverse direction. Therefore, the platform has the ability to move in transverse direction without having to make a turn, unlike a mobile robot with conventional wheels. An omni-directional wheel is shown in the Fig. 2. The kinematic constraints for this wheel is shown below 0i i i i iR r\u03b8 \u03c6 \u2032+ + =v f f  (1) where iv is the velocity vector of the ith wheel center, R is radius of the wheel hub, r is the radius of the roller, i\u03b8 is 978-1-5090-3222-8/16/$31.00 \u00a92016 IEEE 595 the rotational speed of the wheel, i\u03c6 is the rotational speed of the roller in contact with the ground, if and i\u2032f are the unit vectors along the tangential directions of the wheel and the roller, respectively. As can be seen from (1), the wheel rotational velocity can be written in terms of the velocity vector of its center"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0000984_j.ifacol.2016.10.148-Figure2-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0000984_j.ifacol.2016.10.148-Figure2-1.png",
+        "caption": "Fig. 2. Free-body diagram of the forces acting on the UAV in the climbing coordinated turn.",
+        "texts": [
+            " (in print, 2016), the filter gain \u0393 is now a positive definite matrix \u0393 =   \u039311 0 \u00b7 \u00b7 \u00b7 ... . . . 0 0 \u00b7 \u00b7 \u00b7 \u0393mm   (6) where m is the number of sensors. Here m = 7 which accounts for the leader UAV (numbered as 4) and the six follower UAVs in the 3D rigid formation (cross-shape) of Figure 3. Please note that a full gain matrix \u0393 would require information exchange between all UAVs whereas a diagonal gain matrix \u0393 significantly reduces the communication requirements at the possible expense of estimator performance. Figure 2 shows the free-body diagram for a UAV climbing at a flight path angle \u03b3 and a bank angle \u03c6. The derivation of the UAV equations of motion are based on the pointmass model of a fixed-wing aircraft Beard and McLain (2012); Zhao and Tsiotras (2010); Menon et al. (2012). Choosing the coordinate axes X,Y and Z to be directed east, north, and towards the earth center respectively, the equations of motion are given by X\u0307 = Vg cos \u03b3 cos\u03c7 Y\u0307 = Vg cos \u03b3 sin\u03c7 Z\u0307 = \u2212Vg sin \u03b3 V\u0307g = 1 M (T \u2212D(CL)\u2212Mg sin \u03b3) \u03b3\u0307 = 1 MVg (L(CL) cos\u03c6\u2212Mg cos \u03b3) \u03c7\u0307 = L(CL) sin\u03c6 cos(\u03c7\u2212 \u03c8) MVg cos \u03b3 (7) IFAC NOLCOS 2016 August 23-25, 2016",
+            " Using condition (13), the desired ground speed magnitude V d g , the flight path angle \u03b3d, and the course angle \u03c7d for the rigid flying formation are calculated using the first three equations of (7) V d g = \u221a (X\u0307d)2 + (Y\u0307 d)2 + (Z\u0307d)2, \u03b3d = arcsin ( \u2212 Z\u0307d V d g ) , \u03c7d = atan2 ( Y\u0307 d, X\u0307d ) . (15) Using (15), the desired rates of change of V d g , \u03b3 d and \u03c7d are determined based on the first order models V\u0307 d g = bVg (V d g \u2212 Vg), \u03b3\u0307d = b\u03b3(\u03b3 d \u2212 \u03b3), \u03c7\u0307d = b\u03c7(\u03c7 d \u2212 \u03c7), (16) IFAC NOLCOS 2016 August 23-25, 2016. Monterey, California, USA Michael A. Demetriou et al. / IFAC-PapersOnLine 49-18 (2016) 110\u2013115 113 Fig. 2. Free-body diagram of the forces acting on the UAV in the climbing coordinated turn. where M is the mass of the UAV, Vg is the ground speed, \u03c8 is the heading angle, \u03c7 is the course angle and T is the thrust. The forces L(CL) and D(CL) are the lift and drag forces, respectively, and are given by L(CL) = 1 2 \u03c1V 2 a SCL, D(CL) = 1 2 \u03c1V 2 a SCD(CL), (8) where Va is the airspeed, \u03c1 is air density, S is the planform area of the UAV wing, CL and CD are the lift and drag coefficients which are related via CD = CDp + C2 L \u03c0eAR , (9) where CDp is the parasitic drag caused by the shear stress of air moving over the wing, b is the wing span, AR b2 S is the wing aspect ratio, and e is the Oswald efficiency factor Beard and McLain (2012)"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0001847_pime_proc_1994_208_361_02-Figure1-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0001847_pime_proc_1994_208_361_02-Figure1-1.png",
+        "caption": "Fig. 1 Porous bearing geometry, notations and the coordinate system used with an ideal journal",
+        "texts": [
+            " In the theoretical analysis of porous bearings (2-8) the journal has been assumed to be ideal, that is perfectly cylindrical. Since all machined journals have some geometric irregularities, a study is necessary to know their role on the performance of bearings. In the present analysis geometric irregularities of a journal, such as circumferential undulations and barrkl/bellmouth shapes, are taken into account. Similar studies for solid journal bearings have already been reported (9-1 1). 2 ANALYSIS Figure 1 shows the porous bearing geometry with the coordinate system and the notations used in the analysis with an ideal journal. Figure 2 shows various types of geometric irregularities of the journal considered in the analysis. It is assumed that the irregularities vary sinusoidally. In commercial porous bearings, the permeability increases gradually from the ends to the middle and in the present analysis the variation is assumed to be sinusoidal, as per the curves of reference (12), and the variable permeability @,,, = @ (0"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0001147_s1068799816040085-Figure1-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0001147_s1068799816040085-Figure1-1.png",
+        "caption": "Fig. 1. The amplitude-frequency characteristics describing Eqs. (4).",
+        "texts": [
+            " (4) and at these points the closed-loop system with the amplitude phase frequency characteristic (APFC) of the open-loop system IMPROVEMENT OF HANDLING QUALITIES FOR THE AIRCRAFT LONGITUDINAL MOTION RUSSIAN AERONAUTICS Vol. 59 No. 4 2016 491 . ( ) spj o lW j e \u2212 \u03d5 \u03c9 is at the stability boundary, then the stability region with respect to the parameters k \u03d1 , z k \u03c9 is determined on the APFC segments ( )2 , ,sp z je W j k k\u03d5 \u03d1 \u03c9 \u2212 \u03c9 from the side of outer hatching 1( )W j\u03c9 applied to the left as the frequency \u03c9 increases (Fig. 1). Here the locus ( )2 , ,sp z j e W j k k \u03d5 \u03d1 \u03c9 \u2212 \u03c9 is a semicircle (with the radius 2k \u03d1 ) turned through an angle sp\u03d5 . The area with respect to the parameter z k \u03c9 at a fixed value k \u03d1 is determined in accordance with Fig. 1 on the AB arc with the aid of the inequality ( ) ( ) 1 1 2 2 * * 1 1 1 2 2 2 sin cos sin cos sin cos sin cosz sp sp sp sp sp sp sp sp a b a b k k k b a b a \u03d1 \u03c9 \u03d1 \u03d5 \u2212 \u03d5 \u03d5 \u2212 \u03d5 < < \u03c9 \u03d5 + \u03d5 \u03c9 \u03d5 + \u03d5 , (5) where ia , ib , * i\u03c9 ( 1,2i = ) are the values corresponding to the intersection points of the locus; in this case, * * 1 2av\u03c9 < \u03c9 < \u03c9 . Similarly, the condition is determined to select the parameter z k \u03c9 , at which the closed-loop system has a predetermined stability margin in absolute value 0 1spA< < . In this case, at a frequency \u03c0 \u03c9 of the APFC intersection with the real axis, the equality ( ). , , zo l spW j k k A \u03c0 \u03d1 \u03c9 \u03c9 = \u2212 should be fulfilled with the parameter area z k \u03c9 being determined by the condition: 1 2 * * 1 1 2 2 z b b k k k a a\u03d1 \u03c9 \u03d1 \u2212 \u2212 < < \u03c9 \u03c9 , (6) where ia , ib , * i\u03c9 ( 1,2i = ) have the corresponding values obtained at the locus points of intersection (see Fig. 1). When conditions (5), (6) are jointly fulfilled, the area of the parameter z k \u03c9 is determined, at which the specified stability margins sp\u03d5 , spA are attained. Find the gains Hk , Hk , Hq for the external control loop with respect to the flight altitude, assuming that ( )sp spf H H H H\u0394 \u2212 \u0394 = \u0394 \u2212 \u0394 in expression (1). In this case, we will have 1( ) ( ) sp HW p W p\u0394 \u0394\u03d1 = , ( )2 , , ,H H H HH HW p k k q k k p q p= + + in Eq. (3). Then, according to [7], for a given phase margin stability sp\u03d5 , the following condition GARKUSHENKO, VINOGRADOV RUSSIAN AERONAUTICS Vol"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0001235_b978-0-08-099425-3.00005-4-Figure5.54-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0001235_b978-0-08-099425-3.00005-4-Figure5.54-1.png",
+        "caption": "FIGURE 5.54",
+        "texts": [
+            " For ride and vibration studies the tyre model is often required to transmit the effects from a road surface where the inputs are small but of high frequency. In the simplest form the tyre may be represented as a simple compression only spring and Camber stiffness with load. (Courtesy of Dunlop Tyres Ltd.) damper acting between the wheel centre and the surface of the road. The simulation may in fact recreate the physical testing using a four-poster test rig with varying vertical inputs at each wheel. A concept of the tyre model for this type of simulation is provided in Figure 5.54 where for clarity only the right side of the vehicle is shown. In suspension loading or durability studies the tyre model must accurately represent the contact forces generated when the tyre strikes obstacles such as potholes and road bumps. In these applications the deformation of the tyre as it contacts the obstacle is of importance and is a factor in developing the model. These sort of tyre models are often developed for agricultural or construction type vehicles Aligning camber stiffness with load"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0002530_881621-Figure2-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0002530_881621-Figure2-1.png",
+        "caption": "Fig, 2 Test piston for the measurement of piston l a t e r a l motion",
+        "texts": [
+            " MEASUREMENT INSTRUMENTS MEASUREMENT DEVICE FOR PISTON FRICTION AND PISTON LATERAL MOTION - Figure 1 shows a crosssectional view of the test engine and the measurement device used for the present experiment. The test engine was a single- Numbers in parentheses designate references at end of paper. cylinder gasoline engine. For the piston friction measurement, its cylinder block was modified to floating liner structure. The liner was supported by three piezo-type pick-ups, by which the liner support became reliable. By increasing block rigidity and minimizing its deformation, high measurement accuracy was obtained. For the measurement of piston lateral motion, an inductance-type pick-up shown in Figure 2(a) was designed and manufactured, Four of these pick-ups were embedded in four different locations on the piston skirt, the distance between each of the pick-ups and the cylinder wall was measured to trace the piston lateral motion. Since this pick-up was made of copper wire with polyimide coating for the coil and of epoxy resin with high heat resistance for the core, heat resistance of the whole pick-up was increased. The pick-up was attached to the piston skirt by both screws and epoxy paint, which increased rigidity at the connection"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0000020_msn.2015.9-Figure3-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0000020_msn.2015.9-Figure3-1.png",
+        "caption": "Fig. 3. Three Angular Movement Dimensions and Propeller Motor Numbering of a Quadcopter",
+        "texts": [
+            " 400Hz) [2]. Next, let us elaborate the control strategies implemented in the main program. The control strategies include locationangular control and height control, the control outputs of the two control sub-strategies are combined to drive the four propeller motors. The location control of a flying quadcopter is tightly coupled with the quadcopter\u2019s pitch, roll, yaw angular dynamics (sometimes the three angular dynamics are holistically called the \u201cattitude\u201d of the quadcopter) control. As shown in Fig. 3, a quadcopter moves forward/backward iff its pitch angle is non-zero; a quadcopter moves leftward/rightward iff its roll angle is non-zero. Therefore, the location-angular control takes a nested outerinner control loop form, as shown in Fig. 4. The outer control loop is the location control loop; and the inner control loop is the angular control loop. The input to the location (i.e. outer) control loop is the desired location coordinates (in terms of body-oriented (x, y)coordinates, to be explained later) Xb ref = (xb ref , y b ref) T",
+            " Formally, we have aref = k p h(href \u2212 h) + kdh(h\u0307ref \u2212 h\u0307) +kih \u222b t 0 (href(\u03c4)\u2212 h(\u03c4))d\u03c4, \u0394uf = kpa(aref + g \u2212 a) + kda(a\u0307ref + g\u0307 \u2212 a\u0307) +kia \u222b t 0 (aref(\u03c4) + g(\u03c4)\u2212 a(\u03c4))d\u03c4, where k p h (also kpa), kih (also kia), and kdh (also kda) are respectively the proportional, integral, and derivative control coefficients. Finally, all the above control outputs converge to become the control output toward quadcopter propeller motors. The total control output consists of two high level components: throttle and angular control adjustments. The throttle component uf is to control the vertical acceleration (ultimately, height) of the quadcopter. It is the same to each of the four propeller motors (see Fig. 3). The height control output \u0394uf affects the throttle component: uf is updated as per uf (t+ dt) = uf (t) + \u0394uf (t). (1) The angular control adjustments are different to each of the four propeller motors. Without loss of generality, suppose we are adjusting the pitch. Suppose we want to increase the pitch angle (see Fig. 3), then propeller 1 and 3\u2019s motors should speed up, while propeller 2 and 4\u2019s motors should slow down. Meanwhile, we cannot change the total throttle component. Therefore, the pitch angle control signal u\u2032 \u03b8 should be applied positively to motor 1 and 3, but negatively to motor 2 and 4. Combining all the above considerations, suppose U1, U2, U3, and U4 respectively represent the raw total control signal applied to propeller motor 1, 2, 3, and 4, then the update rules are U1(t+ dt) = uf (t+ dt) + u\u2032 \u03b8(t) + u\u2032 \u03c6(t)\u2212 u\u2032 \u03c8(t), (2) U2(t+ dt) = uf (t+ dt)\u2212 u\u2032 \u03b8(t) + u\u2032 \u03c6(t) + u\u2032 \u03c8(t), (3) U3(t+ dt) = uf (t+ dt) + u\u2032 \u03b8(t)\u2212 u\u2032 \u03c6(t) + u\u2032 \u03c8(t), (4) U4(t+ dt) = uf (t+ dt)\u2212 u\u2032 \u03b8(t)\u2212 u\u2032 \u03c6(t)\u2212 u\u2032 \u03c8(t)",
+            " Generally, we first tune the proportional coefficients of the PID controllers to see if the quadcopter can achieve fast enough response time and acceptable overshoot. Then we increase the derivative coefficients to further reduce the overshoots. The integral control We will show the experiment result in the next section to prove our solution. We implemented the Ard-\u03bc-copter architecture described in Section III (see Fig. 1). Next, we carry out various experiments to evaluate the performance of our implementation. First, we fix the pitch axis (see Fig. 3) to a rig, which is in turn fixed on ground. We want to check whether the pitch angle control can keep the quadcopter horizontal on the pitch angular direction. We turn on the IMU sensor at time t = 0 (the IMU sensor needs to be powered on for at least 8 seconds before it can properly work); and turn on the motors at around t = 15 second. The pitch angle trace is shown in Fig. 8. According to the figure, we see the pitch angle control can effectively keep the quadcopter horizontal on the pitch angle dimension"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0000761_978-981-10-2404-7_29-Figure2-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0000761_978-981-10-2404-7_29-Figure2-1.png",
+        "caption": "Fig. 2 CAD model of the single-legged robot with vertical slider",
+        "texts": [
+            " krel;ind \u00bc krel n \u00f032\u00de Here, n equals the number of legs supporting body weight (e.g., for an insect n \u00bc 3, and for a trotting quadruped or a hopper such as kangaroo n \u00bc 2), n \u00bc 1 in this paper. krel denotes the relative stiffness and can be calculated krel \u00bc Fvert=mg Dl=l \u00f033\u00de where Fvert denotes the vertical whole-body ground reaction force; Dl denotes the compression of the whole-body leg spring; l denotes the hip height. So that the absolute spring stiffness can be calculated from (32), (33) ks \u00bc Fvert Dl \u00bc n mg l krel;ind \u00f034\u00de Figure 2 shows the virtual prototype of the single-legged robot constrained by vertical slider with CAD software. By the CAD software, mass and inertia of each segment are estimated with respect to the corresponding joint axis. The main specifications of the robot are shown in Table 1. The kinematics of the robot and the dynamic equations of motion on flight and stance phase presented on Sect. 2 are established in Simulink. As this paper mainly focuses on the dynamic model of hopping motion and the height control scheme, the hydraulic actuator dynamic model is neglected"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0003042_iros.1993.583851-Figure5-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0003042_iros.1993.583851-Figure5-1.png",
