diff --git "a/designv11-35.json" "b/designv11-35.json" new file mode 100644--- /dev/null +++ "b/designv11-35.json" @@ -0,0 +1,6071 @@ +[ + { + "image_filename": "designv11_35_0001672_s0368393100116256-Figure12-1.png", + "original_path": "designv11-35/openalex_figure/designv11_35_0001672_s0368393100116256-Figure12-1.png", + "caption": "Fig. 12. Condition for a stable sustained oscillation.", + "texts": [ + " Then through each block the signal will be multiplied by the transfer function of that block so that, at the cut point X emerges Ftp where F is the transfer function of the whole system and is defined by F=fJJ3 . . . /\u201e where fit f2, etc., are the transfer functions of the individual blocks. Now quite obviously if F = l, the output Ftp may be substituted for the input f and the system will continue to oscillate, without external driving if the loop is closed. This condition is vectorially represented in Fig. 12. Now starting with a system which will just sustain a steady oscillation, it is fairly obvious that, if the total lag is decreased, the motion will decay and if it is increased, the motion will diverge until different relationships prevail. If we now ascertain the individual transfer functions for each stage, including the aero plane itself, we may draw the transfer function for the complete closed system. In general it will resemble the curve of Fig. 13. As sketched in Fig. 13, the system is. stable, because it required an increase in the total lag to satisfy the condition of a stable sustained oscillation", + " The symbol /* is used to designate the relative damping referred to critical damping, /* = !. The symbol r represents the ratio of operating to resonant frequency and constant /\u2022 contours are shown dotted. It will be observed that the transfer function of the aeroplane always introduces a lag in excess of 90\u00b0, that the lag increases 424 damped oscillatory system with simple harmonic forcing. but little from /*=0.3 to /* = 0.2, but that it increases rapidly towards 180\u00b0 as the damp ing approaches zero. Now reverting to Fig. 12, it is immediately obvious that the sum of all the lags must not -be allowed to reach 180\u00b0. Thus the advent ;of aeroplanes approximating to n=Q will require the strictest control of all the other lags. When, in addition, we are faced with a shortening of the yaw resonance period, the magnitude of the problem becomes apparent. If, for example, we specify that the accumu lated phase lag, apart from the aeroplane, is not to exceed 12\u00b0, this will be l/30th of the controlled period, i.e. 100 milliseconds for a three second period and 33 milliseconds for a one second period" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_35_0001745_jada.archive.1940.0343-Figure19-1.png", + "original_path": "designv11-35/openalex_figure/designv11_35_0001745_jada.archive.1940.0343-Figure19-1.png", + "caption": "Fig. 19.\u2014 Lines of force of horizontal cast ing machine.", + "texts": [ + " This crucible is constructed of a highly aluminous clay, and the interior is covered with a thick layer of a highly refractory lining, which will withstand approximately 6ooo\u00b0. After repeated use, this lining undergoes a slight checking, but with reasonable care it will withstand about thirty meltings before any serious cracking occurs. No flux of any kind should ever be used in casting platinum, since it is impossible to oxidize the metal under our working conditions, and the flux itself would be injurious to metal, crucible and investment. Figure 19 illustrates diagrammatically the lines of force involved in making a casting with the horizontal type of cast ing machine. The direction o f rotation is counter clockwise. Regardless of the shape of the arm, the relation of the metal to the center of rotation de termines the forces that will be exerted upon it when the casting machine starts. Since the force of inertia is equal to and opposite that o f rotation, the metal is first thrown against the side of the crucible. Eventually, centrifugal force will operate to neutralize inertia and carry the metal in the direction desired" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_35_0000172_012019-Figure6-1.png", + "original_path": "designv11-35/openalex_figure/designv11_35_0000172_012019-Figure6-1.png", + "caption": "Figure 6. Schematical illustration of a cardan joint shaft in a GHH SLP-3H wheel loader.", + "texts": [ + " Series: Materials Science and Engineering 1097 (2021) 012019 IOP Publishing doi:10.1088/1757-899X/1097/1/012019 During the field tests, various signals were recorded via external sensors and CAN bus. The external sensor data such as vibrations, pressure and shaft velocity of the main shaft were recorded at a high sampling rate. The operational data from the CAN bus was recorded at a low sampling rate due to the limitations of the CAN bus. Figure 5 shows a picture of the vehicle at K+S underground facilities together with sensor positions used for measurements. Figure 6 shows schematically the location of cardan shaft in the wheel loader GHH SLP-3H. The cardan shaft in the middle of the vehicle can transfer the drive torque when vehicle\u2019s front and rear sides have a relative angle with respect to each other (i.e. when steering). Figure 5. A picture of the test wheel loader type GHH SLP-3H together with additional sensors for vibration monitoring of the cardan shaft. In total, three sessions, healthy condition as reference, misaligned cardan shaft, and cardan shaft with bearing axial clearance, all using the same cardan shafts used for test rig investigations in Section 2, were measured" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_35_0000004_j.ijpvp.2021.104314-Figure8-1.png", + "original_path": "designv11-35/openalex_figure/designv11_35_0000004_j.ijpvp.2021.104314-Figure8-1.png", + "caption": "Fig. 8. The FE model of a bi-directional pig running in the straight pipeline.", + "texts": [ + " The friction coefficient in actual pigging operation cannot be obtained directly by the experimental measurement, therefore the method introduced in the reported researches, which calculates the friction coefficient by combining the experimental measurement and the 3D finite element method (FEM) was adopted to determine the friction coefficient in this paper. According to the experimental condition described in section 2, the corresponding 3D FE model was established using the ANSYS 17.0 (ANSYS Software, United States), as shown in Fig. 8(a) and (b). To ensuring the accuracy of the simulation and solving the convergence problem, the pipeline and the sealing discs were meshed into the same part in the circumferential. The detailed setting of the contact pair and the material properties had been introduced in the published paper. Setting the friction coefficient as zero in the FE model, the normal contact force on the sealing disc can be extracted from the ANSYS postprocessing. Following Coulomb\u2019s friction law, the ratio between the simulated normal contact force and the experimental friction force is defined as the friction coefficient" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_35_0000459_10402004.2020.1831674-Figure1-1.png", + "original_path": "designv11-35/openalex_figure/designv11_35_0000459_10402004.2020.1831674-Figure1-1.png", + "caption": "Figure 1. (a) Schematic diagram of the water-lubricated thrust\u2013journal coupled bearing and (b) coordinate system of the coupled bearing.", + "texts": [ + "RTICLE HISTORY Received 7 October 2019 Accepted 29 September 2020 KEYWORDS Coupled bearing; axial microvibration; transient hydrodynamic lubrication; water lubrication; textured surface Journal\u2013thrust coupled bearings are typically used to support the combination of radial and thrust forces occurring in many engineering applications. Recently, a novel shipborne shaftless propulsor (1, 2) was developed to simplify the traditional stern shaft system. In such a marine propulsor, the journal\u2013thrust coupled bearing is integrated with the propeller and, therefore, the coupled bearing rotates with the propeller synchronously, as illustrated in Fig. 1a. It is evident that in such a coupled bearing the hydrodynamic pressure and the flow field of the journal and thrust bearings interact at the common boundary, thereby generating the coupled hydrodynamic pressure. In addition, it can be identified that the coupled bearing will be exposed to periodic excitation caused by the propeller (3), which leads to a transient hydrodynamic response in the coupled bearing. For a single journal bearing system, misalignment (8\u201312) and axial micromotion (13\u201319) typically occur, which affect the transient hydrodynamic performance by the time-varying lubrication gap", + " They found that the positive misalignment mode generates a larger load capacity than the negative misalignment mode. Subsequently, Wang et al. (5) further integrated the mass conserving boundary condition into the previously developed numerical model. In 2006, Wang et al. (6) demonstrated that for oil-lubricated coupled bearings, the elastic deformation cannot be ignored during mixed EHL simulations when the eccentricity ratio is larger than 0.9. However, it is noteworthy that the coupled bearing configuration introduced by Wang et al. (4\u20136) is different from that in the present study, as shown in Fig. 1. To be specific, in Wang et al.\u2019s (4\u20136) studies, the hydrodynamic pressure of the plain thrust bearing was entirely dependent on the journal misalignment. Recently, Xiang et al. (7) developed a mixed EHL model for aligned journal and titling pad thrust coupled bearings. In their study, the interaction between the hydrodynamic behavior of the journal bearing and the thrust bearing was explored. In the mentioned studies, a full understanding of the steady lubrication performance of the coupled bearing has been gained by numerical simulation", + " (17) further compared the time-varying hydrodynamic performance of a misaligned microgrooved journal bearing engineered with three different microgroove distributions, in which the journal performs a periodic axial movement. In 2019, Li et al. (18) carried out a transient thermohydrodynamic lubrication analysis of misaligned journal bearings considering the journal axial movement. It is noteworthy that the abovementioned studies (15\u201318) primarily focused on a single journal bearing considering the axial movement. More recently, Sun et al. (19) performed a stability analysis for journal\u2013thrust coupled bearings considering the disturbing moments (as shown in Fig. 1a). In their study, the transient hydrodynamic force of the journal bearing was simplified as a stiffness\u2013damping model, and the coupled hydrodynamic pressure at the common boundary was absent. For the coupled bearing, it is evident that the additional hydrodynamic pressure spreading from the thrust bearing significantly affects the lubrication performance of the journal bearing, especially when axial microvibration is introduced. The novelty of the present study is to reveal the transient hydrodynamic coupling mechanism of the misaligned journal\u2013thrust coupled bearing subjected to axial microvibration. Nomenclature C: Radial clearance (mm) Dg: Maximum groove depth (lm) e: Eccentricity (see Fig. 1) F: Load capacity (N) GJg: Groove depth of the journal bearing GTg: Groove depth of the thrust bearing h: Film thickness (lm) hdA Axial displacement (lm) hdL Horizontal displacement (lm) hJ0: Geometric clearance of journal bearing (lm) hp: Geometric clearance of thrust bearing (lm) kg: The kgth groove L: Bearing length (mm) DL: 1/2 of the microvibration amplitude (m) m: Circumferential node number of the journal bearing Nr: Rotational speed (rpm) p: Hydrodynamic pressure (MPa) R1: Journal bearing radius (mm) R2: Thrust bearing outer radius (mm) rg: Groove width ratio, Wg/(Wg\u00feWb) Th: Circumferential period (s) Tz: Axial period (s) t: Time (s) U: Circumferential velocity (m/s) V: Axial velocity (m/s) Wb: Bank width (mm) Wg: Groove width (mm) b: Misalignment angle (degree) c: Relation factor D: Grid size (see Fig", + " Furthermore, the parametric studies were conducted as a guide to help identify appropriate operating parameters to both improve the load capacity and decrease the load fluctuation. Transient hydrodynamic pressure for the coupled bearing In the hydrodynamic lubrication regime, the hydrodynamic pressure of the coupled bearing can be calculated by the well-accepted Reynolds equation. It should be mentioned that for the journal bearing, the axial Couette flow should be evaluated in the Reynolds equation due to the axial microvibration. In addition, the transient effects should also be evaluated in the Reynolds equation for both journal and thrust bearings. As shown in Fig. 1, it is evident that the Reynolds equations in the Cartesian and polar coordinate systems are required to determine the hydrodynamic pressure of the journal and thrust bearings, respectively. As illustrated in Eq. [1], the transient hydrodynamic pressure of the journal bearing can be calculated by @ R2 1@h qh3J 12g @pJ @h ! \u00fe @ @z qh3J 12g @pJ @z ! \u00bc U 2 @qhJ R1@h \u00fe V 2 @qhJ @z \u00fe @qhJ @t , [1] where pJ and hJ are the hydrodynamic pressure and the lubrication gap of the journal bearing, respectively; q and g are the density and viscosity of water, respectively; h and z are the circumferential and axial directions, respectively; U and V is the velocity in the circumferential and axial directions, respectively; R1 is the radius of the journal bearing; and t is the operating time", + " For the thrust bearing, the transient hydrodynamic pressure can be governed by @ r@r qh3Tr 12g @pT @r \u00fe 1 r @ @h qh3T 12g @pT @h \u00bc xr 2 @qhT @h \u00fe @qhT @t , [2] where pT and hT are the hydrodynamic pressure and the lubrication gap of the thrust bearing, respectively; r is the radial direction; and x is the rotation speed (rad/s). For the coupled bearing, it is noteworthy that the transient interaction of the hydrodynamic behavior between the journal and thrust bearings is formed by the coupled hydrodynamic pressure at the common boundary (see Fig. 1). More details related to the coupling mechanism will be discussed in the subsection on boundary conditions. Figure 2. Misalignment mode of the coupled bearing. In this study, two misalignment modes\u2014that is, positive misalignment and negative misalignment\u2014are studied, as shown in Fig. 2. For the journal bearing, the lubrication gap is comprised of the geometry gap hJ0, microgroove depth GJg\u00f0h, z\u00de, and misalignment term hJm, which can be expressed as hJ h, z\u00f0 \u00de \u00bc C 1\u00fe e cos h u\u00f0 \u00de\u00f0 \u00de \u00fe tan 6b\u00f0 \u00dez cos h u\u00f0 \u00de \u00fe GJg h, z\u00f0 \u00de, [3] where b is the misalignment angle of the coupled bearing. In this study, a simplification has been made that the misalignment angle remains constant during microvibration; similar treatment can also be found in the literature (14\u201317). In Eq. [3], the plus sign represents positive misalignment and the minus sign represents negative misalignment. u is the attitude angle (as shown in Fig. 1). In the present study, a microgroove with a rectangular bottom shape was considered. Based on the geometric relation shown in Fig. 3, the microgroove depth can be calculated by GJg h, z\u00f0 \u00de \u00bc Dg h kg L h h kg R 0 h > h kg R or h < h kg L , ( [4] where Dg is the maximum groove depth, and h kg L and h kg R are the left and right angular positions of the kgth microgroove, which can be calculated by h kg L \u00bc p Ng 2kg 1 rg h kg R \u00bc p Ng 2kg \u00fe 1 rg , 8>< >: [5] where Ng is the microgroove number, and rg is the microgroove ratio, which can be calculated by rg \u00bc Wg=\u00f0Wg \u00fe Wb\u00de, where Wg and Wb are the width of the groove and bridge, respectively (as shown in Fig. 1). The transient lubrication gap of the thrust bearing can be calculated as the sum of the geometric clearance hp (as shown in Fig. 1), groove depth GTg\u00f0h, r\u00de, misalignment term, and time-varying horizontal displacement hdL\u00f0t\u00de, which can be expressed as hT h, r, t\u00f0 \u00de \u00bc hp \u00fe GTg h, r\u00f0 \u00de \u00fe r sin 6b\u00f0 \u00de cos \u00f0h\u00de \u00fe hdL\u00f0t\u00de, [6] where r is the radial coordinate of the microgrooved thrust bearing, as shown in Fig. 2. For the thrust bearing, the plus sign indicates negative misalignment and the minus sign indicates positive misalignment. In this study, the microgroove depth of the journal bearing is equal to that of the thrust bearing. As depicted in Fig", + " 9b, the load capacity of the thrust bearing yielded by negative misalignment is larger than that yielded by positive misalignment. Generally, it can be concluded that due to the axial microvibration, the time-varying hydrodynamic pressure of the thrust bearing results in hydrodynamic fluctuation of the journal bearing, and the negative misalignment mode has potential to promote the hydrodynamic effect for both the journal and thrust bearings compared to the positive misalignment mode. In the marine propulsor shown in Fig. 1a, the propeller will work in various axial microvibration frequencies and Figure 11. Evolution of the load capacity over time of the (a) journal bearing and (b) thrust bearing at different microvibration amplitudes. Variation in the maximum and minimum predicted results, including the load capacity and the maximum fluid pressure, with microvibration amplitude: (c) journal bearing and (d) thrust bearing. amplitudes (19), thereby generating the transient hydrodynamic performance of the coupled bearing" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_35_0000472_s00170-021-07362-2-Figure20-1.png", + "original_path": "designv11-35/openalex_figure/designv11_35_0000472_s00170-021-07362-2-Figure20-1.png", + "caption": "Fig. 20 Distortions for the canonical part before build plate removal", + "texts": [ + " However, these residual stresses turn tensile at the overhang reaching their maximum value at the top. In this section, the case of the IN718 canonical part is investigated. Unlike the cantilever case where distortions take place after cutting supports, the distortion for this canonical part takes place during the process. The final displacement solution could be obtained by comparing the final deformed mesh to the initial mesh. The predicted distortion for this part with and without adaptive remeshing and at two adaptive remeshing configurations is shown in Fig. 20. This predicted distortion can find an agreement with the experimental solution in [7], though the maximum displacement value is underestimated by 25%. As discussed earlier, the main reason for that deviation might be the assumed yield strength given that microstructure and mechanical properties are different at the thin features. To evaluate the accuracy of using adaptive remeshing, firstly, the distortion solutions at different adaptive remeshing configurations were compared, where it was found that solution slightly changed within the considered adaptive mesh configurations" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_35_0001419_978-94-007-7401-8-Figure6.32-1.png", + "original_path": "designv11-35/openalex_figure/designv11_35_0001419_978-94-007-7401-8-Figure6.32-1.png", + "caption": "Fig. 6.32 3PPPaR-1RPPaPa-type maximaly regular fully-parallel PMs with Sch\u00f6nflies motions defined by MF = SF = 4, (RF) = (v1; v2; v3;xb), TF = 0, NF = 21, limb topology P\\P\\||Pa||R, R||P||Pa||Pa and P\\P\\||Pa\\||R (a), P\\P\\||Pa ??R (b)", + "texts": [ + " 1 5. 3PPaPR-1RPaPaP (Fig. 6.29a) P||Pa\\P\\\\R (Fig. 4.8d) P||Pa\\P\\||R (Fig. 4.8e) R||Pa||Pa||P (Fig. 5.4k) Idem no. 1 6. 3PPaPR-1RPaPatP (Fig. 6.29b) P||Pa\\P\\\\R (Fig. 4.8d) P||Pa\\P\\||R (Fig. 4.8e) R||Pa||Pat||P (Fig. 5.4l) Idem no. 1 7. 2PPaRRR-1PPaRR-1RPaPaP (Fig. 6.30) P||Pa||R||R\\R (Fig. 4.9a) P||Pa||R||R (Fig. 4.8g) R||Pa||Pa||P (Fig. 5.4k) Idem no. 1 8. 2PPaRRR-1PPaRR1RPaPatP (Fig. 6.31) P||Pa||R||R\\R (Fig. 4.9a) P||Pa||R||R (Fig. 4.8g) R||Pa||Pat||P (Fig. 5.4l) Idem no. 1 9. 3PPPaR-1RPPaPa (Fig. 6.32a) P\\P\\||Pa\\||R (Fig. 4.8b) P\\P\\||Pa||R, (Fig. 4.8c) R||P||Pa||Pa (Fig. 5.4m) The last revolute joints of limbs G1, G2 and G3 and the revolute joint between links 7D and 9 have superposed axes. The actuated prismatic joints of limbs G1, G2 and G3 have orthogonal directions 10. 3PPPaR-1PRPaPa (Fig. 6.32b) P\\P\\||Pa\\R (Fig. 4.8a) P\\P\\||Pa||R, (Fig. 4.8c) R||P||Pa||Pa (Fig. 5.4m) Idem no. 9 11. 3PPPaR-1RPPaPat (Fig. 6.33a) P\\P\\||Pa\\||R (Fig. 4.8b) P\\P\\||Pa||R, (Fig. 4.8c) R||P||Pa||Pat (Fig. 5.4n) Idem no. 9 (continued) 642 6 Maximally Regular Topologies with Sch\u00f6nflies Motions Table 6.13 (continued) No. PM type Limb topology Connecting conditions 12. 3PPPaR-1RPPaPat (Fig. 6.33b) P\\P\\||Pa\\\\R (Fig. 4.8a) P\\P\\||Pa||R, (Fig. 4.8c) R||P||Pa||Pat (Fig. 5.4n) Idem no. 9 13. 3PPaPR-1RPPaPa (Fig. 6.34) P||Pa\\P\\R (Fig" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_35_0000504_tmag.2021.3087267-Figure1-1.png", + "original_path": "designv11-35/openalex_figure/designv11_35_0000504_tmag.2021.3087267-Figure1-1.png", + "caption": "Fig. 1. Rendering of a generic stator end-region including the compression system", + "texts": [ + " Index Terms\u2014Clamping System, End-Region Losses, Hydroelectric Power Generation, Synchronous Generator A system matrix Ae element area B\u03020 external flux density amplitude B\u03b4 air-gap flux density B\u0302\u03b4 fundamental air-gap flux density harmonic d thickness H0 external magnetic field HX magnetic field from self-inductance I\u0302\u25e6 loop current Iedge edge current I\u0302\u03b4 virtual air-gap current amplitude lfinger axial length of pressure finger Pedge Losses along one edge of the mesh R\u25e6 diagonal entry in the system matrix Redge edge resistance U i,0 vector of induced voltages from external flux U\u0302 i,0 amplitude phasor of the induced voltage from external flux t pressure finger thickness (axial length) at an edge\u2019s center w width of an edge \u03b4 air-gap thickness \u03c4p pole pitch in meters \u03c3 electric conductivity II. INTRODUCTION Large synchronous machines require a complex stator compression system to apply enough axial force onto the stator teeth to prevent the laminations from vibrating and damaging the stator winding insulation. A rendering of such a compression system for a large salient-pole synchronous machine can be seen in Fig. 1. The compression system consists of clamping bolts applying tension onto the clamping plate through large spring washers. The clamping plate applies force onto the pressure fingers, which in turn apply this force onto the stator teeth, reducing unwanted vibrations of the end laminations. Parasitic eddy current losses inside the compression system are caused by the end-region field. Most losses occur inside the clamping plate and pressure fingers, while losses in the clamping bolts and spring washers are Manuscript received April 6th, 2021; revised June 4th, 2021" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_35_0000027_j.procir.2020.05.200-Figure3-1.png", + "original_path": "designv11-35/openalex_figure/designv11_35_0000027_j.procir.2020.05.200-Figure3-1.png", + "caption": "Fig. 3. Demonstration of shape variability by the example of some demonstrators: (a) workpiece bevel gear, (b) shaft, (c) bevel gear, (d) bearing bush.", + "texts": [ + " Development of a form-flexible gripper with active cooling This section introduces the solution concepts for the gripper echanism and the cooling system. A brief look at the system in ig. 2 shows that the system consists of the gripping unit, which n turn consists of the two formflexible gripper jaws and a parllel jaw gripper. Furthermore, the cooling system, which includes he nozzles and their holder, is shown. The design of the gripper echanics and the active cooling system are presented in the fol- owing. .1. Gripper mechanics In order to meet the demands of the Tailored Forming process, novel gripper was designed. As shown in Fig. 3 , the gripping rinciple of the Omnigripper, pins actuated by air pressure and the atrixgripper, opposing gripper jaws, is combined. The jaws have matrix with 7x5 pins each with a stroke of 24mm, which are ell suited for the demonstrators in Tailored Forming ( Fig. 3 ). The imension of one jaw are 86x66x60 mm and can be extended as equired. The independently movable pins allow differently shaped bjects to be handled. Fig. 3 a shows the workpiece of the bevel ear clamped in the gripper. The other pictures of the figure show he gripped demonstrators of the first period of CRC 1153. The hape variability is given by the gripping principle. The pins are xtended by compressed air. The jaws are then closed by the parllel gripper without compressed air being applied to the pins. The ins, which will face resistance, are retracted during further move- ent of the jaws. When the jaws have reached their position de- ned by the user, compressed air is applied again" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_35_0001419_978-94-007-7401-8-Figure5.13-1.png", + "original_path": "designv11-35/openalex_figure/designv11_35_0001419_978-94-007-7401-8-Figure5.13-1.png", + "caption": "Fig. 5.13 2PaRPRR-1PaRPR-1RUPU-type fully-parallel PMs with uncoupled Sch\u00f6nflies motions defined by MF = SF = 4, RF\u00f0 \u00de \u00bc \u00f0v1; v2; v3;xb\u00de, TF = 0, NF = 11, limb topology Pa\\R\\P\\kR\\R, Pa\\R\\P\\kR and R\\R\\R\\P\\kR\\R", + "texts": [ + " 2PaPRRR1PaPRR1RUPU (Fig. 5.12a) Pa\\P\\\\R||R\\\\R (Fig. 5.3a) The first revolute joint of G4 limb and the last revolute joints of limbs G1, G2 and G3 have parallel axes. The last joints of of limbs G1, G2 and G3 have superposed axes. The revolute joints between links 5 and 6 of of limbs G1, G2 and G3 have orthogonal axes Pa\\P\\\\R||R (Fig. 5.2e) R\\R\\R\\P\\kR\\R (Fig. 5.1b) 10. 2PaPRRR1PaPRR1RUPU (Fig. 5.12b) Pa||P\\R||R\\\\R (Fig. 5.3b) Idem No. 9 Pa||P\\R||R (Fig. 5.2f) R\\R\\R\\P\\kR\\R (Fig. 5.1b) 11. 2PaRPRR1PaRPR1RUPU (Fig. 5.13) Pa\\R\\P\\kR\\R (Fig. 5.3c) The first revolute joint of G4 limb and the last revolute joints of limbs G1, G2 and G3 have parallel axes. The last joints of of limbs G1, G2 and G3 have superposed axes, and their revolute joints between links 4 and 5 have orthogonal axes Pa\\R\\P\\kR (Fig. 5.2g) R\\R\\R\\P\\kR\\R (Fig. 5.1b) 12. 2PaRRRR1PaRRR1RUPU (Fig. 5.14) Pa\\R||R||R\\kR (Fig. 5.3d) Idem No. 11 Pa\\R||R||R (Fig. 5.2h) R\\R\\R\\P\\kR\\R (Fig. 5.1b) 5.1 Fully-Parallel Topologies 527 Table 5.2 Limb topology and connecting conditions of the fully-parallel solutions with no idle mobilities presented in Figs" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_35_0000296_s12206-021-0435-1-Figure2-1.png", + "original_path": "designv11-35/openalex_figure/designv11_35_0000296_s12206-021-0435-1-Figure2-1.png", + "caption": "Fig. 2. Schematic overview of the joints 1-4 assembly [36].", + "texts": [ + " 3 presents the identification method for the dynamic parameters. In Sec. 4, calculation of the optimal excitation trajectory is presented. Experimental results are given and discussed in Sec. 5. Sec. 6 gives the conclusion. Geometrical modeling of robots can be achieved by the Denavit-Hartenberg (DH) formulation based on kinematic frames located on the joint axes. The St\u00e4ubli RX160 is equipped with 6 revolute joints. The first four joints of the robot are directly driven by servo motors via helical gear transmissions as shown in Fig. 2. For the remaining two joints, the servos are mounted inside the fourth link of the robot and a gear transmission transfers the motion from servo 6 through the 5th joint to the 6th joint [37]. This mechanical design causes a kinematic coupling between the servo 5 and joint 6 as shown in Fig. 3. The angular velocity of servo 6 is related to the difference of joint velocities 5q and 6q . Consequently, one needs to define additional coordinate axes in order to take into account this extra friction model" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_35_0001582_s11071-015-2197-8-Figure2-1.png", + "original_path": "designv11-35/openalex_figure/designv11_35_0001582_s11071-015-2197-8-Figure2-1.png", + "caption": "Fig. 2 Frequency-independent DF curve", + "texts": [ + " Hence fundamental frequency response in the DF approach can be considered as y(t) \u2248 A[a1 sin(\u03c9t) + b1 cos(\u03c9t)]G( j\u03c9) (1) where A is the amplitude of the sinusoidal excitation; a1 and b1 are the first-order Fourier coefficients. The output of f [e(t)] can be approximated as u(t) \u2248 N (A, \u03c9)A sin(\u03c9t) (2) where N (A, \u03c9) = j \u03c0 A \u222b \u03c0 \u2212\u03c0 f [A sin(\u03c9t)] \u00b7 e\u2212 j\u03c9t d(\u03c9t) If the nonlinear element is independent of the frequency of input signals,we get the equation N (A, \u03c9) = N (A). The Nyquist plot of the linear plant G( j\u03c9) and the negative inverse describing function\u22121/N (A) are drawn in the complex plane as shown in Fig. 2. Using the small amplitude perturbation method, the traditional DF approach can find an unstable oscillation corresponding to the point a and a stable oscillation corresponding to the point b, and then estimate the oscillation frequency and amplitude of the stable oscillation. With increasing complexity of nonlinear control algorithms, nonlinear elements may be closely related to the frequency of the input signal e(t), called frequency-dependent nonlinearities. Hence, it is hard for the traditional DF approach to plot a cluster of \u22121/N (A, \u03c9)|\u03c9=\u03c91,\u03c92," + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_35_0000246_tte.2021.3068819-Figure13-1.png", + "original_path": "designv11-35/openalex_figure/designv11_35_0000246_tte.2021.3068819-Figure13-1.png", + "caption": "Fig. 13. The topology of the single-tooth model with the concentrated windings, (a) 3D structure with \u03b8=0\u00b0, (b) 3D structure with \u03b8=360\u00b0.", + "texts": [ + " Even though ignoring the modeling time of 3D model and 2D model, and assuming that the computation speed of the computer is constant, 2D models with fewer mesh nodes will take less time to be computed. Authorized licensed use limited to: Univ of Calif Santa Barbara. Downloaded on June 24,2021 at 07:00:45 UTC from IEEE Xplore. Restrictions apply. 2332-7782 (c) 2021 IEEE. Personal use is permitted, but republication/redistribution requires IEEE permission. See http://www.ieee.org/publications_standards/publications/rights/index.html for more information. In this section, the AC loss of a single-tooth model with concentrated windings will be analyzed. Fig. 13 shows the cross section, mesh model and 3D structures with different transposition angles. The single-tooth model is similar in size and structure to the actual motor stator. The stator core is made of amorphous alloy, and the geometry parameters of the stator are shown in Table III. Due to the proximity effect and the skin effect, the difference in the current distribution between the 5 parallel strands is very obvious. As transposition angle and frequency change, the magnitude and phase of the current will be changed distinctly", + " For \u03b8=360\u00b0 or the other angles, n=2 or 3 will be selected. The number of mesh nodes is 2-3 times more than n=1, and the computational efficiency will drop. But compared with the 3DModel, the computational efficiency of 2D-EMM is still much higher. Regardless of the slot opening effect and rotor rotation, single-tooth models were manufactured and their winding shave two different \u03b8. Power loss of two models were tested, shown in Fig. 16. The structure of single-tooth models with concentrated winding are same as shown in Fig. 13. The winding is powered by Programmable AC source 61511@Chroma which can provide 12 kVA from 15Hz to 1500Hz. According to the current test results, loss ratios of two winding samples can be calculated by (18), shown in Fig. 17. It can be seen that, for two winding samples, 2D-EMM has high accuracy. From 50 Hz to 1500Hz, the loss ratio of \u03b8=0\u00b0 will increase sharply. However, the increase in loss ratio of \u03b8=360\u00b0 is not obvious. Due to the neglect of the winding ending effect, the accuracy of 2D-EMM will reduce" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_35_0001419_978-94-007-7401-8-Figure3.83-1.png", + "original_path": "designv11-35/openalex_figure/designv11_35_0001419_978-94-007-7401-8-Figure3.83-1.png", + "caption": "Fig. 3.83 Overactuated PMs with coupled Sch\u00f6nflies motions of types 4CPaP (a) and 4PaCP (b) defined by MF \u00bc SF \u00bc 4; RF\u00f0 \u00de \u00bc v1; v2; v3;xd\u00f0 \u00de, TF = 0, NF = 18, limb topology C||Pa\\P (a) and Pa||C\\P (b)", + "texts": [ + "50m0) Idem No. 37 41. 4PaRPR (Fig. 3.78) Pa\\R\\P\\kR (Fig. 3.50n0) Idem No. 37 42. 4PaPRR (Fig. 3.79a) Pa||P\\R||R (Fig. 3.50o0) Idem No. 37 43. 4PaPRR (Fig. 3.79b) Pa\\P\\\\R||R (Fig. 3.50p0) Idem No. 37 44. 4PaRRP (Fig. 3.80a) Pa\\R||R\\P (Fig. 3.50q0) Idem No. 23 45. 4RPaRR (Fig. 3.80b) R\\Pa\\kR||R (Fig. 3.50s0) Idem No. 37 46. 4PaRRR (Fig. 3.81) Pa\\R||R||R (Fig. 3.50r0) Idem No. 37 47. 4PaPRP (Fig. 3.82a) Pa||P\\R\\P (Fig. 3.50t0) Idem No. 23 48. 4PaPRP (Fig. 3.82b) Pa\\P\\\\R\\P (Fig. 3.50u0) Idem No. 23 49. 4CPaP (Fig. 3.83a) C||Pa\\P (Fig. 3.50v0) The cylindrical joints of the four limbs have parallel directions 50. 4PaCP (Fig. 3.83b) Pa||C\\P (Fig. 3.50w0) Idem No. 49 51. 4CRPa (Fig. 3.84a) C||R||Pa (Fig. 3.50x0) Idem No. 49 52. 4PCPa (Fig. 3.84b) P\\C||Pa (Fig. 3.50y0) Idem No. 49 53. 4PPaC (Fig. 3.85a) P\\Pa||C (Fig. 3.50z0) Idem No. 49 54. 4RCPa (Fig. 3.85b) R||C||Pa (Fig. 3.50z01) Idem No. 49 55. 4CPPa (Fig. 3.86a) C\\P\\kPa (Fig. 3.50a0 0) Idem No. 49 56. 4PaPC (Fig. 3.86b) Pa\\P\\kC (Fig. 3.50b0 0) Idem No. 49 57. 4PaRC (Fig. 3.87a) Pa||R||C (Fig. 3.50c0 0) Idem No. 49 58. 4PaCR (Fig. 3.87b) Pa\\C\\P (Fig. 3.50e0 0) Idem No. 49 59" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_35_0000416_00218464.2021.1926996-Figure2-1.png", + "original_path": "designv11-35/openalex_figure/designv11_35_0000416_00218464.2021.1926996-Figure2-1.png", + "caption": "Figure 2. TAST-like specimen: (a) technical drawing, (b) 3D model, (c) bonded specimen.", + "texts": [ + " In partnership with Tellure R\u00f4ta,[21] a company specialized in manufacturing industrial wheels, Ragni and others[22] developed and experimentally assessed two innovative specimens for the shear strength characterization of bonded joints involving adherends with a remarkably different stiffness and with adhesive in thin film. In particular, the tensile specimen, resembling the Thick Adherend Shear Test (TAST) specimen,[23] is a metal-elastomer-metal sandwich bonded joint, which undergoes shear load; as described in, [22] this configuration, named TAST-like (Figure 2a), provides a nearly uniform shear stress along the bonding line and a negligible peel stress, as occurs for the TAST specimen. Thus, it provides a simple and reliable evaluation of the shear strength between adherends with a remarkably different stiffness. This work uses the TAST-like specimen to investigate the effect of the manufacturing parameters and of the working temperature on the shear strength response of a peculiar metal-polyurethane bonded joint, used in industrial wheels. To this aim, we performed three experimental sub-plans; the first focuses on the effect of thermal conditioning of the adhesive applied to the metallic adherend; the second sub-plan concentrates on the type of shot peening performed on the surface of the metallic adherend; the third sub-plan concerns the type of solvent-based adhesive used in the joint", + " As in the previous work, [22] the manufacturing of the TAST-like specimens involved the following four steps: first, the degreasing of the metallic adherends and shot peening of the surface; second, the application of the solventbased adhesive; third, the conditioning of the adherends with the adhesive at a specific combination of time and temperature into a preheated mold; fourth, the casting of polyurethane after the adherends reached the mold temperature. We highlight that the adhesive thickness is negligible, since the adhesive is liquid, and the solvent evaporates during the conditioning phase. Figure 2a shows the sketch of the main part of the TAST-like specimen, consisting in a sandwich obtained by casting and bonding an inner elastic polyurethane (A) to a couple of adherends made of mild steel (B). The dimensions of the specimens represent the best tradeoff between manufacturing and testing constraints. In particular, assuming a maximum failure shear stress at the metal-polyurethane interface equal to 20 MPa, and considering that the maximum axial force that the available testing machine can apply is 4 kN, it comes a bonded area equal to 20 mm x 10 mm. In Figure 2b, the plates C allow to fasten the TAST-like specimen to the testing machine, while the plates D join the inner sandwich to the plates C. A couple of hinges (E in Figure 3) between the plates C (Figure 2b) and the crossheads of testing machine ensured a perfect alignment of the specimen (Figure 2c) in order to promote a pure shear load on the specimen, by minimizing the bending moment on the joint. We observe that the lower the thickness of the polyurethane the more uniform the shear stress on the joint and the peel stress tends to zero. This was assessed in the design phase[22] both with a FE model of the joint and using the classical theoretical model from Bigwood and Crocombe.[24] The polyurethane thickness in this joint (10 mm) comes from manufacturing constraints due to casting: a lower thickness cannot be easily obtained and would involve a cooling process that significantly differs from the standard industrial process" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_35_0000433_tie.2021.3078396-Figure4-1.png", + "original_path": "designv11-35/openalex_figure/designv11_35_0000433_tie.2021.3078396-Figure4-1.png", + "caption": "Fig. 4: A simple model for jumping motion generation.", + "texts": [ + " Consequently, most walking motion generators consider a reference ZMP. For example, Kajita et al. generated a walking motion with a preview control method [26], and Herdt et al. used the MPC [27]; These methods employed an LIPM, which has a constant height of COM. We generate the horizontal motion of the COM for jumping motion in a similar way to the walking except for a few discrepancies. The LIPM is not proper for jumping motion generation because the height of the COM of the robot changes significantly. So, we introduce a VHIPM as shown in Fig. 4. Further, since the angular momentum of the robot causes the rotating motion of the robot during the flight phase, we consider the angular momentum of the robot about COM. The ZMP equation of the VHIPM is p = x\u2212 hx\u0308 h\u0308+ g \u2212 I\u03b8\u0308 m(h\u0308+ g) , (3) where p is the position of ZMP, x and h are positions of COM along the x- and y-axis, \u03b8 is the rotation angle of the CoM, and g is gravity acceleration, I is the inertia of the robot about COM, and m is the mass of the robot. If the angular momentum about the COM is zero, \u03b8\u0308 = 0 and p is given by: p = x\u2212 h h\u0308+ g x\u0308" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_35_0001419_978-94-007-7401-8-Figure3.6-1.png", + "original_path": "designv11-35/openalex_figure/designv11_35_0001419_978-94-007-7401-8-Figure3.6-1.png", + "caption": "Fig. 3.6 Overactuated PMs with coupled Sch\u00f6nflies motions of types 4RRPR (a) and 4PPRR (b) defined by MF \u00bc SF \u00bc 4; RF\u00f0 \u00de \u00bc v1; v2; v3;xd\u00f0 \u00de, TF = 0, NF = 6, limb topology R||R||P||R (a) and P\\P\\kR||R (b)", + "texts": [ + "1 Limb topology and connecting conditions of the overactuated solutions with no idle mobilities presented in Figs. 3.4, 3.5, 3.6, 3.7, 3.8, 3.9, 3.10, 3.11, 3.12, 3.13, 3.14, 3.15, 3.16, 3.17, 3.18, 3.19, 3.20 No. PM type Limb topology Connecting conditions 1. 4PRRR (Fig. 3.4a) P||R||R||R (Fig. 3.1a) The prismatic joints of the four limbs have parallel directions 2. 4RPRR (Fig. 3.4b) R||P||R||R (Fig. 3.1b) The first revolute joints of the four limbs have parallel axes 3. 4RPRR (Fig. 3.5a) R||P||R||R (Fig. 3.1c) Idem No. 2 4. 4RRPR (Fig. 3.5b) R||R||P||R (Fig. 3.1d) Idem No. 2 5. 4RRPR (Fig. 3.6a) R||R||P||R (Fig. 3.1e) Idem No. 2 6. 4PPRR (Fig. 3.6b) P\\P\\kR||R (Fig. 3.1i) Idem No. 1 7. 4RRRP (Fig. 3.7a) R||R||R||P (Fig. 3.1f) Idem No. 2 8. 4RPRP (Fig. 3.7b) R\\P\\kR||P (Fig. 3.1g) Idem No. 2 9. 4PPRR (Fig. 3.8a) P\\P||R||R (Fig. 3.1h) The last revolute joints of the four limbs have parallel axes 10. 4PRPR (Fig. 3.8b) P\\R||P||R (Fig. 3.1p) Idem No. 9 11. 4RRPP (Fig. 3.9a) R||R||P\\P (Fig. 3.1j) Idem No. 2 12. 4RPRP (Fig. 3.9b) R||P||R\\P (Fig. 3.1k) Idem No. 2 13. 4RPPR (Fig. 3.10a) R||P\\P\\kR (Fig. 3.1l) Idem No. 2 14. 4RPPR (Fig. 3.10b) R\\P\\kP||R (Fig" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_35_0001419_978-94-007-7401-8-Figure4.22-1.png", + "original_path": "designv11-35/openalex_figure/designv11_35_0001419_978-94-007-7401-8-Figure4.22-1.png", + "caption": "Fig. 4.22 1PPn3-3PPn3R-type fully-parallel PM with decoupled Sch\u00f6nflies motions defined by MF = SF = 4, RF\u00f0 \u00de \u00bc v1; v2; v3;xa\u00f0 \u00de, TF = 0, NF = 15, limb topology P||Pn3 and P||Pn3 ? R", + "texts": [ + " The revolute joints between links 4 and 5 of limbs G1, G2 and G3 hvave orthogonal axes. The revolute joints between links 4 and 5 of limbs G3 and G4 have parallel axes 7. 1PRRbR\u20133PRRbRR (Fig. 4.18) P||R||Rb||R (Fig. 4.8i) P||R||Rb||R ? R (Fig. 4.9c) Idem No. 1 8. 1PPn2R\u20133PPn2RR (Fig. 4.19) P||Pn2||R (Fig. 4.8j) P||Pn2||R ? R (Fig. 4.9d) 9. 1PPn2R\u20133PPn2RR (Fig. 4.20) P||Pn2||R (Fig. 4.8k) P||Pn2||R ?R (Fig. 4.9e) Idem No. 1 10. 1PPn3\u20133PPn3R (Fig. 4.21) P||Pn3 (Fig. 4.8l) P||Pn3 ? R (Fig. 4.9f) Idem No. 1 11. 1PPn3\u20133PPn3R (Fig. 4.22) P||Pn3 (Fig. 4.8m) P||Pn3 ? R (Fig. 4.9g) Idem No. 1 (continued) 4.2 Topologies with Complex Limbs 425 Table 4.4 (continued) No. PM type Limb topology Connecting conditions 12. 1CRbR\u20133CRbRR (Fig. 4.23) C||Rb||R (Fig. 4.8n) C||Rb||R ? R (Fig. 4.9h) Idem No. 1 13. 4PPaPaR (Fig. 4.24) P||Pa||Pa||R (Fig. 4.10b) P||Pa||Pa ? R (Fig. 4.10a) Idem No. 1 14. 1PaPaRR\u20133PaPaRRR (Fig. 4.25) Pa ? Pa||R||R (Fig. 4.10c) Pa ? Pa||R||R ?? R (Fig. 4.11a) The last revolute joints of the four limbs have parallel axes" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_35_0001419_978-94-007-7401-8-Figure4.4-1.png", + "original_path": "designv11-35/openalex_figure/designv11_35_0001419_978-94-007-7401-8-Figure4.4-1.png", + "caption": "Fig. 4.4 Fully-parallel PMs with decoupled Sch\u00f6nflies motions of types 1PPRR-3PPRRR (a) and 1PRPR-3PRPRR (b) defined by MF = SF = 4, RF\u00f0 \u00de \u00bc v1; v2; v3;xa\u00f0 \u00de, TF = 0, NF = 3, limb topology P ? P ?||R||R and P ? P ?||R||R ? R (a), P||R ? P ?||R and P||R ? P ?||R ? R (b)", + "texts": [ + "1 Limb topology and connecting conditions of the fully-parallel solutions with no idle mobilities presented in Figs. 4.3, 4.4, 4.5, 4.6, 4.7 No. PM type Limb topology Connecting conditions 1. 4PPPR (Fig. 4.3a) P ?P ??P ??R (Fig. 4.1b) P ?P ??P ?|R (Fig. 4.1a) The last revolute joints of the four limbs have parallel axes. The actuated prismatic joints of limbs G1, G2 and G3 hvave orthogonal directions. The actuated prismatic joints of limbs G3 and G4 have parallel directions 2. 4PPPR (Fig. 4.3b) P ?P ??P ??R (Fig. 2.1b) P ?P ??P||R (Fig. 4.1c) Idem No. 1 3. 1PPRR-3PPRRR (Fig. 4.4a) P ?P ?|R||R (Fig. 4.1d) P ?P ?|R||R ?R (Fig. 4.2a) Idem No. 1 4. 1PRPR-3PRPRR (Fig. 4.4b) P||R ?P ?|R (Fig. 4.1e) P||R ?P ?|R ?R (Fig. 4.2b) Idem No. 1 5. 1PRRP-3PRRPR (Fig. 4.5a) P||R||R ?P (Fig. 4.1f) P||R||R ?P ??R (Fig. 4.2c) Idem No. 1 6. 1PRRR-3PRRRR (Fig. 4.5b) P||R||R||R (Fig. 4.1g) P||R||R||R ?R (Fig. 4.2d) Idem No. 1 7. 1PPPR-3PPC (Fig. 4.6a) P ?P ??P ??R (Fig. 4.1b) P ?P ??C (Fig. 4.1h) The last joints of the four limbs have parallel axes. The actuated prismatic joints of limbs G1, G2 and G3 hvave orthogonal directions. The actuated prismatic joints of limbs G3 and G4 have parallel directions 8" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_35_0001419_978-94-007-7401-8-Figure3.31-1.png", + "original_path": "designv11-35/openalex_figure/designv11_35_0001419_978-94-007-7401-8-Figure3.31-1.png", + "caption": "Fig. 3.31 Overactuated PMs with coupled Sch\u00f6nflies motions of types 4RRRPR (a) and 4RRRRP (b) defined by MF \u00bc SF \u00bc 4; RF\u00f0 \u00de \u00bc v1; v2; v3;xd\u00f0 \u00de, TF = 0, NF = 2, limb topology R||R\\R\\P\\kR (a) and R||R\\R||R\\kP (b)", + "texts": [ + "27b) R||R\\R||R\\R (Fig. 3.3b) Idem No. 14 16. 4RRRRR (Fig. 3.28a) R||R||R\\R||R (Fig. 3.3c) The first revolute joints of the four limbs have parallel axes 17. 4RRRRR (Fig. 3.28b) R||R\\R||R||R (Fig. 3.3d) The last revolute joints of the four limbs have parallel axes 18. 4PRRRR (Fig. 3.29a) P||R||R\\R||R (Fig. 3.3e) Idem No. 17 19. 4RPRRR (Fig. 3.29b) R||P||R\\R||R (Fig. 3.3f) Idem No. 17 20. 4RPRRR (Fig. 3.30a) R||P||R\\R||R (Fig. 3.3g) Idem No. 17 21. 4RRPRR (Fig. 3.30b) R||R||P\\R||R (Fig. 3.3h) Idem No. 17 22. 4RRRPR (Fig. 3.31a) R||R\\R\\P\\kR (Fig. 3.3i) Idem No. 17 23. 4RRRRP (Fig. 3.31b) R||R\\R||R\\kP (Fig. 3.3j) Idem No. 17 24. 4PRRRR (Fig. 3.32a) P\\R\\R||R\\R (Fig. 3.3k) The second and the last joints of the four limbs have parallel axes 25. 4RPRRR (Fig. 3.32b) R\\P\\R||R\\R (Fig. 3.3l) Idem No. 14 26. 4RRPRR (Fig. 3.33a) R\\R\\P\\kR\\R (Fig. 3.3m) Idem No. 14 27. 4RRRPR (Fig. 3.33b) R\\R||R\\P\\kR (Fig. 3.3n) Idem No. 14 28. 4RRRRP (Fig. 3.34a) R\\R||R\\R\\P (Fig. 3.3o) The first and the fourth revolute joints of the four limbs have parallel axes 29. 4PRRRR (Fig. 3.34b) P\\R||R\\R||R (Fig. 3" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_35_0001385_978-981-15-6095-8-Figure5.7-1.png", + "original_path": "designv11-35/openalex_figure/designv11_35_0001385_978-981-15-6095-8-Figure5.7-1.png", + "caption": "Fig. 5.7 Vertical pressure at base of geocell mattress (b/B = 12)\u2014Test series D", + "texts": [ + " It can be seen that the strain in the geocell mattress is maximum in the central region underneath the footing. This is because in the region underneath the footing, the geocell reinforcement actively restrains the stress concentration induced yield in the soil mass and thereby its strength is significantly mobilized leading to enhanced load-carrying capacity. The adjacent portions of the geocell mattress only contribute in a secondary manner through frictional and passive resistance developed at the soil\u2013geocell interfaces. Figure 5.7 shows a typical variation of the contact normal pressure (\u03c3 /q) on the subgrade soil below the geocell mattress. It may be seen that the contact pressure, on the subgrade soil, is maximum at the center of the footing and relatively low in the region beyond the loaded area. This is to be expected due to the load dispersion from the footing edges. The geocell-reinforced foundation bed tends to exhibit elastic behavior as the contact pressure responses are found to be almost falling within a narrow range (i" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_35_0000115_iros45743.2020.9340920-Figure1-1.png", + "original_path": "designv11-35/openalex_figure/designv11_35_0000115_iros45743.2020.9340920-Figure1-1.png", + "caption": "Fig. 1. Three cases of the ground contact with the powered exoskeleton; (a) early contact, (b) late contact, and (c) ideal contact", + "texts": [ + " 3) The beginning of the swing (the initial swing) is defined to be 0% of the gait cycle, and the end of the swing (the terminal swing) is 50% of the gait cycle. The start of the stance (the initial contact, 50% of the gait cycle) occurs at the same time as the end of the swing, and the end of the stance (the terminal stance) is 100% of the gait cycle. Therefore, the swing foot is desired to contact the ground at each 50% of the gait cycle in the ideal situation. The actual ground contact time, however, gets different from the desired ground contact time depending on the body inclination angle at the end of the swing. Figure 1.(a) shows late ground contact by leaning backward, which is not safe because the stance motion starts in the air before the weight of the user is transferred to the next leg. Figure 1.(b) shows the early ground contact by inclining forward, which also ruins gait stability by causing a big impact due to continuous extension of the joints. If the ground contact occurs at the desired time as shown in Fig. 1.(c) by adjusting the body inclination angle during walking appropriately, the powered exoskeleton can effectively assist the paraplegic while maintaining stability. Therefore, the gait pattern for the powered exoskeleton need to consider two factors to fully assist the paraplegics: 1) the body inclination angle during walking, and 2) the compensation method for the individual not to have erroneous ground contact time. Figure. 2 shows the configuration of the powered exoskeleton at the moment of the initial contact" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_35_0001385_978-981-15-6095-8-Figure5.3-1.png", + "original_path": "designv11-35/openalex_figure/designv11_35_0001385_978-981-15-6095-8-Figure5.3-1.png", + "caption": "Fig. 5.3 Test geometry and instrumentation details", + "texts": [ + " The geocell mattresses were prepared by cutting the geogrids to required length and height from full rolls and placing them in transverse and diagonal directions as shown in Fig. 5.1, with bodkin joints at the intersections (Simac 1990; Bush et al. 1990)). The bodkins in the present tests are 6 mm wide and 3 mm thick plastic strips made of low-density polypropylene. The joint strength of geocells, obtained through tensile tests, was found to be 4.75 kN/m. Such low strength of joints was adopted to scale down the overall strength of the geocell reinforcement, making it suitable for the model tests. Figure 5.3 depicts the test geometry considered in the present investigation. In total, five different series tests were carried out the details of which are presented in Table 5.1. Within each series, only one parameter was varied, while the others were kept constant. Tests in series A were conducted on unreinforced soil bed with 70% relative density. Under series B, C, D, and E, tests were conducted by varying the pocket size of geocells (d), height of geocell layer (h), width of the geocell mattress (b), and depth to top of geocell layer below footing (u), respectively", + " Each load increment was maintained constant until the footing settlements under that load increment became constant. Settlements of the footing were measured by two dial gauges situated on diagonally opposite sides of the footing. The deformations (heave/settlement) of the fill surface were also measured by dial gauges on both sides of the footing (Fig. 5.4). The strains in the geocell reinforcement were measured through electrical resistance-type strain gauges fixed horizontally at various locations on the geocell walls (Fig. 5.3). They had a gauge length of 10 mm, gauge factor of 2.1 \u00b1 2%, and resistance of 120 \u00b1 0.2 ohms. At each gauge location, the geogrid surface was cleaned by emery paper to remove dust and oily matter if any and then wiped clean with a clean cloth. Subsequently, the strain gauge was pasted with a quick setting adhesive. Lead wires were soldered with strain gauge leads and were connected to a strain meter. The measured strains in micro-strain units (\u00b5strains) were converted to percentage as \u00b5strains \u00d7 100/106", + " The strain measurements are reported at various normalized footing load levels (BPR). The bearing pressure ratio (BPR) is defined as the ratio between the footing pressure with geocell (q) and the ultimate footing pressure (qult) in tests on unreinforced soil. The compressive strains are reported with negative sign and the tensile strains with positive sign. The vertical pressure transmitted to the subgrade soil below the geocell mattress (\u03c3 ) was measured by placing strain gauge-type earth pressure cells below the geocell layer (Fig. 5.3). The overall diameter and thickness of the pressure cells were 60 mm and 10 mm, respectively. In total, three pressure cells were used. One of the pressure cells was kept at the center of the footing and the other two at a distance of 1.5B from the footing centerline on either side of the footing. In the case of unreinforced earth beds, the earth pressures were measured at various depths (corresponding to the base levels of geocells) in different tests for direct comparison with those measured below the geocells" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_35_0001419_978-94-007-7401-8-Figure6.4-1.png", + "original_path": "designv11-35/openalex_figure/designv11_35_0001419_978-94-007-7401-8-Figure6.4-1.png", + "caption": "Fig. 6.4 Maximaly regular fully-parallel PMs with Sch\u00f6nflies motions of types 2PRPRR1PRPR-1RUPU (a) and 2PRRRR-1PRRR-1RUPU (b) defined by MF = SF = 4, (RF) = (v1; v2; v3;xb), TF = 0, NF = 2, limb topology R\\R\\R\\P\\||R\\R and P||R\\P\\||R\\R, P||R\\P\\||R (a), P||R||R||R\\R, P||R||R||R (b)", + "texts": [], + "surrounding_texts": [ + "In the fully-parallel and maximally regular topologies of PMs with Sch\u00f6nflies motions F G1-G2-G3-G4 presented in this section, the moving platform n : nGi (i = 1, 2, 3, 4) is connected to the reference platform 1 : 1Gi : 0 by four simple limbs with four, five or six degrees of connectivity. One linear actuator is combined in the first prismatic or cylindrical pair of limbs G1, G2 and G3, and one rotary actuator in the first revolute pair of limb G4. Limbs G1, G2 and G3 are used for positioning the moving platform and limb G4 for orienting it. Various maximally regular topologies of PMs with Sch\u00f6nflies motions of the moving platform and no idle mobilities can be obtained by using various limb topologies presented in Figs. 4.1, 4.2 and 5.1. Fully-parallel topologies with at least two identical limbs are illustrated in Figs. 6.1, 6.2, 6.3, 6.4, 6.5, 6.6. The limb topology and connecting conditions of these solutions are systematized in Table 6.1, as are their structural parameters in Tables 6.2, 6.3, 6.4. G. Gogu, Structural Synthesis of Parallel Robots, Solid Mechanics and Its Applications 206, DOI: 10.1007/978-94-007-7401-8_6, Springer Science+Business Media Dordrecht 2014 579 580 6 Maximally Regular Topologies with Sch\u00f6nflies Motions 6.1 Fully-Parallel Topologies with Simple Limbs 581 582 6 Maximally Regular Topologies with Sch\u00f6nflies Motions 6.1 Fully-Parallel Topologies with Simple Limbs 583 584 6 Maximally Regular Topologies with Sch\u00f6nflies Motions 6.1 Fully-Parallel Topologies with Simple Limbs 585 T ab le 6. 1 L im b to po lo gy an d co nn ec ti ng co nd it io ns of th e fu ll y- pa ra ll el so lu ti on s w it h no id le m ob il it ie s pr es en te d in F ig s. 6. 1, 6. 2, 6. 3, 6. 4, 6. 5, 6. 6 N o. P M ty pe L im b to po lo gy C on ne ct in g co nd it io ns 1. 3P P P R -1 R P P P (F ig . 6. 1a ) P \\ P ? ? P \\ || R (F ig .4 .1 a) P \\ P ? ? P ? ? R (F ig . 4. 1b ) R ? P ? ? P ? ? P (F ig . 5. 1a ) T he la st jo in ts of th e fo ur li m bs ha ve su pe rp os ed ax es /d ir ec ti on s. T he ac tu at ed pr is m at ic jo in ts of li m bs G 1 , G 2 an d G 3 ha ve or th og on al di re ct io ns 2. 2P P R R R -1 P P R R -1 R P P P (F ig . 6. 1b ) P ? P ? || R ||R ? R (F ig . 4. 2a ) P ? P ? || R ||R (F ig . 4. 1d ) R ? P ? ? P ? ? P (F ig . 5. 1a ) Id em no . 1 3. 2P R P R R -1 P R P R -1 R P P P (F ig . 6. 2a ) P ||R ? P ? || R ? R (F ig . 4. 2b ) P ||R ? P ? || R (F ig . 4. 1e ) R ? P ? ? P ? ? P (F ig . 5. 1a ) Id em no . 1 4. 2P R R R R -1 P R R R -1 R P P P (F ig . 6. 2b ) P ||R ||R ||R ? R (F ig . 4. 2d ) P ||R ||R ||R (F ig . 4. 1g ) R ? P ? ? P ? ? P (F ig . 5. 1a ) Id em no . 1 5. 3P P P R -1 R U P U (F ig . 6. 3a ) P ? P ? ? P ? || R (F ig . 4. 1a ) P ? P ? ? P ? ? R (F ig .4 .1 b) R ? R ? R ? P ? || R ? R (F ig . 5. 1b ) T he fi rs tr ev ol ut e jo in t of G 4 -l im b an d th e la st la st jo in ts of li m bs G 1 , G 2 an d G 3 ha ve pa ra ll el ax es T he la st re vo lu te jo in ts of li m bs G 1 , G 2 an d G 3 ha ve su pe rp os ed ax es . T he ac tu at ed pr is m at ic jo in ts of li m bs G 1 , G 2 an d G 3 ha ve or th og on al di re ct io ns 6. 2P P R R R -1 P P R R -1 R U P U (F ig . 6. 3b ) P ? P ? || R ||R ? R (F ig . 4. 2a ) P ? P ? || R ||R (F ig . 4. 1d ) R ? R ? R ? P ? || R ? R (F ig . 5. 1b ) Id em no . 5 7. 2P R P R R -1 P R P R -1 R U P U (F ig . 6. 4a ) P ||R ? P ? || R ? R (F ig . 4. 2b ) P ||R ? P ? || R (F ig . 4. 1e ) R ? R ? R ? P ? || R ? R (F ig . 5. 1b ) Id em no . 5 8. 2P R R R R -1 P R R R -1 R U P U (F ig . 6. 4b ) P ||R ||R ||R ? R (F ig . 4. 2d ) P ||R ||R ||R (F ig . 4. 1g ) R ? R ? R ? P ? || R ? R (F ig . 5. 1b ) Id em no . 5 9. 2C P R R -1 C P R -1 R P P P (F ig . 6. 5a ) C ? P ? || R ? R (F ig . 4. 2e ) C ? P ? || R (F ig . 4. 1i ) R ? P ? ? P ? ? P (F ig . 5. 1a ) T he la st jo in ts of th e fo ur li m bs ha ve su pe rp os ed ax es /d ir ec ti on s. T he cy li nd ri ca l jo in ts of li m bs G 1 , G 2 an d G 3 ha ve or th og on al ax es (c on ti nu ed ) 586 6 Maximally Regular Topologies with Sch\u00f6nflies Motions T ab le 6. 1 (c on ti nu ed ) N o. P M ty pe L im b to po lo gy C on ne ct in g co nd it io ns 10 . 2C R R R -1 C R R -1 R P P P (F ig . 6. 5b ) C ||R ||R ? R (F ig . 4. 2g ) C ||R ||R (F ig . 4. 1k ) R ? P ? ? P ? ? P (F ig . 5. 1a ) Id em no . 9 11 . 2C P R R -1 C P R -1 R U P U (F ig . 6. 6a ) C ? P ? || R ? R (F ig . 4. 2e ) C ? P ? || R (F ig . 4. 1i ) R ? R ? R ? P ? || R ? R (F ig . 5. 1b ) T he fi rs tr ev ol ut e jo in t of G 4 -l im b an d th e la st la st jo in ts of li m bs G 1 , G 2 an d G 3 ha ve pa ra ll el ax es T he la st re vo lu te jo in ts of li m bs G 1 , G 2 an d G 3 ha ve su pe rp os ed ax es . T he cy li nd ri ca l jo in ts of li m bs G 1 , G 2 an d G 3 ha ve or th og on al ax es 12 . 2 C R R R -1 C R R -1 R U P U (F ig . 6. 6b ) C ||R ||R ? R (F ig . 4. 2g ) C ||R ||R (F ig . 4. 1k ) R ? R ? R ? P ? || R ? R (F ig . 5. 1b ) Id em no . 11 6.1 Fully-Parallel Topologies with Simple Limbs 587 Table 6.2 Structural parametersa of parallel mechanisms in Figs. 6.1 and 6.2 No. Structural parameter Solution Figure 6.1a Figures 6.1b and 6.2 1. m 14 16 2. pi (i = 1, 3) 4 5 3. p2 4 4 4. p4 4 4 5. p 16 18 6. q 3 3 7. k1 4 4 8. k2 0 0 9. k 4 4 10. (RG1) (v1; v2; v3;xb) (v1; v2; v3;xa;xb) 11. (RG2) (v1; v2; v3;xb) (v1; v2; v3;xb) 12. (RG3) (v1; v2; v3;xb) (v1; v2; v3;xb) 13. (RG4) (v1; v2; v3;xb) (v1; v2; v3;xb) 14. SGi (i = 1, 3) 4 5 15. SG2 4 4 16. SG4 4 4 17. rGi (i = 1, 2, 3) 0 0 18. rG4 0 0 19. MGi (i = 1, 3) 4 5 20. MG2 4 4 21. MG4 4 4 22. (RF) (v1; v2; v3;xb) (v1; v2; v3;xb) 23. SF 4 4 24. rl 0 0 25. rF 12 14 26. MF 4 4 27. NF 6 4 28. TF 0 0 29. Pp1 j\u00bc1 fj 4 5 30. Pp2 j\u00bc1 fj 4 4 31. Pp3 j\u00bc1 fj 4 5 32. Pp4 j\u00bc1 fj 4 4 33. Pp j\u00bc1 fj 16 18 a See footnote of Table 2.2 for the nomenclature of structural parameters 588 6 Maximally Regular Topologies with Sch\u00f6nflies Motions Table 6.3 Structural parametersa of parallel mechanisms in Figs. 6.3 and 6.4 No. Structural parameter Solution Figure 6.3a Figures 6.3b and 6.4 1. m 16 18 2. pi (i = 1,3) 4 5 3. p2 4 4 4. p4 6 6 5. p 18 20 6. q 3 3 7. k1 4 4 8. k2 0 0 9. k 4 4 10. (RG1) (v1; v2; v3;xb) (v1; v2; v3;xa;xb) 11. (RG2) (v1; v2; v3;xb) (v1; v2; v3;xb) 12. (RG3) (v1; v2; v3;xb) (v1; v2; v3;xb;xd) 13. (RG4) (v1; v2; v3;xa;xb;xd) (v1; v2; v3;xa;xb;xd) 14. SGi (i = 1, 3) 4 5 15. SG2 4 4 16. SG4 6 6 17. rGi (i = 1, 2, 3) 0 0 18. rG4 0 0 19. MGi (i = 1, 3) 4 5 20. MG2 4 4 21. MG4 6 6 22. (RF) (v1; v2; v3;xb) (v1; v2; v3;xb) 23. SF 4 4 24. rl 0 0 25. rF 14 16 26. MF 4 4 27. NF 4 2 28. TF 0 0 29. Pp1 j\u00bc1 fj 4 5 30. Pp2 j\u00bc1 fj 4 4 31. Pp3 j\u00bc1 fj 4 5 32. Pp4 j\u00bc1 fj 6 6 33. Pp j\u00bc1 fj 18 20 a See footnote of Table 2.2 for the nomenclature of structural parameters 6.1 Fully-Parallel Topologies with Simple Limbs 589 Table 6.4 Structural parametersa of parallel mechanisms in Figs. 6.5 and 6.6 No. Structural parameter Solution Figure 6.5 Figure 6.6 1. m 13 15 2. pi (i = 1,3) 4 4 3. p2 3 3 4. p4 4 6 5. p 15 17 6. q 3 3 7. k1 4 4 8. k2 0 0 9. k 4 4 10. (RG1) (v1; v2; v3;xa;xb) (v1; v2; v3;xa;xb) 11. (RG2) (v1; v2; v3;xb) (v1; v2; v3;xb) 12. (RG3) (v1; v2; v3;xb;xd) (v1; v2; v3;xb;xd) 13. (RG4) (v1; v2; v3;xb) (v1; v2; v3;xa;xb;xd) 14. SGi (i = 1, 3) 5 5 15. SG2 4 4 16. SG4 4 6 17. rGi (i = 1, 2, 3) 0 0 18. rG4 0 0 19. MGi (i = 1, 3) 5 5 20. MG2 4 4 21. MG4 4 6 22. (RF) (v1; v2; v3;xb) (v1; v2; v3;xb) 23. SF 4 4 24. rl 0 0 25. rF 14 16 26. MF 4 4 27. NF 4 2 28. TF 0 0 29. Pp1 j\u00bc1 fj 5 5 30. Pp2 j\u00bc1 fj 4 4 31. Pp3 j\u00bc1 fj 5 5 32. Pp4 j\u00bc1 fj 4 6 33. Pp j\u00bc1 fj 18 20 a See footnote of Table 2.2 for the nomenclature of structural parameters 590 6 Maximally Regular Topologies with Sch\u00f6nflies Motions" + ] + }, + { + "image_filename": "designv11_35_0000228_j.matpr.2021.02.620-Figure4-1.png", + "original_path": "designv11-35/openalex_figure/designv11_35_0000228_j.matpr.2021.02.620-Figure4-1.png", + "caption": "Fig. 4. Boundary Conditions.", + "texts": [ + "2 KSz = Size factor = 1) Sf \u00bc F:O:S ra 1 F:O:S rm rUl \u00bc Fatigue Strength The build model of NX11 Unigraphics has been imported to ANSYS work bench environment Select simulation for analysis of connecting rod Made the connection between the cup of the connecting rod and the whole parts Select the meshing method of tetrahedral Apply the boundary condition Apply the option of simulation Select the result of simulation result. In order to ascertain the accurate results in the simulation, it is essential that the finite element model is well prepared, with the appropriate element size and connectors. Therefore, the selection of the correct element size should be carefully reviewed. For all practical purposes, the maximum force acting on connecting rod is taken from gas pressure applying on the piston neglecting the inertial effects. Fig. 4 illustrates the maximum gas force about 30.4 KN applying on the piston pin of the connecting rod and constrained at the big end. Further, the analysis of connecting rod has been carried out by assigning the 42CrMo4 material. The Fig. 5 shows the total deformation, Von-Mises stress, life cycle and factor of safety obtained for the connecting rod. The stress-life method usually used for millions of cycles, where the stresses are elastic and this method is stated to infinite design. It is based on the fatigue limit or endurance limit of the material" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_35_0001419_978-94-007-7401-8-Figure6.31-1.png", + "original_path": "designv11-35/openalex_figure/designv11_35_0001419_978-94-007-7401-8-Figure6.31-1.png", + "caption": "Fig. 6.31 2PPaRRR-1PPaRR-1RPaPatP-type maximaly regular fully-parallel PM with Sch\u00f6nflies motions defined by MF = SF = 4, (RF) = (v1; v2; v3;xb), TF = 0, NF = 19, limb topology P||Pa||R||R\\R, P||Pa||R||R\\R and R||Pa||Pat||P", + "texts": [ + " 1 4. 3PPPaR-1RPaPatP (Fig. 6.28b) P\\P\\||Pa\\R (Fig. 4.8a) P\\P\\||Pa||R, (Fig. 4.8c) R||Pa||Pat||P (Fig. 5.4l) Idem no. 1 5. 3PPaPR-1RPaPaP (Fig. 6.29a) P||Pa\\P\\\\R (Fig. 4.8d) P||Pa\\P\\||R (Fig. 4.8e) R||Pa||Pa||P (Fig. 5.4k) Idem no. 1 6. 3PPaPR-1RPaPatP (Fig. 6.29b) P||Pa\\P\\\\R (Fig. 4.8d) P||Pa\\P\\||R (Fig. 4.8e) R||Pa||Pat||P (Fig. 5.4l) Idem no. 1 7. 2PPaRRR-1PPaRR-1RPaPaP (Fig. 6.30) P||Pa||R||R\\R (Fig. 4.9a) P||Pa||R||R (Fig. 4.8g) R||Pa||Pa||P (Fig. 5.4k) Idem no. 1 8. 2PPaRRR-1PPaRR1RPaPatP (Fig. 6.31) P||Pa||R||R\\R (Fig. 4.9a) P||Pa||R||R (Fig. 4.8g) R||Pa||Pat||P (Fig. 5.4l) Idem no. 1 9. 3PPPaR-1RPPaPa (Fig. 6.32a) P\\P\\||Pa\\||R (Fig. 4.8b) P\\P\\||Pa||R, (Fig. 4.8c) R||P||Pa||Pa (Fig. 5.4m) The last revolute joints of limbs G1, G2 and G3 and the revolute joint between links 7D and 9 have superposed axes. The actuated prismatic joints of limbs G1, G2 and G3 have orthogonal directions 10. 3PPPaR-1PRPaPa (Fig. 6.32b) P\\P\\||Pa\\R (Fig. 4.8a) P\\P\\||Pa||R, (Fig. 4.8c) R||P||Pa||Pa (Fig. 5.4m) Idem no" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_35_0001419_978-94-007-7401-8-Figure6.15-1.png", + "original_path": "designv11-35/openalex_figure/designv11_35_0001419_978-94-007-7401-8-Figure6.15-1.png", + "caption": "Fig. 6.15 Maximaly regular fully-parallel PMs with Sch\u00f6nflies motions of types 2CPRR-1CPR1RPaPaP (a) and 2CPRR-1CPR-1RPaPatP (b) defined by MF = SF = 4, (RF) = (v1; v2; v3;xb), TF = 0, NF = 10, limb topology C\\P\\||R\\R, C\\P\\||R and R||Pa||Pa||P (a), R||Pa||Pat||P (b)", + "texts": [ + " 2PRPRR-1PRPR-1RPPaPa (Fig. 6.13a) P||R\\P\\||R\\R (Fig. 4.2b) P||R\\P\\||R (Fig. 4.1e) R||P||Pa||Pa (Fig. 5.4m) Idem no. 9 14. 2PRPRR-1PRPR-1RPPaPat (Fig. 6.13b) P||R\\P\\||R\\R (Fig. 4.2b) P||R\\P\\||R (Fig. 4.1e) R||P||Pa||Pat (Fig. 5.4n) Idem no. 9 15. 2PRRRR-1PRRR-1RPPaPa (Fig. 6.14a) P||R||R||R\\R (Fig. 4.2d) P||R||R||R (Fig. 4.1g) R||P||Pa||Pa (Fig. 5.4m) Idem no. 9 16. 2PRRRR-1PRRR1RPaPatP (Fig. 6.14b) P||R||R||R\\R (Fig. 4.2d) P||R||R||R (Fig. 4.1g) R||P||Pa||Pat (Fig. 5.4n) Idem no. 9 17. 2CPRR-1CPR-1RPaPaP (Fig. 6.15a) C\\P\\||R\\R (Fig. 4.2e) C ?P ?||R (Fig. 4.1i) R||Pa||Pa||P (Fig. 5.4k) Idem no. 1 18. 2CPRR-1CPR1RPaPatP (Fig. 6.15b) C ?P ?||R ?R (Fig. 4.2e) C ?P ?||R (Fig. 4.1i) R||Pa||Pat||P (Fig. 5.4l) Idem no. 1 19. 2CRRR-1CRR-1RPaPaP (Fig. 6.16a) C||R||R ?R (Fig. 4.2g) C||R||R (Fig. 4.1k) R||Pa||Pa||P (Fig. 5.4k) Idem no. 1 20. 2CRRR-1CRR1RPaPatP (Fig. 6.16b) C||R||R ?R (Fig. 4.2g) C||R||R (Fig. 4.1k) R||Pa||Pat||P (Fig. 5.4l) Idem no. 1 Table 6.6 Limb topology and connecting conditions of the fully-parallel solutions with no idle mobilities presented in Figs. 6.17, 6.18, 6.19, 6.20, 6.21, 6.22, 6.23, 6.24, 6.25, 6.26 No" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_35_0001419_978-94-007-7401-8-Figure3.19-1.png", + "original_path": "designv11-35/openalex_figure/designv11_35_0001419_978-94-007-7401-8-Figure3.19-1.png", + "caption": "Fig. 3.19 Overactuated PMs with coupled Sch\u00f6nflies motions of types 4PRPP (a) and 4RPPP (b) defined by MF \u00bc SF \u00bc 4; RF\u00f0 \u00de \u00bc v1; v2; v3;xd\u00f0 \u00de, TF = 0, NF = 6, limb topology P||R\\P\\\\P (a) and R\\P\\\\P\\\\P (b)", + "texts": [ + " 4PPPR (Fig. 3.15a) P\\P\\\\P\\kR (Fig. 3.1h0) Idem No. 9 24. 4PPPR (Fig. 3.15b) P\\P\\\\P\\\\R (Fig. 3.1w) Idem No. 9 25. 4PPPR (Fig. 3.16a) P\\P\\\\P||R (Fig. 3.1x) Idem No. 9 26. 4PPRP (Fig. 3.16b) P\\P||R\\P (Fig. 3.1y) The revolute joints of the four limbs have parallel axes 27. 4PPRP (Fig. 3.17a) P\\P\\\\R||P (Fig. 3.1z) Idem No. 26 28. 4PPRP (Fig. 3.17b) P\\P\\kR\\P (Fig. 3.1a0) Idem No. 26 29. 4PRPP (Fig. 3.18a) P\\R||P\\P (Fig. 3.1b0) Idem No. 26 30. 4PRPP (Fig. 3.18b) P\\R\\P\\kP (Fig. 3.1c0) Idem No. 26 31. 4PRPP (Fig. 3.19a) P||R\\P\\\\P (Fig. 3.1d0) Idem No. 26 32. 4RPPP (Fig. 3.19b) R\\P\\\\P\\\\P (Fig. 3.1e0) Idem No. 26 33. 4RPPP (Fig. 3.20a) R\\P\\kP\\\\P (Fig. 3.1f0) Idem No. 26 34. 4RPPP (Fig. 3.20b) R||P\\P\\\\P (Fig. 3.1g0) Idem No. 26 3.1 Topologies with Simple Limbs 243 Table 3.2 Limb topology and connecting conditions of the overactuated solutions with no idle mobilities presented in Figs. 3.21, 3.22, 3.23, 3.24, 3.25, 3.26, 3.27, 3.28, 3.29, 3.30, 3.31, 3.32, 3.33, 3.34, 3.35, 3.36, 3.37, 3.38, 3.39, 3.40, 3.41, 3.42, 3.43, 3.44, 3.45, 3.46, 3.47, 3.48, 3.49 No. PM type Limb topology Connecting conditions 1" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_35_0001419_978-94-007-7401-8-Figure3.42-1.png", + "original_path": "designv11-35/openalex_figure/designv11_35_0001419_978-94-007-7401-8-Figure3.42-1.png", + "caption": "Fig. 3.42 Overactuated PMs with coupled Sch\u00f6nflies motions of types 4RRRPP (a) and 4PRRPR (b) defined by MF \u00bc SF \u00bc 4; RF\u00f0 \u00de \u00bc v1; v2; v3;xd\u00f0 \u00de, TF = 0, NF = 2, limb topology R||R\\R\\P\\\\P (a) and P||R||R\\P\\\\R (b)", + "texts": [ + "38a) R\\R||P||R\\P (Fig. 3.3w) Idem No. 16 37. 4RPRRP (Fig. 3.38b) R\\P||R||R\\P (Fig. 3.3x) Idem No. 16 38. 4RRPRP (Fig. 3.39a) R\\R||P||R\\P (Fig. 3.3y) Idem No. 16 39. 4RRRPP (Fig. 3.39b) R\\R||R||P\\P (Fig. 3.3z) Idem No. 16 40. 4PRRRP (Fig. 3.40a) P||R||R\\R\\P (Fig. 3.3z1) The fourth joints of the four limbs have parallel axes 41. 4RPRRP (Fig. 3.40b) R||P||R||R\\P (Fig. 3.3a) Idem No. 40 42. 4RPRRP (Fig. 3.41a) R||P||R\\R\\P (Fig. 3.3b0) Idem No. 40 43. 4RRPRP (Fig. 3.41b) R||R||P\\R\\P (Fig. 3.3c0) Idem No. 40 44. 4RRRPP (Fig. 3.42a) R||R\\R\\P\\\\P (Fig. 3.3d0) The third joints of the four limbs have parallel axes 45. 4PRRPR (Fig. 3.42b) P||R||R\\P\\\\R (Fig. 3.3e0) Idem No. 17 46. 4RPRPR (Fig. 3.43a) R||P||R\\P\\\\R (Fig. 3.3f0) Idem No. 17 47. 4RRPPR (Fig. 3.43b) R||R||P\\P\\\\R (Fig. 3.3g0) Idem No. 17 48. 4RPRPR (Fig. 3.44a) R||P||R\\P\\\\R (Fig. 3.3h0) Idem No. 17 49. 4RPRPR (Fig. 3.44b) P\\P\\kR||R\\\\R (Fig. 3.3i0) Idem No. 17 50. 4PPRRR (Fig. 3.44c) P\\P||R||R\\\\R (Fig. 3.3j0) Idem No. 17 51. 4CRRR (Fig. 3.45a) C||R\\R||R (Fig. 3.3k0) Idem No. 17 52. 4RCRR (Fig. 3.45b) R||C\\R||R (Fig. 3.3l0) Idem No. 17 53. 4RRCR (Fig. 3.46a) R||R\\C||R (Fig" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_35_0000497_dese51703.2020.9450784-Figure2-1.png", + "original_path": "designv11-35/openalex_figure/designv11_35_0000497_dese51703.2020.9450784-Figure2-1.png", + "caption": "Fig. 2. (a) An original KUKA IIWA LBR manipulator model; (b) A modified model.", + "texts": [ + " We developed a virtual implementation of KUKA IIWA LBR manipulator [13], [14] based model in Gazebo simulator with soft body features implemented. DART physics engine capabilities were used for the soft body features. The developed system was tested by simulating palpation of abdomen soft model with variable stiffness over its surface. The system successfully performed palpation of the abdomen and was able to detect an area of high stiffness. Simulation environment in the Gazebo (Fig. 1) contains a modified KUKA IIWA LBR manipulator model (Fig. 2) from 978-1-6654-2238-3/20/$31.00 \u00a92020 IEEE 200 20 20 1 3t h In te rn at io na l C on fe re nc e on D ev el op m en ts in e Sy st em s E ng in ee rin g (D eS E) | 97 8- 1- 66 54 -2 23 8- Authorized licensed use limited to: National University of Singapore. Downloaded on July 04,2021 at 00:54:15 UTC from IEEE Xplore. Restrictions apply. our previous work [15]. The manipulator is placed on a cubic supporting base that shifts the workspace in order to cover surgical table surface. A Kinect camera model with Robot Operating System (ROS) depth camera plugin is mounted on a tripod to enable palpated surface geometry data acquisition" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_35_0001419_978-94-007-7401-8-Figure3.67-1.png", + "original_path": "designv11-35/openalex_figure/designv11_35_0001419_978-94-007-7401-8-Figure3.67-1.png", + "caption": "Fig. 3.67 4RPPaP-type overactuated PM with coupled Sch\u00f6nflies motions defined by MF \u00bc SF \u00bc 4; RF\u00f0 \u00de \u00bc v1; v2; v3;xd\u00f0 \u00de, TF = 0, NF = 18, limb topology R||P||Pa\\P", + "texts": [ + "62a) R\\P\\kPa||P (Fig. 3.50p) Idem No. 1 17. 4PPPaR (Fig. 3.62b) P\\P\\kPa||R (Fig. 3.50q) Idem No. 1 18. 4PPPaR (Fig. 3.63a) P\\P||Pa||R (Fig. 3.50r) Idem No. 1 19. 4PPaRR (Fig. 3.63b) P||Pa||R||R (Fig. 3.50s) Idem No. 1 20. 4PPaRP (Fig. 3.64) P\\Pa||R||P (Fig. 3.50t) Idem No. 1 21. 4PPaPR (Fig. 3.65) P\\Pa||P||R (Fig. 3.50u) Idem No. 1 22. 4PPaPR (Fig. 3.66a) P||Pa\\P\\kR (Fig. 3.50v) Idem No. 1 23. 4PPaRP (Fig. 3.66b) P||Pa\\R\\P (Fig. 3.50i0) The before last joints of the four limbs have parallel axes 24. 4RPPaP (Fig. 3.67) R||P||Pa\\P (Fig. 3.50w) Idem No. 1 25. 4RPRPa (Fig. 3.68) R||P||R||Pa (Fig. 3.50x) Idem No. 1 26. 