+        "caption": "Fig. 5: calculated G E P O and C O P O in horizontal gripping plane",
+        "texts": [
+            " In the case where the gripping plane plz and the upper surface of the ,table are parallel, when obstacle is slid 011 the table, each collision set COPO(obstacle,plz, Bh,) is translated in the direction of motion. When object is slid on t.he table, each grasp candidate GGPO(object, p12, 8 h ) is translated too. The sliding dista.nce of object or obstacle can be calculated using GGPO and COPO. Practically the distances to translate GGPO(object,plz, Oh) or COPO(obstacle,p12, 8h.) where a part of the GGPO is not overlapped with any COPO are calculated for every discrete sliding direction. Fig5 shows an example of calculating GPO and COPO in the case where the gripping plane p12 and the upper surface of t.he table are parallel, and Fig.6 shows the results of sliding dist,ances of object of the gripping plane pl, shown in Fig.5 in the 32 directions when GGPO equals the whole GPO. Then in order to escape from a situation where the object cannot be grasped: object is t,ranslat,ed into t,he shaded portion shown in Fig.6. In the case where the gripping plane and the upper surface of the table are not parallel, the cross line between the gripping plane and the upper surface of the table is calculated. Then we solve the problem by means of calculating the sliding distances in the directions both parallel and perpendicular to the cross line"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0001819_jsvi.1997.1111-Figure2-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0001819_jsvi.1997.1111-Figure2-1.png",
+        "caption": "Figure 2. The reference frames of the rotating shaft.",
+        "texts": [
+            " In addition, a finite element model is used to obtain the matrix equations of motion for a rotating shaft, and the effects of axial compressive loads on the stability of a rotating shaft are also studied. The dynamic stabilities of a straight rotating shaft with dissimilar lateral principal moments of inertia and subjected to axial compressive loads are investigated in the present study. Timoshenko beam theory, including rotatory inertia, shear deformations, gyroscopic moments and torsional rigidity, is applied in the formulation. The shaft with a rotating speed V and subjected to axial compressive loads is depicted in Figure 1. The reference frames used in the section are displayed in Figure 2. The rotating frame (oxyz), with unit basis vectors i , j and k , is obtained from the fixed frame (OXYZ) by a rotation of angle Vt about the X-axis. The moving frame (ot1 t2 t3), of unit basis vectors t 1, t 2 and t 3, is attached to shaft at the centroid of a cross-section. The orientation of the moving frame (ot1 t2 t3) with respect to the rotating frame (oxyz), using three Eulerian angles, is depicted in Figure 3. The frame (ox1 y1 z1) is rotated about the y-axis by an angle a with respect to the rotating frame"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0001197_978-3-658-12701-5-Figure1.2-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0001197_978-3-658-12701-5-Figure1.2-1.png",
+        "caption": "Figure 1.2: Industrial robot with seven joints",
+        "texts": [
+            " . . . . . . . . . . . . . . . 85 XV Kinematic redundancy describes a manipulator\u2019s topological property of featuring more joints than necessary to assume any configuration in its task space of given dimension. Figure 1.1 illustrates a planar robot with three joints. Since only the horizontal and the vertical position coordinates of the end-effector but not its orientation are selected to be task space coordinates, the manipulator is kinematically redundant. Similarly, the industrial robot from Figure 1.2 consists of a manipulator with six revolute joints on top of a linear axis which also results in the robot being kinematically redundant since the end-effector pose can be described by means of three coordinates each for its position and for its orientation. 1 \u00a9 Springer Fachmedien Wiesbaden 2016 A. Reiter, Time-Optimal Trajectory Planning for Redundant Robots, BestMasters, DOI 10.1007/978-3-658-12701-5_1 During the last years, the importance of kinematically redundant serial robots has risen due to striking advantages such as their improved flexibility and adaptiveness in structured workspaces and their inherent ability to perform null space motions resulting in remarkable performance in compliance tasks compared to conventional, non-redundant industrial robots",
+            " In Chapter 2 of the present thesis, the theoretical background of NURBS curves is presented. NURBS curves are a special type of mathematical curves with properties suitable for trajectory optimization tasks. Chapter 3 outlines the Projection Equation from [2], a synthetic method for obtaining the equations of motion of a dynamic system such as a robot. The Projection Equation is increasingly advantageous in cases where manipulators consist of a series of similar subsystems such as the planar robot from Figure 1.1 or the industrial robot depicted in Figure 1.2. Chapter 4 discusses the general problem of minimum-time trajectory planning along known geometric paths for kinematically redundant serial robots. Applying the knowledge of NURBS curves for the parameterization of the geometric path, a number of methods that are based on numerical inverse kinematics approaches will be investigated. Additionally, geometric methods to augment the resulting solutions in order to improve specific instantaneous properties of the robot system will be utilized. Chapter 5 introduces a separation approach based on [9] where the kinematic redundancy of the robot is directly exploited by NURBS parameterizations of the path parameter and the trajectory of the redundant coordinate"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0001996_s0736-5845(98)00026-x-Figure9-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0001996_s0736-5845(98)00026-x-Figure9-1.png",
+        "caption": "Fig. 9. Compensation of the deflection caused by elasticity.",
+        "texts": [
+            " Finally, the deflection of the elastic-rigid modeled robot has to be compensated. Therefore, the elastic beams of the robot were considered as Bernoulli beams [17, 18] whose deformation is described by the Ritz\u2014 Raleigh Approach [19]. An efficient approach, which is exactly enough for offline-programming, gives [20]. Herein the elastic deformation is described by a characteristic polynomial. The displacement dP f,r , caused by the elastic deflection, is iterative compensated by inverse kinematics. To compensate the displacement dP f,r , in Fig. 9, we have the following steps from [20]: f Modeling the kinematical structure of the robot as Bernoulli beams. f Calculating the flexible displacement dP f,r at the tool center point by forward kinematics of elastically modeled robot versus rigidly modeled robot. f Incremental compensation of dP f,r , by inverse kinematics for redundant robots as used in Eq. (12). Within this section a motion planning method is presented, which creates a time optimal velocity profile along a specified path with respect to the robots dynamical and kinematical constraints"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0000477_j.ifacol.2015.09.726-Figure1-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0000477_j.ifacol.2015.09.726-Figure1-1.png",
+        "caption": "Fig. 1. Experimental bench with the small scale alternator",
+        "texts": [
+            " The experimental test bench as well as its numerical model are presented in section 2. Section 3 is devoted to the rotor inter-turn short-circuit detection. Finally, the diagnosis and the obtained results are described in section 4. AND ITS NUMERICAL MODEL To validate this approach, an experimental bench, constituted of a reduced scale alternator driven by an induction motor is used (see Richard et al. (1997)). A numerical model based on FEM is built up. Then, the simulation results, at no load, are compared with the measurements. The experimental machine, presented in Fig 1, is similar to a power plant alternator. It is a 3 phases, 4 poles, 50 Hz synchronous machine with 48 stator and 36 rotor slots. The latter receive the rotor winding for the excitation field with also 36 short-circuited damping bars. This machine can be considered as a smooth rotor machine with a constant air gap of 1.6 mm. Furthermore, a radial flux probe is 9th IFAC Symposium on Fault Detection, Supervision and Safety of Technical Processes September 2-4, 2015. Arts et M\u00e9tiers ParisTech, Paris, France Copy ight \u00a9 2015 IFAC 1432 Alternator Rotor Inter-turn Short-Circuit Identification using FEM Based Learning Alexandre Bacchus \u2217,\u2217\u2217 Melisande Biet \u2217\u2217 Ludovic Macaire \u2217 Abdelmounaim Tounzi \u2217 Yvonnick e Menach \u2217 \u2217 L2EP and CRIStAL (UMR CNRS 9189), University of Lille 1, Villeneuve d\u2019Ascq 59650, France (e-mail contact: alexandre",
+            " The experimental test bench as well as its numerical model are presented in section 2. Section 3 is devoted to the rotor inter-tur short-circuit detectio . Finally, the diagnosis and the obtained res lts are described in section 4. AND ITS NUMERICAL MODEL To validate this approach, an experimental bench, constituted of a reduced scale alternator driven by a induction motor is used (see Richard et al. (1997)). A numerical odel based on FEM is built up. Then, the simulation results, at no load, are compared with the measurements. The experimental machine, presented in Fig 1, is similar to a power pla t alternator. It is a 3 phases, 4 poles, 50 Hz synchronous machine with 48 stator and 36 rotor slots. The latter receive the rotor winding for the excitation field with also 36 short-circuited damping bars. This machine can be considered as a smooth rotor machine with a constant air gap of 1.6 mm. Furt ermore, a radial flux probe is 9th IFAC Symposium on Fault Detection, Supervision and Safety of Technical Pr cesses eptember 2-4, 2015. Arts et M\u00e9tiers ParisTech, Paris, France Copyright I AC 1432 lter ator otor I ter-t r ort- irc it I e ti catio si g ase ear i g lexandre Bacchus \u2217,\u2217\u2217 elisande Biet \u2217\u2217 Ludovic acaire \u2217 bdel ounai Tounzi \u2217 vonnick Le enach \u2217 \u2217 L2EP and CRIStAL (U R CNRS 9189), Universi y of Lille 1, Villeneuve d\u2019Ascq 59650, France (e-mail contact: alexandre",
+            " The exp rimental test bench as well as its numerical m del are presented in section 2. S i n 3 is devoted to the rotor inter-turn short-circui detection. Finally, the diagnosis and the obtained results are described in section 4. AND ITS NU ERICAL ODEL To validate this approach, an experimental bench, constituted of a reduced scale alternator driven by an induction m tor is used (see Richard et al. (1997)). A numerical model based on FE is built up. Then, the imulation results, at no load, are compared with the measurements. The exp imen al machine, presented in Fig 1, is similar to a power plant alterna or. It is a 3 phases, 4 poles, 50 Hz synchronous machine with 48 stator and 36 rot r slots. The latter receive the rotor winding for the excitation field with also 36 short-circuited damping b rs. This machine can be considered as a smoo h rot r machine with a constant air gap of 1.6 mm. Furthermore, a radial flux probe is 9th IFAC Symposium on Fault Detection, Supervision and Safety of Technical Processes September 2-4, 2015. Arts et M\u00e9tiers ParisTech, Paris, France Copyright \u00a9 2015 IFAC 1432 Alternator Rotor Inter-turn Short-Circuit Identification using FE Based Learning Alexandre Bacchus \u2217,\u2217\u2217 Melisande Biet \u2217\u2217 Ludovic Macaire \u2217 Abdelmounaim Tounzi \u2217 Yvonnick Le Menach \u2217 \u2217 L2EP and CRIStAL (UMR CNRS 9189), University of Lille 1, Villeneuve d\u2019Ascq 59650, France (e-mail contact: alexandre",
+            " The experimental test bench as well as its numerical model are presented in section 2. Section 3 is devoted to the rotor inter-turn short-circuit detection. Finally, the diagnosis and the obtained results are described in section 4. AND ITS NUMERICAL MODEL To validate this approach, an experimental bench, constituted of a reduced scale alternator driven by an induction motor is used (see Richard et al. (1997)). A numerical model based on FEM is built up. Then, the simulation results, at no load, are compared with the measurements. The experimental machine, presented in Fig 1, is similar to a power plant alternator. It is a 3 phases, 4 poles, 50 Hz synchronous machine with 48 stator and 36 rotor slots. The latter receive the rotor winding for the excitation field with also 36 short-circuited damping bars. This machine can be considered as a smooth rotor machine with a constant air gap of 1.6 mm. Furthermore, a radial flux probe is 9th IFAC Symposium on Fault Detection, Supervision and Safety of Technical Processes September 2-4, 2015. Arts et M\u00e9tiers ParisTech, Paris, France Copyright \u00a9 2015 IFAC 1432 Alexandre Bacchus et al",
+            " Section 3 is devoted to the rotor inter-turn short-circuit detection. Finally, the diagnosis and the obtained results are described in section 4. 2. SMALL SCALE BENCH OF AN ALTERNATOR AND ITS NUMERICAL MODEL To validate this approach, an experimental bench, constituted of a reduced scale alternator driven by an induction motor is used (see Richard et al. (1997)). A numerical model based on FEM is built up. Then, the simulation results, at no load, are compared with the measurements. The experimental machine, presented in Fig 1, is similar to a power plant alternator. It is a 3 phases, 4 poles, 50 Hz synchronous machine with 48 stator and 36 rotor slots. The latter receive the rotor winding for the excitation field with also 36 short-circuited damping bars. This machine can be considered as a smooth rotor machine with a constant air gap of 1.6 mm. Furthermore, a radial flux probe is 9th IFAC Symposium on Fault Detection, Supervision and Safety of Technical Processes September 2-4, 2015. Arts et M\u00e9tiers ParisTech, Paris, France Copyright \u00a9 2015 IFAC 1432 Alternator Rotor Inter-turn Short-Circuit Identification using FEM Based Learning Alexandre Bacchus \u2217,\u2217\u2217 Melisande Biet \u2217\u2217 Ludovic Macaire \u2217 Abdelmounaim Tounzi \u2217 Yvonnick e Menach \u2217 \u2217 L2EP and CRIStAL (UMR CNRS 9189), University of Lille 1, Villeneuve d\u2019Ascq 59650, France (e-mail contact: alexandre",
+            " Section 3 is devoted to the rotor inter-tur short-circuit detectio . Finally, the diagnosis and the obtained res lts are described in section 4. 2. SMALL SCALE BENCH OF AN ALTERNATOR AND ITS NUMERICAL MODEL To validate this approach, an experimental bench, constituted of a reduced scale alternator driven by a induction motor is used (see Richard et al. (1997)). A numerical odel based on FEM is built up. Then, the simulation results, at no load, are compared with the measurements. The experimental machine, presented in Fig 1, is similar to a power pla t alternator. It is a 3 phases, 4 poles, 50 Hz synchronous machine with 48 stator and 36 rotor slots. The latter receive the rotor winding for the excitation field with also 36 short-circuited damping bars. This machine can be considered as a smooth rotor machine with a constant air gap of 1.6 mm. Furt ermore, a radial flux probe is 9th IFAC Symposium on Fault Detection, Supervision and Safety of Technical Pr cesses eptember 2-4, 2015. Arts et M\u00e9tiers ParisTech, Paris, France Copyright \u00a9 2015 IFAC 1432 lter ator otor I ter-t r ort- irc it I e ti catio si g ase ear i g lexandre Bacchus \u2217,\u2217\u2217 elisande Biet \u2217\u2217 Ludovic acaire \u2217 bdel ounai Tounzi \u2217 vonnick Le enach \u2217 \u2217 L2EP and CRIStAL (U R CNRS 9189), Universi y of Lille 1, Villeneuve d\u2019Ascq 59650, France (e-mail contact: alexandre",
+            " S i n 3 is devoted to the rotor inter-turn short-circui detection. Finally, the diagnosis and the obtained results are described in section 4. 2. S ALL SCALE BENCH OF AN ALTERNATOR AND ITS NU ERICAL ODEL To validate this approach, an experimental bench, constituted of a reduced scale alternator driven by an induction m tor is used (see Richard et al. (1997)). A numerical model based on FE is built up. Then, the imulation results, at no load, are compared with the measurements. The exp imen al machine, presented in Fig 1, is similar to a power plant alterna or. It is a 3 phases, 4 poles, 50 Hz synchronous machine with 48 stator and 36 rot r slots. The latter receive the rotor winding for the excitation field with also 36 short-circuited damping b rs. This machine can be considered as a smoo h rot r machine with a constant air gap of 1.6 mm. Furthermore, a radial flux probe is 9th IFAC Symposium on Fault Detection, Supervision and Safety of Technical Processes September 2-4, 2015. Arts et M\u00e9tiers ParisTech, Paris, France Copyright \u00a9 2015 IFAC 1432 Alternator Rotor Inter-turn Short-Circuit Identification using FE Based Learning Alexandre Bacchus \u2217,\u2217\u2217 Melisande Biet \u2217\u2217 Ludovic Macaire \u2217 Abdelmounaim Tounzi \u2217 Yvonnick Le Menach \u2217 \u2217 L2EP and CRIStAL (UMR CNRS 9189), University of Lille 1, Villeneuve d\u2019Ascq 59650, France (e-mail contact: alexandre",
+            " Section 3 is devoted to the rotor inter-turn short-circuit detection. Finally, the diagnosis and the obtained results are described in section 4. 2. SMALL SCALE BENCH OF AN ALTERNATOR AND ITS NUMERICAL MODEL To validate this approach, an experimental bench, constituted of a reduced scale alternator driven by an induction motor is used (see Richard et al. (1997)). A numerical model based on FEM is built up. Then, the simulation results, at no load, are compared with the measurements. The experimental machine, presented in Fig 1, is similar to a power plant alternator. It is a 3 phases, 4 poles, 50 Hz synchronous machine with 48 stator and 36 rotor slots. The latter receive the rotor winding for the excitation field with also 36 short-circuited damping bars. This machine can be considered as a smooth rotor machine with a constant air gap of 1.6 mm. Furthermore, a radial flux probe is 9th IFAC Symposium on Fault Detection, Supervision and Safety of Technical Processes September 2-4, 2015. Arts et M\u00e9tiers ParisTech, Paris, France Copyright \u00a9 2015 IFAC 1432 located in the airgap in order to locally measure the radial magnetic flux density",
+            " The healthy class is correctly classified, which is expected because of the translation. Moreover, B4 class is almost found. The only data that is misclassified is affected to a similar class. The misclassification comes from the high variability of experimental classes compared to the simulated ones. Nevertheless, A1 class is affected to two simulated classes. Regarding to Table 1, both faults induces a decrease of 3.7% of ampere-turns. Indeed, only 6 turns are short-circuited for both faults (see Fig 1). For these classes, only slots are different, this is why the classes are distinct. Consequently the error made by the fault identification is reasonable. Finally, class B2B4 is classified as B2B4 and A1B2B4. A1 fault is not relevant in front of B2B4 which is why both classes are very close in the PCA representation. In this paper, we have proposed a new alternator rotor inter-turn short-circuit identification procedure. Is is based on a numerical model of the simulated machine by FEM that generates prototypes of the different faulty states"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0001054_cgncc.2016.7828871-Figure2-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0001054_cgncc.2016.7828871-Figure2-1.png",
+        "caption": "Fig. 2: The 3D structure view of the pointing mirror",
+        "texts": [
+            " The remainder of this paper is organized as follows. The VCM control model is introduced in Section II. In Section III, the design procedure is constructed. In Section IV, stability analysis is verified with Lyapunov theory. In Section V, Simulation results are provided to illustrate the effectiveness of the control scheme. Finally, the conclusions are drawn in Section VI. The pointing mirror is attached to the two orthogonal flexibility shafts which are driven by two VCMs respectively as shown in Fig 2. So the controller for each axis can be designed separately. As the output torque of the VCM is generated by applying a current to the coil of the VCM, so we consider the VCM to be a mechanical damper system that can be described by an equivalent electrical circuit, by writing the voltage and torque balance equations for the system, the model dynamic can be described by the following three-coupled differential equations m n b K i K Li E Ri K (1) For convenience, we define m mK K J , n nK K J , (rad) and (rad/s) are the angular position and angular velocity of the motor, respectively"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0001463_978-94-6239-082-9_1-Figure1.2-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0001463_978-94-6239-082-9_1-Figure1.2-1.png",
+        "caption": "Fig. 1.2 Inverted pendulum system",
+        "texts": [
+            "4, we should taken into consideration that the factor \u03b3 of the non iterative method weights the norm of the parameters in front of the quadratic sum of errors resulted for the modeling of samples. On the other hand, the factor weights the norm of parameters infront of each one of the errors of samples. If m is the number of used samples, then the equivalence is produced supposing that \u03b4 = \u03b3 /m and in general \u03b4 \u03b3 should be fulfilled. In this section the proposed estimation method is illustrated by an example of an inverted pendulum (see Fig. 1.2). The inverted pendulum can be represented as follows: \u03b8\u0308 = gsen\u03b8 \u2212 cos \u03b8( u+ml \u03b8\u03072sen\u03b8 M+m ) l( 4 3 \u2212 m cos2 \u03b8 M+m ) (1.99) where \u03b8 denotes the angular position (in radians) deviated from the equilibrium position (vertical axis) of the pendulum and \u03b8\u0307 is the angular velocity, g(gravity acceleration) = 9.8 m s2 , M(mass) of the cart=1 kg, m(mass) of the pole=0.1 kg, l is the distance from the center of the mass (m) of the pole to the cart=0.5 m. Assuming that x1 = \u03b8 and x2 = \u03b8\u0307 , then (1.99) can be rewritten in state space form as follows: x1 = \u03b8 (1"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0000054_oceans.2015.7404427-Figure2-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0000054_oceans.2015.7404427-Figure2-1.png",
+        "caption": "Fig. 2: WAM-V USV16 pr",
+        "texts": [],
+        "surrounding_texts": [
+            "The propulsion system is custom peak thrust of 240N. This inclu thrusters, each powered by a 12V 160 N, 15 cm stroke, linear actuat thrusters through an azimuthal ang the vehicle\u2019s longitudinal directio actuated. The propulsion system an be seen in Fig. 2and Fig. 3 respectiv"
+        ]
+    },
+    {
+        "image_filename": "designv11_60_0002923_978-94-011-3514-6-Figure3.11-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0002923_978-94-011-3514-6-Figure3.11-1.png",
+        "caption": "Fig. 3.11 Free state of the damped oscillator. Displacement as a function of time.",
+        "texts": [
+            "41) in the sense that the hyperbolic functions are simply replaced by the trigonometric functions. These functions can be combined (see figure 3.1 page 10) and the displacement takes the simple form with { X tgll> AXa + Vo Wl Xo In this way, the quantity (3.57) (3.58) x is always larger than the initial displacement Xo , except in the very specific case Vo = -A Xo . Likewise, the phase shift II> is always different from zero, except if Vo = -A Xo . wave enclosed between two envelopes \u00b1 X e-At (figure 3.11). - 33 - The period of x(t) or more precisely the pseudo-period since the amplitude diminishes - has the value The damping reduces the angular frequency and increases the period of the oscillations. In effect, using again relation (3.52). (3.60) concept of logarithaic decrement, defined as follows A=l~n x(t) n x(t + nT l ) (3.61) - 34 - x(t + nTl) e-n\"Tl x(t) whence 1\\ (independent of n) n <lIO and T 1 2rr the logarithmic decrement is only a function of the damping factor 1\\ 2 11 n (3.62) ~ and vice versa 1\\ (3"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0003684_20.996240-Figure1-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0003684_20.996240-Figure1-1.png",
+        "caption": "Fig. 1. Basic principle of the moving band.",
+        "texts": [
+            " to are matrices that depend on the converter topology and change at each switch commutation, so they have to be constructed at each operation sequence of the converter. This is made in an entirely automatic way, according to the switches commutation characteristics. The moving band technique was chosen to consider the movement during the field calculation. In this method, the air gap is meshed and the rotation is taken into account by means of a layer of finite elements placed in the air gap [Fig. 1(a)]. The rotation angle is a function of the time step and the speed. At each step, the moving part displaces [Fig. 1(b)], so the elements of the moving band are distorted and, when the distortion is large enough, the air gap is remeshed [Fig. 1(c)]. Periodicity or antiperiodicity conditions (boundaries and ) are used to perform a dynamic allocation of domain nodes, consequently the size of the corresponding matrices is not increased [15]. During the simulation, a large number of switches commutations can occur. According to the switch type, this commutation occurs at the zero crossing of their current or at the zero crossing of the voltage across their terminals. This instant has to be determined accurately because it may introduce numerical problems, and consequently physical errors in the iterative solution of the equations system (1)"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0000882_s1061933x16060181-Figure2-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0000882_s1061933x16060181-Figure2-1.png",
+        "caption": "Fig. 2. Distribution diagram for different forms of folic acid: (1) H3L neutral molecule and (2) H2L\u2013, (3) HL2\u2013, and (4) L3\u2013 anions.",
+        "texts": [
+            "00 was related to the deprotonation of the N3H\u2013C4O moiety of pterin ring. Thus, a folic acid molecule is a triacid from which two protons of carboxyl groups of glutamic 2CeOH+ int 1log SK int 2log SK 2 + int CllogK \u2212 int NalogK + 3 + COLLOID JOURNAL Vol. 78 No. 6 2016 ADSORPTION OF AMINO ACIDS 749 acid and, then, the proton of N3H\u2013C4O moiety are detached. Different forms of folic acid are present in aqueous solutions depending on pH. They are neutral molecules and mono-, di- and trivalent anions. Figure 2 shows the distribution diagram for differently protonated forms of folic acid. The data on adsorption of the amino acids as depending on pH are presented in Fig. 3. Adsorption of folic acid was studied only for solutions with pH > 4, because it is almost insoluble in water at lower pH values. The adsorption curves exhibit slight inflections, the positions of which in the pH scale nearly correspond to the dissociation constants of the additional carboxyl groups in the side chains of amino acids. The adsorption value of folic acid is markedly higher than the values for the two simple amino acids"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0003689_robot.2001.932951-Figure4-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0003689_robot.2001.932951-Figure4-1.png",
+        "caption": "Fig. 4. P1ana.r enveloping grasp of a. pentagonal object using two fingers",
+        "texts": [
+            "3 normal contact forces a t Cki, is necessary to check the feasibility as well as the stability, because the friction cone constraint is given by where p k z is the frictional coefficient a t c k z . We also assumed that (A3) m k = n k for all k = 1 , 2 , . . . , K and a k ( i ) = i for i = 1 , 2 , . . . , m k . These assumptions amount to the conditicm that there is one contact for every link of the limbs, and no contact with base limb. In addition, (A4) the limb-tils contact with the object is not allowed. One such example is given by Fig. 4. 3.1 Geometrical decomposition of contact force The proposed statical decomposition of the total contact force (4) has the following geometrical decomposition property. Theorem 1: Given a grasp and contact configuration E and C, then the contact force at each contact C k , can be decomposed as which is called the base-ward accumulation property. Fiirthermore, each contact force is geometrically decomposed by f k z = f k l t + f k z n r denoting the tangential force and normal force at the i-th contact point of the k-th limb by f kzt E R2 and f kzn E R2 respectively, where f k i t = (10) which is called the base-ward normal accumulation property",
+            "R + SO,lO,} r b Stable grasp synthesis To find the region of joint torques T satisfying the grasp feasibility inequality, i.e. solve AT7 I bw, where A, = -AR - SflF'fl,, b, = -Sfl ; 'm~ b Grasp robustness analysis To find the region of the object wrenches W O satisfying the grasp feasibility inequality, i.e. solve Awwo I I , , , where A, = Sf2F1, b, = { hR + SCl;'fl,} T 5 Numerical Examples Consider the planar two-fingers consisting of three degrees-of-freedom one and two degrees-of-freedom one, grasping a pentagonal object whose side lengths are all 0.2(m), as shown in Fig. 4. The link lengths of the first finger are all 0.2(m), and those of the second one are all 0.16(m). The joint angles are: q1 = (-117.5561,53.7940, G9.0056)T(0) and q2 = (-24.5561,-55.0610)T(0). Each link has one contact with each vertex of the pentagonal object with p1 = (0.1,0.12393,0.1!5971)T(m) and pa = (0.04523, 0.11076)T(m). At current configuration, the object is rotated by 38.4439(\"). Let us assume that all friction coefficignts are 1.0. The indeterminate friction force is chosen as E = (613 , I - 2 0 i;; I--~::::: 1 -30 5"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0000153_imece2015-50907-Figure2-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0000153_imece2015-50907-Figure2-1.png",
+        "caption": "FIGURE 2. THE JOINTS OF ABB IRB 5500.",
+        "texts": [
+            " Since the end effector should always maintain normal direction to the surface according to the requirements of painting process, it is difficult for the calculation of the workspace and to optimize the stop position with the original point set on the base. To make a better determination of the stop position, we propose a creative model to set the base frame at the center wrist point O and a new configuration is formulated. The scheme is mainly implemented on ABB IRB 5500 in this paper. For its special non-spherical wrist, the original point of the base frame is actually at the virtual wrist point which is the projection of the intersection point of axis 5-1 and 5-2 onto axis 4 when the joint angles of axis 4 and 5 are both 0, as shown in Fig. 2. Meanwhile, the link 1 and 2 of the new configuration are the forearm and upper arm of the real manipulator. We also leave out the offsets between Axis 1, 2 and Axis 3, 4 which does not impact on the position of the real base and simplifies the calcula- tion. And because the motion ranges of axis 4, 5, 6 are extensively wide, a 3-DOF spherical joint is designed at the virtual wrist point in addition. Coordinate systems are formulated and noted as S1, S2, S3, S4, S5, S6 which are shown as Fig. 3 "
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0001777_951293-Figure12-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0001777_951293-Figure12-1.png",
+        "caption": "Figure 12. Static Hysteresis Loop of Dry Friction Clutch",
+        "texts": [
+            " The total force resisting the relative motion of bodies D and H is RF = - kxT F ...( 13) and exemplifies the reaction torque exerted by the input shaft on the engine flywheel. Let us consider that the clutch damper has a constant stiffness and experiences dry friction damping determined by different but constant (independent of the relative speed) static and dynamic coefficieiits of friction. If the damper relative movement is slow (static), one can present the resisting force as a function of displacement as shown in Fig. 12 (compare to Fig. 2). Now, let us assume that the above system osculates about a certain equilibrium position E (see Fig. 13) with certain considerable frequency, and that the amplitude of its motion in a steady state is 1x1. During the motion, the torquedisplacement relationship will be a closed curve as shown in Fig. 13 (compare Fig. 9). One can expect that the size and shape of the hysteresis loop will depend on many parameters of the driveline such as: the amplitude and frequency of oscillations, location of the equilibrium position, friction characteristics, inertias of bodies D and H, stiffness of the damper, etc"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0000647_med.2016.7536060-Figure1-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0000647_med.2016.7536060-Figure1-1.png",
+        "caption": "Fig. 1. Hexacopter",
+        "texts": [