4PaPPR (Fig. 3.69a) Pa||P\\P\\kR (Fig. 3.50y) Idem No. 1 27. 4PaPPR (Fig. 3.69b) Pa\\P\\kP||R (Fig. 3.50z) Idem No. 1 28. 4PaPRP (Fig. 3.70) Pa\\P\\kR||P (Fig. 3.50z1) Idem No. 1 29. 4PaRRP (Fig. 3.71) Pa||R||R||P (Fig. 3.50a0) Idem No. 1 30. 4PaRPR (Fig. 3.72) Pa||R||P||R (Fig. 3.50b0) Idem No. 1 31. 4PaPRP (Fig. 3.73a) Pa||P||R\\P (Fig. 3.50c0) Idem No. 1 32. 4PaRPP (Fig. 3.73b) Pa||R\\P\\kP (Fig. 3.50d0) Idem No" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_35_0001419_978-94-007-7401-8-Figure3.79-1.png", + "original_path": "designv11-35/openalex_figure/designv11_35_0001419_978-94-007-7401-8-Figure3.79-1.png", + "caption": "Fig. 3.79 4PaPRR-type overactuated PMs with coupled Sch\u00f6nflies motions defined by MF \u00bc SF \u00bc 4; RF\u00f0 \u00de \u00bc v1; v2; v3;xd\u00f0 \u00de, TF = 0, NF = 18, limb topology Pa||P\\R||R (a) and Pa\\P\\\\R||R (b)", + "texts": [ + "50h0) The third joints of the four limbs have parallel axes 37. 4PPPaR (Fig. 3.76a) P\\P||Pa\\\\R (Fig. 3.50j0) The last joints of the four limbs have parallel axes (continued) 3.2 Topologies with Complex Limbs 359 Table 3.5 (continued) No. PM type Limb topology Connecting conditions 38. 4PPPaR (Fig. 3.76b) P\\P\\kPa\\\\R (Fig. 3.50k0) Idem No. 37 39. 4PaPPR (Fig. 3.77a) Pa||P\\P\\\\R (Fig. 3.50l0) Idem No. 37 40. 4PaPPR (Fig. 3.77b) Pa\\P\\kP\\\\R (Fig. 3.50m0) Idem No. 37 41. 4PaRPR (Fig. 3.78) Pa\\R\\P\\kR (Fig. 3.50n0) Idem No. 37 42. 4PaPRR (Fig. 3.79a) Pa||P\\R||R (Fig. 3.50o0) Idem No. 37 43. 4PaPRR (Fig. 3.79b) Pa\\P\\\\R||R (Fig. 3.50p0) Idem No. 37 44. 4PaRRP (Fig. 3.80a) Pa\\R||R\\P (Fig. 3.50q0) Idem No. 23 45. 4RPaRR (Fig. 3.80b) R\\Pa\\kR||R (Fig. 3.50s0) Idem No. 37 46. 4PaRRR (Fig. 3.81) Pa\\R||R||R (Fig. 3.50r0) Idem No. 37 47. 4PaPRP (Fig. 3.82a) Pa||P\\R\\P (Fig. 3.50t0) Idem No. 23 48. 4PaPRP (Fig. 3.82b) Pa\\P\\\\R\\P (Fig. 3.50u0) Idem No. 23 49. 4CPaP (Fig. 3.83a) C||Pa\\P (Fig. 3.50v0) The cylindrical joints of the four limbs have parallel directions 50. 4PaCP (Fig. 3.83b) Pa||C\\P (Fig. 3.50w0) Idem No" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_35_0001737_8.10872-Figure2-1.png", + "original_path": "designv11-35/openalex_figure/designv11_35_0001737_8.10872-Figure2-1.png", + "caption": "FIG. 2. External loads.", + "texts": [ + " In the graphic integration the effect of a variable moment of inertia can easily be taken into account. Solving Eqs. (1), one obtains the values of the unknown moments that act at the imaginary joints. The above is the procedure to be followed when the ring is not symmetric. If symmetric or antisymmetric loads act upon a symmetric ring, the calculations are materially simplified. In many cases it is advantageous to resolve nonsymmetric loads into symmetric and anti symmetric groups of loads, as discussed in a recent paper published by Newell.3 This can be done as shown in Fig. 2, where load (a) is equivalent to the simultaneous action of loads (b) and (c). In such a case, however, little is gained if the bending moments caused by the un knowns are nonsymmetric. To remedy this situation it is suggested that the unknowns also be combined into symmetric and antisymmetric groups. In Fig. 3 the three imaginary joints are located at points A, 3, and C, respectively. The three unknown moments are denoted MA} MB, and Mc, respectively. The combination MA = MB = Mc = 1 is symmetric and causes a uniform bending moment distribution over the ring", + " If, how ever, Bo in the movable system is just a linear combi nation of Bo in the completely cut system and of Bm, BPf and Bv, then Eqs. (1) and (9) determine the values of Xm, Xp, and Xv which make the strain energy a minimum. It follows from Eqs. (1) and (la) (or (9)) that in the case of symmetric rings all coefficients 8jk vanish in whichever one of the two moments Mj} Mk (or Bjf Bk) is symmetric and the other antisymmetric. Also, it is allowable to calculate all the nonvanishing coefficients for one-half of the ring. TORSION The first numerical example calculated here is shown in Fig. 2b. For the sake of simplicity it is assumed that the distance between the sheet covering and the center line of the ring is negligibly small and that the ring has a constant cross-section. The area A en closed by the center line of the ring is A = 2752 sq.in. The counterclockwise external torque is T = 3000 in. lbs. Consequently, the uniform shear flow in the covering becomes s = T/2A = 0.544 lb. per in. The imaginary pin-joints are assumed at points A, B, and C. The former two are 30 in. above C as shown in Fig", + "30 Xv = 0 or Xv = -2 .75 . The resultant moment curve in the complete structure has the ordinates B0 \u2014 2.75 Bv and is denoted Bres in Fig. 5. It is seen that the guess for the location of the in flection point was a good one, and the final moments Bres. differ but little from those represented by the B0 curve. The conversion factor is 0.544 X (102). Con sequently, the greatest positive moment is M = 666 in. lbs., and the greatest negative moment is M = \u2014319 in. lbs. BENDING The second numerical example is shown in Fig. 2c. For the sake of simplicity it is assumed that the fuselage is a single shell of a (fully effective) wall thickness t. A numerical calculation gives I/t = 55,000 cu.in. for one-half of the fuselage section. The shear flow 5 in a section x is calculated by the known formula 5 = (S/I)F, where S = 100 lbs. is the total shear force; / , the moment of inertia; and F, the first moment of the portion beyond the section x of the cross-section of the fuselage. Consequently, 5 = (100/55,000) (F/t). As a reference shear flow sref^ = s*y the total shear force divided by the total length L = 96" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_35_0001419_978-94-007-7401-8-Figure5.14-1.png", + "original_path": "designv11-35/openalex_figure/designv11_35_0001419_978-94-007-7401-8-Figure5.14-1.png", + "caption": "Fig. 5.14 2PaRRRR-1PaRRR-1RUPU-type fully-parallel PMs with uncoupled Sch\u00f6nflies motions defined by MF = SF = 4, RF\u00f0 \u00de \u00bc \u00f0v1; v2; v3;xb\u00de, TF = 0, NF = 11, limb topology Pa\\R||R||R\\kR, Pa\\R||R||R and R\\R\\R\\P\\kR\\R", + "texts": [ + " 5.1b) 10. 2PaPRRR1PaPRR1RUPU (Fig. 5.12b) Pa||P\\R||R\\\\R (Fig. 5.3b) Idem No. 9 Pa||P\\R||R (Fig. 5.2f) R\\R\\R\\P\\kR\\R (Fig. 5.1b) 11. 2PaRPRR1PaRPR1RUPU (Fig. 5.13) Pa\\R\\P\\kR\\R (Fig. 5.3c) The first revolute joint of G4 limb and the last revolute joints of limbs G1, G2 and G3 have parallel axes. The last joints of of limbs G1, G2 and G3 have superposed axes, and their revolute joints between links 4 and 5 have orthogonal axes Pa\\R\\P\\kR (Fig. 5.2g) R\\R\\R\\P\\kR\\R (Fig. 5.1b) 12. 2PaRRRR1PaRRR1RUPU (Fig. 5.14) Pa\\R||R||R\\kR (Fig. 5.3d) Idem No. 11 Pa\\R||R||R (Fig. 5.2h) R\\R\\R\\P\\kR\\R (Fig. 5.1b) 5.1 Fully-Parallel Topologies 527 Table 5.2 Limb topology and connecting conditions of the fully-parallel solutions with no idle mobilities presented in Figs. 5.15, 5.16, 5.17, 5.18, 5.19, 5.20, 5.21, 5.22, 5.23, 5.24, 5.25, 5.26, 5.27, 5.28, 5.29, 5.30 No. PM type Limb topology Connecting conditions 1. 3PaPPR-1RPaPaP (Fig. 5.15a) Pa||P\\P\\kR (Fig. 5.2a) The last joints of the four limbs have superposed axes/directions Pa||P\\P\\\\R (Fig" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_35_0001730_s0025557200233147-Figure4-1.png", + "original_path": "designv11-35/openalex_figure/designv11_35_0001730_s0025557200233147-Figure4-1.png", + "caption": "FIG. 4.", + "texts": [ + " Fix A OB horizontally and let AOD, BOG make angles 8 and with the vertical, then (DAB = (Q0+6), \\ABC = (90+ with the fixed plane A OB. Applying VI to &AXB, e ' d> i t an - t an ^ = = constant. 2 2 cos a Thus from relation (2), which reduces to * J\u00b1d cos a dd 1 - sin2 a . sin2 \\Q The velocity ratio fluctuates between l'/cos a and cos a. When a = 30\u00b0, sin a = \\, cos a = i^/3, w' \u2022\u2022 a = 45\u00b0, sin a = COS a = 1 IJ2 , a / \u2022\u2022 7 + cos 8 - 272a. , 3 + cos 0' a = 60\u00b0, sin a = 473 , cos a = A, oS =\u2022 \u2014 ~- T;- ^v J ;>+3cos0 Consider also the case in which y = a, S=j8" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_35_0000596_s43452-021-00255-x-Figure20-1.png", + "original_path": "designv11-35/openalex_figure/designv11_35_0000596_s43452-021-00255-x-Figure20-1.png", + "caption": "Fig. 20 Fracture, failure mechanism, and modes in the riveted and hybrid joints; a basic failure modes for the pinned-joint configuration, b mixed failure mode in riveted woven composites [82], c fracture surface of hybrid PR and hybrid SPR joints [83], d failure modes of Z-pins reinforced composites SLJ [84]", + "texts": [ + " 1 3 It can be seen from Table\u00a013 that the percentage of changes in the dissipated energy based on the fracture force connection in the hybrid joints has increased significantly compared with simple rivet joints except for samples 53 and 31 in other connections, which indicates more strength of hybrid joints. Therefore, the use of hybrid joints with 4- and 5-rivet layouts with codes Sra53 and Sra31 is not recommended. An example of failure in the double-lap adhesive joints is shown in Fig.\u00a021b. Mechanically fastened joints under tensile loads generally fail in three basic modes, referred to as net-tension mode, shear-out mode, and bearing mode. These modes are shown in Fig.\u00a020a, b shows the damage state and the fracture modes. One can notice the final deformation of the rivet due to shearing. Experimental tests of the hybrid adhesive-rivet joints for the same specimens show similar behavior. After the initial linear increase, the specimens attained the maximum force at the displacement of approximately 3.5\u00a0mm. The glass fiber/epoxy hybrid composite joint [70] showed a higher sensitivity to degradation in salt-fog environmental conditions than the carbon fiber/epoxy one", + " As concerns the damage mechanisms, this joint experienced catastrophic cleavage/net tension failure mode regardless of the aging time whereas the progressive bearing failure was the mechanism observed in the carbon fiber/epoxy hybrid composite joint during the entire experimental campaign. Corrosion phenomena not significantly affected the degradation performances of both joints. In the mechanical tests of riveted woven composites, the joint\u2019s failure mode was a mixed failure mode because the rivet pull-through, rivet shear failure, and bearing damage were observed as the predominant failure mechanisms (Fig.\u00a020b)[82]. Fracture surface for hybrid PR and hybrid SPR joints in Fig.\u00a020c illustrated [83]. In the newest research by Yang et\u00a0al. [85], failure modes of Z-pins reinforced composites SLJ (FTF-fiber failure, IF-interface failure, CF-cohesive failure) presented in Fig.\u00a020d. When the joint is subjected to tensile loading, part of the force is tolerated by the rivet, and in simple rivet joints because the rivet strength is greater than the strength of the nanocomposite plates (adherends) due to the stress concentration in the holes, the adherends were produced using Plastic injection method and the distribution of nano-clay particles in the polymer field is more likely to be uniform, so the initial crack is most likely caused by the holes and as the sample continues to stretch by the machine test, it grows Archives of Civil and Mechanical Engineering (2021) 21:105 1 3 105 Page 26 of 29 and failures in the holes" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_35_0001419_978-94-007-7401-8-Figure6.14-1.png", + "original_path": "designv11-35/openalex_figure/designv11_35_0001419_978-94-007-7401-8-Figure6.14-1.png", + "caption": "Fig. 6.14 Maximaly regular fully-parallel PMs with Sch\u00f6nflies motions of types 2PRRRR1PRRR-1RPPaPa (a) and 2PRRRR-1PRRR-1RPPaPat (b) defined by MF = SF = 4, (RF) = (v1; v2; v3;xb), TF = 0, NF = 10, limb topology P||R||R||R\\R, P||R||R||R and R||P||Pa||Pa (a), R||P||Pa||Pat (b)", + "texts": [ + " 9 (continued) 612 6 Maximally Regular Topologies with Sch\u00f6nflies Motions Table 6.5 (continued) No. PM type Limb topology Connecting conditions 12. 2PPRRR-1PPRR-1RPPaPat (Fig. 6.12b) P\\P\\||R||R\\R (Fig. 4.2a) P\\P\\||R||R (Fig. 4.1d) R||P||Pa||Pat (Fig. 5.4n) Idem no. 9 13. 2PRPRR-1PRPR-1RPPaPa (Fig. 6.13a) P||R\\P\\||R\\R (Fig. 4.2b) P||R\\P\\||R (Fig. 4.1e) R||P||Pa||Pa (Fig. 5.4m) Idem no. 9 14. 2PRPRR-1PRPR-1RPPaPat (Fig. 6.13b) P||R\\P\\||R\\R (Fig. 4.2b) P||R\\P\\||R (Fig. 4.1e) R||P||Pa||Pat (Fig. 5.4n) Idem no. 9 15. 2PRRRR-1PRRR-1RPPaPa (Fig. 6.14a) P||R||R||R\\R (Fig. 4.2d) P||R||R||R (Fig. 4.1g) R||P||Pa||Pa (Fig. 5.4m) Idem no. 9 16. 2PRRRR-1PRRR1RPaPatP (Fig. 6.14b) P||R||R||R\\R (Fig. 4.2d) P||R||R||R (Fig. 4.1g) R||P||Pa||Pat (Fig. 5.4n) Idem no. 9 17. 2CPRR-1CPR-1RPaPaP (Fig. 6.15a) C\\P\\||R\\R (Fig. 4.2e) C ?P ?||R (Fig. 4.1i) R||Pa||Pa||P (Fig. 5.4k) Idem no. 1 18. 2CPRR-1CPR1RPaPatP (Fig. 6.15b) C ?P ?||R ?R (Fig. 4.2e) C ?P ?||R (Fig. 4.1i) R||Pa||Pat||P (Fig. 5.4l) Idem no. 1 19. 2CRRR-1CRR-1RPaPaP (Fig. 6.16a) C||R||R ?R (Fig. 4.2g) C||R||R (Fig. 4.1k) R||Pa||Pa||P (Fig. 5.4k) Idem no. 1 20. 2CRRR-1CRR1RPaPatP (Fig. 6.16b) C||R||R ?R (Fig. 4.2g) C||R||R (Fig" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_35_0001419_978-94-007-7401-8-Figure6.20-1.png", + "original_path": "designv11-35/openalex_figure/designv11_35_0001419_978-94-007-7401-8-Figure6.20-1.png", + "caption": "Fig. 6.20 Maximaly regular fully-parallel PMs with Sch\u00f6nflies motions of types 2CRRR-1CRR1CPaPa (a) and 2CRRR-1CRR-1CPaPat (b) defined by MF = SF = 4, (RF) = (v1; v2; v3;xb), TF = 0, NF = 10, limb topology C||R||R\\R, C||R||R and C||Pa||Pa (a), C||Pa||Pat (b)", + "texts": [ + "2 Fully-Parallel Topologies with Simple and Complex Limbs 613 Table 6.6 (continued) No. PM type Limb topology Connecting conditions 4. 2PPRRR-1PPRR-1CPaPat (Fig. 6.18b) P ?P ?||R||R ?R (Fig. 4.2a) P ?P ?||R||R (Fig. 4.1d) C||Pa||Pat (Fig. 5.4p) Idem no. 1 5. 2CPRR-1CPR-1CPaPa (Fig. 6.19a) C ?P ?||R ?R (Fig. 4.2e) C ?P ?||R (Fig. 4.1i) C||Pa||Pa (Fig. 5.4o) Idem no. 1 6. 2CPRR-1CPR-1CPaPat (Fig. 6.19b) C ?P ?||R ?R (Fig. 4.2e) C ?P ?||R (Fig. 4.1i) C||Pa||Pat (Fig. 5.4p) Idem no. 1 7. 2CRRR-1CRR-1CPaPa (Fig. 6.20a) C||R||R ?R (Fig. 4.2g) C||R||R (Fig. 4.1k) C||Pa||Pa (Fig. 5.4o) Idem no. 1 8. 2CRRR-1CRR-1CPaPat (Fig. 6.20b) C||R||R ?R (Fig. 4.2g) C||R||R (Fig. 4.1k) C||Pa||Pat (Fig. 5.4p) Idem no. 1 9. 3PPPaR-1RPPP (Fig. 6.21a) P ?P ?||Pa ?||R (Fig. 4.8b) P ?P ?||Pa||R, (Fig. 4.8c) R ?P\\\\P\\\\P (Fig. 5.1a) The last joints of the four limbs have superposed axes/ directions. The actuated prismatic joints of limbs G1, G2 and G3 have orthogonal directions 10. 3PPPaR-1RPPP (Fig. 6.21b) P ?P ?||Pa\\\\R (Fig. 4.8a) P ?P ?||Pa||R, (Fig. 4.8c) R\\\\P\\\\P\\\\P (Fig. 5.1a) Idem no. 9 11. 3PPaPR-1RPPP (Fig. 6.22a) P||Pa ?P\\\\R (Fig. 4" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_35_0001419_978-94-007-7401-8-Figure5.71-1.png", + "original_path": "designv11-35/openalex_figure/designv11_35_0001419_978-94-007-7401-8-Figure5.71-1.png", + "caption": "Fig. 5.71 3PaPPR-1CPaPa-type fully-parallel PMs with uncoupled Sch\u00f6nflies motions defined by MF = SF = 4, RF\u00f0 \u00de \u00bc \u00f0v1; v2; v3;xb\u00de, TF = 0, NF = 21, limb topology C||Pa||Pa and Pa||P\\P\\kR, Pa||P\\P\\\\R (a), Pa\\P\\kP\\kR, Pa\\P\\kP\\\\R (b)", + "texts": [ + "79, 5.80, 5.81, 5.82, 5.83, 5.84, 5.85, 5.86, 5.87, 5.88, 5.89, 5.90, 5.91, 5.92 No. PM type Limb topology Connecting conditions 1. 3PaPaPaR1RPPaPat (Fig. 5.69) Pa\\Pa||Pa\\kR (Fig. 5.6a) The revolute joint between links 7D and 11 of G4 limb and the last revolute joints of limbs G1, G2 and G3 have superposed axes Pa\\Pa||Pa||R (Fig. 5.6b) R||P||Pa||Pat (Fig. 5.4n) 2. 3PaPaPaR1RPPaPat (Fig. 5.70) Pa\\Pa||Pa\\\\R (Fig. 5.6c) Idem No. 1 Pa\\Pa||Pa||R (Fig. 5.6b) R||P||Pa||Pat (Fig. 5.54n) 3. 3PaPPR1CPaPa (Fig. 5.71a) Pa||P\\P\\kR (Fig. 5.2a) The revolute joint between links 6D and 8 of G4 limb and the last revolute joints of limbs G1, G2 and G3 have superposed axes Pa||P\\P\\\\R (Fig. 5.2d) C||Pa||Pa (Fig. 5.4o) 4. 3PaPPR1CPaPa (Fig. 5.71b) Pa\\P\\kP\\kR (Fig. 5.2b) Idem No. 3 Pa\\P\\kP\\\\R (Fig. 5.2c) C||Pa||Pa (Fig. 5.40) 5. 3PaPPR1CPaPat (Fig. 5.72a) Pa||P\\P\\kR (Fig. 5.2a) Idem No. 3 Pa||P\\P\\\\R (Fig. 5.2d) C||Pa||Pat (Fig. 5.4p) 6. 3PaPPR1CPaPat (Fig. 5.72b) Pa\\P\\kP\\kR (Fig. 5.2b) Idem No. 3 Pa\\P\\kP\\\\R (Fig. 5.2c) C||Pa||Pat (Fig. 5.4p) (continued) 534 5 Topologies with Uncoupled Sch\u00f6nflies Motions Table 5.5 (continued) No. PM type Limb topology Connecting conditions 7. 2PaPRRR1PaPRR1CPaPa (Fig. 5.73a) Pa\\P\\\\R||R\\\\R (Fig. 5.3a) Idem No" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_35_0001419_978-94-007-7401-8-Figure2.37-1.png", + "original_path": "designv11-35/openalex_figure/designv11_35_0001419_978-94-007-7401-8-Figure2.37-1.png", + "caption": "Fig. 2.37 Fully-parallel PMs with coupled Sch\u00f6nflies motions of types 4PCPa (a) and 4PPaC (b) defined by MF \u00bc SF \u00bc 4; RF\u00f0 \u00de \u00bc \u00f0v1; v2; v3;xd\u00de, TF = 0, NF = 18, limb topology P\\C\\Pa (a) and P||Pa\\C (b)", + "texts": [ + "33a) R\\P||Pa\\kP (Fig. 2.21o) The first revolute joints of the four limbs have parallel axes 16. 4RPPPa (Fig. 2.33b) R||P\\P||Pa (Fig. 2.21p) Idem No. 15 17. 4PaPPR (Fig. 2.34a) Pa||P\\P||R (Fig. 2.21q) Idem No. 4 18. 4PaPPR (Fig. 2.34b) Pa\\P\\kP\\kR (Fig. 2.21r) Idem No. 4 19. 4PaRPP (Fig. 2.35a) Pa\\R||P\\P (Fig. 2.21s) Idem No. 11 20. 4PaPRP (Fig. 2.35b) Pa||P\\R||P (Fig. 2.21t) Idem No. 5 21. 4PaPRP (Fig. 2.36a) Pa\\P||R\\P (Fig. 2.21u) Idem No. 11 22. 4PaRPP (Fig. 2.36b) Pa\\R\\P\\kP (Fig. 2.21v) Idem No. 5 23. 4PCPa (Fig. 2.37a) P\\C\\Pa (Fig. 2.21w) Idem No. 5 24. 4PPaC (Fig. 2.37b) P||Pa\\C (Fig. 2.21x) The last joints of the four limbs have parallel axes 25. 4CPPa (Fig. 2.38a) C\\P||Pa (Fig. 2.21y) The first joints of the four limbs have parallel axes 26. 4PaPC (Fig. 2.38b) Pa||P\\C (Fig. 2.21z) Idem No. 24 27. 4PaCP (Fig. 2.39) Pa\\C\\P (Fig. 2.21z1) Idem No. 11 28. 4PPaPaR (Fig. 2.40) P\\Pa\\\\Pa\\kR (Fig. 2.22a) Idem No. 4 29. 4PPaRPa (Fig. 2.41) P\\Pa||R\\Pa (Fig. 2.22b) The axes of the revolute joints connecting the two parallelogram loops of the four limbs are parallel (continued) 2" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_35_0001419_978-94-007-7401-8-Figure2.18-1.png", + "original_path": "designv11-35/openalex_figure/designv11_35_0001419_978-94-007-7401-8-Figure2.18-1.png", + "caption": "Fig. 2.18 Fully-parallel PMs with coupled Sch\u00f6nflies motions of types 4RRCR (a) and 4RCRR (b) defined by MF \u00bc SF \u00bc 4; RF\u00f0 \u00de \u00bc \u00f0v1; v2; v3;xd\u00de, TF = 0, NF = 2, limb topology R||R||C\\R (a) and R||C||R\\R (b)", + "texts": [ + "1x) Idem No. 21 26. 4RRPRR (Fig. 2.14b) R\\R||P||R||R||R (Fig. 2.1y) Idem No. 21 27. 4RPRRP (Fig. 2.15a) R\\P\\kR\\R\\P (Fig. 2.1z) Idem No. 19 28. 4PRRRP (Fig. 2.15b) P\\R||R\\R\\P (Fig. 2.1z1) Idem No. 19 29. 4PRRPR (Fig. 2.16) P\\R\\R\\P\\kR (Fig. 2.1a0) Idem No. 5 30. 4CRRR (Fig. 2.17a) C\\R||R\\R (Fig. 2.1d0) Idem No. 1 (continued) 2.1 Topologies with Simple Limbs 59 Table 2.1 (continued) No. PM type Limb topology Connecting conditions 31. 4CRRR (Fig. 2.17b) C||R||R\\R (Fig. 2.1e0) Idem No. 13 32. 4RRCR (Fig. 2.18a) R||R||C\\R (Fig. 2.1f0) Idem No. 13 33. 4RCRR (Fig. 2.18b) R||C||R\\R (Fig. 2.1g0) Idem No. 13 34. 4RRRC (Fig. 2.19a) R\\R||R||C (Fig. 2.1h0) Idem No. 21 35. 4RCRR (Fig. 2.19b) R\\C||R||R (Fig. 2.1l0) Idem No. 21 36. 4RRCR (Fig. 2.20a) R\\R||C||R (Fig. 2.1b0) Idem No. 21 37. 4RCRR (Fig. 2.20b) R||C\\R||R (Fig. 2.1c0) Idem No. 21 60 2 Fully-Parallel Topologies with Coupled Sch\u00f6nflies Motions In the fully-parallel topologies of PMs with coupled Sch\u00f6nflies motions F / G1G2-G3-G4 presented in this section, the moving platform n : nGi (i = 1, 2, 3, 4) is connected to the reference platform 1 : 1Gi : 0 by four spatial complex limbs with four or five degrees of connectivity" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_35_0000114_iros45743.2020.9341171-Figure1-1.png", + "original_path": "designv11-35/openalex_figure/designv11_35_0000114_iros45743.2020.9341171-Figure1-1.png", + "caption": "Fig. 1. Conceptual illustration of the curvature-dependent distributed friction/normal force in TD-CMs and its interrelated relation with curvature.", + "texts": [ + " For instance, various types of TD-CMs have been utilized for minimally invasive treatment of orthopedic applications [2], [3], single-incision laparoscopic [4]. This is mainly due to the excellent features of a tendon-driven actuation mechanism that enable safe and remote transmission of power to the robot\u2019s end-effector using a lightweight and miniaturized tendon-sheath mechanism [1], [5]. However, the adverse effects of tendon-sheath friction along the transmission line (typically, the body of CM as shown in Fig. 1) may result in significant non-uniform tendon tension and subsequently tension/motion losses [5], [6]. These considerable losses affect the deformation behavior of a TD-CM, which need to be considered during the modeling and design phases before fabrication of the robot. Additionally, appropriate TDCM deformation behavior modeling considering the tendonsheath friction can mitigate challenges associated with the CM\u2019s accurate control, shape sensing, and contact detection with the environment [1]" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_35_0000395_978-3-030-67750-3_16-Figure30.10-1.png", + "original_path": "designv11-35/openalex_figure/designv11_35_0000395_978-3-030-67750-3_16-Figure30.10-1.png", + "caption": "Fig. 30.10 aMotion study of wheelbarrow loader bucket. b Load applied on the loader bucket and wheelbarrow tray", + "texts": [ + " These simulations are carried out to ensure the structure design is sufficient to withstand the load applied. There are a few analysis obtained by carrying out an analysis and simulation such as stress, strain, displacement, and motor force. Firstly, the motion study was done to the design. It is very crucial to study the movement of each part and to observe where parts may collide each other so that the parts can be modified to move free as intended without any problems when the fabricating process is done. The motion study is shown in Fig. 30.10a and the results show no collision happened and the design is ready for the next level of analysis which is the structural analysis. During the structural analysis, the load of 500 N is applied to the loader bucket as in Fig. 30.10b and the deformation at certain parts is observed and analyzed to find the affected area of the developed model so that the model can be modified to have more strength and durability during the real-time use. The stress\u2013strain analysis results show that the highest value of stress is 6.297e + 07 N/m2.. In contrast, 6.247e-04 for a strain that occurs at the linkage between the loader bucket and body bracket that is attached to the wheelbarrow body tray and this value is still considered acceptable and can be seen in Figs" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_35_0000016_j.ijsolstr.2021.01.024-Figure7-1.png", + "original_path": "designv11-35/openalex_figure/designv11_35_0000016_j.ijsolstr.2021.01.024-Figure7-1.png", + "caption": "Fig. 7. Illustration of the single-node system evolution in the space of external forces Fi , for no tangential coupling and load case I(F2 \u00bc 0). (a) Plane F2; F3 showing the projection of the separation surface with unit normal d \u00bc 0; 0:6402; 0:7682\u00bd T . (b) Tangential plane F3 \u00bc P0 \u00bc 1. The relevant points are related to: P, normal loading only; E, stick limit k \u00bc k0; S, dissipative limit k \u00bc kD; U, shakedown upper bound k \u00bc k2. (For interpretation of the references to colour in this figure legend, the reader is referred to the web version of this article.)", + "texts": [ + " In this simple single-node system, we can derive both the limits k2 and kD analytically and provide an explanation to the fact that the optimal slip array has a non-zero component along the x2 direction, in contrast to the predictions of Gauss\u2013Seidel results (orange circles in Fig. 6a). Let us consider the space defined by the components of the external force vector F1; F2; F3. As previously mentioned, the condition of separation F d > 0 is represented by a plane passing through the origin F \u00bc 0, with unit normal d (Klarbring et al., 2007). When there is no tangential coupling the unit normal d has a null component in the F1 direction, so that if we consider the vertical plane F2; F3 its projection is represented by a line, whose slope is tana \u00bc d2=d3 (Fig. 7a). Incidentally, note that in the case of no normal-tangential coupling the unit normal would be parallel to the axis F3. In the tangential plane F3 \u00bc Po the diameter of Coulomb\u2019s circle is equal to 2fPo, while the tangent lines are inclined of an angle b \u00bc arcsin fd2=d3\u00f0 \u00de (Fig. 7b). Such an angle is defined by the line belonging to the plane F3 \u00bc Po, passing through point C and tangent to the circle centred in O. In order to prevent separation, the vertex of any cone has to lie on the separation plane. Initially, the system is brought to point P upon application of normal loading only (according to the sign convention, F3 < 0). Considering load case I, tangential loading is then applied in direction x1 and the stick limit is encountered on the border of the blue circle centred in O (point E), when k \u00bc k0 \u00bc f=2. Since F2 \u00bc 0 and w2 \u00bc 0 at any stage due to the lack of coupling between the tangential directions, we can only have vertical shifts of the cone circles in Fig. 7b as the load multiplier k is increased. Moreover, for the initial assumption of j31 \u00bc 0 in Eq. (24), the radius cannot increase because there is no coupling between tangential slip w1 and normal force F3. Hence, further loading requires the circle to shift to a new one in the tangential plane (yellow circle centred in E), whose border represents the limit k \u00bc kD \u00bc f (point S). However, the optimisation shows that there exists an upper limit k2 which is larger than the multiplier kD, but it can be reached only under specific conditions, namely, the existence of w2 \u2013 0, as found in Fig", + " It was out of the scope of our work to explore the impact of different initial conditions within the conditional region, whose size appears to increase with the level of friction (Fig. 5). In general cases, the impossibility of employing limit analyses theorems has made iterative approaches, such as that proposed by Flicek et al. (2017), inevitable. In the case of the single-node system, we have illustrated how the size of the conditional region can be approximated through a didactic example (Fig. 7). Taking advantage of the amenable case of two-dimensional loading and no tangential coupling (j21 \u00bc 0), we have performed an analytical computation of the limits enclosing all the dissipative scenarios explored with incremental analyses (Fig. 5c). Specifically, we could show that the system is able to sustain cyclic loads up to the dissipative limit kD \u00bc f , whereas reaching the shakedown upper bound k2 requires different initial conditions with a non-zero slip displacement along the other tangential direction" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_35_0001393_ijhm.2020.109916-Figure2-1.png", + "original_path": "designv11-35/openalex_figure/designv11_35_0001393_ijhm.2020.109916-Figure2-1.png", + "caption": "Figure 2 Representation of spur gear (see online version for colours)", + "texts": [], + "surrounding_texts": [ + "The geometrical model of spur gear was constructed using SOLIDWORKS software for the designing process (Figures 1 and 2). After creating the model in SOLIDWORKS software, the design is imported into ANSYS workbench software for further analysis of the key factors such as stresses, deflections and modal analysis of the spur gear. The dimensions and the material properties of the spur gear are given Tables 1 and 2. Figure 3 shows the meshed view of the spur gear. Accuracy and efficiency of finite element method depends upon the meshing size of the model. A very fine mesh is used to design the model so that the results would be more accurate." + ] + }, + { + "image_filename": "designv11_35_0000686_j.mechmachtheory.2021.104433-Figure17-1.png", + "original_path": "designv11-35/openalex_figure/designv11_35_0000686_j.mechmachtheory.2021.104433-Figure17-1.png", + "caption": "Fig. 17. Metamorphic epicyclic gear-rack train based on geometrical constraints.", + "texts": [], + "surrounding_texts": [ + "By combining geometrical constraint and force constraint, a set of metamorphic epicyclic gear trains can be obtained based on a combination constraint. As shown in Figs. Figures 21-23, the combination constraint can provide both geometrical limitation and constraint forces. The metamorphic epicyclic gear-rack train in Fig. 23 is used to demonstrate the configuration transformation. In the rotation configuration, the driving force drives the planet rack overcoming the constraint force, and the spring in combination constraint is compressed or released. The angular velocity \u03c9r2 is not equal to \u03c9r1 the corresponding motion branch has mobility 2 as discussed in Fig. 14(b). When the driving force is equal to the constraint force or the planet rack is locked by geometrical limitation, the metamorphic epicyclic gear-rack train transforms to revolution configuration. The planet rack is rigidified with the sun rack-gear, and the angular velocity \u03c9r2 is equal to \u03c9r1, which means the corresponding motion branch has degenerated from mobility 2 to mobility 1. In practical design, we can adjust the constraint force to determine the constraint condition and control the configuration transformation. By inserting three types of constraints including geometrical constraint, force constraint, and combination constraint separately on the mechanism, one mobility is restricted that the metamorphic epicyclic gear train and initiates one active mobility. In practical design, we use external forces, such as gravity, magnetic force, constraint force from the target object, to overcome the inner constraint to release the mobility. These proposed metamorphic mechanisms usually operate with one active mobility and change their configuration and state of motion when the intervention of external force. With two mobility states, when the metamorphosis is initiated, either geometrical constraint or force constraint is utilized to restrict one mobility. An example describing the way of controlling the transformation is demonstrated in the following section." + ] + }, + { + "image_filename": "designv11_35_0001419_978-94-007-7401-8-Figure3.46-1.png", + "original_path": "designv11-35/openalex_figure/designv11_35_0001419_978-94-007-7401-8-Figure3.46-1.png", + "caption": "Fig. 3.46 Overactuated PMs with coupled Sch\u00f6nflies motions of types 4RRCR (a) and 4RRRC (b) defined by MF \u00bc SF \u00bc 4; RF\u00f0 \u00de \u00bc v1; v2; v3;xd\u00f0 \u00de, TF = 0, NF = 2, limb topology R||R\\C||R (a) and R||R\\R||C (b)", + "texts": [ + "3d0) The third joints of the four limbs have parallel axes 45. 4PRRPR (Fig. 3.42b) P||R||R\\P\\\\R (Fig. 3.3e0) Idem No. 17 46. 4RPRPR (Fig. 3.43a) R||P||R\\P\\\\R (Fig. 3.3f0) Idem No. 17 47. 4RRPPR (Fig. 3.43b) R||R||P\\P\\\\R (Fig. 3.3g0) Idem No. 17 48. 4RPRPR (Fig. 3.44a) R||P||R\\P\\\\R (Fig. 3.3h0) Idem No. 17 49. 4RPRPR (Fig. 3.44b) P\\P\\kR||R\\\\R (Fig. 3.3i0) Idem No. 17 50. 4PPRRR (Fig. 3.44c) P\\P||R||R\\\\R (Fig. 3.3j0) Idem No. 17 51. 4CRRR (Fig. 3.45a) C||R\\R||R (Fig. 3.3k0) Idem No. 17 52. 4RCRR (Fig. 3.45b) R||C\\R||R (Fig. 3.3l0) Idem No. 17 53. 4RRCR (Fig. 3.46a) R||R\\C||R (Fig. 3.3m0) Idem No. 16 54. 4RRRC (Fig. 3.46b) R||R\\R||C (Fig. 3.3n0) Idem No. 16 55. 4RCRP (Fig. 3.47a) R\\C||R\\\\P (Fig. 3.3o0) Idem No. 16 56. 4RRCP (Fig. 3.47b) R\\R||C\\\\P (Fig. 3.3p0) Idem No. 16 57. 4CRRP (Fig. 3.48a) C||R\\R\\P (Fig. 3.3q0) Idem No. 44 58. 4RCRP (Fig. 3.48b) R||C\\R\\P (Fig. 3.3r0) Idem No. 44 59. 4CRPR (Fig. 3.49a) C||R\\P\\\\R (Fig. 3.3s0) Idem No. 17 60. 4PCRR (Fig. 3.49b) P\\C||R\\\\R (Fig. 3.3t0) Idem No. 17 3.1 Topologies with Simple Limbs 245 Table 3.3 Structural parametersa of parallel mechanisms in Figs. 3.4, 3.5, 3.6, 3" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_35_0001419_978-94-007-7401-8-Figure2.2-1.png", + "original_path": "designv11-35/openalex_figure/designv11_35_0001419_978-94-007-7401-8-Figure2.2-1.png", + "caption": "Fig. 2.2 Fully-parallel PMs with coupled Sch\u00f6nflies motions of types 4RRRRR (a) and 4RRRRR (b) defined by MF \u00bc SF \u00bc 4; RF\u00f0 \u00de \u00bc \u00f0v1; v2; v3;xd\u00de, TF = 0, NF = 2, limb topology R\\R||R||R\\k R (a) and R||R\\R||R\\R (b)", + "texts": [ + " 54 2 Fully-Parallel Topologies with Coupled Sch\u00f6nflies Motions 2.1 Topologies with Simple Limbs 55 56 2 Fully-Parallel Topologies with Coupled Sch\u00f6nflies Motions 2.1 Topologies with Simple Limbs 57 58 2 Fully-Parallel Topologies with Coupled Sch\u00f6nflies Motions Table 2.1 Limb topology and connecting conditions of the fully-parallel solutions with no idle mobilities presented in Figs. 2.2, 2.3, 2.4, 2.5, 2.6, 2.7, 2.8, 2.9, 2.10, 2.11, 2.12, 2.13, 2.14, 2.15, 2.16, 2.17, 2.18, 2.19, 2.20 No. PM type Limb topology Connecting conditions 1. 4RRRRR (Fig. 2.2a) R\\R||R||R\\k R (Fig. 2.1a) The first and the last revolute joints of the four limbs have parallel axes 2. 4RRRRR (Fig. 2.2b) R||R\\R||R\\R (Fig. 2.1b) The first, the second and the last revolute joints of the four limbs have parallel axes 3. 4RRRRR (Fig. 2.3a) R||R||R\\R||R (Fig. 2.1c) The two last revolute joints of the four limbs have parallel axes 4. 4RRRRR (Fig. 2.3b) R||R||R\\R||R (Fig. 2.1c) The three first revolute joints of the four limbs have parallel axes 5. 4RRRRR (Fig. 2.4a) R||R\\R||R||R (Fig. 2.1d) The two first revolute joints of the four limbs have parallel axes. 6. 4PRRRR (Fig. 2.4b) P\\R\\R||R||R (Fig. 2.1e) The second joints of the four limbs have parallel axes 7" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_35_0000686_j.mechmachtheory.2021.104433-Figure3-1.png", + "original_path": "designv11-35/openalex_figure/designv11_35_0000686_j.mechmachtheory.2021.104433-Figure3-1.png", + "caption": "Fig. 3. Epicyclic external gear train.", + "texts": [ + " A link perpendicular to the instantaneous screw axis and collinear with the mesh line is used to replace the higher pair by calculating the position of the instantaneous screw axis [48] to obtain the mesh points [49]. In this way, equivalent mechanisms are established in this section. Following the principle of gear meshing, the instantaneous screw axis is the axis of the velocity screw of relative rotation between the drive gear and driven gear, and its position vector points to the mesh point. Fig. 3 shows the coordinate systems for the epicyclic external gear train. Note that the z-axis of all coordinate systems in Fig. 3(a) are omitted because they are all perpendicular to the moving plane. The base coordinate system Oe0 \u2212 xe0ye0ze0 is fixed to the ground with its ze0-axis along the rotating axis of the sun gear. The sun gear coordinate system Oe1 \u2212 xe1ye1ze1 is rigidly connected to the sun gear. The rotation angle of the sun gear \u03c6e1 is the angle between the sun gear coordinate system Oe1 and the base coordinate system Oe0. The gear arm coordinate system Oe2 \u2212 xe2ye2ze2 is rigidly attached to the gear arm. The xe2-axis is along the gear arm" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_35_0000576_j.commatsci.2021.110674-Figure1-1.png", + "original_path": "designv11-35/openalex_figure/designv11_35_0000576_j.commatsci.2021.110674-Figure1-1.png", + "caption": "Fig. 1. (a) The physical model established for the powder paving step of TiNbTa powder mixture alloys during LPBF through discrete element modelling, and (b) the three-dimensional model of LPBF powder-bed established for laser fusing the mixture of discrete Ti, Nb and Ta powder employing the \u201cS\u201d scanning strategy.", + "texts": [ + " Compared to the conventional counter-rolling method, the scraping method is well accepted as an efficient type of powder layering and commonly applied in industry as it needs a relatively simple geometry structure, and hence it is selected as the powder layering method in this study. According to the inherent feature of LPBF process, the numerical model for powder paving was established through DEM modelling, containing the paving blade, delivery chamber filled with composite powder packing, powder-bed, etc, as shown in Fig. 1a. The composite powder system consisted of Ti, Ta and Nb particles, and the number of CP-Ti, Ta and Nb particles was generated on the basis of the weight percent of 25 wt% Nb, 10 wt% Ta and balance of Ti, respectively. With M. Xia et al. Computational Materials Science 198 (2021) 110674 respect to powder layering of LPBF process, the powder particles are required to be nearly sphere-shaped as possible to ensure a desired flowability and a high-quality powder-bed. In this case, the CP-Ti powder particles are assumed to be near-spherical, and meanwhile, the used Ta and Nb powder particles are also simplified to be spherical with a mean size of 10 \u03bcm", + " 30 \u03bcm) to ensure the sufficient supply for the powder layering. Considering the complex interaction among the multi-element powder particles, the blade paved powder particles towards the powder-bed with a constant velocity of 40 mm/s in order to maintain a stable powder paving process. Meanwhile, a newly developed mesoscopic powder-bed model involving Ti, Nb and Ta powder particles was also established through FVM modelling according to the packing state of composite powder particles within the powderbed, as shown in Fig. 1b. Taking into consideration of the calculation ability of computer workstation, the physical model for LPBF process was established with a three dimension size of 300 \u00d7 250 \u00d7 50 mm3, considering the transition of solid/ liquid phases, the surface tension and thermo-capillary force of liquid, the recoil pressure and buoyancy force. To further ensure the complicated algorithm mathematically convergent, the thermal physical properties of powder particles, such as viscosity, thermal conductivity and surface tension, were considered as temperature-dependent and given in Table 1, but some other thermal physical parameters were assumed to be temperature independent" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_35_0001419_978-94-007-7401-8-Figure6.17-1.png", + "original_path": "designv11-35/openalex_figure/designv11_35_0001419_978-94-007-7401-8-Figure6.17-1.png", + "caption": "Fig. 6.17 Maximaly regular fully-parallel PMs with Sch\u00f6nflies motions of types 3PPPR1CPaPa (a) and 3PPPR-1CPaPat (b) defined by MF = SF = 4, (RF) = (v1; v2; v3;xb), TF = 0, NF = 12, limb topology P\\P ??P\\||R, P\\P ??P ??R and C||Pa||Pa (a), C||Pa||Pat (b)", + "texts": [ + " 5.4l) Idem no. 1 19. 2CRRR-1CRR-1RPaPaP (Fig. 6.16a) C||R||R ?R (Fig. 4.2g) C||R||R (Fig. 4.1k) R||Pa||Pa||P (Fig. 5.4k) Idem no. 1 20. 2CRRR-1CRR1RPaPatP (Fig. 6.16b) C||R||R ?R (Fig. 4.2g) C||R||R (Fig. 4.1k) R||Pa||Pat||P (Fig. 5.4l) Idem no. 1 Table 6.6 Limb topology and connecting conditions of the fully-parallel solutions with no idle mobilities presented in Figs. 6.17, 6.18, 6.19, 6.20, 6.21, 6.22, 6.23, 6.24, 6.25, 6.26 No. PM type Limb topology Connecting conditions 1. 3PPPR-1CPaPaP (Fig. 6.17a) P ?P\\\\P ?||R (Fig. 4.1a) P ?P\\\\P\\\\R (Fig. 4.1b) C||Pa||Pa (Fig. 5.4o) The last revolute joints of limbs G1, G2 and G3 and the revolute joint between links 6D and 8 have superposed axes. The actuated prismatic joints of limbs G1, G2 and G3 have orthogonal directions 2. 3PPPR-1CPaPatP (Fig. 6.17b) P ?P\\\\P ?||R (Fig. 4.1a) P ?P\\\\P\\\\R (Fig. 4.1b) CPa||Pat (Fig. 5.4p) Idem no. 1 3. 2PPRRR-1PPRR-1CPaPa (Fig. 6.18a) P ?P ?||R||R ?R (Fig. 4.2a) P ?P ?||R||R (Fig. 4.1d) C||Pa||Pa (Fig. 5.4o) Idem no. 1 (continued) 6.2 Fully-Parallel Topologies with Simple and Complex Limbs 613 Table 6.6 (continued) No. PM type Limb topology Connecting conditions 4. 2PPRRR-1PPRR-1CPaPat (Fig. 6.18b) P ?P ?||R||R ?R (Fig. 4.2a) P ?P ?||R||R (Fig. 4.1d) C||Pa||Pat (Fig. 5.4p) Idem no. 1 5. 2CPRR-1CPR-1CPaPa (Fig. 6", + " Structural parameter Solution Figures 6.7 and 6.11 Figures 6.8, 6.9, 6.10, 6.12, 6.13, 6.14 20. MG2 4 4 21. MG4 4 4 22. (RF) (v1; v2; v3;xb) (v1; v2; v3;xb) 23. SF 4 4 24. rl 6 6 25. rF 18 20 26. MF 4 4 27. NF 12 10 28. TF 0 0 29. Pp1 j\u00bc1 fj 4 5 30. Pp2 j\u00bc1 fj 4 4 31. Pp3 j\u00bc1 fj 4 5 32. Pp4 j\u00bc1 fj 10 10 33 Pp j\u00bc1 fj 22 24 a See footnote of Table 2.2 for the nomenclature of structural parameters 616 6 Maximally Regular Topologies with Sch\u00f6nflies Motions Table 6.8 (continued) No. Structural parameter Solution Figures 6.15 and 6.16 Figure 6.17 21. MG4 4 4 22. (RF) (v1; v2; v3;xb) (v1; v2; v3;xb) 23. SF 4 4 24. rl 6 6 25. rF 20 18 26. MF 4 4 27. NF 10 12 28. TF 0 0 29. Pp1 j\u00bc1 fj 5 4 30. Pp2 j\u00bc1 fj 4 4 31. Pp3 j\u00bc1 fj 5 4 32. Pp4 j\u00bc1 fj 10 10 33. Pp j\u00bc1 fj 24 22 a See footnote of Table 2.2 for the nomenclature of structural parameters 6.2 Fully-Parallel Topologies with Simple and Complex Limbs 617 Table 6.9 (continued) No. Structural parameter Solution Figure 6.18 Figures 6.19 and 6.20 23. SF 4 4 24. rl 6 6 25. rF 20 20 26. MF 4 4 27. NF 10 10 28" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_35_0000381_tro.2021.3070102-Figure11-1.png", + "original_path": "designv11-35/openalex_figure/designv11_35_0000381_tro.2021.3070102-Figure11-1.png", + "caption": "Fig. 11. Experiment setup for grasp success prediction. Left: Deformable gripper jaws (blue) grasp a 3-D-printed object with nonplanar surfaces. The 3-D-printed assembly (pink) attached to the grasped object generates external disturbances. Right: Ten 3-D-printed rigid objects that create two types of contact surfaces.", + "texts": [ + " The low standard deviations of both models suggest that the proposed 6DLS models achieve consistent performance and are suitable for a large variety of contact profiles. We apply the 6DLS models to predict physical grasp success for a vertical lifting task. Given the external wrench disturbance wext, the friction coefficient\u03bc, the gripper pose, and an estimated contact profile for each gripper jaw, the algorithm predicts if the grasp can counterbalance wext by checking if the opposite of wext is in the GWS, as summarized in Algorithm 2. Fig. 11 (left) depicts the setup of a KUKA robot arm and a SCHUNK parallel-jaw gripper mounted with customized deformable fin-ray jaws [45] (blue). We 3-D printed rigid grasped objects to control the contact surface. Estimation of the contact profiles is described in Section VIII-B. We further attached a 3-D-printed mechanical assembly (pink) to the grasped object to create different wrench disturbanceswext by mounting weight plates at various locations. Such Authorized licensed use limited to: Carleton University", + " We define an object frame, as the GWS andwext are computed with respect to the origin of the object frame. We select the COM of the grasped object as the origin, instead of the COM of both the grasped object and the wrench disturbance assembly. This frame selection enables easier comparison between the GWSs constructed with different contact models and does not affect the predictions as one can select an arbitrary reference point to compute torques, and therefore, the GWS. We used two approach directions to create wext. Fig. 11 (left) illustrates the x-, y-, z-axes of the object frame for a representative vertical and horizontal grasp with the x-axis parallel to the grasp axis. The wrench disturbance assembly generates wext in the (fy, \u03c4x, \u03c4z)- and (fz, \u03c4x, \u03c4y)-space with the vertical and horizontal grasp direction, respectively. We selected the locations of weight plates so that the disturbances are well-scattered in each space. A force sensor is mounted on each gripper jaw to measure the grasp force along thex-axis. We also mounted an Intel RealSense SR300 RGBD camera (green) on the gripper to label the grasp success by tracking the object pose with the PCL library [46]", + " If the thresholds are high, the grasps will be labeled as a success even if there is a relative motion; whereas with low thresholds, the grasps will be labeled as a failure even if there is no relative motion but the gripper jaws deformed during the manipulation. We also discuss the prediction results with different threshold pairs in Section VIII-E, as robot applications have different tolerance of object motion during the manipulation. While assembly tasks require minimal object motion, bin-picking allows larger object pose change. In future work, we plan to use a tactile sensor to label grasp success by detecting slips. We estimate the contact profile, including the contact surface S and the pressure distribution p, for each gripper jaw. Fig. 11 (right) shows two types of contact surfaces created by ten 3-D-printed rigid objects. As illustrated on the top, the five object models of type I are cut from elliptic cylinders, whose horizontal radii are identical, whereas the vertical radii vary to change the surface curvature. The cylinders are cut so that the contact surface is the same when the grasp force of each jaw is higher than a threshold (20 N). The contact surface is completely defined by the radii and the contact length l1, which is depicted in Fig. 11 (right). If we directly use elliptic cylinders as the object model, we need to measure the contact surface for each trial as the surface increases with the grasp force. As shown on the bottom, each of the type II objects creates five or eight narrow planar contact surfaces with 3\u20135 mm width. We define the contact length l2 of type II as the length of each narrow surface. The direction of frictional forces are constrained to lie in each planar surface, as described in Appendix B. Type II objects show that the contact surface can be nonplanar, even if the local contact surfaces are planar", + " We observed that the pressure values are close to the curve, which suggests that the power-law model is an applicable approximation for the nonplanar contact surfaces used in these experiments. However, we also observed that the exact k value varies from 2.4 to 5.5 for the elliptic cylinders under different loads. Therefore, we discuss the grasp success prediction results with different k values in Section VIII-E. We scaled p\u0302 so that the normal force of each contact matches the force sensor reading Fs. As shown in Fig. 11 (right), Fs measures \u2016f\u22a5x \u2016, the magnitude of the x component of the normal force; therefore, p\u0302 is scaled so that \u2016f\u22a5x \u2016 = Fs. Thus, we computed \u2016f\u0302\u22a5x \u2016 with p\u0302 using (7) and obtained the pressure distribution p(r) = \u03bbp \u00b7 p\u0302(r)with \u03bbp = Fs/\u2016f\u0302\u22a5x \u2016. We precomputed the limit surface models with the normalized power-law pressure distribution p\u0302 for each contact surface. For each grasp, we scaled the contact wrench constraints for each contact with \u03bbp, instead of with the sum magnitude F of the normal forces in (17) to match the force sensor readings. We consider the following baseline contact models to predict the grasp success. 1) 3DLS-planar: The traditional planar surface contact models. As shown in Fig. 11 (right), a planar contact surface (orange line) is created by projecting the nonplanar surface (blue line) along the x-axis onto the yz-plane. We computed the frictional wrenches in the (fy, fz, \u03c4x)-space and fit the 3DLS models to the wrenches. 2) 6DFW: We computed 6DFWs for a nonplanar surface contact and used the 6DFWs of each jaw to construct the GWS without a LS model. 3) 3DLS-nonplanar: The 3DLS models are fit to the three major components, fy, fz, \u03c4x, of the 6DFWs, while the remaining three components are set to zero [16]", + " Downloaded on June 05,2021 at 10:41:34 UTC from IEEE Xplore. Restrictions apply. sensitive to\u03bc as a LS linearly scales with\u03bc. For scenarios with an unknown friction coefficient, one can select a lower \u03bc value for conservative predictions as each model predicts fewer positives. 4) Effect of Contact Lengths: Fig. 14(d) illustrates the results with different contact length offsets. The symbol\u00b10% indicates that the models used the measured contact length l1 and l2 for the two contact types illustrated in Fig. 11 (right). The recall of each baseline model increases with the contact length as the surface area and the frictional torque of each contact also increases. The power-law pressure distribution described in Section VIII-B is based on the assumption that the pressure is symmetric about the object center. However, we observed in our FEM simulations and the results shown in [49] that the pressure distribution can be asymmetric depending on the object pose relative to the fin-ray jaw. With the vertical grasp direction shown in Fig. 11 (left), an asymmetric pressure distribution leads to a component of the normal force that is parallel to the gravity direction, and therefore, affects the prediction results. Furthermore, contact profiles can change during the manipulation due to the jaws\u2019 deformation. One way to address the two limitations is to relax the assumption of a constant symmetric contact profile and to constantly predict grasp success with updated profiles captured with deformable tactile sensors such as GelSlim [33] or the tactile fingertip sensors by Romero et al" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_35_0000467_j.robot.2021.103815-Figure2-1.png", + "original_path": "designv11-35/openalex_figure/designv11_35_0000467_j.robot.2021.103815-Figure2-1.png", + "caption": "Fig. 2. Leg Schematic; Links are referred with upper case letters and joints with lower case letters. In accordance with [27], the leg itself weights a total of 1.5 kg and the rotary compliance weight is considered 0.3 kg.", + "texts": [ + " (7), uk can be computed as: k (\u03c6) = ur (\u03c6, \u03c91) \u2212 ur (\u03c6, \u03c92) \u03c92 1 \u2212 \u03c92 2 (8) Also, substituting Eq. (8) in either Eq. (6) or Eq. (7) gives us Ures. ow in order to design the VPEA, one should use Uk = \u03c92uk (\u03c6) o extract the cam profile and input it into Eq. (4) in place of ur or adaptation. And as previously mentioned, Ures is assumed to e compensated beforehand. This design method will be named he \u2018\u2018frequency-based method\u2019\u2019 hereinafter. . Simulations For the case study, we have selected the robotic leg introduced n [31] to demonstrate the applicability of the proposed VPEA. s depicted in Fig. 2, this 1-DoF leg is comprised of a four-bar inkage mechanism actuated by a motor at joint \u2018h\u2019. The morhology of this robotic leg is designed in a way that it surmounts hose of conventional joint-actuated leg mechanisms in terms of sing fewer actuators (i.e., light-weight structure), uni-directional ctuation (i.e., more energy efficient at high speeds), and simler control strategy. In our previous work [27], we investigated arious types of parallel compliances and experimentally showed hat using a proper fixed rotary compliance at joint \u2018h\u2019 could otably improve energy efficiency of this robotic leg" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_35_0001706_pime_proc_1950_163_014_02-Figure9-1.png", + "original_path": "designv11-35/openalex_figure/designv11_35_0001706_pime_proc_1950_163_014_02-Figure9-1.png", + "caption": "Fig. 9. Test Apparatus", + "texts": [ + ", said that the tests he had carried out in the years 1934-5 in co-operation with sections of the industry concerned at the Research Laboratory for Machine Elements of the Technical University of Berlin were still apparently of some interest. Many different types of seals and packing materials had been tested, some of which had become obsolete j some, however (such as leather and rubber U-section and chevron type seals), were still in use at the present time. The testing equipment had had to be specially developed and constructed for the testing of seals and packings under working conditions, as no apparatus for the purpose was known at the time. Fig. 9 showed the testing apparatus, consisting of a fabricated frame with prismatic slides in which a hydraulic cylinder could be moved up and down by a crank drive, the speed of which was controllable within wide limits. The side loads were taken up by the slides, while any effect on the seals due to the weight of the piston was excluded by the vertical arrangement. When the cylinder was moved, the volume of liquid contained in it remained constant, and as the piston passed straight through the cylinder no resultant force was exerted on it by the hydraulic pressure" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_35_0001419_978-94-007-7401-8-Figure5.74-1.png", + "original_path": "designv11-35/openalex_figure/designv11_35_0001419_978-94-007-7401-8-Figure5.74-1.png", + "caption": "Fig. 5.74 2PaPRRR-1PaPRR-1CPaPat-type fully-parallel PMs with uncoupled Sch\u00f6nflies motions defined by MF = SF = 4, RF\u00f0 \u00de \u00bc \u00f0v1; v2; v3;xb\u00de, TF = 0, NF = 19, limb topology C||Pa||Pat and Pa\\P\\\\R||R\\\\R, Pa\\P\\\\R||R (a), Pa||P\\R||R\\\\R, Pa||P\\R||R (b)", + "texts": [ + " 3PaPPR1CPaPat (Fig. 5.72b) Pa\\P\\kP\\kR (Fig. 5.2b) Idem No. 3 Pa\\P\\kP\\\\R (Fig. 5.2c) C||Pa||Pat (Fig. 5.4p) (continued) 534 5 Topologies with Uncoupled Sch\u00f6nflies Motions Table 5.5 (continued) No. PM type Limb topology Connecting conditions 7. 2PaPRRR1PaPRR1CPaPa (Fig. 5.73a) Pa\\P\\\\R||R\\\\R (Fig. 5.3a) Idem No. 3 Pa\\P\\\\R||R (Fig. 5.2e) C||Pa||Pa (Fig. 5.40) 8. 2PaPRRR1PaPRR1CPaPa (Fig. 5.73b) Pa||P\\R||R\\\\R (Fig. 5.3b) Idem No. 3 Pa||P\\R||R (Fig. 5.2f) C||Pa||Pa (Fig. 5.40) 9. 2PaPRRR1PaPRR1CPaPat (Fig. 5.74a) Pa\\P\\\\R||R\\\\R (Fig. 5.3a) Idem No. 3 Pa\\P\\\\R||R (Fig. 5.2e) C||Pa||Pat(Fig. 5.4p) 10. 2PaPRRR1PaPRR1CPaPat (Fig. 5.74b) Pa||P\\R||R\\\\R (Fig. 5.3b) Idem No. 3 Pa||P\\R||R (Fig. 5.2f) C||Pa||Pat (Fig. 5.4p) 11. 2PaRPRR1PaRPR1CPaPa (Fig. 5.75) Pa\\R\\P\\kR\\R (Fig. 5.3c) Idem No. 3 Pa\\R\\P\\kR (Fig. 5.2g) C||Pa||Pa (Fig. 5.40) 12. 2PaRPRR1PaRPR1CPaPat (Fig. 5.76) Pa\\R\\P\\kR\\R (Fig. 5.3c) Idem No. 3 Pa\\R\\P\\kR (Fig. 5.2g) C||Pa||Pat (Fig. 5.4p) 13. 2PaRRRR1PaRRR1CPaPa (Fig. 5.77) Pa\\R||R||R\\kR (Fig. 5.3d) Idem No. 3 Pa\\R||R||R (Fig. 5.2h) C||Pa||Pa (Fig. 5.40) 14. 2PaRRRR1PaRRR1CPaPat (Fig. 5.78) Pa\\R||R||R\\kR (Fig. 5.3d) Idem No" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_35_0000008_j.autcon.2021.103573-Figure18-1.png", + "original_path": "designv11-35/openalex_figure/designv11_35_0000008_j.autcon.2021.103573-Figure18-1.png", + "caption": "Fig. 18. Configuration of the gantry crane and the block.", + "texts": [ + " 17 shows the configuration of the gantry crane and the block and the desired trajectory that the block was required to follow. The crane H.-W. Lee et al. consisted of a controllable crane girder, two trolleys, and three wire ropes. Transportation and turnover of the block were conducted simultaneously using two trolleys and three wire ropes connected to the hooks. Three hooks and equalizers were connected to the trolley via wire ropes. The upper trolley connected two hooks that held one side of the block, and the lower trolley held the other side with one hook. The properties of the block are shown in Fig. 18. The weight of the block was assumed 300 tons, and the lifting capacity of the gantry crane was 900 tons. The initial and final position of the block is illustrated in Fig. 17. The initial position of the block was (0,0,20), and the final position was (25,10,10), rotated by 90\u25e6. The block was required to be simultaneously transported by 25 m in the x-direction, 10 m in the y-direction, and rotated by 90\u25e6. In feedforward control, the inverse dynamics solver calculates the control inputs required for the target to follow the desired trajectory, provided there are no external disturbances" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_35_0001419_978-94-007-7401-8-Figure2.33-1.png", + "original_path": "designv11-35/openalex_figure/designv11_35_0001419_978-94-007-7401-8-Figure2.33-1.png", + "caption": "Fig. 2.33 Fully-parallel PMs with coupled Sch\u00f6nflies motions of types 4RPPaP (a) and 4RPPPa (b) defined by MF \u00bc SF \u00bc 4; RF\u00f0 \u00de \u00bc \u00f0v1; v2; v3;xd\u00de, TF = 0, NF = 18, limb topology R\\P||Pa\\kP (a) and R||P\\P||Pa (b)", + "texts": [ + "29a) P\\R||P\\Pa (Fig. 2.21g) Idem No. 5 8. 4PRPaP (Fig. 2.29b) P\\R\\Pa \\kP (Fig. 2.21h) Idem No. 5 9. 4PPPaR (Fig. 2.30a) P\\P\\kPa\\kR (Fig. 2.21i) Idem No. 4 10. 4PPPaR (Fig. 2.30b) P\\P||Pa\\R (Fig. 2.21j) Idem No. 4 11. 4PPaRP (Fig. 2.31a) P||Pa\\R||P (Fig. 2.21k) The before last joints of the four limbs have parallel axes 12. 4PPaPR (Fig. 2.31b) P||Pa\\P|| R (Fig. 2.21l) Idem No. 4 13. 4PPaPR (Fig. 2.32a) P\\Pa||P\\\\R (Fig. 2.21m) Idem No. 4 14. 4PPaRP (Fig. 2.32b) P\\Pa\\kR\\kP (Fig. 2.21n) Idem No. 4 15. 4RPPaP (Fig. 2.33a) R\\P||Pa\\kP (Fig. 2.21o) The first revolute joints of the four limbs have parallel axes 16. 4RPPPa (Fig. 2.33b) R||P\\P||Pa (Fig. 2.21p) Idem No. 15 17. 4PaPPR (Fig. 2.34a) Pa||P\\P||R (Fig. 2.21q) Idem No. 4 18. 4PaPPR (Fig. 2.34b) Pa\\P\\kP\\kR (Fig. 2.21r) Idem No. 4 19. 4PaRPP (Fig. 2.35a) Pa\\R||P\\P (Fig. 2.21s) Idem No. 11 20. 4PaPRP (Fig. 2.35b) Pa||P\\R||P (Fig. 2.21t) Idem No. 5 21. 4PaPRP (Fig. 2.36a) Pa\\P||R\\P (Fig. 2.21u) Idem No. 11 22. 4PaRPP (Fig. 2.36b) Pa\\R\\P\\kP (Fig. 2.21v) Idem No. 5 23. 4PCPa (Fig. 2.37a) P\\C\\Pa (Fig. 2.21w) Idem No. 5 24. 4PPaC (Fig. 2.37b) P||Pa\\C (Fig. 2.21x) The last joints of the four limbs have parallel axes 25" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_35_0001419_978-94-007-7401-8-Figure5.77-1.png", + "original_path": "designv11-35/openalex_figure/designv11_35_0001419_978-94-007-7401-8-Figure5.77-1.png", + "caption": "Fig. 5.77 2PaRRRR-1PaRRR-1CPaPa-type fully-parallel PMs with uncoupled Sch\u00f6nflies motions defined by MF = SF = 4, RF\u00f0 \u00de \u00bc \u00f0v1; v2; v3;xb\u00de, TF = 0, NF = 19, limb topology Pa\\R||R||R\\kR, Pa\\R||R||R and C||Pa||Pa", + "texts": [ + "40) 9. 2PaPRRR1PaPRR1CPaPat (Fig. 5.74a) Pa\\P\\\\R||R\\\\R (Fig. 5.3a) Idem No. 3 Pa\\P\\\\R||R (Fig. 5.2e) C||Pa||Pat(Fig. 5.4p) 10. 2PaPRRR1PaPRR1CPaPat (Fig. 5.74b) Pa||P\\R||R\\\\R (Fig. 5.3b) Idem No. 3 Pa||P\\R||R (Fig. 5.2f) C||Pa||Pat (Fig. 5.4p) 11. 2PaRPRR1PaRPR1CPaPa (Fig. 5.75) Pa\\R\\P\\kR\\R (Fig. 5.3c) Idem No. 3 Pa\\R\\P\\kR (Fig. 5.2g) C||Pa||Pa (Fig. 5.40) 12. 2PaRPRR1PaRPR1CPaPat (Fig. 5.76) Pa\\R\\P\\kR\\R (Fig. 5.3c) Idem No. 3 Pa\\R\\P\\kR (Fig. 5.2g) C||Pa||Pat (Fig. 5.4p) 13. 2PaRRRR1PaRRR1CPaPa (Fig. 5.77) Pa\\R||R||R\\kR (Fig. 5.3d) Idem No. 3 Pa\\R||R||R (Fig. 5.2h) C||Pa||Pa (Fig. 5.40) 14. 2PaRRRR1PaRRR1CPaPat (Fig. 5.78) Pa\\R||R||R\\kR (Fig. 5.3d) Idem No. 3 Pa\\R||R||R (Fig. 5.2h) C||Pa||Pat (Fig. 5.4p) 15. 3PaPPaR1CPaPa (Fig. 5.79a) Pa||P\\Pa\\kR (Fig. 5.4a) The revolute joint between links 6D and 9 of G4 limb and the last revolute joints of limbs G1, G2 and G3 have superposed axes Pa||P\\Pa||R (Fig. 5.4d) C||Pa||Pa (Fig. 5.40) 16. 3PaPPaR1CPaPa (Fig. 5.79b) Pa\\P\\\\Pa\\kR (Fig. 5.4b) Idem No. 15 Pa\\P\\\\Pa||R (Fig" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_35_0000100_iros45743.2020.9341746-Figure1-1.png", + "original_path": "designv11-35/openalex_figure/designv11_35_0000100_iros45743.2020.9341746-Figure1-1.png", + "caption": "Fig. 1. a) The structure of the presented deployable mechanism for tooltip camera, b) 2-DOF movement of the deployable arm and the direction of camera.", + "texts": [ + " DESIGN OF THE INSTRUMENT Three main considerations in the design of an endoscopic wrist mechanism: 1) The deployable mechanism can pass through the trocar, 2) the mechanism needs 2 degrees of freedom (extension and rotation), and 3) the mechanism is 2.5 mm or less in diameter, both in terms of manufacturing and assembling. A. 2-DOF Deployable Arm Mechanism To simplify the design, the 2-DOF movement can be achieved by using an S-curved nitinol arm. This allows the arm to adjust the FOV, leading to a simple and scalable structure. The arm has translational and rotational movement along the x-axis like Fig. 1. The camera is located at the distal tip of the arm. The arm\u2019s rotational movement rotates the direction of the camera about the x-axis and the translational movement changes the x-coordinate of the camera. Once the arm advances further and has reached the edge of the main shaft\u2019s hole pattern, the wrist naturally bends upwards from the main tool shaft. The translational advancement along the +x-axis allows the arm to function as a bending wrist. The degree of translational advancement determines the bending degree of the continuum arm according to the arm\u2019s bending stiffness, \ud835\udc38\ud835\udc3c\ud835\udc4e\ud835\udc5f\ud835\udc5a , where \ud835\udc38 and \ud835\udc3carm are elastic modulus and the moment of inertia of the deployable arm, respectively" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_35_0000686_j.mechmachtheory.2021.104433-Figure6-1.png", + "original_path": "designv11-35/openalex_figure/designv11_35_0000686_j.mechmachtheory.2021.104433-Figure6-1.png", + "caption": "Fig. 6. Equivalent mechanism of the epicyclic bevel gear train.", + "texts": [ + " \u23a7 \u23aa \u23a8 \u23aa \u23a9 rib13 = [ 0 0 0 ]T sib13 = [ \u03c9b3cos\u03c6b2 \u03c9b3sin\u03c6b2 rb Rb \u03c9b3 ]T (12) Following the direction vector and position vector, the instantaneous screw axis is coplanar and intersects with the sun bevel gear axis and planet bevel gear axis; moreover, it points to the mesh point of the epicyclic bevel gear train. In addition, it rotates about the axis of the sun bevel gear with a bevel gear arm. As same as the epicyclic external gear train, a closed-loop spatial 5-bar linkage of mobility 2 is constructed as the equivalent mechanism of the epicyclic bevel gear train in Fig. 6. The links lbBC, lbDE, and lbAE are equivalent to the sun bevel gear, the planet bevel gear, and the bevel gear arm, respectively. The Hooke joint D and the ball joint C are H. Yuan et al. Mechanism and Machine Theory 166 (2021) 104433 the centers of curvature of the tooth profile curve at the mesh point. The higher pair at the mesh point is replaced by link lbCD, which is perpendicular to links lbBC and lbDE,. Instantaneously, the equivalent mechanism has the same mobility, instantaneous velocity, and instantaneous acceleration as the epicyclic bevel gear train. Its equivalent geometrical relationships are obtained in Eq. (13). As shown in Fig. 6, the angle \u03b1b between link lbCD and the yb0-axis is the pressure angle, \u03b4b is the pitch angle, \u03b1b1 and \u03b1b2 are the angles between link lbCD and the xbOybO plane and ybOzbO plane, respectively, and \u03b1b3 is the angle between link lbBC and the zbO-axis. This gives the following geometrical relationship between the length of the links lbDE, lbAC, lbCD in the equivalent mechanism and the radiuses Rb, rb of the pitch circles in the epicyclic bevel gear train. \u23a7 \u23aa \u23aa \u23aa \u23aa \u23a8 \u23aa \u23aa \u23aa \u23aa \u23a9 tan\u03b4b = rb/Rb \u03b1b1 = arcsin(sin\u03b1bcos\u03b4b) \u03b1b2 = arctan(tan\u03b1bsin\u03b4b \u03b1b3 = arcsin(sin\u03b1bsin\u03b4b) lbDE = rbcos\u03b1b1 lbBC = Rbcos\u03b1b3 lbCD = rbsin\u03b1b1 + Resin\u03b1b3 (13) For the epicyclic gear-rack train that is a new design by assuming the planet gear having an infinite diameter, a base coordinate system, sun rack-gear coordinate system, rack arm coordinate system, and planet rack coordinate system are built, as shown in Fig", + " The v is the number of virtual constraints. The virtual constraints are the constraint that restricts the helical joint F to rotate about xO-axis with joint E, the constraint that restricts the claw n the helical joint F to rotate about xO-axis with joint E, and the constraint that restricts the claw in the helical joint F to move along the xO-axis with the prismatic joint G. The novel metamorphic clamping equivalent mechanism is designed based on the equivalent mechanism of the epicyclic bevel gear train in Fig. 6. The links lAC, lCD, lDE, lBE form the equivalent mechanism of the epicyclic bevel gear train. The helical joint F is the lead screw nut and the corresponding lead screw is on the claw. The helical joint F is fixed on the link lDE and rotates about xO-axis with link lDE. Restricted by the prismatic joint G on the link lBE, the claw translates along the xO-axis driven by the joint F and rotates about zOaxis driven by the joint B. Thus, the novel metamorphic clamping equivalent mechanism has mobility 2" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_35_0000395_978-3-030-67750-3_16-Figure28.7-1.png", + "original_path": "designv11-35/openalex_figure/designv11_35_0000395_978-3-030-67750-3_16-Figure28.7-1.png", + "caption": "Fig. 28.7 Outlet 2\u2014pressure reading of Mode 2 nozzle", + "texts": [ + " The simulation was implemented to all nine steps in order to find the impact toward the outlet 2 pressure and velocity that affected the spray angle and distance. Figure 28.9 presents the pressure curve that is affected by the length of the nozzle. It clearly can be seen that the pressure at outlet 2 is inversely proportional to the nozzle length. As the nozzle length increases, the pressure will be decreased. The graph pattern is in a gradually decreasing manner. This is because, as the powder goes through along the nozzle, the pressure will continue to decrease due to the loss of potential energy. Figure 28.7 shows the pressure reading during Mode 2 nozzle was set. While, the velocity reading of Mode 2 nozzle set as illustrated in Fig. 28.8. Presented in Fig. 28.10 is the velocity of the outlet 2 against the nozzle length. It shows that as the length of the nozzle increases the velocity happening on outlet two also increases. This is because the velocity is directly proportional to the length of the nozzle. Based on this graph, it is concluded that the higher the length and the diameter of the outlet, the faster the velocity" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_35_0000008_j.autcon.2021.103573-Figure33-1.png", + "original_path": "designv11-35/openalex_figure/designv11_35_0000008_j.autcon.2021.103573-Figure33-1.png", + "caption": "Fig. 33. Configuration of the floating crane and the block.", + "texts": [ + " As the direction of the wind load in the x-direction was positive, the trolleys were controlled to move in the opposite direction. In the final stage, the position of the upper and lower trolley coincided as the block rotated by 90\u25e6. Similar to Case 1, the wind load did not have a great effect on the length of the wire rope. In this section, the floating crane was controlled for block erection using the same control method. The configuration of the floating crane and the block is illustrated in Fig. 33. The floating crane had two booms and four wire ropes to lift the target block. The length, breadth, and height of the block was 23 m, 43 m, and 11 m, and its weight was 300 tons. The lifting capacity of the floating crane was 3600 tons. The floating crane was moored to the quay and the seabed with mooring lines, which were modeled by massless springs. The properties of the floating crane are shown in Table 2. The desired trajectory of the block is shown in Fig. 34. The floating crane lifted the block by 10 m, and moved it 1 m in the x-direction by standing the booms" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_35_0001419_978-94-007-7401-8-Figure2.39-1.png", + "original_path": "designv11-35/openalex_figure/designv11_35_0001419_978-94-007-7401-8-Figure2.39-1.png", + "caption": "Fig. 2.39 4PaCP-type fully-parallel PM with coupled Sch\u00f6nflies motions defined by MF \u00bc SF \u00bc 4; RF\u00f0 \u00de \u00bc \u00f0v1; v2; v3;xd\u00de, TF = 0, NF = 18, limb topology Pa\\C\\P", + "texts": [ + "21s) Idem No. 11 20. 4PaPRP (Fig. 2.35b) Pa||P\\R||P (Fig. 2.21t) Idem No. 5 21. 4PaPRP (Fig. 2.36a) Pa\\P||R\\P (Fig. 2.21u) Idem No. 11 22. 4PaRPP (Fig. 2.36b) Pa\\R\\P\\kP (Fig. 2.21v) Idem No. 5 23. 4PCPa (Fig. 2.37a) P\\C\\Pa (Fig. 2.21w) Idem No. 5 24. 4PPaC (Fig. 2.37b) P||Pa\\C (Fig. 2.21x) The last joints of the four limbs have parallel axes 25. 4CPPa (Fig. 2.38a) C\\P||Pa (Fig. 2.21y) The first joints of the four limbs have parallel axes 26. 4PaPC (Fig. 2.38b) Pa||P\\C (Fig. 2.21z) Idem No. 24 27. 4PaCP (Fig. 2.39) Pa\\C\\P (Fig. 2.21z1) Idem No. 11 28. 4PPaPaR (Fig. 2.40) P\\Pa\\\\Pa\\kR (Fig. 2.22a) Idem No. 4 29. 4PPaRPa (Fig. 2.41) P\\Pa||R\\Pa (Fig. 2.22b) The axes of the revolute joints connecting the two parallelogram loops of the four limbs are parallel (continued) 2.2 Topologies with Complex Limbs 175 Table 2.3 (continued) No. PM type Limb topology Connecting conditions 30. 4PRPaPa (Fig. 2.42) P\\R||Pa\\Pa (Fig. 2.22c) Idem No. 5 31. 4RPPaPa (Fig. 2.43) R\\P\\kPa\\\\Pa (Fig. 2.22d) Idem No. 15 32. 4PaPPaR (Fig" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_35_0001661_pi-a.1955.0031-Figure13-1.png", + "original_path": "designv11-35/openalex_figure/designv11_35_0001661_pi-a.1955.0031-Figure13-1.png", + "caption": "Fig. 13.