+            " This reduces overparameterization due to overactuation since \u0398b1 \u2208 Rr\u00d7(n+l+k+1) and \u0398 \u2208 Rm\u00d7(n+l+k+1) with r < m. The control law can then be written as u = V1\u0398b1\u03d5. (15) In the next section, the example of a hexarotor will be introduced. It will be demonstrated that in this example failure of a rotor and even change in the direction of the thrust can be handled with the standard MRAC update law. The description of the hexarotor dynamics corresponds to the previous work [12], [13]. For describing the dynamics of the hexacopter we use an inertial frame I and a body-fixed frame B as in Figure 1, such that the origin is at the center of gravity. The rotational dynamics are given in the body-fixed frame B by the Euler\u2019s equation J\u03c9\u0307 = \u2212\u03c9 \u00d7 J\u03c9 +Mp +Md, (16) where \u03c9 \u2208 R3 is the angular rate of the body-fixed frame relative to the inertial frame, J \u2208 R3\u00d73 is the moment of inertia, Mp \u2208 R3 is the propulsion moment and Md \u2208 R3 is the disturbance moment. The disturbance moment is a result of unmodeled aerodynamics, wind, parameter errors, etc. Note that Mp can be seen as the virtual control of the system (16). The roll and pitch moments depend on the geometrical arrangement of the rotors. The yaw moment depends on the rotational direction of the rotors. Considering the configuration as in Figure 1, it is possible to write their relationship to the angular velocities of the rotors as [13] Mp = BM\u03c9u, (17) where u = ( \u03c92 1 \u03c92 2 \u03c92 3 \u03c92 4 \u03c92 5 \u03c92 6 )T is the vector of the squared rotor speeds and BM\u03c9 \u2208 R3\u00d76 is defined as [13]\u23a1 \u23a3\u22120.5lkT \u2212lkT \u22120.5lkT 0.5lkT lkT 0.5lkT\u221a 3 2 lkT 0 \u2212 \u221a 3 2 lkT \u2212 \u221a 3 2 lkT 0 \u221a 3 2 lkT \u2212kM kM \u2212kM kM \u2212kM kM \u23a4 \u23a6 The constants kT , kM > 0 \u2208 R are specific rotor parameters and l > 0 \u2208 R is the arm length. For a detailed derivation of the actuator model (17) refer to [13]",
+            " By defining the plant state as xp = \u03c9, rewrite the rotation dynamics (16) using (17) and by introducing the diagonal effectiveness matrix \u039b \u2208 R6\u00d76 x\u0307p = \u2212J\u22121 (xp \u00d7 Jxp) + J\u22121Mp + J\u22121Md, = \u2212J\u22121 (xp \u00d7 Jxp)\ufe38 \ufe37\ufe37 \ufe38 :=f(xp) +J\u22121BM\u03c9\ufe38 \ufe37\ufe37 \ufe38 :=Bp \u039bu+ J\u22121Md\ufe38 \ufe37\ufe37 \ufe38 :=d . The nominal value of \u039b is \u039b = I. By assuming a diagonal moment of inertia we can parameterize the nonlinear function as f (xp) = \u2212J\u22121 (xp \u00d7 Jxp) , = \u23a1 \u23a2\u23a3 Jyy\u2212Jzz Jxx 0 0 0 Jzz\u2212Jxx Jyy 0 0 0 Jxx\u2212Jyy Jzz \u23a4 \u23a5\u23a6 \ufe38 \ufe37\ufe37 \ufe38 :=Kf \u239b \u239d\u03c9z\u03c9y \u03c9z\u03c9x \u03c9x\u03c9y \u239e \u23a0 \ufe38 \ufe37\ufe37 \ufe38 :=\u03c6(xp) . In this example, the disturbance moment comes from the moment produced by an externally commanded normalized thrust Tn \u2208 R as Md = BM\u03c9\u039b [ 1 1 1 1 1 1 ]T Tn. For the hexacopter as in Figure 1, this disturbance moment is zero in the nominal case, but unequal zero in the case of degradation. We choose the reference model as x\u0307m = Amxm +Bmr, = \u23a1 \u23a3\u22127 0 0 0 \u22127 0 0 0 \u22124 \u23a4 \u23a6xm + \u23a1 \u23a37 0 0 0 7 0 0 0 4 \u23a4 \u23a6 r. (18) Hence, we have \u03d5 \u2208 R10 and \u0398 \u2208 R3\u00d710. In this section, the simulation results are presented. The implemented plant dynamics are the ones described by (16). The moment of inertia is diagonal and Jxx = Jyy = 10\u22123[kg \u00b7 m2] and Jzz = 18 \u00b7 10\u22123[kg \u00b7 m2]. The arm length is l = 0.215[m], the motor constants are kT = 6"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0003872_6.2001-2433-Figure1-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0003872_6.2001-2433-Figure1-1.png",
+        "caption": "Figure 1. Geometric Characteristics of Wind Tunnel Model with Canards and Nose Flaps",
+        "texts": [
+            " However, the location of the controls in the expansion region along the missile nose may result in unknown aerodynamic and thermodynamic effects occurring in a high dynamic pressure environment. Such effects include the interference of the nose shock with the control surface, flow choking and increased pressure on the nose due to mutually deflected controls, non-uniform variations in local conditions along the control surface, flow separation, and gap heating. Canards The geometric characteristics of the missile model detailing the nose-mounted canards and flaps are provided in Figure 1. The canards incorporated on this missile configuration were designed during previous studies6 to provide a desired level of control for a similar missile airframe of increased length. In comparison to the canard detailed in Reference 6, the canard presented herein was designed to lie within the nose bow shock at Mach 6.0 and was previously wind tunnel tested at these conditions. Nose Flaps The three nose flaps, also shown in Figure 1, were designed from both a theoretical and experimental basis. The initial nose flap 1 was designed from a purely theoretical approach. Missile DATCOM7 and NEAR ZEUS8 predictions of the aerodynamic stability for the missile body with tail fins (no flaps) were obtained at Mach 6 and angles of attack from 0 to 15 deg. Incremental changes in CN, Cm, and CAP due to pitch control (ACN, ACm, and ACAF) were determined for nose flap deflections of 20 American Institute of Aeronautics and Astronautics and 35 deg from integration of the pressure coefficient given by Modified Newtonian theory9 applied to an inclined flat plate, (i) where Cf",
+            "00 in. The wind tunnel is equipped with a model injection system that includes a remote-controlled, sting support system capable of rotating the model through an angle-of-attack range from -5 deg to 55 deg and an angle-of-sideslip range from -10 deg to 10 deg (precision of \u00b10.05 deg). A detailed description of the tunnel is presented in Reference 10. Wind Tunnel Model and Instrumentation The geometric characteristics of the missile model detailing the nose-mounted canards and flaps are provided in Figure 1. The missile consists of a 3.0-caliber tangent ogive nose, a cylindrical aftbody with a maximum diameter of 2.214 inches, and four stabilizing tail fins. A single canard geometry and three distinct nose flaps of varying flap area were designed to provide pitch control. The model measures 17.367 in. in length with a body diameter of American Institute of Aeronautics and Astronautics 2.214 in. Model Station 0.000 is defined as the tip of the pointed ogive nose. The canards and nose flaps are mounted along the nose at a location 4",
+            "50 to that of flap 1, respectively. The aspect ratios of the flaps are maintained at 2.93. The nose flaps are located 90- deg apart and inline with the tail fins. The configuration nomenclature provided for the plot legends is given as: B (missile body, nose+aftbody), C (canards), T (tail fins), NF1 (nose flap 1 - small flap), NF2 (nose flap 2 - medium flap), and NF3 (nose flap 3 - large flap). A nominal moment reference point (MRP) was located 8.836 in. (3.991 calibers) aft of the model nose as shown in Figure 1. The wind tunnel model was instrumented with a 0.75-in., six-component main balance (2019-D) that utilized a water-cooling jacket for temperature stabilization. No boundary layer transition strips were applied to the model or its components during the test. Figure 3 presents the canard and nose flap pitch deflection orientation. As shown a positive pitch deflection provides a positive normal force (up) and positive pitching moment (nose up) about the moment reference point for the nose-controlled missile at 0- deg angle of attack"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0001091_icsengt.2016.7849648-Figure3-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0001091_icsengt.2016.7849648-Figure3-1.png",
+        "caption": "Fig 3. Kinematics Frame of Hexacopter",
+        "texts": [],
+        "surrounding_texts": [
+            "Keywords\u2014 Chemical Hazard Monitoring, Hardware-In-TheLoop Simulation, Hexacopter, Unmanned Air Vehicle.\nI. INTRODUCTION\nThe development phase of UAV requires huge costs and have a high risk, especially in the field trials [3]. The main causes are pricey components, sensors, and the replacement parts if crash occurred caused by system failure or human error. To avoid and minimize the risk of failure, a methods is used to meet the development process safety standards.\nHardware-In-The-Loop Simulation (HILS) is a technique used in the development and testing phase of complex embedded real-time systems [4]. HILS provides advantages such as reduce the development costs, improve safety standards, and accelerate the development process [3]. With HILS, UAV development process can be executed without using the actual UAV platform, but by using mathematical model to represent the platform that will be used [4], [6].\nUAV simulated using java based software. Kinematics and dynamics equations are used to evaluate the response that occurs in UAV Hexacopter models. Physical parameters are also included in the equation to match the model with actual\nconditions. HILS then simulate the UAV embedded system as if it is on the real flight [5].\nThe hardware used as flight control is Pixhawk. Pixhawk is a microcontroller (32-bit processors) that have capability to operate automated flight and is oftenly used as UAV flight controller among hobbyists and professionals [8].\nHILS also aims to simulate the flight missions that will be performed by the UAV [5], in this case is for chemical contamination monitoring of the air. The pricey chemical sensor device requires the success of simulation before testing the actual flight mission. Simulations with the integration of other modules should also be made to ensure the success of the monitoring mission.\nII. HILS DESIGN\nTo accelerate the HILS development process, a software simulator developed by the Pixhawk engineering team is used, namely jMAVSim. jMAVSim is chosen for its ability to simulate the flight characteristics of multicopter, visualize the motion of multicopter on the environment, and do fairly accurate modeling of the sensor with some noise that might be encountered by the sensor. jMAVSim works by receiving data transmitted by Pixhawk MAVLink via Serial UART communication protocol. MAVLink is a standard communication protocol used on UAVs as packets transmitted from the flight control hardware to the ground control station (GCS) and vice versa. Simulator PC and GCS run in the same PC so that the previous data obtained by MAVLink can be forwarded from jMAVSim to GCS software through UDP communication protocols.\nGCS software then display various data such as the aircraft attitude and position as in the real flight. MAVLink data can be used by the GCS to analyze the flight characteristics, sensors, absolute position of hexacopter to the earth, and others. HILS data flow can be seen in Figure 1. Manual flight control commands is given by the operator via the remote control, while the autonomous commands given through the waypoint commands entered by the operator via GCS software. After the waypoint command is sent to the hexacopter control system, then the autonomous flying mode can be activated via remote control.\n978-1-5090-5089-5/16/$31.00 \u00a92016 IEEE 189",
+            "MAVLink data that is foremost used in HILS is HIL_CONTROLS data structure which is the output rotor speed command of each rotor, HIL_SENSOR data structure which is the modeled sensor in response to UAV flight characteristics as a control input, and HILS_GPS data structure which is modeling the GPS sensor to get the position and speed of the UAV relative to the earth.\nA. Hexacopter Modeling\nKinematics and dynamics of UAV hexacopter are derived from the Newton\u2019s Laws of motion and the principle of conservation momentum as well as additional air resistance. Hexacopter referenced frame orientation refers to the Flat Earth inertial frame (X, Y, Z)0, [9].\nThe motion equations of Newton-Euler model of the rigid body is influenced by force and torque from the outside written as follows,\n(1) where m [kg] is mass, I [Nm2] hexacopter inertia, VB = [u v w ] [m/s] is linear velocity of hexacopter frame, \u03c9B = [ p q r ] [rad/s] is angular velocity of hexacopter frame, FB [N] is working force in hexacopter frame, \u03c4B [Nm] is working torque in hexacopter frame, 03x3 is zero matrix with size of 3x3, and I3x3 is 3x3 identity matrix [1], [2], [7]. While the force of FB and torque of \u03c4B defined according to [7], [8], and [9]. There are two main components of force consisted of gravitational force and lift force generated by rotor thrust. To make the model characteristics more realistic, rotor drag and air resistance components are added to the equation. Gravitational force FB\ngravity will always be pointing downward along the Z axis in Flat Earth reference frame. In relation to hexacopter frame written as follows,\n(2) where RB\nE is flat earth rotational matrix in relation with hexacopter body frame, g [m/s] gravitational force, \u03c6, \u03b8, \u03c8 [rad/s] are roll, pitch, and yaw in consecutive [1]. Thrust is the power that caused the hexacopter to fly. Thrust is produced by speed of each rotor so that it can generate enough power to lift the hexacopter. To maintain the stability and keep the total lift of generated thrust, the rotational speed of one rotor always lower as much as the higher speed of the opposite rotor in terms of roll, pitch, and yaw motion. Because of the fixed rotor, therefore the lift will always attract upward along the Z axis of hexacopter body frame. Thrust Fthrust while hovering can be approximated as\nFB thrust = \u03a3 (kf \u03a9rotor), 0 \u2264 \u03a9rotor \u2264 1 (3)\nwhere kf is thrust maximum generated by hexacopter and \u03a9rotor is the rotor rotational velocity represented as control signal of each rotor with simple low pass filter for modeling the rotor response time in initial acceleration and deceleration condition. Kf is obtained from rotor datasheet. Approximation of air resistance FB\nar is performed with the following calculation\nFB ar = \u2212Cdrag VB |VB| (4)\nwhere Cdrag is drag force constant obtained from hexacopter geometry. Acceleration and deceleration of six independent rotor will produce torque along the X, Y, and Z axis, creating rotational motion of roll, pitch, and yaw. Torque is the force generated by speed of the rotor multiplied by the arm length that will affect the total rotation on a particular axis according to the distance of the center of mass to each rotor. Torque is given as follows\n\u03c4B = \u03a3 (kt \u03a9rotor), 0 \u2264 \u03a9rotor \u2264 1 (5)\nwhere kt is maximum torque generated by hexacopter. Each rotor that rotates clockwise must be balanced with the other rotor spins counter-clockwise to counter the torque generated so that the hexacopter can fly stably [8].\nB. Selection of Hardware The hardware chosen as flight controller is Pixhawk. Pixhawk has several advantages compared to other flight",
+            "controllers including the use of the 32 bit of primary and backup processors, larger memory, equipped with 10DOF IMU sensor, lightweight and small dimensions, and also affordable price.\nC. Software Design The software that is chosen to simulate HILS is jMAVSim, because it has the ability to visualize the hexacopter model and already been using MAVLink communication protocol. QGroundControl (QGC) is used as the GCS software that is also using MAVLink communication protocol. Therefore, the UAV flight characteristic such as attitude and position can be monitored directly by the operator through the GCS PC.\njMAVSim hexacopter parameters adjusted to real hexcopter parameters to gather the actual representation of the model. jMAVSim works by receiving all the MAVLink packets transmitted by the Pixhawk through serial UART. To run the simulation, HILS need to gather specific data structure from the MAVLink packets which are HIL_CONTROLS that serve as rotor speed, HIL_SENSOR that serve as IMU sensor value, and HIL_GPS that serve as GPS value to obtain position relative to the earth.\nD. Hexacopter Control\nUAV movement is controlled by regulating yaw, pitch and roll of the aircraft. The resulting rotor speed will generate the lift needed to make the UAV move upward or downward. While the speed differences of the rotor will produce a force\nthat can change the attitude and direction of the aircraft. The hexacopter can be activated to do autonomous mission by using remote control command or GCS command and can changed back to manual control by the same way.\nThe block diagram then implemented on the Pixhawk hardware.\nIII. HILS IMPLEMENTATION AND TEST\nThe UAV platform used to perform monitoring mission of chemical contamination is Tarot 680PRO hexacopter. The parameter of the hexacopter is as follows.\nPixhawk and jMAVSim will communicate to each other through serial port with MAVLink protocol. jMAVSim then forwards all the MAVLink packets to the QGroundControl software through UDP port 14550 and vice versa. Operator can enter commands in the form of waypoints through QGC to run the mission. The data is then sent to the jMAVSim for the next process which is combining to MAVLink packets and send back to Pixhawk."