\u2014Relevant to the analytical representation of an approximate equivalent system. (a) Alternator sections at retaining ring.", + "texts": [ + " = /, \\B* Uyk 0 \u2014 cos y cos \u2022 - cos y 0 \u2014 cos a \u2014 cos j8 cos a 0 \"I Vbn bl2 b{{ 1 1 J L^31 *32 ^33. h A where bn . . . i above. Substituting eqn. (14) in eqn. (13) gives /CM 12 21 22 (14) result from the calculations described A (15) + Ar13/1/3, etc., where kn = \u2014 b2l cos yi.e. Fx = knlf + k + b3l cos j8, etc. The terms kn . . . k33 were evaluated at a number of points along the reference coil and the results are shown graphically in Fig. 3. (10.2) Analysis of an Equivalent Two-Dimensional System Fig. 13(a) shows a drawing of the end-winding region of a turbo-alternator. The system analysed represents a section through the stator-coil knuckles as indicated. Fig. 13(6) shows the equivalent system assumed. The effect of the stator winding is represented by a current sheet located at the mean radius i?3 of the conductors. This current sheet has a sinusoidally distributed m.m.f. equal to the fundamental component associated with the a.c. stator current, /, and given by M = IM0cosd (16) where Mo is a function of the stator-winding parameters and has a value of 15-7 gilberts/amp for the machine considered. The magnetic retaining-ring and the proximity of the stator iron are represented by iron surfaces at, respectively, a known radius R{ and a radius R2 to be given a suitable value consistent with test results. It is required to determine the flux distribution in 1 he annular air-space between the two cylindrical iron surfaces in the presence of the concentric current sheet. In regions (i) and (ii) of Fig. 13(6) the following equations can be written relating tkz radial, BR, and circumferential, BQ, components of magnetic flux density at any point (R, 6) = 0 \u2022 \u2022 (17) The general solutions5 for eqns. (17) are BR = (CiR\"-1 + DiR-\"-*) sin n(9 - 6 BQ \u2014 (C2R n~1 + D2R~n~ 0 sin n(fi \u2014 6 the constants having different values in regions (i) and (ii). (18) YOUNG AND TOMPSETT: SHORT-CIRCUIT FORCES ON TURBO-ALTERNATOR END-WINDINGS 111 (b) Two-dimensional model to same scale. The boundary conditions are B'Q = 0 B'o' = 0 B'R = BR (R = R{) (R = i", + "2> h \u2022 (19)* J Considering these conditions and the current sheet at R3 with m.m.f. of value JM0 cos 6, the solutions to eqns. (18) take the form * * - - \u2022 (20) R] R R-2 R+ W From search-coil records on the alternator, values of BR (at R = RR) and BQ (R = RQ) were available corresponding to low currents and steady-state conditions. A particular value of R2 was found which when substituted in eqns. (20) with the values of Rlt R3, R, MQ and / gave results corresponding closely with the test figures. It can be seen from Fig. 13(6) that the value of R2 represents approximately a mean path to the end of the stator iron. Using this value of R2 and setting Rx = 0 to represent the removal of the rotor resulted in values of BR and BQ agreeing reasonably well with those measured on the replica stator tests. In order to represent the effect of saturation and eddy currents in the retaining ring, a current sheet is introduced at R \u2014 Ri having an m.m.f. of value \u2014AM0 COS 6. This results in modification of the results in eqn. (20) for region (i), these now taking the form \u2014Mn (21) or, substituting the numerical values of the parameters BR = 0-260/ - 0-276A r" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_35_0001584_978-1-4614-8544-5_15-Figure15.21-1.png", + "original_path": "designv11-35/openalex_figure/designv11_35_0001584_978-1-4614-8544-5_15-Figure15.21-1.png", + "caption": "FIGURE 15.21. Behavior of 4 as a function of , .", + "texts": [ + " Example 578 Frequency response at invariant frequencies. The frequency response is a function of , , and . Damping always diminishes the amplitude of vibration, so at rst we set = 0 and plot the behavior of as a function of , . Figure 15.19 illustrates the behavior of at the second invariant frequency 2. Because lim 0 2 = 1 (15.77) 2 starts at one, regardless of the value of . The value of 2 is always greater than one. Figure 15.20 shows that 3 is not a function of and is a decreasing function of . Figure 15.21 shows that 4 1 regardless of the value of and . The relative behavior of 2, 3, and 4 is shown in Figures 15.22 and 15.23. Example 579 Natural frequencies and vibration isolation of a quarter car. For a modern typical passenger car, the values of natural frequencies are around 1Hz and 10Hz respectively. The former is due to the bounce of sprung mass and the latter belongs to the unsprung mass. At average speed, bumps with wavelengths much greater than the wheelbase of the vehicle, will excite bounce motion of the body" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_35_0001480_978-3-642-22647-2_16-Figure6-1.png", + "original_path": "designv11-35/openalex_figure/designv11_35_0001480_978-3-642-22647-2_16-Figure6-1.png", + "caption": "Fig. 6 Centraized lubrication system for a large diesel engine", + "texts": [ + " The extent of the yield shift will depend on the feed and operating conditions. As a metric for the quality of a base oil hydrotreating operation, one can use run length, ppm N, and ppm S of the feed for the ISODEWAXING\u00ae catalyst and the degree of aromatics hydrogenation. At a given wax content, the VI is an easily accessible measure for the degree of hydrogenation because aromatic molecules exhibit a lower VI than the corresponding naphthenic molecules (Fig. 4). Lube Hydrotreating and Hydrocracking, Fig. 6 Hydrocracking can dramatically improve the VI of feeds derived from crudes that exhibit only a low concentration of paraffins If vacuum gas oils based on highly aromatic or highly naphthenic crudes are subjected to deep hydrotreating only, the resulting hydrotreated product is frequently not paraffinic enough to merit further processing with an ISODEWAXING\u00ae or other dewaxing catalyst. Due to this lack of paraffins, the viscosity of these gas oils is highly responsive to changes in temperature (i", + " Combining the impact of base oil hydrotreating and base oil hydrocracking, one can see that the VI of the fraction in the base oil boiling L range increases through a combination of aromatics hydrogenation (Fig. 4) and paraffin concentration (Fig. 5). Thus, hydrocracking affords the profitable upgrading of fairly low-paraffin vacuum gas oils like those derived from the Alaskan North Slope. However, it has its limitations; upgrading of a gas oil derived from San Joaquin Valley crude is less economical (Fig. 6). It is of interest to note that not all hydrocracking catalysts are suited for base oil production. Thus, typically fuel-directed hydrocracking catalysts tend to selectively turn high-VI molecules into fuels, whereas tailored base oil-directed hydrocracking catalysts will leave these high-VI molecules in the unconverted oil fraction (Fig. 7). Hydrotreating, hydrocracking, or a combination thereof greatly enhances the flexibility of upgrading feeds into base oil. It allows upgrading gas oils from various crude oil sources into base oil, and it allows using various refinery streams, such as solvent-extracted raffinates" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_35_0000008_j.autcon.2021.103573-Figure14-1.png", + "original_path": "designv11-35/openalex_figure/designv11_35_0000008_j.autcon.2021.103573-Figure14-1.png", + "caption": "Fig. 14. Rotational motion of the block from initial to final orientation.", + "texts": [ + " c(t) = \u23a7 \u23a8 \u23a9 cI(t) for t0 \u2264 t < \u03c40 cII(t) for \u03c40 \u2264 t < tf \u2212 \u03c40 cIII(t) for tf \u2212 \u03c40 \u2264 t < tf ( \u03c4 = tf \u2212 t0 \u03c40 : accleration/deceleration time ) (28) The rotational motion of the block can also be described by its initial and final orientation, ORa and ORe. The rotational transformation matrix aRe can be calculated by the given value: aRe(\u03c6a,\u03c6e) = ORT a(\u03c6a) ORe(\u03c6e) (29) For rotational motion planning, we assume the block rotates from its initial to the final orientation about a certain axis of rotation, ere, and angle, sre, as illustrated in Fig. 14. By using the Rodrigues\u2019 rotation formula, the rotational transformation matrix is described by the axis of rotation and angle: aRe(er, sre) = cos(sre)I3 + sin(sre)e\u0303r +(1 \u2212 cos(sre) )erer T (30) Then, using Eqs. (28) and (29), the rotational axis and angle for obtaining the final orientation can be calculated as: sre = arccos ( 1 2 ( aRe(1, 1) + aRe ( 2 , 2 ) + aRe ( 3, 3) \u2212 1 ) ) er = 1 sin(sre) \u23a1 \u23a2 \u23a2 \u23a3 aRe(3, 2) \u2212 aRe(2, 3) aRe(1, 3) \u2212 aRe(3, 1) aRe(2, 1) \u2212 aRe(1, 2) \u23a4 \u23a5 \u23a5 \u23a6, (31) where aRe(i, j) is ith row and jth column of aRe (\u03c6a, \u03c6e)" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_35_0000255_9781119246213.ch2-Figure2.1-1.png", + "original_path": "designv11-35/openalex_figure/designv11_35_0000255_9781119246213.ch2-Figure2.1-1.png", + "caption": "Figure 2.1 Network structure of a network with a directed spanning tree", + "texts": [ + "37 was used in [23, 151], where pinning control synchronization of complex networks was considered: a reference node v as an isolated node is first chosen, and some pinning controllers are then designed to ensure that the whole network is synchronized with the state of node v. In this section, a more general discussion on the consensus of a virtual network has been carried out. If the graph with node set \ud835\udcb1 \u2212 {v} has a spanning tree, then similar analytical results can be obtained by using Theorem 2.35. 2.2.5 Simulation Examples In this subsection, a simulation example is given to verify the theoretical analysis. Consider the multi-agent system (2.2), where the network structure is shown in Fig. 2.1, the coupling strength c = 11, and the nonlinear function f is described by Chua\u2019s circuit [24] f (xi(t)) = \u239b \u239c\u239c\u239c\u239d \ud835\udefc(\u2212xi1 + xi2 \u2212 l(xi1)), xi1 \u2212 xi2 + xi3, \u2212\ud835\udefdxi2, \u239e \u239f\u239f\u239f\u23a0 , (2.43) where l(xi1) = bxi1 + 0.5(a \u2212 b)(|xi1 + 1| \u2212 |xi1 \u2212 1|). The system (2.43) is chaotic when \ud835\udefc = 10, \ud835\udefd = 18, a = \u22124\u22153, and b = \u22123\u22154, as shown in [142]. In view of Assumption 2.21, by computations, one obtains \ud835\udf03 = 10.3246. From Fig. 2.1, it is easy to see that the network contains a directed spanning tree where the nodes 1\u20134 and 5\u20137 belong to the first and second strongly connected components, respectively. By Lemma 2.28 and Definition 2.33, one has a(L\u030311) = 1.8118 and b(L\u030322) = 1.0206, where \ud835\udf091 = (0.2727, 0.1818, 0.1364, 0.4091)T and \ud835\udf092 = (0.4615, 0.3077, 0.2308)T . By Theorem 2.35, one has that cmin{a(L\u030311), b(L\u030322)} = 11.2266 > \ud835\udf03 = 10.3246. Therefore, consensus can be achieved in this example of the multi-agent system (2.2)" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_35_0001419_978-94-007-7401-8-Figure5.15-1.png", + "original_path": "designv11-35/openalex_figure/designv11_35_0001419_978-94-007-7401-8-Figure5.15-1.png", + "caption": "Fig. 5.15 3PaPPR-1RPaPaP-type fully-parallel PMs with uncoupled Sch\u00f6nflies motions defined by MF = SF = 4, (RF) = ( :v1; v2; v3;xb:), TF = 0, NF = 21, limb topology R||Pa||Pa||P and Pa||P\\P\\kR, Pa||P\\P\\\\R (a), Pa\\P\\kP\\kR, Pa\\P\\kP\\\\R (b)", + "texts": [ + " The last joints of of limbs G1, G2 and G3 have superposed axes, and their revolute joints between links 4 and 5 have orthogonal axes Pa\\R\\P\\kR (Fig. 5.2g) R\\R\\R\\P\\kR\\R (Fig. 5.1b) 12. 2PaRRRR1PaRRR1RUPU (Fig. 5.14) Pa\\R||R||R\\kR (Fig. 5.3d) Idem No. 11 Pa\\R||R||R (Fig. 5.2h) R\\R\\R\\P\\kR\\R (Fig. 5.1b) 5.1 Fully-Parallel Topologies 527 Table 5.2 Limb topology and connecting conditions of the fully-parallel solutions with no idle mobilities presented in Figs. 5.15, 5.16, 5.17, 5.18, 5.19, 5.20, 5.21, 5.22, 5.23, 5.24, 5.25, 5.26, 5.27, 5.28, 5.29, 5.30 No. PM type Limb topology Connecting conditions 1. 3PaPPR-1RPaPaP (Fig. 5.15a) Pa||P\\P\\kR (Fig. 5.2a) The last joints of the four limbs have superposed axes/directions Pa||P\\P\\\\R (Fig. 5.2d) R||Pa||Pa||P (Fig. 5.4k) 2. 3PaPPR-1RPaPaP (Fig. 5.15b) Pa\\P\\kP\\kR (Fig. 5.2b) Idem No. 1 Pa\\P\\kP\\\\R (Fig. 5.2c) R||Pa||Pa||P (Fig. 5.4k) 3. 3PaPPR-1RPPaPa (Fig. 5.16a) Pa||P\\P\\kR (Fig. 5.2a) The revolute joint between links 7D and 9 of G4 limb and the last revolute joints of limbs G1, G2 and G3 have superposed axes Pa||P\\P\\\\R (Fig. 5.2d) R||P||Pa||Pa (Fig. 5.4m) 4. 3PaPPR-1RPPaPa (Fig. 5.16b) Pa\\P\\kP\\kR (Fig. 5.2b) Idem No. 3 Pa\\P\\kP\\\\R (Fig. 5.2c) R||P||Pa||Pa (Fig. 5.4m) 5. 3PaPPR1RPaPatP (Fig. 5.17a) Pa||P\\P\\kR (Fig. 5.2a) Idem No. 1 Pa||P\\P\\\\R (Fig" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_35_0000132_lra.2021.3062583-Figure1-1.png", + "original_path": "designv11-35/openalex_figure/designv11_35_0000132_lra.2021.3062583-Figure1-1.png", + "caption": "Fig. 1. Structure of the wristed percutaneous robot consisting mainly of a needle, a flexible wrist, and a rigid arm. Magnified views of the (a) distal fixture and (b) proximal fixture. (c) Cross-section of the rigid arm. (d) Silicone sheath that fits over the flexible wrist.", + "texts": [ + " 3) The robot should be able to bend at its distal tip to steer the tip away from the beating heart and facilitate the guidewire to stay inside the pericardial space after passing through the robot lumen. The flexible tip of the robot should have large bending curvature (>150 m\u22121) to allow 90\u25e6 bending with less than 10 mm bending length to cater to pericardial effusion depth between 10 mm-20 mm [26]. The percutaneous robot can be considered as a combination of a 12-Fr dilator and a 18-gauge needle, with the largest outer diameter being 4 mm. As seen in Fig. 1, it consists of three major parts, namely a 18-gauge bevel-tip stainless steel needle, a flexible wrist, and a rigid arm. The wrist is composed of a Nitinol miro-spring (Kellogg\u2019s Research Labs, USA) and two fixtures on two ends of the spring, namely the distal fixture (DF) and proximal fixture (PF). The DF, shown in two views in Fig. 1(a), features a needle hub on one end and four prongs on the other end to interface with the needle and the SMA spring, respectively. The PF, shown in two views in Fig. 1(b), features four prongs on one end and four straight legs on the other end to interface with the SMA spring and the rigid arm, respectively. An elastic sheath made of silicone (see Fig. 1(d)) is wrapped around the SMA spring as a protective cover that separates the cables and SMA spring from the surrounding biological tissues. The rigid arm is a hollow tube with a length of 80 mm and a 1.4 mm diameter lumen that is sufficiently large to accommodate a heat shrink tube and a Polytetrafluoroethylene (PTFE) tube, as shown in Fig. 1(c). The PTFE tube allows drainage of the pericardial fluid and passage of guidewire into the pericardial space. Its other function as a heating element will be discussed in Section II(B). There are four equidistantly-spaced 0.6 mm diameter rings at two ends of the rigid arm to hold four long rigid stainless steel tubes of 0.6 mm outer diameter along the periphery of the lumen of the rigid arm. The long tubes act as channels (refer to Fig. 1(c)) for the passage of actuation cables (Osaka Coat Rope Co., Ltd., Japan) with 0.21 mm diameter. This is a low-cost yet effective solution to compensate for the inability of commonly accessible manufacturing techniques to produce sub-millimeter diameter channels with large aspect ratios (>100). These channels serve as dedicated pathways for the cables, preventing cable tangling and more importantly, keeping the center of the lumen clear. As shown in Fig. 2, the SMA spring in the flexible wrist has a length ls of 6 mm, mean coil radius r of 0" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_35_0001419_978-94-007-7401-8-Figure2.116-1.png", + "original_path": "designv11-35/openalex_figure/designv11_35_0001419_978-94-007-7401-8-Figure2.116-1.png", + "caption": "Fig. 2.116 4PaRRRPa-type fully-parallel PM with coupled Sch\u00f6nflies motions defined by MF \u00bc SF \u00bc 4; RF\u00f0 \u00de \u00bc \u00f0v1; v2; v3;xd\u00de, TF = 0, NF = 26, limb topology Pa||R||R\\R||Pa", + "texts": [], + "surrounding_texts": [ + "170 2 Fully-Parallel Topologies with Coupled Sch\u00f6nflies Motions", + "2.2 Topologies with Complex Limbs 171", + "172 2 Fully-Parallel Topologies with Coupled Sch\u00f6nflies Motions" + ] + }, + { + "image_filename": "designv11_35_0000170_14763141.2021.1880619-Figure2-1.png", + "original_path": "designv11-35/openalex_figure/designv11_35_0000170_14763141.2021.1880619-Figure2-1.png", + "caption": "Figure 2. Critical instants and phases of discus throwing.", + "texts": [ + " The angular momentum of the system about the left-right axis represents the amount of top-to-front (positive) or top-to-back (negative) rotation of the system. The angular momentum of the system about the up-down axis represents the amount of right-to-left (negative) or right-to-left (positive) rotation of the system. All data reductions were performed using an MSDiscusAnalysis computer programme package (MotionSoft Inc., Chapel Hill, NC, USA). Five critical instants in discus throwing procedure were identified as (1) right foot takeoff, (2) left foot takeoff, (3) right touchdown, (4) left foot touchdown, and (5) release of discus (Figure 2; Hay & Yu, 1995). Paired t-tests were performed to compare official distance and partial distances, and normalised system angular momentum and three direction cosines of whole-body angular momentum vector at right foot takeoff, left foot takeoff, left foot touchdown, and release between long and short trials. System angular momentum at right foot touchdown was not analysed because system angular momentum does not change during airborne between left foot takeoff and right foot touchdown. A Type I error rate less than or equal to 0" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_35_0001419_978-94-007-7401-8-Figure3.15-1.png", + "original_path": "designv11-35/openalex_figure/designv11_35_0001419_978-94-007-7401-8-Figure3.15-1.png", + "caption": "Fig. 3.15 4PPPR-type overactuated PMs with coupled Sch\u00f6nflies motions defined by MF \u00bc SF \u00bc 4; RF\u00f0 \u00de \u00bc v1; v2; v3;xd\u00f0 \u00de, TF = 0, NF = 6, limb topology P\\P\\\\P\\kR (a) and P\\P\\\\P\\R (b)", + "texts": [ + "1n) The last prismatic joints of the four limbs have parallel directions 16. 4PRPR (Fig. 3.11b) P\\R||P||R (Fig. 3.1o) Idem No. 9 17. 4PRRP (Fig. 3.12a) P||R||R\\P (Fig. 3.1v) The first prismatic joints of the four limbs have parallel directions 18. 4RPRP (Fig. 3.12b) R||P||R\\P (Fig. 3.1q) Idem No. 2 19. 4PRPR (Fig. 3.13a) P||R\\P\\kR (Fig. 3.1r) Idem No. 1 20. 4RPPR (Fig. 3.13b) R||P\\P\\kR (Fig. 3.1s) Idem No. 2 21. 4RPPR (Fig. 3.14a) R\\P\\kP||R (Fig. 3.1t) Idem No. 2 22. 4RPRP (Fig. 3.14b) R\\P\\kR||P (Fig. 3.1u) Idem No. 2 23. 4PPPR (Fig. 3.15a) P\\P\\\\P\\kR (Fig. 3.1h0) Idem No. 9 24. 4PPPR (Fig. 3.15b) P\\P\\\\P\\\\R (Fig. 3.1w) Idem No. 9 25. 4PPPR (Fig. 3.16a) P\\P\\\\P||R (Fig. 3.1x) Idem No. 9 26. 4PPRP (Fig. 3.16b) P\\P||R\\P (Fig. 3.1y) The revolute joints of the four limbs have parallel axes 27. 4PPRP (Fig. 3.17a) P\\P\\\\R||P (Fig. 3.1z) Idem No. 26 28. 4PPRP (Fig. 3.17b) P\\P\\kR\\P (Fig. 3.1a0) Idem No. 26 29. 4PRPP (Fig. 3.18a) P\\R||P\\P (Fig. 3.1b0) Idem No. 26 30. 4PRPP (Fig. 3.18b) P\\R\\P\\kP (Fig. 3.1c0) Idem No. 26 31. 4PRPP (Fig. 3.19a) P||R\\P\\\\P (Fig. 3.1d0) Idem No. 26 32. 4RPPP (Fig. 3.19b) R\\P\\\\P\\\\P (Fig" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_35_0001419_978-94-007-7401-8-Figure6.34-1.png", + "original_path": "designv11-35/openalex_figure/designv11_35_0001419_978-94-007-7401-8-Figure6.34-1.png", + "caption": "Fig. 6.34 3PPaPR-1RPPaPa-type maximaly regular fully-parallel PM with Sch\u00f6nflies motions defined by MF = SF = 4, (RF) = (v1; v2; v3;xb), TF = 0, NF = 21, limb topology P||Pa\\P ??R, P||Pa\\P\\||R and R||P||Pa||Pa", + "texts": [ + " 3PPPaR-1PRPaPa (Fig. 6.32b) P\\P\\||Pa\\R (Fig. 4.8a) P\\P\\||Pa||R, (Fig. 4.8c) R||P||Pa||Pa (Fig. 5.4m) Idem no. 9 11. 3PPPaR-1RPPaPat (Fig. 6.33a) P\\P\\||Pa\\||R (Fig. 4.8b) P\\P\\||Pa||R, (Fig. 4.8c) R||P||Pa||Pat (Fig. 5.4n) Idem no. 9 (continued) 642 6 Maximally Regular Topologies with Sch\u00f6nflies Motions Table 6.13 (continued) No. PM type Limb topology Connecting conditions 12. 3PPPaR-1RPPaPat (Fig. 6.33b) P\\P\\||Pa\\\\R (Fig. 4.8a) P\\P\\||Pa||R, (Fig. 4.8c) R||P||Pa||Pat (Fig. 5.4n) Idem no. 9 13. 3PPaPR-1RPPaPa (Fig. 6.34) P||Pa\\P\\R (Fig. 4.8d) P||Pa\\P\\||R (Fig. 4.8e) R||P||Pa||Pa (Fig. 5.4m) Idem no. 9 14. 3PPaPR-1RPPatPa (Fig. 6.35) P||Pa\\P\\\\R (Fig. 4.8d) P||Pa\\P\\||R (Fig. 4.8e) R||P||Pa||Pat (Fig. 5.4n) Idem no. 9 15. 2PPaRRR-1PPaRR-1RPPaPa (Fig. 6.36) P||Pa||R||R\\R (Fig. 4.9a) P||Pa||R||R (Fig. 4.8g) R||P||Pa||Pa (Fig. 5.4m) Idem no. 9 16. 2PPaRRR-1PPaRR-1RPPaPat (Fig. 6.37) P||Pa||R||R\\R (Fig. 4.9a) P||Pa||R||R (Fig. 4.8g) R||P||Pa||Pat (Fig. 5.4n) Idem no. 9 17. 3PPaPaR-1RPaPaP (Fig. 6.38) P||Pa||Pa\\R (Fig" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_35_0001419_978-94-007-7401-8-Figure5.56-1.png", + "original_path": "designv11-35/openalex_figure/designv11_35_0001419_978-94-007-7401-8-Figure5.56-1.png", + "caption": "Fig. 5.56 2PaPaRRR-1PaPaRR-1RPPaPa-type fully-parallel PM with uncoupled Sch\u00f6nflies motions defined by MF = SF = 4, RF\u00f0 \u00de \u00bc \u00f0v1; v2; v3;xb\u00de, TF = 0, NF = 28, limb topology Pa\\Pa||R||R\\kR, Pa\\Pa||R||R and R||P||Pa||Pa", + "texts": [ + " 2PaPaRRR-1PaPaRR1RPaPatP (Fig. 5.53) Pa\\Pa||R||R\\\\R (Fig. 5.5a) Idem No. 1 Pa\\Pa||R||R (Fig. 5.4i) R||Pa||Pat||P (Fig. 5.4l) 20. 2PaPaRRR-1PaPaRR1RPaPatP (Fig. 5.54) Pa\\Pa||R||R\\kR (Fig. 5.5b) Idem No. 1 Pa\\Pa||R||R (Fig. 5.4i) R||Pa||Pat||P (Fig. 5.4l) 21. 2PaPaRRR-1PaPaRR1RPPaPa (Fig. 5.55) Pa\\Pa||R||R\\\\R (Fig. 5.5a) The revolute joint between links 7D and 10 of G4 limb and the last revolute joints of limbs G1, G2 and G3 have superposed axes Pa\\Pa||R||R (Fig. 5.4i) R||P||Pa||Pa (Fig. 5.4m) 22. 2PaPaRRR-1PaPaRR1RPPaPa (Fig. 5.56) Pa\\Pa||R||R\\kR (Fig. 5.5b) Idem No. 21 Pa\\Pa||R||R (Fig. 5.4i) R||P||Pa||Pa (Fig. 5.4m) 23. 2PaPaRRR-1PaPaRR1RPPaPat (Fig. 5.57) Pa\\Pa||R||R\\\\R (Fig. 5.5a) Idem No. 21 Pa\\Pa||R||R (Fig. 5.4i) R||P||Pa||Pat (Fig. 5.4n) 24. 2PaPaRRR-1PaPaRR1RPPaPat (Fig. 5.58) Pa\\Pa||R||R\\kR (Fig. 5.5b) Idem No. 21 Pa\\Pa||R||R (Fig. 5.4i) R||P||Pa||Pat (Fig. 5.4n) 25. 3PaPaPaR-1RPPP (Fig. 5.59) Pa\\Pa||Pa\\kR (Fig. 5.6a) Idem No. 1 Pa\\Pa||Pa||R (Fig. 5.6b) R\\P\\\\P\\\\P (Fig. 5.1a) 26. 3PaPaPaR-1RPPP (Fig. 5.60) Pa\\Pa||Pa\\\\R (Fig" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_35_0001419_978-94-007-7401-8-Figure5.68-1.png", + "original_path": "designv11-35/openalex_figure/designv11_35_0001419_978-94-007-7401-8-Figure5.68-1.png", + "caption": "Fig. 5.68 3PaPaPaR-1RPPaPa-type fully-parallel PM with uncoupled Sch\u00f6nflies motions defined by MF = SF = 4, RF\u00f0 \u00de \u00bc \u00f0v1; v2; v3;xb\u00de, TF = 0, NF = 39, limb topology Pa\\Pa||Pa\\\\R, Pa\\Pa||Pa||R and R||P||Pa||Pa", + "texts": [ + "54k) 31. 3PaPaPaR1RPaPatP (Fig. 5.65) Pa\\Pa||Pa\\kR (Fig. 5.6a) Idem No. 1 Pa\\Pa||Pa||R (Fig. 5.6b) R||Pa||Pat||P (Fig. 5.4l) 32. 3PaPaPaR1RPaPatP (Fig. 5.66) Pa\\Pa||Pa\\\\R (Fig. 5.6c) Idem No. 1 Pa\\Pa||Pa||R (Fig. 5.6b) R||Pa||Pat||P (Fig. 5.54l) 33. 3PaPaPaR-1RPPaPa (Fig. 5.67) Pa\\Pa||Pa\\kR (Fig. 5.6a) The revolute joint between links 7D and 11 of G4 limb and the last revolute joints of limbs G1, G2 and G3 have superposed axes Pa\\Pa||Pa||R (Fig. 5.6b) R||P||Pa||Pa (Fig. 5.4m) 34. 3PaPaPaR-1RPPaPa (Fig. 5.68) Pa\\Pa||Pa\\\\R (Fig. 5.6c) Idem No. 33 Pa\\Pa||Pa||R (Fig. 5.6b) R||P||Pa||Pa (Fig. 5.54m) Table 5.5 Limb topology and connecting conditions of the fully-parallel solutions with no idle mobilities presented in Figs. 5.69, 5.70, 5.71, 5.72, 5.73, 5.74, 5.75, 5.76, 5.77, 5.78, 5.79, 5.80, 5.81, 5.82, 5.83, 5.84, 5.85, 5.86, 5.87, 5.88, 5.89, 5.90, 5.91, 5.92 No. PM type Limb topology Connecting conditions 1. 3PaPaPaR1RPPaPat (Fig. 5.69) Pa\\Pa||Pa\\kR (Fig. 5.6a) The revolute joint between links 7D and 11 of G4 limb and the last revolute joints of limbs G1, G2 and G3 have superposed axes Pa\\Pa||Pa||R (Fig" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_35_0000360_j.matpr.2021.04.147-Figure3-1.png", + "original_path": "designv11-35/openalex_figure/designv11_35_0000360_j.matpr.2021.04.147-Figure3-1.png", + "caption": "Fig. 3. Side Impact test deformation.", + "texts": [], + "surrounding_texts": [ + "Off road vehicle is shaped to race and steer on different terrains. ORV is designed in such a way that it can endure off-roading terrains. In off-terrain circumstances, the vehicle bears dynamic loads and all that is sustained through the chassis frame. Chassis frame bears every mountings and assembly, so it is expected from an ORV chassis frame to sustain both static and dynamic loads. The selection of materials for chassis greatly depends on the high tensile strength and material light weight. The majority of manufacturers favour lightweight, cost-effective, safe, and recyclable materials." + ] + }, + { + "image_filename": "designv11_35_0000472_s00170-021-07362-2-Figure7-1.png", + "original_path": "designv11-35/openalex_figure/designv11_35_0000472_s00170-021-07362-2-Figure7-1.png", + "caption": "Fig. 7 IN718 square canonical part. aMesh configuration. b Part thin walls", + "texts": [ + " These elements are high order elements and supported by the advanced Ansys features adopted in the proposed framework (e.g., remeshing, remapping). Because these elements have 10 nodes and 4 integration points, transferring solution field variables between different mesh configurations is more accurate than using linear elements. Moreover, the use of tetrahedral mesh is advantageous as it allows having a high mesh growth rate and representing curvatures and complex features with less elements compared to the voxel-based remeshing approach. In this work, the IN718 canonical part (Fig. 7) and the IN625 cantilever (Fig. 8) were considered for validating the 3D modeling technique [7, 18]. Throughout the modeling steps, the mesh configuration for each part being built is divided into three regions: (1) N layers being solved before changing the mesh configuration, (2) previously solved layers, and (3) deactivated future layers. The meshmust be kept fine at the layers being solved and coarse at both the previously solved and future layers to reduce the number of nodes and stiffness matrix size", + "43 min 0.81 min Fig. 22 Effect of number of layers per task on computational time for NIST cantilever (with 1 CPU core, C=1.75) small regions having high strain gradients. The solution with C=1.75 could keep the highlighted details that could not be captured with C=3.5. Thus, depending on strain gradients at the different geometrical features, appropriate C values must be chosen. Low C values are recommended for parts that include complex features and thin walls, like the canonical example shown in Fig. 7. The adaptive remeshing parameter effects on the solution are further demonstrated by the final displacement solution. Figures 28 and 29 show the differences between the predicted cantilever distortions and experimental results. It is obvious all predictions are underestimated from the experimental solution, and the difference within the predicted results at different N and C combinations can hardly be observed. Thus, the Fig. 24 Effect of coarsening level on the total computational time for different N values (with 1 CPU core) model is valid within the range of remeshing parameters utilized in this work" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_35_0001419_978-94-007-7401-8-Figure5.29-1.png", + "original_path": "designv11-35/openalex_figure/designv11_35_0001419_978-94-007-7401-8-Figure5.29-1.png", + "caption": "Fig. 5.29 2PaRRRR-1PaRRR-1RPaPatP-type fully-parallel PMs with uncoupled Sch\u00f6nflies motions defined by MF = SF = 4, RF\u00f0 \u00de \u00bc \u00f0v1; v2; v3;xb\u00de, TF = 0, NF = 19, limb topology Pa\\R||R||R\\kR, Pa\\R||R||R and R||Pa||Pat||P", + "texts": [ + " The revolute joints between links 4 and 5 of limbs G1, G2 and G3 have orthogonal axes Pa\\R\\P\\kR (Fig. 5.2g) R||P||Pa||Pat (Fig. 5.4n) (continued) 5.1 Fully-Parallel Topologies 529 Table 5.2 (continued) No. PM type Limb topology Connecting conditions 21. 2PaRRRR-1PaRRR1RPaPaP (Fig. 5.27) Pa\\R||R||R\\kR (Fig. 5.3d) Idem No. 17 Pa\\R||R||R (Fig. 5.2h) R||Pa||Pa||P (Fig. 5.4k) 22. 2PaRRRR-1PaRRR1RPPaPa (Fig. 5.28) Pa\\R||R||R\\kR (Fig. 5.3d) Idem No. 20 Pa\\R||R||R (Fig. 5.2h) R||P||Pa||Pa (Fig. 5.4m) 23. 2PaRRRR-1PaRRR1RPaPatP (Fig. 5.29) Pa\\R||R||R\\kR (Fig. 5.3d) Idem No. 17 Pa\\R||R||R (Fig. 5.2h) R||Pa||Pat||P (Fig. 5.4l) 24. 2PaRRRR-1PaRRR1RPPaPat (Fig. 5.30) Pa\\R||R||R\\kR (Fig. 5.3d) Idem No. 20 Pa\\R||R||R (Fig. 5.2h) R||P||Pa||Pat (Fig. 5.4n) Table 5.3 Limb topology and connecting conditions of the fully-parallel solutions with no idle mobilities presented in Figs. 5.31, 5.32, 5.33, 5.34, 5.35, 5.36, 5.37, 5.38 No. PM type Limb topology Connecting conditions 1. 3PaPPaR1RPPP (Fig. 5.31a) Pa||P\\Pa\\kR (Fig. 5.4a) The last joints of the four limbs have superposed axes/directionsPa||P\\Pa||R (Fig" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_35_0001419_978-94-007-7401-8-Figure5.7-1.png", + "original_path": "designv11-35/openalex_figure/designv11_35_0001419_978-94-007-7401-8-Figure5.7-1.png", + "caption": "Fig. 5.7 3PaPPR-1RPPP-type fully-parallel PMs with uncoupled Sch\u00f6nflies motions defined by MF = SF = 4, RF\u00f0 \u00de \u00bc \u00f0v1; v2; v3;xb\u00de), TF = 0, NF = 15, limb topology R\\P\\\\P\\\\P and Pa||P\\P\\kR, Pa||P\\P\\\\R (a), Pa\\P\\kP\\kR, Pa\\P\\kP\\\\R (b)", + "texts": [ + "6a) The revolute joint between links 6D and 11 of G4 limb and the last revolute joints of limbs G1, G2 and G3 have superposed axes Pa\\Pa||Pa||R (Fig. 5.6b) C||Pa||Pa (Fig. 5.4o) 28. 3PaPaPaR1CPaPa (Fig. 5.90) Pa\\Pa||Pa\\\\R (Fig. 5.6c) Idem No. 27 Pa\\Pa||Pa||R (Fig. 5.6b) C||Pa||Pa (Fig. 5.54o) 29. 3PaPaPaR1CPaPat (Fig. 5.91) Pa\\Pa||Pa\\kR (Fig. 5.6a) Idem No. 27 Pa\\Pa||Pa||R (Fig. 5.6b) C||Pa||Pat (Fig. 5.4p) 30. 3PaPaPaR1CPaPat (Fig. 5.92) Pa\\Pa||Pa\\\\R (Fig. 5.6c) Idem No. 27 Pa\\Pa||Pa||R (Fig. 5.6b) C||Pa||Pat (Fig. 5.54p) 536 5 Topologies with Uncoupled Sch\u00f6nflies Motions No. Structural parameter Solution Figure 5.7 Figures 5.8, 5.9, 5.10 Figures 5.11 1. m 20 22 22 2. pi (i = 1, 3) 7 8 7 3. p2 7 7 7 4. p4 4 4 6 5. p 25 27 27 6. q 6 6 6 7. k1 1 1 1 8. k2 3 3 3 9. k 4 4 4 10. (RG1) (v1; v2; v3;xb) (v1; v2; v3;xa;xb) (v1; v2; v3;xb) 11. (RG2) (v1; v2; v3;xb) (v1; v2; v3;xb) (v1; v2; v3;xb) 12. (RG3) (v1; v2; v3;xb) (v1; v2; v3;xb;xd) (v1; v2; v3;xb) 13. (RG4) (v1; v2; v3;xb) (v1; v2; v3;xb) (v1; v2; v3;xa;xb;xd) 14. SGi (i = 1, 3) 4 5 4 15. SG2 4 4 4 16. SG4 4 4 6 17. rGi (i = 1, 2, 3) 3 3 3 18. rG4 0 0 0 19. MGi (i = 1, 3) 4 5 4 20" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_35_0000116_er.6543-Figure8-1.png", + "original_path": "designv11-35/openalex_figure/designv11_35_0000116_er.6543-Figure8-1.png", + "caption": "FIGURE 8 Structure of battery pack [Colour figure can be viewed at wileyonlinelibrary.com]", + "texts": [ + " TABLE 8 Parameters after correction Baseline 4_12_3 10_5_3 8_6_3 First Second Third Fourth Fifth 4_12_3 Experimental result 141 159 237 267 315 This simplification can greatly reduce the number of elements and nodes of battery module, so the FEM model size of battery pack assembly can be apparently smaller than the model using other simplified methods. In this section, the multibody model of battery module is connected with the FEM model of enclosure to combine into a simplified model of the battery pack assembly. By comparing with the FEM model using the cuboid simplified module, the superiority of this method is demonstrated. The simulation environment is Abaqus 6.13-4. The structure of the battery pack is shown in Figure 8A. The module in the battery pack is only constrained along the direction of the axis of the batteries. Figure 8B is a section view of the direction along the battery axis of the battery pack (the red dotted line), which can help to describe the assembly relationship within the battery pack as follows: 1. The battery modules are bolted together with end panels and current bus as described in section 2. In the FEM model, the free nodes of each bushing at the end of the battery module are connected with the corresponding end panel by the rigid beam element. 2. After the battery module is placed in the enclosure, wedge blocks are inserted at one end of the module to compress the module in the axial direction", + " The ratio of transverse and vertical dimension of these battery modules are all close to 2, so the parameters of module-2 (10-5-3) of which the ratio is 2 in section 4.4 are used in this FEM model for convenience. Another model is established using the traditional simplified method, which simplifies the module into a cuboid structure with orthotropic material, the parameter of this material is obtained also by optimization method. Battery module model using this method in Abaqus is shown in the bottom right of Figure 8A. The element size of the simplified module is 5 mm, increasing the element size of the simplified module can reduce the number of elements, but will increase the difference with the element size of battery enclosure model, which will lead to an additional error caused by the connected area between battery module and enclosure. Table 10 shows the comparison between the FEM models using the method this article proposed to the traditional method. These two models adopted the same FEM model of battery enclosure, and the difference between these two simplified methods will be presented by the FEM model of battery pack assembly removing the enclosure part" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_35_0000548_s0263574721000424-Figure3-1.png", + "original_path": "designv11-35/openalex_figure/designv11_35_0000548_s0263574721000424-Figure3-1.png", + "caption": "Figure 3. Phase portrait of the biped walker in periodic gait.", + "texts": [ + " After updating all agents\u2019 PBA, the memory of ith agent\u2019s neighbor, NBAi, is updated using the updated PBA. The updated NBAi is selected as the non-dominated agents in the PBA of ith agent\u2019s neighbors. The updating procedure is repeated until the iteration reaches the threshold, which is defined as the maximum number of iterations. The Pareto set that contains the solutions of MOMM optimization problem is the non-dominated agents in the NBA when the iteration ends. Inspecting several outputs from the walker help us to understand the characteristics of periodic gait. Figure 3 shows the phase trajectory of the biped walker and four instant markers (I\u2013IV) display the moments when the biped walker changes its phase. The instant marker I corresponds to time t = 0+, the instant moment after the rear and the swing legs are contacted with the ground (t = 0). Then, the rear leg that was contacted with the ground, supporting the walker in the previous step, loses contact with the ground and becomes the swing leg. The stick diagram shows a red leg as a swing leg and blue leg as a rear leg" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_35_0001419_978-94-007-7401-8-Figure5.48-1.png", + "original_path": "designv11-35/openalex_figure/designv11_35_0001419_978-94-007-7401-8-Figure5.48-1.png", + "caption": "Fig. 5.48 3PaPaPR-1RPPaPa-type fully-parallel PM with uncoupled Sch\u00f6nflies motions defined by MF = SF = 4, RF\u00f0 \u00de \u00bc \u00f0v1; v2; v3;xb\u00de, TF = 0, NF = 30 limb topology Pa\\Pa\\\\P\\\\R, Pa\\Pa\\\\P\\kR and R||P||Pa||Pa", + "texts": [ + "4h) R||Pa||Pa||P (Fig. 5.4k) 10. 3PaPaPR-1RPaPaP (Fig. 5.44) Pa\\Pa\\\\P\\\\R (Fig. 5.4g) Idem No. 1 Pa\\Pa\\\\P\\kR (Fig. 5.4e) R||Pa||Pa||P (Fig. 5.4k) 11. 3PaPaPR1RPaPatP (Fig. 5.45) Pa\\Pa\\kP\\\\R (Fig. 5.4f) Idem No. 1 Pa\\Pa\\kP\\kR (Fig. 5.4h) R||Pa||Pat||P (Fig. 5.4l) 12. 3PaPaPR1RPaPatP (Fig. 5.46) Pa\\Pa\\\\P\\\\R (Fig. 5.4g) Idem No. 1 Pa\\Pa\\\\P\\kR (Fig. 5.4e) R||Pa||Pat||P (Fig. 5.4l) 13. 3PaPaPR-1RPPaPa (Fig. 5.47) Pa\\Pa\\kP\\\\R (Fig. 5.4f) Idem No. 3 Pa\\Pa\\kP\\kR (Fig. 5.4h) R||P||Pa||Pa (Fig. 5.4m) 14. 3PaPaPR-1RPPaPa (Fig. 5.48) Pa\\Pa\\\\P\\\\R (Fig. 5.4g) Idem No. 3 Pa\\Pa\\\\P\\kR (Fig. 5.4e) R||P||Pa||Pa (Fig. 5.4m) 15. 3PaPaPR-1RPPaPat (Fig. 5.49) Pa\\PakP\\\\R (Fig. 5.4f) Idem No. 3 Pa\\Pa\\kP\\kR (Fig. 5.4h) R||P||Pa||Pat (Fig. 5.4n) 16. 3PaPaPR-1RPPaPat (Fig. 5.50) Pa\\Pa\\\\P\\\\R (Fig. 5.4g) Idem No. 3 Pa\\Pa\\\\P\\kR (Fig. 5.4e) R||P||Pa||Pat (Fig. 5.4n) 17. 2PaPaRRR-1PaPaRR1RPaPaP (Fig. 5.51) Pa\\Pa||R||R\\\\R (Fig. 5.5a) Idem No. 1 Pa\\Pa||R||R (Fig. 5.4i) R||Pa||Pa||P (Fig. 5.4k) (continued) 532 5 Topologies with Uncoupled Sch\u00f6nflies Motions Table 5" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_35_0000522_j.biosystems.2021.104451-Figure1-1.png", + "original_path": "designv11-35/openalex_figure/designv11_35_0000522_j.biosystems.2021.104451-Figure1-1.png", + "caption": "Fig. 1. Two-wheel differential drive model of swarm robot.", + "texts": [ + " Obstacle Detection is performed with the help of ultrasonic, IR-based and camera sensors in swarm robotics (Hedjar ve Bounkhel, 2014; Pandey, 2017). Robots may use these sensors to move around while avoiding obstacles and each other. The method developed in the present study supports Optical-based and Radio-based systems for the robot detector RBH system as well as the IR and ultrasonic based sensors for obstacle detection. A non-holonomic, 2 wheel driven differential mobile robot is taken as the reference of the swarm robot model in the present study. Fig. 1 shows the model of a two-wheel differential drive mobile robot. L2 has a circular structure with a half diameter. Two R-diameter wheels placed in parallel can be driven independently of each other. Thus, the swarm robot can rotate 360 degrees. The swarm robot can move forward at a velocity of v along the direction it is headed to. Swarm robot can move in the bounded GRF area expressed in equation (1). Here, vr is the velocity of the robot in the x direction and \u03c9r denotes the angular velocity of the robot. Swarm robot moves at a velocity of vr, angular velocity of \u03c9r rotating xr units in the xdirection and yr units in the y direction with an angle of \u03b8r. \u23a1 \u23a3 x\u0307r y\u0307r \u03b8\u0307r \u23a4 \u23a6= \u23a1 \u23a3 cos(\u03b8r) 0 sin(\u03b8r) 0 0 1 \u23a4 \u23a6 [ vr \u03c9r ] (1) where, xr, yr and \u03b8r are the parameters that must be added to the current position of the driving robot as expressed in equation (2). \u23a1 \u23a2 \u23a3 \u02d9xr(t + \u0394t) \u02d9yr(t + \u0394t) \u02d9\u03b8r(t + \u0394t) \u23a4 \u23a5 \u23a6 = \u23a1 \u23a2 \u23a3 \u02d9xr(t) \u02d9yr(t) \u02d9\u03b8r(t) \u23a4 \u23a5 \u23a6+ \u23a1 \u23a3 x\u0307r y\u0307r \u03b8\u0307r \u23a4 \u23a6 (2) The model presented in Fig. 1 Is used to determine the angular velocity of the left and right wheels (\u03c9L and \u03c9R) subject to the angular velocity vr and \u03c9r angular velocity. Thus, the swarm robot determines its direction based on \u03c9Left and \u03c9Right. In Equation (3), relationships for robot velocity vr and angular velocity \u03c9r are given for the left wheel velocity \u03c9L and right wheel velocity \u03c9R, respectively. vr = R ( \u03c9Left + \u03c9Right ) 2 \u03c9r = R ( \u03c9Left \u2212 \u03c9Right ) L (3) Subject to equation (3), left and right wheel angular velocities are expressed in equation (4), respectively as: \u03c9Left = vr \u2212 ( L 2 ) \u03c9r R \u03c9Right = vr + ( L 2 ) \u03c9r R (4) This section describes the control strategy of the DISA method" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_35_0000164_tie.2021.3060653-Figure2-1.png", + "original_path": "designv11-35/openalex_figure/designv11_35_0000164_tie.2021.3060653-Figure2-1.png", + "caption": "Fig. 2. Evolution of the two motion parts.", + "texts": [ + " In section \u2163, the influences of the HMCE on two motion parts are evaluated and compared. Meanwhile, the key parameters relating to the coupling effect are also analyzed. Then, two coupling effect suppression strategies are discussed in Section V, and the 2DoFDDIM machine prototype is implemented and tested to verify the effectiveness of the proposed techniques. Finally, section \u2165 draws the conclusions. II. HELICAL MOTION COUPLING EFFECT The topology of the linear motion stator is evolved from the rotary motion stator as shown in Fig. 2, where their outer stator diameter Dso and inner stator diameter Dsi of the two parts are respectively identical. The mechanical pole pitch of the linear motion stator, \u03c4pl, is the same as that of the rotary motion stator, \u03c4pr, which can be calculated as: _ 1 2 4 a r pr pl si p p l D N N (1) where la_r is the stack length of the rotary motion stator and Np is the number of pole pairs. electromagnetic pole pitch \u03c4epr is longer than that of the mechanical size \u03c4pr, as shown in Fig. 3. The air-gap flux density distributions of rotary motion components at rated sr = 0" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_35_0001385_978-981-15-6095-8-Figure2.3-1.png", + "original_path": "designv11-35/openalex_figure/designv11_35_0001385_978-981-15-6095-8-Figure2.3-1.png", + "caption": "Fig. 2.3 Mohr circle for calculation of the apparent cohesion for geocell-soil composite. Sourced from Hegde (2017)", + "texts": [ + " Bathurst and Karpurapu (1993) carried out a series of large-scale triaxial tests on 200 mm high isolated geocell specimen. From the test results, the significant improvement in the apparent cohesion was observed in the presence of geocell reinforcement. Rajagopal et al. (1999) conducted the triaxial compression tests on granular soil encased in a single and multiple geocells. Both geocell reinforced and unreinforced samples exhibited same frictional strength, but significant increment in apparent cohesion (Cr) was observed in the reinforced case as shown in Fig. 2.3. In figure, the small circle refers to the Mohr circle of the unreinforced soil. Due to the provision of geocell reinforcement, the confining stress increases from \u03c3 3 to \u03c3 3 + \u03c3 3 due to which the ultimate normal stress increases to \u03c3 1 from \u03c3 1u. The intermediate circle in the figure indicates the Mohr circle corresponding to this state. The same ultimate stress can also be represented with the larger Mohr circle, which has a confining pressure of \u03c3 3 with an apparent cohesion of Cr (Rajagopal et al" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_35_0000174_j.matpr.2021.01.640-Figure15-1.png", + "original_path": "designv11-35/openalex_figure/designv11_35_0000174_j.matpr.2021.01.640-Figure15-1.png", + "caption": "Fig. 15. Mode shapes 3.", + "texts": [ + "3411e-9 Engineering analysis Stress on Metal Matrix Composites as shown in Fig. 12.Stress analysis on metal matrix composite, von mises elastic theory is applied, the stress value obtained as maximum 74.301 MPa and minimum as 0.38088 MPa. etal Matrix Composites. By Modal analysis the characteristic frequencies and mode shapes shown in Fig. 13. In the event that the shat revolves at its characteristic recurrence, it very well may be seriously vibrated. The modal investigation performed to locate the normal frequencies Fig. 14 Fig. 15. Modal analysis in structural mechanics is determined the natural mode shapes and frequencies of an object or structure during free vibration. In this analysis the frequency varies 196.67 Hz, the mode shape level is 1. The mode shape can be extended on level 1, due to the rotations. The mode shapes level can be extend upto level 5. After applying boundary conditions, the modal analysis is to be perform. The free vibrations can be oscillated due to the drive shaft, the natural frequencies are obtained" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_35_0001385_978-981-15-6095-8-Figure5.1-1.png", + "original_path": "designv11-35/openalex_figure/designv11_35_0001385_978-981-15-6095-8-Figure5.1-1.png", + "caption": "Fig. 5.1 Geocell reinforcement with base geogrid", + "texts": [ + " Keywords Reinforced soil \u00b7 Geocell reinforcement \u00b7 Strip footing Owing to rapid urbanization and industrialization in the present days, the requirement for in situ treatment of foundation soil to improve its bearing capacity has risen markedly. Among the various ground improvement techniques used, geosynthetic reinforcement is probably themost popular one. This is primarily due to its simplicity, ease of construction and overall economy that find favor with the practicing engineers. The more recent advancement in this field is to provide three-dimensional confinement to the soil by using geocells. The geocell reinforcement consists of a series of interlocking cells constructed from polymer grid reinforcement (Fig. 5.1), which contains and confines the soil within its pockets. S. K. Dash (B) Indian Institute of Technology Kharagpur, Kharagpur 721302, India e-mail: sujit@civil.iitkgp.ac.in \u00a9 Springer Nature Singapore Pte Ltd. 2020 T. G. Sitharam et al. (eds.), Geocells, Springer Transactions in Civil and Environmental Engineering, https://doi.org/10.1007/978-981-15-6095-8_5 131 Several studies have been reported highlighting the beneficial use of geocells. Rea andMitchell (1978) andMitchell et al. (1979) conducted a series of model plate load tests on circular footings supported over sand-filled square-shaped paper grid cells", + " It has square-shaped apertures with opening size of 35 mm \u00d7 35 mm. Geometrical details of the geogrid are shown in Fig. 5.2. The properties of the geogrid determined from standard wide width tension tests (ASTMD-4595) are: ultimate tensile strength = 20 kN/m, elongation at yield = 23%, Secant modulus at 5% strain = 160 kN/m and Secant modulus at 10% strain = 125 kN/m. The geocell mattresses were prepared by cutting the geogrids to required length and height from full rolls and placing them in transverse and diagonal directions as shown in Fig. 5.1, with bodkin joints at the intersections (Simac 1990; Bush et al. 1990)). The bodkins in the present tests are 6 mm wide and 3 mm thick plastic strips made of low-density polypropylene. The joint strength of geocells, obtained through tensile tests, was found to be 4.75 kN/m. Such low strength of joints was adopted to scale down the overall strength of the geocell reinforcement, making it suitable for the model tests. Figure 5.3 depicts the test geometry considered in the present investigation. In total, five different series tests were carried out the details of which are presented in Table 5" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_35_0001419_978-94-007-7401-8-Figure6.1-1.png", + "original_path": "designv11-35/openalex_figure/designv11_35_0001419_978-94-007-7401-8-Figure6.1-1.png", + "caption": "Fig. 6.1 Maximaly regular fully-parallel PMs with Sch\u00f6nflies motions of types 3PPPR-1RPPP (a) and 2PPRRR-1PPRR-1RPPP (b) defined by MF = SF = 4, (RF) = (v1; v2; v3;xb), TF = 0, NF = 6 (a), NF = 4 (b), limb topology R\\P\\\\P\\\\P and P\\P\\\\P\\||R, P\\P\\\\P\\\\R (a), P\\P\\||R||R\\R, P\\P\\||R||R (b)", + "texts": [ + " 1b ) T he fi rs tr ev ol ut e jo in t of G 4 -l im b an d th e la st la st jo in ts of li m bs G 1 , G 2 an d G 3 ha ve pa ra ll el ax es T he la st re vo lu te jo in ts of li m bs G 1 , G 2 an d G 3 ha ve su pe rp os ed ax es . T he cy li nd ri ca l jo in ts of li m bs G 1 , G 2 an d G 3 ha ve or th og on al ax es 12 . 2 C R R R -1 C R R -1 R U P U (F ig . 6. 6b ) C ||R ||R ? R (F ig . 4. 2g ) C ||R ||R (F ig . 4. 1k ) R ? R ? R ? P ? || R ? R (F ig . 5. 1b ) Id em no . 11 6.1 Fully-Parallel Topologies with Simple Limbs 587 Table 6.2 Structural parametersa of parallel mechanisms in Figs. 6.1 and 6.2 No. Structural parameter Solution Figure 6.1a Figures 6.1b and 6.2 1. m 14 16 2. pi (i = 1, 3) 4 5 3. p2 4 4 4. p4 4 4 5. p 16 18 6. q 3 3 7. k1 4 4 8. k2 0 0 9. k 4 4 10. (RG1) (v1; v2; v3;xb) (v1; v2; v3;xa;xb) 11. (RG2) (v1; v2; v3;xb) (v1; v2; v3;xb) 12. (RG3) (v1; v2; v3;xb) (v1; v2; v3;xb) 13. (RG4) (v1; v2; v3;xb) (v1; v2; v3;xb) 14. SGi (i = 1, 3) 4 5 15. SG2 4 4 16. SG4 4 4 17. rGi (i = 1, 2, 3) 0 0 18. rG4 0 0 19. MGi (i = 1, 3) 4 5 20. MG2 4 4 21. MG4 4 4 22. (RF) (v1; v2; v3;xb) (v1; v2; v3;xb) 23. SF 4 4 24. rl 0 0 25. rF 12 14 26" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_35_0001419_978-94-007-7401-8-Figure2.67-1.png", + "original_path": "designv11-35/openalex_figure/designv11_35_0001419_978-94-007-7401-8-Figure2.67-1.png", + "caption": "Fig. 2.67 4RPaPPa-type fully-parallel PM with coupled Sch\u00f6nflies motions defined by MF \u00bc SF \u00bc 4; RF\u00f0 \u00de \u00bc \u00f0v1; v2; v3;xd\u00de, TF = 0, NF = 30, limb topology R\\Pa\\kP\\\\Pa", + "texts": [ + " 4 47. 4RPPaPa (Fig. 2.59) R\\P||Pa||Pa (Fig. 2.22t) Idem No. 15 48. 4PaPPaR (Fig. 2.60) Pa\\P\\\\Pa\\kR (Fig. 2.22u) Idem No. 4 49. 4PaRPaP (Fig. 2.61) Pa\\R\\Pa\\kR (Fig. 2.22v) Idem No. 34 50. 4PaRPPa (Fig. 2.62) Pa\\R||P\\Pa (Fig. 2.22w) Idem No. 34 51. 4PaPRPa (Fig. 2.63) Pa\\P||R\\Pa (Fig. 2.22x) Idem No. 33 52. 4PaPaPR (Fig. 2.64) Pa\\Pa\\\\P||R (Fig. 2.22y) Idem No. 4 53. 4PaPaRP (Fig. 2.65) Pa\\Pa\\\\R||P (Fig. 2.22z) Idem No. 41 54. 4RPaPaP (Fig. 2.66) R\\Pa\\\\Pa\\\\P (Fig. 2.22a0) Idem No. 15 55. 4RPaPPa (Fig. 2.67) R\\Pa\\kP\\\\Pa (Fig. 2.22b0) Idem No. 15 56. 4RPPaPa (Fig. 2.68) R||P\\Pa\\\\Pa (Fig. 2.22c0) Idem No. 15 57. 4PRPaPa (Fig. 2.69) P||R\\Pa\\\\Pa (Fig. 2.22d0) Idem No. 5 58. 4PaCPa (Fig. 2.70) Pa\\C\\Pa (Fig. 2.22e0) The axes of the cylindrical joints of the four limbs are parallel 59. 4CPaPa (Fig. 2.71) C\\Pa\\\\Pa (Fig. 2.22f 0) Idem No. 58 60. 4PaPaC (Fig. 2.72) Pa\\Pa\\\\C (Fig. 2.22g0) Idem No. 58 176 2 Fully-Parallel Topologies with Coupled Sch\u00f6nflies Motions T ab le 2. 4 L im b to po lo gy an d co nn ec ti ng co nd it io ns of th e fu ll y- pa ra ll el so lu ti on s w it h no id le m ob il it ie s pr es en te d in F ig s" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_35_0001419_978-94-007-7401-8-Figure6.16-1.png", + "original_path": "designv11-35/openalex_figure/designv11_35_0001419_978-94-007-7401-8-Figure6.16-1.png", + "caption": "Fig. 6.16 Maximaly regular fully-parallel PMs with Sch\u00f6nflies motions of types 2CRRR-1CRR1RPaPaP (a) and 2CRRR-1CRR-1RPaPatP (b) defined by MF = SF = 4, (RF) = (v1; v2; v3;xb), TF = 0, NF = 10, limb topology C||R||R\\R, C||R||R and R||Pa||Pa||P (a), R||Pa||Pat||P (b)", + "texts": [ + " 2PRRRR-1PRRR-1RPPaPa (Fig. 6.14a) P||R||R||R\\R (Fig. 4.2d) P||R||R||R (Fig. 4.1g) R||P||Pa||Pa (Fig. 5.4m) Idem no. 9 16. 2PRRRR-1PRRR1RPaPatP (Fig. 6.14b) P||R||R||R\\R (Fig. 4.2d) P||R||R||R (Fig. 4.1g) R||P||Pa||Pat (Fig. 5.4n) Idem no. 9 17. 2CPRR-1CPR-1RPaPaP (Fig. 6.15a) C\\P\\||R\\R (Fig. 4.2e) C ?P ?||R (Fig. 4.1i) R||Pa||Pa||P (Fig. 5.4k) Idem no. 1 18. 2CPRR-1CPR1RPaPatP (Fig. 6.15b) C ?P ?||R ?R (Fig. 4.2e) C ?P ?||R (Fig. 4.1i) R||Pa||Pat||P (Fig. 5.4l) Idem no. 1 19. 2CRRR-1CRR-1RPaPaP (Fig. 6.16a) C||R||R ?R (Fig. 4.2g) C||R||R (Fig. 4.1k) R||Pa||Pa||P (Fig. 5.4k) Idem no. 1 20. 2CRRR-1CRR1RPaPatP (Fig. 6.16b) C||R||R ?R (Fig. 4.2g) C||R||R (Fig. 4.1k) R||Pa||Pat||P (Fig. 5.4l) Idem no. 1 Table 6.6 Limb topology and connecting conditions of the fully-parallel solutions with no idle mobilities presented in Figs. 6.17, 6.18, 6.19, 6.20, 6.21, 6.22, 6.23, 6.24, 6.25, 6.26 No. PM type Limb topology Connecting conditions 1. 3PPPR-1CPaPaP (Fig. 6.17a) P ?P\\\\P ?||R (Fig. 4.1a) P ?P\\\\P\\\\R (Fig. 4.1b) C||Pa||Pa (Fig. 5.4o) The last revolute joints of limbs G1, G2 and G3 and the revolute joint between links 6D and 8 have superposed axes" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_35_0000477_j.procir.2021.05.085-Figure7-1.png", + "original_path": "designv11-35/openalex_figure/designv11_35_0000477_j.procir.2021.05.085-Figure7-1.png", + "caption": "Figure 7. Non-assembly without any support", + "texts": [ + " It has been observed that the non-assembly mechanism made without any support offers minimum build time (127 min) as compared to previously discussed cases due to the elimination of the support construction time. It has been observed that the absence of support at the build plate prevents the non-assembly to adhere to the build plate, therefore, upcoming layers are failed to comply with the previous layers. The shape of the part deteriorates due to the distortion created at the base features as shown in Fig. 7. This distortion also affects the movability of joints. For this reason, the multiarticulated joints in the non-assembly mechanisms build without any support are not observed as an appropriate alternative. This study aims to minimize the support material and build time in the construction of the multi-articulated joints in the nonassembly mechanism of a moveable horse structure. For this mean, support structure type and support placement are considered as key parameters to reduce the support structure without creating any distortion in the profiles and affecting the movability of the joints" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_35_0001419_978-94-007-7401-8-Figure3.5-1.png", + "original_path": "designv11-35/openalex_figure/designv11_35_0001419_978-94-007-7401-8-Figure3.5-1.png", + "caption": "Fig. 3.5 Overactuated PMs with coupled Sch\u00f6nflies motions of types 4RPRR (a) and 4RRPR (b) defined by MF \u00bc SF \u00bc 4; RF\u00f0 \u00de \u00bc v1; v2; v3;xd\u00f0 \u00de, TF = 0, NF = 6, limb topology R||P||R||R (a) and R||R||P||R (b)", + "texts": [ + "1 Topologies with Simple Limbs 241 242 3 Overactuated Topologies with Coupled Sch\u00f6nflies Motions Table 3.1 Limb topology and connecting conditions of the overactuated solutions with no idle mobilities presented in Figs. 3.4, 3.5, 3.6, 3.7, 3.8, 3.9, 3.10, 3.11, 3.12, 3.13, 3.14, 3.15, 3.16, 3.17, 3.18, 3.19, 3.20 No. PM type Limb topology Connecting conditions 1. 4PRRR (Fig. 3.4a) P||R||R||R (Fig. 3.1a) The prismatic joints of the four limbs have parallel directions 2. 4RPRR (Fig. 3.4b) R||P||R||R (Fig. 3.1b) The first revolute joints of the four limbs have parallel axes 3. 4RPRR (Fig. 3.5a) R||P||R||R (Fig. 3.1c) Idem No. 2 4. 4RRPR (Fig. 3.5b) R||R||P||R (Fig. 3.1d) Idem No. 2 5. 4RRPR (Fig. 3.6a) R||R||P||R (Fig. 3.1e) Idem No. 2 6. 4PPRR (Fig. 3.6b) P\\P\\kR||R (Fig. 3.1i) Idem No. 1 7. 4RRRP (Fig. 3.7a) R||R||R||P (Fig. 3.1f) Idem No. 2 8. 4RPRP (Fig. 3.7b) R\\P\\kR||P (Fig. 3.1g) Idem No. 2 9. 4PPRR (Fig. 3.8a) P\\P||R||R (Fig. 3.1h) The last revolute joints of the four limbs have parallel axes 10. 4PRPR (Fig. 3.8b) P\\R||P||R (Fig. 3.1p) Idem No. 9 11. 4RRPP (Fig. 3.9a) R||R||P\\P (Fig. 3.1j) Idem No. 2 12. 4RPRP (Fig. 3.9b) R||P||R\\P (Fig" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_35_0001419_978-94-007-7401-8-Figure6.35-1.png", + "original_path": "designv11-35/openalex_figure/designv11_35_0001419_978-94-007-7401-8-Figure6.35-1.png", + "caption": "Fig. 6.35 3PPaPR-1RPPaPat-type maximaly regular fully-parallel PM with Sch\u00f6nflies motions defined by MF = SF = 4, (RF) = (v1; v2; v3;xb), TF = 0, NF = 21, limb topology P||Pa\\P ??R, P||Pa\\P\\||R and R||P||Pa||Pat", + "texts": [ + " 3PPPaR-1RPPaPat (Fig. 6.33a) P\\P\\||Pa\\||R (Fig. 4.8b) P\\P\\||Pa||R, (Fig. 4.8c) R||P||Pa||Pat (Fig. 5.4n) Idem no. 9 (continued) 642 6 Maximally Regular Topologies with Sch\u00f6nflies Motions Table 6.13 (continued) No. PM type Limb topology Connecting conditions 12. 3PPPaR-1RPPaPat (Fig. 6.33b) P\\P\\||Pa\\\\R (Fig. 4.8a) P\\P\\||Pa||R, (Fig. 4.8c) R||P||Pa||Pat (Fig. 5.4n) Idem no. 9 13. 3PPaPR-1RPPaPa (Fig. 6.34) P||Pa\\P\\R (Fig. 4.8d) P||Pa\\P\\||R (Fig. 4.8e) R||P||Pa||Pa (Fig. 5.4m) Idem no. 9 14. 3PPaPR-1RPPatPa (Fig. 6.35) P||Pa\\P\\\\R (Fig. 4.8d) P||Pa\\P\\||R (Fig. 4.8e) R||P||Pa||Pat (Fig. 5.4n) Idem no. 9 15. 2PPaRRR-1PPaRR-1RPPaPa (Fig. 6.36) P||Pa||R||R\\R (Fig. 4.9a) P||Pa||R||R (Fig. 4.8g) R||P||Pa||Pa (Fig. 5.4m) Idem no. 9 16. 2PPaRRR-1PPaRR-1RPPaPat (Fig. 6.37) P||Pa||R||R\\R (Fig. 4.9a) P||Pa||R||R (Fig. 4.8g) R||P||Pa||Pat (Fig. 5.4n) Idem no. 9 17. 3PPaPaR-1RPaPaP (Fig. 6.38) P||Pa||Pa\\R (Fig. 4.10a) P||Pa||Pa||R (Fig. 4.10b) R||Pa||Pa||P (Fig. 5.4k) Idem no. 1 18. 3PPaPaR-1RPaPatP (Fig. 6.39) P||Pa||Pa\\R (Fig" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_35_0000636_j.mechmachtheory.2021.104432-Figure2-1.png", + "original_path": "designv11-35/openalex_figure/designv11_35_0000636_j.mechmachtheory.2021.104432-Figure2-1.png", + "caption": "Fig. 2. Suspension concept of the multi-link torsion axle (MLTA). (a) Generating the instantaneous centre of motion in side-view (b) Isometric view.", + "texts": [ + " As mentioned above, the advantages regarding the design space and the toe-in behaviour of the reversed installation, as a result of the relocated cross beam and body mounts, should be maintained. On the other hand, the amplification of the pitch angle during braking, as a result of the new IC, should be avoided. Thus, the location of the cross beam and the IC should be decoupled. For this purpose, an additional longitudinal link and flexible joints to the knuckle are introduced. With these, the reversed TBA is integrated into a longitudinal Watt\u2019s linkage (see Fig. 2(a) and (b)). The trailing links and arms are designed to create an IC that is located in front of the wheels. The projections of these trailing links and arms intersect at the new centre of motion. Additionally, this virtual IC should be located above the WC to ensure a desirable positive WC recession during jounce. Fig. 1(b) shows the increased and connected package space in the underbody that can be provided by the use of a reversed TBA or the MLTA, respectively. In this work, a Ford Fiesta as a typical B-segment vehicle was chosen as a reference", + " The revolute joint can, for example, be positioned at the point RL instead of RU. With this alternative design, the distance between the wheel patch and the revolute joint is reduced, leading to an overall reduction of the loads. On the other hand, the redundant DOFs r are converted into additional rotational DOFs for the wheels, which must be locked with supplementary links or joints (e.g. cardanic joints or integral links). This results in a more complex suspension design, and for this reason, the current work will focus on the initial mechanism design shown in Fig. 2(b) and Fig. 3. Similar to a conventional TBA, the integration of the reversed TBA into the longitudinal Watt\u2019s linkage also leads to two different motion states for the MLTA (see Eq. (1)). While the parallel wheel travel of the MLTA is mainly dependant on the design of the Watt\u2019s linkage and the orientation of the revolute axes, the opposite wheel travel is highly influenced by the nonlinear twisting and bending of the flexible cross beam and the deformations in the body bushings [21,22]; that is why simplifications are used for the ideal-kinematic modelling", + " \u2022 L must be located underneath the cross section of the rear rail of the body to ensure a connection with sufficient local stiffness. The position is defined by a straight line. As can be found in Figs. 6(b)\u2013(d), the hardpoints are considered as the centre points of their respective joints. For this reason, offsets are used to take the outer dimensions of the bushing geometry into account. T. Niessing and X. Fang Mechanism and Machine Theory 165 (2021) 104432 To integrate the MLTA into a vehicle, the ideal kinematic configurations of joints, shown in Fig. 3 or 4 (a), are replaced by a set of rubber bushings as shown in Fig. 2. The reasons for this are the overall lower production costs of the rubber bushings, the damping of road-induced vibrations, and the possibility of implementing different stiffnesses in the local directions [13,20]. In addition, the former simplification by the usage of rigid equivalent bodies representing the twist-beam substructure (see Section 3) needs to be reversed. These elasticities are modelled with two different approaches. In the first approach, the kinematically optimised hardpoints are implemented into the so-called FastCon-tool in which the components of a regular TBA, such as the trailing arms and the cross beam are modelled as simplified beams elements [34\u201336]", + " The local geometric constraints of the hardpoints L(linear) and RL(circular) can easily be seen in the Figure. Also, it can be observed that the scatter radius of RL is particularly high compared to the other hardpoints. This is, on the one hand, a consequence of the large cuboid volume (see Fig. 6(a)) but also implies a low sensitivity of that hardpoint on the total error function of the local minima. Also, in combination with the large range for L, multiple locally optimal positions and alignments of longitudinal link l2 (see Fig. 2(a)) are obtained. For the hardpoint U, on the other hand, 11 out of the 17 optimal points are within a 1 mm margin at the upper limit of the xcoordinate. For the z-coordinate of U, the 15 best solutions are within a range of 1.5 mm at the top limitation. This implies that U should be located as high and as far back as possible. Its limitations in the vehicle are defined by the rear floor and the rear crash structures of the BIW. For the hardpoint RU, the optimal solutions are found on the outside spindle radius (also see Fig" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_35_0000225_j.addma.2021.101955-Figure12-1.png", + "original_path": "designv11-35/openalex_figure/designv11_35_0000225_j.addma.2021.101955-Figure12-1.png", + "caption": "Fig. 12. Sample part model with dimensions.", + "texts": [ + " The model needs to be split into more layers if the part has dramatic changes in geometry internally. For a one time calculation, the Backward Interpolation algorithm may not have a significant computational advantage but shines when multiple distortions have to be calculated for the same part model with different hatch patterns. The distortion prediction method described was validated using a 316 L stainless steel plate built with 0.02 mm layer thickness. The dimensions of the sample are shown in Fig. 12, and the sample was built using the EOS M280 machine. The hatch angle of the sample is as shown in Fig. 13, where a raster pattern was utilized for all layers with a 90-degree rotation between consequent layers. The hatch spacing between the two scan paths was 0.1 mm. The laser power, the laser travel speed and the laser spot diameter were set at 195 W, 1083 mm/s, and 0.1 mm, respectively. The as-built sample before cutoff is shown in Fig. 14. The inherent strain of this sample part was calculated as the average residual plastic strain inside the interest area of a heat affected zone model simulated by ANSYS [46] with temperature history generated by a thermal simulation model developed by the authors [52]" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_35_0001419_978-94-007-7401-8-Figure3.126-1.png", + "original_path": "designv11-35/openalex_figure/designv11_35_0001419_978-94-007-7401-8-Figure3.126-1.png", + "caption": "Fig. 3.126 4RPRPaR-type overactuated PM with coupled Sch\u00f6nflies motions defined by MF \u00bc SF \u00bc 4; RF\u00f0 \u00de \u00bc v1; v2; v3;xd\u00f0 \u00de, TF = 0, NF = 14, limb topology R||P||R\\Pa||R", + "texts": [ + " 3.120) P||R||R\\R||Pa (Fig. 3.52m) Idem No. 15 34. 4RPRRPa (Fig. 3.121) R||P||R\\R||Pa (Fig. 3.52n) Idem No. 15 35. 4RPRRPa (Fig. 3.122) R||P||R\\R||Pa (Fig. 3.52o) Idem No. 15 36. 4RRPRPa (Fig. 3.123) R||R||P\\R||Pa (Fig. 3.52p) Idem No. 15 37. 4RRRPPa (Fig. 3.124) R||R\\R\\P\\kPa (Fig. 3.52q) Idem No. 15 (continued) 3.2 Topologies with Complex Limbs 361 Table 3.6 (continued) No. PM type Limb topology Connecting conditions 38. 4PRRPaR (Fig. 3.125) P||R||R\\Pa||R (Fig. 3.52r) Idem No. 13 39. 4RPRPaR (Fig. 3.126) R||P||R\\Pa||R (Fig. 3.52s) Idem No. 13 40. 4RPRPaR (Fig. 3.127) R||P||R\\Pa||R (Fig. 3.52t) Idem No. 13 41. 4RRPPaR (Fig. 3.128) R||R||P\\Pa||R (Fig. 3.52u) Idem No. 13 42. 4PPaRRR (Fig. 3.129a) P\\Pa\\kR||R||\\kR (Fig. 3.52v) Idem No. 13 43. 4PaPRRR (Fig. 3.129b) Pa\\P\\\\R||R\\\\R (Fig. 3.52w) Idem No. 13 44. 4PPaRRR (Fig. 3.130) P\\Pa\\\\R||R\\kR (Fig. 3.52x) Idem No. 13 45. 4PaPRRR (Fig. 3.131) Pa\\P||R||R\\kR (Fig. 3.52y) Idem No. 13 46. 4PaRPRR (Fig. 3.132) Pa\\R||P||R\\kR (Fig. 3.52z) Idem No. 13 47. 4PaRPRR (Fig" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_35_0001385_978-981-15-6095-8-Figure13.12-1.png", + "original_path": "designv11-35/openalex_figure/designv11_35_0001385_978-981-15-6095-8-Figure13.12-1.png", + "caption": "Fig. 13.12 a and b Skeleton view of the FLAC3D model: a unreinforced case; b geocell and geogrid reinforced case. Sourced from Hegde and Sitharam (2015c)", + "texts": [ + " The elastic-perfectly plastic Mohr\u2013Coulomb model was used to simulate the behavior of the subgrade soil and the infill soil. The geocell was modeled using the geogrid structural element while the pipe was modeled using the shell structural element available in FLAC3D. Linear elastic model was used to simulate the behavior of the geocell and the pipe. The rigid nature of the geocell joint was simulated by fixing the nodes representing the joints. The interface between the geocell and the soil was linearly modeled with Mohr\u2013Coulomb yield criterion. Figure 13.12a, b shows the skeleton view of the FLAC3Dmodel for the unreinforced and reinforced cases. Analyses were carried out under controlled velocity loading of 2.5\u00d7 E-5 m/step. Only a quarter portion of the test bed was modeled making use of the symmetry to reduce the computational effort. The quarter symmetric model of size 0.45 m \u00d7 0.45 m \u00d7 0.6 m was discretized into 10,320 zones. Sensitivity analyses were carried out to determine the mesh density and based on which, the relatively coarse mesh was chosen for the analysis" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_35_0001419_978-94-007-7401-8-Figure3.65-1.png", + "original_path": "designv11-35/openalex_figure/designv11_35_0001419_978-94-007-7401-8-Figure3.65-1.png", + "caption": "Fig. 3.65 4PPaPR-type overactuated PM with coupled Sch\u00f6nflies motions defined by MF \u00bc SF \u00bc 4; RF\u00f0 \u00de \u00bc v1; v2; v3;xd\u00f0 \u00de, TF = 0, NF = 18, limb topology P\\Pa||P||R", + "texts": [ + "50l) Idem No. 1 13. 4RPPPa (Fig. 3.60b) R||P\\P\\kPa (Fig. 3.50m) Idem No. 1 14. 4PaRPP (Fig. 3.61a) Pa||R||P\\P (Fig. 3.50n) Idem No. 1 15. 4RPaPP (Fig. 3.61b) R||Pa||P\\P (Fig. 3.50o) Idem No. 1 16. 4RPPaP (Fig. 3.62a) R\\P\\kPa||P (Fig. 3.50p) Idem No. 1 17. 4PPPaR (Fig. 3.62b) P\\P\\kPa||R (Fig. 3.50q) Idem No. 1 18. 4PPPaR (Fig. 3.63a) P\\P||Pa||R (Fig. 3.50r) Idem No. 1 19. 4PPaRR (Fig. 3.63b) P||Pa||R||R (Fig. 3.50s) Idem No. 1 20. 4PPaRP (Fig. 3.64) P\\Pa||R||P (Fig. 3.50t) Idem No. 1 21. 4PPaPR (Fig. 3.65) P\\Pa||P||R (Fig. 3.50u) Idem No. 1 22. 4PPaPR (Fig. 3.66a) P||Pa\\P\\kR (Fig. 3.50v) Idem No. 1 23. 4PPaRP (Fig. 3.66b) P||Pa\\R\\P (Fig. 3.50i0) The before last joints of the four limbs have parallel axes 24. 4RPPaP (Fig. 3.67) R||P||Pa\\P (Fig. 3.50w) Idem No. 1 25. 4RPRPa (Fig. 3.68) R||P||R||Pa (Fig. 3.50x) Idem No. 1 26. 4PaPPR (Fig. 3.69a) Pa||P\\P\\kR (Fig. 3.50y) Idem No. 1 27. 4PaPPR (Fig. 3.69b) Pa\\P\\kP||R (Fig. 3.50z) Idem No. 1 28. 4PaPRP (Fig. 3.70) Pa\\P\\kR||P (Fig. 3.50z1) Idem No. 1 29" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_35_0000596_s43452-021-00255-x-Figure15-1.png", + "original_path": "designv11-35/openalex_figure/designv11_35_0000596_s43452-021-00255-x-Figure15-1.png", + "caption": "Fig. 15 Stress countor a the double-lap adhesive joint, b adhesive area in the double-lap adhesive-rivet joint", + "texts": [ + " The von Mises stress contour and stress distribution in the internal layer of the joint, 9-rivet layout, are shown in Figs.\u00a012 and 13. According to Fig.\u00a0 14a observed that the maximum von Mises stress on the rivet equals 74.25\u00a0MPa. Also, in Fig.\u00a014b, the highest amount of the elastic stress distribution occurs in a lateral surface of holes (53.66\u00a0MPa). Results of simulation for the double-lap adhesive joint using ABAQUS software was showed that the highest amount of stress in the adhesive layer equal 17.61\u00a0MPa and occurs around holes and external edges (Fig.\u00a015a, b). 1 3 1 3 In the simulation of an adhesive joint, what is seen is the uniformity of tension that occurs at the surface of the joint. Also, creating tension concentration around the rivet holes in the structure is a feature of these rivet joints. In the composite bonding of the adhesive layer between the nanocomposite plates, it has withstood a large amount of stress due to loading. Due to its flexibility in the direction of tension, increasing the length of the adhesive layer is expected and, in the end, the damage that occurs is not from the adhesive area" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_35_0000395_978-3-030-67750-3_16-Figure28.3-1.png", + "original_path": "designv11-35/openalex_figure/designv11_35_0000395_978-3-030-67750-3_16-Figure28.3-1.png", + "caption": "Fig. 28.3 Nozzle exterior design improvement", + "texts": [ + " The aim is to create a high pressure and velocity stream of the fine particles flow out through the outlet, thus creating a narrow spray distance. As for mode two, as presented in Fig. 28.2. Figure 28.2(b), the dry powder particles for this mode flow two direction through outlet 1 and outlet 2. For the current nozzle, outlet 2 has a bigger diameter with 52 mm. The working principle of these outlets requires the nozzle to turn counter clockwise to allow the particles travel through two pathways. This is to allow the fluid to pass through a more significant diameter outlet (outlet 2) that create a wider spray angle with a slow velocity. Figure 28.3 illustrate the evolution of the bi-nozzle design to the existing one. The design begins with the setting up the project in SolidWorks. In the study on the nozzle, an internal analysis type was selected to simulate an internal flow. A list of pre-defined fluid already exists in SolidWorks but ammonium dihydrogen phosphate (Senthilkumaran et al. 2012) was not included in the library. This needs to be added manually. Thewall was assumed to be adiabatic. The roughness of the internal nozzle is also assumed to be zero" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_35_0001572_978-3-319-05371-4_7-Figure7.2-1.png", + "original_path": "designv11-35/openalex_figure/designv11_35_0001572_978-3-319-05371-4_7-Figure7.2-1.png", + "caption": "Fig. 7.2 Schematic of bicycle model", + "texts": [ + " The generalization to the case when both vehicles collaboratively try to avoid the collision will probably involve some vehicle-to-vehicle (V2V) communication and it is left for future investigation. Henceforth, we thus only consider the problem when only one (the bullet) vehicle is actively controlled. This problem was posed in [6, 7] as a time-optimal control problem, and it was solved using pseudospectral methods [25]. In the next two sections we briefly summarize the problem definition and its numerical solution. The model used in this chapter is the so-called \u201cbicycle model\u201d [26], augmented with wheel dynamics. The nomenclature and conventions regarding this model are shown in Fig. 7.2. The state is given by x = [u, v, r,\u03c8,\u03c9 f ,\u03c9r ]T, where u and v are, respectively, the body-fixed longitudinal and lateral velocities, r is the vehicle yaw rate, \u03c8 is the vehicle heading, and \u03c9 f \u2265 0 and \u03c9r \u2265 0 are the angular speeds of the front and rear wheels, respectively. The system is controlled by u = [\u03b4, Tb, Thb]T, where \u03b4 is the steering angle and Tb, Thb denote the torques generated by the footbrake and handbrake, respectively. The equations of motion of the vehicle can be written as shown in (7" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_35_0001584_978-1-4614-8544-5_15-Figure15.23-1.png", + "original_path": "designv11-35/openalex_figure/designv11_35_0001584_978-1-4614-8544-5_15-Figure15.23-1.png", + "caption": "FIGURE 15.23. Behavior of 3 2 as a function of , .", + "texts": [], + "surrounding_texts": [ + "The most employed and useful model of a vehicle suspension system is a quarter car model, shown in Figure 15.1. We introduce, examine, and optimize the quarter car model in this chapter. 15.1 Mathematical Model We may represent the vertical vibration of a vehicle using a quarter car model made of two solid masses and denoted as sprung and unsprung masses, respectively. The sprung mass represents 1 4 of the body of the vehicle, and the unsprung mass represents one wheel of the vehicle. A spring of sti ness , and a shock absorber with viscous damping coe cient , support the sprung mass and are called the main suspension. The unsprung mass is in direct contact with the ground through a spring , representing the tire sti ness. The governing di erential equations of motion for the quarter car model shown in Figure 15.1, are: \u00a8 + ( ) + ( ) = 0 (15.1) \u00a8 + ( ) + ( + ) = (15.2) R.N. Jazar, Vehicle Dynamics: Theory and Application, DOI 10.1007/978-1-4614-8544-5_15 \u00a9 Springer Science+Business Media New York 2014 985 Proof. The kinetic energy, potential energy, and dissipation function of the system are: = 1 2 2 + 1 2 2 (15.3) = 1 2 ( ) 2 + 1 2 ( ) 2 (15.4) = 1 2 ( ) 2 (15.5) Employing the Lagrange method,\u03bc \u00b6 + + = 0 (15.6)\u03bc \u00b6 + + = 0 (15.7) we nd the equations of motion \u00a8 = ( ) ( ) (15.8) \u00a8 = ( ) + ( ) ( ) (15.9) which can be expressed in a matrix form [ ] x\u0308+ [ ]x+ [ ]x = F (15.10) 0 0 \u00b8 \u00a8 \u00a8 \u00b8 + \u00b8 \u00b8 + + \u00b8 \u00b8 = 0 \u00b8 (15.11) Example 570 Tire damping. We may add a damper in parallel to , as shown in Figure 15.1, to model any damping in tires. However, the value of for tires, compared to , are very small and hence, we may ignore to simplify the model. Having the damper in parallel to makes the equation of motion the same as Equations (12.44) and (12.45) with a matrix form as Equation (12.47). Example 571 Mathematical model\u2019s limitations. The quarter car model contains no representation of the geometric e ects of the full car and o ers no possibility of studying longitudinal and lateral interconnections. However, it contains the most basic features of the real problem and includes a proper representation of the problem of controlling wheel and wheel-body load variations. In the quarter car model, we assume that the tire is always in contact with the ground, which is true at low frequency but might not be true at high frequency. A better model must be able to include the possibility of separation between the tire and ground. Optimal design of two-DOF vibration systems, including a quarter car model, is the subject of numerous investigations since the invention of the vibration absorber theory by Frahm in 1909. It seems that the rst analytical investigation on the damping properties of two-DOF systems is due to Den Hartog (1901 1989). 15.2 Frequency Response To nd the frequency response, we consider a harmonic excitation, = cos (15.12) and look for a harmonic solution in the form = 1 sin + 1 cos = sin ( ) (15.13) = 2 sin + 2 cos = sin ( ) (15.14) = = 3 sin + 3 cos = sin ( ) (15.15) where , , and are complex amplitudes. By introducing the following dimensionless characters: = (15.16) = r (15.17) = r (15.18) = (15.19) = (15.20) = 2 (15.21) we search for the absolute and relative frequency responses: = \u00af\u0304\u0304\u0304 \u00af\u0304\u0304\u0304 (15.22) = \u00af\u0304\u0304\u0304 \u00af\u0304\u0304\u0304 (15.23) = \u00af\u0304\u0304\u0304 \u00af\u0304\u0304\u0304 (15.24) and obtain the following functions: 2 = 4 2 2 + 1 2 1 + 2 2 (15.25) 2 = 4 2 2 + 1 + 2 \u00a1 2 2 \u00a2 2 1 + 2 2 (15.26) 2 = 4 2 1 + 2 2 (15.27) 1 = \u00a3 2 \u00a1 2 2 1 \u00a2 + \u00a1 1 (1 + ) 2 2 \u00a2\u00a4 (15.28) 2 = 2 \u00a1 1 (1 + ) 2 2 \u00a2 (15.29) The absolute acceleration of sprung mass and unsprung mass may be de ned by the following equations: = \u00af\u0304\u0304\u0304 \u00af \u00a8 2 \u00af\u0304\u0304\u0304 \u00af = 2 2 (15.30) = \u00af\u0304\u0304\u0304 \u00af \u00a8 2 \u00af\u0304\u0304\u0304 \u00af = 2 2 (15.31) Proof. To nd the frequency responses, let us apply a harmonic excitation = cos (15.32) and assume that the solutions are harmonic functions with unknown coefcients. = 1 sin + 1 cos (15.33) = 2 sin + 2 cos (15.34) Substituting the solutions in the equations of motion (15.1)-(15.2) and collecting the coe cients of sin and cos in both equations provides the following set of algebraic equations for 1, 1, 2, 2: [ ] 1 2 1 2 = 0 0 0 (15.35) where [ ] is the coe cient matrix. [ ] = 2 2 + 2 + 2 (15.36) The unknowns may be found by matrix inversion 1 2 1 2 = [ ] 1 0 0 0 (15.37) and therefore, the amplitudes and can be found. 2 = 2 1 + 2 1 = \u00a1 2 2 + 2 \u00a2 2 3 + 2 4 2 (15.38) 2 = 2 2 + 2 2 = \u00a1 4 2 + 2 2 2 2 + 2 \u00a2 2 3 + 2 4 2 (15.39) 3 = \u00a1 2 ( + + ) 4 \u00a2 (15.40) 4 = \u00a1 3 ( + ) \u00a2 (15.41) Having and helps us to calculate and its amplitude . = = ( 1 2) sin + ( 1 2) cos = 3 sin + 3 cos = sin ( ) (15.42) 2 = 2 3 + 2 3 = 4 2 2 3 + 2 4 2 (15.43) Taking derivative from and provides the acceleration frequency responses and for the unsprung and sprung masses. Equations (15.30)- (15.31) express and . Using the de nitions (15.16)-(15.21), we may transform Equations (15.38), (15.39), (15.43) to (15.25), (15.26), (15.27). Figures 15.2, 15.3, 15.4, are samples of the frequency responses , , and for = 3, and = 0 2. Example 572 Average value of parameters for street cars. Equations (15.25)-(15.27) indicate that the frequency responses , , and are functions of four parameters: mass ratio , damping ratio , natural frequency ratio , and excitation frequency ratio . The average, minimum, and maximum of practical values of the parameters are indicated in Table 14 1. For a quarter car model, it is known that , and therefore, 1. Typical mass ratio, , for vehicles lies in the range 3 to 8, with small cars closer to 8 and large cars near 3. The excitation frequency is equal to , when = 1 , and equal to , when = 1. For a real model, the order of magnitude of the sti ness is , so , and 1. Therefore, 1 at = . So, we expect to have two resonant frequencies greater than = 1. Example 573 F Three-dimensional visualization for frequency responses. To get a sense about the behavior of di erent frequency responses of a quarter car model, Figures 15.5 to 15.8 are plotted for = 375kg = 35000N m = 75kg = 193000N m (15.44) 15.3 F Natural and Invariant Frequencies The quarter car system is a two-DOF system and therefore it has two natural frequencies 1 , 2 : 1 = s 1 2 2 \u03bc 1 + (1 + ) 2 q (1 + (1 + ) 2) 2 4 2 \u00b6 (15.45) 2 = s 1 2 2 \u03bc 1 + (1 + ) 2 + q (1 + (1 + ) 2) 2 4 2 \u00b6 (15.46) The family of response curves for the displacement frequency response of the sprung mass, , are obtained by keeping and constant, and varying . This family has several points in common, which are at frequencies 1, 2, 3, 4, and 1, 2, 3, 4, 1 = 0 1 = 1 3 = 1 3 = 1 2 2 = 1 1 (1 + ) 2 2 2 4 4 = 1 1 (1 + ) 2 2 2 (15.47) 2 = s 1 2 2 \u03bc 1 + 2 (1 + ) 2 q (1 + 2 (1 + ) 2) 2 8 2 \u00b6 (15.48) 4 = s 1 2 2 \u03bc 1 + 2 (1 + ) 2 + q (1 + 2 (1 + ) 2) 2 8 2 \u00b6 (15.49) where 1 (= 0) 2 1 1 + 3 \u03bc = 1 \u00b6 4 (15.50) The corresponding transmissivities at 2 and 4 are 2 = 1 1 (1 + ) 2 2 2 (15.51) 4 = 1 1 (1 + ) 2 2 2 (15.52) The frequencies 1, 2, 3, and 4 are called invariant frequencies, and their corresponding amplitudes are called invariant amplitudes because they are not dependent on . However, they are dependent on the values of and . The order of magnitude of the natural and invariant frequencies are: 1 (= 0) 1 2 1 1 + 3 \u03bc = 1 2 \u00b6 2 4 (15.53) The curves for have no other common points except 1, 2, 3, 4. The order of frequencies along with the order of corresponding amplitudes can be used to predict the shape of the frequency response curves of the sprung mass . Figure 15.9 shows schematically the shape of the amplitude versus excitation frequency ratio . Proof. The natural and resonant frequencies of a system are at positions where the amplitude goes to in nity when damping is zero. Hence, the natural frequencies would be the roots of the denominator of the function. \u00a1 2 \u00a2 = 2 \u00a1 2 2 1 \u00a2 + \u00a1 1 (1 + ) 2 2 \u00a2 = 2 2 \u00a1 1 + (1 + ) 2 \u00a2 + 1 = 0 (15.54) The solution of this equation are the natural frequencies given in Equations (15.45) and (15.46). The invariant frequencies are independent of , so they can be found by intersecting the curves for = 0 and = . lim 0 2 = \u00b1 1 ( 2 ( 2 2 1) 2 2 ( + 1) + 1) 2 (15.55) lim 2 = \u00b1 1 ( 2 2 ( + 1) 1) 2 (15.56) Therefore, the invariant frequencies, , can be determined by solving the following equation: 2 \u00a1 2 2 1 \u00a2 + \u00a1 1 (1 + ) 2 2 \u00a2 = \u00b1 \u00a11 (1 + ) 2 2 \u00a2 (15.57) Using the (+) sign, we nd 1 and 3 with their corresponding transmissivities 1, and 3, 1 = 0 1 = 1 (15.58) 3 = 1 3 = 1 (15.59) and, with the ( ) sign, we nd the following equation for 2, and 4: 2 4 \u00a1 1 + 2 (1 + ) 2 \u00a2 2 + 2 = 0 (15.60) Equation (15.60) has two real positive roots, 2 and 4, 2 = s 1 2 2 \u03bc 1 + 2 (1 + ) 2 q (1 + 2 (1 + ) 2) 2 8 2 \u00b6 (15.61) 4 = s 1 2 2 \u03bc 1 + 2 (1 + ) 2 + q (1 + 2 (1 + ) 2) 2 8 2 \u00b6 (15.62) with the following relative order of magnitude: 1 (= 0) 2 1 1 + 3 \u03bc = 1 \u00b6 4 (15.63) The corresponding amplitudes at 2, and 4 can be found by substituting Equations (15.61) and (15.62) in (15.25). 2 = 1 1 (1 + ) 2 2 2 (15.64) 4 = 1 1 (1 + ) 2 4 2 (15.65) It can be checked that (1 + ) 2 2 4 1 1 (15.66) and hence, | 4| 1 (= 3) 1 | 2| (15.67) and therefore, 2 1 (15.68) 4 1 (15.69) Using Equation (15.54), we can evaluate \u00a1 2 2 \u00a2 , \u00a1 2 4 \u00a2 , and \u00a1 2 3 \u00a2 as\u00a1 2 2 \u00a2 = (1 + ) 2 2 2 1 0 (15.70)\u00a1 2 4 \u00a2 = (1 + ) 2 2 4 1 0 (15.71)\u00a1 2 3 \u00a2 = \u03bc 1 2 \u00b6 (15.72) therefore, the two positive roots of Equation (15.54), 1 and 2 \u00a1 2 2 \u00a2 , have the order of magnitudes as: 1 (= 0) 1 2 1 1 + 3 \u03bc = 1 2 \u00b6 2 4 (15.73) Example 574 F Nodes of the absolute frequency response . There are four nodes in the absolute displacement frequency response of a quarter car. The rst node is at a trivial point ( 1 = 0 1 = 1), which shows that = when the excitation frequency is zero. The fourth node is at ( 4 4 1). There are also two middle nodes at ( 2 2 1) and \u00a1 3 = 1 3 = 1 \u00a2 . Because 1 1 and 4 1, the middle nodes are important in optimization. To have a better view at the middle nodes, Figure 15.10 illustrates a magni cation of the sprung mass displacement frequency response, = \u00af\u0304 \u00af\u0304 around the middle nodes. Example 575 F There is no Frahm optimal quarter car. Reduction in absolute amplitude is the rst attempt for optimization. If the amplitude frequency response = ( ) contains xed points with respect to some parameters, then using the Frahm method, the optimization process is carried out in two steps: 1 We select the parameters that control the position of the invariant points to equalize the corresponding height at the invariant frequencies, and minimize the height of the xed points as much as possible. 2 We nd the remaining parameters such that the maximum amplitude coincides precisely at the invariant points. For a real problem, the values of mass ratio , and wheel frequency are xed and we are trying to nd the optimum values of and . The parameters and include the unknown sti ness of the main spring and the unknown damping of the main shock absorber, respectively. The amplitude at invariant frequencies , show that the rst invariant point ( 1 = 0 1 = 1) is always xed, and the fourth one ( 4 4 1) happens after the natural frequencies. Therefore, the second and third nodes are the suitable nodes for applying the above optimization steps. However, 2 1 3 1 (15.74) and hence, we cannot apply the above optimization method. It is because 2 and 3 can never be equated by varying . Ever so, we can still nd the optimum value of by evaluating based on other constraints. Example 576 Natural frequency variation. The natural frequencies 1 and 2 , as given in Equations (15.45) and (15.46), are functions of and . Figures 15.11 and 15.12 illustrate the e ect of these two parameters on the variation of the natural frequencies. The rst natural frequency 1 1 decreases by increasing the mass ratio . 1 is close to the natural frequency of a 1 8 car model and indicates the principal natural frequency of a car. Hence, it is called the body bounce natural frequency. The second natural frequency, 2 , approaches in nity when decreases. However, 2 10Hz for street cars with acceptable ride comfort. 1 relates to the unsprung mass, and is called the wheel hop natural frequency. Figure 15.13 that plots the natural frequency ratio 1 2 shows their relative behavior. Example 577 Invariant frequencies variation. The invariant frequencies 2, 3, and 4, as given in Equation (15.47), are functions of and . Figures 15.14 to 15.18 illustrate the e ect of these two parameters on the invariant frequencies. The second invariant frequency 2, as shown in Figure 15.14, is always less than 2 because lim 0 2 = 2 (15.75) So, whatever the value of the mass ratio is, 2 cannot be greater than 2. Such a behavior does not let us control the position of second node freely. The third invariant frequency 3 as shown in Figure 15.15 is not a function of the mass ratio and may have any value depending on . The fourth invariant frequency 4 is shown in Figure 15.15. 4 increases when decreases. However, 4 settles when & 0 6. lim 0 2 = (15.76) To have a better picture about the behavior of invariant frequencies, Figures 15.17 and 15.18 depict the relative frequency ratio 4 3 and 3 2. Example 578 Frequency response at invariant frequencies. The frequency response is a function of , , and . Damping always diminishes the amplitude of vibration, so at rst we set = 0 and plot the behavior of as a function of , . Figure 15.19 illustrates the behavior of at the second invariant frequency 2. Because lim 0 2 = 1 (15.77) 2 starts at one, regardless of the value of . The value of 2 is always greater than one. Figure 15.20 shows that 3 is not a function of and is a decreasing function of . Figure 15.21 shows that 4 1 regardless of the value of and . The relative behavior of 2, 3, and 4 is shown in Figures 15.22 and 15.23. Example 579 Natural frequencies and vibration isolation of a quarter car. For a modern typical passenger car, the values of natural frequencies are around 1Hz and 10Hz respectively. The former is due to the bounce of sprung mass and the latter belongs to the unsprung mass. At average speed, bumps with wavelengths much greater than the wheelbase of the vehicle, will excite bounce motion of the body. At at higher speed, wavelength of the bumps become shorter than a wheelbase length and cause heavy vibrations of the unsprung. Therefore, when the wheels hit a single bump on the road, the impulse will set the wheels into oscillation at the natural frequency of the unsprung mass around 10Hz. In turn, for the sprung mass, the excitation will be the frequency of vibration of the unsprung around 10Hz. Because the natural frequency of the sprung is approximately 1Hz, the excellent isolation for sprung mass occurs and the frequency range around 10Hz has no essential in uence on the sprung discomfort. When the wheel runs over a rough undulating surface, the excitation will consists of a wide range of frequencies. Again, high excitation frequency at 5Hz to 20Hz means high frequency input to the sprung mass, which can e ectively be isolated. Low frequency excitation, however, will cause resonance in the sprung mass. 15.4 F RMS Optimization Figure 15.24 is a design chart for optimal suspension parameters of a base excited two-DOF system such as a quarter car model. The horizontal axis is the root mean square of relative displacement, = ( ), and the vertical axis is the root mean square of absolute acceleration, = ( ). There are two sets of curves that make a mesh. The rst set, which is almost parallel at the right end, are constant damping ratio , and the second set is constant natural frequency ratio . There is a curve, called the optimal design curve, which indicates the optimal main suspension parameters: The optimal design curve is the result of the RMS optimization strategy \u00a8 (15.78) which states that the minimum absolute acceleration with respect to the relative displacement, if there is any, makes the suspension of a quarter car optimal. Mathematically, it is equivalent to the following minimization problem: = 0 (15.79) 2 2 0 (15.80) To use the design curve and determine optimal sti ness and damping for the main suspension of the system, we start from an estimate value for on the horizontal axis and draw a vertical line to hit the optimal curve. The intersection point indicates the optimal and for the . Figure 15.25 illustrates a sample application for = 0 75, which indicates 0 3 and 0 35 for optimal suspension. Having and , determines the optimal value of and . = 2 (15.81) = 2 p (15.82) Proof. The RMS of a continues function ( ) is de ned by ( ) = s 1 2 1 Z 2 1 2 ( ) (15.83) where 2 1 is called the working frequency range. Let us consider a working range for the excitation frequency 0 \u00a1 = 2 \u00a2 20Hz to include almost all ground vehicles, especially road vehicles, and show the RMS of and by = ( ) (15.84) = ( ) (15.85) In applied vehicle dynamics, we usually measure frequencies in [ Hz], instead of [ rad s], we perform design calculations based on cyclic frequencies and in [ Hz], and we do analytic calculation based on angular frequencies and in [ rad s]. To calculate and over the working frequency range = s 1 40 Z 40 0 2 (15.86) = s 1 40 Z 40 0 2 = 2 s 1 40 Z 40 0 2 2 (15.87) we rst nd integrals of 2 and 2.Z 2 = 1 2 6 \u03bc 1 1 + 1 5 \u00b6 ln \u03bc 1 + 1 \u00b6 + 1 2 7 \u03bc 1 2 + 2 5 \u00b6 ln \u03bc 2 + 2 \u00b6 + 1 2 8 \u03bc 1 3 + 3 5 \u00b6 ln \u03bc 3 + 3 \u00b6 + 1 2 9 \u03bc 1 4 + 4 5 \u00b6 ln \u03bc 4 + 4 \u00b6 (15.88) Z 2 = 3 1 2 6 ln \u03bc 1 + 1 \u00b6 ++ 3 2 2 7 ln \u03bc 2 + 2 \u00b6 + 3 3 2 8 ln \u03bc 3 + 3 \u00b6 + 3 4 2 9 ln \u03bc 4 + 4 \u00b6 (15.89) The parameters 1 through 9 are: 1 = 1 2 19 + 23 19 1 4 15 14 (15.90) 2 = 1 2 19 23 19 1 4 15 14 (15.91) 3 = 1 2 19 + 24 19 1 4 15 14 (15.92) 4 = 1 2 19 24 19 1 4 15 14 (15.93) 5 = 4 2 (15.94) 6 = \u00a1 2 1 2 2 \u00a2 \u00a1 2 1 2 3 \u00a2 \u00a1 2 1 2 4 \u00a2 (15.95) 7 = \u00a1 2 2 2 3 \u00a2 \u00a1 2 2 2 3 \u00a2 \u00a1 2 2 2 1 \u00a2 (15.96) 8 = \u00a1 2 3 2 4 \u00a2 \u00a1 2 3 2 1 \u00a2 \u00a1 2 3 2 2 \u00a2 (15.97) 9 = \u00a1 2 4 2 1 \u00a2 \u00a1 2 4 2 2 \u00a2 \u00a1 2 4 2 3 \u00a2 (15.98) 10 = 1 6 3 p 20 + 8 13 + 2 3 2 11 3 20 + 1 3 11 (15.99) 11 = 8 16 14 3 3 15 8 3 14 (15.100) 12 = 4 16 14 15 3 15 8 2 14 17 8 3 14 (15.101) 13 = 64 2 14 17 15 + 256 3 14 18 + 16 14 2 15 16 3 4 15 256 4 14 (15.102) 14 = 4 (15.103) 15 = 2 4 (1 + ) 2 2 + 4 (1 + ) 2 4 2 (15.104) 16 = 8 2 2 (1 + ) + (1 + ) 2 4 2 2 (2 + ) + 1 (15.105) 17 = 4 2 2 2 (1 + ) 2 (15.106) 18 = 1 (15.107) 19 = 10 11 (15.108) 20 = 21 + 12 p 22 (15.109) 21 = 288 11 13 + 108 2 12 + 8 3 11 (15.110) 22 = 768 3 13 + 384 2 11 2 13 48 13 4 11 432 11 2 12 13 + 81 4 12 + 12 3 11 2 12 (15.111) 23 = 19 ( 11 10) 2 12 3 2 19 (15.112) 24 = 19 ( 11 + 10) + 2 12 3 2 19 (15.113) Now the required RMS, , and , over the frequency range 0 20Hz, can be calculated analytically from Equations (15.86) and (15.87). Equations (15.86) and (15.87) show that both and are functions of only three variables: , , and . = ( ) (15.114) = ( ) (15.115) In applied vehicle dynamics, is usually a xed parameter, so, any pair of design parameters ( ) determines and uniquely. Let us set = 3 (15.116) Using Equations (15.86) and (15.87), we may draw Figure 15.26 to illustrate how behaves with respect to when and vary. Keeping constant and varying , it is possible to minimize with respect to . The minimum points make the optimal curve and determine the best and . The way to use the optimal design curve is to estimate a value for or and nd the associated point on the design curve. A magni ed picture is shown in Figure 15.24. The horizontal axis is the root mean square of relative displacement, = ( ), and the vertical axis is the root mean square of absolute acceleration, = ( ). The optimal curve indicates that softening a suspension decreases the body acceleration, however, it requires a large room for relative displacement. Due to physical constraints, the wheel travel is limited, and hence, we must design the suspension such that to use the available suspension travel, and decrease the body acceleration as low as possible. Mathematically it is equivalent to (15.79) and (15.80). Example 580 Examination of the optimal quarter car model. To examine the optimal design curve and compare practical ways to make a suspension optimal, we assume that there is a quarter car with an o - optimal suspension, indicated by point 1 in Figure 15.27. = 3 = 0 35 = 0 4 (15.117) To optimize the suspension practically, we may keep the sti ness constant and change the damper to a corresponding optimal value, or keep the damping constant and change the sti ness to a corresponding optimal value. However, if it is possible, we may change both, sti ness and damping to a point on the optimal curve depending on the physical constraints and requirements. Point 2 in Figure 15.27 has the same as point 1 with an optimal damping ratio 0 3. Point 3 in Figure 15.27 has the same as point 1 with an optimal natural frequency ratio 0 452. Hence, points 2 and 3 are two alternative optimal designs for the o -optimal point 1. Figure 15.28 compares the acceleration frequency response log for the three points 1, 2, and 3. Point 3 has the minimum acceleration frequency response. Figure 15.29 depicts the absolute displacement frequency response log and Figure 15.30 compares the relative displacement frequency response log for the there points 1, 2, 3. These gures show that both points 2 and 3 introduce better suspension than point 1. Suspension 2 has a higher level of acceleration but needs less relative suspension travel than suspension 3. Suspension 3 has a lower level of acceleration, but it needs more room for suspension travel than suspension 2. Example 581 Comparison of an o -optimal quarter car with two optimals. An alternative method to optimize an o -optimal suspension is to keep the RMS of relative displacement or absolute acceleration constant and nd the associated point on the optimal design curve. Figure 15.31 illustrates two alternative optimal designs, points 2 and 3, for an o - optimal design at point 1. The mass ratio is assumed to be = 3 (15.118) and the suspension characteristics at 1 are = 0 0465 = 0 265 = 2 = 0 15 (15.119) The optimal point corresponding to 1 with the same is at 2 with the characteristics = 0 23 = 0 45 = 0 543 = 0 15 (15.120) and the optimal point with the same as point 1 is a point at 2 with the characteristics: = 0 0949 = 0 1858 = 2 = 0 0982 (15.121) Figure 15.32 depicts the sprung mass vibration amplitude , which shows that both points 2 and 3 have lower overall amplitude specially at second resonance. Figure 15.33 shows the amplitude of relative displacement between sprung and unsprung masses. The amplitude of absolute acceleration of the sprung mass is shown in Figure 15.34. Example 582 F Natural frequencies and vibration isolation requirements. Road irregularities are the most common source of excitation for passenger cars. Therefore, the natural frequencies of vehicle system are the primary factors in determining design requirements for conventional isolators. The natural frequency of the vehicle body supported by the primary suspension is usually between 0 2Hz and 2Hz, and the natural frequency of the unsprung mass, called wheel hop frequency, usually is between 2Hz and 20Hz. The higher values generally apply to military vehicles. The isolation of sprung mass from the uneven road can be improved by using a soft spring, which reduces the primary natural frequency. Lowering the natural frequency always improves the ride comfort, however it causes a design problem due to the large relative motion between the sprung and unsprung masses. One of the most important constraints that suspension system designers have to consider is the rattle-space constraint, the maximum allowable relative displacement. Additional factors are imposed by the overall stability, reliability, and economic or cost factors. Example 583 Optimal characteristics variation. We may collect the optimal and and plot them as shown in Figures 15.35 and 15.36. These gures illustrate the trend of their variation. The optimal value of both and are decreasing functions of relative displacement RMS . So, when more room is available, we may reduce and and have a softer suspension for better ride comfort. Figure 15.37 shows how the optimal and change with each other. 15.5 F Optimization Based on Natural Frequency and Wheel Travel Assume a xed value for the mass ratio and natural frequency ratio to x the position of the nodes in the frequency response plot. Then, an optimal value for damping ratio is F = 35 36 sq 2 37 8 2 + 37 8 2 35 (15.122) where 35 = 2 (1 + ) + 1 (15.123) 36 = 4 1 + (15.124) 37 = \u00a1 2 2 (1 + ) + 1 \u00a2 (15.125) The optimal damping ratio F causes the second resonant amplitude 2 to occur at the second invariant frequency 2. The value of relative displacement at = 2 for = F is, 2 = vuuut \u00b3p 2 37 8 2 35 \u00b4 1 + 2 2 \u00b3 28 p 2 37 29 \u00b4 (15.126) where, 28 = 4 4 (1 + ) 4 4 2 (1 + ) 2 (1 ) + \u00a1 1 + 2 \u00a2 (15.127) 29 = 8 6 (1 + ) 5 + 12 4 (1 + ) 3 (1 ) 2 4 (1 + ) \u00a1 1 + 3 2 2 \u00a2 + \u00a1 1 + 2 \u00a2 (15.128) Proof. Natural frequencies of the sprung and unsprung masses, as given in Equations (15.45) and (15.46), are related to and . When is given, we can evaluate by considering the maximum permissible static de ection, which in turn adjusts the value of natural frequencies. If the values of and are determined and kept xed, then the value of damping ratio which cause the rst resonant amplitude to occur at the second node, can be determined as optimum damping. For a damping ratio less or greater than the optimum, the resonant amplitude would be greater. The frequencies related to the maximum of are obtained by di erentiating with respect to and setting the result equal to zero = 1 2 2 = 1 2 25 \u00a1 8 2 25 26 27 \u00a2 = 0 (15.129) where 25 = \u00a3 2 \u00a1 2 2 1 \u00a2 + \u00a1 1 (1 + ) 2 2 \u00a2\u00a42 +4 2 2 \u00a1 1 (1 + ) 2 2 \u00a22 (15.130) 26 = 8 2 \u00a1 4 2 2 + 1 \u00a2 \u00a1 3 2 2 (1 + ) 1 \u00a2 \u00d7 \u00a1 2 2 (1 + ) 1 \u00a2 (15.131) 27 = 4 \u00a1 4 2 2 + 1 \u00a2 \u00a3 2 2 (1 + ) + 2 \u00a1 1 2 2 \u00a2 1 \u00a4 \u00d7 \u00a3 2 (1 + ) 2 2 2 + 1 \u00a4 (15.132) Now, the optimal value F in Equation (15.122) is obtained if the frequency ratio in Equation (15.129) is replaced with 2 given by Equation (15.61). The optimal damping ratio F makes have a maximum at the second invariant frequency 2. Figure 15.38 illustrates an example of frequency response for di erent including = F. Figure 15.39 shows the sensitivity of F to and . Substituting F in the general expression of , the absolute maximum value of would be equal to 2 given by equation (15.51). Substituting = 2 and = F in Equation (15.25) gives us Equation (15.126) for 2. The lower the natural frequency of the suspension, the more e ective the isolation from road irregularities. So, the sti ness of the main spring must be as low as possible. Figure 15.40 shows the behavior of 2 for = F. Example 584 Nodes in 2 for = F. The relative displacement at second node, 2, is a monotonically increasing function of and has two invariant points. The invariant points of may be found from \u00b1 2 \u00a3\u00a1 2 2 1 \u00a2 + \u00a1 1 (1 + ) 2 \u00a2\u00a4 = \u00a3 2 \u00a1 2 2 1 \u00a2 + \u00a1 1 (1 + ) 2 2 \u00a2\u00a42 + 2 \u00a1 1 (1 + ) 2 2 \u00a2 (15.133) that are, 1 = 0 1 = 0 (15.134) 0 = 1 1 + 0 = 1 + 1 (15.135) The value of at 0 is, 0 = 2 2 (1 + ) 3 \u00a3 4 2 + 2 (1 + ) \u00a4 (15.136) Example 585 F Maximum value of . Figure 15.4 shows that has a node at the intersection of the curves for = 0 and = . There might be a speci c damping ratio to make have a maximum at the node. To nd the maximum value of , we have to solve the following equation for : = 1 2 2 = 1 2 23 \u00a1 4 3 23 30 31 \u00a2 (15.137) where 23 = \u00a3 2 \u00a1 2 2 1 \u00a2 + \u00a1 1 (1 + ) 2 2 \u00a2\u00a42 +4 2 2 \u00a1 1 (1 + ) 2 2 \u00a22 (15.138) 30 = 8 2 5 \u00a3 3 2 2 (1 + ) 1 \u00a4 \u00a1 2 2 (1 + ) 1 \u00a2 (15.139) 31 = 4 5 \u00a3 2 2 (1 + ) + 2 \u00a1 1 2 2 \u00a2 1 \u00a4 \u00d7 \u00a3 2 (1 + ) 2 2 2 + 1 \u00a4 (15.140) Therefore, the maximum occurs at the roots of the equation: 32 8 + 33 6 + 34 2 1 = 0 (15.141) where 32 = 4 (15.142) 33 = 2 4 2 (1 + ) 2 + 4 (1 + ) 2 (15.143) 34 = 2 (1 + ) + 1 2 2 (15.144) Equation (15.141) has two positive roots when is less than a speci c value of damping ratio, , and one positive root when is greater than , where, = ( ) (15.145) The positive roots of Equation (15.141) are 5 and 6, and the corresponding relative displacements are denoted by 5 and 6, where 5 6. The invariant frequencies 5 and 6 would be equal when , and they approach 0 when goes to in nity. The invariant frequency 6 is greater than 5 as long as , and they are equal when . The relative displacements 5 and 6 are monotonically decreasing functions of and they approach 0 when goes to in nity. It is seen from (15.135) that the invariant point at 0 depends on and but the value of 0 depends only on . If is given, then 0 is xed. Therefore, the maximum value of the relative displacement, , cannot be less than 0 and we cannot nd any real value for that causes the maximum of to occur at 0 . The optimum value of could be found when we adjust the maximum value of 6, to be equal to the allowed wheel travel. 15.6 Summary The vertical vibration of vehicles may be modeled by a two-DOF linear system called quarter car model. One-fourth of the body mass, known as sprung mass, is suspended by the main suspension of the vehicle and . The main suspension and are mounted on a wheel of the vehicle, known as unsprung mass. The wheel is sitting on the road by a tire with sti ness . Assuming the vehicle is running on a harmonically bumped road we are able to nd the frequency responses of the sprung and unsprung masses, and relative displacement can be found analytically by taking advantage of the linearity of the system. The frequency response of the sprung mass has four nodes. The rst and fourth nodes are usually out of resonance or out of working frequency range. The middle nodes sit at di erent sides of = 1, and therefore, they cannot be equated and Frahm optimization cannot be applied. The root mean square of the absolute acceleration and relative displacement can be found analytically by applying the RMS optimization method. The RMS optimization method is based on minimizing the absolute acceleration RMS with respect to the relative displacement RMS. The result of RMS optimization introduces an optimal design curve for a xed mass ratio. 15.7 Key Symbols \u00a8 acceleration damping main suspension damper [ ] damping matrix 1 road wave length 2 road wave amplitude dissipation function F force = 1 cyclic frequency [ Hz] damper force spring force cyclic natural frequency [ Hz]\u00a1 2 \u00a2 characteristic equation sti ness main suspension spring sti ness tire sti ness equivalent sti ness [ ] sti ness matrix kinetic energy L Lagrangean mass sprung mass unsprung mass [ ] mass matrix = excitation frequency ratio nodal frequency ratio = natural frequency ratio = ( ) RMS of = ( ) RMS of time period = 2 2 sprung mass acceleration frequency response = 2 2 unsprung mass acceleration frequency response potential energy absolute displacement sprung mass displacement unsprung mass displacement steady-state amplitude of steady-state amplitude of steady-state amplitude of base excitation displacement steady-state amplitude of relative displacement steady-state amplitude of short notation parameter = sprung mass ratio = sprung mass ratio = | | sprung mass relative frequency response = | | sprung mass frequency response = \u00a1 2 \u00a2 damping ratio F optimal damping ratio = | | unsprung mass frequency response = 2 angular frequency [ rad s] = p sprung mass frequency = p unsprung mass frequency natural frequency Subscript node number natural sprung unsprung Exercises 1. Quarter car natural frequencies. Determine the natural frequencies of a quarter car with the following characteristics: = 275kg = 45 kg = 200000N m = 10000N m 2. Equations of motion. Derive the equations of motion for the quarter car model that is shown in Figure 15.1, using the relative coordinates: (a) = = (b) = = (c) = = 3. F Natural frequencies for di erent coordinates. Determine and compare the natural frequencies of the three cases in Exercise 2 and check their equality by employing the numerical data of Exercise 1. 4. Quarter car nodal frequencies. Determine the nodal frequencies of a quarter car with the following characteristics: = 275kg = 45 kg = 200000N m = 10000N m Check the order of the nodal frequencies with the natural frequencies found in Exercise 1. 5. Frequency responses of a quarter car. A car is moving on a wavy road with a wave length 1 = 20m and wave amplitude 2 = 0 08m. = 200 kg = 40 kg = 220000N m = 8000N m = 1000N s m Determine the steady-state amplitude , , and if the car is moving at: (a) = 30km h (b) = 60km h (c) = 120km h 6. Quarter car suspension optimization. Consider a car with = 200 kg = 40 kg = 220000N m = 0 75 and determine the optimal suspension parameters. 7. A quarter car has = 0 45 and = 0 4. What is the required wheel travel if the road excitation has an amplitude = 1cm? 8. F Quarter car and time response. Find the optimal suspension of a quarter car with the following characteristics: = 220 kg = 42 kg = 150000N m = 0 75 and determine the response of the optimal quarter car to a unit step excitation. 9. F Quarter car mathematical model. In the mathematical model of the quarter car, we assumed the tire is always sticking to the road. Determine the condition at which the tire leaves the surface of the road. 10. Optimal damping. Consider a quarter car with = 0 45 and = 0 4. Determine the optimal damping ratio F." + ] + }, + { + "image_filename": "designv11_35_0001419_978-94-007-7401-8-Figure2.20-1.png", + "original_path": "designv11-35/openalex_figure/designv11_35_0001419_978-94-007-7401-8-Figure2.20-1.png", + "caption": "Fig. 2.20 Fully-parallel PMs with coupled Sch\u00f6nflies motions of types 4RRCR (a) and 4RCRR (b) defined by MF \u00bc SF \u00bc 4; RF\u00f0 \u00de \u00bc \u00f0v1; v2; v3;xd\u00de, TF = 0, NF = 2, limb topology R\\R||C||R (a) and R||C\\R||R (b)", + "texts": [ + "16) P\\R\\R\\P\\kR (Fig. 2.1a0) Idem No. 5 30. 4CRRR (Fig. 2.17a) C\\R||R\\R (Fig. 2.1d0) Idem No. 1 (continued) 2.1 Topologies with Simple Limbs 59 Table 2.1 (continued) No. PM type Limb topology Connecting conditions 31. 4CRRR (Fig. 2.17b) C||R||R\\R (Fig. 2.1e0) Idem No. 13 32. 4RRCR (Fig. 2.18a) R||R||C\\R (Fig. 2.1f0) Idem No. 13 33. 4RCRR (Fig. 2.18b) R||C||R\\R (Fig. 2.1g0) Idem No. 13 34. 4RRRC (Fig. 2.19a) R\\R||R||C (Fig. 2.1h0) Idem No. 21 35. 4RCRR (Fig. 2.19b) R\\C||R||R (Fig. 2.1l0) Idem No. 21 36. 4RRCR (Fig. 2.20a) R\\R||C||R (Fig. 2.1b0) Idem No. 21 37. 4RCRR (Fig. 2.20b) R||C\\R||R (Fig. 2.1c0) Idem No. 21 60 2 Fully-Parallel Topologies with Coupled Sch\u00f6nflies Motions In the fully-parallel topologies of PMs with coupled Sch\u00f6nflies motions F / G1G2-G3-G4 presented in this section, the moving platform n : nGi (i = 1, 2, 3, 4) is connected to the reference platform 1 : 1Gi : 0 by four spatial complex limbs with four or five degrees of connectivity. The complex limbs combine only revolute, prismatic and cylindrical joints. One actuator is combined in each limb. The actuated joint is underlined in the structural graph" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_35_0000108_iros45743.2020.9340689-Figure4-1.png", + "original_path": "designv11-35/openalex_figure/designv11_35_0000108_iros45743.2020.9340689-Figure4-1.png", + "caption": "Fig. 4: Capability of the robot arm in the grasping region. The colors correspond to the HSV scale, with blue and red being the highest and lowest local dexterity, respectively.", + "texts": [ + " For a top grasp, for instance, the grasp node is added only if it is reachable by the robot. The reachability is evaluated using a capability map [15], which is computed offline for a given robot kinematics. For a given hand pose, a query of the reachability map indicates if the pose is reachable or not. If the top grasp node is added, it also stores the capability index associated with the grasp pose; this index is an indirect measure of local dexterity. The capability of the robot arm in the bin region for our evaluation setup is shown in Fig. 4, and it clearly indicates that some regions of the bin have better dexterity compared to others. More details on reachability and capability analysis can be found in [15]. The top grasp node also stores information related to how cluttered the environment is around the given object; for this, the MPV magnitudes of all the neighbors are added and this total MPV score is also saved as a property of the grasp node. Other properties stored for all grasp nodes include the required grasp pose, pre-grasp pose, vision score, and object fragility index" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_35_0000125_s00170-021-07485-6-Figure1-1.png", + "original_path": "designv11-35/openalex_figure/designv11_35_0000125_s00170-021-07485-6-Figure1-1.png", + "caption": "Fig. 1 The model for power skiving system", + "texts": [ + " Aiming to investigate the dynamic machining process for power skiving by overcoming the shortcoming of [26], this work developed a discrete enveloping-based mathematic modelling method for power skiving, which is the effect of the general profile tasks with strong robustness. The remainder is organized as follows. The basic mathematic model for skiving is introduced in Section 2. The numerical simulation method for power skiving is studied in Section 3. Then in Section 4, the deviation estimation is described. In Section 5, several skiving tasks are simulated and concluded at last. In a general power skiving system, the cutter and workpiece is set up as a pair of cross-axis in Fig. 1. Workpiece coordinate system Sw: Ow-XwYwZw is attached to the workpiece, and its Zw-axis is coincided with the cylindrical workpiece axis. Similarly, cutter coordinate system Sc: Oc-XcYcZc is attached to the cutter, and its Zc-axis is coincided with the cutter axis. In initial, we define that Xw-axis is going through the origin Oc and Xc-axis is coinciding with Xw-axis. Meanwhile, with respect to the meshing of cutter and workpiece, the shaft angle \u03a9 between their axes and the nearest distance E0 between their axes are given as follows: \u03a9 \u00bc \u2212 jw\u03b2w\u2212kio jc\u03b2c \u00f01\u00de E0 \u00bc Rw \u00fe kioRc \u00f02\u00de where jw denotes the helix direction of workpice (jw = 0 indicates spur, jw = 1 indicates right hand, and jw = \u2212 1 indicates left hand), and jc denotes the helix direction of cutter (jc = 0 indicates spur, jc = 1 indicates right hand, and jc = \u2212 1 indicates left hand); \u03b2w and \u03b2c are the helix angles on the pitch circles, respectively, for the workpiece and the cutter; and kio denotes the type of skiving (kio = 1 denotes external skiving, and kio = \u2212 1 denotes internal skiving). Moreover, in practical applications, the cutter takes an eccentricity Ec along the Zc-axis from the origin Oc, looking to improve the working condition of cutting edges and to avoid the interferences during the machining process. Power skiving consists of two kinds of coupled motion as illustrated in Fig. 1, i.e., the meshing motion and the feeding motion. The cutter and workpiece rotate \u03c6c and \u03c6w around Zcaxis and Zw-axis, respectively, with constant transmission ratio, which is termed as main meshing motion, contributing to material cut off. Meanwhile, a differential rotation \u0394\u03c6c is implemented on the cutter to ensure the meshing motion when the cutter performs a synchronous linear feed motion f along Zw-axis, which contributes to produce the complete slot. The rotation angles of the power skiving system are given as below: \u03c6c \u00bc \u2212kio\u03c6w zw zc \u00fe\u0394\u03c6c \u00f03\u00de where zw and zc are the slot number of workpiece and teeth number of cutter, respectively;\u0394\u03c6c is the differential rotation angle of the skiving cutter, and it is determined by both the axial feeding \u0394f of the skiving cutter along Zw-axis and the configuration of power skiving system as follows: \u0394\u03c6c \u00bc kio jw \u0394f sin \u03b2w\u00f0 \u00de Rccos \u03b2c\u00f0 \u00de \u00f04\u00de Aiming to satisfy the diverse requirements in practical applications of skiving, such as tooth profile crown correction and machining error sensitivity analysis, we involve an additional feed motion Ke to enhance this virtual kinematic model of power skiving as in [38], which expresses the constant radial distance E0 between the cutter and the part" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_35_0001419_978-94-007-7401-8-Figure5.110-1.png", + "original_path": "designv11-35/openalex_figure/designv11_35_0001419_978-94-007-7401-8-Figure5.110-1.png", + "caption": "Fig. 5.110 1PPn3-2PPn3R-1PPPn3R type redundantly actuated PM with uncoupled Sch\u00f6nflies motions defined by MF = 5, SF = 4, RF\u00f0 \u00de \u00bc \u00f0v1; v2; v3;xa\u00de, TF = 1, NF = 15, limb topology P||Pn3, P||Pn3\\R, P||P||Pn3\\R", + "texts": [ + " 1PRRbR-2PRRbRR-1PPRRbRR (Fig. 5.106) P||R||Rb||R (Fig. 4.1i) Idem No. 1 P||R||Rb||R\\R (Fig. 4.9c) P||P||R||Rb||R\\R (Fig. 5.99f) 7. 1PPn2R-2PPn2RR-1PPPn2RR (Fig. 5.107) P||Pn2||R (Fig. 4.8j) Idem No. 1 P||Pn2||R\\R (Fig. 4.9d) P||P||Pn2||R\\R (Fig. 5.99g) 8. 1PPn2R-2PPn2RR-1PPPn2RR (Fig. 5.108) P||Pn2||R (Fig. 4.8k) Idem No. 1 P||Pn2||R\\R (Fig. 4.9e) P||P||Pn2||R\\R (Fig. 5.99h) 9. 1PPn3-2PPn3R-1PPPn3R (Fig. 5.109) P||Pn3 (Fig. 4.8l) Idem No. 1 P||Pn3\\R (Fig. 4.9f) P||P||Pn3\\R (Fig. 5.99i) 10. 1PPn3-2PPn3R-1PPPn3R (Fig. 5.110) P||Pn3 (Fig. 4.8m) Idem No. 1 P||Pn3\\R (Fig. 4.9g) P||P||Pn3\\R (Fig. 5.99j) 11. 1PPaPR-2PPaC-1PPPaC (Fig. 5.111) P||Pa\\P\\kR (Fig. 4.8e) Idem No. 1 P||Pa\\C (Fig. 4.8o) P||P||Pa\\C (Fig. 5.99k) 12. 1CRbR-2CRbRR-1PCRbRR (Fig. 5.112) C||Rb||R (Fig. 4.8n) Idem No. 1 C||Rb||R\\R (Fig. 4.9 h) P||C||Rb||R\\R (Fig. 5.99l) 13. 3PPaPaR-1PPPaPaR (Fig. 5.113) P||Pa||Pa||R (Fig. 4.10b) Idem No. 1 P||Pa||Pa\\R (Fig. Fig. 4.10a) P||P||Pa||Pa\\R (Fig. 5.100a) 14. 1PRRbRbR-2PRRbRbRR1PPRRbRbRR (Fig. 5.114) P||R||Rb||Rb||R (Fig" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_35_0001419_978-94-007-7401-8-Figure3.112-1.png", + "original_path": "designv11-35/openalex_figure/designv11_35_0001419_978-94-007-7401-8-Figure3.112-1.png", + "caption": "Fig. 3.112 4PaRRRR-type overactuated PM with coupled Sch\u00f6nflies motions defined by MF \u00bc SF \u00bc 4; RF\u00f0 \u00de \u00bc v1; v2; v3;xd\u00f0 \u00de, TF = 0, NF = 14, limb topology Pa||R||R\\R||R", + "texts": [ + "51q) The revolute joints of the parallelogram loops connecting the four limbs to the fixed base have parallel axes 18. 4CPaPa (Fig. 3.106) C||Pa||Pa (Fig. 3.51r) The cylindrical joints of the four limbs have parallel axes 19. 4PaPaC (Fig. 3.107) Pa||Pa||C (Fig. 3.51s) Idem No. 18 20. 4PaCPa (Fig. 3.108) Pa||C||Pa (Fig. 3.51t) Idem No. 18 21. 4RPaRRR (Fig. 3.109) R\\Pa\\kR\\R||R (Fig. 3.52a) Idem No. 13 22. 4RRRRPa (Fig. 3.110a) R||R\\R||R||Pa (Fig. 3.52b) Idem No. 15 23. 4RRRRPa (Fig. 3.110b) R\\R||R\\R||Pa (Fig. 3.52d) Idem No. 15 24. 4PaRRRR (Fig. 3.111) Pa||R\\R||R\\kR (Fig. 3.52c) Idem No. 13 25. 4PaRRRR (Fig. 3.112) Pa||R||R\\R||R (Fig. 3.52e) Idem No. 17 26. 4RRRPaR (Fig. 3.113) R||R\\R\\Pa\\kR (Fig. 3.52f) Idem No. 12 27. 4RRRRPa (Fig. 3.114) R||R\\R||R\\kPa (Fig. 3.52g) Idem No. 12 28. 4PRRRPa (Fig. 3.115) P\\R\\R||R\\kPa (Fig. 3.52h) The second joints of the four limbs have parallel axes 29. 4RRPRPa (Fig. 3.116) R\\R||P||R\\kPa (Fig. 3.52i) Idem No. 12 30. 4RPRRPa (Fig. 3.117) R\\P||R||R\\kPa (Fig. 3.52j) Idem No. 12 31. 4RRPRPa (Fig. 3.118) R\\R||P||R\\kPa (Fig. 3.52k) Idem No. 12 32. 4RRRPPa (Fig. 3.119) R\\R||R||P\\kPa (Fig" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_35_0001419_978-94-007-7401-8-Figure2.85-1.png", + "original_path": "designv11-35/openalex_figure/designv11_35_0001419_978-94-007-7401-8-Figure2.85-1.png", + "caption": "Fig. 2.85 4PaPaRPa-type fully-parallel PM with coupled Sch\u00f6nflies motions defined by MF \u00bc SF \u00bc 4; RF\u00f0 \u00de \u00bc \u00f0v1; v2; v3;xd\u00de, TF = 0, NF = 42, limb topology Pa||Pa\\R||Pa", + "texts": [], + "surrounding_texts": [ + "2.2 Topologies with Complex Limbs 139", + "140 2 Fully-Parallel Topologies with Coupled Sch\u00f6nflies Motions", + "2.2 Topologies with Complex Limbs 141" + ] + }, + { + "image_filename": "designv11_35_0001419_978-94-007-7401-8-Figure5.20-1.png", + "original_path": "designv11-35/openalex_figure/designv11_35_0001419_978-94-007-7401-8-Figure5.20-1.png", + "caption": "Fig. 5.20 2PaPRRR-1PaPRR-1RPPaPa-type fully-parallel PMs with uncoupled Sch\u00f6nflies motions defined by MF = SF = 4, RF\u00f0 \u00de \u00bc \u00f0v1; v2; v3;xb\u00de, TF = 0, NF = 19, limb topology R||P||Pa||Pa and Pa\\P\\\\R||R\\\\R, Pa\\P\\\\R||R (a), Pa||P\\R||R\\\\R, Pa||P\\R||R (b)", + "texts": [ + "19a) Pa\\P\\R||R\\\\R (Fig. 5.3a) The last joints of the four limbs have superposed axes/directions. The revolute joints between links 5 and 6 of limbs G1, G2 and G3 have orthogonal axes Pa\\P\\\\R||R (Fig. 5.2e) R||Pa||Pa||P (Fig. 5.4k) 10. 2PaPRRR-1PaPRR1RPaPaP (Fig. 5.19b) Pa||P\\R||R\\\\R (Fig. 5.3b) Idem No. 9 Pa||P\\R||R (Fig. 5.2f) R||Pa||Pa||P (Fig. 5.4k) (continued) 528 5 Topologies with Uncoupled Sch\u00f6nflies Motions Table 5.2 (continued) No. PM type Limb topology Connecting conditions 11. 2PaPRRR-1PaPRR1RPPaPa (Fig. 5.20a) Pa\\P\\\\R||R\\\\R (Fig. 5.3a) Idem No. 9 Pa\\P\\\\R||R (Fig. 5.2e) R||P||Pa||Pa (Fig. 5.4m) 12. 2PaPRRR-1PaPRR1RPPaPa (Fig. 5.20b) Pa||P\\R||R\\\\R (Fig. 5.3b) Idem No. 9 Pa||P\\R||R (Fig. 5.2f) R||P||Pa||Pa (Fig. 5.4m) 13. 2PaPRRR-1PaPRR1RPaPatP (Fig. 5.21a) Pa\\P\\\\R||R\\\\R (Fig. 5.3a) Idem No. 9 Pa\\P\\\\R||R (Fig. 5.2e) R||Pa||Pat||P (Fig. 5.4l) 14. 2PaPRRR-1PaPRR1RPaPatP (Fig. 5.21b) Pa||P\\R||R\\\\R (Fig. 5.3b) Idem No. 9 Pa||P\\R||R (Fig. 5.2f) R||Pa||Pat||P (Fig. 5.4l) 15. 2PaPRRR-1PaPRR1RPPaPat (Fig. 5.22a) Pa\\P\\\\R||R\\\\R (Fig. 5.3a) Idem No. 9 Pa\\P\\\\R||R (Fig. 5.2e) R||P||Pa||Pat (Fig. 5.4n) 16. 2PaPRRR-1PaPRR1RPPaPat (Fig. 5.22b) Pa||P\\R||R\\\\R (Fig" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_35_0001385_978-981-15-6095-8-Figure11.1-1.png", + "original_path": "designv11-35/openalex_figure/designv11_35_0001385_978-981-15-6095-8-Figure11.1-1.png", + "caption": "Fig. 11.1 Ballasted track structure", + "texts": [ + " The settlement of the track must also be within the acceptable limits. The ballasted railway tracks employ multiple layers of unbound granular material to transfer the train-induced loads safely to the subgrade. These tracks consist of two essential components: superstructure and substructure. The superstructure comprises rails, rail pads, sleepers (or ties), and the fasteners.Moreover, the substructure constitutes ballast, subballast (capping), structural fill, general fill, and soil subgrade (prepared and natural subgrade or formation). Figure 11.1 shows a typical cross-section of the ballasted track. The rail is a longitudinal steel member which is supported by sleepers at regular intervals. It provides a firm base for the movement of trains. It must possess adequate strength and stiffness to resist the forces exerted by the rolling stock without undergoing significant deformation. The rail primarily accommodates the wheel and transfers the load from the train to the sleepers. Moreover, it may also serve as an electric signal conductor in an electrified line (Indraratna et al" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_35_0000165_s40430-021-02894-w-Figure5-1.png", + "original_path": "designv11-35/openalex_figure/designv11_35_0000165_s40430-021-02894-w-Figure5-1.png", + "caption": "Fig. 5 Sketch of the coordinate transformation from Sc to S1", + "texts": [ + " where the coordinate transformation from Sb to Sc is represented by The unit normal of the active surfaces of the cutters is represented as The unit normal of the tip relief surfaces of the cutters is represented as Similarly, the unit normal of the fillet surfaces of the cutters is represented as (12) P,i c ( 1, u1) = c,b ( 1) b,ib ib(u1) (13) P,it c ( 1, ut) = c,b ( 1) b,ib ib(ut) (14) P,if c ( 1, 1) = c,b ( 1) b,ifb ifb( 1) (15) P,o c ( 1, u1) = c,b ( 1) b,ob ob(u1) (16) P,ot c ( 1, ut) = c,b ( 1) b,ob ob(ut) (17) P,of c ( 1, 1) = c,b ( 1) b,ofb ofb( 1) (18) c,b ( 1) = \u23a1 \u23a2\u23a2\u23a2\u23a2\u23a3 cos 1 0 \u2212 sin 1 \u2212(R + m 4 ) cos 1 0 1 0 0 sin 1 0 cos 1 \u2212(R + m 4 ) sin 1 0 0 0 1 \u23a4\u23a5\u23a5\u23a5\u23a5\u23a6 (19) P,s c = P,s c 1 \u00d7 P,s c u1 ||| P,s c 1 \u00d7 P,s c u1 ||| (20) P,st c = P,st c 1 \u00d7 P,st c ut ||| P,st c 1 \u00d7 P,st c ut ||| (21) P,sf c = P,sf c 1 \u00d7 P,sf c 1 |||| P,sf c 1 \u00d7 P,sf c 1 ||||Cutter pitch plane B yb xbzb ob yc xc zc k oc R\u00b1\u03c0m/4 \u03b81 B/2 \u03c9c Fig. 4 Relations between the coordinate systems Sb and Sc Journal of the Brazilian Society of Mechanical Sciences and Engineering (2021) 43:183 1 3 183 Page 6 of 18 Figure\u00a05 shows the relationship among the coordinate systems Sc(xc, yc, zc), Sq(xq, yq, zq), Sp(xp, yp, zp) and S1(x1, y1, z1) for the generation of the gears. The coordinate systems Sq and Sp denote the reference coordinate systems, and the coordinate systems Sc and S1 are rigidly connected to the rotational centres of the cutter and gear blank, respectively. When the pitch cylinder of the gear relative to the pitch plane of the cutter is pure rolling, the cutter pitch plane remains tangent to the gear pitch cylinder at point I, which is the instantaneous centre of rotation" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_35_0001553_978-3-319-14705-5_14-Figure5-1.png", + "original_path": "designv11-35/openalex_figure/designv11_35_0001553_978-3-319-14705-5_14-Figure5-1.png", + "caption": "Fig. 5 Linearization of the robot model. a Strict motel. b Linearized model", + "texts": [ + " mxu is the position vector of the shoulder with respect to the neck frame. mxt and mxw are the position vectors of the neck and the waist with respect to the moving coordinate frame \u2211 m , respectively. In case the moments are not generated by the fictitious forces, the moment M can be divided as three moment components such as pitch, roll and yaw moments. We assume that both the waist and the trunk particles do not move vertically (m z\u0308w = 0, m z\u0308q = 0), and the trunk arm rotates on only the horizontal plane as shown in Fig. 5. We put the terms relating to the motion of the upper-body particles on the left-hand side as unknown variables, and the terms relating to the moments generated by the lower-limb particles on the right-hand side as known parameters. The decoupled and linearized ZMP equations can be obtained as follows: M\u0302yt + M\u0302yw = M\u0302y(t) M\u0302xt + M\u0302xw = M\u0302x (t) mtl2\u03b8\u0308t = M\u0302z(t) \u23ab \u23ac \u23ad (4) where M\u0302yt = mt ( m zt \u2212 m zzmp ) m x\u0308t \u2212 mt gz m xt M\u0302yw = mw ( m zw \u2212 m zzmp ) m x\u0308w \u2212 mwgz m xw M\u0302y(t) = \u2212My \u2212 ( mt ( m zt \u2212 m zzmp ) x\u0308q + mt gz m xzmp +mw ( m zw \u2212 m zzmp ) x\u0308q + mwgz m xzmp ) M\u0302xt = \u2212mt ( m zt \u2212 m zzmp ) m y\u0308t + mt gz m yt M\u0302xw = \u2212mw ( m zw \u2212 m zzmp ) m y\u0308w + mwgz m yw M\u0302x (t) = \u2212Mx \u2212 (\u2212mt ( m zt \u2212 m zzmp ) y\u0308q \u2212 mt gz m yzmp \u2212mw ( m zw \u2212 m zzmp ) y\u0308q \u2212 mwgz m yzmp ) M\u0302z(t) = \u2212Mz \u2212 ( mt ( m xt \u2212 m xzmp ) ( m y\u0308t + y\u0308q ) \u2212 mt ( m yt \u2212 m yzmp ) ( m x\u0308t + x\u0308q ) +mw ( m xw \u2212 m xzmp ) ( m y\u0308w + y\u0308q ) \u2212 mw ( m yw \u2212 m yzmp ) ( m x\u0308w + x\u0308q ) ) \u23ab \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\u23ad (5) where l is the length between the neck and the shoulder" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_35_0000775_j.engfailanal.2021.105672-Figure3-1.png", + "original_path": "designv11-35/openalex_figure/designv11_35_0000775_j.engfailanal.2021.105672-Figure3-1.png", + "caption": "Fig. 3. Fractured aft cross tube and a cross-section of fractured surface.", + "texts": [ + " The left part of the fractured cross tube (viewed in the direction of the flight) felt to the concrete substrate. The right part of the rebound D. Rakovic\u0301 et al. Engineering Failure Analysis 129 (2021) 105672 remained under the fuselage in the tunnel section for the tubes attached to the fuselage by coupling. The post-accidental visual inspection of the failed landing gear revealed that the aft cross tube was fractured into two pieces. The broken, separated part was removed and taken for examination. Fig. 3 shows fractured aft cross tube and fracture surface. Visual analysis of fractured surface performed on macroscopic level, Fig. 3., revealed that the fracture was initiated at the bottom, position 6o\u2019clock, on the inner surface of the tube. The characteristic beach marks are observed indicating fatigue fracture (areas A in the Fig. 4 and Fig. 5), where crack propagated up to a half of the wall thickness. This region is clearly separate from rougher region with V-shape chevron marks (area B in the Fig. 4). Tracing back the observed chevron marks, the fracture origin area was also indicated to be in the sector A of fracture surface" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_35_0001419_978-94-007-7401-8-Figure4.14-1.png", + "original_path": "designv11-35/openalex_figure/designv11_35_0001419_978-94-007-7401-8-Figure4.14-1.png", + "caption": "Fig. 4.14 4PPPaR-type fully-parallel PMs with decoupled Sch\u00f6nflies motions defined by MF = SF = 4, RF\u00f0 \u00de \u00bc v1; v2; v3;xa\u00f0 \u00de, TF = 0, NF = 18, limb topology P ? P ?|Pa||R and P ? P ?||Pa ?||R (a), P ? P ?||Pa ?? R (b)", + "texts": [ + "2 Topologies with Complex Limbs 421 422 4 Fully-Parallel Topologies with Decoupled Sch\u00f6nflies Motions 4.2 Topologies with Complex Limbs 423 424 4 Fully-Parallel Topologies with Decoupled Sch\u00f6nflies Motions Table 4.4 Limb topology and connecting conditions of the fully-parallel solutions with no idle mobilities presented in Figs. 4.14, 4.15, 4.16, 4.17, 4.18, 4.19, 4.20, 4.21, 4.22, 4.23, 4.24, 4.25, 4.26, 4.27, 4.28, 4.29, 4.30, 4.31, 4.32, 4.33, 4.34, 4.35, 4.36, 4.37, 4.38, 4.39, 4.40, 4.41, 4.42 No. PM type Limb topology Connecting conditions 1. 4PPPaR (Fig. 4.14a) P ?? P ? Pa||R (Fig. 4.8c) P ? P ?|Pa ?? R (Fig. 4.8b) The last revolute joints of the four limbs have parallel axes. The actuated prismatic joints of limbs G1, G2 and G3 have orthogonal directions. The actuated prismatic joints of limbs G3 and G4 have parallel directions 2. 4PPPaR (Fig. 4.14b) P ? P ??Pa||R (Fig. 4.8c) P ? P ??Pa ?? R (Fig. 4.8a) Idem No. 1 3. 4PPaPR (Fig. 4.15a) P||Pa ? P ??R (Fig. 4.8e) P||Pa ?P ?? R (Fig. 4.8d) Idem No. 1 4. 4PPaPR (Fig. 4.15b) P||Pa ? P ??R (Fig. 4.8e) P||Pa ? P||R (Fig. 4.8f) Idem No. 1 5. 1PPaRR-3PPaRRR (Fig. 4.16) P||Pa||R||R (Fig. 4.8g) P||Pa||R||R ? R (Fig. 4.9a) Idem No. 1 6. 1PaRPR\u20133PaRPRR (Fig. 4.17) Pa ? R ? P ??R (Fig. 4.8h) Pa ? R ? P ??R ? R (Fig. 4.9b) The last revolute joints of the four limbs have parallel axes. The revolute joints between links 4 and 5 of limbs G1, G2 and G3 hvave orthogonal axes" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_35_0000213_j.triboint.2021.106993-Figure14-1.png", + "original_path": "designv11-35/openalex_figure/designv11_35_0000213_j.triboint.2021.106993-Figure14-1.png", + "caption": "Fig. 14. Shear Stress Contour by CFD simulation (10,000 N load, 9000 rpm); (a) Smooth TPJB, (b) Pocketed TPJB.", + "texts": [ + " Further simulations have demonstrated that a high level of cavitation occurs for a wide range of operating conditions. This is good since it suggests that the proposed novel features for the TPJB may be beneficial for a wide range of applications. In the pockets, the evaporation significantly lessens the dynamic viscosity (\u03bcf ) of the fluid, and the larger film thickness decreases the velocity gradient (\u2202ucir/\u2202r). Therefore, the pockets reduce the shear stress (\u03bcf \u2202ucir/\u2202r). A substantial shear stress reduction is confirmed in Fig. 14, which implies a corresponding significant drop in drag torque and power loss by Eqs. (18) and (19), respectively. Fig. 14(a) shows a significant level of shear stress on the upper pads of the conventional TPJB despite the low pad loading. So, reducing the power loss from the upper pads can yield a significant overall power savings for the TPJB. Fig. 15 shows the temperature distribution on the pad and journal at 9000 rpm and 10,000 N load, for both the smooth and pocketed bearings. Figs. 15(a-1) shows the temperature varying in the axial direction due to the three nozzle oil injection cooling flows. The pad surface temperature increases along the rotation direction, and the bottom pads have relatively higher temperatures because of the larger viscous heat generation in the thinner film" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_35_0001419_978-94-007-7401-8-Figure3.103-1.png", + "original_path": "designv11-35/openalex_figure/designv11_35_0001419_978-94-007-7401-8-Figure3.103-1.png", + "caption": "Fig. 3.103 4PaRRPa-type overactuated PM with coupled Sch\u00f6nflies motions defined by MF \u00bc SF \u00bc 4; RF\u00f0 \u00de \u00bc v1; v2; v3;xd\u00f0 \u00de, TF = 0, NF = 30, limb topology Pa\\R||R||Pa", + "texts": [ + " PM type Limb topology Connecting conditions 9. 4PaPaRP (Fig. 3.97) Pa||Pa||R||P (Fig. 3.51i) Idem No. 1 10. 4PaRPaP (Fig. 3.98) Pa||R||Pa||P (Fig. 3.51j) Idem No. 1 11. 4PaRPPa (Fig. 3.99) Pa||R||P||Pa (Fig. 3.51k) Idem No. 1 12. 4RPaPaP (Fig. 3.100) R||Pa\\Pa\\\\P (Fig. 3.51l) The first revolute joints of the four limbs have parallel axes 13. 4PaRPaR (Fig. 3.101) Pa||R\\Pa\\kR (Fig. 3.51m) The last revolute joints of the four limbs have parallel axes 14. 4PaPaRR (Fig. 3.102) Pa\\Pa||R||R (Fig. 3.51n) Idem No. 13 15. 4PaRRPa (Fig. 3.103) Pa\\R||R||Pa (Fig. 3.51o) The revolute joints of the parallelogram loops connecting the four limbs to the moving plateform have parallel axes 16. 4PaPPaR (Fig. 3.104) Pa\\P\\\\Pa||R (Fig. 3.51p) Idem No. 13 17. 4PaRPaP (Fig. 3.105) Pa||R\\Pa\\\\P (Fig. 3.51q) The revolute joints of the parallelogram loops connecting the four limbs to the fixed base have parallel axes 18. 4CPaPa (Fig. 3.106) C||Pa||Pa (Fig. 3.51r) The cylindrical joints of the four limbs have parallel axes 19. 4PaPaC (Fig. 3.107) Pa||Pa||C (Fig" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_35_0001419_978-94-007-7401-8-Figure5.58-1.png", + "original_path": "designv11-35/openalex_figure/designv11_35_0001419_978-94-007-7401-8-Figure5.58-1.png", + "caption": "Fig. 5.58 2PaPaRRR-1PaPaRR-1RPPaPat-type fully-parallel PM with uncoupled Sch\u00f6nflies motions defined by MF = SF = 4, RF\u00f0 \u00de \u00bc \u00f0v1; v2; v3;xb\u00de, TF = 0, NF = 28, limb topology Pa\\Pa||R||R\\kR, Pa\\Pa||R||R and R||P||Pa||Pat", + "texts": [ + " 2PaPaRRR-1PaPaRR1RPPaPa (Fig. 5.55) Pa\\Pa||R||R\\\\R (Fig. 5.5a) The revolute joint between links 7D and 10 of G4 limb and the last revolute joints of limbs G1, G2 and G3 have superposed axes Pa\\Pa||R||R (Fig. 5.4i) R||P||Pa||Pa (Fig. 5.4m) 22. 2PaPaRRR-1PaPaRR1RPPaPa (Fig. 5.56) Pa\\Pa||R||R\\kR (Fig. 5.5b) Idem No. 21 Pa\\Pa||R||R (Fig. 5.4i) R||P||Pa||Pa (Fig. 5.4m) 23. 2PaPaRRR-1PaPaRR1RPPaPat (Fig. 5.57) Pa\\Pa||R||R\\\\R (Fig. 5.5a) Idem No. 21 Pa\\Pa||R||R (Fig. 5.4i) R||P||Pa||Pat (Fig. 5.4n) 24. 2PaPaRRR-1PaPaRR1RPPaPat (Fig. 5.58) Pa\\Pa||R||R\\kR (Fig. 5.5b) Idem No. 21 Pa\\Pa||R||R (Fig. 5.4i) R||P||Pa||Pat (Fig. 5.4n) 25. 3PaPaPaR-1RPPP (Fig. 5.59) Pa\\Pa||Pa\\kR (Fig. 5.6a) Idem No. 1 Pa\\Pa||Pa||R (Fig. 5.6b) R\\P\\\\P\\\\P (Fig. 5.1a) 26. 3PaPaPaR-1RPPP (Fig. 5.60) Pa\\Pa||Pa\\\\R (Fig. 5.6c) Idem No. 1 Pa\\Pa||Pa||R (Fig. 5.6b) R\\P\\\\P\\\\P (Fig. 5.1a) 27. 3PaPaPaR-1RUPU (Fig. 5.61) Pa\\Pa||Pa\\kR (Fig. 5.6a) The first revolute joint of G4 limb and the last revolute joints of limbs G1, G2 and G3 have parallel axes. The last joints of of limbs G1, G2 and G3 have superposed axes Pa\\Pa||Pa||R (Fig" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_35_0000571_j.bprint.2021.e00155-Figure1-1.png", + "original_path": "designv11-35/openalex_figure/designv11_35_0000571_j.bprint.2021.e00155-Figure1-1.png", + "caption": "Fig. 1. Schematic representation of Biopen version 2. (A) Schematic representation of the handheld bioprinter. (B) Core-Shell arrangement of the extruder nozzle. (C) Picture of Core-Shell deposition. (D) Confocal image of Core-Shell arrangement where the Shell consisted of GelMa/HAMa plus LAP 0.1% in red channel and Core consisted of Gelma/HAMA in the green channel. The image was obtained from Duchi et al. [67]. (For interpretation of the references to color in this figure legend, the reader is referred to the Web version of this article.)", + "texts": [ + " Connell and the team further perfected the Biopen with a new chassis and more ergonomic design with dual bioink chambers, dual inlet T. Bharadwaj and D. Verma Bioprinting 23 (2021) e00155 extrusion nozzle, UV light source of 365 nm wavelength, and a motorized extrusion system [64]. The upgraded design was used to test the Biopen in a pre-clinical setting for addressing a full-thickness chondral defect in a large animal ovine model. Unlike the previous model of Biopen, which had a collinear dual extruder arrangement, this model had core/shell distribution, as shown in Fig. 1. This new arrangement enabled a bioink carrying stem cells to be coated with another protective layer of bioink, thus creating a core-shell arrangement. The shell of the bioink consisted of GelMa/HAMa and LAP 0.1%, while the core consisted of GelMa/HAMa only. The study proved that the device could be used in a real-time intraoperative scenario to treat cartilage defects. This experiment was conducted with proper safety and clinical efficacy. The bioprinted cartilage showed good macroscopic and microscopic properties in comparison to the bioprinted 3D scaffold" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_35_0000170_14763141.2021.1880619-Figure1-1.png", + "original_path": "designv11-35/openalex_figure/designv11_35_0000170_14763141.2021.1880619-Figure1-1.png", + "caption": "Figure 1. Angular momentum components in discus throwing.", + "texts": [ + " As a vector, the angular momentum of the thrower-discus system has three independent components representing system rotations about three orthogonal axes: (1) angular momentum of top-to-left or top-to-right rotation about a horizontal axis aligned with the midline of the throwing sector (front-back axis); (2) angular momentum of top-tofront or top-to-back rotation about a horizontal axis aligned with a foul line the throwing ring (left-right axis); and (3) angular momentum of right-to-left or left-toright rotation about a vertical axis (up-down axis) (Figure 1). These angular momentum components may affect the vacuum flight distance through their relationships with discus speeds. The system angular momentum about the up-down and left-right axes is related to the horizontal speed of the discus, while the angular momentum about the front-back and axis is related to the vertical speed of the discus (Dapena & McDonald, 1989). The angular momentum of the thrower-discus system may also affect the aerodynamic distance through their relationships with discus rotations" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_35_0001419_978-94-007-7401-8-Figure3.131-1.png", + "original_path": "designv11-35/openalex_figure/designv11_35_0001419_978-94-007-7401-8-Figure3.131-1.png", + "caption": "Fig. 3.131 4PaPRRR-type overactuated PM with coupled Sch\u00f6nflies motions defined by MF \u00bc SF \u00bc 4; RF\u00f0 \u00de \u00bc v1; v2; v3;xd\u00f0 \u00de, TF = 0, NF = 14, limb topology Pa\\P||R||R\\kR", + "texts": [ + " PM type Limb topology Connecting conditions 38. 4PRRPaR (Fig. 3.125) P||R||R\\Pa||R (Fig. 3.52r) Idem No. 13 39. 4RPRPaR (Fig. 3.126) R||P||R\\Pa||R (Fig. 3.52s) Idem No. 13 40. 4RPRPaR (Fig. 3.127) R||P||R\\Pa||R (Fig. 3.52t) Idem No. 13 41. 4RRPPaR (Fig. 3.128) R||R||P\\Pa||R (Fig. 3.52u) Idem No. 13 42. 4PPaRRR (Fig. 3.129a) P\\Pa\\kR||R||\\kR (Fig. 3.52v) Idem No. 13 43. 4PaPRRR (Fig. 3.129b) Pa\\P\\\\R||R\\\\R (Fig. 3.52w) Idem No. 13 44. 4PPaRRR (Fig. 3.130) P\\Pa\\\\R||R\\kR (Fig. 3.52x) Idem No. 13 45. 4PaPRRR (Fig. 3.131) Pa\\P||R||R\\kR (Fig. 3.52y) Idem No. 13 46. 4PaRPRR (Fig. 3.132) Pa\\R||P||R\\kR (Fig. 3.52z) Idem No. 13 47. 4PaRPRR (Fig. 3.133) Pa\\R||P||R\\kR (Fig. 3.52a0) Idem No. 13 48. 4PaRRPR (Fig. 3.134) Pa\\R||R||P\\kR (Fig. 3.52b0) Idem No. 13 49. 4RCRPa (Fig. 3.135) R\\C||R\\kPa (Fig. 3.52c0) Idem No. 12 50. 4RRCPa (Fig. 3.136) R\\R||C\\kPa (Fig. 3.52d0) Idem No. 12 51. 4CRRPa (Fig. 3.137) C||R\\R||Pa (Fig. 3.52e0) Idem No. 15 52. 4RCRPa (Fig. 3.138) R||C\\R||Pa (Fig. 3.52f0) Idem No. 15 53. 4CRPaR (Fig. 3.139) C||R\\Pa||R (Fig" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_35_0000533_17445302.2021.1938809-Figure1-1.png", + "original_path": "designv11-35/openalex_figure/designv11_35_0000533_17445302.2021.1938809-Figure1-1.png", + "caption": "Figure 1. Synchronization motion diagram of vessels during underway replenishment.", + "texts": [ + " Section 3 gives the output feedback robust adaptive neural synchronization control design and its stability analysis. Section 4 contains the simulations on a supply vessel. Section 5 contains the conclusions. 2. Problem formulation In order to realize the synchronization control between the supply and the main vessels, we introduce the concept of the virtual trajectory hr , which is shifted by a distance lm at an angle gm relative to the trajectory hm of the main vessel, with gm counter clockwise positive. As is shown in Figure 1, the coordinate XoOYo represents the north-east-down (NED) frame. The axis OXo is directed to the east, and the axis OYo is directed to the north. VM represents the main vessel and VS represents the supply vessel. XMOMYM and XSOSYS are the body-fixed frames of the main vessel and the supply vessel, respectively. The originsOM andOS are chosen at the centers of gravity of the main vessel and the supply vessel, respectively. The axes OMXM and OSXS are directed to fore, respectively. The axes OMYM and OSYS are directed to starboard, respectively. According to Figure 1, the geometric relation between the virtual trajectory and the trajectory of the main vessel is expressed as follows xr = xm + lm cos (gm + cm) (1) yr = ym + lm sin (gm + cm) (2) cr = cm (3) Equations (1)\u2013(3) can be written in the vector form as hr = hm + J(cm)L (4) where hr = [xr , yr , cr] T , hm = [xm, ym, cm] T , L = lm cos gm lm sin gm 0 \u23a1 \u23a3 \u23a4 \u23a6, and J(cm) = cos (cm) \u2212 sin (cm) 0 sin (cm) cos (cm) 0 0 0 1 \u23a1 \u23a3 \u23a4 \u23a6 called the rotation matrix. The 3-DOF nonlinear mathematical model of vessel motion can be expressed as h\u0307 = J(c)y (5) My\u0307+ C(y)y+ D(y)y = t+ d(t) (6) where h = [x, y, c]T represents the position vector in the NED frame consisting of the position (x, y) and yaw angle c, y = [u, v, r]T represents the velocity vector in the body-fixed frame consisting of velocities u in surge and v in sway, and yaw rate r, and t = [t1, t2, t3] T is the actual control input vector produced by the vessel propulsion system" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_35_0001419_978-94-007-7401-8-Figure2.100-1.png", + "original_path": "designv11-35/openalex_figure/designv11_35_0001419_978-94-007-7401-8-Figure2.100-1.png", + "caption": "Fig. 2.100 4PaRRRR-type fully-parallel PM with coupled Sch\u00f6nflies motions defined by MF \u00bc SF \u00bc 4; RF\u00f0 \u00de \u00bc \u00f0v1; v2; v3;xd\u00de, TF = 0, NF = 14, limb topology Pa\\R\\R||R\\R (a) and Pa||R||R\\R||R (b)", + "texts": [], + "surrounding_texts": [ + "154 2 Fully-Parallel Topologies with Coupled Sch\u00f6nflies Motions", + "2.2 Topologies with Complex Limbs 155", + "156 2 Fully-Parallel Topologies with Coupled Sch\u00f6nflies Motions" + ] + }, + { + "image_filename": "designv11_35_0000246_tte.2021.3068819-Figure2-1.png", + "original_path": "designv11-35/openalex_figure/designv11_35_0000246_tte.2021.3068819-Figure2-1.png", + "caption": "Fig. 2. 3D model turns into 2D model at \u03b8 = 180\u00b0, (a) with same axial length, (b) with different axial lengths.", + "texts": [ + " In addition, the physical relationship among these models need to be determined according to winding axial transposition angle \u03b8. 5- parallel-strands model can clearly reveal the effects of skin effect and proximity effect and it is convenient to observe the different cases of different winding axial transposition angles. Therefore, taking 5-parallel-strands model as an example, shown in the Fig. 1, if the strands are evenly displaced along the axial direction in a certain unit length, 3D model could be equivalent to several simple 2D models connecting in the circuit, which have same axial length, see the Fig. 2(a). The AC loss could be analyzed through these 2D models in a single simulation project. From the engineering point of view, multiple models could be analyzed by 2D-FEM software simultaneously in a single simulation project. As well as, models need to be mechanically or electrically connected between each other. These models in the same simulation project must be the same axial length. If models with different axial lengths and electrical connection, shown in Fig.2(b), it could not be solved in one simulation project. 1 2 1 1 : : 0 1 : Z Z Z Z Z t A A J x x y y A A H n + = \u2212 = = \u2212 (1) Usually, Co-simulation may be needed, such as Simulink, although it is really time consuming. The proposed method in this paper was used to equivalent these models with different axial lengths and different transposition angles to several models with the same axial length, so that they can be solved in one simulation project instead of 3D FEA and Co-simulation", + " The position error can be expressed as: ( ) 11 0 1 , in n i i n n d n n +\u2212 = = \u2212 (9) So, the maximum error is: ( ) 2 0 1 1, = 2 d = \u2212 (10) We can get other errors, just like: ( ) 11 22 0 2 1 1 2, = 2 2 8 i i i d + = = \u2212 (11) ( ) 12 23 0 3 1 1 3, = 3 3 18 i i i d + = = \u2212 (12) From (10)-(12), it is clear that the position accumulative error is small enough and will not affect the calculation accuracy when number of separated models is 3 (n=3). A. Verification and analysis of 5-parallel-strand model In this section, the 5-parallel-strands model, shown in Fig. 2 and Fig. 3, is taken as an example to verify the correctness of the equivalent permeability modeling method and to analyze the winding AC loss. For analysis and verification, this section takes n=3 as an example for analysis. Fig. 4 is a schematic diagram of an equivalent permeability modeling method. Taking \u03b8=0\u00b0 as an example, Fig. 5(a) shows the physical circuit of 2D-EMM. The axial lengths of Model 1, Model 2, and Model 3 are l1, l2, and l3, respectively. When \u03be1=\u03be2=\u03be3, the strands are evenly displaced in the axial direction; When \u03be1\u2260\u03be2\u2260\u03be3, strands are unevenly displaced in the axial direction" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_35_0001419_978-94-007-7401-8-Figure3.75-1.png", + "original_path": "designv11-35/openalex_figure/designv11_35_0001419_978-94-007-7401-8-Figure3.75-1.png", + "caption": "Fig. 3.75 Overactuated PMs with coupled Sch\u00f6nflies motions of types 4PRPaP (a) and 4PPRPa (b) defined by MF \u00bc SF \u00bc 4; RF\u00f0 \u00de \u00bc v1; v2; v3;xd\u00f0 \u00de, TF = 0, NF = 18, limb topology P\\R\\Pa\\\\P (a) and P\\P\\\\R\\Pa (b)", + "texts": [ + "69b) Pa\\P\\kP||R (Fig. 3.50z) Idem No. 1 28. 4PaPRP (Fig. 3.70) Pa\\P\\kR||P (Fig. 3.50z1) Idem No. 1 29. 4PaRRP (Fig. 3.71) Pa||R||R||P (Fig. 3.50a0) Idem No. 1 30. 4PaRPR (Fig. 3.72) Pa||R||P||R (Fig. 3.50b0) Idem No. 1 31. 4PaPRP (Fig. 3.73a) Pa||P||R\\P (Fig. 3.50c0) Idem No. 1 32. 4PaRPP (Fig. 3.73b) Pa||R\\P\\kP (Fig. 3.50d0) Idem No. 1 33. 4PRPPa (Fig. 3.74a) P\\R\\P||Pa (Fig. 3.50e0) The second joints of the four limbs have parallel axes 34. 4PRPPa (Fig. 3.74b) P\\R\\P\\Pa (Fig. 3.50f0) Idem No. 33 35. 4PRPaP (Fig. 3.75a) P\\R\\Pa\\\\P (Fig. 3.50g0) Idem No. 33 36. 4PPRPa (Fig. 3.75b) P\\P\\\\R\\Pa (Fig. 3.50h0) The third joints of the four limbs have parallel axes 37. 4PPPaR (Fig. 3.76a) P\\P||Pa\\\\R (Fig. 3.50j0) The last joints of the four limbs have parallel axes (continued) 3.2 Topologies with Complex Limbs 359 Table 3.5 (continued) No. PM type Limb topology Connecting conditions 38. 4PPPaR (Fig. 3.76b) P\\P\\kPa\\\\R (Fig. 3.50k0) Idem No. 37 39. 4PaPPR (Fig. 3.77a) Pa||P\\P\\\\R (Fig. 3.50l0) Idem No. 37 40. 4PaPPR (Fig. 3.77b) Pa\\P\\kP\\\\R (Fig. 3.50m0) Idem No. 37 41. 