+        ]
+    },
+    {
+        "image_filename": "designv11_60_0001235_b978-0-08-099425-3.00005-4-Figure5.18-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0001235_b978-0-08-099425-3.00005-4-Figure5.18-1.png",
+        "caption": "FIGURE 5.18",
+        "texts": [
+            " It is generally sufficient to treat the tyre as a linear spring and damper when computing the vertical force component, notwithstanding the reservations expressed in Chapter 4 about hysteretic damping and ways to model it. The tyre is quite lightly damped and in any case in the running vehicle its motions are dominated by wheel hop, so the use of an equivalent viscous damping term with the equivalence point taken at wheel hop is satisfactory for simple handling models. The calculation of the vertical force Fz acting at point P in the tyre contact patch has a contribution due to stiffness Fzk and a contribution due to damping Fzc. These forces act in the direction of the {Zsae}1 vector shown in Figure 5.18 Fz = Fzk + Fzc \u00f05:30\u00de Fzk = -kz \u03b4z \u00f05:31\u00de Vertical tyre force model based on a linear spring damper. Fzc = -cz VZ \u00f05:32\u00de where km2.0c ztz \u22c5= \u03b6 \u00f05:33\u00de and mt \u00bc mass of tyre kz \u00bc radial tyre stiffness z \u00bc radial damping ratio dz \u00bc tyre penetration Vz \u00bc rate of change of tyre penetration A linear model of tyre vertical force may need to be extended to a nonlinear model for applications involving very heavy vehicles or studies where the tyre encounters obstacles in the road or terrain of a similar size to the contact patch or smaller"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0001768_a:1008839925389-Figure4-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0001768_a:1008839925389-Figure4-1.png",
+        "caption": "Figure 4. This figure shows the phase plane of the onset of rhythmic movement for any position state variable and its derivative (segment OA), and the stopping of rhythmic movement (segment BO).",
+        "texts": [
+            " (8) by a set of independent vectors, N , such that the \u2202C \u2202W T 0 term is eliminated. N \u2202CT \u2202W = 0 (10) NG(Ws) = \u2212 N \u2202LT \u2202W F (11) with F as computed, 0 is derived from Eq. (9). P1: KCU/RKB P2: STR/SRK P3: STR/SRK QC: Autonomous Robots KL465-03 July 1, 1997 15:43 248 Jalics, Hemami and Clymer The muscle forces, F , in standing have two functions: 1. to render the state, Ws , to be an equilibrium, and 2. to guarantee stability through the use of coactiva- tion. The phase plane of a rhythmic movement is shown in Fig. 4. The starting phase is portrayed by segment OA. In this segment, the system is moved to its rhythmic trajectory in a short time. The starting phase can be implemented by reflex action (Herman et al., 1987), or also by toe-off, i.e., by the action of the foot. The system is given a burst of power to start the walk. Since the biped considered here has no feet, this starting phase is abstracted by application of two short impulses in the horizontal and vertical direction at both feet that would give the system an initial velocity in zero time"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0003019_iros.1996.568948-Figure4-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0003019_iros.1996.568948-Figure4-1.png",
+        "caption": "Fig. 4 Grasp with five frictionless contacts",
+        "texts": [
+            " m = 2, n = 4 and the corresponding matrices are -1 -- Jz h A A 0 - 2 0 0 JT=[o 0 2 01 Since rank(W) = 2, rank(JT) = 2 and rank(A) = 4, by using Eq. (6), we obtain Ni = 1, N, = Oand dimrpl = 0, dimrp2 = 2, dimrhl = 2, d imrh2 = 0. Therefore, the contact force is decomposed into f = fp2 + fh1 and can be determined as Fig. 3b shows the meaning of the decomposition, The contact force can also be formulated with respect to the external force and the joint torque Note that, in this case G1 = 0 , which means the grasp can only passively resist the applied external forces. 4.3 Case III: fpl = 0 In the grasp shown in Fig. 4a, the fingers contact the object at five points, m = 2,n = 5. The corresponding matrices are determined as I J = o [ 0 2 4 7 0 0 - 2 0 0 0 In this case, rank(A) = 4 is less than the full rank. Hence, there exists uncontrollable components ofthe contact force. Since rank(W) = 2, rank(J) = 2, from Eq. (7) we have Ni = 1, N, = 0. Substituting these into Eq. (6), we obtain dim rpl = 0 , dimrp2 = 2 , d imrhl = 2 and dim rh2 = 1. Therefore, the contact force is decomposed into f = fp2 + fhl + fh2 and its components can be determined by Eq"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0001129_icrom.2016.7886773-Figure2-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0001129_icrom.2016.7886773-Figure2-1.png",
+        "caption": "Figure 2. Geometric parameters of the 3-RPS parallel manipulator",
+        "texts": [
+            " Where The Jacobian matrix for the Homotopy function is defined as: ( , ) = \u22ef\u2026\u22ee \u22ee \u22f1 \u22ee\u2026 (11) III. KINEMATICS MODEL OF 3-RPS MANIPULATOR Forward kinematic analysis of the 3-RPS parallel manipulators includes finding the position and orientation of the moving platform with respect to the fixed base while the rotation angle of the revolute actuators ( = 1, 2, 3) is given. Obviously, it is equivalent to the calculating of coordinates of the centers of the spherical joints ( = 1, 2, 3), attached to the moving platform, see Fig. 2. With regard to Fig. 2, the reference coordinate frame O{x, y, z} is attached to the fixed base while the z-axis is perpendicular to the plane which is passing through ( = 1, 2, 3). Coordinates of the points ( = 1, 2, 3) can be written as: = ( , , ) (12) Where , and are the known values. The smallest angle between the positive side of the z-axis and the line passing through the leg is denoted by (1, 2, 3). Also ( = 1, 2, 3) is the angle from positive side of the x-axis to the line segment where is the projection of the point on O-xy plane"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0003913_s1474-6670(17)54642-x-Figure2-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0003913_s1474-6670(17)54642-x-Figure2-1.png",
+        "caption": "Fig. 2: The friction area",
+        "texts": [
+            "1) where ~ is the limit friction angle. Therefore, even if N is known, the value of F is unknown, since T is indeterminate. In fact, as indicated in (3.1), T is given at the limit condition only, being dependent, in other cases, on the global force system acting on the body. At any rate, once the contact points have been fixed, it is possible to recognize an area (friction area) on the plane in which F lies: this area is delimited by the two limit conditions for each contact, expressed in (3 . 1) (see Fig.2) (Abel, Holzman, McCarthy, 1985). Considering three forces, the equilibrium can exist only if one point common to the three directions exists. Hence, having fixed three contacts, the equilibrium can be possible if at least one point common to the friction areas of the contacts can be found. This condition is the first that must be respected by the three potential contact points. In general, if three points of the object surface can be contacts (in the above-said sense), a finite area is common to the friction areas of each contact (see Fig. 2). Every point H belonging to A represents, as potential intersection of the direction of the Fi at each contact, a particular solution for the autoequilibrated grasp. Then, choosing H means fixing the direction of the F forces and then the real friction angle at the contacts and the ratio between Ni and Ti . All the possible autoequilibrated grasps corresponding to a given triple of contact points can be found: from the knowledge of the object features, fixing the contacts Pi and a point H \u00a3 A, the action line of F and the current values of ai (friction angle) can be found"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0000487_j.proeng.2016.06.329-Figure1-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0000487_j.proeng.2016.06.329-Figure1-1.png",
+        "caption": "Fig. 1. Illustrating the gear effect for a toe impact.",
+        "texts": [
+            " Nomenclature CoF centre of face b bulge radius CG centre of gravity e coefficient of restitution CGx distance from the CoF to the CG along the \u2018x\u2019 axis X actual carry distance MOIy moment of inertia about the vertical \u2018y\u2019 axis M weighted carry distance \u03c9 angular velocity W weighting parameter R horizontal distance from the CG to the contact point Z deviation r golf ball radius Zmax maximum acceptable deviation \u00a9 2016 The Authors. Published by Elsevier Ltd. This is an open access article under the CC BY-NC-ND license (http://creativecommons.org/licenses/by-nc-nd/4.0/). Peer-revie under responsibility of the organizing co mittee of ISEA 2016 A clubhead\u2019s MOIy refers to its resistance to rotation about its vertical axis. A higher MOIy provides more forgiveness during off-centre impacts with the golf ball. The mechanism behind this increase in forgiveness is due to a phenomenon known as the \u201cgear effect,\u201d illustrated in Fig. 1 [1]. When the clubhead strikes the ball near the toe, the impact generates a moment about the clubhead\u2019s CG, causing it to rotate clockwise (right-handed golfer). During the impact, the clubhead and ball can be thought of as two gears that are meshed together. When two gears are meshed, the contact point on both gears must share the same tangential velocity, vcontact, equal to the radius of the gear multiplied by its angular velocity. Following this analogy, the side-spin imparted on the ball due to the gear effect can be approximated by Eq",
+            " (1), where R is the horizontal distance from the clubhead CG to the contact point (equivalent to CGx), r is the radius of the golf ball, and \u03c9 is the angular speed, or spin. ballc rR lub (1) Given that r is a constant, the side-spin resulting from the gear effect is proportional to R\u03c9club. Increasing the MOIy of the clubhead reduces \u03c9club during off-centre impacts, thus reducing the amount of spin imparted to the ball, which leads to longer and straighter drives. Conversely, increasing R by moving the CG away from the clubface increases the spin imparted to the ball. In Fig. 1, the counter-clockwise ball spin generated from a toe impact causes the ball to curl to the left [2]. This ball flight is known as a \u201cdraw.\u201d A heel impact causes the ball to spin in the opposite direction, and creates a ball flight that moves to the right, known as a \u201cfade.\u201d Due to the forgiving nature of high MOIy, the sport\u2019s governing bodies have limited the clubhead\u2019s MOIy to a value of 5,900 \u00b1 100 g-cm2 [3]. The bulge b of the clubhead is the radius of curvature of the clubface in the horizontal plane",
+            " The bulge indirectly counter-acts the gear effect by altering the normal direction of the impact thus changing the direction of the ball\u2019s initial velocity. For example, the trajectory of a toed ball would start towards the right because of the bulge, and draw back towards the centre of the fairway due to the side-spin generated by the gear effect. The bulge directly counter-acts the gear effect spin by generating an opposing moment on the ball caused by the horizontal contact angle, assuming that the clubhead velocity is parallel to the X axis in Fig.1; this is analogous to the backspin produced by club loft. If the bulge radius is too small, the counter-spin generated can overpower the gear effect, causing the ball trajectory to remain straight or even curl the opposite way. If the bulge radius is too large, the spin generated from the gear effect will be overpowering, causing the ball to curl too much. Selecting the correct bulge radius can be the difference between a driver that is forgiving and one that is not. To simulate golf drives with varying clubhead parameters, an impulse-momentum impact model validated with finite-element analysis [4] is used in conjunction with an aerodynamic ball flight model [5]"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0001235_b978-0-08-099425-3.00005-4-Figure5.16-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0001235_b978-0-08-099425-3.00005-4-Figure5.16-1.png",
+        "caption": "FIGURE 5.16",
+        "texts": [
+            " The existence of run-flat tyres with a braced sidewall suggests this is so, and recent advances in contact pressure distribution measurements have resulted in a proliferation of data supporting the assertion. Despite this, it is still regarded by some as a matter for debate. The local pressures and stresses distributed over the tyre contact patch can be integrated to produce forces and moments referenced to a local coordinate system within the contact patch. Using the SAE tyre axis system the full set of forces and moments are as shown in Figure 5.16. The following section will explain the mechanical characteristics of each force and moment component. The order in which these components are described will be that which most facilitates an understanding of the mechanisms and dependencies rather than following the local order of the SAE tyre axis system. The tractive force Fx and lateral force Fy depend on the magnitude of the normal force component Fz. Hence the normal force is described first. It should also be noted that more than one mechanism will be involved in the generation of each component"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0001129_icrom.2016.7886773-Figure1-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0001129_icrom.2016.7886773-Figure1-1.png",
+        "caption": "Figure 1. The 3-RPS parallel manipulator",
+        "texts": [],
+        "surrounding_texts": [
+            "As we saw in the previous section, there have been several studies that combined local methods with HCM such as Newton-HCM, Secant-HCM. Nor [19] combined local Ostrowski methods with HCM. Ostrowski-HCM is as follows: = \u2212 ( )( ) (9) = \u2212 ( )( ) \u2212 2 ( ) ( )( ) = 0,1,2, \u2026 , \u2212 1 Where = 0,1,2, \u2026 , \u2212 1 and t \u2208 [0, 1]. To solve a system of nonlinear algebraic equations, we have: = \u2212 [ ( , )] ( , ) (10) = \u2212 [ ( , )] ( ) ( )( ) \u2212 2 ( ) = 0,1,2, \u2026 , \u2212 1 Where , and ( , ) are supposed to be vectors with dimensions n\u00d71, = ( \u22ef ) , = ( \u22ef ) , and the ( , ) Jacobian matrix of size n\u00d7n and the ( ) and ( ) are scalars. Where The Jacobian matrix for the Homotopy function is defined as: ( , ) = \u22ef\u2026\u22ee \u22ee \u22f1 \u22ee\u2026 (11) III. KINEMATICS MODEL OF 3-RPS MANIPULATOR Forward kinematic analysis of the 3-RPS parallel manipulators includes finding the position and orientation of the moving platform with respect to the fixed base while the rotation angle of the revolute actuators ( = 1, 2, 3) is given. Obviously, it is equivalent to the calculating of coordinates of the centers of the spherical joints ( = 1, 2, 3), attached to the moving platform, see Fig. 2. With regard to Fig. 2, the reference coordinate frame O{x, y, z} is attached to the fixed base while the z-axis is perpendicular to the plane which is passing through ( = 1, 2, 3). Coordinates of the points ( = 1, 2, 3) can be written as: = ( , , ) (12) Where , and are the known values. The smallest angle between the positive side of the z-axis and the line passing through the leg is denoted by (1, 2, 3). Also ( = 1, 2, 3) is the angle from positive side of the x-axis to the line segment where is the projection of the point on O-xy plane. Considering legs of the manipulator as three 3D lines, the parametric equations of the lines can be expressed as: : = += += + : = += += + (13) : = += += + In which = sin( ) cos( ) = sin( ) cos( ) (14) = cos ( ) And , and are independent variables. Geometric constraints due to the manipulator architecture can be expressed as follows: ( \u2212 ) + ( \u2212 ) + ( \u2212 ) = ( \u2212 ) + ( \u2212 ) + ( \u2212 ) = (15) ( \u2212 ) + ( \u2212 ) + ( \u2212 ) = Where = ( = 1, 2, 3). Introducing Eq. (2) into Eq. (3) Leads to + + = 0 + + = 0 (16) + + = 0 In which: = \u2206 , = \u2206 + \u2206 , = \u2206 + \u2206 + \u2206 = \u2206 , = \u2206 + \u2206 , = \u2206 + \u2206 + \u2206 (17) = \u2206 , = \u2206 + \u2206 , = \u2206 + \u2206 + \u2206 And \u2206 ( = 1,\u2026 , 18) are the coefficients which depend on the geometric parameters of the manipulator, see appendix A. Thus, we can write the Homotopy continuation function as follows: = ( + + ) \u00d7 + (1 \u2212 ) \u00d7 (18-a) = ( + S + ) \u00d7 + (1 \u2212 ) \u00d7 = 0 (18-b) = ( + P + ) \u00d7 + (1 \u2212 ) \u00d7 = 0 (18-c) We solve Equations (18a)-(18b)-(18c) by Ostrowski-HCM method and change the Homotopy parameter t from 0 to 1. Finally, S, P and T are calculated (table 2), and by substituting them in equation (13) the coordinates of the points Bi (i=1, 2, 3) are calculated (table 1). Direct position kinematics of the 3-RPS parallel manipulator is solved with the followings values, which are taken from [22]: = (0, 0, 0) , = (25, 0, 0) , = (12.5, 21.65, 0) = 20 ( , = 1, 2, 3) , = 25 , =30, =150, =270 \u2206 ( = 1,\u2026 , 18)are calculated based on the values of above parameters. (See appendix A) We choose the initial guesses of unknown parameters as: ( , T, S) = (\u2212500 , \u221250, 2) By adjusting the auxiliary Homotopy function appropriately, seven sets of roots for this problem will be obtained without any divergence. The auxiliary Homotopy functions of these results are given in Table 2. In the simulation, the criterion used for measuring the iteration number of Ostrowski method is the following term: ( ) \u2212 ( )< \u2212 Fi is obtained by substituting the solutions in equation (15). The simulation is done using MATLAB program, the Intel\u00ae Core\u2122 i5, and 4GB RAM in 64-bit Operating system. The results of Ostrowski-HCM in comparison with NewtonHCM method, not only shows less computational time; but also, it provides a higher accuracy. This method reduces the computational time for solving direct kinematic problems of the proposed manipulator up to 97 %, (Table 2), in comparison with Newton-HCM method. Hence, this approach can be implemented in real-time applications. In this paper, the direct kinematic analysis of the 3-RPS parallel manipulator has been carried out by Ostrowski-HCM. The analysis shows that there are 7 real solutions for non-linear direct kinematic equations of 3RPS manipulator. Results prove that Ostrowski-HCM method reduces computational time for kinematic analysis of proposed parallel manipulator. Moreover, it is more effective and faster than other iterative methods such as Newton-Raphson and conventional HCM methods. Furthermore, solutions do not depend on the initial guess. Therefore, this method is highly advantageous for kinematic analysis of many parallel robots, especially, in applications in which fastness and accuracy of the method is greatly important such as Real-Time obstacles. However, one of the challenges of this method is to choose the appropriate Homotopy functions to reach the desirable set of solution, which will be investigated in the future jobs. G"