4PaRPR (Fig" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_35_0000141_tasc.2021.3062485-Figure1-1.png", + "original_path": "designv11-35/openalex_figure/designv11_35_0000141_tasc.2021.3062485-Figure1-1.png", + "caption": "Fig. 1. Reference cross section for the baseline motor.", + "texts": [ + " The field winding on the rotor is powered with a flux pump type brushless exciter [10] that helps reduce the thermal conduction by eliminating current leads spanning room-temperature and cryogenic environments. However, the stator winding employing current leads represents a significant thermal load for the refrigeration system. Each phase has independent current leads, i.e., 2 leads/phase\u2014an arrangement preferred for electric drives characterized with low harmonic content. Finally, no environmental EM shield is specified outside the motor as the fan duct housing could serve as the EM shield\u2014analysis assumes the fan duct is located at a radius of 560 mm. Fig. 1 shows the magnetic cross section for the baseline motor with conductor areas highlighted. Mechanical support components are not shown in the figure. The rotor and stator winding are included within radii R1 and R2 for the rotor and R3 and R4 for the stator, respectively. All windings, including the rotor EM shield, have the same active axial length. The radial built of the field winding on the rotor is 7 mm and the ac stator winding is 9\u201311 mm. Rotor EM shield thickness is 5 mm. The radial space between the inside radius of the EM shield and the outside radius of the field winding includes torque tube, and thermal insulation between the field coils and the torque tube" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_35_0000117_icccr49711.2021.9349398-Figure1-1.png", + "original_path": "designv11-35/openalex_figure/designv11_35_0000117_icccr49711.2021.9349398-Figure1-1.png", + "caption": "Figure 1. D-H coordinate of 6 DOF manipulator.", + "texts": [ + " Based on the basic RRT algorithm, an improved target probability offset algorithm and a variable step size control optimization algorithm are used to plan the collision-free path, and quasi-uniform Bspline is used to smooth the path. II. MANIPULATOR MODELING AND COLLISION DETECTION In this paper, the Denavit-Hartenberg (D-H) method [13] is used to model the SR4B 6R space robot manipulator. Joint variables and link parameter of 6 degrees of freedom (DOF) manipulator are shown in Table I. According to the Table I, we can build the D-H coordinate of the manipulator like Fig. 1 below. 104 2021 International Conference on Computer, Control and Robotics 978-1-7281-9035-8/21/$31.00 \u00a92021 IEEE 20 21 In te rn at io na l C on fe re nc e on C om pu te r, C on tro l a nd R ob ot ic s ( IC C C R ) | 9 78 -1 -7 28 1- 90 35 -8 /2 0/ $3 1. 00 \u00a9 20 21 IE EE | D O I: 10 .1 10 9/ IC C C R 49 71 1. 20 21 .9 34 93 98 Authorized licensed use limited to: BOURNEMOUTH UNIVERSITY. Downloaded on June 23,2021 at 14:41:08 UTC from IEEE Xplore. Restrictions apply. Positive kinematics refers to the process of finding the end position and posture of variables of each joint of a manipulator" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_35_0001419_978-94-007-7401-8-Figure2.7-1.png", + "original_path": "designv11-35/openalex_figure/designv11_35_0001419_978-94-007-7401-8-Figure2.7-1.png", + "caption": "Fig. 2.7 Fully-parallel PMs with coupled Sch\u00f6nflies motions of types 4RRPRR (a) and 4PRRRR (b) defined by MF \u00bc SF \u00bc 4; RF\u00f0 \u00de \u00bc \u00f0v1; v2; v3;xd\u00de, TF = 0, NF = 2, limb topology R\\R\\P\\kR\\R (a) and P\\R||R\\R||R (b)", + "texts": [ + " 2.4a) R||R\\R||R||R (Fig. 2.1d) The two first revolute joints of the four limbs have parallel axes. 6. 4PRRRR (Fig. 2.4b) P\\R\\R||R||R (Fig. 2.1e) The second joints of the four limbs have parallel axes 7. 4RRRPR (Fig. 2.5a) R||R\\R\\P\\kR (Fig. 2.1f) Idem No. 5 8. 4RRPRR (Fig. 2.5b) R||R||P\\R||R (Fig. 2.1g) Idem No. 5 9. 4PRRRR (Fig. 2.6a) P\\R\\R||R\\R (Fig. 2.1h) The second and the last joints of the four limbs have parallel axes 10. 4PRRRR (Fig. 2.6b) P||R\\R||R\\R (Fig. 2.1i) Idem No. 9 11. 4RRPRR (Fig. 2.7a) R\\R\\P\\kR\\R (Fig. 2.1j) Idem No. 1 12. 4PRRRR (Fig. 2.7b) P\\R||R\\R||R (Fig. 2.1k) Idem No. 3 13. 4PRRRR (Fig. 2.8a) P||R||R||R||R\\R (Fig. 2.1l) The last revolute joints of the four limbs have parallel axes 14. 4RPRRR (Fig. 2.8b) R||P||R||R\\R (Fig. 2.1m) Idem No. 13 15. 4RRPRR (Fig. 2.9a) R||R||P||R\\R (Fig. 2.1n) Idem No. 13 16. 4RRRPR (Fig. 2.9b) R||R||R||P\\R (Fig. 2.1o) Idem No. 13 17. 4RPRRR (Fig. 2.10a) R||P||R||R\\R (Fig. 2.1p) Idem No. 13 18. 4RRPRR (Fig. 2.10b) R||R||P||R\\R (Fig. 2.1q) Idem No. 13 19. 4RRRRP (Fig. 2.11a) R||R||R\\R\\P (Fig. 2.1r) The before last revolute joints of the four limbs have parallel axes 20" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_35_0000168_j.mechatronics.2021.102514-Figure4-1.png", + "original_path": "designv11-35/openalex_figure/designv11_35_0000168_j.mechatronics.2021.102514-Figure4-1.png", + "caption": "Fig. 4. The mechanical diagram and the definition of the generalized coordinates of the Delta manipulator. (a) The scheme of the Delta manipulator. (b) The mechanical model of the Delta manipulator. (c) The simplified mechanical model of the Delta manipulator using the property of parallelogram. Let |\ud835\udc5c\ud835\udc341| = |\ud835\udc5c\ud835\udc342| = |\ud835\udc5c\ud835\udc343| = \ud835\udc5f\ud835\udc34, |\ud835\udc5d\ud835\udc361| = |\ud835\udc5d\ud835\udc362| = |\ud835\udc5d\ud835\udc363| = \ud835\udc5f\ud835\udc35 , |\ud835\udc341\ud835\udc351| = |\ud835\udc342\ud835\udc352| = |\ud835\udc343\ud835\udc353| = \ud835\udc591, |\ud835\udc351\ud835\udc361| = |\ud835\udc352\ud835\udc362| = |\ud835\udc353\ud835\udc363| = \ud835\udc592. (d) The definition of the generalized coordinates of the Delta manipulator. Point \ud835\udc36 \u2032 \ud835\udc56 is the projection of point \ud835\udc36\ud835\udc56 in plane \ud835\udc5c\ud835\udc34\ud835\udc56\ud835\udc35\ud835\udc56. It should be noted that \ud835\udefc\ud835\udc56 is a constant which satisfies \ud835\udefc\ud835\udc56 = \ud835\udf0b\u22156 + 2\ud835\udf0b(\ud835\udc56 \u2212 1)\u22153, \ud835\udc56 = 1, 2, 3.", + "texts": [ + " Therefore, we have \ud835\udc52\ud835\udc47 \ud835\udc52 \u22642\ud835\udc49 (0)e\u2212\ud835\udefe\ud835\udc61 + 2\ud835\udf12\u2215\ud835\udefe(1 \u2212 e\u2212\ud835\udefe\ud835\udc61) + 2\ud835\udeefe\ud835\udefe\ud835\udee5\ud835\udc61 e\ud835\udefe\ud835\udee5\ud835\udc61 \u2212 1 . (40) It can be seen from Eq. (40) that \ud835\udc52\ud835\udc47 \ud835\udc52 can be upper bounded such that arbitrary small tracking error can be achieved by choosing \ud835\udefe sufficiently large and \ud835\udf12 and \ud835\udeef sufficiently small. This ends the proof of Theorem 1. A simulation example based on a Delta manipulator is given in this section. The structure of a typical Delta manipulator is shown in Fig. 3. The mechanical model of the Delta manipulator and the definition of the generalized coordinates are given in Fig. 4. The position vector of the end-effector is denoted by [\ud835\udc65\ud835\udc5d, \ud835\udc66\ud835\udc5d, \ud835\udc67\ud835\udc5d]\ud835\udc47 . To build the dynamics equation of the Delta manipulator, the PM is cut into three open-chain systems \ud835\udc5c \u2212 \ud835\udc34\ud835\udc56 \u2212 \ud835\udc35\ud835\udc56 \u2212 \ud835\udc36\ud835\udc56 where \ud835\udc56 \u2208 1, 2, 3 and a moving platform. Each open-chain system has three degrees of freedom, i.e. \ud835\udc5e\ud835\udc561, \ud835\udc5e\ud835\udc562 and \ud835\udc5e . Thus, \ud835\udc5e = [\ud835\udc5e , \ud835\udc5e , \ud835\udc5e , \ud835\udc5e , \ud835\udc5e , \ud835\udc5e , \ud835\udc5e , \ud835\udc5e , \ud835\udc5e , \ud835\udc65 , \ud835\udc66 , \ud835\udc67 ]\ud835\udc47 \u2208 R12. It \ud835\udc563 11 12 13 21 22 23 31 32 33 \ud835\udc5d \ud835\udc5d \ud835\udc5d should be noted that motors are only placed at \ud835\udc341, \ud835\udc342 and \ud835\udc343 such that \ud835\udc5e\ud835\udc4e = [\ud835\udc5e11, \ud835\udc5e21, \ud835\udc5e31]\ud835\udc47 " + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_35_0001419_978-94-007-7401-8-Figure3.128-1.png", + "original_path": "designv11-35/openalex_figure/designv11_35_0001419_978-94-007-7401-8-Figure3.128-1.png", + "caption": "Fig. 3.128 4RRPPaR-type overactuated PM with coupled Sch\u00f6nflies motions defined by MF \u00bc SF \u00bc 4; RF\u00f0 \u00de \u00bc v1; v2; v3;xd\u00f0 \u00de, TF = 0, NF = 14, limb topology R||R||P\\Pa||R", + "texts": [ + " 3.122) R||P||R\\R||Pa (Fig. 3.52o) Idem No. 15 36. 4RRPRPa (Fig. 3.123) R||R||P\\R||Pa (Fig. 3.52p) Idem No. 15 37. 4RRRPPa (Fig. 3.124) R||R\\R\\P\\kPa (Fig. 3.52q) Idem No. 15 (continued) 3.2 Topologies with Complex Limbs 361 Table 3.6 (continued) No. PM type Limb topology Connecting conditions 38. 4PRRPaR (Fig. 3.125) P||R||R\\Pa||R (Fig. 3.52r) Idem No. 13 39. 4RPRPaR (Fig. 3.126) R||P||R\\Pa||R (Fig. 3.52s) Idem No. 13 40. 4RPRPaR (Fig. 3.127) R||P||R\\Pa||R (Fig. 3.52t) Idem No. 13 41. 4RRPPaR (Fig. 3.128) R||R||P\\Pa||R (Fig. 3.52u) Idem No. 13 42. 4PPaRRR (Fig. 3.129a) P\\Pa\\kR||R||\\kR (Fig. 3.52v) Idem No. 13 43. 4PaPRRR (Fig. 3.129b) Pa\\P\\\\R||R\\\\R (Fig. 3.52w) Idem No. 13 44. 4PPaRRR (Fig. 3.130) P\\Pa\\\\R||R\\kR (Fig. 3.52x) Idem No. 13 45. 4PaPRRR (Fig. 3.131) Pa\\P||R||R\\kR (Fig. 3.52y) Idem No. 13 46. 4PaRPRR (Fig. 3.132) Pa\\R||P||R\\kR (Fig. 3.52z) Idem No. 13 47. 4PaRPRR (Fig. 3.133) Pa\\R||P||R\\kR (Fig. 3.52a0) Idem No. 13 48. 4PaRRPR (Fig. 3.134) Pa\\R||R||P\\kR (Fig. 3.52b0) Idem No. 13 49. 4RCRPa (Fig" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_35_0001419_978-94-007-7401-8-Figure5.65-1.png", + "original_path": "designv11-35/openalex_figure/designv11_35_0001419_978-94-007-7401-8-Figure5.65-1.png", + "caption": "Fig. 5.65 3PaPaPaR-1RPaPatP-type fully-parallel PM with uncoupled Sch\u00f6nflies motions defined by MF = SF = 4, RF\u00f0 \u00de \u00bc \u00f0v1; v2; v3;xb\u00de, TF = 0, NF = 39, limb topology Pa\\Pa||Pa\\kR, Pa\\Pa||Pa||R and R||Pa||Pat||P", + "texts": [ + "1b) 28. 3PaPaPaR-1RUPU (Fig. 5.62) Pa\\Pa||Pa\\\\R (Fig. 5.6c) Idem No. 27 Pa\\Pa||Pa||R (Fig. 5.6b) R\\R\\R\\P\\kR\\R (Fig. 5.1b) 29. 3PaPaPaR-1RPaPaP (Fig. 5.63) Pa\\Pa||Pa\\kR (Fig. 5.6a) Idem No. 1 Pa\\Pa||Pa||R (Fig. 5.6b) R||Pa||Pa||P (Fig. 5.4k) (continued) 5.1 Fully-Parallel Topologies 533 Table 5.4 (continued) No. PM type Limb topology Connecting conditions 30. 3PaPaPaR-1RPaPaP (Fig. 5.64) Pa\\Pa||Pa\\\\R (Fig. 5.6c) Idem No. 1 Pa\\Pa||Pa||R (Fig. 5.6b) R||Pa||Pa||P (Fig. 5.54k) 31. 3PaPaPaR1RPaPatP (Fig. 5.65) Pa\\Pa||Pa\\kR (Fig. 5.6a) Idem No. 1 Pa\\Pa||Pa||R (Fig. 5.6b) R||Pa||Pat||P (Fig. 5.4l) 32. 3PaPaPaR1RPaPatP (Fig. 5.66) Pa\\Pa||Pa\\\\R (Fig. 5.6c) Idem No. 1 Pa\\Pa||Pa||R (Fig. 5.6b) R||Pa||Pat||P (Fig. 5.54l) 33. 3PaPaPaR-1RPPaPa (Fig. 5.67) Pa\\Pa||Pa\\kR (Fig. 5.6a) The revolute joint between links 7D and 11 of G4 limb and the last revolute joints of limbs G1, G2 and G3 have superposed axes Pa\\Pa||Pa||R (Fig. 5.6b) R||P||Pa||Pa (Fig. 5.4m) 34. 3PaPaPaR-1RPPaPa (Fig. 5.68) Pa\\Pa||Pa\\\\R (Fig. 5.6c) Idem No" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_35_0001596_isscc.1959.1157072-Figure7-1.png", + "original_path": "designv11-35/openalex_figure/designv11_35_0001596_isscc.1959.1157072-Figure7-1.png", + "caption": "Figure 7-Two-array coupling providing an inverting, or complementing, transfer. Solid arrows in A show ZERO state, dotted show ONE state. Sequence as in Figure 6 here leaves B cmtaining inverse of initial A-contained", + "texts": [ + "operation nor during the subsequent clear-A operation does a disturbance propagate out of this two-array system which is serious enough to affect the information states of adjoining arrays. No extra decoupling devices are thus needed, and it is found that extension of this transfer-clear procedure to a chain of arrays yields a serial shift register (Figure 6) . Gain characteristics are readily achievable such that the register can be closed upon itself to obtain continuous circulation of stable, arbitrary ONE-ZERO patterns. Reversing the output connections (Figure 7 ) of an array provides a view, looking left from the driven nodes, of a low threshold in CLEARED state, and a high threshold in SET or RESET states. This results in a complementary output from the array. Furthemore, it is easy to demonstrate that a multi-input array ( e .g . , A or B of Figure 8 ) assumes the state, after a succession of input transfers, which is the logical sum of the variables being transferred into it. Judicious use of the abilities outlined above can provide networks capable in principle of realizing any logical function" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_35_0001419_978-94-007-7401-8-Figure3.100-1.png", + "original_path": "designv11-35/openalex_figure/designv11_35_0001419_978-94-007-7401-8-Figure3.100-1.png", + "caption": "Fig. 3.100 4RPaPaP-type overactuated PM with coupled Sch\u00f6nflies motions defined by MF \u00bc SF \u00bc 4; RF\u00f0 \u00de \u00bc v1; v2; v3;xd\u00f0 \u00de, TF = 0, NF = 30, limb topology R||Pa\\Pa\\\\P", + "texts": [ + "94) Pa||P||R||Pa (Fig. 3.51f) Idem No. 1 7. 4RPaPPa (Fig. 3.95) R||Pa||P||Pa (Fig. 3.51g) Idem No. 1 8. 4PaPaPR (Fig. 3.96) Pa||Pa||P||R (Fig. 3.51h) Idem No. 1 (continued) 360 3 Overactuated Topologies with Coupled Sch\u00f6nflies Motions Table 3.6 (continued) No. PM type Limb topology Connecting conditions 9. 4PaPaRP (Fig. 3.97) Pa||Pa||R||P (Fig. 3.51i) Idem No. 1 10. 4PaRPaP (Fig. 3.98) Pa||R||Pa||P (Fig. 3.51j) Idem No. 1 11. 4PaRPPa (Fig. 3.99) Pa||R||P||Pa (Fig. 3.51k) Idem No. 1 12. 4RPaPaP (Fig. 3.100) R||Pa\\Pa\\\\P (Fig. 3.51l) The first revolute joints of the four limbs have parallel axes 13. 4PaRPaR (Fig. 3.101) Pa||R\\Pa\\kR (Fig. 3.51m) The last revolute joints of the four limbs have parallel axes 14. 4PaPaRR (Fig. 3.102) Pa\\Pa||R||R (Fig. 3.51n) Idem No. 13 15. 4PaRRPa (Fig. 3.103) Pa\\R||R||Pa (Fig. 3.51o) The revolute joints of the parallelogram loops connecting the four limbs to the moving plateform have parallel axes 16. 4PaPPaR (Fig. 3.104) Pa\\P\\\\Pa||R (Fig. 3.51p) Idem No. 13 17. 4PaRPaP (Fig" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_35_0001419_978-94-007-7401-8-Figure2.4-1.png", + "original_path": "designv11-35/openalex_figure/designv11_35_0001419_978-94-007-7401-8-Figure2.4-1.png", + "caption": "Fig. 2.4 Fully-parallel PMs with coupled Sch\u00f6nflies motions of types 4RRRRR (a) and 4PRRRR (b) defined by MF \u00bc SF \u00bc 4; RF\u00f0 \u00de \u00bc \u00f0v1; v2; v3; xd\u00de, TF = 0, NF = 2, limb topology R||R\\R||R||R (a) and P\\R\\R||R||R (b)", + "texts": [ + " PM type Limb topology Connecting conditions 1. 4RRRRR (Fig. 2.2a) R\\R||R||R\\k R (Fig. 2.1a) The first and the last revolute joints of the four limbs have parallel axes 2. 4RRRRR (Fig. 2.2b) R||R\\R||R\\R (Fig. 2.1b) The first, the second and the last revolute joints of the four limbs have parallel axes 3. 4RRRRR (Fig. 2.3a) R||R||R\\R||R (Fig. 2.1c) The two last revolute joints of the four limbs have parallel axes 4. 4RRRRR (Fig. 2.3b) R||R||R\\R||R (Fig. 2.1c) The three first revolute joints of the four limbs have parallel axes 5. 4RRRRR (Fig. 2.4a) R||R\\R||R||R (Fig. 2.1d) The two first revolute joints of the four limbs have parallel axes. 6. 4PRRRR (Fig. 2.4b) P\\R\\R||R||R (Fig. 2.1e) The second joints of the four limbs have parallel axes 7. 4RRRPR (Fig. 2.5a) R||R\\R\\P\\kR (Fig. 2.1f) Idem No. 5 8. 4RRPRR (Fig. 2.5b) R||R||P\\R||R (Fig. 2.1g) Idem No. 5 9. 4PRRRR (Fig. 2.6a) P\\R\\R||R\\R (Fig. 2.1h) The second and the last joints of the four limbs have parallel axes 10. 4PRRRR (Fig. 2.6b) P||R\\R||R\\R (Fig. 2.1i) Idem No. 9 11. 4RRPRR (Fig. 2.7a) R\\R\\P\\kR\\R (Fig. 2.1j) Idem No. 1 12. 4PRRRR (Fig. 2.7b) P\\R||R\\R||R (Fig. 2.1k) Idem No. 3 13. 4PRRRR (Fig. 2" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_35_0000360_j.matpr.2021.04.147-Figure1-1.png", + "original_path": "designv11-35/openalex_figure/designv11_35_0000360_j.matpr.2021.04.147-Figure1-1.png", + "caption": "Fig. 1. Front impact deformation.", + "texts": [], + "surrounding_texts": [ + "Off road vehicle is shaped to race and steer on different terrains. ORV is designed in such a way that it can endure off-roading terrains. In off-terrain circumstances, the vehicle bears dynamic loads and all that is sustained through the chassis frame. Chassis frame bears every mountings and assembly, so it is expected from an ORV chassis frame to sustain both static and dynamic loads. The selection of materials for chassis greatly depends on the high tensile strength and material light weight. The majority of manufacturers favour lightweight, cost-effective, safe, and recyclable materials." + ] + }, + { + "image_filename": "designv11_35_0000636_j.mechmachtheory.2021.104432-Figure3-1.png", + "original_path": "designv11-35/openalex_figure/designv11_35_0000636_j.mechmachtheory.2021.104432-Figure3-1.png", + "caption": "Fig. 3. Kinematic representation of the MLTA using a cylindrical joint at the shear centre.", + "texts": [ + " (1), using the Gr\u00fcblerKutzbach criterion [19]: DOF = 6 \u22c5 (k+ l) \u2212 \u2211g i (6 \u2212 fi) \u2212 r = 6 \u22c5 (2+ 4) \u2212 \u22119 i (6 \u2212 fi) \u2212 r = 36 \u2212 6 \u22c5 3 \u2212 2 \u22c5 5 \u2212 1 \u22c5 4 \u2212 2 = 2 (1) where: k, l Number of knuckles and links or bodies i Index for the joint g Number of overall joints fi DOFs for joint i r Number of redundant DOFs In this approach, the twist-beam substructure of the MLTA is modelled as proposed by Matschinsky in [20]. In this approximation, the flexible cross beam and the trailing arms are replaced by two rigid equivalent bodies, which are interconnected by a cylindrical joint located in the shear centre (SC) of the cross beam (see Fig. 3). Two spherical joints at the hardpoints U are used as the rear body attachments for the trailing arms. With this type of configuration, the motion of one equivalent body can momentarily be approximated by two instantaneous rotational axes for the parallel apar and opposite aopp wheel travel, respectively. While the axis of rotation for the parallel travel is defined by both body mounts U, the axis for the opposite wheel travel is generated by the corresponding body mount and the shear centre SC (see Fig. 3) [20]. Furthermore, the knuckles are attached to the equivalent bodies via the revolute joints at the hardpoints RU. Compared to a regular TBA, this leaves each knuckle with one additional DOF, which are then compensated by the longitudinal links. These are attached to the knuckle via a spherical joint, located at RL and fixed to the BIW by another spherical joint at the position L. In both longitudinal links, one redundant DOF occurs, which represents a rotation about the longitudinal axis of the corresponding link", + " With this alternative design, the distance between the wheel patch and the revolute joint is reduced, leading to an overall reduction of the loads. On the other hand, the redundant DOFs r are converted into additional rotational DOFs for the wheels, which must be locked with supplementary links or joints (e.g. cardanic joints or integral links). This results in a more complex suspension design, and for this reason, the current work will focus on the initial mechanism design shown in Fig. 2(b) and Fig. 3. Similar to a conventional TBA, the integration of the reversed TBA into the longitudinal Watt\u2019s linkage also leads to two different motion states for the MLTA (see Eq. (1)). While the parallel wheel travel of the MLTA is mainly dependant on the design of the Watt\u2019s linkage and the orientation of the revolute axes, the opposite wheel travel is highly influenced by the nonlinear twisting and bending of the flexible cross beam and the deformations in the body bushings [21,22]; that is why simplifications are used for the ideal-kinematic modelling", + " In Section 3.2 these are iteratively integrated so that the trajectories of each hardpoint can be obtained and thus used in the optimisation process shown in Section 4. The given calculation approach is based on the works of T. Niessing and X. Fang Mechanism and Machine Theory 165 (2021) 104432 T. Niessing and X. Fang Mechanism and Machine Theory 165 (2021) 104432 Matschinsky and Schnelle carried out in [20,23,24] and applied to the mechanism of the MLTA. In contrast to the kinematic model shown in Fig. 3, Schnelle proposes a revolute joint instead of a cylindrical one located in the SC (see, also, Fig. 4(a)) [23]. With that, the former spherical joints used for both body mounts are replaced by two joints with a translational DOF in the lateral direction. In the further course of this work, these types of joints will be referred to as inline joints. This approach can be considered a more realistic representation of the elasto-kinematic deformations in the bushings and the cross beam during body roll [20]", + " It can be written as: v K,r = v RU,r + \u03a9 K,r \u00d7 r RUWC (10) With that the angular and translational velocities of the twist-beam equivalent body as well as the knuckle are fully defined. The velocities of the remaining hardpoints result from these and the corresponding geometry. A main advantage of this approach is that it can be expanded to include angled inline joints which may be used to represent angled body bushing, that are commonly used for TBA to compensate for elasto-kinematic compliances [25]. Furthermore, only minor changes are required to apply this calculation approach to the mechanism shown in Fig. 3. In case of a cylindrical joint located at the SC, the isolated right equivalent body undergoes an additional translational motion relative to the left body. Its direction is the same as the axis of the cylindrical joint. Due to the now fixed spherical joints (see Fig. 3), the distance between these body mounts cannot change. However, the torsional rotation increases the distance of the isolated mounts (see Eq. (4)). Therefore, this must now be compensated by the translational motion of the SC. Thus, the following condition needs to be fulfilled: d\u0307Ur = v Ur \u22c5 e lr = e lr \u22c5 ( \u03c8\u0307 r \u00d7 U r + d\u0307Sc,r \u22c5 yR ) = 0 d\u0307Sc,r = \u2212 e lr \u22c5 \u03c8\u0307r \u00d7 U r elr \u22c5 yR (11) T. Niessing and X. Fang Mechanism and Machine Theory 165 (2021) 104432 where: d\u0307SC,r Magnitude of the velocity at the shear centre using a cylindrical joint With this additional equation, the velocity of the right body mount is completely defined", + " \u2022 L must be located underneath the cross section of the rear rail of the body to ensure a connection with sufficient local stiffness. The position is defined by a straight line. As can be found in Figs. 6(b)\u2013(d), the hardpoints are considered as the centre points of their respective joints. For this reason, offsets are used to take the outer dimensions of the bushing geometry into account. T. Niessing and X. Fang Mechanism and Machine Theory 165 (2021) 104432 To integrate the MLTA into a vehicle, the ideal kinematic configurations of joints, shown in Fig. 3 or 4 (a), are replaced by a set of rubber bushings as shown in Fig. 2. The reasons for this are the overall lower production costs of the rubber bushings, the damping of road-induced vibrations, and the possibility of implementing different stiffnesses in the local directions [13,20]. In addition, the former simplification by the usage of rigid equivalent bodies representing the twist-beam substructure (see Section 3) needs to be reversed. These elasticities are modelled with two different approaches", + " 7(b) and (c)) and also meet the boundary condition of the maximum y-coordinate of RUi during the roll motion. The span of 2.5 mm is a result of a compromise between the characteristics of the roll steer and camber as well as the camber compliance. The position of SC scattered within a range of about 58 mm in the x-direction and 18 mm in the z-direction. Keeping the overall low range of U in mind, this indicates that the instantaneous rotational axis to approximate the opposite wheel travel aopp (see Fig. 3) varies more heavily in the x-direction than in the z-direction. Since it is directly connected to the roll steer and camber gradients, it leads to the conclusion that the roll steer is limited more strictly by the maximum limit to the y-coordinate of RUi than the roll camber gradient. Fig. 8(b) shows a spider diagram of the first local minima in comparison to the \u201cdesired\u201d as well as the \u201crequired\u201d maximum and minimum values of the suspension characteristics. When looking at the objective values, it is noticeable that the two weighted variables for the roll motion, the roll steer and camber, fulfil their desired target values", + " While the roll steer gradient of the TBA increases with a growing mass, the corresponding value of the MLTA decreases (Fig. 10(b)). The reason for this can be found in the position of the SC with respect to the rear body mount U. For the MLTA the SC is typically positioned below the hardpoint U (see Fig. 8 (a)). Therefore, a parallel jounce motion reduces the distance between both hardpoints, and consequently the orientation of the instantaneous axis of the opposite wheel travel aopp flattens (see Fig. 3). For the conventional TBA, the body mount is located underneath the SC and, therefore, an increasing wheel travel also leads to an increasing roll steer gradient. Since the MLTA has a structural toe-in behaviour under lateral forces, the deviations to the TBA regarding the roll steer can be addressed by the elastic design as well as the bushing setup of the twist-beam substructure. T. Niessing and X. Fang Mechanism and Machine Theory 165 (2021) 104432 While the anti-lift of the MLTA is at the same level as the TBA for the first three loading states, the corresponding value for the gross vehicle weight differs by about 20% (see Fig" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_35_0001419_978-94-007-7401-8-Figure2.45-1.png", + "original_path": "designv11-35/openalex_figure/designv11_35_0001419_978-94-007-7401-8-Figure2.45-1.png", + "caption": "Fig. 2.45 4PaPRPa-type fully-parallel PM with coupled Sch\u00f6nflies motions defined by MF \u00bc SF \u00bc 4; RF\u00f0 \u00de \u00bc \u00f0v1; v2; v3;xd\u00de, TF = 0, NF = 30, limb topology Pa\\P\\kR\\Pa", + "texts": [ + "40) P\\Pa\\\\Pa\\kR (Fig. 2.22a) Idem No. 4 29. 4PPaRPa (Fig. 2.41) P\\Pa||R\\Pa (Fig. 2.22b) The axes of the revolute joints connecting the two parallelogram loops of the four limbs are parallel (continued) 2.2 Topologies with Complex Limbs 175 Table 2.3 (continued) No. PM type Limb topology Connecting conditions 30. 4PRPaPa (Fig. 2.42) P\\R||Pa\\Pa (Fig. 2.22c) Idem No. 5 31. 4RPPaPa (Fig. 2.43) R\\P\\kPa\\\\Pa (Fig. 2.22d) Idem No. 15 32. 4PaPPaR (Fig. 2.44) Pa\\P\\\\Pa\\\\R (Fig. 2.22e) Idem No. 4 33. 4PaPRPa (Fig. 2.45) Pa\\P\\kR\\Pa (Fig. 2.22f) The axes of the revolute joints connecting links 5 and 6 of the four limbs are parallel 34. 4PaRPPa (Fig. 2.46) Pa||R\\P\\\\Pa (Fig. 2.22g) The axes of the revolute joints connecting links 4 and 5 of the four limbs are parallel 35. 4RPaPPa (Fig. 2.47) R||Pa\\P\\\\Pa (Fig. 2.22h) Idem No. 15 36. 4PaPaPR (Fig. 2.48) Pa\\Pa\\\\P\\\\R (Fig. 2.22i) Idem No. 4 37. 4RPaPaR (Fig. 2.49) R\\Pa||Pa\\kR (Fig. 2.22j) The first and the last revolute joints of the four limbs are parallel 38. 4RPaPaR (Fig" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_35_0000019_cac51589.2020.9327633-Figure1-1.png", + "original_path": "designv11-35/openalex_figure/designv11_35_0000019_cac51589.2020.9327633-Figure1-1.png", + "caption": "Fig. 1. coordinate of the rotorcraft aerial manipulator", + "texts": [ + " Finally, the performance of the designed controller is tested through the simulation experiments in MATLAB. II. DYNAMIC MODEL In this paper, the rotorcraft aerial manipulator includes a two-degree-of-freedom manipulator with revolute-revolute (RR) structure and a carbon fiber frame with length of 450 mm. Considering the installation and operation of the manipulator, the structure of the quadrotor UAV choose \u2018X\u2019 type. Before establishing the kinematics model of the RAM, the coordinate of the system should be established firstly, which is shown in Fig. 1. ( )O X Y Zb b b bB is the body coordinate, ( )O X Y Zn n n nN is the inertial coordinate. ( )O X Y Z0 0 0 0J is base coordinate of the manipulator and its origin coincides with the origin of the airframe coordinate. After establishing the coordinate system, the state of the UAV is described by Euler angle [ ]T , including roll angle , pitch angle and yaw angle . Therefore, the rotation matrix N BR of the attitude of UAV can be expressed by equation (1). Through the rotation matrix N BR , the rotation and translation motion of the UAV can be described, and then the kinematic equation of the UAV can be derived" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_35_0000228_j.matpr.2021.02.620-Figure5-1.png", + "original_path": "designv11-35/openalex_figure/designv11_35_0000228_j.matpr.2021.02.620-Figure5-1.png", + "caption": "Fig. 5. Ansys results of Connecting rod with 42CrMo4 material.", + "texts": [ + " In order to ascertain the accurate results in the simulation, it is essential that the finite element model is well prepared, with the appropriate element size and connectors. Therefore, the selection of the correct element size should be carefully reviewed. For all practical purposes, the maximum force acting on connecting rod is taken from gas pressure applying on the piston neglecting the inertial effects. Fig. 4 illustrates the maximum gas force about 30.4 KN applying on the piston pin of the connecting rod and constrained at the big end. Further, the analysis of connecting rod has been carried out by assigning the 42CrMo4 material. The Fig. 5 shows the total deformation, Von-Mises stress, life cycle and factor of safety obtained for the connecting rod. The stress-life method usually used for millions of cycles, where the stresses are elastic and this method is stated to infinite design. It is based on the fatigue limit or endurance limit of the material. For fatigue strength at different life cycle (see Table 4); S N\u00f0 \u00de \u00bc a Nb \u00f01\u00de Fig. 3. Meshed Connecting Rod. Fig. 2. Drafting of Connecting Rod. where a \u00bc \u00f00:9 SUt\u00de2 Se ; b \u00bc 1 3 log 0:9 SUt Se ; Table 6 and Fig" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_35_0001419_978-94-007-7401-8-Figure5.49-1.png", + "original_path": "designv11-35/openalex_figure/designv11_35_0001419_978-94-007-7401-8-Figure5.49-1.png", + "caption": "Fig. 5.49 3PaPaPR-1RPPaPat-type fully-parallel PM with uncoupled Sch\u00f6nflies motions defined by MF = SF = 4, RF\u00f0 \u00de \u00bc \u00f0v1; v2; v3;xb\u00de, TF = 0, NF = 30 limb topology Pa\\Pa\\kP\\\\R, Pa\\Pa\\kP\\kR and R||P||Pa||Pat", + "texts": [ + "4e) R||Pa||Pa||P (Fig. 5.4k) 11. 3PaPaPR1RPaPatP (Fig. 5.45) Pa\\Pa\\kP\\\\R (Fig. 5.4f) Idem No. 1 Pa\\Pa\\kP\\kR (Fig. 5.4h) R||Pa||Pat||P (Fig. 5.4l) 12. 3PaPaPR1RPaPatP (Fig. 5.46) Pa\\Pa\\\\P\\\\R (Fig. 5.4g) Idem No. 1 Pa\\Pa\\\\P\\kR (Fig. 5.4e) R||Pa||Pat||P (Fig. 5.4l) 13. 3PaPaPR-1RPPaPa (Fig. 5.47) Pa\\Pa\\kP\\\\R (Fig. 5.4f) Idem No. 3 Pa\\Pa\\kP\\kR (Fig. 5.4h) R||P||Pa||Pa (Fig. 5.4m) 14. 3PaPaPR-1RPPaPa (Fig. 5.48) Pa\\Pa\\\\P\\\\R (Fig. 5.4g) Idem No. 3 Pa\\Pa\\\\P\\kR (Fig. 5.4e) R||P||Pa||Pa (Fig. 5.4m) 15. 3PaPaPR-1RPPaPat (Fig. 5.49) Pa\\PakP\\\\R (Fig. 5.4f) Idem No. 3 Pa\\Pa\\kP\\kR (Fig. 5.4h) R||P||Pa||Pat (Fig. 5.4n) 16. 3PaPaPR-1RPPaPat (Fig. 5.50) Pa\\Pa\\\\P\\\\R (Fig. 5.4g) Idem No. 3 Pa\\Pa\\\\P\\kR (Fig. 5.4e) R||P||Pa||Pat (Fig. 5.4n) 17. 2PaPaRRR-1PaPaRR1RPaPaP (Fig. 5.51) Pa\\Pa||R||R\\\\R (Fig. 5.5a) Idem No. 1 Pa\\Pa||R||R (Fig. 5.4i) R||Pa||Pa||P (Fig. 5.4k) (continued) 532 5 Topologies with Uncoupled Sch\u00f6nflies Motions Table 5.4 (continued) No. PM type Limb topology Connecting conditions 18. 2PaPaRRR-1PaPaRR1RPaPaP (Fig. 5.52) Pa\\Pa||R||R\\kR (Fig" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_35_0001419_978-94-007-7401-8-Figure6.2-1.png", + "original_path": "designv11-35/openalex_figure/designv11_35_0001419_978-94-007-7401-8-Figure6.2-1.png", + "caption": "Fig. 6.2 Maximaly regular fully-parallel PMs with Sch\u00f6nflies motions of types 2PRPRR1PRPR-1RPPP (a) and 2PRRRR-1PRRR-1RPPP (b) defined by MF = SF = 4, (RF) = (v1; v2; v3;xb), TF = 0, NF = 4, limb topology R\\P\\\\P\\\\P and P||R\\P\\||R\\R, P||R\\P\\||R (a), P||R||R||R\\R, P||R||R||R (b)", + "texts": [], + "surrounding_texts": [ + "In the fully-parallel and maximally regular topologies of PMs with Sch\u00f6nflies motions F G1-G2-G3-G4 presented in this section, the moving platform n : nGi (i = 1, 2, 3, 4) is connected to the reference platform 1 : 1Gi : 0 by four simple limbs with four, five or six degrees of connectivity. One linear actuator is combined in the first prismatic or cylindrical pair of limbs G1, G2 and G3, and one rotary actuator in the first revolute pair of limb G4. Limbs G1, G2 and G3 are used for positioning the moving platform and limb G4 for orienting it. Various maximally regular topologies of PMs with Sch\u00f6nflies motions of the moving platform and no idle mobilities can be obtained by using various limb topologies presented in Figs. 4.1, 4.2 and 5.1. Fully-parallel topologies with at least two identical limbs are illustrated in Figs. 6.1, 6.2, 6.3, 6.4, 6.5, 6.6. The limb topology and connecting conditions of these solutions are systematized in Table 6.1, as are their structural parameters in Tables 6.2, 6.3, 6.4. G. Gogu, Structural Synthesis of Parallel Robots, Solid Mechanics and Its Applications 206, DOI: 10.1007/978-94-007-7401-8_6, Springer Science+Business Media Dordrecht 2014 579 580 6 Maximally Regular Topologies with Sch\u00f6nflies Motions 6.1 Fully-Parallel Topologies with Simple Limbs 581 582 6 Maximally Regular Topologies with Sch\u00f6nflies Motions 6.1 Fully-Parallel Topologies with Simple Limbs 583 584 6 Maximally Regular Topologies with Sch\u00f6nflies Motions 6.1 Fully-Parallel Topologies with Simple Limbs 585 T ab le 6. 1 L im b to po lo gy an d co nn ec ti ng co nd it io ns of th e fu ll y- pa ra ll el so lu ti on s w it h no id le m ob il it ie s pr es en te d in F ig s. 6. 1, 6. 2, 6. 3, 6. 4, 6. 5, 6. 6 N o. P M ty pe L im b to po lo gy C on ne ct in g co nd it io ns 1. 3P P P R -1 R P P P (F ig . 6. 1a ) P \\ P ? ? P \\ || R (F ig .4 .1 a) P \\ P ? ? P ? ? R (F ig . 4. 1b ) R ? P ? ? P ? ? P (F ig . 5. 1a ) T he la st jo in ts of th e fo ur li m bs ha ve su pe rp os ed ax es /d ir ec ti on s. T he ac tu at ed pr is m at ic jo in ts of li m bs G 1 , G 2 an d G 3 ha ve or th og on al di re ct io ns 2. 2P P R R R -1 P P R R -1 R P P P (F ig . 6. 1b ) P ? P ? || R ||R ? R (F ig . 4. 2a ) P ? P ? || R ||R (F ig . 4. 1d ) R ? P ? ? P ? ? P (F ig . 5. 1a ) Id em no . 1 3. 2P R P R R -1 P R P R -1 R P P P (F ig . 6. 2a ) P ||R ? P ? || R ? R (F ig . 4. 2b ) P ||R ? P ? || R (F ig . 4. 1e ) R ? P ? ? P ? ? P (F ig . 5. 1a ) Id em no . 1 4. 2P R R R R -1 P R R R -1 R P P P (F ig . 6. 2b ) P ||R ||R ||R ? R (F ig . 4. 2d ) P ||R ||R ||R (F ig . 4. 1g ) R ? P ? ? P ? ? P (F ig . 5. 1a ) Id em no . 1 5. 3P P P R -1 R U P U (F ig . 6. 3a ) P ? P ? ? P ? || R (F ig . 4. 1a ) P ? P ? ? P ? ? R (F ig .4 .1 b) R ? R ? R ? P ? || R ? R (F ig . 5. 1b ) T he fi rs tr ev ol ut e jo in t of G 4 -l im b an d th e la st la st jo in ts of li m bs G 1 , G 2 an d G 3 ha ve pa ra ll el ax es T he la st re vo lu te jo in ts of li m bs G 1 , G 2 an d G 3 ha ve su pe rp os ed ax es . T he ac tu at ed pr is m at ic jo in ts of li m bs G 1 , G 2 an d G 3 ha ve or th og on al di re ct io ns 6. 2P P R R R -1 P P R R -1 R U P U (F ig . 6. 3b ) P ? P ? || R ||R ? R (F ig . 4. 2a ) P ? P ? || R ||R (F ig . 4. 1d ) R ? R ? R ? P ? || R ? R (F ig . 5. 1b ) Id em no . 5 7. 2P R P R R -1 P R P R -1 R U P U (F ig . 6. 4a ) P ||R ? P ? || R ? R (F ig . 4. 2b ) P ||R ? P ? || R (F ig . 4. 1e ) R ? R ? R ? P ? || R ? R (F ig . 5. 1b ) Id em no . 5 8. 2P R R R R -1 P R R R -1 R U P U (F ig . 6. 4b ) P ||R ||R ||R ? R (F ig . 4. 2d ) P ||R ||R ||R (F ig . 4. 1g ) R ? R ? R ? P ? || R ? R (F ig . 5. 1b ) Id em no . 5 9. 2C P R R -1 C P R -1 R P P P (F ig . 6. 5a ) C ? P ? || R ? R (F ig . 4. 2e ) C ? P ? || R (F ig . 4. 1i ) R ? P ? ? P ? ? P (F ig . 5. 1a ) T he la st jo in ts of th e fo ur li m bs ha ve su pe rp os ed ax es /d ir ec ti on s. T he cy li nd ri ca l jo in ts of li m bs G 1 , G 2 an d G 3 ha ve or th og on al ax es (c on ti nu ed ) 586 6 Maximally Regular Topologies with Sch\u00f6nflies Motions T ab le 6. 1 (c on ti nu ed ) N o. P M ty pe L im b to po lo gy C on ne ct in g co nd it io ns 10 . 2C R R R -1 C R R -1 R P P P (F ig . 6. 5b ) C ||R ||R ? R (F ig . 4. 2g ) C ||R ||R (F ig . 4. 1k ) R ? P ? ? P ? ? P (F ig . 5. 1a ) Id em no . 9 11 . 2C P R R -1 C P R -1 R U P U (F ig . 6. 6a ) C ? P ? || R ? R (F ig . 4. 2e ) C ? P ? || R (F ig . 4. 1i ) R ? R ? R ? P ? || R ? R (F ig . 5. 1b ) T he fi rs tr ev ol ut e jo in t of G 4 -l im b an d th e la st la st jo in ts of li m bs G 1 , G 2 an d G 3 ha ve pa ra ll el ax es T he la st re vo lu te jo in ts of li m bs G 1 , G 2 an d G 3 ha ve su pe rp os ed ax es . T he cy li nd ri ca l jo in ts of li m bs G 1 , G 2 an d G 3 ha ve or th og on al ax es 12 . 2 C R R R -1 C R R -1 R U P U (F ig . 6. 6b ) C ||R ||R ? R (F ig . 4. 2g ) C ||R ||R (F ig . 4. 1k ) R ? R ? R ? P ? || R ? R (F ig . 5. 1b ) Id em no . 11 6.1 Fully-Parallel Topologies with Simple Limbs 587 Table 6.2 Structural parametersa of parallel mechanisms in Figs. 6.1 and 6.2 No. Structural parameter Solution Figure 6.1a Figures 6.1b and 6.2 1. m 14 16 2. pi (i = 1, 3) 4 5 3. p2 4 4 4. p4 4 4 5. p 16 18 6. q 3 3 7. k1 4 4 8. k2 0 0 9. k 4 4 10. (RG1) (v1; v2; v3;xb) (v1; v2; v3;xa;xb) 11. (RG2) (v1; v2; v3;xb) (v1; v2; v3;xb) 12. (RG3) (v1; v2; v3;xb) (v1; v2; v3;xb) 13. (RG4) (v1; v2; v3;xb) (v1; v2; v3;xb) 14. SGi (i = 1, 3) 4 5 15. SG2 4 4 16. SG4 4 4 17. rGi (i = 1, 2, 3) 0 0 18. rG4 0 0 19. MGi (i = 1, 3) 4 5 20. MG2 4 4 21. MG4 4 4 22. (RF) (v1; v2; v3;xb) (v1; v2; v3;xb) 23. SF 4 4 24. rl 0 0 25. rF 12 14 26. MF 4 4 27. NF 6 4 28. TF 0 0 29. Pp1 j\u00bc1 fj 4 5 30. Pp2 j\u00bc1 fj 4 4 31. Pp3 j\u00bc1 fj 4 5 32. Pp4 j\u00bc1 fj 4 4 33. Pp j\u00bc1 fj 16 18 a See footnote of Table 2.2 for the nomenclature of structural parameters 588 6 Maximally Regular Topologies with Sch\u00f6nflies Motions Table 6.3 Structural parametersa of parallel mechanisms in Figs. 6.3 and 6.4 No. Structural parameter Solution Figure 6.3a Figures 6.3b and 6.4 1. m 16 18 2. pi (i = 1,3) 4 5 3. p2 4 4 4. p4 6 6 5. p 18 20 6. q 3 3 7. k1 4 4 8. k2 0 0 9. k 4 4 10. (RG1) (v1; v2; v3;xb) (v1; v2; v3;xa;xb) 11. (RG2) (v1; v2; v3;xb) (v1; v2; v3;xb) 12. (RG3) (v1; v2; v3;xb) (v1; v2; v3;xb;xd) 13. (RG4) (v1; v2; v3;xa;xb;xd) (v1; v2; v3;xa;xb;xd) 14. SGi (i = 1, 3) 4 5 15. SG2 4 4 16. SG4 6 6 17. rGi (i = 1, 2, 3) 0 0 18. rG4 0 0 19. MGi (i = 1, 3) 4 5 20. MG2 4 4 21. MG4 6 6 22. (RF) (v1; v2; v3;xb) (v1; v2; v3;xb) 23. SF 4 4 24. rl 0 0 25. rF 14 16 26. MF 4 4 27. NF 4 2 28. TF 0 0 29. Pp1 j\u00bc1 fj 4 5 30. Pp2 j\u00bc1 fj 4 4 31. Pp3 j\u00bc1 fj 4 5 32. Pp4 j\u00bc1 fj 6 6 33. Pp j\u00bc1 fj 18 20 a See footnote of Table 2.2 for the nomenclature of structural parameters 6.1 Fully-Parallel Topologies with Simple Limbs 589 Table 6.4 Structural parametersa of parallel mechanisms in Figs. 6.5 and 6.6 No. Structural parameter Solution Figure 6.5 Figure 6.6 1. m 13 15 2. pi (i = 1,3) 4 4 3. p2 3 3 4. p4 4 6 5. p 15 17 6. q 3 3 7. k1 4 4 8. k2 0 0 9. k 4 4 10. (RG1) (v1; v2; v3;xa;xb) (v1; v2; v3;xa;xb) 11. (RG2) (v1; v2; v3;xb) (v1; v2; v3;xb) 12. (RG3) (v1; v2; v3;xb;xd) (v1; v2; v3;xb;xd) 13. (RG4) (v1; v2; v3;xb) (v1; v2; v3;xa;xb;xd) 14. SGi (i = 1, 3) 5 5 15. SG2 4 4 16. SG4 4 6 17. rGi (i = 1, 2, 3) 0 0 18. rG4 0 0 19. MGi (i = 1, 3) 5 5 20. MG2 4 4 21. MG4 4 6 22. (RF) (v1; v2; v3;xb) (v1; v2; v3;xb) 23. SF 4 4 24. rl 0 0 25. rF 14 16 26. MF 4 4 27. NF 4 2 28. TF 0 0 29. Pp1 j\u00bc1 fj 5 5 30. Pp2 j\u00bc1 fj 4 4 31. Pp3 j\u00bc1 fj 5 5 32. Pp4 j\u00bc1 fj 4 6 33. Pp j\u00bc1 fj 18 20 a See footnote of Table 2.2 for the nomenclature of structural parameters 590 6 Maximally Regular Topologies with Sch\u00f6nflies Motions" + ] + }, + { + "image_filename": "designv11_35_0000371_tmag.2021.3076134-Figure7-1.png", + "original_path": "designv11-35/openalex_figure/designv11_35_0000371_tmag.2021.3076134-Figure7-1.png", + "caption": "Fig. 7. Open-circuit flux distributions of 12-stator-pole ST-PMTFLGs having different translator pole numbers (\u03b8e = 0\u25e6). (a) 10-translator-pole. (b) 11-translator-pole. (c) 13-translator-pole. (d) 14-translator-pole.", + "texts": [ + " In this section, the 12-stator-pole ST-PMTFLG with 10-, 11-, 13-, and 14-translator-pole translators are designed and compared to maximize the average electromagnetic thrust. The machine performance compared including open-circuit flux linkage, back EMF, and ON-load electromagnetic thrust characteristics, are as follows. The open-circuit flux distributions of 12-stator-pole ST-PMTFLGs having different translator pole numbers (10, 11, 13, and 14) when translator positions in electric \u03b8e = 0 are shown in Fig. 7. With higher translator pole number, 12-stator-pole ST-PMTFLG has lower phase flux linkage, as shown in Fig. 8(a). However, the 12/11 stator/translator-pole Authorized licensed use limited to: California State University Fresno. Downloaded on June 26,2021 at 18:41:18 UTC from IEEE Xplore. Restrictions apply. ST-PMTFLG has the highest fundamental back EMF among these four machines, as shown in Fig. 8(b). The 12/10, 12/11, and 12/13 stator/translator-pole ST-PMTFLG has the third- and fifth-order phase back EMF harmonics" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_35_0001419_978-94-007-7401-8-Figure6.22-1.png", + "original_path": "designv11-35/openalex_figure/designv11_35_0001419_978-94-007-7401-8-Figure6.22-1.png", + "caption": "Fig. 6.22 Maximaly regular fully-parallel PMs with Sch\u00f6nflies motions of types 3PPaPR1RPPP (a) and 2PPaRRR-1PPaRR-1RPPP (b) defined by MF = SF = 4, (RF) = (v1; v2; v3;xb), TF = 0, NF = 15 (a), NF = 13 (b), limb topology R\\P ??P ??P and P||Pa\\P ??R, P||Pa\\P\\||R (a), P||Pa||R||R\\R, P||Pa||R||R (b)", + "texts": [ + " 2CRRR-1CRR-1CPaPat (Fig. 6.20b) C||R||R ?R (Fig. 4.2g) C||R||R (Fig. 4.1k) C||Pa||Pat (Fig. 5.4p) Idem no. 1 9. 3PPPaR-1RPPP (Fig. 6.21a) P ?P ?||Pa ?||R (Fig. 4.8b) P ?P ?||Pa||R, (Fig. 4.8c) R ?P\\\\P\\\\P (Fig. 5.1a) The last joints of the four limbs have superposed axes/ directions. The actuated prismatic joints of limbs G1, G2 and G3 have orthogonal directions 10. 3PPPaR-1RPPP (Fig. 6.21b) P ?P ?||Pa\\\\R (Fig. 4.8a) P ?P ?||Pa||R, (Fig. 4.8c) R\\\\P\\\\P\\\\P (Fig. 5.1a) Idem no. 9 11. 3PPaPR-1RPPP (Fig. 6.22a) P||Pa ?P\\\\R (Fig. 4.8d) P||Pa ?P ?||R (Fig. 4.8e) R ?P\\\\P\\\\P (Fig. 5.1a) Idem no. 9 12. 2PPaRRR-1PPaRR-1RPPP (Fig. 6.22b) P||Pa||R||R ?R (Fig. 4.9a) P||Pa||R||R (Fig. 4.8g) R ?P\\\\P\\\\P (Fig. 5.1a) Idem no. 9 13. 3PPPaR-1RUPU (Fig. 6.23a) P ?P ?||Pa ?||R (Fig. 4.8b) P ?P ?||Pa||R, (Fig. 4.8c) R ?R ?R ?P ?||R ?R (Fig. 5.1b) The first revolute joint of G4-limb and the last last joints of limbs G1, G2 and G3 have parallel axes. The last revolute joints of limbs G1, G2 and G3 have superposed axes. The actuated prismatic joints of limbs G1, G2 and G3 have orthogonal directions (continued) 614 6 Maximally Regular Topologies with Sch\u00f6nflies Motions Table 6", + "2 Fully-Parallel Topologies with Simple and Complex Limbs 617 Table 6.9 (continued) No. Structural parameter Solution Figure 6.18 Figures 6.19 and 6.20 23. SF 4 4 24. rl 6 6 25. rF 20 20 26. MF 4 4 27. NF 10 10 28. TF 0 0 29. Pp1 j\u00bc1 fj 5 5 30. Pp2 j\u00bc1 fj 4 4 31. Pp3 j\u00bc1 fj 5 5 32. Pp4 j\u00bc1 fj 10 10 33. Pp j\u00bc1 fj 24 24 a See footnote of Table 2.2 for the nomenclature of structural parameters 618 6 Maximally Regular Topologies with Sch\u00f6nflies Motions Table 6.10 (continued) No. Structural parameter Solution Figures 6.21 and 6.22a Figure 6.22b 24. rl 9 9 25. rF 21 23 26. MF 4 4 27. NF 15 13 28. TF 0 0 29. Pp1 j\u00bc1 fj 7 8 30. Pp2 j\u00bc1 fj 7 7 31. Pp3 j\u00bc1 fj 7 8 32. Pp4 j\u00bc1 fj 4 4 33. Pp j\u00bc1 fj 25 27 a See footnote of Table 2.2 for the nomenclature of structural parameters 6.2 Fully-Parallel Topologies with Simple and Complex Limbs 619 Table 6.11 (continued) No. Structural parameter Solution Figures 6.23 and 6.24a Figure 6.24b 25. rF 23 25 26. MF 4 4 27. NF 13 11 28. TF 0 0 29. Pp1 j\u00bc1 fj 7 8 30. Pp2 j\u00bc1 fj 7 7 31. Pp3 j\u00bc1 fj 7 8 32. Pp4 j\u00bc1 fj 6 6 33" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_35_0000100_iros45743.2020.9341746-Figure3-1.png", + "original_path": "designv11-35/openalex_figure/designv11_35_0000100_iros45743.2020.9341746-Figure3-1.png", + "caption": "Fig. 3. a) The diagram and the design parameters of non-uniform through-hole pattern on the nitinol tube. A pair of columns were engraved and each column has \u03b8 = 134\u00b0 of the central angle. b) the automized laser machining set-up with RT1000 Laser Tube Cutting Machine, Preco Inc., USA. c) the jig of the bending wrist for the heat treatment. d) the microscopic photograph of the laser-machined nitinol tube. e) the microscopic photograph of the tube\u2019s distal part after heat treatment.", + "texts": [ + " the proximal section is a 40-degree curve with a radius of curvature of 15 mm and the second part made a 30-degree curve with 25 mm. The wrist is made of a nitinol tube with 1.83 mm of the outer diameter and 1.54 mm of the inner diameter. To achieve a minimum radius of curvature is 15 mm, a specialized manufacturing process is required. It is as follows: 1) the nitinol tube was asymmetrically laser patterned to have higher curvature, and 2) the deployable arm was shape-set through the two-step heat treatment [23]. The asymmetric patterns illustrated in Fig. 3a reduces the bending stiffness along the bending direction, so it enables the higher curvature of the tube. At the same time, the asymmetric pattern maintains torsional rigidity and the flexural rigidity along the non-patterned area of Fig. 3a compared to the universal patterning [22], [24], [25]. Thus, the asymmetrically patterned deployable arm is more stable as it has a higher threshold for buckling, as opposed to the universal patterns. The through-hole patterns on the nitinol tube were created by laser cutting (RT1000 Laser Tube Cutting Machine, Preco Inc., KS, USA), where feed rate is 127 mm/min, the duty cycle is 12%, power is 250 watts, and pulse frequency is 500Hz, as shown in Fig. 3b. For the shape setting, an aluminum jig was manufactured as seen in Fig. 3c. Since the high curvature cannot be shaped by a single heat treatment process, the heat treatment process was performed twice: the first used a low curvature jig and the second used a high curvature jig. After the two heat treatment steps, the deployable arm was shaped into the 3050 Authorized licensed use limited to: University of Prince Edward Island. Downloaded on May 16,2021 at 23:26:55 UTC from IEEE Xplore. Restrictions apply. designated design. For each heat treatment, the patterned tube was placed in the jig, annealed in the furnace (3-1750, Vulcan Muffle Furnace, Neytech, USA) for 28 minutes at 530\u00b0C, and then quenched in room temperature water" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_35_0001706_pime_proc_1950_163_014_02-Figure4-1.png", + "original_path": "designv11-35/openalex_figure/designv11_35_0001706_pime_proc_1950_163_014_02-Figure4-1.png", + "caption": "Fig. 4. Enlarged Model of a Moving Packing", + "texts": [ + " Under these conditions escape of the liquid is then prevented only by viscous stresses in the film, and the nature of the surfaces plays no further part in friction. The behaviour of packings under film conditions was studied in a large model which was lightly loaded so that the film was built up at quite low speeds. The packing was represented by a soft rubber block slotted to form two separated sealing pads t See reference at foot of p. 98. (as shown in Fig. 3) which were then pressed on to a rotating glass-topped table (Fig. 4) in such a way that the interference was constant. Liquid at a pressure of 20 Ib. per sq. in. or less was applied to the slot, leakage being prevented by the seal formed between the rubber and glass. As the glass passed beneath the rubber block, a uniform layer of lubricant remained on it, whose thickness varied greatly with speed and loading. The level of liquid in the pipette tube (Fig. 4) remained constant even when there was excess liquid on the table, showing that the same quantity was passing beneath the leading (motoring) pad as was escaping from beneath the trailing one. (To obtain this balance the widths of the pads had to be precisely the same otherwise the liquid rose or fell steadily in the pipette.) Certain tests confirmed that the pads were sufficiently long to avoid error from leakage at the ends. The frictional characteristics of the film were found to be like those of the film in journal bearings despite the flexible nature of the pads and the differing pressures at the two sides" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_35_0001419_978-94-007-7401-8-Figure6.28-1.png", + "original_path": "designv11-35/openalex_figure/designv11_35_0001419_978-94-007-7401-8-Figure6.28-1.png", + "caption": "Fig. 6.28 3PPPaR-1RPaPatP-type maximaly regular fully-parallel PMs with Sch\u00f6nflies motions defined by MF = SF = 4, (RF) = (v1; v2; v3;xb), TF = 0, NF = 21, limb topology P\\P\\||Pa||R, R||Pa||Pat||P and P\\P\\||Pa\\||R (a), P\\P\\||Pa\\R (b)", + "texts": [ + "31, 6.32, 6.33, 6.34, 6.35, 6.36, 6.37, 6.38, 6.39, 6.40, 6.41 No. PM type Limb topology Connecting conditions 1. 3PPPaR-1RPaPaP (Fig. 6.27a) P\\P\\||Pa\\||R (Fig. 4.8b) P\\P\\||Pa||R, (Fig. 4.8c) R||Pa||Pa||P (Fig. 5.4k) The last joints of the four limbs have superposed axes/directions. The actuated prismatic joints of limbs G1, G2 and G3 have orthogonal directions 2. 3PPPaR-1RPaPaP (Fig. 6.27b) P\\P\\||Pa\\R (Fig. 4.8a) P\\P\\||Pa||R, (Fig. 4.8c) R||Pa||Pa||P (Fig. 5.4k) Idem no. 1 3. 3PPPaR-1RPaPatP (Fig. 6.28a) P\\P\\||Pa\\||R (Fig. 4.8b) P\\P\\||Pa||R, (Fig. 4.8c) R||Pa||Pat||P (Fig. 5.4l) Idem no. 1 4. 3PPPaR-1RPaPatP (Fig. 6.28b) P\\P\\||Pa\\R (Fig. 4.8a) P\\P\\||Pa||R, (Fig. 4.8c) R||Pa||Pat||P (Fig. 5.4l) Idem no. 1 5. 3PPaPR-1RPaPaP (Fig. 6.29a) P||Pa\\P\\\\R (Fig. 4.8d) P||Pa\\P\\||R (Fig. 4.8e) R||Pa||Pa||P (Fig. 5.4k) Idem no. 1 6. 3PPaPR-1RPaPatP (Fig. 6.29b) P||Pa\\P\\\\R (Fig. 4.8d) P||Pa\\P\\||R (Fig. 4.8e) R||Pa||Pat||P (Fig. 5.4l) Idem no. 1 7. 2PPaRRR-1PPaRR-1RPaPaP (Fig. 6.30) P||Pa||R||R\\R (Fig. 4.9a) P||Pa||R||R (Fig. 4.8g) R||Pa||Pa||P (Fig. 5.4k) Idem no. 1 8. 2PPaRRR-1PPaRR1RPaPatP (Fig. 6.31) P||Pa||R||R\\R (Fig. 4" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_35_0001419_978-94-007-7401-8-Figure3.60-1.png", + "original_path": "designv11-35/openalex_figure/designv11_35_0001419_978-94-007-7401-8-Figure3.60-1.png", + "caption": "Fig. 3.60 Overactuated PMs with coupled Sch\u00f6nflies motions of types 4RPaPP (a) and 4RPPPa (b) defined by MF \u00bc SF \u00bc 4; RF\u00f0 \u00de \u00bc v1; v2; v3;xd\u00f0 \u00de, TF = 0, NF = 18, limb topology R||Pa\\P\\kP (a) and R||P\\P\\kPa (b)", + "texts": [ + " 3.50c) Idem No. 1 4. 4PRRPa (Fig. 3.55b) P||R||R||Pa (Fig. 3.50d) Idem No. 1 5. 4PPRPa (Fig. 3.56a) P\\P||R||Pa (Fig. 3.50e) Idem No. 1 6. 4PPRPa (Fig. 3.56b) P\\P\\kR||Pa (Fig. 3.50f) Idem No. 1 7. 4PRPPa (Fig. 3.57a) P\\R||P||Pa (Fig. 3.50g) Idem No. 1 8. 4PRPaP (Fig. 3.57b) P\\R||Pa||P (Fig. 3.50h) Idem No. 1 9. 4RPRPa (Fig. 3.58a) R||P||R||Pa (Fig. 3.50i) Idem No. 1 10. 4RRPPa (Fig. 3.58b) R||R||P||Pa (Fig. 3.50j) Idem No. 1 11. RRPaP (Fig. 3.59) R||R||Pa||P (Fig. 3.50k) Idem No. 1 12. 4RPaPP (Fig. 3.60a) R||Pa\\P\\kP (Fig. 3.50l) Idem No. 1 13. 4RPPPa (Fig. 3.60b) R||P\\P\\kPa (Fig. 3.50m) Idem No. 1 14. 4PaRPP (Fig. 3.61a) Pa||R||P\\P (Fig. 3.50n) Idem No. 1 15. 4RPaPP (Fig. 3.61b) R||Pa||P\\P (Fig. 3.50o) Idem No. 1 16. 4RPPaP (Fig. 3.62a) R\\P\\kPa||P (Fig. 3.50p) Idem No. 1 17. 4PPPaR (Fig. 3.62b) P\\P\\kPa||R (Fig. 3.50q) Idem No. 1 18. 4PPPaR (Fig. 3.63a) P\\P||Pa||R (Fig. 3.50r) Idem No. 1 19. 4PPaRR (Fig. 3.63b) P||Pa||R||R (Fig. 3.50s) Idem No. 1 20. 4PPaRP (Fig. 3.64) P\\Pa||R||P (Fig. 3.50t) Idem No. 1 21. 4PPaPR (Fig. 3.65) P\\Pa||P||R (Fig. 3.50u) Idem No" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_35_0000141_tasc.2021.3062485-Figure4-1.png", + "original_path": "designv11-35/openalex_figure/designv11_35_0000141_tasc.2021.3062485-Figure4-1.png", + "caption": "Fig. 4. Magnetic field distribution due to field excitation coil current.", + "texts": [ + " This case is included to study the effect of the iron yoke on the ac losses and machine mass. The machine with iron yoke has the highest mass and lowest power density. All machine size and mass estimates include the contribution of end-turns and other inactive components. The radial build of the field winding on the rotor is selected as 7 mm ( = R2-R1 in Table II) for maximizing magnetic coupling between the field winding and stator winding. In the past, flat HTS tapes were used for the construction of the field winding, [14, Fig. 4.10], but the use of such construction in a small radius rotor increases radial build of the field winding, which reduces magnetic coupling between the field and stator windings. To accommodate necessary amp-turns in the allocated space on the rotor it becomes necessary to employ a conductor that could be bent in 3-D without impacting the performance of the HTS tapes. The CORC [11] cable satisfying these requirements is selected for the dc field winding. The cable made with 2 mm wide REBCO tapes is detailed in Table III", + " The 2-D analysis is complemented with correction factors to account for end-turn contributions. The correction factors were derived by comparing 2-D models with 3-D finite element models for similarly sized machines. The 2-D analytical models aided with correction factors are ideal for rapid scanning of a wide range of design options. Once an attractive design is achieved it could be further optimized using 3-D modeling. The purpose of the 2-D analysis here is simply to compare competing design concepts. Magnetic field distribution due to the field winding current is shown in Fig. 4, with armature Phase-A coil aligned with the field excitation coil on the D-axis. In the case of G10 teeth (air-core) stator, flux lines are not constrained in the winding region. However, in iron tooth stator, radially directed flux lines mostly pass through the iron teeth, but the tangentially directed flux lines go across slots. The concentration of flux lines in the slots has an adverse effect on the MgB2 cable critical current and ac losses. Fig. 5 shows the field in individual slots of iron and G10 teeth stators for a \u00bd pole-pitch\u2013slot numbers begin from the D-axis", + " Loss management in superconductors is critical. Losses in the superconducting windings and their support structure are removed with cryocoolers, which are not efficient at cryogenic temperatures. The machines, operating at 20 K, are cooled with LH2 available on board in the plane. However, LH2 cannot be used directly for cooling the winding due to safety and other concerns. An intermediate gas, like gaseous helium, is used for transporting the thermal load from the windings to LH2 reservoir. Kalsi et al. [16, Fig. 4] show a possible schematic for such a system. Rotor and stator winding losses are summarized in Table XI. Rotor field excitation coils carry dc and are protected from stator ac fields with an EM shield. These coils operate at high current (3\u20134 kA). Supplying these coils with current leads spanning room-temperature to 20 K coil temperature would cause a large thermal conduction load. Since the rotor cooling is challenging, it is preferable to minimize the thermal load on the rotor. Keeping this in mind, the field coils are supplied with a flux pump exciter, for keeping coil leads within a 20-K environment" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_35_0001419_978-94-007-7401-8-Figure2.31-1.png", + "original_path": "designv11-35/openalex_figure/designv11_35_0001419_978-94-007-7401-8-Figure2.31-1.png", + "caption": "Fig. 2.31 Fully-parallel PMs with coupled Sch\u00f6nflies motions of types 4PPaRP (a) and 4PPaPR (b) defined by MF \u00bc SF \u00bc 4; RF\u00f0 \u00de \u00bc \u00f0v1; v2; v3;xd\u00de, TF = 0, NF = 18, limb topology P||Pa\\R||P (a) and P||Pa\\P|| R (b)", + "texts": [ + "21d) The last revolute joints of the four limbs have parallel axes 5. 4PRPPa (Fig. 2.28a) P||R\\P||Pa (Fig. 2.21e) The second joints of the four limbs have parallel axes 6. 4PPRPa (Fig. 2.28b) P\\P||R\\Pa (Fig. 2.21f) The third joints of the four limbs have parallel axes 7. 4PRPPa (Fig. 2.29a) P\\R||P\\Pa (Fig. 2.21g) Idem No. 5 8. 4PRPaP (Fig. 2.29b) P\\R\\Pa \\kP (Fig. 2.21h) Idem No. 5 9. 4PPPaR (Fig. 2.30a) P\\P\\kPa\\kR (Fig. 2.21i) Idem No. 4 10. 4PPPaR (Fig. 2.30b) P\\P||Pa\\R (Fig. 2.21j) Idem No. 4 11. 4PPaRP (Fig. 2.31a) P||Pa\\R||P (Fig. 2.21k) The before last joints of the four limbs have parallel axes 12. 4PPaPR (Fig. 2.31b) P||Pa\\P|| R (Fig. 2.21l) Idem No. 4 13. 4PPaPR (Fig. 2.32a) P\\Pa||P\\\\R (Fig. 2.21m) Idem No. 4 14. 4PPaRP (Fig. 2.32b) P\\Pa\\kR\\kP (Fig. 2.21n) Idem No. 4 15. 4RPPaP (Fig. 2.33a) R\\P||Pa\\kP (Fig. 2.21o) The first revolute joints of the four limbs have parallel axes 16. 4RPPPa (Fig. 2.33b) R||P\\P||Pa (Fig. 2.21p) Idem No. 15 17. 4PaPPR (Fig. 2.34a) Pa||P\\P||R (Fig. 2.21q) Idem No. 4 18. 4PaPPR (Fig. 2.34b) Pa\\P\\kP\\kR (Fig. 2.21r) Idem No. 4 19. 4PaRPP (Fig. 2.35a) Pa\\R||P\\P (Fig. 2.21s) Idem No. 11 20" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_35_0001419_978-94-007-7401-8-Figure6.19-1.png", + "original_path": "designv11-35/openalex_figure/designv11_35_0001419_978-94-007-7401-8-Figure6.19-1.png", + "caption": "Fig. 6.19 Maximaly regular fully-parallel PMs with Sch\u00f6nflies motions of types 2CPRR-1CPR1CPaPa (a) and 2CPRR-1CPR-1CPaPat (b) defined by MF = SF = 4, (RF) = (v1; v2; v3;xb), TF = 0, NF = 10, limb topology C\\P\\||R\\R, C\\P\\||R and C||Pa||Pa (a), C||Pa||Pat (b)", + "texts": [ + "17b) P ?P\\\\P ?||R (Fig. 4.1a) P ?P\\\\P\\\\R (Fig. 4.1b) CPa||Pat (Fig. 5.4p) Idem no. 1 3. 2PPRRR-1PPRR-1CPaPa (Fig. 6.18a) P ?P ?||R||R ?R (Fig. 4.2a) P ?P ?||R||R (Fig. 4.1d) C||Pa||Pa (Fig. 5.4o) Idem no. 1 (continued) 6.2 Fully-Parallel Topologies with Simple and Complex Limbs 613 Table 6.6 (continued) No. PM type Limb topology Connecting conditions 4. 2PPRRR-1PPRR-1CPaPat (Fig. 6.18b) P ?P ?||R||R ?R (Fig. 4.2a) P ?P ?||R||R (Fig. 4.1d) C||Pa||Pat (Fig. 5.4p) Idem no. 1 5. 2CPRR-1CPR-1CPaPa (Fig. 6.19a) C ?P ?||R ?R (Fig. 4.2e) C ?P ?||R (Fig. 4.1i) C||Pa||Pa (Fig. 5.4o) Idem no. 1 6. 2CPRR-1CPR-1CPaPat (Fig. 6.19b) C ?P ?||R ?R (Fig. 4.2e) C ?P ?||R (Fig. 4.1i) C||Pa||Pat (Fig. 5.4p) Idem no. 1 7. 2CRRR-1CRR-1CPaPa (Fig. 6.20a) C||R||R ?R (Fig. 4.2g) C||R||R (Fig. 4.1k) C||Pa||Pa (Fig. 5.4o) Idem no. 1 8. 2CRRR-1CRR-1CPaPat (Fig. 6.20b) C||R||R ?R (Fig. 4.2g) C||R||R (Fig. 4.1k) C||Pa||Pat (Fig. 5.4p) Idem no. 1 9. 3PPPaR-1RPPP (Fig. 6.21a) P ?P ?||Pa ?||R (Fig. 4.8b) P ?P ?||Pa||R, (Fig. 4.8c) R ?P\\\\P\\\\P (Fig. 5.1a) The last joints of the four limbs have superposed axes/ directions. The actuated prismatic joints of limbs G1, G2 and G3 have orthogonal directions 10" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_35_0000387_tmag.2021.3079978-Figure9-1.png", + "original_path": "designv11-35/openalex_figure/designv11_35_0000387_tmag.2021.3079978-Figure9-1.png", + "caption": "Fig. 9 (a) Stator with round wire field winding and rectangular wire armature winding. (b) Experimental platform of the DSBLM drive system.", + "texts": [ + " It shows that the conduction loss of BLDC driving mode is greater than that of BLAC driving mode, while the switching loss of BLDC driving mode is less than that of BLAC driving mode. In general, the total converter loss in the BLAC driving mode is less than that in the BLDC driving mode, and the converter loss is approximately proportional to the effective value of phase current in the two driving modes. At the same time, the simulation analysis also shows that the converter loss is not affected by the motor speed basically. Fig. 9(a) shows the stator of the DSBLM prototype. The DC field winding adopts round wire and the armature winding adopts rectangular wire. The experimental platform of the DSBLM drive system has been built, as shown in Fig. 9(b). In order to compare the loss characteristics of DSBLM with rectangular wire armature winding in BLDC and BLAC driving modes, the experiment of the two driving modes has been carried out under the conditions of the same speed, excitation current and generator load. Fig. 10 shows the experiment waveforms of phase current in BLDC and BLAC driving modes at 2000r/min. Since the generator load is same, the difference of the motor input power can be considered as the difference of the motor loss between the two driving modes" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_35_0001419_978-94-007-7401-8-Figure6.30-1.png", + "original_path": "designv11-35/openalex_figure/designv11_35_0001419_978-94-007-7401-8-Figure6.30-1.png", + "caption": "Fig. 6.30 2PPaRRR-1PPaRR-1RPaPaP-type maximaly regular fully-parallel PM with Sch\u00f6nflies motions defined by MF = SF = 4, (RF) = (v1; v2; v3;xb), TF = 0, NF = 19, limb topology P||Pa||R||R\\R, P||Pa||R||R\\R and R||Pa||Pa||P", + "texts": [ + "4k) Idem no. 1 3. 3PPPaR-1RPaPatP (Fig. 6.28a) P\\P\\||Pa\\||R (Fig. 4.8b) P\\P\\||Pa||R, (Fig. 4.8c) R||Pa||Pat||P (Fig. 5.4l) Idem no. 1 4. 3PPPaR-1RPaPatP (Fig. 6.28b) P\\P\\||Pa\\R (Fig. 4.8a) P\\P\\||Pa||R, (Fig. 4.8c) R||Pa||Pat||P (Fig. 5.4l) Idem no. 1 5. 3PPaPR-1RPaPaP (Fig. 6.29a) P||Pa\\P\\\\R (Fig. 4.8d) P||Pa\\P\\||R (Fig. 4.8e) R||Pa||Pa||P (Fig. 5.4k) Idem no. 1 6. 3PPaPR-1RPaPatP (Fig. 6.29b) P||Pa\\P\\\\R (Fig. 4.8d) P||Pa\\P\\||R (Fig. 4.8e) R||Pa||Pat||P (Fig. 5.4l) Idem no. 1 7. 2PPaRRR-1PPaRR-1RPaPaP (Fig. 6.30) P||Pa||R||R\\R (Fig. 4.9a) P||Pa||R||R (Fig. 4.8g) R||Pa||Pa||P (Fig. 5.4k) Idem no. 1 8. 2PPaRRR-1PPaRR1RPaPatP (Fig. 6.31) P||Pa||R||R\\R (Fig. 4.9a) P||Pa||R||R (Fig. 4.8g) R||Pa||Pat||P (Fig. 5.4l) Idem no. 1 9. 3PPPaR-1RPPaPa (Fig. 6.32a) P\\P\\||Pa\\||R (Fig. 4.8b) P\\P\\||Pa||R, (Fig. 4.8c) R||P||Pa||Pa (Fig. 5.4m) The last revolute joints of limbs G1, G2 and G3 and the revolute joint between links 7D and 9 have superposed axes. The actuated prismatic joints of limbs G1, G2 and G3 have orthogonal directions 10" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_35_0001419_978-94-007-7401-8-Figure2.110-1.png", + "original_path": "designv11-35/openalex_figure/designv11_35_0001419_978-94-007-7401-8-Figure2.110-1.png", + "caption": "Fig. 2.110 Fully-parallel PMs with coupled Sch\u00f6nflies motions of types 4RPRRPa (a) and 4PRPaRR (b) defined by MF \u00bc SF \u00bc 4; RF\u00f0 \u00de \u00bc \u00f0v1; v2; v3;xd\u00de, TF = 0, NF = 14, limb topology R\\P\\kR\\R||Pa (a) and P\\R\\Pa\\kR\\R (b)", + "texts": [], + "surrounding_texts": [ + "164 2 Fully-Parallel Topologies with Coupled Sch\u00f6nflies Motions", + "2.2 Topologies with Complex Limbs 165", + "166 2 Fully-Parallel Topologies with Coupled Sch\u00f6nflies Motions" + ] + }, + { + "image_filename": "designv11_35_0001419_978-94-007-7401-8-Figure2.71-1.png", + "original_path": "designv11-35/openalex_figure/designv11_35_0001419_978-94-007-7401-8-Figure2.71-1.png", + "caption": "Fig. 2.71 4CPaPa-type fully-parallel PM with coupled Sch\u00f6nflies motions defined by MF \u00bc SF \u00bc 4; RF\u00f0 \u00de \u00bc \u00f0v1; v2; v3;xd\u00de, TF = 0, NF = 30, limb topology C\\Pa\\\\Pa", + "texts": [ + "22x) Idem No. 33 52. 4PaPaPR (Fig. 2.64) Pa\\Pa\\\\P||R (Fig. 2.22y) Idem No. 4 53. 4PaPaRP (Fig. 2.65) Pa\\Pa\\\\R||P (Fig. 2.22z) Idem No. 41 54. 4RPaPaP (Fig. 2.66) R\\Pa\\\\Pa\\\\P (Fig. 2.22a0) Idem No. 15 55. 4RPaPPa (Fig. 2.67) R\\Pa\\kP\\\\Pa (Fig. 2.22b0) Idem No. 15 56. 4RPPaPa (Fig. 2.68) R||P\\Pa\\\\Pa (Fig. 2.22c0) Idem No. 15 57. 4PRPaPa (Fig. 2.69) P||R\\Pa\\\\Pa (Fig. 2.22d0) Idem No. 5 58. 4PaCPa (Fig. 2.70) Pa\\C\\Pa (Fig. 2.22e0) The axes of the cylindrical joints of the four limbs are parallel 59. 4CPaPa (Fig. 2.71) C\\Pa\\\\Pa (Fig. 2.22f 0) Idem No. 58 60. 4PaPaC (Fig. 2.72) Pa\\Pa\\\\C (Fig. 2.22g0) Idem No. 58 176 2 Fully-Parallel Topologies with Coupled Sch\u00f6nflies Motions T ab le 2. 4 L im b to po lo gy an d co nn ec ti ng co nd it io ns of th e fu ll y- pa ra ll el so lu ti on s w it h no id le m ob il it ie s pr es en te d in F ig s. 2. 73 , 2. 74 , 2. 75 , 2. 76 , 2. 77 , 2. 78 ,2 .7 9, 2. 80 ,2 .8 1, 2. 82 ,2 .8 3, 2. 84 ,2 .8 5, 2. 86 ,2 .8 7, 2. 88 ,2 .8 9, 2. 90 ,2 .9 1, 2. 92 ,2 .9 3, 2. 94 ,2 .9 5, 2" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_35_0001419_978-94-007-7401-8-Figure2.70-1.png", + "original_path": "designv11-35/openalex_figure/designv11_35_0001419_978-94-007-7401-8-Figure2.70-1.png", + "caption": "Fig. 2.70 4PaCPa-type fully-parallel PM with coupled Sch\u00f6nflies motions defined by MF \u00bc SF \u00bc 4; RF\u00f0 \u00de \u00bc \u00f0v1; v2; v3;xd\u00de, TF = 0, NF = 30, limb topology Pa\\C\\Pa", + "texts": [ + " 34 50. 4PaRPPa (Fig. 2.62) Pa\\R||P\\Pa (Fig. 2.22w) Idem No. 34 51. 4PaPRPa (Fig. 2.63) Pa\\P||R\\Pa (Fig. 2.22x) Idem No. 33 52. 4PaPaPR (Fig. 2.64) Pa\\Pa\\\\P||R (Fig. 2.22y) Idem No. 4 53. 4PaPaRP (Fig. 2.65) Pa\\Pa\\\\R||P (Fig. 2.22z) Idem No. 41 54. 4RPaPaP (Fig. 2.66) R\\Pa\\\\Pa\\\\P (Fig. 2.22a0) Idem No. 15 55. 4RPaPPa (Fig. 2.67) R\\Pa\\kP\\\\Pa (Fig. 2.22b0) Idem No. 15 56. 4RPPaPa (Fig. 2.68) R||P\\Pa\\\\Pa (Fig. 2.22c0) Idem No. 15 57. 4PRPaPa (Fig. 2.69) P||R\\Pa\\\\Pa (Fig. 2.22d0) Idem No. 5 58. 4PaCPa (Fig. 2.70) Pa\\C\\Pa (Fig. 2.22e0) The axes of the cylindrical joints of the four limbs are parallel 59. 4CPaPa (Fig. 2.71) C\\Pa\\\\Pa (Fig. 2.22f 0) Idem No. 58 60. 4PaPaC (Fig. 2.72) Pa\\Pa\\\\C (Fig. 2.22g0) Idem No. 58 176 2 Fully-Parallel Topologies with Coupled Sch\u00f6nflies Motions T ab le 2. 4 L im b to po lo gy an d co nn ec ti ng co nd it io ns of th e fu ll y- pa ra ll el so lu ti on s w it h no id le m ob il it ie s pr es en te d in F ig s. 2. 73 , 2. 74 , 2. 75 , 2. 76 , 2. 77 , 2. 78 ,2 .7 9, 2. 80 ,2 " + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_35_0000181_s10409-021-01064-4-Figure6-1.png", + "original_path": "designv11-35/openalex_figure/designv11_35_0000181_s10409-021-01064-4-Figure6-1.png", + "caption": "Fig. 6 a, i Geometrical dimensions of the workpiece used by Treloar [19] for uniaxial tension experiments, a,ii details of the one quarter 3D model used in FE simulation, b the meshed 3D model, and c the geometrical details of the workpiece used by Fujikawa et\u00a0al. [41] for equibiaxial loading and its meshed one quarter 3D model", + "texts": [ + " It is worth noting that for hyperelastic materials in Abaqus, the stress should be returned in a co-rotational coordinate system in which the basis system rotates with the material. For this reason, the stress must be rotated back before updating the new stresses as shown in the second last step of the algorithm. To compare the simulation and the experimental results for uniaxial loading, the workpiece geometry was modeled according to the dimensions of the specimen used by Treloar [19] in his experiments (see Fig.\u00a06a, i). The simulation set-up was simplified by modeling only one-quarter of the workpiece as shown in Fig.\u00a06a, ii and utilizing symmetry boundary conditions. The model was meshed with a total of 6000 full integration 8-node linear brick elements 1 3 (C3D8) as shown in Fig.\u00a06b. As for the equibiaxial loading, a notched workpiece proposed by Fujikawa et\u00a0al. [41] in their experiments with geometrical details given in Fig.\u00a06c was used. According to the authors, the notched workpiece helps mitigate the non-uniform deformation between the clamps. The model parameters for the equibiaxial loading ( 1 = 801198.03, 2 = 98.58, 3 = \u2212527655.25, K = 5 \u00d7 108 ) were obtained by fitting the modified strain energy equation to the strain energy density data calculated from the experimental uniaxial tension data by the authors. For each simulation set-up, a reference point was set to capture both the displacement and force in the direction of the deformation" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_35_0001419_978-94-007-7401-8-Figure2.13-1.png", + "original_path": "designv11-35/openalex_figure/designv11_35_0001419_978-94-007-7401-8-Figure2.13-1.png", + "caption": "Fig. 2.13 Fully-parallel PMs with coupled Sch\u00f6nflies motions of types 4RRPRR (a) and 4RPRRR (b) defined by MF \u00bc SF \u00bc 4; RF\u00f0 \u00de \u00bc \u00f0v1; v2; v3;xd\u00de, TF = 0, NF = 2, limb topology R\\R||P||R||R (a) and R\\P||R||R||R (b)", + "texts": [ + " 2.1o) Idem No. 13 17. 4RPRRR (Fig. 2.10a) R||P||R||R\\R (Fig. 2.1p) Idem No. 13 18. 4RRPRR (Fig. 2.10b) R||R||P||R\\R (Fig. 2.1q) Idem No. 13 19. 4RRRRP (Fig. 2.11a) R||R||R\\R\\P (Fig. 2.1r) The before last revolute joints of the four limbs have parallel axes 20. 4RPRRR (Fig. 2.11b) R\\P\\kR\\R||R (Fig. 2.1s) Idem No. 3 21. 4RRRRP (Fig. 2.12a) R\\R||R||R||P (Fig. 2.1t) The first revolute joints of the four limbs have parallel axes 22. 4RRRPR (Fig. 2.12b) R\\R||R||P||R Fig. 2.1u) Idem No. 21 23. 4RRPRR (Fig. 2.13a) R\\R||P||R||R (Fig. 2.1v) Idem No. 21 24. 4RPRRR (Fig. 2.13b) R\\P||R||R||R (Fig. 2.1w) Idem No. 21 25. 4RRRPR (Fig. 2.14a) R\\R||R||P||R (Fig. 2.1x) Idem No. 21 26. 4RRPRR (Fig. 2.14b) R\\R||P||R||R||R (Fig. 2.1y) Idem No. 21 27. 4RPRRP (Fig. 2.15a) R\\P\\kR\\R\\P (Fig. 2.1z) Idem No. 19 28. 4PRRRP (Fig. 2.15b) P\\R||R\\R\\P (Fig. 2.1z1) Idem No. 19 29. 4PRRPR (Fig. 2.16) P\\R\\R\\P\\kR (Fig. 2.1a0) Idem No. 5 30. 4CRRR (Fig. 2.17a) C\\R||R\\R (Fig. 2.1d0) Idem No. 1 (continued) 2.1 Topologies with Simple Limbs 59 Table 2.1 (continued) No. PM type Limb topology Connecting conditions 31" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_35_0001516_cp.2016.0271-Figure3-1.png", + "original_path": "designv11-35/openalex_figure/designv11_35_0001516_cp.2016.0271-Figure3-1.png", + "caption": "Fig. 3. 5-phase optimized permanent magnet synchronous machine", + "texts": [ + " The designed machines have to satisfy both specifications, the electro-mechanical specification and the DC bus voltage constraint. Several structures have been optimized and tested to satisfy both criteria. The final choice has been made on a permanent magnet synchronous machine with an optimized copper repartition on the stator and a combination between flux focusing permanent magnet configuration and buried magnet on the rotor. In the Figure 2, the 3-phase optimized permanent magnet synchronous machine and in the Figure 3, the corresponding 5-phase optimized permanent magnet synchronous machine. Please notice that the 3-phase machine was initially designed and optimized, and from this initial design, the 5- phase machine was deduced and re-optimized. Firstly, the number of pole pairs was identical for both machines, but the performances reached by the 5-phase were reduced. Then, the number of pole pairs was increased (from 2 to 4 pole pairs) in order to guarantee the performances. Both machines have the same dimensions (stator external radius, air gap, stator internal radius), however the 3-phase machine has 8 wires by phase (2 in serie composed by 4 in parallel) and the 5-phase machine has 4 wires (2 in serie composed by 2 in parallel)" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_35_0001419_978-94-007-7401-8-Figure2.32-1.png", + "original_path": "designv11-35/openalex_figure/designv11_35_0001419_978-94-007-7401-8-Figure2.32-1.png", + "caption": "Fig. 2.32 Fully-parallel PMs with coupled Sch\u00f6nflies motions of types 4PPaPR (a) and 4PPaRP (b) defined by MF \u00bc SF \u00bc 4; RF\u00f0 \u00de \u00bc \u00f0v1; v2; v3;xd\u00de, TF = 0, NF = 18, limb topology P\\Pa||P\\\\R (a) and P\\Pa\\kR\\kP (b)", + "texts": [ + " 4PPRPa (Fig. 2.28b) P\\P||R\\Pa (Fig. 2.21f) The third joints of the four limbs have parallel axes 7. 