+        ]
+    },
+    {
+        "image_filename": "designv11_60_0003872_6.2001-2433-Figure3-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0003872_6.2001-2433-Figure3-1.png",
+        "caption": "Figure 3. Canard and Nose Flap Pitch Deflection Orientation",
+        "texts": [
+            " From the prescribed flap areas the length and width of flap 2 and 3 were then calculated to maintain an aspect ratio (flap length/flap width) of 2.93, equivalent to that of flap 1. The geometric relations of nose flap 2 and 3 to the Mach 6.0 predicted nose shock, as well as nose flap 1, are shown in Figure 2. While the nose flaps were designed for a single flap deflected in the vertical, windward plane for pitch control, each nose flap was tested in both the single flap and double flap orientation as shown in Figure 3. For the double flap case, two identical flaps (located 90-deg apart) were mutually deflected in the windward plane for the x-orientation. Test Conditions A wind tunnel test has been conducted in the NASA-Langley Research Center 20-Inch Mach 6 Wind Tunnel to acquire aerodynamic force and moment data for a missile configuration with nosemounted canards and flaps. Experimental data were obtained at Mach 6.0 with a test gas of air. Freestream test conditions were maintained at a Reynolds number of 2",
+            " The configuration nomenclature provided for the plot legends is given as: B (missile body, nose+aftbody), C (canards), T (tail fins), NF1 (nose flap 1 - small flap), NF2 (nose flap 2 - medium flap), and NF3 (nose flap 3 - large flap). A nominal moment reference point (MRP) was located 8.836 in. (3.991 calibers) aft of the model nose as shown in Figure 1. The wind tunnel model was instrumented with a 0.75-in., six-component main balance (2019-D) that utilized a water-cooling jacket for temperature stabilization. No boundary layer transition strips were applied to the model or its components during the test. Figure 3 presents the canard and nose flap pitch deflection orientation. As shown a positive pitch deflection provides a positive normal force (up) and positive pitching moment (nose up) about the moment reference point for the nose-controlled missile at 0- deg angle of attack. Data presented for the canards were obtained only for two horizontal canards equally deflected to produce the desired pitch command at 0-deg missile roll angle (^-orientation) as shown in Figure 3. Test data were obtained for canard pitch deflections of 0, 15, and 30 deg at angles of attack from 0 to 15 deg. Data presented for the nose flap were obtained for both the single and double flap orientation. The single flap orientation is defined by a single flap deflected in the windward plane to produce positive pitch (0-deg missile roll angle). The double flap orientation is defined by two flaps equally deflected in the windward plane to produce positive pitch (45- deg roll angle, x-orientation). Test data were obtained for nose flap deflections of 0, 20, and 35 deg relative to the missile centerline for both single and double flap orientations shown in Figure 3. For all cases and roll angles, the canards and flaps were located inline with the tail fins. Data presented are provided in the missile (non-rolled) axis system. Experimental Data Accuracy Force and moment coefficient accuracy for the experimental data were provided-by NASA-Langley as 0.5% of maximum balance loads. Table 1 provides the calculated accuracy of the experimental data expressed in coefficient form for the nominal operating dynamic pressure of 2.1 psia at Mach 6.0. Aerodynamic force and moment data were reduced to coefficient form using a reference length of 2",
+            " It is observed that theoretical predictions of normal force and axial force coefficients tend to offer good agreement with test data for angles of attack up to 6 deg, although theoretical data tend to slightly overpredict the missile trim angle of attack for 20- and 35-deg flap deflections. With increasing angle of attack, predictions of pitch control characteristics agree less favorably with test data, since nonlinearities in the flap control increments due to downwash are not modeled. Figure 10 presents a comparison of missile stability and pitch control for the canards and single nose flaps at Mach 6,0 and <|)=0 deg (+-orientation) with 15-deg canard deflection and 20-deg flap deflection. The orientation of the canards, single nose flaps, and tail fins is provided in Figure 3. Data for each of the three nose flap designs are plotted. Stability data for the nose flap configuration appears as the body+tail case (6f=0 deg), since the flap is conformal with the nose surface when undeflected. Stability data for the canard configuration appears as the body+canard+tail case with four canards undeflected (6C=0 deg). It is noted that the presence of the canards alone (5C=0 deg) provides a 15-20% increase in total missile drag. The nose flaps show increasing pitch control authority and axial force with increasing flap area from nose flap 1 to 3",
+            " It is observed that ACN, ACm, and AC^ for nose flap 2 and 3 correlate well with the flap planform area ratios (i.e. ratio of flap 2 planform area to flap 1 planform area). Presented in Figure 14 is a comparison of missile stability and pitch control for the canards and double nose flaps at Mach 6.0 with 15-deg canard deflection and 20-deg flap deflection. It is noted that the canards are in the H-orientation and the nose flaps are in the x-orientation, both controls inline with the tail fins as shown in Figure 3. The nose flaps show increasing pitch control authority and axial force with increasing flap area from nose flap 1 to 3. Both canard and nose flap configurations provide a very stable airframe over the entire angle-of-attack range. In particular, double nose flap 3 provides comparable pitch effectiveness and trim to the canard-controlled configuration at all angles of attack, although the axial force coefficient is greater. For this case with double flaps deflected in the xorientation, flap AC^ and ACm are shown in Figure 15 to be comparable to those of the canards, more well behaved with less nonlinearities than those observed in Figure 11 for the single nose flaps"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0002235_0924-0136(95)02159-0-Figure3-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0002235_0924-0136(95)02159-0-Figure3-1.png",
+        "caption": "Fig. 3. Profile of the rolled trapezoidal thread.",
+        "texts": [
+            " + [-xl(io~ cos p + 1) + z~iol sin p cos ~o~ + Aiot cos p cos \u00a2p~ - S~io~ sin p sin ~o~]N o, + (-xl io~ sin p cos q~ -y~iok sin p sin ~o~ + Aiol sin It - ~)Nl: = 0 (7) Xo = x~(cos \u00a2Po cos ~o~ -- cos p sin \u00a2Po sin \u00a2Pl) +y~(cos \u00a2Po sin \u00a2p~ + cos l t sin q~o cos \u00a2p~ ) +z~ sin p sin ~o~ -- S~ sin i t sin ~o~ - A cos q~o Yo = -x~(s in ~Po coq q~ + cos p cos q~o sin q~t ) --y~(sin \u00a2Po sin ~Pt - cos tt cos \u00a2Po cos qg~) +z t sin It cos ~Oo + A sin q~o Zo = x~ sin i t sin ~Ol +y~ sin p cos ~o~ -~-'~1 COS fl - - S I c o s It The solution of this system (7) at Yo = 0 gives the axial profile, and at Zo = 0 gives the frontal profile, of the shaping surface of the rollers. As the equations in the system are transcendental, its solution is achieved through excluding one of the three parameters \u00a2p~, p~ or 0t and applying Newton's method for solving a system of aon-linear equations [5]. 3. Profi|ing of rol|ers for trapezoidal thread The developed mathematical model is applied to the profiling of rollers for the rolling of trapezoidal thread Tr24 x 20 (P5) [5] (Fig. 3). The solution of the problem is carried out in the following sequence: (a) The defining of the equation of the given surface of the part. In this case, the surface is Archimedean, and the equation of its right profile in parametrical form is: x~ = Pt cos 0 i y~ = p~ sin 0~ zl = (Pl -- P3) cotg7 +pO~ where P3 = (d3/2)=9.25 mm is a minimum radius profile; p = (S/2rc)- 3.183 mm is a screw parameter; 7 = (zt/2) - (~/2) = 1.308997 rad; S = 20 mm is the lead of the thread; 0 is an angular parameter; and p is a parameter"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0003347_s0022-5193(87)80219-9-Figure9-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0003347_s0022-5193(87)80219-9-Figure9-1.png",
+        "caption": "FIG. 9. The coordinate system used for a right limb. For simplicity, the intercalary region is approximated by a cylinder of length L and radius R. Its origin is at the centre of the limb with the z-axis aligned along the proximo-distal (PD) axis and increasing towards the distal. The x-axis is aligned along the AP axis and increases towards the anterior; the y-axis is aligned along the DV axis and increases towards the dorsal. Points on the limb surface are specified by the cylindrical coordinates z and ~b. z = 0 corresponds to the middle of the intercalary region. For a left limb, a similar choice is made except the y-axis increases towards the ventral.",
+        "texts": [
+            " This is the qualitative basis for thinking of the ipsilateral supernumerary as a symmetry breaking bifurcation. F I E L D M O D E L O F S U P E R N U M E R A R Y P R O D U C T I O N 427 Although the motivation for viewing the production of ipsilateral supernumeraries as.a bifurcation in the limb field came fron treating 0 as a \"control parameter\", it has turned out to be more tractable in solving the model to set 0 = 180 \u00b0 (corresponding to APDV inversion) and look for bifurcations as the length L (see Fig. 9) of the intercalary region is varied. We will see in Section 4(B) that as L is varied and 0 is held fixed in our model, bifurcations do indeed occur giving rise to supernumeraries. Although one might have expected solutions representing supernumeraries to be less likely than those not representing supernumeraries because of the greater field gradients associated with the supernumerary pattern, it turns out that as L increases the field gradients become small enough that solutions representing supernumeraries are possible",
+            " F I E L D M O D E L O F S U P E R N U M E R A R Y P R O D U C T I O N 429 (B) T H E S T U M P A N D B L A S T E M A B O U N D A R I E S The field t~ associated with the intercalary region is assumed to develop so as to adjust to the stable field values o f the b o u n d a r y tissues. We now want to obtain the analytic expressions for the bounda ry fields. Since these boundar ies have the normal limb pattern, we first consider the form of the field for a normal limb and then apply this to the boundar ies o f the graft ing experiments. The limb, for the purposes o f our model is described in terms o f a mapp ing between biological and spatial coordinates. The spatial coordinates are illustrated in Fig. 9. It has been assumed for simplicity that the limb surface is a cylinder o f fixed radius R descr ibed by 'cylindrical coordina tes (~b, z) where q~ is defined as shown in Fig. 9. A por t ion o f the surface o f the normal limb indicating the field vectors at several points is illustrated in Fig. 10(a) using the p lanar representat ion o f the l imb's surface. For the normal limb along lines o f constant ~b, the vectors have the same direct ion since the axial values are the same but they decrease in magni tude distally. As is evident f rom the figure, for a normal limb ~(~b, z) varies smooth ly over the surface. A useful ana logy for the field 6(th, z) is to think o f it as a t w o - c o m p o n e n t local order parameter similar to the average spin o f an Ising model (Ma, 1976, p",
+            " This ensures that there is no bunching or spreading out of the field values (contrast with Bryant & Iten, 1976, Figs 20, 21 and 22) which is necessary for the model in its present form to be consistent with the ipsilateral and contralateral data. The values of the field of a normal limb at a given proximo-distal level are thus u = m cos ~b, v = rn sin ~b (1) for a right limb (the handedness of the limb of Fig. 10(a)) and u = m cos ~, v = - m sin ~b (2) for a left limb, where [t~ I is denoted by rn. m is a monotonically decreasing function of z for the normal limb. We now focus on the intercalary region and its boundary fields. For convenience, a similar spatial coordinate system is used as illustrated in Fig. 9. The field 6 over the surface of the intercalary region is again a function of the cylindrical coordinates (~b, z), with the mid-point of the intercalary region at z = 0 and with the stump and blastema boundaries at z = - L / 2 and z = +L/2, respectively, where L is the length of the intercalary region. For definiteness, we consider the stump in the grafting experiments to be a right limb. Under these assumptions, the field values of the s tump boundary are again given by equation (1). Unless indicated otherwise, the magnitude of ~ at the stump will be normalized to 1 so that the stump boundary is given by u = cos ~b, v = sin ~b",
+            " Thirdly, the expectation that supernumerary production involves bifurcation (see Section 2(E)) is most easily satisfied by including the non-linear fourth order term in 161. The field equations constitute the phenomenological rules governing the behaviour of the field ~ and therefore the pattern formation process. In deriving these equations, we assume that the various time scales (release of cells, cell division and migration, pattern formation, redifferentiation and outgrowth) are such that the intercalary region can be treated effectively as having a fixed length L between non-moving boundaries (see Fig. 9). Starting with some arbitrary initial field 6 over the intercalary region, we require that the field adjust so as to de6rease the free energy F. A stable steady state is reached when F is a local minimum. It can be shown (Totafurno, 1985, pp. 105-106) that the time development of 6 is such that F{t~} decreases with time if t~ = (u, v) satisfies the equations I ~.2au= V~u + au -/3u(u2+ 02) at (9) \u00a22 av l -~ = v % + a v - / 3 v ( u z + v 2 ) where 1 8 2 a 2 V 2~ R2 a(b-------- ~ ~ az 2 (10) and z, a and/3 are constants"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0002030_20.560098-Figure9-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0002030_20.560098-Figure9-1.png",
+        "caption": "Fig. 9. (a) Current waveform. (b) Flux density waveform. (c) Hysteresis loop.",
+        "texts": [
+            " a path like VWX, as shown in Fig. 8. If the core is magnetized to the level corresponding to the point U , a small value of H is applied and varied between positive and negative values around a small loop like US. The curves are characterized by \u201cminor\u201d loop or \u201csubsidiary\u201d (VX and U S ) hysteresis loops [12]. If a sinusoidal flux density waveform is applied to the test core with a voltage large enough to saturate the core, the sequence of events occurring during the time intervals of one cycle of the waveform shown in Fig. 9 can be explained as follows. 1) During the time intervals ( t l - t2 ) and (t3-t4), the core is unsaturated, and the current flowing in the winding around the test core [Fig. 9(a)] causes the field intensity H to change with time in a manner similar to its change with time. The flux density during time interval t l - t 2 represents the ascending path of the hysteresis loop, and the flux density during the interval t3-44 represents the descending path. 2) During the time intervals ( t 2 - t 3 ) and (t4-t5), the flux density B has its maximum value which causes the core to be saturated. The field intensity H changes, depending on the current, during these intervals. The result of applying a nonsinusoidal excitation waveform can be explained as follows"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0001021_ichve.2016.7800914-Figure2-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0001021_ichve.2016.7800914-Figure2-1.png",
+        "caption": "Fig. 2. Typical working position on suspension tower",
+        "texts": [],
+        "surrounding_texts": [
+            "Keywords\u2014 1000kV; UHV; live working; compact transmission line; minimum approach distance\nI. INTRODUCTION The 1,000kV UHV AC demonstration project of southeastern Shanxi-Nanyang-Jingmen has been put into commercial operation since the end of 2008. According to the\nTwelfth Five-Year Guideline of State Grid Corporation of China, the three vertical and three across UHV backbone net rack that connect the large-scale energy base and the main load center will be built in the following years. At that time, there will be strong smart power grids which include the UHV backbone net rack and all levels power grid. [1]Live working is an important technical means to ensure the power grid safety and reliability operation; the live working technology has been researched and applied for more than fifty years which has a mature technical system for live working on these lines below 500kV and developed a complete live work standardization system in China[2-9]. Recent years, the research achievement on live working projects such as AC 750kV, DC 660kV 800kV have stepped into popularization and application[10-15]. There\u2019s a correlation research at the beginning of the UHV demonstration project, it provided basic parameters for the\nexternal insulation designing of operation gaps for the tower window. In April 2008, the live line worker entered the 1000 kV equal potential conductors successfully in Wuhan UHV test base which marked that China has the ability to conduct live working on UHV live working[16-20].\nHowever, there\u2019s a significant difference between the towers structure of UHV compact transmission line and demonstration project. The three-phase conductor of UHV compact transmission line has a compact arrangement and there\u2019s no component between phase-phase, which proposes a higher requirement for the safety of live working. The tower of UHV compact transmission line is high whose tower windows are large. There are no related reports about the key technology investigation about live working for UHV compact transmission line in both China and aboard. So, the live working parameter of EHV / UHV transmission lines can\u2019t be applied to the live working for UHV compact transmission line.\nBased on the structure of suspension and tension towers for UHV compact transmission line, conducts the impulsive discharge tests of air gaps for typical live working condition were conducted , the discharge characteristic were obtained and the minimum minimum approach distances were calculated in this article. It can provide the parameter for external insulation of the lines\u2019 tower window and the live working.\nII. TEST METHOD The test was conducted in the UHV outdoor testing field of State Grid. The test equipment: 5,400kV, 527kJ impulse voltage generator (figure 1-a); 5,400kV low damping cascade resistancecapacitance divider; 64M type peak voltmeter and Tek TDS340 oscilloscope. The measuring system was checked by the national high voltage metering station, the total uncertainty of the measured voltage peak is less than 3%. There\u2019s another set of 3,000kV mobile impulse voltage generator (figure 1-b) in the testing field that activate synchronized with the 5,400kV impulse voltage generator and output the impulse voltage, which was used to conduct the phase-phase switching impulse voltage discharge test together with the 5,400kV impulse voltage generator. The test adopted the the up-and-down methods to apply impulse voltage, during the phase-ground switching impulse voltage discharge test, a +250/2,500 s standard switching impulse wave was applied to the operating phase conductor and the other phase conductors were grounded; While in the test of phase-phase switching impulse voltage discharge\n978-1-5090-0496-6/16/$31.00 \u00a92016 IEEE",
+            "test, a switching impulse wave of +250/2,500 s and - 250/2,500 s with a form factor that =0.4 was applied to the phase conductor, the other phase conductors were grounded.\nb 3000kV impulse voltage generator According to engineering economy and electromagnetic environment control condition, a 12\u00d7400mm2 conductor is proposed to use in the 1,000kV UHV compact transmission line ,whose bundle space is 350mm , bundle circle diameter is 1.35m and conductor overhang angle is 10 degree (consider the conductor is sagged at the suspension position of tower by its own weight, and formed an angle with horizontal plane). According to the requirement, the test adopted 12-bundles subconductor as model to make the simulated conductor of stainless 25mm steel tube.\nThe dummy man made of aluminum alloy was applied to simulate the live line works at actual working condition in the test, which had the same form and structure with actual electric workers whose legs and arms could bend freely so that it adjusted all kinds of postures based on typical live working conditions. The dummy man was 1.8m high and 0.5m wide in standing posture.\nIII. TESTS ON AIR GAPS FOR LIVE WORKING The air gaps for live working discharge tests were concentrated on the switching impulse discharge test in gaps during live working process,which was applied to checked whether the gaps of all working positions could satisfy the live working safety rules.\nBased on the live working research results of China and abroad, when the live line workers at the conductor (equal potential), the switching impulse 50% discharge voltage (U50) to the tower are lower than that he in the tower truss (ground potential) to conductor. So, the air gaps discharge test for live working on compact UHV alternating current transmission line, only considered the equal potential working condition. All the test data were amended to standard meteorological condition in accordance with GB16927.1-1997 in this article [21].\nA. air gaps discharge tests for live working on suspension tower\nThe minimum approach distance discharge test for the typical live working condition in suspension tower included three kinds of discharge gap: equal potential worker to the upper cross arm (position 1), equal potential worker in the upper phase conductor to the side tower frame (position 2), equal potential worker in the low-phase conductor to the upper phase conductor (position 3). The dummy man wore the full set of shielding clothes in the test and adopted the stand posture in position 1 and 3, the sitting posture with his face to the conductor and back to the tower truss in position 2 as shown in Fig.3, which kept equal potential with the simulated conductor. When the gaps distance was adjusted completely, conduct the switching impulse voltage discharge tests were conducted and discharge paths were observed by the high speed camera meanwhile.\nAccording to the result of minimum approach distance discharge tests for live working in a suspension tower, the discharge voltage U50 in position 1 was obviously lower than that in position 2 at the same gap distance. Therefore,when the equal potential worker was in the upper phase conductor with his head to the upper frame, the discharge voltage U50 was the lowest.\nObserving the discharge path by high speed camera, when the distances between the simulated conductor and the upper frame or the side frame of the tower were equal, all the discharge happened between the dummy man\u2019s head and the upper frame in test of position 1, while most discharge paths happened between the dummy man\u2019s head or grading ring of conductor and upper frame in the test of position 2. At this time, in order to make sure that the discharge generated between dummy man a d side frame of tower, test arrangement was adjusted by making the distance of the simulated conductor to upper frame be 1.0~1.5m bigger than that to the side frame of the tower, and then the discharge were realized at the side frame. But there were still some discharge to the upper cross arm. When the phase-ground distance was 1.0~1.5m bigger than the phase-phase distance, most of the discharge paths were between the two testing phases in the test of position 3. By this time, due to the limitation from the synchronicity of applied positive and negative polarity switching impulse wave for the two phases and field test arrangement, some discharges still happened between phase and ground.\nCompared the discharge characteristics and discharge path results of the three working conditions, it was concluded that the most dangerous live working condition for worker in upper",
+            "phase with equal potential. Therefore the equal potential worker should control the human body\u2019s position in operation.\nThe minimum approach distance for tension tower (angle tower) was mainly concentrated on the gap discharge characteristics when the worker between the conductor in the tension string terminal and the cross arms. ,The discharge voltage U50, the equal potential worker\u2019s discharge to tower frame was obviously lower than those high voltage conductor to the ground potential worker , which were similar as the suspension tower minimum approach distance discharge feature. So, discharge tests for the tension tower only considered the condition that the worker located at the tension string close to the conductor.\nThe minimum approach distance discharge test arrangement for tension insulator string was shown in Fig.4 . The dummy man wore shielding clothes by using striding no more than two insulators method at bend-down posture and squat at the linking position of the tension string and the conductor grading ring. The XZP-300 insulators with structure height 195mm was selected, and the width of the dummy man\u2019s shoulders was 0.5m. The witching impulse discharge tests were conducted by adjusting the distance between cross arms and dummy man .\nTABLE I. TEST RESULTS OF MINIMUM APPROACH DISTANCE FOR A SUSPENSION TOWER\nTABLE II. TEST RESULTS OF MINIMUM APPROACH DISTANCE FOR A\nTENSION INSULATOR STRING POSITION 4\nInsulator number\nGap distance /m\nU50 kV\nStandard deviation [ d] %\n25 4.4 1606 5.7\n30 5.4 1781 4.4\n35 6.3 2001 4.8\n40 7.3 2212 5.9\n45 8.3 2304 5.0\nTABLE III. MINIMUM MINIMUM APPROACH DISTANCES\nThe test results in TABLE II were compared with the minimum approach distance test results for the tension tower of UHV regular lines. It indicated that discharge characteristics for the tension tower of the two different UHV transmission lines were almost the same due to the towers having the similar structure designing. The discharge path is between the dummy man in high potential and cross arms of a tension tower.\nIV. MINIMUM MINIMUM APPROACH DISTANCES Minimum minimum approach distance means that the required distance of phase-ground or phase-phase to ensure the safe of the worker in live working. According to GB/T191852003 calculation method of minimum approach\nFig. 3. Discharge characteristics of minimum approach distance for a suspension tower\nFig. 4. Test layout of minimum approach distance for a tension insulator string position 4"
+        ]
+    },
+    {
+        "image_filename": "designv11_60_0003035_dftvs.1994.630030-Figure6-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0003035_dftvs.1994.630030-Figure6-1.png",
+        "caption": "Figure 6: Compensation paths for PE1 and PE2.",
+        "texts": [
+            " the distribution of the compensation paths covering all the faulty processors such that the paths satisfy the conditions in Theorem 3, the maximum independent graph theory in [9] is now extended to 3D. A completely connected graph with 6 vertices is formed per faulty P E to represent the six possible compensation path directions from the PE. Then inter-graph edges are drawn between two vertices that belong to two different graphs. An inter-graph edge represents directions of compensation paths that violate the conditions in Theorem 3. For example, in Figure 6 let U; be a vertex in the graph for a faulty P E denoted by PE1 and vj be a vertex in the graph for another faulty PE, P E * , then an edge < v;,vj > represents the fact that a Reconfiguration Techniques 201 compensation path passing in direction v; through PE1 and a path passing in direction v j through PE2 will violate a compensation path condition given in Theorem 3. Specifically, if PE1 has a north compensation path, then PE1 cannot have a south compensation path since they are in near-miss",
+            " Also < N, 2, > are intersecting, < 2-, Z+ > are in near-miss and < 2-, S > are intersecting. The maximal set of independent vertices, that is vertices that are not connected, is then found. Since each graph is completely connected, one vertex from each graph is chosen. If the number of vertices so found is equal to the number of faulty PE\u2019s then the fault distribution is said to be reconfigurable. The directions that should be assigned to a faulty PE is given by the direction represented by the vertex chosen from the graph of the faulty PE. For example in Figure 6 some of the paths allowed are {S, E}, {N, 2-} and {Z+, N}. It is well known that the solution to the maximum independent set problem is NPcomplete. In [15, 141, the authors give a polynomial time algorithm in which the routing problem in the 2D 1; model is treated as the problem of finding non-intersection straight lines from faulty PE\u2019s (vertices in a grid) to the the boundary of the grid. The extension of the polynomial routing algorithm to 3D is not obvious and is the subject of future research"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0003173_s0141-6359(02)00117-4-Figure15-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0003173_s0141-6359(02)00117-4-Figure15-1.png",
+        "caption": "Fig. 15. Finite element analysis of moment-carrying coupling made of aluminum.",
+        "texts": [
+            " \u03c6 = \u03c0 2 \u2212 \u03b2 (22) Therefore, \u2223\u2223\u2223tan (\u03c0 2 \u2212 \u03c6 )\u2223\u2223\u2223 \u00b5s > 1 (23) \u00b5s > |tan(\u03c6)| (24) For a given \u00b5s, the maximum value of the inclination angle is given by \u03c6 < tan\u22121(\u00b5s) (25) Finite element models were created and evaluated using ProENGINEERTM and ProMECHANICATM. Torsional loads were applied as point loads. Note that ProMECHANICATM automatically distributes a point load over a small circular area; this area is small relative to the size of the components, and does not significantly change our results. Initial models were done on single fingers or finger pairs to save computation time. After these models demonstrated the feasibility of the idea, models of entire couplings were constructed and analyzed. Sample results are shows in Figs. 15 and 16. Fig. 15 is a plot of displacement magnitude in an aluminum model of the perpendicular adjustment moment carrying variation; displacement is expressed in units of milli-inches. Fig. 16 shows strain energy for the three-dimensional printed perpendicular adjustment non-moment carrying variation. These figures show good strain and stress distribution indicating a robust design. Physical models were tested to validate analytical and finite element models of torsional stiffness. Fig. 17 shows a test fixture that was constructed to permit the testing of three-dimensional printed couplings, as well as an aluminum model fabricated specifically for the tests"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0003547_0954406001523074-Figure8-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0003547_0954406001523074-Figure8-1.png",
+        "caption": "Fig. 8 (a) Involutes arrived at as in Fig. 7, by tracing points as they move down the conjugate line. The PAT-N line",
+        "texts": [
+            " The meshing involutes are tangent to each other and perpendicular to the conjugate line where they cross that line. Where the base line extends outside the pitch line, there are reversals in the involutes, which indicates one of the many practical problems. By coincidence, in Fig. 7, the angle of rotation at which the pitch outlines are shown, the conjugate line, appears to meet both the tangent and normal conditions (E and N). This condition occurs twice per cycle, and Fig. 7 shows just one of them. Figure 8 is similar to Fig. 7, except that here a pressure angle PAT-N of 45 was used. That is, the line of action made a \u00aexed angle of 45 to the tangent at the down the conjugate line from one base outline to the other. The points are traced on rotating coordinate systems \u00aexed in each gear centre. The PAC-N line of action and base outline of the second-order ellipse as in Fig. 4c are used. (b) Enlargement of the region about the pitch point, showing that normal conjugate action is taking place between the opposite-handed involutes of action and base outlines of the same second-order ellipse shown above is used",