4PRPPa (Fig. 2.29a) P\\R||P\\Pa (Fig. 2.21g) Idem No. 5 8. 4PRPaP (Fig. 2.29b) P\\R\\Pa \\kP (Fig. 2.21h) Idem No. 5 9. 4PPPaR (Fig. 2.30a) P\\P\\kPa\\kR (Fig. 2.21i) Idem No. 4 10. 4PPPaR (Fig. 2.30b) P\\P||Pa\\R (Fig. 2.21j) Idem No. 4 11. 4PPaRP (Fig. 2.31a) P||Pa\\R||P (Fig. 2.21k) The before last joints of the four limbs have parallel axes 12. 4PPaPR (Fig. 2.31b) P||Pa\\P|| R (Fig. 2.21l) Idem No. 4 13. 4PPaPR (Fig. 2.32a) P\\Pa||P\\\\R (Fig. 2.21m) Idem No. 4 14. 4PPaRP (Fig. 2.32b) P\\Pa\\kR\\kP (Fig. 2.21n) Idem No. 4 15. 4RPPaP (Fig. 2.33a) R\\P||Pa\\kP (Fig. 2.21o) The first revolute joints of the four limbs have parallel axes 16. 4RPPPa (Fig. 2.33b) R||P\\P||Pa (Fig. 2.21p) Idem No. 15 17. 4PaPPR (Fig. 2.34a) Pa||P\\P||R (Fig. 2.21q) Idem No. 4 18. 4PaPPR (Fig. 2.34b) Pa\\P\\kP\\kR (Fig. 2.21r) Idem No. 4 19. 4PaRPP (Fig. 2.35a) Pa\\R||P\\P (Fig. 2.21s) Idem No. 11 20. 4PaPRP (Fig. 2.35b) Pa||P\\R||P (Fig. 2.21t) Idem No. 5 21. 4PaPRP (Fig. 2.36a) Pa\\P||R\\P (Fig. 2.21u) Idem No. 11 22" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_35_0001603_301-Figure3-1.png", + "original_path": "designv11-35/openalex_figure/designv11_35_0001603_301-Figure3-1.png", + "caption": "Fig. 3. Crossed cylinder friction apparatus AA\u2019, cylindrical specimens; B, turntable; C, direction of sliding; D, friction measuring element; E, point of application", + "texts": [ + " From this it follows that the frictional forces at the onset of sliding are typical of contaminated surfaces, and that as sliding proceeds the coefficient of friction in general changes. Different frictional phenomena and different degrees of surface damage are therefore to be expected in mechanical systems in which the sliding distance is short, as with a pair of spur gears, from those in which it is long, as with a piston in a cylinder. The effect of sliding distance is conveniently studied in a friction apparatus in which the specimens are cylinders, one inclined at 45\u2019 to the direction of sliding, the other being mounted on a turntable (Fig. 3). The sliding distance is varied by rotating the turntable. Such an apparatus has been designed by Dr. Archard and results obtained with Friction records corresponding to the specimen arrangements shown at A , B and C. With A and C the sliding distance is long, wth,B it is short. Upper specimen 0.4% carbon steel; lower specimen 1 \u2018 5 ;4 nickel, 1 % chromium steel unlubricatcd. Speed 0.05 cmis. Loads: A , 10 kg; B, 7 . 5 to 25 kg: C. 10 kg. it show that the influence of the sliding distance can indeed be very great (Fig" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_35_0000690_j.jmst.2021.05.038-Figure7-1.png", + "original_path": "designv11-35/openalex_figure/designv11_35_0000690_j.jmst.2021.05.038-Figure7-1.png", + "caption": "Fig. 7. Schematic diagram of the top and bottom of the LSFed sample under different heat input conditions: (a) Melt pool & Heat affected zone and (b) thermal cycle curve.", + "texts": [ + " After many thermal cycles during the deposition process, he bainite gradually transformed into a mixed ferrite and carbide tructure. By comparing the three groups of samples, it can be seen that nder different heat input conditions, the same material (34CrN- Mo6) exhibits different characteristics after LSF. To clarify the in- uence of heat input on the microstructure, three schematic diarams of the heat-affected zone of the LSFed samples fabricated nder different dimensionless heat inputs, and the thermal cycle urves are drawn. Fig. 7 (a) divides the heat-affected zone into a ully austenitized zone and a tempering zone [28] . The structure f the fully austenitized zone was mainly affected by the solidifiation and cooling rate of the molten pool, whereas the structure f the tempering zone was related to the temperature experienced uring the thermal cycle. For the samples in group A ( Q \u2217 = 1) with he lowest heat input, although the size of the molten pool and eat-affected zone at the top of the sample was larger than that at he bottom, its spread was still not noticeable, except in the temering zone. For group B ( Q \u2217 = 1.9) with a higher heat input, the ize of the molten pool and the range of the heat-affected zone ncreased significantly. It was wider and deeper when the heat in- ut reached the condition of group C ( Q \u2217 = 2.9). With a low heat nput, the difference in the thermal cycle curve between the botom and the top of the sample was insignificant under the cold ubstrate and the faster cooling rate, as shown in Fig. 7 (b). Then, ith increasing heat input, the bottom of the sample experienced longer thermal cycle time and higher temperature extrema (both he peak and valley temperatures increased). The top of the sample howed a rapid increase and a rapid decrease owing to the lack of ubsequent heat input. .2. Microhardness of the LSFed samples with different Q \u2217 Fig. 8 shows the hardness distribution of the LSFed samples abricated under different heat input conditions. From the top to he middle, the microhardness of the three groups of samples had similar variation: gradually decreasing from the top toward the iddle, after which it was stable; the average value decreased as Q 9 d c a s t i a 6 v w s 3 m t s p l f b c w t t a u c S 2 o w t m t t a 3 c s p h m p F u p u t \u2217 increased" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_35_0001419_978-94-007-7401-8-Figure2.12-1.png", + "original_path": "designv11-35/openalex_figure/designv11_35_0001419_978-94-007-7401-8-Figure2.12-1.png", + "caption": "Fig. 2.12 Fully-parallel PMs with coupled Sch\u00f6nflies motions of types 4RRRRP (a) and 4RRRPR (b) defined by MF \u00bc SF \u00bc 4; RF\u00f0 \u00de \u00bc \u00f0v1; v2; v3;xd\u00de, TF = 0, NF = 2, limb topology R\\R||R||R||P (a) and R\\R||R||P||R (b)", + "texts": [ + "1l) The last revolute joints of the four limbs have parallel axes 14. 4RPRRR (Fig. 2.8b) R||P||R||R\\R (Fig. 2.1m) Idem No. 13 15. 4RRPRR (Fig. 2.9a) R||R||P||R\\R (Fig. 2.1n) Idem No. 13 16. 4RRRPR (Fig. 2.9b) R||R||R||P\\R (Fig. 2.1o) Idem No. 13 17. 4RPRRR (Fig. 2.10a) R||P||R||R\\R (Fig. 2.1p) Idem No. 13 18. 4RRPRR (Fig. 2.10b) R||R||P||R\\R (Fig. 2.1q) Idem No. 13 19. 4RRRRP (Fig. 2.11a) R||R||R\\R\\P (Fig. 2.1r) The before last revolute joints of the four limbs have parallel axes 20. 4RPRRR (Fig. 2.11b) R\\P\\kR\\R||R (Fig. 2.1s) Idem No. 3 21. 4RRRRP (Fig. 2.12a) R\\R||R||R||P (Fig. 2.1t) The first revolute joints of the four limbs have parallel axes 22. 4RRRPR (Fig. 2.12b) R\\R||R||P||R Fig. 2.1u) Idem No. 21 23. 4RRPRR (Fig. 2.13a) R\\R||P||R||R (Fig. 2.1v) Idem No. 21 24. 4RPRRR (Fig. 2.13b) R\\P||R||R||R (Fig. 2.1w) Idem No. 21 25. 4RRRPR (Fig. 2.14a) R\\R||R||P||R (Fig. 2.1x) Idem No. 21 26. 4RRPRR (Fig. 2.14b) R\\R||P||R||R||R (Fig. 2.1y) Idem No. 21 27. 4RPRRP (Fig. 2.15a) R\\P\\kR\\R\\P (Fig. 2.1z) Idem No. 19 28. 4PRRRP (Fig. 2.15b) P\\R||R\\R\\P (Fig. 2.1z1) Idem No. 19 29. 4PRRPR (Fig. 2.16) P\\R\\R\\P\\kR (Fig. 2.1a0) Idem No. 5 30. 4CRRR (Fig. 2.17a) C\\R||R\\R (Fig. 2" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_35_0000395_978-3-030-67750-3_16-Figure30.7-1.png", + "original_path": "designv11-35/openalex_figure/designv11_35_0000395_978-3-030-67750-3_16-Figure30.7-1.png", + "caption": "Fig. 30.7 Side vie of the wheelbarrow loader bucket", + "texts": [ + " The simple 2D sketches equipped with the specifications and parts from the data collected in the previous design method process had been analyzed later by using a weighted objective analysis method. Therefore, the final design concept is available after finishing up all the designs steps. The final conceptual design was made by using SolidWorks, Fig. 30.6a displays the final conceptual design of the wheelbarrow loader bucket, while Fig. 30.6b is the exploded view of the design. The side view of wheelbarrow bucket loader is shown in Fig. 30.7. The design is based on the existing wheelbarrow but has a modification by using the six-point linkage based on the existing loader bucket design. The pneumatic actuator attached at the loader linkages is to make ease the loading and unloading process of the bucket. Themainmaterial used for the loader bucket is using Aluminiumwhile the wheelbarrow body is maintains to use the existing material, which is steel. The wheelbarrow loader bucket is attached to a loader linkage that is attached to shafts that connect to a linear actuator that acts as the arm for the bucket to move downwards or upwards" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_35_0001419_978-94-007-7401-8-Figure5.18-1.png", + "original_path": "designv11-35/openalex_figure/designv11_35_0001419_978-94-007-7401-8-Figure5.18-1.png", + "caption": "Fig. 5.18 3PaPPR-1RPPaPat-type fully-parallel PMs with uncoupled Sch\u00f6nflies motions defined by MF = SF = 4, RF\u00f0 \u00de \u00bc \u00f0v1; v2; v3;xb\u00de, TF = 0, NF = 21, limb topology R||P||Pa||Pat and Pa||P\\P\\kR, Pa||P\\P\\\\R (a), Pa\\P\\kP\\kR, Pa\\P\\kP\\\\R (b)", + "texts": [ + "2a) The revolute joint between links 7D and 9 of G4 limb and the last revolute joints of limbs G1, G2 and G3 have superposed axes Pa||P\\P\\\\R (Fig. 5.2d) R||P||Pa||Pa (Fig. 5.4m) 4. 3PaPPR-1RPPaPa (Fig. 5.16b) Pa\\P\\kP\\kR (Fig. 5.2b) Idem No. 3 Pa\\P\\kP\\\\R (Fig. 5.2c) R||P||Pa||Pa (Fig. 5.4m) 5. 3PaPPR1RPaPatP (Fig. 5.17a) Pa||P\\P\\kR (Fig. 5.2a) Idem No. 1 Pa||P\\P\\\\R (Fig. 5.2d) R||Pa||Pat||P (Fig. 5.4l) 6. 3PaPPR1RPaPatP (Fig. 5.17b) Pa\\P\\kP\\kR (Fig. 5.2b) Idem No. 1 Pa\\P\\kP\\\\R (Fig. 5.2c) R||Pa||Pat||P (Fig. 5.4l) 7. 3PaPPR-1RPPaPat (Fig. 5.18a) Pa||P\\P\\kR (Fig. 5.2a) Idem No. 3 Pa||P\\P\\\\R (Fig. 5.2d) R||P||Pa||Pat (Fig. 5.4n) 8. 3PaPPR-1RPPaPat (Fig. 5.18b) Pa\\P\\kP\\kR (Fig. 5.2b) Idem No. 3 Pa\\P\\kP\\\\R (Fig. 5.2c) R||P||Pa||Pat (Fig. 5.4n) 9. 2PaPRRR-1PaPRR1RPaPaP (Fig. 5.19a) Pa\\P\\R||R\\\\R (Fig. 5.3a) The last joints of the four limbs have superposed axes/directions. The revolute joints between links 5 and 6 of limbs G1, G2 and G3 have orthogonal axes Pa\\P\\\\R||R (Fig. 5.2e) R||Pa||Pa||P (Fig. 5.4k) 10. 2PaPRRR-1PaPRR1RPaPaP (Fig. 5.19b) Pa||P\\R||R\\\\R (Fig. 5.3b) Idem No. 9 Pa||P\\R||R (Fig. 5.2f) R||Pa||Pa||P (Fig. 5.4k) (continued) 528 5 Topologies with Uncoupled Sch\u00f6nflies Motions Table 5" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_35_0000089_tmag.2021.3060767-Figure9-1.png", + "original_path": "designv11-35/openalex_figure/designv11_35_0000089_tmag.2021.3060767-Figure9-1.png", + "caption": "Fig. 9. Manufactured PMSG prototype: shaft connected PMSG prototype.", + "texts": [ + " A rectangular tooth surface meshing is considered to calculate the radial force vectors. Two scenarios, namely, no-load and Id = 120.2 A loading case, are shown in Fig. 8. Id = 120.2 A is considered as it produces the highest eighth-order radial force density harmonic. As shown in Fig. 8, compared with the no-load, for Id = 120.2 A loading, the radial force increases by 50%. With this idea in mind, in the following section, a prototype of the FCPMSG is developed and tested to validate its performance. The prototype of the FCPMSG along with its test bed is shown in Fig. 9. The prototype is developed using the dimensions listed in Table I. The stator and rotor used Pohang Iron and Steel Company (POSCO) Steel\u2019s laminated steel plate 50PN400 and the PM used is NdFeB with a residual flux density of 1.25 T and relative permeability of 1.05. The windings use an AWG 14-grade enamel copper wire. The dynamometer used for the testing is a three-phase induction motor with a rated output of 150 kW and a maximum rotational speed of 3500 r/min. The test is performed under no-load and loadcondition" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_35_0000690_j.jmst.2021.05.038-Figure1-1.png", + "original_path": "designv11-35/openalex_figure/designv11_35_0000690_j.jmst.2021.05.038-Figure1-1.png", + "caption": "Fig. 1. Schematic diagram of the LSF process & the power density distribution at the horizontal position under different \u0192 values.", + "texts": [ + " The power density disributions of these heat sources often follow axisymmetric Gausian profiles [19] . d = f P \u03c0 r 2 b exp ( \u2212 f r 2 r 2 b ) (1) here \u0192 is the distribution factor, P is the total power of the heat ource, r b is the radius of the heat source and r is the radial disance of any point from the axis of the heat source. Eq. (1) indiates the power density distribution of the heat source on the surace. A higher value of \u0192 indicates higher power density at the heat ource axis and vice versa, and a larger r b indicates lower power ensity at radial locations and vice versa, as shown Fig. 1 . The power density distribution can also be uniform, depending n the nature of the heat source [22\u201324] . According to Eq. (1) , the ower density can be adjusted by controlling the value of r b . Under he condition of uniform power density, the heat input arc section f the molten pool can be simulated as a rectangle, as shown in ig. 1 . The heat input per unit bead length ( Q (J/mm)), as defined n ISO 25901-4, was determined using the following equation. The quivalent deformation of this formula can be used to obtain the eat input value SE (specific energy per unit area)" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_35_0001579_2015-24-2526-Figure1-1.png", + "original_path": "designv11-35/openalex_figure/designv11_35_0001579_2015-24-2526-Figure1-1.png", + "caption": "Figure 1. Engine mass balancing unit (MBU)", + "texts": [ + " \u2022 How to select gear parameters (in case of a gear drive) - number of teeth, backlash, helix angle (axial vibration component)? \u2022 Include scissors or damper gears to reduce rattle, w/o badly affect gear tooth durability? In this paper this process is discussed for a FIAT I4 Diesel engine under development. The engine is equipped with a standard Lanchester drive, including scissors gear between the crankshaft and the first balancer shaft. The shafts are supported by roller bearings to reduce losses. The first balancer shaft is driven by gears from the second crankshaft web (Figure 1), close to the engine front, as a compromise between bending and torsional vibration problems and enabling simple packaging and maintenance procedure. This chapter describes the engineering tasks for analyzing a balancer drive. This work is started as integrated part of the crank train design analysis. The investigation starts with a kinematic analysis for designing the crankshaft counter weights and the MBU parameters (mass/inertia, position) as pure rigid system. This approach is used for the baseline layout and it solves vibration problems as kinematic problem neglecting interaction mechanisms between crank train components and real operating condition" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_35_0000648_2050-7038.13001-Figure5-1.png", + "original_path": "designv11-35/openalex_figure/designv11_35_0000648_2050-7038.13001-Figure5-1.png", + "caption": "FIGURE 5 (A) Starting pulses generated from initial noise in place of back EMF signals. (B) Frequency variation of pulses during startup", + "texts": [ + " During the instant tst when the pulse frequency fp is equal to synchronous mode startup frequency fs as well as when pulse frequency ramp up rate, dfp/dt is nearly equal to synchronous mode allowable ramp up rate for startup dfs/ dt, the enable pulse is released for switching on DC supply. Rotor starts to move and at time t1st, rotor locks with field, and at time t2st, regular commutation continues after startup and the pulse frequency of the pulse contributed by one of line-to-line back EMF reduces to twice of fundamental frequency of back EMF signal. In order to get higher pulse frequency to limit the startup current, the threshold band is reduced gradually at startup. Figure 5 (a) shows the noise signal and the threshold band for generating the commutation pulses. Figure 5(b) shows the variation of frequency with time during startup. Variation in the commutation angle has greater impact on the current wave shape. The waveshapes of phase currents for lagged, precise and advanced commutation modes are shown in Figure 6. Here, t0 is the time of precise commutation; t1 is the actual time of start of commutation in first 600 portion of phase current, tp1 is the point of minimum current in this region; t2 is the start time of next commutation sequence. The time deviation from precise commutation is td" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_35_0001419_978-94-007-7401-8-Figure5.12-1.png", + "original_path": "designv11-35/openalex_figure/designv11_35_0001419_978-94-007-7401-8-Figure5.12-1.png", + "caption": "Fig. 5.12 2PaPRRR-1PaPRR-1RUPU-type fully-parallel PMs with uncoupled Sch\u00f6nflies motions defined by MF = SF = 4, RF\u00f0 \u00de \u00bc \u00f0v1; v2; v3;xb\u00de, TF = 0, NF = 11, limb topology R\\R\\R\\P\\kR\\R and Pa\\P\\\\R||R\\\\R, Pa\\P\\\\R||R (a), Pa||P\\R||R\\\\R, Pa||P\\R||R (b)", + "texts": [ + "1 Fully-Parallel Topologies 523 524 5 Topologies with Uncoupled Sch\u00f6nflies Motions 5.1 Fully-Parallel Topologies 525 526 5 Topologies with Uncoupled Sch\u00f6nflies Motions 7. 3PaPPR-1RUPU (Fig. 5.11a) Pa||P\\P\\kR (Fig. 5.2a) The first revolute joint of G4 limb and the last revolute joints of limbs G1, G2 and G3 have parallel axes. The last joints of limbs G1, G2 and G3 have superposed axes Pa||P\\P\\\\R (Fig. 5.2d) R\\R\\R\\P\\kR\\R (Fig. 5.1b) 8. 3PaPPR-1RUPU (Fig. 5.11b) Pa\\P\\kP\\kR (Fig. 5.2b) Idem No. 7 Pa\\P\\kP\\\\R (Fig. 5.2c) R\\R\\R\\P\\kR\\R (Fig. 5.1b) 9. 2PaPRRR1PaPRR1RUPU (Fig. 5.12a) Pa\\P\\\\R||R\\\\R (Fig. 5.3a) The first revolute joint of G4 limb and the last revolute joints of limbs G1, G2 and G3 have parallel axes. The last joints of of limbs G1, G2 and G3 have superposed axes. The revolute joints between links 5 and 6 of of limbs G1, G2 and G3 have orthogonal axes Pa\\P\\\\R||R (Fig. 5.2e) R\\R\\R\\P\\kR\\R (Fig. 5.1b) 10. 2PaPRRR1PaPRR1RUPU (Fig. 5.12b) Pa||P\\R||R\\\\R (Fig. 5.3b) Idem No. 9 Pa||P\\R||R (Fig. 5.2f) R\\R\\R\\P\\kR\\R (Fig. 5.1b) 11. 2PaRPRR1PaRPR1RUPU (Fig. 5.13) Pa\\R\\P\\kR\\R (Fig. 5.3c) The first revolute joint of G4 limb and the last revolute joints of limbs G1, G2 and G3 have parallel axes. The last joints of of limbs G1, G2 and G3 have superposed axes, and their revolute joints between links 4 and 5 have orthogonal axes Pa\\R\\P\\kR (Fig. 5.2g) R\\R\\R\\P\\kR\\R (Fig. 5.1b) 12. 2PaRRRR1PaRRR1RUPU (Fig. 5.14) Pa\\R||R||R\\kR (Fig. 5.3d) Idem No" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_35_0000277_14644207211003321-Figure2-1.png", + "original_path": "designv11-35/openalex_figure/designv11_35_0000277_14644207211003321-Figure2-1.png", + "caption": "Figure 2. Numerical model of asymmetric gear pair: (a) kinematic model with boundary conditions, (b) overall view of the discretized model, (c) discretization around the root region, and (d) discretization near the flanks.", + "texts": [ + " Simulations involving contact and large deformation are time-intensive, which encourage the use of two-dimensional models (2D) to reduce the computational effort. In quasi-static numerical analyses, the gear bending stress can be predicted using a 2D three-teeth model with a minimum error of 1% compared to three-dimensional (3D) models of the complete gear.31 Hence, a 2D model with three teeth is preferable compared to a comprehensive 3D gear model. Accordingly, in this study, 2D, multipair contact models of symmetric and asymmetric gears with three teeth and full-rim were used. Figure 2(a) to (d) shows the kinematic and discretized models of asymmetric gear. In the model, the driver and driven gears are positioned at the top and the bottom, respectively. The parts of symmetric and asymmetric gears were created and discretized in ABAQUS/Standard. An ideal involute curve was considered for the profile, and the profile deviations which occur during fabrication were disregarded. The involute profile for the teeth was generated using the co-ordinates determined using a MATLAB code. The gear materials were assumed to exhibit linear elastic behavior and isotropic properties" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_35_0000141_tasc.2021.3062485-Figure6-1.png", + "original_path": "designv11-35/openalex_figure/designv11_35_0000141_tasc.2021.3062485-Figure6-1.png", + "caption": "Fig. 6. Magnetic field distribution for machine with iron yoke.", + "texts": [ + " Losses for the stator winding are summarized in Table VII. The two high-loss components are due to hysteresis and transport current. The magnetic field experienced by the stator windings is intentionally kept low for limiting the hysteresis loss component and for ease of cooling the winding. The table has an additional column for an iron tooth machine with a 30-mm thick (radially) iron yoke. The addition of iron yoke reduces the amp-turns requirement on the rotor by about 10% but adds significant weight to the machine. Fig. 6 shows field distribution for such a machine\u2013there is essentially no field inside the slots. The iron yoke significantly reduces field in slots as compared with the machine without the iron yoke and the hysteresis loss is reduced very significantly. However, any savings in hysteresis loss in the superconductor is replaced by core losses in the iron Authorized licensed use limited to: University of Cape Coast. Downloaded on May 14,2021 at 10:25:07 UTC from IEEE Xplore. Restrictions apply. yoke. Table VII shows that total stator losses increase with the addition of the iron yoke" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_35_0001419_978-94-007-7401-8-Figure3.38-1.png", + "original_path": "designv11-35/openalex_figure/designv11_35_0001419_978-94-007-7401-8-Figure3.38-1.png", + "caption": "Fig. 3.38 Overactuated PMs with coupled Sch\u00f6nflies motions of types 4RRPRP (a) and 4RPRRP (b) defined by MF \u00bc SF \u00bc 4; RF\u00f0 \u00de \u00bc v1; v2; v3;xd\u00f0 \u00de, TF = 0, NF = 2, limb topology R\\R||P||R\\P (a) and R\\P||R||R\\P (b)", + "texts": [ + "3p) The second joints of the four limbs have parallel axes (continued) 244 3 Overactuated Topologies with Coupled Sch\u00f6nflies Motions Table 3.2 (continued) No. PM type Limb topology Connecting conditions 30. 4RPRRR (Fig. 3.35a) R\\P\\kR\\R||R (Fig. 3.3q) Idem No. 16 31. 4RRRPR (Fig. 3.35b) R||R\\R||P||R (Fig. 3.3r) Idem No. 16 32. 4RRRPR (Fig. 3.36a) R||R\\R||P||R (Fig. 3.3s) Idem No. 16 33. 4RRRRP (Fig. 3.36b) R||R\\R||R||P (Fig. 3.3t) Idem No. 16 34. 4RRRRP (Fig. 3.37a) R||R\\R||R\\P (Fig. 3.3u) Idem No. 16 35. 4PRRRP (Fig. 3.37b) P\\R\\R||R\\P (Fig. 3.3v) Idem No. 29 36. 4RRPRP (Fig. 3.38a) R\\R||P||R\\P (Fig. 3.3w) Idem No. 16 37. 4RPRRP (Fig. 3.38b) R\\P||R||R\\P (Fig. 3.3x) Idem No. 16 38. 4RRPRP (Fig. 3.39a) R\\R||P||R\\P (Fig. 3.3y) Idem No. 16 39. 4RRRPP (Fig. 3.39b) R\\R||R||P\\P (Fig. 3.3z) Idem No. 16 40. 4PRRRP (Fig. 3.40a) P||R||R\\R\\P (Fig. 3.3z1) The fourth joints of the four limbs have parallel axes 41. 4RPRRP (Fig. 3.40b) R||P||R||R\\P (Fig. 3.3a) Idem No. 40 42. 4RPRRP (Fig. 3.41a) R||P||R\\R\\P (Fig. 3.3b0) Idem No. 40 43. 4RRPRP (Fig. 3.41b) R||R||P\\R\\P (Fig. 3.3c0) Idem No. 40 44. 4RRRPP (Fig. 3.42a) R||R\\R\\P\\\\P (Fig. 3.3d0) The third joints of the four limbs have parallel axes 45" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_35_0001419_978-94-007-7401-8-Figure6.7-1.png", + "original_path": "designv11-35/openalex_figure/designv11_35_0001419_978-94-007-7401-8-Figure6.7-1.png", + "caption": "Fig. 6.7 Maximaly regular fully-parallel PMs with Sch\u00f6nflies motions of types 3PPPR1RPaPaP (a) and 3PPPR-1RPaPatP (b) defined by MF = SF = 4, (RF) = (v1; v2; v3;xb), TF = 0, NF = 12, limb topology P\\P ??P\\||R, P\\P ??P ??R and R||Pa||Pa||P (a), R||Pa||Pat||P (b)", + "texts": [ + "2 Fully-Parallel Topologies with Simple and Complex Limbs 607 608 6 Maximally Regular Topologies with Sch\u00f6nflies Motions 6.2 Fully-Parallel Topologies with Simple and Complex Limbs 609 610 6 Maximally Regular Topologies with Sch\u00f6nflies Motions 6.2 Fully-Parallel Topologies with Simple and Complex Limbs 611 Table 6.5 Limb topology and connecting conditions of the fully-parallel solutions with no idle mobilities presented in Figs. 6.7, 6.8, 6.9, 6.10, 6.11, 6.12, 6.13, 6.14, 6.15, 6.16 No. PM type Limb topology Connecting conditions 1. 3PPPR-1RPaPaP (Fig. 6.7a) P\\P ??P\\||R (Fig. 4.1a) P\\P ??P ??R (Fig. 4.1b) R||Pa||Pa||P (Fig. 5.4k) The last joints of the four limbs have superposed axes/ directions. The actuated prismatic joints of limbs G1, G2 and G3 have orthogonal directions 2. 3PPPR-1RPaPatP (Fig. 6.7b) P\\P ??P\\||R (Fig. 4.1a) P\\P ??P ??R (Fig. 4.1b) R||Pa||Pat||P (Fig. 5.4l) Idem no. 1 3. 2PPRRR-1PPRR-1RPaPaP (Fig. 6.8a) P ?P\\||R||R\\R (Fig. 4.2a) P\\P\\||R||R (Fig. 4.1d) R||Pa||Pa||P (Fig. 5.4k) Idem no. 1 4. 2PPRRR-1PPRR1RPaPatP (Fig. 6.8b) P\\P\\||R||R\\R (Fig. 4.2a) P\\P\\||R||R (Fig. 4.1d) R||Pa||Pat||P (Fig. 5.4l) Idem no. 1 5. 2PRPRR-1PRPR-1RPaPaP (Fig. 6.9a) P||R\\P\\||R\\R (Fig. 4.1b) P||R\\P\\||R (Fig. 4.1e) R||Pa||Pa||P (Fig. 5.4k) Idem no. 1 6. 2PRPRR-1PRPR1RPaPatP (Fig. 6.9b) P||R\\P\\||R\\R (Fig" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_35_0001419_978-94-007-7401-8-Figure5.76-1.png", + "original_path": "designv11-35/openalex_figure/designv11_35_0001419_978-94-007-7401-8-Figure5.76-1.png", + "caption": "Fig. 5.76 2PaRPRR-1PaRPR-1CPaPat-type fully-parallel PMs with uncoupled Sch\u00f6nflies motions defined by MF = SF = 4, RF\u00f0 \u00de \u00bc \u00f0v1; v2; v3;xb\u00de, TF = 0, NF = 19, limb topology Pa\\R\\P\\kR\\R, Pa\\R\\P\\kR and C||Pa||Pat", + "texts": [ + "40) 8. 2PaPRRR1PaPRR1CPaPa (Fig. 5.73b) Pa||P\\R||R\\\\R (Fig. 5.3b) Idem No. 3 Pa||P\\R||R (Fig. 5.2f) C||Pa||Pa (Fig. 5.40) 9. 2PaPRRR1PaPRR1CPaPat (Fig. 5.74a) Pa\\P\\\\R||R\\\\R (Fig. 5.3a) Idem No. 3 Pa\\P\\\\R||R (Fig. 5.2e) C||Pa||Pat(Fig. 5.4p) 10. 2PaPRRR1PaPRR1CPaPat (Fig. 5.74b) Pa||P\\R||R\\\\R (Fig. 5.3b) Idem No. 3 Pa||P\\R||R (Fig. 5.2f) C||Pa||Pat (Fig. 5.4p) 11. 2PaRPRR1PaRPR1CPaPa (Fig. 5.75) Pa\\R\\P\\kR\\R (Fig. 5.3c) Idem No. 3 Pa\\R\\P\\kR (Fig. 5.2g) C||Pa||Pa (Fig. 5.40) 12. 2PaRPRR1PaRPR1CPaPat (Fig. 5.76) Pa\\R\\P\\kR\\R (Fig. 5.3c) Idem No. 3 Pa\\R\\P\\kR (Fig. 5.2g) C||Pa||Pat (Fig. 5.4p) 13. 2PaRRRR1PaRRR1CPaPa (Fig. 5.77) Pa\\R||R||R\\kR (Fig. 5.3d) Idem No. 3 Pa\\R||R||R (Fig. 5.2h) C||Pa||Pa (Fig. 5.40) 14. 2PaRRRR1PaRRR1CPaPat (Fig. 5.78) Pa\\R||R||R\\kR (Fig. 5.3d) Idem No. 3 Pa\\R||R||R (Fig. 5.2h) C||Pa||Pat (Fig. 5.4p) 15. 3PaPPaR1CPaPa (Fig. 5.79a) Pa||P\\Pa\\kR (Fig. 5.4a) The revolute joint between links 6D and 9 of G4 limb and the last revolute joints of limbs G1, G2 and G3 have superposed axes Pa||P\\Pa||R (Fig" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_35_0001419_978-94-007-7401-8-Figure3.11-1.png", + "original_path": "designv11-35/openalex_figure/designv11_35_0001419_978-94-007-7401-8-Figure3.11-1.png", + "caption": "Fig. 3.11 Overactuated PMs with coupled Sch\u00f6nflies motions of types 4PRRP (a) and 4PRPR (b) defined by MF \u00bc SF \u00bc 4; RF\u00f0 \u00de \u00bc v1; v2; v3;xd\u00f0 \u00de, TF = 0, NF = 6, limb topology P\\R||R||P (a) and P\\R||P||R (b)", + "texts": [ + " 4RRRP (Fig. 3.7a) R||R||R||P (Fig. 3.1f) Idem No. 2 8. 4RPRP (Fig. 3.7b) R\\P\\kR||P (Fig. 3.1g) Idem No. 2 9. 4PPRR (Fig. 3.8a) P\\P||R||R (Fig. 3.1h) The last revolute joints of the four limbs have parallel axes 10. 4PRPR (Fig. 3.8b) P\\R||P||R (Fig. 3.1p) Idem No. 9 11. 4RRPP (Fig. 3.9a) R||R||P\\P (Fig. 3.1j) Idem No. 2 12. 4RPRP (Fig. 3.9b) R||P||R\\P (Fig. 3.1k) Idem No. 2 13. 4RPPR (Fig. 3.10a) R||P\\P\\kR (Fig. 3.1l) Idem No. 2 14. 4RPPR (Fig. 3.10b) R\\P\\kP||R (Fig. 3.1m) Idem No. 2 15. 4PRRP (Fig. 3.11a) P\\R||R||P (Fig. 3.1n) The last prismatic joints of the four limbs have parallel directions 16. 4PRPR (Fig. 3.11b) P\\R||P||R (Fig. 3.1o) Idem No. 9 17. 4PRRP (Fig. 3.12a) P||R||R\\P (Fig. 3.1v) The first prismatic joints of the four limbs have parallel directions 18. 4RPRP (Fig. 3.12b) R||P||R\\P (Fig. 3.1q) Idem No. 2 19. 4PRPR (Fig. 3.13a) P||R\\P\\kR (Fig. 3.1r) Idem No. 1 20. 4RPPR (Fig. 3.13b) R||P\\P\\kR (Fig. 3.1s) Idem No. 2 21. 4RPPR (Fig. 3.14a) R\\P\\kP||R (Fig. 3.1t) Idem No. 2 22. 4RPRP (Fig. 3.14b) R\\P\\kR||P (Fig. 3.1u) Idem No. 2 23. 4PPPR (Fig. 3.15a) P\\P\\\\P\\kR (Fig. 3.1h0) Idem No. 9 24. 4PPPR (Fig" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_35_0001419_978-94-007-7401-8-Figure5.45-1.png", + "original_path": "designv11-35/openalex_figure/designv11_35_0001419_978-94-007-7401-8-Figure5.45-1.png", + "caption": "Fig. 5.45 3PaPaPR-1RPaPatP-type fully-parallel PM with uncoupled Sch\u00f6nflies motions defined by MF = SF = 4, RF\u00f0 \u00de \u00bc \u00f0v1; v2; v3;xb\u00de, TF = 0, NF = 30 limb topology Pa\\Pa\\kP\\\\R, Pa\\Pa\\kP\\kR and R||Pa||Pat||P", + "texts": [ + "4c) R||Pa||Pat||P (Fig. 5.4l) 7. 3PaPPaR-1RPPaPat (Fig. 5.42a) Pa||P\\Pa\\kR (Fig. 5.4a) Idem No. 3 Pa||P\\Pa||R (Fig. 5.4d) R||P||Pa||Pat (Fig. 5.4n) 8. 3PaPPaR-1RPPaPat (Fig. 5.42b) Pa\\P\\\\Pa\\kR (Fig. 5.4b) Idem No. 3 Pa\\P\\\\Pa||R (Fig. 5.4c) R||P||Pa||Pat (Fig. 5.4n) 9. 3PaPaPR-1RPaPaP (Fig. 5.43) Pa\\Pa\\kP\\\\R (Fig. 5.4f) Idem No. 1 Pa\\Pa\\kP\\kR (Fig. 5.4h) R||Pa||Pa||P (Fig. 5.4k) 10. 3PaPaPR-1RPaPaP (Fig. 5.44) Pa\\Pa\\\\P\\\\R (Fig. 5.4g) Idem No. 1 Pa\\Pa\\\\P\\kR (Fig. 5.4e) R||Pa||Pa||P (Fig. 5.4k) 11. 3PaPaPR1RPaPatP (Fig. 5.45) Pa\\Pa\\kP\\\\R (Fig. 5.4f) Idem No. 1 Pa\\Pa\\kP\\kR (Fig. 5.4h) R||Pa||Pat||P (Fig. 5.4l) 12. 3PaPaPR1RPaPatP (Fig. 5.46) Pa\\Pa\\\\P\\\\R (Fig. 5.4g) Idem No. 1 Pa\\Pa\\\\P\\kR (Fig. 5.4e) R||Pa||Pat||P (Fig. 5.4l) 13. 3PaPaPR-1RPPaPa (Fig. 5.47) Pa\\Pa\\kP\\\\R (Fig. 5.4f) Idem No. 3 Pa\\Pa\\kP\\kR (Fig. 5.4h) R||P||Pa||Pa (Fig. 5.4m) 14. 3PaPaPR-1RPPaPa (Fig. 5.48) Pa\\Pa\\\\P\\\\R (Fig. 5.4g) Idem No. 3 Pa\\Pa\\\\P\\kR (Fig. 5.4e) R||P||Pa||Pa (Fig. 5.4m) 15. 3PaPaPR-1RPPaPat (Fig. 5.49) Pa\\PakP\\\\R (Fig. 5.4f) Idem No. 3 Pa\\Pa\\kP\\kR (Fig" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_35_0001594_s11051-015-3102-6-Figure1-1.png", + "original_path": "designv11-35/openalex_figure/designv11_35_0001594_s11051-015-3102-6-Figure1-1.png", + "caption": "Fig. 1 Discretization of the nanowire to characterize DEP force, DEP torque, and the torque generated by DEP forces", + "texts": [ + " This model assumes the length of the nanowire is smaller than the non-uniformity of the electric field, which is practically inappropriate since the gap of electrodes usually is smaller than the nanowire\u2019s length. Therefore, similar to our analysis on 2D model (Tao et al. 2014), we divide the whole nanowire into N small segments such that the electric field around each part can be treated as uniform. Each segment is modeled as an ellipsoid, such that classical formulas can be used to calculate DEP force and torque (Jones 1995; Morgan and Green 2003), as shown in Fig. 1. Both the DEP force and the DEP torque on the ith segment have three base components, respectively; FDEPxh ii \u00bc 1 2 pr2lemRe ~Ka r Eixj j2; \u00f01a\u00de FDEPy i \u00bc 1 2 pr2lemRe ~Kb r Eiy 2; \u00f01b\u00de FDEPzh ii \u00bc 1 2 pr2lemRe ~Kc r Eizj j2; \u00f01c\u00de TDEPxh ii \u00bc 1 3 pr2lem Lc Lb\u00f0 \u00deEiyEizRe ~Kb ~Kc ; \u00f02a\u00de TDEPy i \u00bc 1 3 pr2lem La Lc\u00f0 \u00deEizEixRe ~Kc ~Ka ; \u00f02b\u00de TDEPzh ii \u00bc 1 3 pr2lem Lb La\u00f0 \u00deEixEiyRe ~Ka ~Kb ; \u00f02c\u00de in which em is the permittivity of the surrounding medium. r and l are the radius and length of the small segment, respectively" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_35_0000467_j.robot.2021.103815-Figure1-1.png", + "original_path": "designv11-35/openalex_figure/designv11_35_0000467_j.robot.2021.103815-Figure1-1.png", + "caption": "Fig. 1. Proposed variable parallel elastic actuator. (a) 3D CAD design, (b) General schematic of VPEA for implementing rotary nonlinear compliances.", + "texts": [ + " Then, in the remainder of his section, the mathematics regarding the computation of the ctuator\u2019s output torque along with the derivation of the VPEA daptation rule are discussed. Thereafter, Section 3 elaborates the teps to design the VPEA for a given system. In Section 4, the erformance of the VPEA with the proposed adaptation rule is nvestigated in terms of energy efficiency on a robotic leg model s the case study. Finally, we conclude with a discussion and rovide conclusions in Section 5. . Variable parallel elastic actuator Displayed in Fig. 1 is the variable parallel elastic actuator esign for employment in systems comprised of rotary motion p mechanisms. The setup consists of three main parts: (1) a noncircular cam coupled to the motor shaft, (2) a linear compression spring along with a follower and (3) a secondary nonbackdrivable servomotor for changing the spring preset via a ballscrew mechanism which creates a linear drive for stiffness adjustment [21]. In this actuator, the cam with its non-circular profile in combination with the spring produces a torque profile accordant with the cam shape", + " In order to design a cam shape for the unidirectional rotary movements of a rotary joint (\u03c6)1 based on a given torque profile (uc) some design parameters such as the spring stiffness (k) and the initial spring length (l\u2212 l0) should e predefined. The next step is to solve the following differential quation with j(0) being its initial condition: dj(\u03c6) d\u03c6 = uc(\u03c6) k(l(\u03c6) \u2212 l0 \u2212 l\u03020(0)) = uc(\u03c6) k(l(\u03c6) \u2212 l0) (1) where j(\u03c6) is the distance between the center of the cam and oller and is defined as j = L\u2212D\u2212 (l\u2212 l\u03020); see Fig. 1(b). According to this figure, L and D are two lengths representing the distance between the spring and center of the cam, and the distance between the fixed end of the spring and center of the roller, respectively. Also, l\u03020 is the spring precompression adjusted by the motor, which at the time of cam profile design, it is considered 1 In rotary movements: \u222b T \u03c6\u0307dt = 2\u03c0, \u03c6\u0307 > 0; where T denotes one time eriod of motion. t s e i 2 e c t d u \u02c6 w t r p f r t o t j ( t m a t c j m i f b t U N o w f c w c t d p F E m f u u T s u N t f b t 4 i A l p t u a p v t n o be zero; i.e., l\u03020(0) = 0. With this definition, l(\u03c6)\u2212 l\u03020 is then the pring length and l0 denotes the spring rest length. For detailed xplanations of the remaining steps to design the cam shape, the nterested reader is referred to [27] and [28]. .2. Compliance adjustment approach As mentioned previously, in our proposed variable parallel lastic actuator (Fig. 1), the secondary motor is responsible for hanging the precompression of the spring (l\u03020(t)) so as to adapt he designed compliance torque (uc) to the required torque. Hereinafter, the adapted compliance torque is denoted by u\u0303c and efined as: \u02dc c(\u03c6) = k dj(\u03c6) d\u03c6 ( l(\u03c6) \u2212 l0 \u2212 l\u03020(t) ) = uc(\u03c6) \u2212 k dj(\u03c6) d\u03c6 l\u03020(t) (2) As from now, for the sake of simplicity, unless otherwise specified, we denote by j\u2032(\u03c6) the derivative of j with respect to \u03c6. Now, the degree to which the compliance preset (l\u03020) must be adjusted in order to minimize the actuator\u2019s squared torque (i" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_35_0001647_h0055877-Figure1-1.png", + "original_path": "designv11-35/openalex_figure/designv11_35_0001647_h0055877-Figure1-1.png", + "caption": "FIG. 1. Arrangement of blocks and available culs-de-sac in goal units to illustrate setting for Group VI in Part I (see Table 1)", + "texts": [ + " By imposing mechanical guidance for two successive units near the goal, one should be able to determine in the free units whether there is more, equal, or less effect than for a comparable pair of culs-de-sac. To establish the relative dominance one can also pit mechanical guidance for one pair of goal units against trialand-error avoidance of culs-de-sac on the other side for an adjacent pair. Finally, after a pattern has been established in the free area of the maze, how resistant is it to change by a shift of culs-de-sac or guided paths in the goal region ? The elevated rectilinear maze of 12 units, the last four of which are shown in Fig. 1, is an adaptation of the Heron-Drake pattern (8) and is described in detail in (7). The doors which prevent retracing can be left unlocked to make the unit free, or one can be locked to make a culde-sac on that 6ide. For mechanical guidance 431 432 GEORGE M. HASLERUD a 9 X 12-in. metal sheet, painted black like the rest of the apparatus, was fastened 1 in. from the T joining to block a cul-de-sac. The four goal units were prescribed for various combinations of pairs of guided and problem units as shown in Table 1" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_35_0001419_978-94-007-7401-8-Figure2.112-1.png", + "original_path": "designv11-35/openalex_figure/designv11_35_0001419_978-94-007-7401-8-Figure2.112-1.png", + "caption": "Fig. 2.112 4CPaRR-type fully-parallel PM with coupled Sch\u00f6nflies motions defined by MF \u00bc SF \u00bc 4; RF\u00f0 \u00de \u00bc \u00f0v1; v2; v3;xd\u00de, TF = 0, NF = 14, limb topology C\\Pa\\kR\\R", + "texts": [ + " Gogu G (2009) Structural synthesis of parallel robots: part 2-translational topologies with two and three degrees of freedom. Springer, Dordrecht 3. Gogu G (2010) Structural synthesis of parallel robots: part 3-topologies with planar motion of the moving platform. Springer, Dordrecht 4. Gogu G (2012) Structural synthesis of parallel robots: part 4-other topologies with two and three degrees of freedom. Springer, Dordrecht Table 2.8 Structural parametersa of parallel mechanisms in Figs. 2.112, 2.113, 2.114, 2.115, 2.116, 2.117, 2.118, 2.119 No. Structural parameter Solution Figure 2.112 Figures 2.113, 2.114, 2.115, 2.116, 2.117, 2.118, 2.119 1. m 22 34 2. pi (i = 1,\u2026,4) 7 11 3. p 28 44 4. q 7 11 5. k1 0 0 6. k2 4 4 7. k 4 4 8. (RG1) (v1; v2; v3;xb;xd) (v1; v2; v3;xb;xd) 9. (RG2) (v1; v2; v3;xa;xd) (v1; v2; v3;xa;xd) 10. (RG3) (v1; v2; v3;xb;xd) (v1; v2; v3;xb;xd) 11. (RG4) (v1; v2; v3;xa;xd) (v1; v2; v3;xa;xd) 12. SGi (i = 1,\u2026,4) 5 5 13. rGi (i = 1,\u2026,4) 3 6 14. MGi (i = 1,\u2026,4) 5 5 15. (RF) (v1; v2; v3;xd) (v1; v2; v3;xd) 16. SF 4 4 17. rl 12 24 18. rF 28 40 19. MF 4 4 20. NF 14 26 21" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_35_0000118_icmeae51770.2020.00022-Figure1-1.png", + "original_path": "designv11-35/openalex_figure/designv11_35_0000118_icmeae51770.2020.00022-Figure1-1.png", + "caption": "Figure 1: Reference frames and principal forces convention for a X configuration Parrot Bebop 2 R\u00a9 quadrotor, [17].", + "texts": [ + " The remaining of the paper is divided as follows. In Section II the mathematical tools for the aircraft dynamics are presented. The overall controller scheme is developed in Section III, with simulation results in Section IV. In Section V some conclusions and findings are described. The mathematical model of the UAV considers two reference frames. A ground-fixed inertial framed defined as I={xI , yI , zI}, and a body frame attached to the center of gravity of the aircraft as B={xB, yB, zB}, see Fig. 1. In order to describe the equations of motion, a Newton-Euler formulation is used. The model is defined as \u03be\u0307 = V (1) mV\u0307 = R(\u2212TT ) +mge3 (2) R\u0307 = R\u03a9\u0302 (3) J\u03a9\u0307 = \u2212\u03a9 \u00d7 J\u03a9 + \u03c4a (4) where \u03be = (x, y, z) \u2208 R 3 are the position coordinates relative to the inertial frame and \u03b7 = (\u03c6, \u03b8, \u03c8) \u2208 R 3 describes the rotation coordinates for UAV; this given by an orthogonal rotation matrix R \u2208 SO(3) : B \u2192 I parameterized by the Euler angles \u03c6 roll, \u03b8 pitch and \u03c8 yaw or heading, defined as R = \u239b \u239d c\u03b8c\u03c8 s\u03c6s\u03b8c\u03c8 \u2212 c\u03c6s\u03c8 c\u03c6s\u03b8c\u03c8 + s\u03c6s\u03c8 c\u03b8s\u03c8 s\u03c6s\u03b8s\u03c8 + c\u03c6c\u03c8 c\u03c6s\u03b8s\u03c8 \u2212 s\u03c6c\u03c8 \u2212s\u03b8 s\u03c6c\u03b8 c\u03c6c\u03b8 \u239e \u23a0 (5) where c\u03b8, c\u03c6, c\u03c8, s\u03b8, s\u03c6, s\u03c8 are cosine and sine operations, respectively" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_35_0001419_978-94-007-7401-8-Figure6.39-1.png", + "original_path": "designv11-35/openalex_figure/designv11_35_0001419_978-94-007-7401-8-Figure6.39-1.png", + "caption": "Fig. 6.39 3PPaPaR-1RPaPatP-type maximaly regular fully-parallel PM with Sch\u00f6nflies motions defined by MF = SF = 4, (RF) = (v1; v2; v3;xb), TF = 0, NF = 30, limb topology P||Pa||Pa\\R, P||Pa||Pa||R and R||Pa||Pat||P", + "texts": [ + " 3PPaPR-1RPPatPa (Fig. 6.35) P||Pa\\P\\\\R (Fig. 4.8d) P||Pa\\P\\||R (Fig. 4.8e) R||P||Pa||Pat (Fig. 5.4n) Idem no. 9 15. 2PPaRRR-1PPaRR-1RPPaPa (Fig. 6.36) P||Pa||R||R\\R (Fig. 4.9a) P||Pa||R||R (Fig. 4.8g) R||P||Pa||Pa (Fig. 5.4m) Idem no. 9 16. 2PPaRRR-1PPaRR-1RPPaPat (Fig. 6.37) P||Pa||R||R\\R (Fig. 4.9a) P||Pa||R||R (Fig. 4.8g) R||P||Pa||Pat (Fig. 5.4n) Idem no. 9 17. 3PPaPaR-1RPaPaP (Fig. 6.38) P||Pa||Pa\\R (Fig. 4.10a) P||Pa||Pa||R (Fig. 4.10b) R||Pa||Pa||P (Fig. 5.4k) Idem no. 1 18. 3PPaPaR-1RPaPatP (Fig. 6.39) P||Pa||Pa\\R (Fig. 4.10a) P||Pa||Pa||R (Fig. 4.10b) R||Pa||Pat||P (Fig. 5.4l) Idem no. 1 19. 3PPaPaR-1RPPaPa (Fig. 6.40) P||Pa||Pa\\R (Fig. 4.10a) P||Pa||Pa||R (Fig. 4.10b) R||P||Pa||Pa (Fig. 5.4m) Idem no. 9 20. 3PPaPaR-1RPaPatP (Fig. 6.41) P||Pa||Pa\\R (Fig. 4.10a) P||Pa||Pa||R (Fig. 4.10b) R||P||Pa||Pat (Fig. 5.4n) Idem no. 9 Table 6.14 Limb topology and connecting conditions of the fully-parallel solutions with no idle mobilities presented in Figs. 6.42, 6.43, 6.44, 6.45, 6.46 No. PM type Limb topology Connecting conditions 1" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_35_0001411_978-3-030-55061-5-Figure20-1.png", + "original_path": "designv11-35/openalex_figure/designv11_35_0001411_978-3-030-55061-5-Figure20-1.png", + "caption": "Fig. 20. Motion simulation with cross link", + "texts": [ + " The simulation of the three degrees of freedom of the end-effector in the first cycle is presented in Figs. 11, 12 and 13. Fig. 9. Parallelogram mechanism modeled with ADAMS For the second cycle (Fig. 14) the mobility number is 2. A rotation around the Z axis will be possible (Fig. 15) and the two independent translations, X (Fig. 16) and Y (Fig. 17), will appear as impossible (for example in Fig. 18). the base and an actuator was placed along the X axis, between the cross element and the base (Fig. 19). Following the simulation, it is observed visually (Fig. 20), but also in the superimposed kinematic diagrams (Fig. 21), the existence of a degree of freedom and the appearance of two dependent movements, translation along the X and Y axis. Computer-Assisted Learning Used to Overconstrained Mechanism\u2019s Mobility 525 The authors use computer-assisted learning/teaching based on the ADAMS software and develop a new teaching method for calculating the mobility of overconstrained mechanismswith a structural formula. The algorithm is based on the closed chain segmentation method, developed on the TRIZ inventive method" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_35_0000491_aero50100.2021.9438380-Figure1-1.png", + "original_path": "designv11-35/openalex_figure/designv11_35_0000491_aero50100.2021.9438380-Figure1-1.png", + "caption": "Figure 1. OctArm Continuum Manipulator", + "texts": [ + " Once calculated, we can use equation 3 to find the desired primitive parameters \u03c4p(p\u0302) in b1. One final adjustment is the weighting of the k-NN primitives according to the distance between the current endeffector location and the starting location of each primitive. By adding this set of weights, wi. . .wk, we obtain a weighted least squares solution when solving 7. Thus, equations 3 and 7 are adjusted as follows: { A1Wx = b1 A2Wx = b2 , (8) whereW = diag(1,w1, w2, . . . , wk). The OctArm, illustrated in Figure 1, is a 3-section, 9 DoF continuum robot, actuated by pneumatically driven McKibben actuators. Pressuring the McKibben muscle creates extension along the length of an individual muscle. By connecting three of these muscles (or sets of them at equal intervals) in parallel to form a continuum section, we can extend said section by pressurizing all muscles simultaneously by the same amount or create bending by differentially pressurizing the muscles. The OctArm is comprised of three of these serially connected sections, with each section actuated by three independently controlled pressures, and capable of independent extension and two DoF of spatial bending, providing the OctArm 9 DoF overall" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_35_0001419_978-94-007-7401-8-Figure5.75-1.png", + "original_path": "designv11-35/openalex_figure/designv11_35_0001419_978-94-007-7401-8-Figure5.75-1.png", + "caption": "Fig. 5.75 2PaRPRR-1PaRPR-1CPaPa-type fully-parallel PMs with uncoupled Sch\u00f6nflies motions defined by MF = SF = 4, RF\u00f0 \u00de \u00bc \u00f0v1; v2; v3;xb\u00de, TF = 0, NF = 19, limb topology Pa\\R\\P\\kR\\R, Pa\\R\\P\\kR and C||Pa||Pa", + "texts": [ + " 2PaPRRR1PaPRR1CPaPa (Fig. 5.73a) Pa\\P\\\\R||R\\\\R (Fig. 5.3a) Idem No. 3 Pa\\P\\\\R||R (Fig. 5.2e) C||Pa||Pa (Fig. 5.40) 8. 2PaPRRR1PaPRR1CPaPa (Fig. 5.73b) Pa||P\\R||R\\\\R (Fig. 5.3b) Idem No. 3 Pa||P\\R||R (Fig. 5.2f) C||Pa||Pa (Fig. 5.40) 9. 2PaPRRR1PaPRR1CPaPat (Fig. 5.74a) Pa\\P\\\\R||R\\\\R (Fig. 5.3a) Idem No. 3 Pa\\P\\\\R||R (Fig. 5.2e) C||Pa||Pat(Fig. 5.4p) 10. 2PaPRRR1PaPRR1CPaPat (Fig. 5.74b) Pa||P\\R||R\\\\R (Fig. 5.3b) Idem No. 3 Pa||P\\R||R (Fig. 5.2f) C||Pa||Pat (Fig. 5.4p) 11. 2PaRPRR1PaRPR1CPaPa (Fig. 5.75) Pa\\R\\P\\kR\\R (Fig. 5.3c) Idem No. 3 Pa\\R\\P\\kR (Fig. 5.2g) C||Pa||Pa (Fig. 5.40) 12. 2PaRPRR1PaRPR1CPaPat (Fig. 5.76) Pa\\R\\P\\kR\\R (Fig. 5.3c) Idem No. 3 Pa\\R\\P\\kR (Fig. 5.2g) C||Pa||Pat (Fig. 5.4p) 13. 2PaRRRR1PaRRR1CPaPa (Fig. 5.77) Pa\\R||R||R\\kR (Fig. 5.3d) Idem No. 3 Pa\\R||R||R (Fig. 5.2h) C||Pa||Pa (Fig. 5.40) 14. 2PaRRRR1PaRRR1CPaPat (Fig. 5.78) Pa\\R||R||R\\kR (Fig. 5.3d) Idem No. 3 Pa\\R||R||R (Fig. 5.2h) C||Pa||Pat (Fig. 5.4p) 15. 3PaPPaR1CPaPa (Fig. 5.79a) Pa||P\\Pa\\kR (Fig. 5.4a) The revolute joint between links 6D and 9 of G4 limb and the last revolute joints of limbs G1, G2 and G3 have superposed axes Pa||P\\Pa||R (Fig" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_35_0000797_9781119526483.ch14-Figure14.37-1.png", + "original_path": "designv11-35/openalex_figure/designv11_35_0000797_9781119526483.ch14-Figure14.37-1.png", + "caption": "Figure 14.37 Diatom ion cyclotron resonance (1987). A take on diatoms and cyclotrons [14.207] (public domain image). Accelerating diatoms are Cymbella cistula [14.244] (reprinted with permission of the Mus\u00e9um national d\u2019Histoire naturelle under a Creative Commons license).", + "texts": [ + " The basic idea is that ions that cross cell membranes, such as calcium, potassium and magnesium, can have their rates of movement affected by changing magnetic fields that cause them to go into resonance at certain frequencies that are related to how cyclotrons work. The conflicting results for diatom motility have been tabulated [14.109] [14.200] [14.216], and they are hard to interpret because of \u201cthe failure to find a reasonable physical explanation\u201d [14.216] (cf. [14.2] [14.289]). This also makes for lack of a model for the relationship between ion cyclotron resonance and the diatom motor, which stands with no specific hypothesis (Figure 14.37). The initial work showed a positive effect [14.217] [14.236] [14.237] [14.238] [14.343] [14.344]. It was followed by negative results reported by a number of groups [14.53] [14.54] [14.278] [14.293] [14.312] and some positive results [14.302]. The two microfilaments adjacent to a raphe may not be static structures. For example, if actin were added at one end and came off the other end, the microfilament would exhibit treadmilling [14.43] (Figure 14.38). Actin treadmilling has been hypothesized for the growth of diatom setae [14" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_35_0001419_978-94-007-7401-8-Figure3.62-1.png", + "original_path": "designv11-35/openalex_figure/designv11_35_0001419_978-94-007-7401-8-Figure3.62-1.png", + "caption": "Fig. 3.62 Overactuated PMs with coupled Sch\u00f6nflies motions of types 4RPPaP (a) and 4PPPaR (b) defined by MF \u00bc SF \u00bc 4; RF\u00f0 \u00de \u00bc v1; v2; v3;xd\u00f0 \u00de, TF = 0, NF = 18, limb topology R\\P\\kPa||P (a) and P\\P\\kPa||R (b)", + "texts": [ + "50g) Idem No. 1 8. 4PRPaP (Fig. 3.57b) P\\R||Pa||P (Fig. 3.50h) Idem No. 1 9. 4RPRPa (Fig. 3.58a) R||P||R||Pa (Fig. 3.50i) Idem No. 1 10. 4RRPPa (Fig. 3.58b) R||R||P||Pa (Fig. 3.50j) Idem No. 1 11. RRPaP (Fig. 3.59) R||R||Pa||P (Fig. 3.50k) Idem No. 1 12. 4RPaPP (Fig. 3.60a) R||Pa\\P\\kP (Fig. 3.50l) Idem No. 1 13. 4RPPPa (Fig. 3.60b) R||P\\P\\kPa (Fig. 3.50m) Idem No. 1 14. 4PaRPP (Fig. 3.61a) Pa||R||P\\P (Fig. 3.50n) Idem No. 1 15. 4RPaPP (Fig. 3.61b) R||Pa||P\\P (Fig. 3.50o) Idem No. 1 16. 4RPPaP (Fig. 3.62a) R\\P\\kPa||P (Fig. 3.50p) Idem No. 1 17. 4PPPaR (Fig. 3.62b) P\\P\\kPa||R (Fig. 3.50q) Idem No. 1 18. 4PPPaR (Fig. 3.63a) P\\P||Pa||R (Fig. 3.50r) Idem No. 1 19. 4PPaRR (Fig. 3.63b) P||Pa||R||R (Fig. 3.50s) Idem No. 1 20. 4PPaRP (Fig. 3.64) P\\Pa||R||P (Fig. 3.50t) Idem No. 1 21. 4PPaPR (Fig. 3.65) P\\Pa||P||R (Fig. 3.50u) Idem No. 1 22. 4PPaPR (Fig. 3.66a) P||Pa\\P\\kR (Fig. 3.50v) Idem No. 1 23. 4PPaRP (Fig. 3.66b) P||Pa\\R\\P (Fig. 3.50i0) The before last joints of the four limbs have parallel axes 24. 4RPPaP (Fig. 3.67) R||P||Pa\\P (Fig. 3.50w) Idem No. 1 25. 4RPRPa (Fig" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_35_0000296_s12206-021-0435-1-Figure1-1.png", + "original_path": "designv11-35/openalex_figure/designv11_35_0000296_s12206-021-0435-1-Figure1-1.png", + "caption": "Fig. 1. Coordinate frames of the St\u00e4ubli RX160.", + "texts": [ + " The joint torques of the robot, computed through of the identified dynamic model, are subtracted from the measured joint torques in order to obtain the additional torques due to the external interaction. The wrench at the end-effector is calculated by means of the inverse transpose of the Jacobian matrix. This approach has been used in different robotic systems [32-34]. The method proposed in this paper can be regarded as an extension of Refs. [32, 33], with the consideration of a complex friction model and joint torques due to the external interaction. The main goal of this work is to identify a consistent dynamic model of the St\u00e4ubli RX-160 industrial robot (Fig. 1) as a case study. The base parameters, defined in a linear parameter vector of minimal order [35, 36], are used in the modeling equations of the robot. Base parameters are linear combinations of the dynamic parameters of each link composing the robot and their use results in a well-conditioned overdetermined regression matrix. A friction model including the complex Stribeck effect is used and the mechanical coupling between the 5th and 6th joints of the robot is taken into account in order to establish a realistic dynamic model", + " For the remaining two joints, the servos are mounted inside the fourth link of the robot and a gear transmission transfers the motion from servo 6 through the 5th joint to the 6th joint [37]. This mechanical design causes a kinematic coupling between the servo 5 and joint 6 as shown in Fig. 3. The angular velocity of servo 6 is related to the difference of joint velocities 5q and 6q . Consequently, one needs to define additional coordinate axes in order to take into account this extra friction model. The velocity 7q associated with the additional coordinate axes is defined as follows: 7 5 6 q q q= + . (1) The frame definition of the robot is shown in Fig. 1, and the corresponding DH parameters in Table 1. For each link in the Table 1, the transformation matrix 1 i iT \u2212 of the frame attached to the link i with respect to the frame attached to link 1i \u2212 can be given as follows: 1 0 0 0 0 1 i i i i i i i i i i i i i ii i i i i i c s c s s a c s c c c s a s T s c d \u03b8 \u03b8 \u03b1 \u03b8 \u03b1 \u03b8 \u03b8 \u03b8 \u03b1 \u03b8 \u03b1 \u03b8 \u03b1 \u03b1\u2212 \u2212\u23a1 \u23a4 \u23a2 \u23a5\u2212\u23a2 \u23a5= \u23a2 \u23a5 \u23a2 \u23a5 \u23a2 \u23a5\u23a3 \u23a6 (2) where ic\u03b8 and is\u03b8 are the standard abbreviations of cos( )i\u03b8 and ( )sin .i\u03b8 The basic equations of motion are derived by using the Euler-Lagrange equations [38]" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_35_0001419_978-94-007-7401-8-Figure6.8-1.png", + "original_path": "designv11-35/openalex_figure/designv11_35_0001419_978-94-007-7401-8-Figure6.8-1.png", + "caption": "Fig. 6.8 Maximaly regular fully-parallel PMs with Sch\u00f6nflies motions of types 2PPRRR1PPRR-1RPaPaP (a) and 2PPRRR-1PPRR-1RPaPatP (b) defined by MF = SF = 4, (RF) = (v1; v2; v3;xb), TF = 0, NF = 10, limb topology P\\P\\||R||R\\R, P\\P\\||R||R and R||Pa||Pa||P (a), R||Pa||Pat||P (b)", + "texts": [ + "5 Limb topology and connecting conditions of the fully-parallel solutions with no idle mobilities presented in Figs. 6.7, 6.8, 6.9, 6.10, 6.11, 6.12, 6.13, 6.14, 6.15, 6.16 No. PM type Limb topology Connecting conditions 1. 3PPPR-1RPaPaP (Fig. 6.7a) P\\P ??P\\||R (Fig. 4.1a) P\\P ??P ??R (Fig. 4.1b) R||Pa||Pa||P (Fig. 5.4k) The last joints of the four limbs have superposed axes/ directions. The actuated prismatic joints of limbs G1, G2 and G3 have orthogonal directions 2. 3PPPR-1RPaPatP (Fig. 6.7b) P\\P ??P\\||R (Fig. 4.1a) P\\P ??P ??R (Fig. 4.1b) R||Pa||Pat||P (Fig. 5.4l) Idem no. 1 3. 2PPRRR-1PPRR-1RPaPaP (Fig. 6.8a) P ?P\\||R||R\\R (Fig. 4.2a) P\\P\\||R||R (Fig. 4.1d) R||Pa||Pa||P (Fig. 5.4k) Idem no. 1 4. 2PPRRR-1PPRR1RPaPatP (Fig. 6.8b) P\\P\\||R||R\\R (Fig. 4.2a) P\\P\\||R||R (Fig. 4.1d) R||Pa||Pat||P (Fig. 5.4l) Idem no. 1 5. 2PRPRR-1PRPR-1RPaPaP (Fig. 6.9a) P||R\\P\\||R\\R (Fig. 4.1b) P||R\\P\\||R (Fig. 4.1e) R||Pa||Pa||P (Fig. 5.4k) Idem no. 1 6. 2PRPRR-1PRPR1RPaPatP (Fig. 6.9b) P||R\\P\\||R\\R (Fig. 4.1b) P||R\\P\\||R (Fig. 4.1e) R||Pa||Pat||P (Fig. 5.4l) Idem no. 1 7. 2PRRRR-1PRRR-1RPaPaP (Fig. 6.10a) P||R||R||R\\R (Fig. 4.2d) P||R||R||R (Fig. 4.1g) R||Pa||Pa||P (Fig. 5.4k) Idem no. 1 8. 2PRRRR-1PRRR1RPaPatP (Fig. 6.10b) P||R||R||R\\R (Fig" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_35_0000248_09596518211007294-Figure1-1.png", + "original_path": "designv11-35/openalex_figure/designv11_35_0000248_09596518211007294-Figure1-1.png", + "caption": "Figure 1. Shaking table morphology: (a) one horizontal DoF, (b) two horizontal DoF, (c) three linear DoF, and (d) six DoF.", + "texts": [ + "haking table control, feedback linearization, system identification, dynamic inversion, state-space modeling Date received: 28 January 2021; accepted: 3 March 2021 Regardless of the advances in numerical methods, computation power, and constitutive laws for engineering materials, experimental dynamic testing of structures is still indispensable for behavior investigation, model validation, and product certification purposes. Shaking table testing is one of the most representative laboratory structural testing methods together with quasistatic testing and hybrid simulation (HS).1,2 A shaking table is a very stiff platform onto which the structure under test (SuT) is mounted and vibrated. Depending on the spatial arrangement of the actuators which drive the table, it will be able to move in one or several degrees of freedom (DoF). Figure 1 shows different morphologies of shaking tables identifying their main components. These testing systems are often powered by hydraulic servoactuators due to their high force density and velocity and stroke capabilities.3 High-performance servovalves govern the motion of the actuators, which are often equipped with advanced features such as low friction (hydrostatic or hydrodynamic) rod bearings and adjustable backlash swivels. Motion to be reproduced by the platform is usually specified in terms of acceleration time histories (waveform replication); however, other reference specifications such as swept sine, resonance search and tracked dwell, random (acceleration spectral density), sine on random, random on random, shock, or shock response spectrum are utilized depending on the purpose of the test to be carried out.4 The tight tolerances demanded in reference tracking, and the high frequency content of motion profiles along with the low hydraulic resonance frequencies and the inherent non-linearity of hydraulic actuation systems make the design of control systems for shaking tables a challenging task. This circumstance is more exacerbated when dealing with multi-axis, over-constrained (more actuators than DoFs) systems (see Figure 1(d)). Techniques to control the motion of shaking table systems fall into two main categories, namely, iterative or frequency domain Escuela Te\u0301cnica Superior de Ingenieros de Caminos, Canales y Puertos, Universidad Polite\u0301cnica de Madrid, Madrid, Spain Corresponding author: Jose\u0301 Ram\u0131\u0301rez-Senent, Escuela Te\u0301cnica Superior de Ingenieros de Caminos, Canales y Puertos, Universidad Polite\u0301cnica de Madrid, Campus Ciudad Universitaria, Calle del Prof. Aranguren, 3, 28040 Madrid, Spain. Email: jose.ramirez", + " Other methods make use of SuT reaction force to cope more adequately with table\u2013structure interaction;21 however, according to the authors\u2019 knowledge, a real-time dynamic inversion approach, accounting simultaneously for the effects of the force exerted by the SuT and the shake table on the actuator and of the servovalve command on table acceleration, has not yet been addressed. In this article, a new MBC methodology based on the real-time inversion of the state-space model of the shaking table servoactuator is presented. A parallel PID controller complements the model inversion by accounting for non-modeled dynamics and external perturbations. The proposed technique is applied to a one DoF planar shaking table (see Figure 1(a)), and its effectiveness is demonstrated by means of numerical simulations. The output of the actuator state-space model is rod acceleration, and its inputs are, on one hand, the servovalve control command and the external force exerted on actuator rod, on the other. The latter provides all the relevant dynamic information related to the shaking table and the SuT with which it is loaded and, eventually, on the external forces exerted on them. The suggested method belongs to the time domain\u2014 MBC family, does not require iterations, and is able to cope with SuT non-linear behavior, external forces acting on the system, and acceleration references, which change during the test, as in the case of complex HS test scenarios, in which both shaking tables and actuators are used", + " The \u2018\u2018Control methodology description\u2019\u2019 section addresses the control system which comprises the servovalve feedback linearization scheme described in the \u2018\u2018Servovalve flow feedback linearization\u2019\u2019 section, the system identification module covered in the \u2018\u2018State-space model identification\u2019\u2019 section, and the real-time dynamic inversion module and parallel PID controller, which are dealt with in the \u2018\u2018Real-time dynamic inversion and parallel PID controller\u2019\u2019 section. Simulation results obtained are shown in the \u2018\u2018Simulation results and discussion\u2019\u2019 section and are compared with the performance of a generic iterative method. Finally, conclusions, potential applications, and further work suggestions for a successful actual implementation are outlined in the \u2018\u2018Conclusions and further work\u2019\u2019 section. A model of the one DoF shaking table system (see Figure 1(a)) and the SuT, consisting in a two-story shear building, has been implemented to assess the performance of the proposed control methodology. The shaking table under study features a bare table mass of 3000 kg, a maximum payload of 10,000 kg. It is powered by a servoactuator with a stroke of 300mm and a static load capacity of 165 kN at a supply pressure of 280bar. Figure 4 illustrates the model elements and variables along with sign criteria followed. Servovalve spool motion has been approximated by the following first-order dynamic system24 Cspusv= tsp _ysp+ ysp \u00f01\u00de where usv is the input voltage to the servovalve, ysp is the spool position, Csp is the spool static gain, and tsp is the time constant of the system" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_35_0000421_lcomm.2021.3078469-Figure1-1.png", + "original_path": "designv11-35/openalex_figure/designv11_35_0000421_lcomm.2021.3078469-Figure1-1.png", + "caption": "Fig. 1: System Model", + "texts": [ + "cn) dimension, deep deterministic policy gradient (DDPG) proposed in the paper can search the best movements for UAVs in a continuous domain and optimize the trajectory even when there exist obstacles. \u2022 Game theory is introduced to solve the service assignment problem. Nash equilibrium (NE) is proved to be exist by distributed updating each UAV\u2019s decision, which greatly reduces the computation complexity and convergence time of DDPG. \u2022 UAVs\u2019 fairness is also guaranteed during the dynamic data offloading. As a result, the total offloaded data size by each UAV is approximately the same. II. MULTI-UAV ASSISTED OFFLOADING SCHEME As shown in Fig.1, there are N = {1, 2, \u00b7 \u00b7 \u00b7 , N} UAVs and M = {1, 2, \u00b7 \u00b7 \u00b7 , M} GUs in the system. Their position is measured by Cartesian coordinate, which is denoted by Cn = [cxn, cyn, czn] and C \u2032m = [c \u2032 xm, c \u2032 ym ] respectively. The total data size of GUs is L = {`1, `2, \u00b7 \u00b7 \u00b7 , `m}. Suppose that the communication channels follow quasi-static block fading, i.e., channels\u2019 state remains unchanged in a transmission frame. Let U = [un,m] be the service assignment, where un,m \u2208 {0, 1} and un,m = 1 means UAV n provide data relay to GU m", + " UAVs and GUs are equipped with one antenna each, which means data can only be transmitted and received asynchronously, and there are constraints \u2211M m=1 un,m \u2264 1 and \u2211N n=1 un,m \u2264 1. A. Data Offloading In the wireless communication networks, since the computation result of an intensive task executed by the remote server, possibly at ground base station or cloud, is considered to be very small, it is insignificant compared with the size of task itself [7]. So the paper only considers the offloading delay, i.e., ground-to-air (G2A) link in Fig.1. The distance between GU m and UAV n in time slot k is: dn,m[k] = \u221a cz2 n+ \u2016 [cxn, cyn] \u2212 [c \u2032 xm, c\u2032ym] \u2016 2, (1) where \u2016 denotes the Euclidean norm. For LoS communication, the channel gain of UAV n to GU m in a frame is hn,m[k] = \u03c10dn,m[k]\u2212\u03b1, where \u03c10 denotes the reference channel power gain at d0 = 1m, and \u03b1 \u2265 2 is the path loss exponent. The corresponding received signal-to-interference-plus-noise ratio Authorized licensed use limited to: Carleton University. Downloaded on June 01,2021 at 02:34:16 UTC from IEEE Xplore" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_35_0000116_er.6543-Figure1-1.png", + "original_path": "designv11-35/openalex_figure/designv11_35_0000116_er.6543-Figure1-1.png", + "caption": "FIGURE 1 Structure of soft-connected battery module (A, structure of battery module assembly; B, structure of battery module multibody model; C, multibody model with end panel in Abaqus) [Colour figure can be viewed at wileyonlinelibrary.com]", + "texts": [ + " Thereafter, parameters of multibody model are identified by means of optimization method based on the modal test. In the same section, the error between multibody model simulation and experimental result is discussed and a correction of parameters according to structure feature is proposed to make the model more accurate. Finally, a battery pack assembly is modeled using this novel simplified method, and the superiority of the simplified method is presented by comparing the novel method with the traditional one. The battery module in this research (Figure 1A) is composed of 18 650 cylindrical batteries, which are connected by pluggable structure. This connecting structure makes it convenient for the battery module to replace faulted batteries. When the battery pack is severely deformed by accident, the structure will disconnect batteries to avoid short circuit. The structure of soft-connected battery module can be divided into the following parts: (a) positive and negative current collector plate at both ends; (b) positive and negative panels for clamping battery module in axis direction; (c) anode and cathode bus board for parallel connection at both ends; (d) plastic fixture to support batteries in each layer; (e) claws to hold batteries and connect batteries in series; (f) metal sheet used for connecting batteries in parallel in each layer", + " It should be noted that the multibody method ignores the shape of batteries, so the method is only suitable for cases in which the battery pack does not deform greatly and batteries do not touch each other. According to the force-displacement curve of 18 650 battery studied by Elham Sahraei,6-8 the radial stiffness of the cell at initial deformation is 300 N/mm approximately, which is much larger than the stiffness of sheet spring structure in claw. This means that the batteries will only translate and rotate without deformation when the battery pack vibrates, so it is appropriate to treat the battery as a rigid body. The structure of the multibody model of battery module is presented in Figure 1B. The equivalent structure is a rigid body, which integrates the mass of battery, claw, and plastic fixture. In FEM software, the equivalent structure is composed of mass element, rigid surface, and some nodes that mark the location of the connection. The elastic property of claw and plastic fixture is realized by elastic element called bushing in the multibody model. The bushing has six parameters to present the stiffness in six directions and has no detailed structure. Nodes of equivalent structure are connected by bushing in series and parallel, corresponding to the series and parallel connection of the batteries in the battery module", + " To make it convenient for combining the FEM model of enclosure and analyzing the assembly of battery pack, the multibody model is realized in Abaqus 6.13-4. In addition to the battery and connection components, the battery module contains panels at the end. The FEM model of panel at the end of battery module is connected with multibody model to get the simplified model of battery module. The panel is simplified to a flat plate, and the free node of the end bushing in the multibody model is connected with the closest nodes of the panel by the rigid beam element. The visualization of the model in Abaqus is shown in Figure 1C. Figure 4A shows the parts of vibration modes from simulation using this model, which can accurately correspond to the results obtained by modal tests in the next section. Free modal experiment of the battery module was carried out to obtain the modal characteristics of the battery module and get the stiffness parameters of the multibody model. There are three kinds of small size battery modules selected as specimens, represented by the number of rows, columns, and layers: module-1: 4_12_3; module-2: 10_5_3; and module-3: 8_6_3" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_35_0001419_978-94-007-7401-8-Figure5.86-1.png", + "original_path": "designv11-35/openalex_figure/designv11_35_0001419_978-94-007-7401-8-Figure5.86-1.png", + "caption": "Fig. 5.86 2PaPaRRR-1PaPaRR-1CPaPa-type fully-parallel PM with uncoupled Sch\u00f6nflies motions defined by MF = SF = 4, RF\u00f0 \u00de \u00bc \u00f0v1; v2; v3;xb\u00de, TF = 0, NF = 28, limb topology Pa\\Pa||R||R\\kR, Pa\\Pa||R||R and C||Pa||Pa", + "texts": [ + " 5.40) 21. 3PaPaPR1CPaPat (Fig. 5.83) Pa\\Pa\\kP\\\\R (Fig. 5.4f) Idem No. 15 Pa\\Pa\\kP\\kR (Fig. 5.4h) C||Pa||Pat (Fig. 5.4p) 22. 3PaPaPR1CPaPat (Fig. 5.84) Pa\\Pa\\\\P\\\\R (Fig. 5.4g) Idem No. 15 Pa\\Pa\\\\P\\kR (Fig. 5.4e) C||Pa||Pat (Fig. 5.4p) 23. 2PaPaRRR1PaPaRR1CPaPa (Fig. 5.85) Pa\\Pa||R||R\\\\R (Fig. 5.5a) The revolute joint between links 6D and 10 of G4 limb and the last revolute joints of limbs G1, G2 and G3 have superposed axes Pa\\Pa||R||R (Fig. 5.4i) C||Pa||Pa (Fig. 5.40) 24. 2PaPaRRR1PaPaRR1CPaPa (Fig. 5.86) Pa\\Pa||R||R\\kR (Fig. 5.5b) Idem No. 23 Pa\\Pa||R||R (Fig. 5.4i) C||Pa||Pa (Fig. 5.40) 25. 2PaPaRRR1PaPaRR1CPaPat (Fig. 5.87) Pa\\Pa||R||R\\\\R (Fig. 5.5a) Idem No. 23 Pa\\Pa||R||R (Fig. 5.4i) C||Pa||Pat (Fig. 5.4p) 26. 2PaPaRRR1PaPaRR1CPaPat (Fig. 5.88) Pa\\Pa||R||R\\kR (Fig. 5.5b) Idem No. 23 Pa\\Pa||R||R (Fig. 5.4i) C||Pa||Pat (Fig. 5.4p) 27. 3PaPaPaR1CPaPa (Fig. 5.89) Pa\\Pa||Pa\\kR (Fig. 5.6a) The revolute joint between links 6D and 11 of G4 limb and the last revolute joints of limbs G1, G2 and G3 have superposed axes Pa\\Pa||Pa||R (Fig" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_35_0000149_s11548-021-02338-9-Figure4-1.png", + "original_path": "designv11-35/openalex_figure/designv11_35_0000149_s11548-021-02338-9-Figure4-1.png", + "caption": "Fig. 4 Forceps handle shaped operating device (Leader): a photograph of the prototype device with embedded encoders with five DOFs, and b schematic diagram of the device and the joystick coordinates", + "texts": [ + " The free movement of the leader device functions 1 3 as the clutch for the differential workspace between devices. Point-to-point commands to control the individual axes of the follower device driven by motors are input by the appropriate axes of the leader device rotating encoder, and these commands are processed by the provider, which outputs the appropriate signals to the follower device. The forceps handle shaped operating device with embedded encoders with five DOFs in the joystick coordinates (Leader) is shown in Fig.\u00a04. The mechanism of the device, which can be draped with a sterilized cover, is a magnetic ball joint and fixed gimbals for the pitch and yaw axes, a linear guide for the insertion axis, a circular dial for the roll axis, and a handle for the grasp axis. The linear displacement of the insertion axis and each rotary angle of the other four axes is detected by a linear encoder (resolution: 0.04\u00a0mm, MLS-12-1500-250, MTL, Sagamihara, Japan) and rotary encoders (resolution: 32,768 ppr, MES-12-2048PST16)" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_35_0001419_978-94-007-7401-8-Figure2.14-1.png", + "original_path": "designv11-35/openalex_figure/designv11_35_0001419_978-94-007-7401-8-Figure2.14-1.png", + "caption": "Fig. 2.14 Fully-parallel PMs with coupled Sch\u00f6nflies motions of types 4RRRPR (a) and 4RRPRR (b) defined by MF \u00bc SF \u00bc 4; RF\u00f0 \u00de \u00bc \u00f0v1; v2; v3;xd\u00de, TF = 0, NF = 2, limb topology R\\R||R||P||R (a) and R\\R||P||R||R||R (b)", + "texts": [ + " 2.1q) Idem No. 13 19. 4RRRRP (Fig. 2.11a) R||R||R\\R\\P (Fig. 2.1r) The before last revolute joints of the four limbs have parallel axes 20. 4RPRRR (Fig. 2.11b) R\\P\\kR\\R||R (Fig. 2.1s) Idem No. 3 21. 4RRRRP (Fig. 2.12a) R\\R||R||R||P (Fig. 2.1t) The first revolute joints of the four limbs have parallel axes 22. 4RRRPR (Fig. 2.12b) R\\R||R||P||R Fig. 2.1u) Idem No. 21 23. 4RRPRR (Fig. 2.13a) R\\R||P||R||R (Fig. 2.1v) Idem No. 21 24. 4RPRRR (Fig. 2.13b) R\\P||R||R||R (Fig. 2.1w) Idem No. 21 25. 4RRRPR (Fig. 2.14a) R\\R||R||P||R (Fig. 2.1x) Idem No. 21 26. 4RRPRR (Fig. 2.14b) R\\R||P||R||R||R (Fig. 2.1y) Idem No. 21 27. 4RPRRP (Fig. 2.15a) R\\P\\kR\\R\\P (Fig. 2.1z) Idem No. 19 28. 4PRRRP (Fig. 2.15b) P\\R||R\\R\\P (Fig. 2.1z1) Idem No. 19 29. 4PRRPR (Fig. 2.16) P\\R\\R\\P\\kR (Fig. 2.1a0) Idem No. 5 30. 4CRRR (Fig. 2.17a) C\\R||R\\R (Fig. 2.1d0) Idem No. 1 (continued) 2.1 Topologies with Simple Limbs 59 Table 2.1 (continued) No. PM type Limb topology Connecting conditions 31. 4CRRR (Fig. 2.17b) C||R||R\\R (Fig. 2.1e0) Idem No. 13 32. 4RRCR (Fig. 2.18a) R||R||C\\R (Fig. 2.1f0) Idem No" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_35_0000027_j.procir.2020.05.200-Figure5-1.png", + "original_path": "designv11-35/openalex_figure/designv11_35_0000027_j.procir.2020.05.200-Figure5-1.png", + "caption": "Fig. 5. Simplified simulation of the temperature distribution in the semi-finished bearing bushing during inductive heating (a) without and (b) with active cooling.", + "texts": [ + " The temperature distribution is important, especially or the bearing bushing blank, because it is inductively heated rom inside. The problem is that the aluminium on the outside s heated strongly by the thermal conduction, as can be seen in ig. 5 a. The aluminium comes close to its melting temperature of 50 \u25e6C . By heating with activated cooling, however, the temperaure of the aluminium remains within a normal range of tempera- Initial temperature of material (before heating): 20 \u00b0C t b r t T \u03c1 4 I n o s o w j T s ure ( Fig. 5 b). The disadvantage of this, is that the heat conduction etween steel an aluminium leads to cooling of the steel side. This esults in an optimization problem for the right settings to cool he aluminum while keeping the steel at the required temperature. his problem must be investigated further. c v \u2202T \u2202t \u2212 \u2207 (\u03bb\u2207 T ) = N \u2211 \u02d9 qi (1) i =0 t . Automated process design The system presented so far is not able to grip independently. n order to automate the process of handling, a control system is eeded" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_35_0001419_978-94-007-7401-8-Figure5.11-1.png", + "original_path": "designv11-35/openalex_figure/designv11_35_0001419_978-94-007-7401-8-Figure5.11-1.png", + "caption": "Fig. 5.11 3PaPPR-1RUPU-type fully-parallel PMs with uncoupled Sch\u00f6nflies motions defined by MF = SF = 4, RF\u00f0 \u00de \u00bc \u00f0v1; v2; v3;xb\u00de, TF = 0, NF = 13, limb topology R\\R\\R\\P\\kR\\R and Pa||P\\P\\kR, Pa||P\\P\\\\R (a), Pa\\P\\kP\\kR, Pa\\P\\kP\\\\R (b)", + "texts": [ + "1 Fully-Parallel Topologies 515 516 5 Topologies with Uncoupled Sch\u00f6nflies Motions 5.1 Fully-Parallel Topologies 517 518 5 Topologies with Uncoupled Sch\u00f6nflies Motions 5.1 Fully-Parallel Topologies 519 520 5 Topologies with Uncoupled Sch\u00f6nflies Motions 5.1 Fully-Parallel Topologies 521 522 5 Topologies with Uncoupled Sch\u00f6nflies Motions 5.1 Fully-Parallel Topologies 523 524 5 Topologies with Uncoupled Sch\u00f6nflies Motions 5.1 Fully-Parallel Topologies 525 526 5 Topologies with Uncoupled Sch\u00f6nflies Motions 7. 3PaPPR-1RUPU (Fig. 5.11a) Pa||P\\P\\kR (Fig. 5.2a) The first revolute joint of G4 limb and the last revolute joints of limbs G1, G2 and G3 have parallel axes. The last joints of limbs G1, G2 and G3 have superposed axes Pa||P\\P\\\\R (Fig. 5.2d) R\\R\\R\\P\\kR\\R (Fig. 5.1b) 8. 3PaPPR-1RUPU (Fig. 5.11b) Pa\\P\\kP\\kR (Fig. 5.2b) Idem No. 7 Pa\\P\\kP\\\\R (Fig. 5.2c) R\\R\\R\\P\\kR\\R (Fig. 5.1b) 9. 2PaPRRR1PaPRR1RUPU (Fig. 5.12a) Pa\\P\\\\R||R\\\\R (Fig. 5.3a) The first revolute joint of G4 limb and the last revolute joints of limbs G1, G2 and G3 have parallel axes. The last joints of of limbs G1, G2 and G3 have superposed axes. The revolute joints between links 5 and 6 of of limbs G1, G2 and G3 have orthogonal axes Pa\\P\\\\R||R (Fig. 5.2e) R\\R\\R\\P\\kR\\R (Fig. 5.1b) 10. 2PaPRRR1PaPRR1RUPU (Fig. 5.12b) Pa||P\\R||R\\\\R (Fig" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_35_0001419_978-94-007-7401-8-Figure2.15-1.png", + "original_path": "designv11-35/openalex_figure/designv11_35_0001419_978-94-007-7401-8-Figure2.15-1.png", + "caption": "Fig. 2.15 Fully-parallel PMs with coupled Sch\u00f6nflies motions of types 4RPRRP (a) and 4PRRRP (b) defined by MF \u00bc SF \u00bc 4; RF\u00f0 \u00de \u00bc \u00f0v1; v2; v3;xd\u00de, TF = 0, NF = 2, limb topology R\\P\\kR\\R\\P (a) and P\\R||R\\R\\P (b)", + "texts": [ + "1r) The before last revolute joints of the four limbs have parallel axes 20. 4RPRRR (Fig. 2.11b) R\\P\\kR\\R||R (Fig. 2.1s) Idem No. 3 21. 4RRRRP (Fig. 2.12a) R\\R||R||R||P (Fig. 2.1t) The first revolute joints of the four limbs have parallel axes 22. 4RRRPR (Fig. 2.12b) R\\R||R||P||R Fig. 2.1u) Idem No. 21 23. 4RRPRR (Fig. 2.13a) R\\R||P||R||R (Fig. 2.1v) Idem No. 21 24. 4RPRRR (Fig. 2.13b) R\\P||R||R||R (Fig. 2.1w) Idem No. 21 25. 4RRRPR (Fig. 2.14a) R\\R||R||P||R (Fig. 2.1x) Idem No. 21 26. 4RRPRR (Fig. 2.14b) R\\R||P||R||R||R (Fig. 2.1y) Idem No. 21 27. 4RPRRP (Fig. 2.15a) R\\P\\kR\\R\\P (Fig. 2.1z) Idem No. 19 28. 4PRRRP (Fig. 2.15b) P\\R||R\\R\\P (Fig. 2.1z1) Idem No. 19 29. 4PRRPR (Fig. 2.16) P\\R\\R\\P\\kR (Fig. 2.1a0) Idem No. 5 30. 4CRRR (Fig. 2.17a) C\\R||R\\R (Fig. 2.1d0) Idem No. 1 (continued) 2.1 Topologies with Simple Limbs 59 Table 2.1 (continued) No. PM type Limb topology Connecting conditions 31. 4CRRR (Fig. 2.17b) C||R||R\\R (Fig. 2.1e0) Idem No. 13 32. 4RRCR (Fig. 2.18a) R||R||C\\R (Fig. 2.1f0) Idem No. 13 33. 4RCRR (Fig. 2.18b) R||C||R\\R (Fig. 2.1g0) Idem No. 13 34. 4RRRC (Fig. 2.19a) R\\R||R||C (Fig. 2.1h0) Idem No. 21 35" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_35_0000246_tte.2021.3068819-Figure18-1.png", + "original_path": "designv11-35/openalex_figure/designv11_35_0000246_tte.2021.3068819-Figure18-1.png", + "caption": "Fig. 18. 3D temperature field analysis of stator system, (a) thermal model, (b) cross section of stator system.", + "texts": [ + " However, the increase in loss ratio of \u03b8=360\u00b0 is not obvious. Due to the neglect of the winding ending effect, the accuracy of 2D-EMM will reduce. Through the comparison between simulation result and test result of \u03b8=360\u00b0, the loss ratio in test result will higher at high frequency. In this section, a 20kW high-speed SRM thermal model was build and finite element method was used to analyze temperature rise under different conditions. 3D thermal model and cross section of stator system are shown in Fig. 18. Temperature of hot spot in slot, stator tooth and stator yoke were analyzed and tested by PT-100. T1 represents the temperature of the winding in the slot center, the slot fill factor can reach 70.4%, T2 and T3 represent the temperature of the amorphous alloy stator tooth and yoke, respectively. Fig. 19 shows thermal model calculation results at different speeds. Temperature rise of the winding is always the highest in the stator system. When the motor operates at the rated speed (12000r/min, 1600Hz), there will be a difference of 22\u2103 with or without considering AC loss" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_35_0001419_978-94-007-7401-8-Figure3.33-1.png", + "original_path": "designv11-35/openalex_figure/designv11_35_0001419_978-94-007-7401-8-Figure3.33-1.png", + "caption": "Fig. 3.33 Overactuated PMs with coupled Sch\u00f6nflies motions of types 4RRPRR (a) and 4RRRPR (b) defined by MF \u00bc SF \u00bc 4; RF\u00f0 \u00de \u00bc v1; v2; v3;xd\u00f0 \u00de, TF = 0, NF = 2, limb topology R\\R\\P\\kR\\R (a) and R\\R||R\\P\\kR (b)", + "texts": [ + "3e) Idem No. 17 19. 4RPRRR (Fig. 3.29b) R||P||R\\R||R (Fig. 3.3f) Idem No. 17 20. 4RPRRR (Fig. 3.30a) R||P||R\\R||R (Fig. 3.3g) Idem No. 17 21. 4RRPRR (Fig. 3.30b) R||R||P\\R||R (Fig. 3.3h) Idem No. 17 22. 4RRRPR (Fig. 3.31a) R||R\\R\\P\\kR (Fig. 3.3i) Idem No. 17 23. 4RRRRP (Fig. 3.31b) R||R\\R||R\\kP (Fig. 3.3j) Idem No. 17 24. 4PRRRR (Fig. 3.32a) P\\R\\R||R\\R (Fig. 3.3k) The second and the last joints of the four limbs have parallel axes 25. 4RPRRR (Fig. 3.32b) R\\P\\R||R\\R (Fig. 3.3l) Idem No. 14 26. 4RRPRR (Fig. 3.33a) R\\R\\P\\kR\\R (Fig. 3.3m) Idem No. 14 27. 4RRRPR (Fig. 3.33b) R\\R||R\\P\\kR (Fig. 3.3n) Idem No. 14 28. 4RRRRP (Fig. 3.34a) R\\R||R\\R\\P (Fig. 3.3o) The first and the fourth revolute joints of the four limbs have parallel axes 29. 4PRRRR (Fig. 3.34b) P\\R||R\\R||R (Fig. 3.3p) The second joints of the four limbs have parallel axes (continued) 244 3 Overactuated Topologies with Coupled Sch\u00f6nflies Motions Table 3.2 (continued) No. PM type Limb topology Connecting conditions 30. 4RPRRR (Fig. 3.35a) R\\P\\kR\\R||R (Fig. 3.3q) Idem No. 16 31. 4RRRPR (Fig. 3.35b) R||R\\R||P||R (Fig" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_35_0000165_s40430-021-02894-w-Figure3-1.png", + "original_path": "designv11-35/openalex_figure/designv11_35_0000165_s40430-021-02894-w-Figure3-1.png", + "caption": "Fig. 3 Blade profiles for the generation of virtual surfaces", + "texts": [ + " If the translational velocity of the virtual surfaces satisfies equation vc = rp1\u03c9b, then the tooth surfaces of the curvilinear gear will be generated. All convex sides of the pinion are processed by the inner blade, after which the cutter is replaced with the outer blade to generate the concave side. The centre distance between the central axes of the two cutters is \u03c0m/2 along the direction of the translational velocity of the cutter (see Fig.\u00a02). The same method is applied to the wheel tooth surfaces. 1 3 Figure\u00a03 shows the profiles of the blades that are employed for the generation of the cutter surfaces, where u1 and \u03bb1 represent the variable position of a point on the straight profile and edge profile, respectively. The coordinate systems Sib(xib, y ib, zib) and Sob(xob, yob, zob) are fixed to the inner blade and outer blade, respectively. The system axes y ib and yob are directed along the blade profiles towards the addendum of the blades, and their origins Oib and Oob are located on the pitch planes", + " The position vector of a point on the straight profile can be expressed in the coordinate systems Sib and Sob as follows: (1) ib(u1) = ob(u1) = \u23a1\u23a2\u23a2\u23a2\u23a3 0 u1 0 0 \u23a4\u23a5\u23a5\u23a5\u23a6 The coordinate transformations from Sib and Sob to Sb are represented by the matrix The position vector of a point on the edge profile can be expressed in the coordinate systems Sifb and Sofb, as follows: The coordinate transformations from Sifb and Sofb to Sb are represented by the matrix To avoid severe contact stresses caused by edge contact, the tip relief parabola geometry of the blade profiles is applied to the design of the curvilinear gear sets. In Fig.\u00a03, at represents the tip relief parabola coefficient. When the parabola coefficient at = 0, the gear has not been modified on the addendum. ht represents the tip relief height, and ut (2) b,ib = \u23a1 \u23a2\u23a2\u23a2\u23a3 cos n sin n 0 m 4 \u2212 sin n cos n 0 0 0 0 1 0 0 0 0 1 \u23a4 \u23a5\u23a5\u23a5\u23a6 (3) b,ob = \u23a1\u23a2\u23a2\u23a2\u23a3 cos n \u2212 sin n 0 \u2212 m 4 sin n cos n 0 0 0 0 1 0 0 0 0 1 \u23a4\u23a5\u23a5\u23a5\u23a6 (4) ifb( 1) = \u23a1\u23a2\u23a2\u23a2\u23a3 m cos 1 \u2212 m sin 1 0 1 \u23a4\u23a5\u23a5\u23a5\u23a6 (5) ofb( 1) = \u23a1\u23a2\u23a2\u23a2\u23a3 m cos 1 \u2212 m sin 1 0 1 \u23a4\u23a5\u23a5\u23a5\u23a6 (6) b,ifb = \u23a1 \u23a2\u23a2\u23a2\u23a3 1 0 0 m 4 \u2212 tan n( m sin n + bm \u2212 m) \u2212 m cos n 0 1 0 \u2212(bm \u2212 m) 0 0 1 0 0 0 0 1 \u23a4\u23a5\u23a5\u23a5\u23a6 (7) b,ofb = \u23a1\u23a2\u23a2\u23a2\u23a3 1 0 0 m 4 \u2212 tan n( m sin n + bm \u2212 m) \u2212 m cos n 0 1 0 \u2212(bm \u2212 m) 0 0 1 0 0 0 0 1 \u23a4\u23a5\u23a5\u23a5\u23a6 Journal of the Brazilian Society of Mechanical Sciences and Engineering (2021) 43:183 1 3 Page 5 of 18 183 represents the variable position of a point on the parabola profile. In Fig.\u00a03 (a), the three axes of the coordinate system Sit(xit, yit, zit) are parallel to the xib, yib and zib axes of the coordinate system Sib(xib, yib, zib), respectively. The origin Oit of the coordinate system Sot is separated with respect to Oib a distance (am-ht)/cos\u03b1n along axis yib. The definition of the coordinate systems in Fig.\u00a03 (b) is similar to that in Fig.\u00a03 (a). A point laying on the parabola profile is expressed in the coordinate system Skt(xkt, ykt, zkt) as follows: The upper sign in Eq.\u00a0(8) represents a point on the outer blade, and the lower sign represents a point on the inner blade. The position vector of a point on the parabola profile is expressed in the coordinate systems Sib and Sob as follows: Here, The virtual surfaces of the cutters are generated by the profiles of the blades rotating around the centre axes of the cutters. Figure\u00a04 shows the coordinate transformation from Sb(xb, yb, zb) to Sc(xc, yc, zc)" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_35_0000362_j.ijthermalsci.2021.107000-Figure4-1.png", + "original_path": "designv11-35/openalex_figure/designv11_35_0000362_j.ijthermalsci.2021.107000-Figure4-1.png", + "caption": "Fig. 4. Thermal modeling of the sub- and main case for the estimation of heat transfer at the static and moving interface.", + "texts": [ + " The rotation causes forced convection of the surrounding air and leads to heat losses at the front surface, but as velocities in magnitude or 1 m s occur, these are expected to have minor impact and are neglected. As the housing is not rotating, corresponding x- and y-velocity components are set to zero. An exemplary temperature and velocity field is visualized in Fig. 3 showing velocities of about 0.2 m s . u\u2192=\u03c9 \u00d7 r\u2192 (4) Besides the velocity components, also the boundary conditions need to be specified, in this case four boundary types (\u03a91 - \u03a94) are defined for the thermal modeling of the bearing system (Fig. 4). The first boundary type \u03a91 uses the temperature data from the IRrecording to define the inlet temperature of the convective fluxes. Also the outer boundary temperature of the housing and rig are set according to the recorded temperatures at this location. This temperature information is only used as a boundary condition and is not further considered in the objective function. T(x, y, t) =Ybj (x, y, t) at \u03a91 (5) At \u03a92, an outlet boundary condition is applied which is a common approach in numerical methods to model the convective outflow" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_35_0001730_s0025557200233147-Figure15-1.png", + "original_path": "designv11-35/openalex_figure/designv11_35_0001730_s0025557200233147-Figure15-1.png", + "caption": "FIG. 15.", + "texts": [], + "surrounding_texts": [ + "MECHANISMS IN T H R E E DIMENSIONS 19\nIn Form I, if there is sliding as well as turning at the hinges C, B, (\", B' there may be a slight difference of plane between the two shafts to be joined.\nD\n<0 C\nForm 1. Fid . 14.\n(a is small.'\nA model of Hooke's joint good enough for the drive to the propellor of a model boat, say, can be constructed from wire, preferably in Form I I .\nWe have already found for Hooke's joint tha t tan 8 = tan cj). cos a, where 6 = cot, cf>=a/t.\nDifferentiating, \u2014 = =\u20145 5-r \u2022 to 1 - sin^ct. cos\"1 a\nI t may be convenient to have this velocity ratio in the form of an expansion ; this is most quickly obtained as follows :\nco'jw = (1 - sin2a)s (1 - sin2 a cos2 d)-1\n= (I - \\ sin2 a - \\ sin4 a) (1 + sin2a cos2 6 + sin4 a cos4 6),\nneglecting powers of a above the fourth, since a is small,\n= 1 + sin2 a( - -1- + cos2 0) + s i n 4 \u00ab . { - \\ - \\ cos2 6 + cos46) = 1 + s i n 2 a ( | cos 29) + sin4 a ( \\ cos 26 + } cos id).\nThus\nw'jco = 1 +cos 2a>t(l sin2 a + I sin4 a + ...) + cos 4a)^(| sin4 a + . . . ) .\nI t might interest readers to consider for themselves the remark able motions which occur as a approaches 90\u00b0 in Hooke's joint and also the case in which a =90\u00b0, this being a particular case of those spheric chains in which a = y = (180 -jS) = (180 - 8 ) .\nThe Skew Jsogram. A mechanism for connecting two skew shafts, not coplanar, can be formed from the Skew Isogram ; the case for which the shafts are perpendicular is shown below.\nPlane ABB is fixed, a = 90, 0 = sin-1 {b/a).\nhttps://doi.org/10.2307/3606015 11 Feb 2020 at 10:34:09, subject to the Cambridge Core terms of use, available at https://www.cambridge.org/core/terms. Downloaded from https://www.cambridge.org/core. Conrad Grebel University College, University of Waterloo, on", + "20 T H E MATHEMATICAL GAZETTE\nThe hinge-line at G is always perpendicular to the hinge-line a t / ) .\nThe mechanism is remarkable since it transmits a positive torque, even in the zero positions. (There are no \" dead centres \" and no \" change points \".)\nIf co' is constant (driver) and the fluctuation of a> is not to exceed n% of the mean w, on account of excessive inertia forces,\nl/a+b\\ lfa-b\\ n\nand ajb = J'{4 (100/w)2 + 1} = 200/w, approximately.\nIf n=25, afa8b, which is reasonable. w = 10, \u00abi=a20&, which is possible. ra = 5, a \u00ab 4 0 6 , which is impracticable.\nAs also with Hooke's joint, a constant velocity ratio may be obtained by placing two similar Skew Isograms in series so that the second balances out the inequalities of the first.\nA small working model can quickly be made from wire, using a cardboard box for the bearings.\nAlthough the mechanism is not, perhaps, very suitable for power transmission due to the considerable stresses which are liable to be set up in the connecting rod and elsewhere, one would have thought it would have been more used as a movement in machinery, for there must surely be occasions on which it will provide a required motion in fewer pieces than any other mechanism. Perhaps, with the advent of plastics as materials for construction, engineering design may tend to become less rectangular and find room more easily for mechanisms such as this. R. H. M.\nhttps://doi.org/10.2307/3606015 11 Feb 2020 at 10:34:09, subject to the Cambridge Core terms of use, available at https://www.cambridge.org/core/terms. Downloaded from https://www.cambridge.org/core. Conrad Grebel University College, University of Waterloo, on" + ] + }, + { + "image_filename": "designv11_35_0001419_978-94-007-7401-8-Figure5.19-1.png", + "original_path": "designv11-35/openalex_figure/designv11_35_0001419_978-94-007-7401-8-Figure5.19-1.png", + "caption": "Fig. 5.19 2PaPRRR-1PaPRR-1RPaPaP-type fully-parallel PMs with uncoupled Sch\u00f6nflies motions defined by MF = SF = 4, RF\u00f0 \u00de \u00bc \u00f0v1; v2; v3;xb\u00de, TF = 0, NF = 19, limb topology R||Pa||Pa||P and Pa\\P\\\\R||R\\\\R, Pa\\P\\\\R||R (a), Pa||P\\R||R\\\\R, Pa||P\\R||R (b)", + "texts": [ + "2c) R||P||Pa||Pa (Fig. 5.4m) 5. 3PaPPR1RPaPatP (Fig. 5.17a) Pa||P\\P\\kR (Fig. 5.2a) Idem No. 1 Pa||P\\P\\\\R (Fig. 5.2d) R||Pa||Pat||P (Fig. 5.4l) 6. 3PaPPR1RPaPatP (Fig. 5.17b) Pa\\P\\kP\\kR (Fig. 5.2b) Idem No. 1 Pa\\P\\kP\\\\R (Fig. 5.2c) R||Pa||Pat||P (Fig. 5.4l) 7. 3PaPPR-1RPPaPat (Fig. 5.18a) Pa||P\\P\\kR (Fig. 5.2a) Idem No. 3 Pa||P\\P\\\\R (Fig. 5.2d) R||P||Pa||Pat (Fig. 5.4n) 8. 3PaPPR-1RPPaPat (Fig. 5.18b) Pa\\P\\kP\\kR (Fig. 5.2b) Idem No. 3 Pa\\P\\kP\\\\R (Fig. 5.2c) R||P||Pa||Pat (Fig. 5.4n) 9. 2PaPRRR-1PaPRR1RPaPaP (Fig. 5.19a) Pa\\P\\R||R\\\\R (Fig. 5.3a) The last joints of the four limbs have superposed axes/directions. The revolute joints between links 5 and 6 of limbs G1, G2 and G3 have orthogonal axes Pa\\P\\\\R||R (Fig. 5.2e) R||Pa||Pa||P (Fig. 5.4k) 10. 2PaPRRR-1PaPRR1RPaPaP (Fig. 5.19b) Pa||P\\R||R\\\\R (Fig. 5.3b) Idem No. 9 Pa||P\\R||R (Fig. 5.2f) R||Pa||Pa||P (Fig. 5.4k) (continued) 528 5 Topologies with Uncoupled Sch\u00f6nflies Motions Table 5.2 (continued) No. PM type Limb topology Connecting conditions 11. 2PaPRRR-1PaPRR1RPPaPa (Fig. 5.20a) Pa\\P\\\\R||R\\\\R (Fig. 5.3a) Idem No. 9 Pa\\P\\\\R||R (Fig. 5.2e) R||P||Pa||Pa (Fig. 5.4m) 12. 2PaPRRR-1PaPRR1RPPaPa (Fig. 5.20b) Pa||P\\R||R\\\\R (Fig. 5.3b) Idem No. 9 Pa||P\\R||R (Fig. 5.2f) R||P||Pa||Pa (Fig. 5.4m) 13. 2PaPRRR-1PaPRR1RPaPatP (Fig. 5" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_35_0000174_j.matpr.2021.01.640-Figure1-1.png", + "original_path": "designv11-35/openalex_figure/designv11_35_0000174_j.matpr.2021.01.640-Figure1-1.png", + "caption": "Fig. 1. Conventional steel drive shaft arrangement.", + "texts": [ + " Evaluation of young\u2019s modulus for anisotropic materials [3]. Design of mechanical power transmitting element drive shaft with composite materials and applicable to various vehicles [4]. The spicer u-joint division of dana company for the portage Econoline van models built up the primary composite propeller shaft in 1985. the prospective for composites in supporting car applications as essential from a basic perspective applications [5,6] and one part car half and half drive shaft of composite\\ aluminium is appeared in Fig. 1 was produced among other assembling technique, in which a carbon fiber epoxy composite layer was co-cured inside the aluminium tube as an alternative of wrap over outer surface to protect the composite layer from outside effect due to environmental factors. The spicer u-joint division of company for the portage Econoline van models ng 2020: built up the main composite propeller shaft in 1984. The hypothetical subtle elements of composite materials and composite structures widely evaluated [7] and rapidly fabricate the application oriented parts with potential outcomes of composites in the field of Automotive industry particularly fabricate composite elliptic springs" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_35_0000775_j.engfailanal.2021.105672-Figure4-1.png", + "original_path": "designv11-35/openalex_figure/designv11_35_0000775_j.engfailanal.2021.105672-Figure4-1.png", + "caption": "Fig. 4. Fracture surface of the cross tube: A - initiation site; B - chevron marks; C - area of final fracture.", + "texts": [ + " The post-accidental visual inspection of the failed landing gear revealed that the aft cross tube was fractured into two pieces. The broken, separated part was removed and taken for examination. Fig. 3 shows fractured aft cross tube and fracture surface. Visual analysis of fractured surface performed on macroscopic level, Fig. 3., revealed that the fracture was initiated at the bottom, position 6o\u2019clock, on the inner surface of the tube. The characteristic beach marks are observed indicating fatigue fracture (areas A in the Fig. 4 and Fig. 5), where crack propagated up to a half of the wall thickness. This region is clearly separate from rougher region with V-shape chevron marks (area B in the Fig. 4). Tracing back the observed chevron marks, the fracture origin area was also indicated to be in the sector A of fracture surface. The region opposite to cracks initiation area (marked area C in the Fig. 4) is the final, fast fracture area. Macroscopic examination of the aft cross tube surface revealed numerous deep scratches as the inner surface defects. These preexisting cracks are identified to be in the circumferential and in the longitudinal directions, where some of them are shown in Fig. 6. In Fig. 7 one of the initiation places, is shown which is presented in Fig. 5. Crack initiation has semi elliptical shape, approx. 2.3 mm and 0.7 mm deep in the tube wall (red arrow). Further propagation is revealed by presence of typical ratchet marks (yellow arrows). The crack was initiated in the inner side of the cross tube and propagated through tube wall. Observation of crack propagation area at low magnification by SEM, revealed a numerous secondary cracks (yellow arrows in Fig. 8). The rough mate region identified as overload (Fig. 9) revealed transgranular fracture surface, with secondary cracking on coarse second phase particles (zone C, Fig. 4). Additional, longitudinal cut, of the cross tube in the vicinity of fracture has been conducted in order to identify microstructure, and to perform chemical and hardness analysis. Different sizes of preexisting cracks were detected on this longitudinal cut surface. While their sizes vary from a few hundred microns, to over 3 mm in depth in the tube wall, they are propagating at approximately 45\u25e6 with respect to the cross tube inner wall (Fig. 10). Microstructure of the cross tube at low magnification at additional cut surface can be seen at Fig" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_35_0000395_978-3-030-67750-3_16-Figure28.8-1.png", + "original_path": "designv11-35/openalex_figure/designv11_35_0000395_978-3-030-67750-3_16-Figure28.8-1.png", + "caption": "Fig. 28.8 Outlet 2\u2014velocity reading of Mode 2 nozzle", + "texts": [ + "9 presents the pressure curve that is affected by the length of the nozzle. It clearly can be seen that the pressure at outlet 2 is inversely proportional to the nozzle length. As the nozzle length increases, the pressure will be decreased. The graph pattern is in a gradually decreasing manner. This is because, as the powder goes through along the nozzle, the pressure will continue to decrease due to the loss of potential energy. Figure 28.7 shows the pressure reading during Mode 2 nozzle was set. While, the velocity reading of Mode 2 nozzle set as illustrated in Fig. 28.8. Presented in Fig. 28.10 is the velocity of the outlet 2 against the nozzle length. It shows that as the length of the nozzle increases the velocity happening on outlet two also increases. This is because the velocity is directly proportional to the length of the nozzle. Based on this graph, it is concluded that the higher the length and the diameter of the outlet, the faster the velocity. The Taguchi method was conducted using the Minitab software (Khavekar et al. 2017; Arnold 2006; Zaman et al" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_35_0001419_978-94-007-7401-8-Figure5.81-1.png", + "original_path": "designv11-35/openalex_figure/designv11_35_0001419_978-94-007-7401-8-Figure5.81-1.png", + "caption": "Fig. 5.81 3PaPaPR-1CPaPa-type fully-parallel PM with uncoupled Sch\u00f6nflies motions defined by MF = SF = 4, RF\u00f0 \u00de \u00bc \u00f0v1; v2; v3;xb\u00de, TF = 0, NF = 30 limb topology Pa\\Pa\\kP\\\\R, Pa\\Pa\\kP\\kR and C||Pa||Pa", + "texts": [ + "4d) C||Pa||Pa (Fig. 5.40) 16. 3PaPPaR1CPaPa (Fig. 5.79b) Pa\\P\\\\Pa\\kR (Fig. 5.4b) Idem No. 15 Pa\\P\\\\Pa||R (Fig. 5.4c) C||Pa||Pa (Fig. 5.40) 17. 3PaPPaR1CPaPat (Fig. 5.80a) Pa||P\\Pa\\kR (Fig. 5.4a) Idem No. 15 Pa||P\\Pa||R (Fig. 5.4d) C||Pa||Pat (Fig. 5.4p) 18. 3PaPPaR1CPaPat (Fig. 5.80b) Pa\\P\\\\Pa\\kR (Fig. 5.4b) Idem No. 15 Pa\\P\\\\Pa||R (Fig. 5.4c) C||Pa||Pat (Fig. 5.4p) (continued) 5.1 Fully-Parallel Topologies 535 Table 5.5 (continued) No. PM type Limb topology Connecting conditions 19. 3PaPaPR1CPaPa (Fig. 5.81) Pa\\Pa\\kP\\\\R (Fig. 5.4f) Idem No. 15 Pa\\Pa\\kP\\kR (Fig. 5.4h) C||Pa||Pa (Fig. 5.4o) 20. 3PaPaPR1CPaPa (Fig. 5.82) Pa\\Pa\\\\P\\\\R (Fig. 5.4g) Idem No. 15 Pa\\Pa\\\\P\\kR (Fig. 5.4e) C||Pa||Pa (Fig. 5.40) 21. 3PaPaPR1CPaPat (Fig. 5.83) Pa\\Pa\\kP\\\\R (Fig. 5.4f) Idem No. 15 Pa\\Pa\\kP\\kR (Fig. 5.4h) C||Pa||Pat (Fig. 5.4p) 22. 3PaPaPR1CPaPat (Fig. 5.84) Pa\\Pa\\\\P\\\\R (Fig. 5.4g) Idem No. 15 Pa\\Pa\\\\P\\kR (Fig. 5.4e) C||Pa||Pat (Fig. 5.4p) 23. 2PaPaRRR1PaPaRR1CPaPa (Fig. 5.85) Pa\\Pa||R||R\\\\R (Fig. 5.5a) The revolute joint between links 6D and 10 of G4 limb and the last revolute joints of limbs G1, G2 and G3 have superposed axes Pa\\Pa||R||R (Fig" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_35_0001730_s0025557200233147-Figure13-1.png", + "original_path": "designv11-35/openalex_figure/designv11_35_0001730_s0025557200233147-Figure13-1.png", + "caption": "FIG. 13.", + "texts": [], + "surrounding_texts": [ + "MECHANISMS I N T H R E E DIMENSIONS 17\nFurther, from IV,\nsin (ADB + DBA) _ cos a + cos jS sin DAB 1+oosDB\nSubstituting for cos DB, by I,\nsin# 1 + {cos a cos |3+sin a sin j8 cos (180-0)} d,9 s in cos a + cos /3 d\nHence - ddjd = sec \\x. - t an /M COS 0\n= . / ( l+ fc 2 ) -&cos0 ,\nwhere /x and fc have the values of equation (12). This method might, of course, be applied equally well to the spheric chain above, with =b +acos(7r - 0 ) .\nThus cos4> = {b -acos 0)/(a - b cos 0)\nand so, from relation (3)\nd J(a2-b*)_j(aZ-b*) d0 a-b cos 9 AE\nWhen a=b, a=/3 . Fig. 10 becomes symmetrical about AE. Projecting AGE on to AF,\n(a-a cos 9) cos + a sin 9sin c/> cos BEARINGS\nFOLLOWING SHAFT-\nForm I. FIG. 12. Form II .\nSpecial Cases. (1) Inversion of Hooke's Joint. Most conveni ently made as a combination of I and IT. Plane BOC is fixed horizontal.\nIf AOB has angular velocity w, c/> will fluctuate between iba, as shown previously, and hence we have a means of turning rotary into oscillatory motion.\n(2) Hooke's joint. For joining shafts not quite in alignment, but in the same plane. (Fig. 14.)\nWhen arranged as in Form I I , CDC\"' D' may take the form of a split circular plate, in which case an Oldham's coupling can be fitted between the two halves of it, if the shafts to be joined are not quite coplanar.\nhttps://doi.org/10.2307/3606015 11 Feb 2020 at 10:34:09, subject to the Cambridge Core terms of use, available at https://www.cambridge.org/core/terms. Downloaded from https://www.cambridge.org/core. Conrad Grebel University College, University of Waterloo, on", + "MECHANISMS IN T H R E E DIMENSIONS 19\nIn Form I, if there is sliding as well as turning at the hinges C, B, (\", B' there may be a slight difference of plane between the two shafts to be joined.\nD\n<0 C\nForm 1. Fid . 14.\n(a is small.'\nA model of Hooke's joint good enough for the drive to the propellor of a model boat, say, can be constructed from wire, preferably in Form I I .\nWe have already found for Hooke's joint tha t tan 8 = tan cj). cos a, where 6 = cot, cf>=a/t.\nDifferentiating, \u2014 = =\u20145 5-r \u2022 to 1 - sin^ct. cos\"1 a\nI t may be convenient to have this velocity ratio in the form of an expansion ; this is most quickly obtained as follows :\nco'jw = (1 - sin2a)s (1 - sin2 a cos2 d)-1\n= (I - \\ sin2 a - \\ sin4 a) (1 + sin2a cos2 6 + sin4 a cos4 6),\nneglecting powers of a above the fourth, since a is small,\n= 1 + sin2 a( - -1- + cos2 0) + s i n 4 \u00ab . { - \\ - \\ cos2 6 + cos46) = 1 + s i n 2 a ( | cos 29) + sin4 a ( \\ cos 26 + } cos id).\nThus\nw'jco = 1 +cos 2a>t(l sin2 a + I sin4 a + ...) + cos 4a)^(| sin4 a + . . . ) .\nI t might interest readers to consider for themselves the remark able motions which occur as a approaches 90\u00b0 in Hooke's joint and also the case in which a =90\u00b0, this being a particular case of those spheric chains in which a = y = (180 -jS) = (180 - 8 ) .\nThe Skew Jsogram. A mechanism for connecting two skew shafts, not coplanar, can be formed from the Skew Isogram ; the case for which the shafts are perpendicular is shown below.\nPlane ABB is fixed, a = 90, 0 = sin-1 {b/a).\nhttps://doi.org/10.2307/3606015 11 Feb 2020 at 10:34:09, subject to the Cambridge Core terms of use, available at https://www.cambridge.org/core/terms. Downloaded from https://www.cambridge.org/core. Conrad Grebel University College, University of Waterloo, on" + ] + }, + { + "image_filename": "designv11_35_0000008_j.autcon.2021.103573-Figure17-1.png", + "original_path": "designv11-35/openalex_figure/designv11_35_0000008_j.autcon.2021.103573-Figure17-1.png", + "caption": "Fig. 17. Configuration of the control of the block using gantry crane.", + "texts": [ + " Automation in Construction 124 (2021) 103573 F(X) = \u23a1 \u23a2 \u23a2 \u23a3 f1 f2 f3 f4 \u23a4 \u23a5 \u23a5 \u23a6 = \u23a1 \u23a2 \u23a2 \u23a2 \u23a2 \u23a3 pn+1 \u2212 pn \u2212 \u0394tvn+1 M1(vn+1 \u2212 vn) \u2212 \u0394t(Bun+1 \u2212 G1(pn+1\u03b3n+1)\u03bbn+1 ) M2 ( \u03b3\u0308n+1 \u2212 \u03b1s\u0307 \u2212 \u03b2s \u2212 \u03c7 \u222b sd ) \u2212 f2 + G2(pn+1\u03b3n+1)\u03bbn+1 \u03a6(pn+1\u03b3n+1) \u23a4 \u23a5 \u23a5 \u23a5 \u23a5 \u23a6 = 0 (38) The inverse dynamics solver was applied to two block erection simulations involving a gantry crane and a floating crane; this was done in order to develop and verify the automated crane control methods. With the control input obtained from Eq. (38), the motion of the actual mechanical system was carried out by solving the DELE. The modeling and simulations were performed using the in-house program developed by the authors in C# language. Fig. 17 shows the configuration of the gantry crane and the block and the desired trajectory that the block was required to follow. The crane H.-W. Lee et al. consisted of a controllable crane girder, two trolleys, and three wire ropes. Transportation and turnover of the block were conducted simultaneously using two trolleys and three wire ropes connected to the hooks. Three hooks and equalizers were connected to the trolley via wire ropes. The upper trolley connected two hooks that held one side of the block, and the lower trolley held the other side with one hook. The properties of the block are shown in Fig. 18. The weight of the block was assumed 300 tons, and the lifting capacity of the gantry crane was 900 tons. The initial and final position of the block is illustrated in Fig. 17. The initial position of the block was (0,0,20), and the final position was (25,10,10), rotated by 90\u25e6. The block was required to be simultaneously transported by 25 m in the x-direction, 10 m in the y-direction, and rotated by 90\u25e6. In feedforward control, the inverse dynamics solver calculates the control inputs required for the target to follow the desired trajectory, provided there are no external disturbances. For verification, feedforward control was conducted via the inverse dynamics solver for the gantry crane formulated in this study" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_35_0001419_978-94-007-7401-8-Figure3.40-1.png", + "original_path": "designv11-35/openalex_figure/designv11_35_0001419_978-94-007-7401-8-Figure3.40-1.png", + "caption": "Fig. 3.40 Overactuated PMs with coupled Sch\u00f6nflies motions of types 4PRRRP (a) and 4RPRRP (b) defined by MF \u00bc SF \u00bc 4; RF\u00f0 \u00de \u00bc v1; v2; v3;xd\u00f0 \u00de, TF = 0, NF = 2, limb topology P||R||R\\R\\P (a) and R||P||R||R\\P (b)", + "texts": [ + "3r) Idem No. 16 32. 4RRRPR (Fig. 3.36a) R||R\\R||P||R (Fig. 3.3s) Idem No. 16 33. 4RRRRP (Fig. 3.36b) R||R\\R||R||P (Fig. 3.3t) Idem No. 16 34. 4RRRRP (Fig. 3.37a) R||R\\R||R\\P (Fig. 3.3u) Idem No. 16 35. 4PRRRP (Fig. 3.37b) P\\R\\R||R\\P (Fig. 3.3v) Idem No. 29 36. 4RRPRP (Fig. 3.38a) R\\R||P||R\\P (Fig. 3.3w) Idem No. 16 37. 4RPRRP (Fig. 3.38b) R\\P||R||R\\P (Fig. 3.3x) Idem No. 16 38. 4RRPRP (Fig. 3.39a) R\\R||P||R\\P (Fig. 3.3y) Idem No. 16 39. 4RRRPP (Fig. 3.39b) R\\R||R||P\\P (Fig. 3.3z) Idem No. 16 40. 4PRRRP (Fig. 3.40a) P||R||R\\R\\P (Fig. 3.3z1) The fourth joints of the four limbs have parallel axes 41. 4RPRRP (Fig. 3.40b) R||P||R||R\\P (Fig. 3.3a) Idem No. 40 42. 4RPRRP (Fig. 3.41a) R||P||R\\R\\P (Fig. 3.3b0) Idem No. 40 43. 4RRPRP (Fig. 3.41b) R||R||P\\R\\P (Fig. 3.3c0) Idem No. 40 44. 4RRRPP (Fig. 3.42a) R||R\\R\\P\\\\P (Fig. 3.3d0) The third joints of the four limbs have parallel axes 45. 4PRRPR (Fig. 3.42b) P||R||R\\P\\\\R (Fig. 3.3e0) Idem No. 17 46. 4RPRPR (Fig. 3.43a) R||P||R\\P\\\\R (Fig. 3.3f0) Idem No. 17 47. 4RRPPR (Fig. 3.43b) R||R||P\\P\\\\R (Fig. 3.3g0) Idem No. 17 48. 4RPRPR (Fig. 3.44a) R||P||R\\P\\\\R (Fig. 3.3h0) Idem No" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_35_0000465_tpel.2021.3089611-Figure8-1.png", + "original_path": "designv11-35/openalex_figure/designv11_35_0000465_tpel.2021.3089611-Figure8-1.png", + "caption": "Fig. 8. Experimental setup.", + "texts": [ + " Assuming slow variation of rotor flux as opposed to the stator flux (as rotor time constant is much greater than the usual discrete sampling interval), thereby incorporating the approximations: 1\u02c6\u02c6 kk rr \u03c8\u03c8 and 1\u02c6\u02c6 kk ee , equations (4a), (4b), (4c) and (4d) modifies as 1\u02c6)(\u02c6 112 kkkjkk RTkTk errr ss P sr ss V\u03c8 i i (22a) where, kkkjkk RTkTk errr ss sr ss V\u03c8 i i \u02c6)(\u02c6 11 (22b) 1112 kRkTkk ss P ssss i-V\u03c8\u03c8 (22c) 2i . 2\u03c8( Im43)2( ss kkconjPkT PPP e (22d) m P swn P s P eeu InkkTTg 22\u02c6 * s\u03c8 . (22e) At 20 rad/s of drive speed, Fig. 7 shows the experimental results which depicts that the predicted torque is more oscillatory in case of one-step ahead prediction than that for the two-step ahead prediction. The experimental set-up [see Fig. 8] is built with a 2.2 kW squirrel-cage IM coupled with an eddy current breaker (ECB). The parameters of the IM is given in Table I. A three-phase VSI is used to feed the IM drive. Load torque controller provides required loading to IM through ECB. An incremental encoder is also used to measure the actual speed of the motor. The control algorithm is executed using dSPACE ds1103 R&D controller board with control desk and MATLAB Simulink software packages as interfacing media. The sampling frequency is kept at 20 kHz for the implementation of the scheme in digital domain" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_35_0000225_j.addma.2021.101955-Figure5-1.png", + "original_path": "designv11-35/openalex_figure/designv11_35_0000225_j.addma.2021.101955-Figure5-1.png", + "caption": "Fig. 5. Example sub model for obtaining distortion factors from slicing.", + "texts": [ + " A pyramid is split into four tetrahedrons, while a prism is split into two tetrahedrons and two pyramids. This splitting process may not be the most efficient since a pyramid can be split into two tetrahedrons, and a prism can be split into three tetrahedrons. But the current mode of splitting guarantees the continuity of FEM elements and ensures sharing the same surface between elements since the split lines are the same for neighboring elements. After slicing, the FEA sub-model is used for solving for the distortion factor, as shown in Fig. 5. In this figure, a unit force with the same magnitude and opposite direction is applied at the location of the node (such as node 4 in Fig. 5), which defines the cutting plane at the Z level. The centroid of the force couple coincides with this node (4). The distance between the two forces is the same as the hatch spacing along X and Y directions and the powder layer thickness along the Z direction. Because the sub-model includes the substrate, the corner nodes on the substrate are chosen as the constraint points, and the constraint type is fixed as a simple support, as shown in nodes 1 and 3 of Fig. 5. The displacements of nodes 1\u20135 can be obtained by solving the linear-elastic sub-model of Fig. 5. If the unit force couple is along the X-direction and the displacements of node 2 are d2x, d2y and d2z, then based on Eq. (1), the distortion factors of node 4 along the X-direction that contributes to node 2 is given by Eqs. (2)\u2013(4): w4x2x = d2x (2) w4x2y = d2y (3) w4x2z = d2z (4) The distortion factors along the Y and Z direction follow in the same manner. Following the solution for distortion factors, Backward Interpolation is performed to calculate the final distortion of all outer surface vertex nodes, as depicted in a 2D model in Fig" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_35_0000193_tmag.2021.3067323-Figure2-1.png", + "original_path": "designv11-35/openalex_figure/designv11_35_0000193_tmag.2021.3067323-Figure2-1.png", + "caption": "Fig. 2. Simple pole model. (a) Latitudinal direction (b) Longitudinal direction", + "texts": [ + " (1) where Tx, Ty, and Tz are output torques around x, y, and z-axis, respectively, pn is the position of each coil, Kmx(pn), Kmy(pn), and Kmz(pn) are magnet torque constants around x, y, and z-axis, respectively, Tcogx(pn), Tcogy(pn), and Tcogz(pn) are cogging torques around x, y and z-axis, respectively, i1, \u2026 and in are coil currents of each phase, Km is the magnet torque constant matrix, i is the coil current vector of each phase, and Tcog is the cogging torque vector. C. Torque Constant Map As shown in (1), the spherical actuator torque is obtained by summing up the torques generated by a simple magnetic pole model shown in Fig. 2, which consists of a pair of magnetic poles at each position. A magnetic field analysis using a 3D FEM is performed for each posture of the simple magnetic pole model, and the magnet torque constant matrix Km (pn) and the cogging torque vector Tcog (pn) are calculated by the following equations. (2) (3) where T is output torque, and 1 A and 0 A are coil currents. These torque constant matrices are maps that describe the torque constant around the x, y, and z-axis concerning present magnetic pole positions" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_35_0000802_s41315-021-00190-3-Figure2-1.png", + "original_path": "designv11-35/openalex_figure/designv11_35_0000802_s41315-021-00190-3-Figure2-1.png", + "caption": "Fig. 2 Problem of redundancy", + "texts": [ + "g. the posture of an object to be picked up. In this article, the reachability of a target point tp is guaranteed if the posture of the mobile robot Pend p assures the respect the following equality: The IKMMM is a function calculating the inverse kinematic model of a robot, called for finding the configuration qm of a manipulator, and for testing the reachability of a tp. The tp is reachable if Eq.\u00a0(1) has solution. The space S is a rectangular area defined in two dimensional cartesian space (Fig.\u00a02) by the cartesian coordinate vectors XS = [ Xmin S ,Xmax S ] and YS = [ Ymin S , Ymax S ] . The dimension of S depends on the dimensions of the manipulator links and the position of tp. (1)IKMMM ( Pend p , tp ) = qm = [qm1...qmn] \u2260 \ufffd 1 3 Ring of reachability is included in S. Positioning the base of the manipulator in this ring augments the probability to reach tp. Indeed, the target point tp is the objective to be reached with the EF. For this purpose, the base of the manipulator mounted on the mobile robot should be inside the ring. The redundancy appears at different levels: \u2013 Huge number of postures Pend p allows reaching the tp (Fig.\u00a02a). \u2013 Infinite number of paths and trajectories defined in pose space are generated and connect Pinit and Pend (Fig.\u00a02b). The mobile platform is adapted to take many paths. Every one generates an infinite number of trajectories P(t). The movement of the vehicle have an impact on the movement of the manipulator. Consequently, large number of movements of the manipulator joints are also generated. Non holonomic constraint is a relevant restriction typical for wheeled mobile robots. This restriction limits the movement of mobile robots, and allows their displacements only in certain directions (Eq 2). This kinematic constraint is analytically nonintegrable (Zhang et\u00a0al" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_35_0001603_301-Figure5-1.png", + "original_path": "designv11-35/openalex_figure/designv11_35_0001603_301-Figure5-1.png", + "caption": "Fig. 5. Schematic diagram of crossed cylinders wear machine", + "texts": [ + " Moreover, the stresses which can safely be imposed on sliding elements are very small in comparison with those which can be carried by rollers. C O N T I N U O U S R U B B I N G I N C O N D I T I O N S O F P O I N T C O N T A C T To maintain nominal point contact conditions indefinitely a fairly complicated apparatus is required. Dr. Archard and Mr. Tillen have designed a machine in which the test specimens are cylinders both rotating with their axes crossed at SO\u201d. Simultaneously, one cylinder reciprocates in a direction at 45\u201c to its axis (Fig. 5), the reciprocating motion being deliberately AA\u2019, upper specimen axis; BB\u2019, lower specimen axis; CC\u2019, fixed pivot axis, W, load. unrelated to the speed of rotation. Each point of the surface of each cylinder therefore rubs against the other, so that although the specimens wear they retain their cylindrical form. In this way point contact conditions are maintained and, since wear changes the diameter of the cylinders very little, the stresses are also effectively constant. The sliding distance can be varied by varying the speed of rotation of one cylinder with respect to the other" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_35_0000395_978-3-030-67750-3_16-Figure24.1-1.png", + "original_path": "designv11-35/openalex_figure/designv11_35_0000395_978-3-030-67750-3_16-Figure24.1-1.png", + "caption": "Fig. 24.1 3D DIC (stereo vision) cameras set up at a measured distance relative to the position of specimen", + "texts": [ + " Random speckling on specimen surface are done using matte black and white spray paints. A white matte layer serves as a contrasting background and random tiny black speckles pattern created using black spray paint is uniformly distributed on the entire specimen. The punch test rig consists of die sets to secure the 90 mm circular specimen blanks. The die sets have clearance in the centre to allow specimens to deform when loaded with a 20 mm diameter conical punch. The punch test rig is set up together with DIC cameras on a 100 kN Electric Mayes machine as shown in Fig. 24.1. The general method of assembling the whole rig starts with the clamping of the blanks. The blank is secured in between the dies and tightened before being put on the punch. The die set is supported with a pair of stands from the bottom for a little clearance between the punch tip and the specimen to avoid the punch from damaging the specimen before the start of the experiment. Supporting bars are then placed on top of the die set to provide a hold down force during the test and is tightly bolted at the connection which secures the position of both top supporting bar and bottom supporting bar of the Mayes machine to secure the die set", + " The 3D DIC set up is done by mounting cameras on the boom at the top of the test rig and connecting both cameras to the computer to view the area of interest on the specimen. On the computer, the software being used for the pattern recognition is based on the principles of digital image correlation is the VIC-3D by Correlated Solution. Every pair of images continuously captured by both cameras is processed into stereo images. Cameras orientation is adjusted until the area of interest is at the centre of camera views, then the best focal length is found by adjusting the lenses. Figure 24.1 shows the measurements of cameras set up relative to the specimen. Cameras are adjusted until they are in focus at the same location on the specimen preferably at the centre region. Examples of speckle images with a good focus are shown in Fig. 24.2. Once both cameras are in focus, the calibration process is carried out by using a pre-determined dot-pattern calibration plate. At least 25 images are captured for the calibration process at which later only 15 images with good calibration score are chosen to make sure the calibration error between the two cameras is below 0" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_35_0000529_012088-Figure6-1.png", + "original_path": "designv11-35/openalex_figure/designv11_35_0000529_012088-Figure6-1.png", + "caption": "Figure 6. Initial and Final Position of the Robot in XYZ plane before performing DH frame Assignments", + "texts": [ + " After getting end points (Px, Py and Pz) using forward kinematics, inverse kinematics helps in finding 1 to 5 respectively. According to inverse kinematics F is the function of x,y,z and R. ICASSCT 2021 Journal of Physics: Conference Series 1921 (2021) 012088 IOP Publishing doi:10.1088/1742-6596/1921/1/012088 All the joint angles are found based on the known transformation matrix shown in (3). All the unknown joint angles are found using (16) to (29). These joint angles are shown in Table 3 (Both Initial and Final angles). Figure 6 and 7 shows the initial and final position of the robot as per forward kinematics, which is simulated in MATALAB. After applying inverse kinematic technique, the final ICASSCT 2021 Journal of Physics: Conference Series 1921 (2021) 012088 IOP Publishing doi:10.1088/1742-6596/1921/1/012088 As discussed earlier, robot trajectory planning is to ensure the smooth variation in the robotic joints. Trajectory planning also gives time history of position, velocity and acceleration at the intermediate point as well as final and starting point [25]" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_35_0000686_j.mechmachtheory.2021.104433-Figure25-1.png", + "original_path": "designv11-35/openalex_figure/designv11_35_0000686_j.mechmachtheory.2021.104433-Figure25-1.png", + "caption": "Fig. 25. The equivalent metamorphic clamping mechanism.", + "texts": [ + " Mechanism and Machine Theory 166 (2021) 104433 (1) Rotation motion output to provide the motion of claw to closer or away from the target object. (2) Translation motion output to provide the motion of claw to clamp or release target object. The motion modes of the clamping mechanism include rotation and translation. Therefore, we choose the equivalent mechanism of the epicyclic bevel gear mechanism to replace the rotation link of the clamping mechanism in Fig. 24. An equivalent metamorphic clamping mechanism is constructed in Fig. 25. The motion-screw system Tm of the equivalent metamorphic clamping mechanism is obtained as Eq. (52). H. Yuan et al. Mechanism and Machine Theory 166 (2021) 104433 Tm = \u23a7 \u23aa \u23aa \u23aa \u23aa \u23aa \u23aa \u23aa \u23aa \u23aa \u23aa \u23aa \u23aa \u23aa \u23aa \u23aa \u23aa \u23aa \u23aa \u23aa \u23aa \u23aa \u23aa \u23aa \u23a8 \u23aa \u23aa \u23aa \u23aa \u23aa \u23aa \u23aa \u23aa \u23aa \u23aa \u23aa \u23aa \u23aa \u23aa \u23aa \u23aa \u23aa \u23aa \u23aa \u23aa \u23aa \u23aa \u23aa \u23a9 TA = [ 0 0 \u03c9A 0 0 0 ]T TB = [ 0 0 \u03c9B 0 0 0 ]T TC1 = [ \u2212 sin\u03b1b3\u03c9C1 \u2212 cos\u03b1b\u03c9C1 \u2212 sin\u03b1b1\u03c9C1 0 Rbsin\u03b1b1\u03c9C1 \u2212 Rbcos\u03b1b\u03c9C1 ] T TC2 = [ cos\u03b1b3\u03c9C2 \u2212 cos\u03b1btan\u03b1b3\u03c9C2 \u2212 sin\u03b1b1tan\u03b1b3\u03c9C2 0 0 0 ]T TC3 = [ 0 \u2212 sin\u03b1b1\u03c9C3 cos\u03b1b\u03c9C3 q11\u03c9C3 q12\u03c9C3 q13\u03c9C3 ] T cos\u03b1b3 TD1 = [ \u2212 sin\u03b1b1sin\u03b1b3\u03c9D1 \u2212 cos\u03b1bsin\u03b1b1\u03c9D1 cos2\u03b1b1\u03c9D1 q14\u03c9D1 q15\u03c9D1 q16\u03c9D1 ]T cos\u03b1b1 TD2 = [ cos\u03b1b\u03c9D2 \u2212 sin\u03b1b3\u03c9D2 0 q17\u03c9D2 q18\u03c9D2 q19\u03c9D2 ] T cos\u03b1b1 TE = [\u03c9E 0 0 0 rb\u03c9E 0 ]T TF1 = [ 0 0 0 vF 0 0 ]T TF2 = [\u03c9F2 0 0 0 rb\u03c9F2 0 ]T TG = [ 0 0 0 vG 0 0 ] (52) where the q11 to q19 are given in Eq" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_35_0001419_978-94-007-7401-8-Figure6.37-1.png", + "original_path": "designv11-35/openalex_figure/designv11_35_0001419_978-94-007-7401-8-Figure6.37-1.png", + "caption": "Fig. 6.37 2PPaRRR-1PPaRR-1RPPaPat-type maximaly regular fully-parallel PM with Sch\u00f6nflies motions defined by MF = SF = 4, (RF) = (v1; v2; v3;xb), TF = 0, NF = 19, limb topology P||Pa||R||R\\R, P||Pa||R||R and R||P||Pa||Pat", + "texts": [ + " 3PPPaR-1RPPaPat (Fig. 6.33b) P\\P\\||Pa\\\\R (Fig. 4.8a) P\\P\\||Pa||R, (Fig. 4.8c) R||P||Pa||Pat (Fig. 5.4n) Idem no. 9 13. 3PPaPR-1RPPaPa (Fig. 6.34) P||Pa\\P\\R (Fig. 4.8d) P||Pa\\P\\||R (Fig. 4.8e) R||P||Pa||Pa (Fig. 5.4m) Idem no. 9 14. 3PPaPR-1RPPatPa (Fig. 6.35) P||Pa\\P\\\\R (Fig. 4.8d) P||Pa\\P\\||R (Fig. 4.8e) R||P||Pa||Pat (Fig. 5.4n) Idem no. 9 15. 2PPaRRR-1PPaRR-1RPPaPa (Fig. 6.36) P||Pa||R||R\\R (Fig. 4.9a) P||Pa||R||R (Fig. 4.8g) R||P||Pa||Pa (Fig. 5.4m) Idem no. 9 16. 2PPaRRR-1PPaRR-1RPPaPat (Fig. 6.37) P||Pa||R||R\\R (Fig. 4.9a) P||Pa||R||R (Fig. 4.8g) R||P||Pa||Pat (Fig. 5.4n) Idem no. 9 17. 3PPaPaR-1RPaPaP (Fig. 6.38) P||Pa||Pa\\R (Fig. 4.10a) P||Pa||Pa||R (Fig. 4.10b) R||Pa||Pa||P (Fig. 5.4k) Idem no. 1 18. 3PPaPaR-1RPaPatP (Fig. 6.39) P||Pa||Pa\\R (Fig. 4.10a) P||Pa||Pa||R (Fig. 4.10b) R||Pa||Pat||P (Fig. 5.4l) Idem no. 1 19. 3PPaPaR-1RPPaPa (Fig. 6.40) P||Pa||Pa\\R (Fig. 4.10a) P||Pa||Pa||R (Fig. 4.10b) R||P||Pa||Pa (Fig. 5.4m) Idem no. 9 20. 3PPaPaR-1RPaPatP (Fig. 6.41) P||Pa||Pa\\R (Fig. 4" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_35_0001419_978-94-007-7401-8-Figure5.94-1.png", + "original_path": "designv11-35/openalex_figure/designv11_35_0001419_978-94-007-7401-8-Figure5.94-1.png", + "caption": "Fig. 5.94 3PPPR-1PPPPR type redundantly actuated PMs with uncoupled Sch\u00f6nflies motions defined by MF = 5, SF = 4, RF\u00f0 \u00de \u00bc \u00f0v1; v2; v3;xa\u00de, TF = 1, NF = 6, limb topology P\\P\\\\P\\\\R, P\\P\\\\P\\kR, P||P\\P\\\\P\\kR (a) and P\\P\\\\P\\\\R, P\\P\\\\P||R, P||P\\P\\\\P||R (b)", + "texts": [ + "2 Redundantly Actuated Topologies 545 546 5 Topologies with Uncoupled Sch\u00f6nflies Motions 5.2 Redundantly Actuated Topologies 547 548 5 Topologies with Uncoupled Sch\u00f6nflies Motions 5.2 Redundantly Actuated Topologies 549 550 5 Topologies with Uncoupled Sch\u00f6nflies Motions 5.2 Redundantly Actuated Topologies 551 Table 5.13 Limb topology and connecting conditions of the fully-parallel solutions with no idle mobilities presented in Figs. 5.94, 5.95, 5.96, 5.97, 5.98 No. PM type Limb topology Connectingconditions 1. 3PPPR1PPPPR (Fig. 5.94a) P\\P\\\\P\\\\R (Fig. 4.1b) The last revolute joints of the four limbs have parallel axes. The actuated prismatic joints of limbs G1, G2 and G3 hvave orthogonal directions. The actuated prismatic joints of limbs G3 and G4 have parallel directions P\\P\\\\P\\kR (Fig. 4.1a) P||P\\P\\\\P\\kR (Fig. 5.93a) 2. 3PPPR1PPPPR (Fig. 5.94b) P\\P\\\\P\\\\R (Fig. 4.1b) Item No. 1 P\\P\\\\P||R (Fig. 4.1c) P||P\\P\\\\P||R (Fig. 5.93b) 3. 1PPRR2PPRRR1PPPRR (Fig. 5.95a) P\\P\\kR||R (Fig. 4.1d) Idem No. 1 P\\P\\kR||R\\R (Fig. 4.2a) P||P\\P\\kR||R\\R (Fig. 5.93d) 4. 1PRPR2PRPRR1PPRPRR (Fig. 5.95b) P||R\\P\\kR (Fig. 4.1e) Idem No. 1 P||R\\P\\kR\\R (Fig. 4.2b) P||P||R\\P\\kR\\R (Fig. 5.93e) 5. 1PRRP2PRRPR1PPRRPR (Fig. 5.96a) P||R||R\\P (Fig. 4.1f) Idem No. 1 P||R||R\\P\\\\R (Fig. 4.2c) P||P||R||R\\P\\\\R (Fig. 5.93c) 6. 1PRRR2PRRRR1PPRRRR (Fig. 5.96b) P||R||R||R (Fig. 4.1 g) Idem No", + " 1CPR-2CPRR1PCPRR (Fig. 5.97b) C\\P\\kR (Fig. 4.1i) Idem No. 1 C\\P\\kR\\R (Fig. 4.2e) P|| C\\P\\kR\\R (Fig. 5.93i) 9. 1CRP-2CRPR1PCRPR (Fig. 5.98a) C||R\\P (Fig. 4.1j) Idem No. 1 C||R\\P\\\\R (Fig. 4.2f) P||C||R\\P\\\\R (Fig. 5.93h) 10. 1CRR-2CRRR1PCRRR (Fig. 5.98b) C||R||R (Fig. 4.1 k) Idem No. 1 C||R||R\\R (Fig. 4.2 g) P||C||R||R\\R (Fig. 5.93j) 552 5 Topologies with Uncoupled Sch\u00f6nflies Motions Table 5.14 Structural parametersa of parallel mechanisms in Figs. 5.94, 5.95, 5.96 No. Structural parameter Solution Figure 5.94 Figures 5.95 and 5.96 1. m 15 18 2. p1 4 4 3. pi (i = 2, 3) 4 5 4. p4 5 6 5. p 17 20 6. q 3 3 7. k1 4 4 8. k2 0 0 9. k 4 4 10. (RG1) (v1; v2; v3;xa) (v1; v2; v3;xa) 11. (RG2) (v1; v2; v3;xa) (v1; v2; v3;xa;xb) 12. (RG3) (v1; v2; v3;xa) (v1; v2; v3;xa;xd) 13. (RG4) (v1; v2; v3;xa) (v1; v2; v3;xa;xd) 14. SG1 4 4 15. SGi (i = 2, 3) 4 5 16. SG4 4 5 17. rGi (i = 1,\u2026,4) 0 0 18. MG1 4 4 19. MGi (i = 2, 3) 4 5 20. MG4 5 6 21. (RF) (v1; v2; v3;xa) (v1; v2; v3;xa) 22. SF 4 4 23. rl 0 0 24. rF 12 15 25. MF 5 5 26" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_35_0000371_tmag.2021.3076134-Figure4-1.png", + "original_path": "designv11-35/openalex_figure/designv11_35_0000371_tmag.2021.3076134-Figure4-1.png", + "caption": "Fig. 4. 2-D simplification model of the proposed ST-PMTFLG. (a) Equivalent schematic. (b) 2-D simplification model.", + "texts": [ + " When the spiral translator moves along the z-axis, it can be considered that it is rotating in the xy plane, like a partitioned stator flux-reversal permanent magnet (FRPM) machine [17]. The linear velocity along the z-axis can be converted to the equivalent angular speed of rotation in the xy plane, as in \u03b8twist \u03c9 = lt v (3) where \u03b8twist is the translator twist arc, lt is the translator length, \u03c9 is the equivalent angular speed of rotation in the xy plane, and v is the moving speed of translator along the z-axis. There is a factor needed to consider when using the above 2-D simplification model, as shown in Fig. 4(b). In a partitioned stator FRPM [17], the moving rotor is approximately a rectangle when viewed from the xz plane. However, in the proposed ST-PMTFLG, the moving translator is approximately a parallelogram when viewed from the xz plane. So the equivalent factor k is introduced and defined as k = SST\u2212PMTFLG SFRPM = SBEFC SAEFC = SAEFC \u2212 SABE SAEFC (4) where SST\u2212PMTFLG is the area of the translator curved surface when main flux flow through it in proposed ST-PMTFLG. SFRPM is the area of the rotor curved surface when main flux flows through it in partitioned stator FRPM SAEFC = \u03c0(Rti + Rto)\u03b8st l pm 360 (5) where Rti is the translator inner radius, Rto is the translator outer radius, l pm is the PM length, and \u03b8st is the steel block arc SABE \u2248 l2 pm 2 tan(\u03b8twist) (6) where \u03b8twist is the translator twist arc" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_35_0000360_j.matpr.2021.04.147-Figure2-1.png", + "original_path": "designv11-35/openalex_figure/designv11_35_0000360_j.matpr.2021.04.147-Figure2-1.png", + "caption": "Fig. 2. Equivalent (von mises) stress for Front Impact Test.", + "texts": [], + "surrounding_texts": [ + "Off road vehicle is shaped to race and steer on different terrains. ORV is designed in such a way that it can endure off-roading terrains. In off-terrain circumstances, the vehicle bears dynamic loads and all that is sustained through the chassis frame. Chassis frame bears every mountings and assembly, so it is expected from an ORV chassis frame to sustain both static and dynamic loads. The selection of materials for chassis greatly depends on the high tensile strength and material light weight. The majority of manufacturers favour lightweight, cost-effective, safe, and recyclable materials." + ] + }, + { + "image_filename": "designv11_35_0001762_pime_proc_1924_106_008_02-Figure7-1.png", + "original_path": "designv11-35/openalex_figure/designv11_35_0001762_pime_proc_1924_106_008_02-Figure7-1.png", + "caption": "FIG. 7.", + "texts": [ + " In passing, it might be remarked that the same method would furnish a perfectly rigorous solution of the somewhat similar problem of the anchor ring and chain link. Equations Resulting from Conditions of Equilibrium of a Segment of the Rim.-Let 20 be the angle between two consecutive arms, w the weight of unit volume of the material, w the angular velocity, A the radial section of the rim, and R the radius of the unstrained central line of the rim-not the neutral axis as i t was termed by Professor Pippard (page 2 7 ) . Let N and M be the direct stress (tension), and bending moment respectively a t a section X, Fig. 7, defined by the radial angle /3 measured from the section midway between two arms ; N,, and Mu at WEST VIRGINA UNIV on June 5, 2016pme.sagepub.comDownloaded from JAN. 1924. PLY-WHEEL STRESSES. 46 are the tension and bending moment a t the midway section, where, from considerations of symmetry, it was clear that the shearing force must vanish. First consider the equilibrium of the segment of the rim between the two knidway sections a t A and B, and let T be the tension in the arm a t its junction with the rim" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_35_0000395_978-3-030-67750-3_16-Figure30.6-1.png", + "original_path": "designv11-35/openalex_figure/designv11_35_0000395_978-3-030-67750-3_16-Figure30.6-1.png", + "caption": "Fig. 30.6 a Final conceptual design of wheelbarrow loader bucket. b Exploded view of the design", + "texts": [ + " Based on all the design steps taken before, three possible design sketches have been made to slightly give a picture to develop a possible design for this product. The simple 2D sketches equipped with the specifications and parts from the data collected in the previous design method process had been analyzed later by using a weighted objective analysis method. Therefore, the final design concept is available after finishing up all the designs steps. The final conceptual design was made by using SolidWorks, Fig. 30.6a displays the final conceptual design of the wheelbarrow loader bucket, while Fig. 30.6b is the exploded view of the design. The side view of wheelbarrow bucket loader is shown in Fig. 30.7. The design is based on the existing wheelbarrow but has a modification by using the six-point linkage based on the existing loader bucket design. The pneumatic actuator attached at the loader linkages is to make ease the loading and unloading process of the bucket. Themainmaterial used for the loader bucket is using Aluminiumwhile the wheelbarrow body is maintains to use the existing material, which is steel" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_35_0000686_j.mechmachtheory.2021.104433-Figure5-1.png", + "original_path": "designv11-35/openalex_figure/designv11_35_0000686_j.mechmachtheory.2021.104433-Figure5-1.png", + "caption": "Fig. 5. Epicyclic bevel gear train.", + "texts": [ + " The geometrical relationships between the length of the links leAC, leOB, leOD, leCD, leDB in the equivalent mechanism and the radiuses Re, re of the pitch circles in the epicyclic external gear train are given in Eq. (7), where \u03b1e represents the pressure angle. \u23a7 \u23aa \u23aa \u23aa \u23aa \u23aa \u23aa \u23aa \u23aa \u23aa \u23aa \u23a8 \u23aa \u23aa \u23aa \u23aa \u23aa \u23aa \u23aa \u23aa \u23aa \u23aa \u23a9 leAC = re + Re leOB = Recos\u03b1e \u03b1e1 = arctan ( leACsin\u03b1e leOB ) \u2212 \u03b1e leOD = leOB cos(\u03b1 + \u03b1e1) leCD = recos\u03b1e leDB = (re + Re)sin\u03b1e (7) Similarly, for an epicyclic bevel gear train, a base coordinate system, sun bevel gear coordinate system, bevel gear arm coordinate system, and planet bevel gear coordinate system are built, as shown in Fig. 5. Similar to the epicyclic external gear train, the transmission relationship of the epicyclic bevel gear train can be obtained as Eq. (8). \u03c9b1Rb = \u03c9b2Rb + \u03c9b3rb (8) where \u03c9b1, \u03c9b2 represent the scalar angular velocities of the sun bevel gear, the bevel gear arm rotating around the zb0-axis H. Yuan et al. Mechanism and Machine Theory 166 (2021) 104433 respectively. \u03c9b3 represents the scalar angular velocity of the planet bevel gear rotating around the zb3-axis. Rb, rb refer to the pitch radius of the sun bevel gear and the planet bevel gear, respectively. The screw system of the epicyclic bevel gear train, which is expressed in Fig. 5(b), is written as Eq. (9). Sb = \u23a7 \u23a8 \u23a9 Sb1 = [ 0 0 1 0 0 0 ]T Sb2 = [ 0 0 1 0 0 0 ]T Sb3 = [ cos\u03c6b2 sin\u03c6b2 0 0 0 0 ]T (9) Then, the motion-screw system is derived as \u23a7 \u23a8 \u23a9 Tb1 = [ 0 0 \u03c9b1 0 0 0 ]T Tb2 = [ 0 0 \u03c9b2 0 0 0 ]T Tb3 = [\u03c9b3cos\u03c6b2 \u03c9b3sin\u03c6b2 \u03c9b2 0 0 0 ]T (10) Obviously, dim(Sb)=2, and the mobility of the epicyclic bevel gear train is 2. The instantaneous screw axis Tib13 can be calculated as Eq. (11). Tib13 = Tb3 \u2212 Tb1 = [\u03c9b3cos\u03c6b2 \u03c9b3sin\u03c6b2 \u03c9b2 \u2212 \u03c9b1 0 0 0 ]T (11) The position vector rib13 and the direction vector sib13 can be obtained as Eq" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_35_0001419_978-94-007-7401-8-Figure6.26-1.png", + "original_path": "designv11-35/openalex_figure/designv11_35_0001419_978-94-007-7401-8-Figure6.26-1.png", + "caption": "Fig. 6.26 3PPaPaR-1RUPU-type maximaly regular fully-parallel PM with Sch\u00f6nflies motions defined by MF = SF = 4, (RF) = (v1; v2; v3;xb), TF = 0, NF = 22, limb topology P||Pa||Pa\\R, P||Pa||Pa||R and R\\R\\R\\P\\||R\\R", + "texts": [ + " 3PPPaR-1RUPU (Fig. 6.23b) P ?P ?||Pa\\\\R (Fig. 4.8a) P ?P ?||Pa||R, (Fig. 4.8c) R ?R ?R ?P ?||R ?R (Fig. 5.1b) Idem no. 13 15. 3PPaPR-1RUPU (Fig. 6.24a) P||Pa ?P\\\\R (Fig. 4.8d) P||Pa ?P ?||R (Fig. 4.8e) R ?R ?R ?P ?||R ?R (Fig. 5.1b) Idem no. 13 16. 2PPaRRR-1PPaRR-1RUPU (Fig. 6.24b) P||Pa||R||R ?R (Fig. 4.9a) P||Pa||R||R (Fig. 4.8g) R ?R ?R ?P ?||R ?R (Fig. 5.1b) Idem no. 13 17. 3PPaPaR-1RPPP (Fig. 6.25) P||Pa||Pa ?R (Fig. 4.10a) P||Pa||Pa||R (Fig. 4.10b) R ?P\\\\P\\\\P (Fig. 5.1a) Idem no. 9 18. 3PPaPaR-1RUPU (Fig. 6.26) P||Pa||Pa ?R (Fig. 4.10a) P||Pa||Pa||R (Fig. 4.10b) R ?R ?R ?P ?||R ?R (Fig. 5.1b) Idem no. 13 Table 6.7 Structural parametersa of parallel mechanisms in Figs. 6.7, 6.8, 6.9, 6.10, 6.11, 6.12, 6.13, 6.14 No. Structural parameter Solution Figures 6.7 and 6.11 Figures 6.8, 6.9, 6.10, 6.12, 6.13, 6.14 1. m 18 20 2. pi (i = 1, 3) 4 5 3. p2 4 4 4. p4 10 10 5. p 22 24 6. q 5 5 7. k1 3 3 8. k2 1 1 9. k 4 4 10. (RG1) (v1; v2; v3;xb) (v1; v2; v3;xa;xb) 11. (RG2) (v1; v2; v3;xb) (v1; v2; v3;xb) 12. (RG3) (v1; v2; v3;xb) (v1; v2; v3;xb;xd) 13", + "4o and p to replace a group of revolute and prismatic pairs with coincident axis/direction. Fully-parallel topologies with at least two identical limbs are illustrated in Figs. 6.27, 6.28, 6.29, 6.30, 6.31, 6.32, 6.33, 6.34, 6.35, 6.36, 6.37, 6.38, 6.39, 6.40, 6.41, 6.42, 6.43, 6.44, 6.45, 6.46. The limb:topology and connecting conditions of these solutions are systematized in Tables 6.13 and 6.14, as are their structural parameters in Tables 6.15 and 6.16. Table 6.12 (continued) No. Structural parameter Solution Figure 6.25 Figure 6.26 Pp1 j\u00bc1 fj 10 10 30. Pp2 j\u00bc1 fj 10 10 31. Pp3 j\u00bc1 fj 10 10 32. Pp4 j\u00bc1 fj 4 6 33. Pp j\u00bc1 fj 34 36 a See footnote of Table 2.2 for the nomenclature of structural parameters 6.3 Fully-Parallel Topologies with Complex Limbs 621 622 6 Maximally Regular Topologies with Sch\u00f6nflies Motions 6.3 Fully-Parallel Topologies with Complex Limbs 623 624 6 Maximally Regular Topologies with Sch\u00f6nflies Motions 6.3 Fully-Parallel Topologies with Complex Limbs 625 626 6 Maximally Regular Topologies with Sch\u00f6nflies Motions 6" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_35_0000164_tie.2021.3060653-Figure1-1.png", + "original_path": "designv11-35/openalex_figure/designv11_35_0000164_tie.2021.3060653-Figure1-1.png", + "caption": "Fig. 1. Topology of the designed 2DoFDDIM: (a) structure and (b) distribution of the flux density.", + "texts": [], + "surrounding_texts": [ + "0278-0046 (c) 2021 IEEE. Personal use is permitted, but republication/redistribution requires IEEE permission. See http://www.ieee.org/publications_standards/publications/rights/index.html for more information.\nAbstract\u2014A 2-degree-of-freedom (2DoF) induction machine realizes rotary, linear and helical motions, exhibiting the merits of integrated structure and high material utilization. Compared with traditional electric machines, one of the special factors affecting the electromagnetic performances is the helical motion coupling effect (HMCE). This paper systematically investigates its production mechanism based on the proposed HMCE model firstly. Then a comprehensive comparison of the torque performances between the rotary and linear motion parts is conducted to evaluate the HMCE in two motion parts, and then the results are verified by 3- dimension finite element model. Besides, resistant torque and force ratios are proposed and calculated to assess the extent of the HMCE. It can be concluded that the torque performances of the rotary motion component are weakened more than those of the linear motion component by the HMCE. Further, two methods to suppress the coupling effect, namely adjustment of linear motion slip ratio and change of linear motion frequency, are discussed. Experimental results have validated the analytical results and effectiveness of the suppression strategies.\nIndex Terms\u20142-degree-of-freedom induction motion, helical motion coupling effect, rotary motion, linear motion, resistant torque ratio, resistant force ratio.\nI. INTRODUCTION\nOR INDUSTRY APPLIATIONS, including pick-and-place\nrobots [1], compressors [2], wave energy harvesting [3] and\nManuscript received June 04, 2020; revised July 18, 2020, October 03, 2020, and December 02, 2020; accepted February 05, 2021. This work is supported by National Natural Science Foundation of China under grant 51777060 and 51277054, (Corresponding author: Jikai Si).\nLujia Xie is with the Department of Electrical Engineering and Electronics, University of Liverpool, Liverpool L693BX, U.K. and Department of Electrical Engineering, National Tsing Hua University, Hsinchu 30013, Taiwan (e-mail: xielujia95@gmail.com).\nJikai Si and Shuai Xu are with the Department of Electrical Engineering, Zhengzhou University, Zhengzhou 450006, China (e-mail: sijikai527@126.com).\nTsai-fu Wu is with the Department of Electrical Engineering, National Tsing Hua University, Hsinchu 30013, Taiwan (e-mail: tfwu@ee.nthu.edu.tw)\nJiafeng Zhou is with the Department of Electrical Engineering and Electronics, University of Liverpool, Liverpool, L69 3BX, UK (e-mail: zhouj@liverpool.ac.uk).\nYihua Hu is with the Department of Electronic Engineering, University of York, York YO10 5DD, U.K. (E-mail: yihua.hu@york.ac.uk).\nso on, some special characteristics are needed for the\ncomplicated multi-degree-freedom motions of traction systems,\nsuch as high integration, high reliability and quick response [4-\n5]. Recently, a new type of 2-degree-of-freedom (2DoF)\nmachine was proposed, which can realize 2D rotary motion, 2D\nlinear motion or 3D helical motion by employing only one\nmotor [6]-[11]. The 2DoF machine is much more integrated and\nroom-saving compared with traditional rotary-linear motion\nsystems due to the absence of intermediate mechanical devices,\nwhich can be considered as a promising candidate of traction\nmachines in intelligent industrial systems [12].\nHowever, owing to the high integration of multiple motions\nof the 2DoF machines, there is a helical motion coupling effect\n(HMCE) in the helical motion [13]. It significantly affects the\nelectromagnetic performances, which is one of the key factors\nhindering industrial applications of the machines. Ebrahim\nAmiri combined the transient time-domain finite element model\nwith frequency domain slip frequency technique to model the\ncoupling effect caused by dynamic end effects in the rotary\narmature of the rotary-linear induction motor [14]. Reference\n[15] used the design of short magnetic paths to avoid the\ncoupled effect in both linear motion and rotary movement for a\nrotary-linear switched-reluctance motor. Besides, a control\nalgorithm based on inductance profiles of the motor phases was\nproposed to decouple the rotary and linear movements of\nanother type of rotary-linear switched reluctance motor [16].\nFor the double-stator rotary-linear permanent magnet motor\n[17], the orthogonal crossed magnetic field caused by the\ncoupling effect was investigated to study the degree of the\ncoupling effect. Although some researches have been achieved\n[13], systematic and profound approaches to deal with the\ncoupling effect in 2DoF machines are yet not available.\nMoreover, the analysis of the coupling effect is always based\non 3-dimension (3D) finite element analysis (FEA), which is\naccurate but time-consuming.\nTherefore, this paper is to systematically investigate the\ninfluence of HMCE on electromagnetic performance of a 2DoF\ndirect drive induction motor (2DoFDDIM), and ultimately\npropose HMCE suppression strategies. For the 2DoFDDIM\n[18], characteristic analyses of the rotary, linear and helical\nmotions have been presented focusing on super-\nsynchronization effect, no-load and load characteristics, and\noverall difference between three motions, while the production\nLujia Xie, Jikai Si, Member, IEEE, Tsai-fu Wu, Senior Member, IEEE, Yihua Hu, Senior\nMember, IEEE, Jiafeng Zhou, and Shuai Xu\nAnalysis and Suppression Techniques of Helical Motion Coupling Effect for the 2DoF\nDirect Drive Induction Machine\nF\nAuthorized licensed use limited to: Univ of Calif Santa Barbara. Downloaded on May 15,2021 at 19:03:15 UTC from IEEE Xplore. Restrictions apply.", + "0278-0046 (c) 2021 IEEE. Personal use is permitted, but republication/redistribution requires IEEE permission. See http://www.ieee.org/publications_standards/publications/rights/index.html for more information.