+            " (b) Enlargement of the region about the pitch point, showing that normal conjugate action is taking place between the oppositehanded involutes Proc Instn Mech Engrs Vol 214 Part C C02197 \u00df IMechE 2000 at OhioLink on November 7, 2014pic.sagepub.comDownloaded from contact point. This angle, larger than that used for Fig. 5c, shows that the base outline can be contained within the pitch outline if a su ciently large pressure angle is used. In a coordinate system in which the pitch point is stationary, the instantaneous line of action would rotate about a mean direction, which here would be at 45 to either the X or Y axis. Figure 8, like Fig. 7, shows involutes that meet the requirements of Buckingham's basic law for teeth pro\u00aeles (Sections 3.1 and 3.2 above). It can be seen that, in the region enlarged in Fig. 8b, it would be easy to make excessively long teeth so that they would clash with their opposite numbers. It is also clear that, where the base line comes close to the pitch line, it would be possible easily to cut teeth deeper than the base line. The involutes here are in principle the equivalent of those generated by rolling a rack around the pitch pro\u00aele, but it is clear that the addendum and dedendum vary considerably around the pitch outline. Note that, if such teeth were cut with a rack or hob in the manner that has been described in the literature (Section 2), they might indeed function incorrectly"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0000878_978-3-319-47614-8-Figure2.5-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0000878_978-3-319-47614-8-Figure2.5-1.png",
+        "caption": "Fig. 2.5 Force on the piston",
+        "texts": [
+            " This load is from the pressure of the wind on the roof which therefore is perpendicular to the roof as shown in Fig. 2.3. This is net load shown by integrating the pressures over the surface of the roof. Another example of imposed loads is a weight being carried on a crane that is being hoisted from one place to another see Fig. 2.4, that we commonly see in a workshop. This load acts on the beam of the crane. Yet another familiar example is the piston of an internal combustion engine on which we have a pressure of the burnt gases from combustion, see Fig. 2.5. The pressures integrated over the area of the piston give the load or Force as shown. These forces or loads on a structure cause them to deform and in this process of deformation we determine the effects of deformation (strain and stress) that allow us to design the structure to be safe. The design of the structure to determine the sizes of a chosen material will be the subject matter of ensuing design courses. The design of any structure stationary, rotating or moving in air, water, or in vacuum (space) is dependent on the state of stress, steady or unsteady",
+            " Assume that the body forces can be represented by a vector at the point under consideration and that it can be resolved into components along the directions of the coordinate system chosen. 2:6. Extend the state equation derived in problem 5 above and obtain the three equations of equilibrium written for the arbitrarily chosen the frame. Comment on the number of equations, three in this case, and the number of unknowns, (six or nine?), in the stress tensor. 2:7. Consider the connecting rod in Fig. 2.5 of this chapter, at one crank angle; this rod can be simplified to be carrying 1000 N force. Midway the length of the rod, the cross-sectional area A is 8 cm2, determine the average stress in this cross-section. Comment on this value of average stress as compared to the stress tensor at a point, keeping in view the question 6 above. 2:8. The stress components at a point in a structure are given as 7 1:5 0 1:5 40 3 0 3 0 2 4 3 5 MPa. Find the components of the surface traction vector for an interface having a normal vector~m \u00bc 0:10~i\u00fe 0:3~j\u00fe 0:8~k"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0003296_10402000208982529-Figure1-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0003296_10402000208982529-Figure1-1.png",
+        "caption": "Fig. 1--Coordinate system and clrcumterential grooved Journal bearing.",
+        "texts": [],
+        "surrounding_texts": [
+            "A Study on Nonlinear Frequency Response Analysis of Hydrodynamic Journal Bearings\nWith External ~isturbances~ BYOUNG-HOO RHO and KYUNG-WOONG KIM\nKorea Advanced Institute of Science and Technology (KAIST) Department of Mechanical Engineering\nTaejon, 305-701, Korea\nThe rzonlir~eur vibration characteristics of hydrodynamic jour-\nnal bearings with a circumferential groove are analyzed numeri-\ncully in case tlrat esternal sinusoidal disturbances are given to the\nrotor-bearing system. The cavitarion algorithm implementing the\nJukohssot~-Floberg-Olsson boundary condition is adopted ro pre-\ndict cavitation regions in the fluid film. The comparison of the\nsteady state journal orbits using linear dynanzic coeficients with\nthe trrrnsietzt motion calculated from nonlinear analysis is per-\nformed. The fieqrtency response functions obtained by linear and\n~zotzlitzear analysis are also presented. When an e.rterna1 sinu-\nsoidal disturbance is applied to the bearing, the range of the full\nfilnz region varies periodically and it becomes snzaller than that of\nstatic eqrrilibrirtn~ state. The d~fference between linear and nonlin-\near analysis increases as the excitation amplitude or the frequen-\ncy of the esternal disturbance increases.\nPresented a s a Society of Tribologists and Lubrication Engineers Paper at the STLEIASME Tribology Conference in San Francisco, CA\nOctober 21-24, 2001 Final manuscript approved September 30,2001\nReview led by Alan Palazzolo\nKEY WORDS\nHydrodynamic Journal Bearings; Frequency Response Function\nINTRODUCTION\nWith the advance of the industrial technology, rotary machines like turbines, compressors, generators, etc. have become larger and larger and the operating speed has become higher and higher. Thus an accurate prediction of the vibration characteristics of the rotor-bearing system must be made in order to design and manufacture highly reliable rotary machines. The vibration characteristics of the rotor-bearing system and response to external shocks have been studied mainly by a linear theory (Lund, (1975)). Adams, (1980) and Hori, (1988), pointed out the importance of adequately designing rotating machinery to withstand a shock, such as earthquake. From this point of view, Kato and Hori, (1990) investigated theoretically the stability of Jefcott rotor and two-disklfour-bearing rotor system on which an artificial sinusoidal emulating a seismic disturbance is applied considering nonlinear characteristics of oil film force. Choy, Braun and Hu, (1992) examined the nonlinear effect on the dynamic performance of hydrodynamic journal bearings for several kinds of operating conditions such as small or large eccentricities. They reported that the nonlinear analysis is more realistic than the linear one, and that\n= vibration amplitude of external disturbance = bearing clearance = damping coefficients of bearing = diameter of bearing = excitation frequency of external disturbance = switch function or cavitation index = film thickness = stiffness coefficients of bearing = length of bearing = mass of journal = dimensionless oil supply pressure, p.JDIW\n= oil pressure = amplitude ratio response functions = time = surface velocity of journal in x direction = static load of the journal bearing = coordinates = fractional film content = phase delay response functions = oil viscosity = angular coordinate = fractional film content = oil density = excitation frequency of external disturbance, 2njd\nD ow\nnl oa\nde d\nby [\nFU B\ner lin\n] at\n0 3:\n19 2\n3 O\nct ob\ner 2\n01 4",
+            "B. RHO AND K. KIM\nthc journnl motion exhibits a more severe non-linearity when the jourlial ccccntricity is relalivcly low.\nI-lowcvcr, when an cxterlial disturbance is applied to the rotorbcnring system, the cavitation regions, which is expected to affect tlic oil fil~ii force and the dynamic characteristics of the bearing, vary from position to position of the journal center, and an accuIXIIC prctliction of cavitation region becomes necessary to get the accurate oil film force. The objective of this paper is to examine tlic nonline;~r vibration characteristics of hydrodynamic journal bearings irsing cavitation algorithm (Elrod, (1981)) at various opcrn~ing conditions. For the purpose of this study, continuous sin~~soitlal excitation in the hori~ontal or vertical direction is considcrcd as nn external disturbnlice.\nGOVERNING EQUATIONS\nThc coordiliatc system and the configuration of the circumfercn~i:~lly grooved journal bearing are shown in Fig. I. It is assumed thal tlic journal and the bearing are circular, the load is applied in clic direction of x, and the circumferential groove is filled with the lubricnnt of constanl pressure.\nFor n finite bearing, the two dimensional, unsteady, Reynolds cquation for a Newtonian lubricant in laminar flow, with a fluid compressibility effects, can be written as (Vijayraghavan and Kcirh, (1989))\nwlicrc tlic cnvitation index g , is zero within the cavitation region and unity clsewhere, and the fluid fractional film content Of is tlcfinctl as\nwlicrc p, is the density of the lubricant at the cavitation pressure. In addition, the pressure in the full film region can be cxprcsscd it1 terms of the fractional film content through the fluid bulk modulus and the switch function as [Elrod, (1981)l\nwhere p, is the cavitation pressure of the oil, taken as zero gauge pressure.\nThe following boundary conditions for the Reynolds Eq. [I] are adopted according to the geometric configuration, the feeding condition, and the periodic condition.\np = p, ut oil supply groove 141\np = 0 ut usial ends 151\nThe Elrod algorithm is used to solve Eq. [I], and the finite difference method together with the Gaussian elimination method is also used for the analysis. The number of grid points in the circumferential and axial directions for a half of each fluid film are 127 and 17, respectively.\nLINEAR METHOD\nIn the case of infinitesimal vibration of the rotor around the static equilibrium position, the oil film force in the bearing can be approximated by a linear theory, and the equation of motion for a rotor-bearing system can be written as\nwhere zij and kij are linear damping and stiffness coefficients of the bearing, and (Aj, Ayj) and (Axh, Ayh) are displacement of the journal and bearing center with respect to the static equilibrium position.\nIn order to obtain the frequency response functions, applying the Laplace transform, the equation of motion can be expressed as follows;\nwhere\nNONLINEAR METHOD\nUsing the angular coordinate, 0, which is shown in Fig. I, the film thickness can be expressed as\nD ow\nnl oa\nde d\nby [\nFU B\ner lin\n] at\n0 3:\n19 2\n3 O\nct ob\ner 2\n01 4",
+            "A Study on Nonlinear Frequency Response Analysis of Hydrodynamic Journal Bearings with External Disturbances\nwhere (.v,, y,) and (.v,, y,) are the coordinate of the journal and bearing center, respectively.\nAs an external disturbance, continuous sinusoidal vibrations in the horizontal or vertical direction are assumed independently. That is, the external disturbance is given by\nwhereAd is the amplitude and wl, is the excitation frequency of the external disturbance.\nFor the sake of simplicity, a symmetrical rigid rotor of mass m supported by oil film bearing is considered. The equation of motion are written as\nwhere W is applied load in x-direction, and Af is the area of full film region.\nThe transient motion of the journal can be evaluated by numerical integration of its acceleration at small time intervals. When an external disturbance is applied to the bearing, the nonlinear frequency response functions are calculated from the steady-state limit cycle of the journal center as\nwhere T is the period of the steady-state response, and\nstate responses. In order to obtain the nonlinear vibration characteristics of the rotor-bearing system, the excitation frequency of the external disturbance is varied from 20 to 300 Hz with 5 Hz step. The vibration amplitude of external disturbance is varied from 0.001C to 0. IC.\nRESLILTS AND DISCUSSION\nThe nonlinear vibration characteristics of hydrodynamic journal bearing with a circumferential groove are obtained numerically using the cavitation algorithm. The specifications of the bearing and the value of each parameter are listed in Table 1.\nFigure 2 shows the normalized steady state loci of the journal center, which is obtained by both linear and nonlinear analysis, when the external sinusoidal disturbance is applied to the bearing in x direction for the rotational speed of 3960 rpm. This operating condition corresponds to the eccentricity of 0.5. It can be seen that the center of the loci moves closer to the bearing center as the amplitude or frequency of the external disturbance increases. I t is also found, from the nonlinear steady state orbit distorted with elliptical shape, that the nonlinear orbit contains higher order harmonics which of course are not found in the linear analysis. 'The nonlinearity of the rotor-bearing system increases significantly with the excitation amplitude or excitation frequency of the external disturbance.\nD ow\nnl oa\nde d\nby [\nFU B\ner lin\n] at\n0 3:\n19 2\n3 O\nct ob\ner 2\n01 4"
+        ]
+    },
+    {
+        "image_filename": "designv11_60_0002132_jsvi.1996.0655-Figure4-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0002132_jsvi.1996.0655-Figure4-1.png",
+        "caption": "Figure 4. A general rotor\u2013bearing system.",
+        "texts": [
+            " Thus, equation (11) must be changed to {ST}r =[Td T]{ST}l , {S2T}r =[Td 2T]{S2T}l (22a, b) and equations (18a) and (18b) are changed to {ST}r =[Ts T]{ST}l , {S2T}r =[Ts 2T]{S2T}l , (23a, b) where Ti 0 0 0 \u00b7 \u00b7 \u00b7 Ti 1 0 0 \u00b7 \u00b7 \u00b7 0 Ti 2 0 \u00b7 \u00b7 \u00b7 0 Ti 3 0 \u00b7 \u00b7 \u00b7 [Ti T]=G G G K k 0 0 . . . \u00b7 \u00b7 \u00b7 G G G L l and [Ti 2T]=G G G K k 0 0 . . . \u00b7 \u00b7 \u00b7 G G G L l , i= d, s. \u00b7 \u00b7 \u00b7 \u00b7 \u00b7 \u00b7 \u00b7 \u00b7 \u00b7 \u00b7 \u00b7 \u00b7 \u00b7 \u00b7 \u00b7 \u00b7 \u00b7 \u00b7 \u00b7 \u00b7 \u00b7 \u00b7 \u00b7 \u00b7 Let R and L denote the right and the left sides of the overall system, respectively. The overall transfer matrix of a system, as shown in Figure 4, is obtained by {ST}R = t 1 i=N [TT ]i{ST}L =[To T]{ST}L (24a) for the T-type motion, and {S2T}R = t 1 i=N [T2T ]i{S2T}L =[To 2T]{S2T}L (24b) for the 2T-type motion, where t 1 i=N [TT ]i and t 1 i=N [T2T ]i represent the product of transfer matrices from the first station (denoted by L) to the nth station (denoted by R). Thus, the two Hill\u2019s infinite determinants of the T- and 2T-type motions are obtained independently. The roots of the truncated determinants of [To T] and [To 2T] give approximate solutions of T- and 2T-type transition curves respectively"
+        ],
+        "surrounding_texts": []
+    },
+    {
+        "image_filename": "designv11_60_0001594_bf00411970-Figure5-1.png",
+        "original_path": "designv11-60/openalex_figure/designv11_60_0001594_bf00411970-Figure5-1.png",
+        "caption": "Fig. 5. An e l e m e n t of the f i c t i t i ous spiral sprit.g, the coi ls of w h i c h are c ircular and ident ica l , in the de f l ec ted s tate .",
+        "texts": [
+            " If, for the sake of simplicity, we imagine the (clamped) wire ends to b e adjustable along their respective radii in such a way that the transition from radius R 0 to R can take place without hindrance, then here, too, there will be no forces present in the plane of the spiral. Besides the winding moment M~ and the (small) reaction moments M x and My caused b y the possible tilting of the coils, we thus need only take into account the (likewise small) axial reaction force Z. The point of the central line of the spring wire, lying initially in ELASTIC STABILITY OF FLAT SPIRAL SPRINGS 19 the x~-plane, is given by the cylindrical coordinates R, O and ~. At this point we introduce (see fig. 5) an auxiliary system of axes which is rigidly connected to the normal cross-section of the wire and which has the axes 1, 2 and 3, where axis 1 is directed along the first principal axis of inertia of the wire section, axis 2 is directed along the second principal axis of inertia of the wire section, axis 3 is directed along the normal to the wire section. We confine ourselves to those wire sections of which the first principal axis of inertia in the unloaded state lies in the plane of the spiral, so that in the deflected state the axes 1 and 3 intersect the xy-plane at very small angles. Their direction cosines with the z-axis will be called vl and v3 respectively. To a first approximation the load which at the outer end of the spring wire is given by the components M , My, M. and Z causes the following moments with respect to the axes 1, 2 and 3 of the wire\" section (see fig. 5) M1 = - - Mx cos O - - 'My sin O 4- M~ v 1 , ] M 2 = M,, } (17) M a = - M s s i n O 4 - M y c o s O + M , v a4-ZR. ] Let us call A the rigidity of the wire with respect to bending about the first principal axis of inertia (axis 1), B the rigidity of the wire with respect to bending about the second principal axis of inertia (axis 2), C the rigidity of the wire wKh respect to twisting (about axis 3) ; then, as will be shown presently, the following relations exist between the deflection components and the loading moments dr3 M1R ] dO -- A + vl' 1 1 M 2 i R R o B ' dv 1 MaR d O - C ~ va\" ] (18) d\u00a2/RdO, we can"
+        ],
+        "surrounding_texts": []
+    }
+]
\ No newline at end of file