\nmechanism of the HMCE has not been investigated yet. Besides,\nit is worth exploring and comparing the different degrees of the\nHMCE on each motion, which is useful in developing the\ncoupling effect suppressing schemes. Thus, its production\nmechanism is investigated in section II, and the resistant torque\nand force are calculated based on the HMCE mathematical\nmodel presented in section III. In section \u2163, the influences of\nthe HMCE on two motion parts are evaluated and compared.\nMeanwhile, the key parameters relating to the coupling effect\nare also analyzed. Then, two coupling effect suppression\nstrategies are discussed in Section V, and the 2DoFDDIM\nmachine prototype is implemented and tested to verify the\neffectiveness of the proposed techniques. Finally, section \u2165\ndraws the conclusions.\nII. HELICAL MOTION COUPLING EFFECT\nThe topology of the linear motion stator is evolved from the\nrotary motion stator as shown in Fig. 2, where their outer stator\ndiameter Dso and inner stator diameter Dsi of the two parts are respectively identical. The mechanical pole pitch of the linear\nmotion stator, \u03c4pl, is the same as that of the rotary motion stator, \u03c4pr, which can be calculated as:\n_ 1\n2 4\na r\npr pl si\np p\nl D\nN N (1)\nwhere la_r is the stack length of the rotary motion stator and Np is the number of pole pairs.\nelectromagnetic pole pitch \u03c4epr is longer than that of the mechanical size \u03c4pr, as shown in Fig. 3. The air-gap flux density distributions of rotary motion components at rated sr = 0.2 are derived using FEA. It can be seen that the flux density decays\nAverage no-load\nrotating speed\nAverage locked\ntorque\nUnpowered linear motion stator 792.38 r/min 12 Nm\nLinear velocity 3.9 m/s 651.18 r/min 10.94 Nm\nLinear velocity 3.12 m/s 661.99 r/min 11 Nm Linear velocity 2.34 m/s 669.51 r/min 11.12 Nm Linear velocity 1.56 m/s 678.99 r/min 11.36 Nm Linear velocity 0.78 m/s 682.58 r/min 11.41 Nm\nLinear velocity 0 m/s 689.56 r/min 11.43 Nm\nAuthorized licensed use limited to: Univ of Calif Santa Barbara. Downloaded on May 15,2021 at 19:03:15 UTC from IEEE Xplore. Restrictions apply.", + "0278-0046 (c) 2021 IEEE. Personal use is permitted, but republication/redistribution requires IEEE permission. See http://www.ieee.org/publications_standards/publications/rights/index.html for more information.\n2\u03c4pr\n3\u03c4epr\n-160 -120 -80 -40 0 40 80 120 160\n-0.4\n-0.2\n0.0\n0.2\n0.4\nF lu\nx d\nen si\nty (\nT )\nCircumferential position (deg.)\nLnopower Lpower \u03c4pr_end\nFig.3. Air gap flux density of the rotary motion part with unpowered linear motion stator (sr=0.2) and both powered rotary and linear motion stators (sr =0.2 and sl =0.2).\nHowever, the no-load rotating speed of the 2DoFDDIM with\nhelical motion is smaller than that with corresponding single\nDoF rotary motion as listed in Table \u2161. Moreover, the average\nlocked torque of the motor with the single DoF rotary motion is\nhigher than that with helical motion. Besides, the no-load\nrotating speed and average locked torque of the motor with\nhelical motion increases with the decrease of the axial velocity.\nHence, it concludes that the resistant torque is produced due to\nthe linear motion component, which contributes to deteriorating\nthe rotary motion performances and increases with the increase\nof the axial velocity.\nThe average no-load linear velocity and locked force of the\n2DoFDDIM with a single DoF linear motion (only the linear\nmotion stator is powered) and a 2DoF helical motion with\ndifferent rotating speeds (the power is supplied for both motion\nstators) are listed in Table \u2162. Same as the analysis for the rotary\nmotion, the no-load linear velocity (4.03m/s) is higher than the\nideal synchronous velocity 3.9m/s on account of the effective\nelectromagnetic pole pitch shown in Fig.4. The resistant force\nis generated due to the rotary motion component, which\ndeteriorates the linear motion performances and increases with\nthe increase of the rotary speed.\nFig.4. Air gap flux density of the linear motion part with unpowered rotary motion stator (sl=0.2) and both powered rotary and linear motion stators (sr=0.2 and sl=0.2).\nIII. HELICAL MOTION COUPLING EFFECT MATHEMATICAL\nMODEL\nTo calculate the resistant torque and force, an HMCE mathematical model is derived based on the expanded model shown in Fig. 5. The following assumptions are made [21]: (1) The curvature of the mover is ignored. (2) Only are the fundamental components of armature reaction flux density considered.\nWhen the mover produces a rotary motion, the speed V only\nconsists of the circumferential component Vx (peripheral speed). Besides, the armature reaction flux density Bx, and the induced voltage Ex is produced in the mover. When the mover produces a linear motion, only do the axial components Vz, Bz and Ez exist\nin the mover. For the helical motion, x zV V i V k . Moreover,\nBy and Ey will be induced beside circumferential and axial components. Therefore, it can be derived that\nx y zB B i B j B k and x y zE E i E j E k .\nThe electromagnetic force density F can be calculated as\nshown in (2):\nF J B (2)\nwhere, J is the current density, ( )J E V B and \u03c3 is\nthe conductivity. Thus, it can be derived from (2) that\n2 2\n2\n( ) ( )\n( )\n( )\n( )\n(\nx y z z y x z z x x y\ny x y zx z\nyz\nx z z x z x y y z\nx z\nz x x x y z y z\nz y x\nA F E V B B B V B B\nt\ni j k i j k\nB B B V B V B V B V B\nA B B BA A\nt t t\nAA V B V B B V B B B i\nt t\nA A B B V B B V B B j\nt t\nV B V\n \n\n\n\n\n \n\n \n \n \n \n \n \n \n 2 ) y x\nx z z x x y\nA A B B V B B B k\nt t\n \n \n(3)\nFor the rotary motion, Vz=Bz=Ax=Ay=0, thus the electromagnetic force in the circumferential direction is\n2( ) ( )z z\nx x y y x x x y\nA A F V B B i B V B B j\nt t \n \n (4)\nFor the linear motion, Vx=Bx=Az=Ay=0. Then the electromagnetic force of axial direction is\n2( ) ( )x x\nz z z y z z y y\nA A F B V B B j V B B k\nt t \n \n (5)\nAuthorized licensed use limited to: Univ of Calif Santa Barbara. Downloaded on May 15,2021 at 19:03:15 UTC from IEEE Xplore. Restrictions apply." + ] + }, + { + "image_filename": "designv11_35_0001553_978-3-319-14705-5_14-Figure4-1.png", + "original_path": "designv11-35/openalex_figure/designv11_35_0001553_978-3-319-14705-5_14-Figure4-1.png", + "caption": "Fig. 4 Upper body model as a four-particle model. a Strict model. b Four-particle model", + "texts": [ + " (1) can be modified as follows: All Particles\u2211 i mi (mxi \u2212 mxzmp ) \u00d7 (m x\u0308i + x\u0308q + mG + m\u03c9\u0307 \u00d7 mxi + 2m\u03c9 \u00d7 m x\u0307i + m\u03c9 \u00d7 (m\u03c9 \u00d7 mxi )) \u2212 All Points\u2211 k (mxFk \u2212 mxzmp )\u00d7mFk \u2212 All Points\u2211 j mM j + mT0 = 0 (2) where xq = [ xq , yq , zq ]T is the position vector of the origin of the frame \u2211 m from the origin of the frame \u2211 O . m\u03c9 and m\u03c9\u0307 are the angular velocity and acceleration vectors, respectively. This equation is non-linear because the three-axis motion of the trunk is interferential each other. Therefore, it is difficult to derive analytically the compensatory motion of the trunk and the waist from Eq. (2). To obtain the approximate solution analytically, we assume the followings: (a) The external forces are not considered in the approximate model. (b) The upper body is modelled as a four-particle model (see Fig. 4). (c) The moving coordinate frame \u2211 m does not rotate. (d) The trunk and the waist do not move vertically. The moment generated by the motion of the lower-limb particles, M = [ Mx , My, Mz ]T , can be obtained as follows: mu mxu \u00d7 (m x\u0308u + m\u03c9\u0307 \u00d7 mxu + 2m\u03c9 \u00d7 m x\u0307u + m\u03c9 \u00d7 (m\u03c9 \u00d7 mxu )) + mt (mxt \u2212 mxzmp ) \u00d7 (m x\u0308t + x\u0308q + mG + m\u03c9\u0307 \u00d7 mxt +2m\u03c9 \u00d7 m x\u0307t + m\u03c9 \u00d7 (m\u03c9 \u00d7 mxt )) + mw (mxw \u2212 mxzmp ) \u00d7 (m x\u0308w + x\u0308q + mG + m\u03c9\u0307 \u00d7 mxw +2m\u03c9 \u00d7 m x\u0307w + m\u03c9 \u00d7 (m\u03c9 \u00d7 mxw )) = \u2212M (3) where mu is the mass of both shoulders including the mass of arms" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_35_0001702_1.1707312-Figure1-1.png", + "original_path": "designv11-35/openalex_figure/designv11_35_0001702_1.1707312-Figure1-1.png", + "caption": "FIG. 1.", + "texts": [ + " The experimental apparatus was practically the same as that already described,3 except that copper tubing was so placed about the bearing that lubricant could be introduced either in the loaded or unloaded portion, that is, at the base JOURNAL OF ApPLIED PHYSICS [This article is copyrighted as indicated in the article. Reuse of AIP content is subject to the terms at: http://scitation.aip.org/termsconditions. Downloaded to ] IP: 141.218.1.105 On: Wed, 24 Dec 2014 12:52:17 or crown of the bearing. The cases in which the source, which determines the polar axis, is at the crown is illustrated by Fig. 1. In order to secure better torque characteristics at extremely low speeds, a sliding gear transmission was introduced between the Graham variable speed transmission and the journal. Vibration was greatly reduced by mounting the driving mecha nism on rubber on a separate base, the journal being driven through two universal joints. The bearing length was 9.75 in. and its radius 1.75 in., giving a ratio of length to perimeter of 0.88. The ratio ric of journal radius to radial clearance was 1470. The flow was determined by collecting the oil, that had passed through the bearing, in a burette" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_35_0001419_978-94-007-7401-8-Figure2.3-1.png", + "original_path": "designv11-35/openalex_figure/designv11_35_0001419_978-94-007-7401-8-Figure2.3-1.png", + "caption": "Fig. 2.3 4RRRRR-type fully-parallel PMs with coupled Sch\u00f6nflies motions defined by MF \u00bc SF \u00bc 4; RF\u00f0 \u00de \u00bc \u00f0v1; v2; v3;xd\u00de, TF = 0, NF = 2, limb topology R||R||R\\R||R", + "texts": [ + "1 Limb topology and connecting conditions of the fully-parallel solutions with no idle mobilities presented in Figs. 2.2, 2.3, 2.4, 2.5, 2.6, 2.7, 2.8, 2.9, 2.10, 2.11, 2.12, 2.13, 2.14, 2.15, 2.16, 2.17, 2.18, 2.19, 2.20 No. PM type Limb topology Connecting conditions 1. 4RRRRR (Fig. 2.2a) R\\R||R||R\\k R (Fig. 2.1a) The first and the last revolute joints of the four limbs have parallel axes 2. 4RRRRR (Fig. 2.2b) R||R\\R||R\\R (Fig. 2.1b) The first, the second and the last revolute joints of the four limbs have parallel axes 3. 4RRRRR (Fig. 2.3a) R||R||R\\R||R (Fig. 2.1c) The two last revolute joints of the four limbs have parallel axes 4. 4RRRRR (Fig. 2.3b) R||R||R\\R||R (Fig. 2.1c) The three first revolute joints of the four limbs have parallel axes 5. 4RRRRR (Fig. 2.4a) R||R\\R||R||R (Fig. 2.1d) The two first revolute joints of the four limbs have parallel axes. 6. 4PRRRR (Fig. 2.4b) P\\R\\R||R||R (Fig. 2.1e) The second joints of the four limbs have parallel axes 7. 4RRRPR (Fig. 2.5a) R||R\\R\\P\\kR (Fig. 2.1f) Idem No. 5 8. 4RRPRR (Fig. 2.5b) R||R||P\\R||R (Fig. 2.1g) Idem No. 5 9. 4PRRRR (Fig. 2.6a) P\\R\\R||R\\R (Fig. 2.1h) The second and the last joints of the four limbs have parallel axes 10" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_35_0001445_gt2016-58093-Figure1-1.png", + "original_path": "designv11-35/openalex_figure/designv11_35_0001445_gt2016-58093-Figure1-1.png", + "caption": "Figure 1. BUILD ORIENTATION AND SUPPORT STRUCTURES (RED) OF THE DMLS COUPONS.", + "texts": [ + " Ten coupons were investigated all of which were made with DMLS; details on the coupons can be found in Table 1. Five of the coupons were made using cobaltchrome-molybdenum-based superalloy (CoCr) powder [13] with each having a unique channel width and height combination. The remaining five were made from an InconelTM 718 (Inco) powder [14] and nominally had the same geometry as the five made from CoCr. All ten coupons were 25.4 mm in length and in width. Each coupon was built at a 45\u00b0 angle from the horizontal as shown in Figure 1 using material specific build parameters calculated according to the recommendations of the DMLS machine manufacturer [15]. Snyder et al. [3] showed that a 45\u00b0 build direction results in higher friction drag than a vertical build direction for round channels. This is due to the large roughness on downward-facing surfaces that exists at every build orientation except vertical; vertically oriented surfaces are much smoother. Building all flow channels in a vertical orientation would seem like a simple method to minimize friction loss in a design, but it is highly unlikely that a working design of a gas turbine component would contain channels that are to be built only in the vertical direction" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_35_0000577_s40313-021-00754-5-Figure15-1.png", + "original_path": "designv11-35/openalex_figure/designv11_35_0000577_s40313-021-00754-5-Figure15-1.png", + "caption": "Fig. 15 UX-6 3D model", + "texts": [ + " A flight condition was equated with the UX-6 UAV steady-state flight condition (obtained from flight data). The UX-6 UAV steady-state flight condition was predicted to occur at AoA of 4 degrees, velocity of 15.3\u00a0m/s, and an altitude of 67\u00a0m. Figure\u00a014 shows UX-6 UAV geometry and flight scenario data were input to Datcom + Pro software. The UX-6 UAV geometry data and flight scenario that entered to Datcom + Pro software generated the aircraft 3D model and the aircraft calculated aerodynamic coefficients. The UX-6 UAV 3D model is shown in Fig.\u00a015. The calculated aerodynamic coefficients of Datcom + Pro were divided into static and dynamic aerodynamic coefficients. The static aerodynamic coefficients were derived from the calculation of the plane\u2019s geometry shape. The results from the static aerodynamic coefficients calculation are shown in Table\u00a06. The dynamic aerodynamic coefficient is obtained from the movement or deflection of the plane\u2019s control surface (elevator and aileron). The results from the static aerodynamic coefficients calculation are shown in Table\u00a07" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_35_0000008_j.autcon.2021.103573-Figure20-1.png", + "original_path": "designv11-35/openalex_figure/designv11_35_0000008_j.autcon.2021.103573-Figure20-1.png", + "caption": "Fig. 20. Gantry crane system with the initial condition error.", + "texts": [ + " The control inputs were properly calculated in the inverse dynamics solver. For this application, three cases were used to investigate the effect of the control method under the modeling uncertainties and external disturbances. The total simulation time was 500 s, and the operation time was assumed 400 s, including the acceleration and deceleration time of 200 s. After the block reached the final position, it maintained the position for 100 s. In Case 1 the initial condition error at the position of the block existed as a modeling uncertainty, as illustrated in Fig. 20. The block was initially located at position (1,0,20), a 1 m deviation from the desired trajectory. The resultant motion of the block is shown in Fig. 21. Due to modeling error, the block deviated from the desired trajectory by 1 m in the x-direction, depicted by the dotted black line at the start. Within 40 s, the block was controlled to precisely follow the target trajectory. The rotation angle, y, and z position followed the desired values exactly, with no perturbations. The inverse dynamics solver developed in this study calculated the control inputs to control the motion of the block" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_35_0001385_978-981-15-6095-8-Figure4.13-1.png", + "original_path": "designv11-35/openalex_figure/designv11_35_0001385_978-981-15-6095-8-Figure4.13-1.png", + "caption": "Fig. 4.13 Calculation model of the geocell-reinforced embankment (Zhang et al. 2010a, b)", + "texts": [ + "5) where \u201cEc\u201d is the tensile modulus of the geocell material and can be estimated by an indoor tensile test (ASTMD638-14); \u201c\u03b5\u201d is the tensile strain of the geocell material; \u201chg\u201d is the height of the geocell wall; \u201ca\u201d is the horizontal angle of the tensional force \u201cT\u201d. Before calculating \u201c\u03b5\u201d, the deformation shape of the reinforcement should be determined. Sophisticated numerical analyses have shown that the shape of the deflected geocell is a catenary (BS8006 1995; Yin 2000). However, at relatively small deflections the catenary may be approximated by a parabola which simplifies the analysis procedure for determining the tensile force in the geocell. As shown in Fig. 4.13, the deformation on the road surface is in the form of Eq. (4.6). y0 = \u2212 0 r20 x2 + h0 + s0 (4.6) where \u201cy0\u201d is the deformation on the road surface; \u201c s0\u201d is themaximum differential settlement at the surface; \u201ch0\u201d is the vertical distance from the origin of coordinates shown in Fig. 4.13 to the embankment surface. By differentiating Eqs. (4.5) and (4.7) is obtained. dy0 dx = \u22122 s0 r20 x (4.7) 4 Geocell-Reinforced Foundations 91 When x = r0, dy0/dx = \u22122 s0/r0. Supposing that the normal directions of points A and B on the deformation parabola are the same as the diffusion directions of embankment fill under the external load \u201cp\u201d, then, Eq. (4.8) can be presented. tan \u03b2 = 2 s0 r0 = r0 h0 = rn h0 + h (4.8) where \u201cb\u201d is the angle depicted in Fig. 4.13; \u201crn\u201d is the half of the chord length of parabola depicted in Fig. 4.13 and calculated by Eq. (4.9); \u201ch\u201d is the height of the embankment. rn = r20 + 2 s0h r0 (4.9) The relative deformation equation of the geocell reinforcement shown in Fig. 4.13 is in the form of Eq. (4.10). yn = \u2212 sn r2n x2 + h0 + h + sn (4.10) where \u201cyn\u201d is the deformation of the geocell reinforcement; \u201c sn\u201d is the maximum vertical deformation of the reinforcement. Be similar to Eqs. (4.6) and (4.11) is 92 G. Tavakoli Mehrjardi and S. N. Moghaddas Tafreshi obtained. sin \u03b1 = [ 1 + ( rn 2 sn )2 ]\u2212 1 2 (4.11) Then, the tensile strain of the geocell (\u03b5) is determined as Eq. (4.12). \u03b5 = 1 2 \u03b4 + rn 4 sn ln [ 2 sn rn + \u03b4 ] \u2212 1 (4.12) where \u03b4 is defined as Eq. (4.13). \u03b4 = [ 1 + ( 2 sn rn )2 ] 1 2 (4" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_35_0000750_j.matpr.2021.07.468-Figure1-1.png", + "original_path": "designv11-35/openalex_figure/designv11_35_0000750_j.matpr.2021.07.468-Figure1-1.png", + "caption": "Fig 1. Schematic of Stereolithography (SLA) 3D Printing [16].", + "texts": [ + " In general, there are five classes of 3D printing methods used in soft robotic applications, which include multiple methods that are similar in nature. Working principles of these five categories of 3D printing techniques, namely photopolymerization, powder bed fusion, material jetting, direct ink writing, fused deposition modeling, are briefly discussed, and a schematic diagram is presented. Stereolithography, commonly known as SLA, is one of the earliest forms of vat photopolymerization method (Fig. 1) of 3D printing. In SLA, a vat of resin is selectively photopolymerized using a laser or other UV lights to print a layer of the desired object. The next layer is then printed by moving the light beam on an XY plane as the build platform lowers to allow the fresh layer of resin to rise on the surface, and the process continues several times until the completion of a desired 3D object. Several other vat photopolymerization techniques, such as digital light processing (DLP), light crystal display (LCD), two-photon polymerization (2PP), continuous liquid interface production (CLIP), follow the same principle as SLA" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_35_0001419_978-94-007-7401-8-Figure2.109-1.png", + "original_path": "designv11-35/openalex_figure/designv11_35_0001419_978-94-007-7401-8-Figure2.109-1.png", + "caption": "Fig. 2.109 4PRPaRR-type fully-parallel PM with coupled Sch\u00f6nflies motions defined by MF \u00bc SF \u00bc 4; RF\u00f0 \u00de \u00bc \u00f0v1; v2; v3;xd\u00de, TF = 0, NF = 14, limb topology P||R\\Pa\\kR\\R", + "texts": [], + "surrounding_texts": [ + "2.2 Topologies with Complex Limbs 163", + "164 2 Fully-Parallel Topologies with Coupled Sch\u00f6nflies Motions", + "2.2 Topologies with Complex Limbs 165" + ] + }, + { + "image_filename": "designv11_35_0000213_j.triboint.2021.106993-Figure13-1.png", + "original_path": "designv11-35/openalex_figure/designv11_35_0000213_j.triboint.2021.106993-Figure13-1.png", + "caption": "Fig. 13. Vapor volume fraction contour with applied load and operating speed changes; (a) 5000 RPM, (b) 9000 RPM, (c) 13,000 RPM; (1) 5000 N, (2) 15,000 N.", + "texts": [ + " Thus, the pressure suddenly drops at the leading edge of the pockets, generating cavitation in the pocket. The negative gauge pressure inside the pockets is recovered at the trailing edge of the pocket. In addition, Fig. 12(b) shows that the negative pressure inside the pockets causes the significant phase change (high vapor volume fraction) from the liquid phase to the gas phase, while cavitation is not observed in the smooth TPJB result, as seen in Fig. 12 (a). The vaporization is activated more at the higher speeds, as demonstrated in Fig. 13, because the pressure drop inside the pockets becomes larger at the higher speeds. Fig. 13 also shows that the degree of cavitation is nearly insensitive to the applied load. Further simulations have demonstrated that a high level of cavitation occurs for a wide range of operating conditions. This is good since it suggests that the proposed novel features for the TPJB may be beneficial for a wide range of applications. In the pockets, the evaporation significantly lessens the dynamic viscosity (\u03bcf ) of the fluid, and the larger film thickness decreases the velocity gradient (\u2202ucir/\u2202r)" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_35_0001419_978-94-007-7401-8-Figure6.10-1.png", + "original_path": "designv11-35/openalex_figure/designv11_35_0001419_978-94-007-7401-8-Figure6.10-1.png", + "caption": "Fig. 6.10 Maximaly regular fully-parallel PMs with Sch\u00f6nflies motions of types 2PRRRR1PRRR-1RPaPaP (a) and 2PRRRR-1PRRR-1RPaPatP (b) defined by MF = SF = 4, (RF) = (v1; v2; v3;xb), TF = 0, NF = 10, limb topology P||R||R||R\\R, P||R||R||R and R||Pa||Pa||P (a), R||Pa||Pat||P (b)", + "texts": [ + " 2PPRRR-1PPRR-1RPaPaP (Fig. 6.8a) P ?P\\||R||R\\R (Fig. 4.2a) P\\P\\||R||R (Fig. 4.1d) R||Pa||Pa||P (Fig. 5.4k) Idem no. 1 4. 2PPRRR-1PPRR1RPaPatP (Fig. 6.8b) P\\P\\||R||R\\R (Fig. 4.2a) P\\P\\||R||R (Fig. 4.1d) R||Pa||Pat||P (Fig. 5.4l) Idem no. 1 5. 2PRPRR-1PRPR-1RPaPaP (Fig. 6.9a) P||R\\P\\||R\\R (Fig. 4.1b) P||R\\P\\||R (Fig. 4.1e) R||Pa||Pa||P (Fig. 5.4k) Idem no. 1 6. 2PRPRR-1PRPR1RPaPatP (Fig. 6.9b) P||R\\P\\||R\\R (Fig. 4.1b) P||R\\P\\||R (Fig. 4.1e) R||Pa||Pat||P (Fig. 5.4l) Idem no. 1 7. 2PRRRR-1PRRR-1RPaPaP (Fig. 6.10a) P||R||R||R\\R (Fig. 4.2d) P||R||R||R (Fig. 4.1g) R||Pa||Pa||P (Fig. 5.4k) Idem no. 1 8. 2PRRRR-1PRRR1RPaPatP (Fig. 6.10b) P||R||R||R\\R (Fig. 4.2d) P||R||R||R (Fig. 4.1g) R||Pa||Pat||P (Fig. 5.4l) Idem no. 1 9. 3PPPR-1RPPaPa (Fig. 6.11a) P\\P ??P\\||R (Fig. 4.1a) P\\P ??P ??R (Fig. 4.1b) R||P||Pa||Pa (Fig. 5.4m) The last revolute joints of limbs G1, G2 and G3 and the revolute joint between links 7D and 9 have superposed axes. The actuated prismatic joints of limbs G1, G2 and G3 have orthogonal directions 10. 3PPPR-1RPPaPat (Fig. 6.11b) P\\P ??P\\||R (Fig. 4.1a) P\\P ??P ??R (Fig. 4.1b) R||P||Pa||Pat (Fig. 5.4n) Idem no. 9 11" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_35_0000067_jestpe.2021.3061120-Figure6-1.png", + "original_path": "designv11-35/openalex_figure/designv11_35_0000067_jestpe.2021.3061120-Figure6-1.png", + "caption": "Fig. 6. No-load flux distribution of LSPMVMs. (a) Type I: V-shaped LSPMVM. (b) Type II: Consequent-pole U-shaped LSPMVM. (c) Type III: Dual-stator spoke LSPMVM. (d) Type IV: Consequent-pole SPM LSPMVM. (e) Type V: LSPMSM.", + "texts": [ + " 5, where U, E, I, Id, Iq, Xd, Xq, Rs and \u03b8 are the phase voltage, back-electromotive force (EMF), stator winding current, d-axis current, q-axis current, d-axis inductance, q-axis inductance and stator winding resistance and the power angle between back-EMF and current, respectively. The track of voltage vector is a circular, where the amplitude of which is equal to the grid phase voltage 220 V. The no-load magnetic field distributions of four LSPMVMs and the regular commercial LSPMSM are displayed in Fig. 6. It should be noticed that due to the introduction of rotor starting units, the rotor are saturated at the no-load state. Especially, the rotor saturation is more serious in Type IV, that is, consequent-pole SPM LSPMVM, due to both the rotor windings and PMs are located at the rotor outer circumference. For the LSPMSM, the stator teeth are more saturated due to the increase of stator slot number. Moreover, it can be observed that compared with V-shaped LSPMVM, the consequent-pole U-shaped LSPMVM shows less flux leakage, which is conductive to the high air-gap flux density and back-EMF" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_35_0000776_s10854-021-06832-3-Figure6-1.png", + "original_path": "designv11-35/openalex_figure/designv11_35_0000776_s10854-021-06832-3-Figure6-1.png", + "caption": "Fig. 6 Schematic of an enzyme-free glucose sensor integrated with working electrode and counter electrode", + "texts": [ + " Figure 5c shows the cyclic voltammetry curves at different scan rates. It is obvious that the absolute values of the oxidation and reduction peak current densities increase with the scan rate in the range of 10\u2013100 mV s-1. Besides, the oxidation and reduction peak current densities are proportional to the square root of scan rate, with R2 of 0.99642 and 0.99625, respectively, shown in Fig. 5d, demonstrating a diffusion-controlled process. The enzyme-free glucose sensor consists of a working electrode and a counter electrode, as illustrated in Fig. 6. The working electrode is composed of bimetallic PtNi material, and glucose can be oxidized on the surface of the electrode. The active areas of the working electrodes are designed to be 10 9 10 mm2, 5 9 5 mm2, and 1 9 1 mm2, respectively. The counter electrode is made of Pt, with the width of 400 lm. Si3N4 is served as an insulating layer between the working electrode and counter electrode. The enzyme-free glucose sensors based on bimetallic PtNi materials were fabricated by MEMS techniques, as shown in Fig" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_35_0001419_978-94-007-7401-8-Figure2.34-1.png", + "original_path": "designv11-35/openalex_figure/designv11_35_0001419_978-94-007-7401-8-Figure2.34-1.png", + "caption": "Fig. 2.34 4PaPPR-type fully-parallel PMs with coupled Sch\u00f6nflies motions defined by MF \u00bc SF \u00bc 4; RF\u00f0 \u00de \u00bc \u00f0v1; v2; v3;xd\u00de, TF = 0, NF = 18, limb topology Pa||P\\P||R (a) and Pa\\P\\kP\\kR (b)", + "texts": [ + "30b) P\\P||Pa\\R (Fig. 2.21j) Idem No. 4 11. 4PPaRP (Fig. 2.31a) P||Pa\\R||P (Fig. 2.21k) The before last joints of the four limbs have parallel axes 12. 4PPaPR (Fig. 2.31b) P||Pa\\P|| R (Fig. 2.21l) Idem No. 4 13. 4PPaPR (Fig. 2.32a) P\\Pa||P\\\\R (Fig. 2.21m) Idem No. 4 14. 4PPaRP (Fig. 2.32b) P\\Pa\\kR\\kP (Fig. 2.21n) Idem No. 4 15. 4RPPaP (Fig. 2.33a) R\\P||Pa\\kP (Fig. 2.21o) The first revolute joints of the four limbs have parallel axes 16. 4RPPPa (Fig. 2.33b) R||P\\P||Pa (Fig. 2.21p) Idem No. 15 17. 4PaPPR (Fig. 2.34a) Pa||P\\P||R (Fig. 2.21q) Idem No. 4 18. 4PaPPR (Fig. 2.34b) Pa\\P\\kP\\kR (Fig. 2.21r) Idem No. 4 19. 4PaRPP (Fig. 2.35a) Pa\\R||P\\P (Fig. 2.21s) Idem No. 11 20. 4PaPRP (Fig. 2.35b) Pa||P\\R||P (Fig. 2.21t) Idem No. 5 21. 4PaPRP (Fig. 2.36a) Pa\\P||R\\P (Fig. 2.21u) Idem No. 11 22. 4PaRPP (Fig. 2.36b) Pa\\R\\P\\kP (Fig. 2.21v) Idem No. 5 23. 4PCPa (Fig. 2.37a) P\\C\\Pa (Fig. 2.21w) Idem No. 5 24. 4PPaC (Fig. 2.37b) P||Pa\\C (Fig. 2.21x) The last joints of the four limbs have parallel axes 25. 4CPPa (Fig. 2.38a) C\\P||Pa (Fig. 2.21y) The first joints of the four limbs have parallel axes 26" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_35_0000435_j.cirp.2021.04.050-Figure5-1.png", + "original_path": "designv11-35/openalex_figure/designv11_35_0000435_j.cirp.2021.04.050-Figure5-1.png", + "caption": "Fig. 5. Motion of robot manipulator during camera localization.", + "texts": [ + " The camera takes an image of the tag again, and the transformation from the localized camera frame to the tag frame M0 5 is obtained. Then, the absolute error of the camera pose (Dx, Dy, Dz, Drx, Dry, Drz) is calculated by M2 and M0 5. If each error is below its threshold, the camera localization process finishes, and then the workpiece exchange process starts. If any component is above the threshold, the camera localization process continues following a flowchart shown in Fig. 4. Each threshold is arbitrarily determined depending on the clearance between the chuck and the workpiece. In Fig. 5, the motions of the robot manipulator are illustrated during the camera localization process: the robot manipulator with a broken line shows the initial pose, while the one with a solid line shows the pose after localizing the camera. The transformations M4 andM5 are defined in Section 3.2. If the camera localization process finishes successfully in accordance with the flowchart in Fig. 4, the transformation from the current robot frame to the goal frame is calculated following Eq. (5). current_robot goal M \u00bccurrent_robot taught_camera M \u00a2M0 5 \u00a2 taggoalM \u00f05\u00de Please cite this article as: Y" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_35_0000370_j.sysconle.2021.104931-Figure1-1.png", + "original_path": "designv11-35/openalex_figure/designv11_35_0000370_j.sysconle.2021.104931-Figure1-1.png", + "caption": "Fig. 1. Rotational manipulator arm consisting of two links and a passive joint. Source: The figure is taken from [11].", + "texts": [ + " A passive joint consisting of a linear spring\u2013damper combination couples the two links to each other. Passive in this context means, that there is no input force at this point. We stress, that the linearity of the passive joint does not result in linear equations of motion, see (31). As an output we measure the position of point S on the second link. Using a d ( a F y a N a s \u03b2 c d c R N L t i N s s R t c body fixed coordinate system the point S on the passive link is escribed by 0 \u2264 s \u2264 l. The situation is depicted in Fig. 1. We present the manipulator\u2019s equations of motion. As we will see in a second, it is reasonable to consider the dynamics of the manipulator for \u03b2 \u2208 B := { \u03b2 \u2208 R cos(\u03b2) \u0338= 2l 3s } . Henceforth, we assume \u03b2 \u2208 B and perform the computations. We define U\u03b2 := R \u00d7 B \u00d7 R2, and set M : B \u2192 R2\u00d72, x2 \u21a6\u2192 l2m [ 5 3 + cos(x2) 1 3 + 1 2 cos(x2) 1 3 + 1 2 cos(x2) 1 3 ] , f1 : U\u03b2 \u2192 R x1, x2, x3, x4)\u22a4 \u21a6\u2192 1 2 l 2mx4(2x3 + x4) sin(x2), f2 : U\u03b2 \u2192 R (x1, x2, x3, x4)\u22a4 \u21a6\u2192 \u2212cx2 \u2212 dx4 \u2212 1 2 l2mx23 sin(x2), nd obtain the equations of motion( \u03b1\u0308(t) \u03b2\u0308(t) ) = M(\u03b2(t))\u22121 ( f1(\u03b1(t), \u03b2(t), \u03b1\u0307(t), \u03b2\u0307(t)) f2(\u03b1(t), \u03b2(t), \u03b1\u0307(t), \u03b2\u0307(t)) ) + M(\u03b2(t))\u22121 [ 1 0 ] u(t)" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_35_0001419_978-94-007-7401-8-Figure3.140-1.png", + "original_path": "designv11-35/openalex_figure/designv11_35_0001419_978-94-007-7401-8-Figure3.140-1.png", + "caption": "Fig. 3.140 4RCPaR-type overactuated PM with coupled Sch\u00f6nflies motions defined by MF \u00bc SF \u00bc 4; RF\u00f0 \u00de \u00bc v1; v2; v3;xd\u00f0 \u00de, TF = 0, NF = 14, limb topology R||C\\Pa||R", + "texts": [ + " 4PaRPRR (Fig. 3.132) Pa\\R||P||R\\kR (Fig. 3.52z) Idem No. 13 47. 4PaRPRR (Fig. 3.133) Pa\\R||P||R\\kR (Fig. 3.52a0) Idem No. 13 48. 4PaRRPR (Fig. 3.134) Pa\\R||R||P\\kR (Fig. 3.52b0) Idem No. 13 49. 4RCRPa (Fig. 3.135) R\\C||R\\kPa (Fig. 3.52c0) Idem No. 12 50. 4RRCPa (Fig. 3.136) R\\R||C\\kPa (Fig. 3.52d0) Idem No. 12 51. 4CRRPa (Fig. 3.137) C||R\\R||Pa (Fig. 3.52e0) Idem No. 15 52. 4RCRPa (Fig. 3.138) R||C\\R||Pa (Fig. 3.52f0) Idem No. 15 53. 4CRPaR (Fig. 3.139) C||R\\Pa||R (Fig. 3.52g0) Idem No. 13 54. 4RCPaR (Fig. 3.140) R||C\\Pa||R (Fig. 3.52h0) Idem No. 13 55. 4PaCRR (Fig. 3.141a) Pa\\C||R\\kR (Fig. 3.52i0) Idem No. 13 56. 4PaRCR (Fig. 3.141b) Pa\\R||C\\kR (Fig. 3.52j0) Idem No. 13 57. 4PaRRRPa (Fig. 3.142) Pa\\R||R\\kR||Pa (Fig. 3.53a) Idem No. 15 58. 4RPaRRPa (Fig. 3.143) R\\Pa\\kR\\kR||Pa (Fig. 3.53b) Idem No. 15 59. 4RRRPaPa (Fig. 3.144) R||R\\R||Pa||Pa (Fig. 3.53c) Idem No. 15 60. 4RPaRPaR (Fig. 3.145) R\\Pa\\kR\\kPa||R(Fig. 3.53d) Idem No. 13 61. 4PaPaRRR (Fig. 3.146) Pa||Pa\\R||R\\kR (Fig. 3.53e) Idem No. 13 62. 4PaPaRRR (Fig" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_35_0001419_978-94-007-7401-8-Figure5.90-1.png", + "original_path": "designv11-35/openalex_figure/designv11_35_0001419_978-94-007-7401-8-Figure5.90-1.png", + "caption": "Fig. 5.90 3PaPaPaR-1CPaPa-type fully-parallel PM with uncoupled Sch\u00f6nflies motions defined by MF = SF = 4, RF\u00f0 \u00de \u00bc \u00f0v1; v2; v3;xb\u00de, TF = 0, NF = 39, limb topology Pa\\Pa||Pa\\\\R, Pa\\Pa||Pa||R and C||Pa||Pa", + "texts": [ + "40) 25. 2PaPaRRR1PaPaRR1CPaPat (Fig. 5.87) Pa\\Pa||R||R\\\\R (Fig. 5.5a) Idem No. 23 Pa\\Pa||R||R (Fig. 5.4i) C||Pa||Pat (Fig. 5.4p) 26. 2PaPaRRR1PaPaRR1CPaPat (Fig. 5.88) Pa\\Pa||R||R\\kR (Fig. 5.5b) Idem No. 23 Pa\\Pa||R||R (Fig. 5.4i) C||Pa||Pat (Fig. 5.4p) 27. 3PaPaPaR1CPaPa (Fig. 5.89) Pa\\Pa||Pa\\kR (Fig. 5.6a) The revolute joint between links 6D and 11 of G4 limb and the last revolute joints of limbs G1, G2 and G3 have superposed axes Pa\\Pa||Pa||R (Fig. 5.6b) C||Pa||Pa (Fig. 5.4o) 28. 3PaPaPaR1CPaPa (Fig. 5.90) Pa\\Pa||Pa\\\\R (Fig. 5.6c) Idem No. 27 Pa\\Pa||Pa||R (Fig. 5.6b) C||Pa||Pa (Fig. 5.54o) 29. 3PaPaPaR1CPaPat (Fig. 5.91) Pa\\Pa||Pa\\kR (Fig. 5.6a) Idem No. 27 Pa\\Pa||Pa||R (Fig. 5.6b) C||Pa||Pat (Fig. 5.4p) 30. 3PaPaPaR1CPaPat (Fig. 5.92) Pa\\Pa||Pa\\\\R (Fig. 5.6c) Idem No. 27 Pa\\Pa||Pa||R (Fig. 5.6b) C||Pa||Pat (Fig. 5.54p) 536 5 Topologies with Uncoupled Sch\u00f6nflies Motions No. Structural parameter Solution Figure 5.7 Figures 5.8, 5.9, 5.10 Figures 5.11 1. m 20 22 22 2. pi (i = 1, 3) 7 8 7 3. p2 7 7 7 4" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_35_0000560_lra.2021.3086670-Figure9-1.png", + "original_path": "designv11-35/openalex_figure/designv11_35_0000560_lra.2021.3086670-Figure9-1.png", + "caption": "Fig. 9. A 3D voyage of a mini-propeller. Eleven key waypoints are numbered, and the magnetic field is programmed as parameters in Table II.", + "texts": [ + " Both the splitting force and centrifugal force increase with the frequency (Fig. 8(d)), and the sum of them may exceed the magnetic bond between modules (\u2265 14Hz). Consequently, the assembled structure of a mini-propeller becomes unstable, and these magnetic fields can be adopted to control the disassembly of mini-propellers. V. 3D LOCOMOTION AND UPSTREAM SWIMMING A mini-propeller can manage controllable 3D motion and travel through a 3D maze (Movie S3). Here we designed a 3D trajectory to demonstrate the motion and orientation (Fig. 9). The framework for constructing such a 3D trajectory was fabricated by 3D printing, and it was submerged in the solution. The whole trajectory comprises of horizontal motion (cyan) at different altitudes and vertical motion (orange) for approaching a low-altitude door and high-altitude window. Three nodal positions (magenta) were set during the trajectory as origin and destinations. The magnetic field was reprogrammed to adjust the swimming pose of the mini-propeller for navigation at key waypoints, and detailed parameters are shown in Table II" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_35_0001385_978-981-15-6095-8-Figure1.12-1.png", + "original_path": "designv11-35/openalex_figure/designv11_35_0001385_978-981-15-6095-8-Figure1.12-1.png", + "caption": "Fig. 1.12 Reinforcement function of geosynthetics: a tensile member; and b tensioned membrane. Sourced from Fluet (1988)", + "texts": [ + " For soil\u2013geosynthetics interaction problems, interface shear strength is the significant parameter. It is responsible for the transference of stresses from soil to geosynthetics. The tensile modulus helps to maintain the total deformation of the composite material within the allowable limits of the soil. Initially, the function of reinforcement was divided into two types, namely, tensile element and the tensioned membrane (Fluet 1988). The tensile member is able to support the planar load as shown in Fig. 1.12a. The planar and normal loads can be supported in the case of the tensioned membrane as shown in Fig. 1.12b. Later, Koerner (1998) and Jewell (1996) revealed that the geosynthetics works as a tensile member due to anchorage and shear mechanisms between the soil and reinforcement. Thus, they have suggested three different mechanisms, namely, shear, anchorage, and tensioned membrane for the soil reinforcement function. Shearing is also referred to as sliding, which occurs due to the sliding of soil particles over the geosynthetic. Anchorage is known as pullout and occurs due to the pullout action of geosynthetics from the soil" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_35_0000686_j.mechmachtheory.2021.104433-Figure18-1.png", + "original_path": "designv11-35/openalex_figure/designv11_35_0000686_j.mechmachtheory.2021.104433-Figure18-1.png", + "caption": "Fig. 18. Metamorphic epicyclic external gear train based on force constraint.", + "texts": [ + " When the gear arm reaches the limit position and is rigidified with the geometrical limit block which means the angular velocity \u03c9e2 of the gear arm has been constrained to equal to \u03c9e1. Then the corresponding configuration transforms from configuration 2 to configuration 1, and the mobility has degenerated from 2 to 1. The metamorphic epicyclic external gear train is in revolution configuration (Fig. 15(b)). Likewise, a set of metamorphic epicyclic gear trains are constructed by adding force constraints to control the constraint condition. As shown in Fig. 18\u201320, the spring pin which consists of a block and spring provides a certain constraint force. For instance, the metamorphic epicyclic bevel gear train is in revolution configuration (Fig.19 (b)) when the bevel gear arm has been locked to the sun bevel gear by the spring pin. At this moment, the angular velocity \u03c9b2 is equal to \u03c9b1, which means the corresponding motion branch has mobility 1. Until driving force drive the bevel gear arm to overcome the constraint force (Fig. 19(a)), the metamorphic epicyclic bevel gear train transforms to rotation configuration which is similar to the example discussed above and has mobility 2 H" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_35_0001419_978-94-007-7401-8-Figure5.46-1.png", + "original_path": "designv11-35/openalex_figure/designv11_35_0001419_978-94-007-7401-8-Figure5.46-1.png", + "caption": "Fig. 5.46 3PaPaPR-1RPaPatP-type fully-parallel PM with uncoupled Sch\u00f6nflies motions defined by MF = SF = 4, RF\u00f0 \u00de \u00bc \u00f0v1; v2; v3;xb\u00de, TF = 0, NF = 30 limb topology Pa\\Pa\\\\P\\\\R, Pa\\Pa\\\\P\\kR and R||Pa||Pat||P", + "texts": [ + "4d) R||P||Pa||Pat (Fig. 5.4n) 8. 3PaPPaR-1RPPaPat (Fig. 5.42b) Pa\\P\\\\Pa\\kR (Fig. 5.4b) Idem No. 3 Pa\\P\\\\Pa||R (Fig. 5.4c) R||P||Pa||Pat (Fig. 5.4n) 9. 3PaPaPR-1RPaPaP (Fig. 5.43) Pa\\Pa\\kP\\\\R (Fig. 5.4f) Idem No. 1 Pa\\Pa\\kP\\kR (Fig. 5.4h) R||Pa||Pa||P (Fig. 5.4k) 10. 3PaPaPR-1RPaPaP (Fig. 5.44) Pa\\Pa\\\\P\\\\R (Fig. 5.4g) Idem No. 1 Pa\\Pa\\\\P\\kR (Fig. 5.4e) R||Pa||Pa||P (Fig. 5.4k) 11. 3PaPaPR1RPaPatP (Fig. 5.45) Pa\\Pa\\kP\\\\R (Fig. 5.4f) Idem No. 1 Pa\\Pa\\kP\\kR (Fig. 5.4h) R||Pa||Pat||P (Fig. 5.4l) 12. 3PaPaPR1RPaPatP (Fig. 5.46) Pa\\Pa\\\\P\\\\R (Fig. 5.4g) Idem No. 1 Pa\\Pa\\\\P\\kR (Fig. 5.4e) R||Pa||Pat||P (Fig. 5.4l) 13. 3PaPaPR-1RPPaPa (Fig. 5.47) Pa\\Pa\\kP\\\\R (Fig. 5.4f) Idem No. 3 Pa\\Pa\\kP\\kR (Fig. 5.4h) R||P||Pa||Pa (Fig. 5.4m) 14. 3PaPaPR-1RPPaPa (Fig. 5.48) Pa\\Pa\\\\P\\\\R (Fig. 5.4g) Idem No. 3 Pa\\Pa\\\\P\\kR (Fig. 5.4e) R||P||Pa||Pa (Fig. 5.4m) 15. 3PaPaPR-1RPPaPat (Fig. 5.49) Pa\\PakP\\\\R (Fig. 5.4f) Idem No. 3 Pa\\Pa\\kP\\kR (Fig. 5.4h) R||P||Pa||Pat (Fig. 5.4n) 16. 3PaPaPR-1RPPaPat (Fig. 5.50) Pa\\Pa\\\\P\\\\R (Fig. 5.4g) Idem No" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_35_0001419_978-94-007-7401-8-Figure6.21-1.png", + "original_path": "designv11-35/openalex_figure/designv11_35_0001419_978-94-007-7401-8-Figure6.21-1.png", + "caption": "Fig. 6.21 3PPPaR-1RPPP-type maximaly regular fully-parallel PMs with Sch\u00f6nflies motions defined by MF = SF = 4, (RF) = (v1; v2; v3;xb), TF = 0, NF = 15, limb topology P\\P\\||Pa||R, R\\P ??P ??P and P\\P\\||Pa\\||R (a), P\\P\\||Pa ??R (b)", + "texts": [ + "1d) C||Pa||Pat (Fig. 5.4p) Idem no. 1 5. 2CPRR-1CPR-1CPaPa (Fig. 6.19a) C ?P ?||R ?R (Fig. 4.2e) C ?P ?||R (Fig. 4.1i) C||Pa||Pa (Fig. 5.4o) Idem no. 1 6. 2CPRR-1CPR-1CPaPat (Fig. 6.19b) C ?P ?||R ?R (Fig. 4.2e) C ?P ?||R (Fig. 4.1i) C||Pa||Pat (Fig. 5.4p) Idem no. 1 7. 2CRRR-1CRR-1CPaPa (Fig. 6.20a) C||R||R ?R (Fig. 4.2g) C||R||R (Fig. 4.1k) C||Pa||Pa (Fig. 5.4o) Idem no. 1 8. 2CRRR-1CRR-1CPaPat (Fig. 6.20b) C||R||R ?R (Fig. 4.2g) C||R||R (Fig. 4.1k) C||Pa||Pat (Fig. 5.4p) Idem no. 1 9. 3PPPaR-1RPPP (Fig. 6.21a) P ?P ?||Pa ?||R (Fig. 4.8b) P ?P ?||Pa||R, (Fig. 4.8c) R ?P\\\\P\\\\P (Fig. 5.1a) The last joints of the four limbs have superposed axes/ directions. The actuated prismatic joints of limbs G1, G2 and G3 have orthogonal directions 10. 3PPPaR-1RPPP (Fig. 6.21b) P ?P ?||Pa\\\\R (Fig. 4.8a) P ?P ?||Pa||R, (Fig. 4.8c) R\\\\P\\\\P\\\\P (Fig. 5.1a) Idem no. 9 11. 3PPaPR-1RPPP (Fig. 6.22a) P||Pa ?P\\\\R (Fig. 4.8d) P||Pa ?P ?||R (Fig. 4.8e) R ?P\\\\P\\\\P (Fig. 5.1a) Idem no. 9 12. 2PPaRRR-1PPaRR-1RPPP (Fig. 6.22b) P||Pa||R||R ?R (Fig. 4.9a) P||Pa||R||R (Fig. 4.8g) R ?P\\\\P\\\\P (Fig. 5.1a) Idem no. 9 13. 3PPPaR-1RUPU (Fig. 6.23a) P ?P ?||Pa ?||R (Fig. 4.8b) P ?P ?||Pa||R, (Fig. 4.8c) R ?R ?R ?P ?||R ?R (Fig. 5.1b) The first revolute joint of G4-limb and the last last joints of limbs G1, G2 and G3 have parallel axes" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_35_0000745_j.apm.2021.07.014-Figure1-1.png", + "original_path": "designv11-35/openalex_figure/designv11_35_0000745_j.apm.2021.07.014-Figure1-1.png", + "caption": "Fig. 1. Schematic view of the QUAV system.", + "texts": [ + " Initially, dynamic models (linearization or other forms) of the attitude of the QUAV were developed, ignoring one or more of the effects of air resistance, internal force and gyroscopic effects. The attitude model of the quadrotor is given in [10] as follows: { I x \u02d9 p = ( I y \u2212 I z ) qr \u2212 I r q \u2212 k 4 p + \u03c4\u03c6 I y \u0307 q = ( I z \u2212 I x ) pr + I r p \u2212 k 5 q + \u03c4\u03b8 I z \u0307 r = ( I x \u2212 I y ) pq \u2212 k 6 r + \u03c4\u03c8 . (1) which takes into account the gyroscopic effect and air resistance. It is a complete dynamic model of the attitude system of the QUAV without disturbance. Where i ( i = 1, 2, 3, 4) is the i th propeller rotor speed of the QUAV ( Fig. 1 ), = 2 + 4 \u2212 1 \u2212 3 the sum speed of four rotors, I x , I y , I z the three moments of inertia of the three axes x , y , z , respectively, I r the total inertia moment of the four rotors, k i ( i = 4, 5, 6) drag coefficients, \u03c9 = [ p q r ] T the attitude angular velocity vector with respect to the B-frame and f = [ \u03c4\u03c6 \u03c4\u03b8 \u03c4\u03c8 ] = [ bl( 2 4 \u2212 2 2 ) bl( 2 3 \u2212 2 1 ) d( 2 4 + 2 2 \u2212 2 3 \u2212 2 1 ) ] , (2) where b the lift coefficient, l the distance from QUAV center to the propeller center, d the drag coefficient, and \u03c4\u03c6 = bl( 2 4 \u2212 2 2 ) , \u03c4\u03b8= bl( 2 3 \u2212 2 1 ) , \u03c4\u03c8 = d( 2 4 + 2 2 \u2212 2 3 \u2212 2 1 ) , (3) the lift torques", + " Set = [ \u03c6 \u03b8 \u03c8 ] T the angular vector with respect to the E-frame, \u03d5 the roll angle, \u03b8 the pitch angle, \u03c8 the yaw angle, \u02d9 = [ \u02d9 \u03c6 \u02d9 \u03b8 \u02d9 \u03c8 ] T the attitude angular velocity vector. The relationship between \u02d9 = [ \u02d9 \u03c6 \u02d9 \u03b8 \u02d9 \u03c8 ] T and \u03c9 = [ p q r ] T is expressed as \u02d9 = A ( ) \u03c9, (4) where A ( ) = [ 1 sin \u03c6 tan \u03b8 cos \u03c6 tan \u03b8 0 cos \u03c6 \u2212 sin \u03c6 0 sin \u03c6/ cos \u03b8 cos \u03c6/ cos \u03b8 ] . (5) When the QUAV performs yaw maneuver, the front rotor speed 1 is equal to the rear rotor speed 3 , and the right rotor speed 2 is equal to the left rotor speed 4 , i.e., 1 = 3 , 2 = 4 ( Fig. 1 ). Let the speed difference between the front rotor and the left rotor is c, c = 2 \u2212 1 = 0. Then, the total speed of the four rotors under yaw maneuver is = 2 c , \u03c4\u03d5 = \u03c4 \u03b8 = 0, \u03c4\u03c8 = 2 d ( c 2 + 2 c 1 ) and the system (1) is simplified as { I x \u02d9 p = ( I y \u2212 I z ) qr \u2212 2 c I r q \u2212 k 4 p I y \u0307 q = ( I z \u2212 I x ) pr + 2 c I r p \u2212 k 5 q I z \u0307 r = ( I x \u2212 I y ) pq \u2212 k 6 r + 2 d( c 2 + 2 c 1 ) . (6) Eq. (6) is called a no-disturbed attitude system. The developed dynamic models of the QUAV usually ignore the effects of disturbance" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_35_0001419_978-94-007-7401-8-Figure5.79-1.png", + "original_path": "designv11-35/openalex_figure/designv11_35_0001419_978-94-007-7401-8-Figure5.79-1.png", + "caption": "Fig. 5.79 3PaPPaR-1CPaPa-type fully-parallel PMs with uncoupled Sch\u00f6nflies motions defined by MF = SF = 4, RF\u00f0 \u00de \u00bc \u00f0v1; v2; v3;xb\u00de, TF = 0, NF = 30 limb topology C||Pa||Pa and Pa||P\\Pa\\kR, Pa||P\\Pa||R (a), Pa\\P\\\\Pa\\kR, Pa\\P\\\\Pa||R (b)", + "texts": [ + "2f) C||Pa||Pat (Fig. 5.4p) 11. 2PaRPRR1PaRPR1CPaPa (Fig. 5.75) Pa\\R\\P\\kR\\R (Fig. 5.3c) Idem No. 3 Pa\\R\\P\\kR (Fig. 5.2g) C||Pa||Pa (Fig. 5.40) 12. 2PaRPRR1PaRPR1CPaPat (Fig. 5.76) Pa\\R\\P\\kR\\R (Fig. 5.3c) Idem No. 3 Pa\\R\\P\\kR (Fig. 5.2g) C||Pa||Pat (Fig. 5.4p) 13. 2PaRRRR1PaRRR1CPaPa (Fig. 5.77) Pa\\R||R||R\\kR (Fig. 5.3d) Idem No. 3 Pa\\R||R||R (Fig. 5.2h) C||Pa||Pa (Fig. 5.40) 14. 2PaRRRR1PaRRR1CPaPat (Fig. 5.78) Pa\\R||R||R\\kR (Fig. 5.3d) Idem No. 3 Pa\\R||R||R (Fig. 5.2h) C||Pa||Pat (Fig. 5.4p) 15. 3PaPPaR1CPaPa (Fig. 5.79a) Pa||P\\Pa\\kR (Fig. 5.4a) The revolute joint between links 6D and 9 of G4 limb and the last revolute joints of limbs G1, G2 and G3 have superposed axes Pa||P\\Pa||R (Fig. 5.4d) C||Pa||Pa (Fig. 5.40) 16. 3PaPPaR1CPaPa (Fig. 5.79b) Pa\\P\\\\Pa\\kR (Fig. 5.4b) Idem No. 15 Pa\\P\\\\Pa||R (Fig. 5.4c) C||Pa||Pa (Fig. 5.40) 17. 3PaPPaR1CPaPat (Fig. 5.80a) Pa||P\\Pa\\kR (Fig. 5.4a) Idem No. 15 Pa||P\\Pa||R (Fig. 5.4d) C||Pa||Pat (Fig. 5.4p) 18. 3PaPPaR1CPaPat (Fig. 5.80b) Pa\\P\\\\Pa\\kR (Fig. 5.4b) Idem No. 15 Pa\\P\\\\Pa||R (Fig. 5.4c) C||Pa||Pat (Fig. 5.4p) (continued) 5.1 Fully-Parallel Topologies 535 Table 5.5 (continued) No. PM type Limb topology Connecting conditions 19. 3PaPaPR1CPaPa (Fig. 5.81) Pa\\Pa\\kP\\\\R (Fig. 5.4f) Idem No. 15 Pa\\Pa\\kP\\kR (Fig" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_35_0001563_9781119682684.ch1-Figure1.6-1.png", + "original_path": "designv11-35/openalex_figure/designv11_35_0001563_9781119682684.ch1-Figure1.6-1.png", + "caption": "Figure 1.6 The current coming out of page (a) counter clockwise \u222e B \u22c5 dl = \ud835\udf07oi. Here, the angle between B and dl is 0\u2218 (b) clockwise \u222e B \u22c5 dl = \u2212\ud835\udf07oi. Here, the angle is 180\u2218 (c) Contour abcda not enclosing the current \u222e B \u22c5 dl = 0 because fields along curves get cancelled and field at da and bc are zero since the angle is 90\u2218.", + "texts": [ + " (d) 45\u00b0 45\u00b0 B1 B2 B3 C Consider a long conductor carrying current i and, according to Biot-Sarvat law, sets up a magnetic field H or B around it at a radius of R. Ampere\u2019s law is derived using Biot-Savart law. The line integral of B around a closed path is denoted by \u222e B \u22c5 dl (1.32) To evaluate the product B \u22c5 dl consider a small length element dl on the circular path and sum the products for all elements over the closed circular path. Along this path, the vectors dl and B are parallel at each point (Figure 1.6c), so B \u22c5 dl=Bdl. And also, the magnitude of B is constant on this circle and is given by Eq. (1.25). Therefore, the sum of the products Bdl over the closed path, which is equivalent to the line integral of B \u22c5 dl is \u222e B \u22c5 dl = \u222e B\u2225 \u22c5 dl = B\u222e dl = \ud835\udf07oi 2\ud835\udf0bR (2\ud835\udf0bR) = \ud835\udf07oi where B\u2225 is the component of B which is parallel to dl or \u222e B \u22c5 dl = \ud835\udf07oi (1.33) In free space B=\ud835\udf07oH \u222e H \u22c5 dl = i (1.34) Concerning Figure 1.6, one can show that the enclosed current \u2018i\u2019 in Eq. (1.34) is equal to the integral of the current density J over any surface bounded by the closed path C. \u222e C H \u22c5 dl = \u222b S J \u22c5 dS (1.35) where J and S are the current density (A/m2) and surface area (m2) vectors respectively. The path need not always be circular. It can take any shape but it should enclose the current line. Equations (1.34) and (1.35) are applicable for static magnetic fields. It can be said that the line integral of H around a closed path is equal to the total current flowing through the surface bounded by that path" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_35_0000089_tmag.2021.3060767-Figure4-1.png", + "original_path": "designv11-35/openalex_figure/designv11_35_0000089_tmag.2021.3060767-Figure4-1.png", + "caption": "Fig. 4. Sections highlighting slot shape, geometric parameters, and no-load flux densities in different sections.", + "texts": [ + " 2 and 3, it is evident that compared with the open slot and open slot with the wedge, in the semi-open slot, the magnitude of the 28th, 68th, and 76th harmonics producing the eighth-order radial force harmonic producing flux harmonics mentioned in (6) and (7) has smaller magnitude. Thus, even though the open slot is desired for the large generators, for the proposed FCPMSG with hydraulic bearing, considering the radial force stability semi-open slot is considered. The slot shape of the optimum model along with the no-load flux density is shown in Fig. 4 and other geometric parameters are mentioned in Table I. The optimal slot opening length is determined considering the fundamental component of no-load-induced electromotive force (EMF) and the space required to manually wound the winding into the stator. For a PMSG, it is important to study the characteristic considering the load attached to its output terminal. For a PMSG, the load can either be uncontrolled passive impedance load, diode rectifier load, or controlled converter load. Thus, this section aims to generate the eighth-order radial force density harmonic distribution locus over the Id \u2013Iq plane under loading" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_35_0001419_978-94-007-7401-8-Figure3.12-1.png", + "original_path": "designv11-35/openalex_figure/designv11_35_0001419_978-94-007-7401-8-Figure3.12-1.png", + "caption": "Fig. 3.12 Overactuated PMs with coupled Sch\u00f6nflies motions of types 4PRRP (a) and 4RPRP (b) defined by MF \u00bc SF \u00bc 4; RF\u00f0 \u00de \u00bc v1; v2; v3;xd\u00f0 \u00de, TF = 0, NF = 6, limb topology P||R||R\\P (a) and R||P||R\\P (b)", + "texts": [ + "1h) The last revolute joints of the four limbs have parallel axes 10. 4PRPR (Fig. 3.8b) P\\R||P||R (Fig. 3.1p) Idem No. 9 11. 4RRPP (Fig. 3.9a) R||R||P\\P (Fig. 3.1j) Idem No. 2 12. 4RPRP (Fig. 3.9b) R||P||R\\P (Fig. 3.1k) Idem No. 2 13. 4RPPR (Fig. 3.10a) R||P\\P\\kR (Fig. 3.1l) Idem No. 2 14. 4RPPR (Fig. 3.10b) R\\P\\kP||R (Fig. 3.1m) Idem No. 2 15. 4PRRP (Fig. 3.11a) P\\R||R||P (Fig. 3.1n) The last prismatic joints of the four limbs have parallel directions 16. 4PRPR (Fig. 3.11b) P\\R||P||R (Fig. 3.1o) Idem No. 9 17. 4PRRP (Fig. 3.12a) P||R||R\\P (Fig. 3.1v) The first prismatic joints of the four limbs have parallel directions 18. 4RPRP (Fig. 3.12b) R||P||R\\P (Fig. 3.1q) Idem No. 2 19. 4PRPR (Fig. 3.13a) P||R\\P\\kR (Fig. 3.1r) Idem No. 1 20. 4RPPR (Fig. 3.13b) R||P\\P\\kR (Fig. 3.1s) Idem No. 2 21. 4RPPR (Fig. 3.14a) R\\P\\kP||R (Fig. 3.1t) Idem No. 2 22. 4RPRP (Fig. 3.14b) R\\P\\kR||P (Fig. 3.1u) Idem No. 2 23. 4PPPR (Fig. 3.15a) P\\P\\\\P\\kR (Fig. 3.1h0) Idem No. 9 24. 4PPPR (Fig. 3.15b) P\\P\\\\P\\\\R (Fig. 3.1w) Idem No. 9 25. 4PPPR (Fig. 3.16a) P\\P\\\\P||R (Fig. 3.1x) Idem No. 9 26. 4PPRP (Fig. 3.16b) P\\P||R\\P (Fig. 3.1y) The revolute joints of the four limbs have parallel axes 27" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_35_0000468_j.matpr.2021.05.248-Figure1-1.png", + "original_path": "designv11-35/openalex_figure/designv11_35_0000468_j.matpr.2021.05.248-Figure1-1.png", + "caption": "Fig. 1. Experimental setup of pin on disc tribometer.", + "texts": [ + " Each experiment was run three times in order to aver- Table 1 Process Parameter Levels. Parameters Unit Level I Level II Level III Load N 75 115 150 Speed rpm 800 1140 1440 Operating Time sec 1800 2700 3600 Table 2 L9-Orthogonal Array. Exp. no. Load (N) Speed (rpm) Time (sec) 1 75 800 1800 2 75 1140 2700 3 75 1440 3600 4 115 800 2700 5 115 1140 3600 6 115 1440 1800 7 150 800 3600 8 150 1140 1800 9 150 1440 2700 age the results and reduce the pure experimental error. Pin on Disc Tribometer follows the ASTM G 99 standard. Fig. 1 depicts a pin on a disc tribometer. The load, speed and time was set according to design of experiment on pin on disc tribometer. The results of response parameters was documented for further analysis.The information from the experimental observation of tribological parameters is presented in Table 3. It was used to determine which set of experiments had the least amount of frictional force and wear.Grey Relational Analysis is the method for evaluating the degree of connection between series of experiments and optimizing process parameters with multiple outputs using grey relational ranking" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_35_0001419_978-94-007-7401-8-Figure2.61-1.png", + "original_path": "designv11-35/openalex_figure/designv11_35_0001419_978-94-007-7401-8-Figure2.61-1.png", + "caption": "Fig. 2.61 4PaRPaP-type fully-parallel PM with coupled Sch\u00f6nflies motions defined by MF \u00bc SF \u00bc 4; RF\u00f0 \u00de \u00bc \u00f0v1; v2; v3;xd\u00de, TF = 0, NF = 30, limb topology Pa\\R\\Pa\\kR", + "texts": [ + "22n) The axes of the revolute joints connecting links 7 and 8 of the four limbs are parallel 42. 4PaPaPR (Fig. 2.54) Pa||Pa||P\\R (Fig. 2.22o) Idem No. 4 43. 4PRPaPa (Fig. 2.55) P\\R\\Pa||Pa (Fig. 2.22p) Idem No. 5 44. 4PPaRPa (Fig. 2.56) P||Pa\\R\\Pa (Fig. 2.22q) Idem No. 33 45. 4PPaPaR (Fig. 2.57) P||Pa||Pa\\R (Fig. 2.22r) Idem No. 4 46. 4PaPPaR (Fig. 2.58) Pa||P||Pa\\R (Fig. 2.22s) Idem No. 4 47. 4RPPaPa (Fig. 2.59) R\\P||Pa||Pa (Fig. 2.22t) Idem No. 15 48. 4PaPPaR (Fig. 2.60) Pa\\P\\\\Pa\\kR (Fig. 2.22u) Idem No. 4 49. 4PaRPaP (Fig. 2.61) Pa\\R\\Pa\\kR (Fig. 2.22v) Idem No. 34 50. 4PaRPPa (Fig. 2.62) Pa\\R||P\\Pa (Fig. 2.22w) Idem No. 34 51. 4PaPRPa (Fig. 2.63) Pa\\P||R\\Pa (Fig. 2.22x) Idem No. 33 52. 4PaPaPR (Fig. 2.64) Pa\\Pa\\\\P||R (Fig. 2.22y) Idem No. 4 53. 4PaPaRP (Fig. 2.65) Pa\\Pa\\\\R||P (Fig. 2.22z) Idem No. 41 54. 4RPaPaP (Fig. 2.66) R\\Pa\\\\Pa\\\\P (Fig. 2.22a0) Idem No. 15 55. 4RPaPPa (Fig. 2.67) R\\Pa\\kP\\\\Pa (Fig. 2.22b0) Idem No. 15 56. 4RPPaPa (Fig. 2.68) R||P\\Pa\\\\Pa (Fig. 2.22c0) Idem No. 15 57. 4PRPaPa (Fig. 2.69) P||R\\Pa\\\\Pa (Fig" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_35_0001419_978-94-007-7401-8-Figure6.41-1.png", + "original_path": "designv11-35/openalex_figure/designv11_35_0001419_978-94-007-7401-8-Figure6.41-1.png", + "caption": "Fig. 6.41 3PPaPaR-1RPPaPat-type maximaly regular fully-parallel PM with Sch\u00f6nflies motions defined by MF = SF = 4, (RF) = (v1; v2; v3;xb), TF = 0, NF = 30, limb topology P||Pa||Pa\\R, P||Pa||Pa||R and R||P||Pa||Pat", + "texts": [ + " 2PPaRRR-1PPaRR-1RPPaPat (Fig. 6.37) P||Pa||R||R\\R (Fig. 4.9a) P||Pa||R||R (Fig. 4.8g) R||P||Pa||Pat (Fig. 5.4n) Idem no. 9 17. 3PPaPaR-1RPaPaP (Fig. 6.38) P||Pa||Pa\\R (Fig. 4.10a) P||Pa||Pa||R (Fig. 4.10b) R||Pa||Pa||P (Fig. 5.4k) Idem no. 1 18. 3PPaPaR-1RPaPatP (Fig. 6.39) P||Pa||Pa\\R (Fig. 4.10a) P||Pa||Pa||R (Fig. 4.10b) R||Pa||Pat||P (Fig. 5.4l) Idem no. 1 19. 3PPaPaR-1RPPaPa (Fig. 6.40) P||Pa||Pa\\R (Fig. 4.10a) P||Pa||Pa||R (Fig. 4.10b) R||P||Pa||Pa (Fig. 5.4m) Idem no. 9 20. 3PPaPaR-1RPaPatP (Fig. 6.41) P||Pa||Pa\\R (Fig. 4.10a) P||Pa||Pa||R (Fig. 4.10b) R||P||Pa||Pat (Fig. 5.4n) Idem no. 9 Table 6.14 Limb topology and connecting conditions of the fully-parallel solutions with no idle mobilities presented in Figs. 6.42, 6.43, 6.44, 6.45, 6.46 No. PM type Limb topology Connecting conditions 1. 3PPPaR-1CPaPa (Fig. 6.42a) P\\P\\||Pa\\||R (Fig. 4.8b) P\\P\\||Pa||R, (Fig. 4.8c) C||Pa||Pa (Fig. 5.4o) The last revolute joints of limbs G1, G2 and G3 and the revolute joint between links 6D and 8 have superposed axes" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_35_0000370_j.sysconle.2021.104931-Figure2-1.png", + "original_path": "designv11-35/openalex_figure/designv11_35_0000370_j.sysconle.2021.104931-Figure2-1.png", + "caption": "Fig. 2. Extended mass on a car system. Source: The original figure is taken from [20] and edited for the purpose of the present article.", + "texts": [ + " The two cars are coupled via a spring\u2013damper combination with characteristics K2 and D2, resp. On the second car a ramp with constant angle 0 < \u03b1 < \u03c0/2 is mounted, on which a third mass m3 is lying and is coupled to the second car via a spring with characteristic K3, and a damping with characteristic D3. The first car is driven by a force u1, and the second car individually is driven by a force u2. We measure the horizontal position of the first and the second car. The situation is depicted in Fig. 2. For convenience we assume the constant force on m3 due to ravity, namely m3g sin(\u03b1), where g is the gravitational constant, to be compensated via a linear coordinate transformation, such that K3(0) = 0. Then, according to [20, Sec. 4.2] with s := s1, s2, s3)\u22a4 \u2208 R3 the equations of motion for that system are iven by[m1 + m2 + m3 m2 + m3 m3 cos(\u03b1) m2 + m3 m2 + m3 m3 cos(\u03b1) m3 cos(\u03b1) m3 cos(\u03b1) m3 ] =:M (s\u03081(t) s\u03082(t) s\u03083(t) ) = ( 0 \u2212K2(s(t)) \u2212 D2(s\u0307(t)) \u2212K3(s(t)) \u2212 D3(s\u0307(t)) ) + [1 0 0 1 0 0 ] =:B ( u1(t) u2(t) ) , y(t) = ( y1(t) y2(t) ) = ( s1(t) s1(t) + s2(t) ) = [ 1 0 0 1 1 0 ] =:H (s1(t) s2(t) s3(t) ) " + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_35_0000750_j.matpr.2021.07.468-Figure3-1.png", + "original_path": "designv11-35/openalex_figure/designv11_35_0000750_j.matpr.2021.07.468-Figure3-1.png", + "caption": "Fig 3. Schematic of Ink-based 3D Printing Technique [25].", + "texts": [ + " It was made out of stretchable polyurethane TPU92A-1 and capable of performing bending, rotating, and sensing as well as bi-directional actuation [23]. Roppenecker et al. used a polyamide-based material called PA 2200 (PA 12, Nylon) with the SLS technique to achieve a multi-arm snake-like robot. They fabricated several soft objects which were based on flexure hinges capable of performing endoscopic operations. These structures were capable of bearing ~ 800 g of weight [24]. In the inkjet printing process (Fig 3), the print heads selectively jet the liquid resins through narrow nozzles onto the platform. The liquid resins are photopolymer materials; thus, they get solidified as soon as they come into contact with UV lights. Several 3D printing methods use the same principle of droplet-based printing, such as powder bed, direct inkjet, and hot metal printing. These printing methods share similarities with both SLA and SLS processes, mainly being light-based and photocurable methods. The small difference in inkjet printing is that the inkjet head is not stationary" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_35_0000213_j.triboint.2021.106993-Figure1-1.png", + "original_path": "designv11-35/openalex_figure/designv11_35_0000213_j.triboint.2021.106993-Figure1-1.png", + "caption": "Fig. 1. Schematic of a TPJB System with three nozzles; (a) Overview, (b) Step (green) and Pocket (yellow) included on a Pad. (For interpretation of the references to color in this figure legend, the reader is referred to the Web version of this article.)", + "texts": [ + " The phase change from liquid to vapor in the pockets, referred to as cavitation, decreases the shear stress at the journal surface due to the low gas viscosity. The cavitation reduces power loss proportional to the shear stress reduction. The static and dynamic performances of a conventional and the novel TPJB are compared to show the benefits of the novel TPJB. These simulations are performed using a high fidelity TEHD-CFD model to accurately include the pocket and step features. This section explains the approach to reduce power loss, as validated by TEHD-CFD simulations. The TPJB geometry including all computational domains is depicted in Fig. 1(a). The TPJB shown has five pads, with directed lubrication provided by three nozzles. This arrangement could be generalized for different numbers of pads and nozzles. Direct lubrication lowers power loss significantly, and it is included to show that the proposed TPJB changes can even further reduce power loss. The proposed method is simply to include a pocket and a step in each upper J. Yang and A. Palazzolo Tribology International 159 (2021) 106993 pad, as illustrated in Fig. 1(b). The TPJB is used to support the shaft, and to provide desired stiffness and damping for vibration control. Oil flows from the three nozzles to both side seals and to the thin fluid-film between the journal and the downstream pad. The pressure and shear stress acting on the journal and pad are illustrated in Fig. 2(a). These cause the lifting force and drag force on the shaft, respectively. Important geometric parameters include the bearing clearance (Cl,b), pad clearance (Cl,p), and preload (mpr), as shown in Fig" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_35_0001672_s0368393100116256-Figure15-1.png", + "original_path": "designv11-35/openalex_figure/designv11_35_0001672_s0368393100116256-Figure15-1.png", + "caption": "Fig. 15. Derivation of the transfer function of the aero plane considered as a simple damped oscillatory system.", + "texts": [ + " Before assessing the total tolerable lag in the case of a controlled aeroplane, we require 423> F. W. MEREDITH to know the lag of the aeroplane. With the simplifying assumption, indicated earlier, that the short period yawing oscillation may be considered as the motion of a weather vane with, of course, means representing the rudder for forcing it, the relationships are shown in Fig. 14. In Fig. 14 the control moment is derived as a leading vector from the displacement vector. The locus of the control vector as w is varied is clearly a parabola. Figure 15 shows, in the lower full line loop, the transfer function of the aeroplane in response to the rudder and how it may be simply derived from the parabola of Fig. 14. To invert the ratio, control moment: dis placement, we must derive the reciprocal vector\u2014the upper dotted loop and reflect this curve to transfer phase lead\" to phase lag. I have reproduced this construction to facilitate judgment, by inspection, of the effect of reducing the inherent damping of the aeroplane. The relevant portion of the transfer function locus is the left-hand portion, between resonance and cos a. + cos p From this expression we can obtain d8/d(/> by a method precisely similar to tha t employed for the spheric chain in which a=y, j8 = 8. I t will be of greater interest, however, to treat the problem of the skew isogram also by spherical trigonometry. Draw from the centre of the unit sphere lines parallel to the four hinge-lines and join their intersections with the surface by great circle arcs ; these will form a \" spherical parallelogram \" with the angles 9 and (f> as shown in Fig. 11. Join DB by a great circle arc. Then LADB = L DEC. Then, by applying I I I to the triangle ABJ), tan UADB + DBA) = \u00b0 0 S ' f | a ~ o l cot DAB. 2 ' cos-\u00a3(a+)3) i.e., tan ^0 tan J^ = cos |-(a +/3)/cos | ( a -/?) (11) https://doi.org/10.2307/3606015 11 Feb 2020 at 10:34:09, subject to the Cambridge Core terms of use, available at https://www.cambridge.org/core/terms. Downloaded from https://www.cambridge.org/core. Conrad Grebel University College, University of Waterloo, on MECHANISMS I N T H R E E DIMENSIONS 17 Further, from IV, sin (ADB + DBA) _ cos a + cos jS sin DAB 1+oosDB Substituting for cos DB, by I, sin# 1 + {cos a cos |3+sin a sin j8 cos (180-0)} d,9 s in cos a + cos /3 d Hence - ddjd = sec \\x" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_35_0001419_978-94-007-7401-8-Figure5.9-1.png", + "original_path": "designv11-35/openalex_figure/designv11_35_0001419_978-94-007-7401-8-Figure5.9-1.png", + "caption": "Fig. 5.9 2PaRPRR-1PaRPR-1RPPP-type fully-parallel PMs with uncoupled Sch\u00f6nflies motions defined by MF = SF = 4, RF\u00f0 \u00de \u00bc \u00f0v1; v2; v3;xb\u00de, TF = 0, NF = 13, limb topology Pa\\R\\P\\kR\\R, Pa\\R\\P\\kR and R\\P\\\\P\\\\P", + "texts": [], + "surrounding_texts": [ + "442 5 Topologies with Uncoupled Sch\u00f6nflies Motions", + "5.1 Fully-Parallel Topologies 443", + "444 5 Topologies with Uncoupled Sch\u00f6nflies Motions" + ] + }, + { + "image_filename": "designv11_35_0001385_978-981-15-6095-8-Figure2.6-1.png", + "original_path": "designv11-35/openalex_figure/designv11_35_0001385_978-981-15-6095-8-Figure2.6-1.png", + "caption": "Fig. 2.6 Stresses acting on the surface of the geocell. Sourced from Hegde and Sitharam (2015)", + "texts": [ + " Hence, the confinement effect of the geocell is based on three main mechanisms: active earth pressure within loaded cell, passive earth pressure in the adjacent cells and the Hoop stress within the cell wall (Emersleben and Meyer 2008, 2015). The different stresses developed in the geocell walls under the action of compression loads are shown in Fig. 2.5. The Hoop stress will lead to the deformation of the cell wall. The cell wall deformations can be measured in terms of Hoop strains and the volumetric strains. Hegde and Sitharam (2015) developed the expression for the Hoop stress, Hoop strain and the volumetric strains in the geocell surface using the theory of thin cylinder formulations. Figure 2.6 represents the stresses acting on the surface of the deformed geocell as reported by Hegde and Sitharam (2015). The only half portion of the geocell was considered by the researchers in the formulation due to the symmetry. P is the active earth pressure exerted by the infill soil on the geocell wall. Researchers have considered the a small element of length, l on the periphery of the geocell, making an angle d\u03b8 with the center to obtain the expression for Hoop stress (\u03c3 h),Hoop strain (\u03b5h) and volumetric strain (\u03b5l)" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_35_0001419_978-94-007-7401-8-Figure3.84-1.png", + "original_path": "designv11-35/openalex_figure/designv11_35_0001419_978-94-007-7401-8-Figure3.84-1.png", + "caption": "Fig. 3.84 Overactuated PMs with coupled Sch\u00f6nflies motions of types 4CRPa (a) and 4PCPa (b) defined by MF \u00bc SF \u00bc 4; RF\u00f0 \u00de \u00bc v1; v2; v3;xd\u00f0 \u00de, TF = 0, NF = 18, limb topology C||R||Pa (a) and P\\C||Pa (b)", + "texts": [ + " 3.50p0) Idem No. 37 44. 4PaRRP (Fig. 3.80a) Pa\\R||R\\P (Fig. 3.50q0) Idem No. 23 45. 4RPaRR (Fig. 3.80b) R\\Pa\\kR||R (Fig. 3.50s0) Idem No. 37 46. 4PaRRR (Fig. 3.81) Pa\\R||R||R (Fig. 3.50r0) Idem No. 37 47. 4PaPRP (Fig. 3.82a) Pa||P\\R\\P (Fig. 3.50t0) Idem No. 23 48. 4PaPRP (Fig. 3.82b) Pa\\P\\\\R\\P (Fig. 3.50u0) Idem No. 23 49. 4CPaP (Fig. 3.83a) C||Pa\\P (Fig. 3.50v0) The cylindrical joints of the four limbs have parallel directions 50. 4PaCP (Fig. 3.83b) Pa||C\\P (Fig. 3.50w0) Idem No. 49 51. 4CRPa (Fig. 3.84a) C||R||Pa (Fig. 3.50x0) Idem No. 49 52. 4PCPa (Fig. 3.84b) P\\C||Pa (Fig. 3.50y0) Idem No. 49 53. 4PPaC (Fig. 3.85a) P\\Pa||C (Fig. 3.50z0) Idem No. 49 54. 4RCPa (Fig. 3.85b) R||C||Pa (Fig. 3.50z01) Idem No. 49 55. 4CPPa (Fig. 3.86a) C\\P\\kPa (Fig. 3.50a0 0) Idem No. 49 56. 4PaPC (Fig. 3.86b) Pa\\P\\kC (Fig. 3.50b0 0) Idem No. 49 57. 4PaRC (Fig. 3.87a) Pa||R||C (Fig. 3.50c0 0) Idem No. 49 58. 4PaCR (Fig. 3.87b) Pa\\C\\P (Fig. 3.50e0 0) Idem No. 49 59. 4PaCR (Fig. 3.88) Pa||C||R (Fig. 3.50d0 0) Idem No. 49 Table 3.6 Limb topology and connecting conditions of the overactuated solutions with no idle mobilities presented in Figs" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_35_0001419_978-94-007-7401-8-Figure5.53-1.png", + "original_path": "designv11-35/openalex_figure/designv11_35_0001419_978-94-007-7401-8-Figure5.53-1.png", + "caption": "Fig. 5.53 2PaPaRRR-1PaPaRR-1RPaPatP-type fully-parallel PM with uncoupled Sch\u00f6nflies motions defined by MF = SF = 4, RF\u00f0 \u00de \u00bc \u00f0v1; v2; v3;xb\u00de, TF = 0, NF = 28, limb topology Pa\\Pa||R||R\\\\R, Pa\\Pa||R||R and R||Pa||Pat||P", + "texts": [ + "50) Pa\\Pa\\\\P\\\\R (Fig. 5.4g) Idem No. 3 Pa\\Pa\\\\P\\kR (Fig. 5.4e) R||P||Pa||Pat (Fig. 5.4n) 17. 2PaPaRRR-1PaPaRR1RPaPaP (Fig. 5.51) Pa\\Pa||R||R\\\\R (Fig. 5.5a) Idem No. 1 Pa\\Pa||R||R (Fig. 5.4i) R||Pa||Pa||P (Fig. 5.4k) (continued) 532 5 Topologies with Uncoupled Sch\u00f6nflies Motions Table 5.4 (continued) No. PM type Limb topology Connecting conditions 18. 2PaPaRRR-1PaPaRR1RPaPaP (Fig. 5.52) Pa\\Pa||R||R\\kR (Fig. 5.5b) Idem No. 1 Pa\\Pa||R||R (Fig. 5.4i) R||Pa||Pa||P (Fig. 5.4k) 19. 2PaPaRRR-1PaPaRR1RPaPatP (Fig. 5.53) Pa\\Pa||R||R\\\\R (Fig. 5.5a) Idem No. 1 Pa\\Pa||R||R (Fig. 5.4i) R||Pa||Pat||P (Fig. 5.4l) 20. 2PaPaRRR-1PaPaRR1RPaPatP (Fig. 5.54) Pa\\Pa||R||R\\kR (Fig. 5.5b) Idem No. 1 Pa\\Pa||R||R (Fig. 5.4i) R||Pa||Pat||P (Fig. 5.4l) 21. 2PaPaRRR-1PaPaRR1RPPaPa (Fig. 5.55) Pa\\Pa||R||R\\\\R (Fig. 5.5a) The revolute joint between links 7D and 10 of G4 limb and the last revolute joints of limbs G1, G2 and G3 have superposed axes Pa\\Pa||R||R (Fig. 5.4i) R||P||Pa||Pa (Fig. 5.4m) 22. 2PaPaRRR-1PaPaRR1RPPaPa (Fig. 5.56) Pa\\Pa||R||R\\kR (Fig" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_35_0001419_978-94-007-7401-8-Figure3.87-1.png", + "original_path": "designv11-35/openalex_figure/designv11_35_0001419_978-94-007-7401-8-Figure3.87-1.png", + "caption": "Fig. 3.87 Overactuated PMs with coupled Sch\u00f6nflies motions of types 4PaRC (a) and 4PaCR (b) defined by MF \u00bc SF \u00bc 4; RF\u00f0 \u00de \u00bc v1; v2; v3;xd\u00f0 \u00de, TF = 0, NF = 18, limb topology Pa||R||C (a) and Pa\\C\\P (b)", + "texts": [ + "83a) C||Pa\\P (Fig. 3.50v0) The cylindrical joints of the four limbs have parallel directions 50. 4PaCP (Fig. 3.83b) Pa||C\\P (Fig. 3.50w0) Idem No. 49 51. 4CRPa (Fig. 3.84a) C||R||Pa (Fig. 3.50x0) Idem No. 49 52. 4PCPa (Fig. 3.84b) P\\C||Pa (Fig. 3.50y0) Idem No. 49 53. 4PPaC (Fig. 3.85a) P\\Pa||C (Fig. 3.50z0) Idem No. 49 54. 4RCPa (Fig. 3.85b) R||C||Pa (Fig. 3.50z01) Idem No. 49 55. 4CPPa (Fig. 3.86a) C\\P\\kPa (Fig. 3.50a0 0) Idem No. 49 56. 4PaPC (Fig. 3.86b) Pa\\P\\kC (Fig. 3.50b0 0) Idem No. 49 57. 4PaRC (Fig. 3.87a) Pa||R||C (Fig. 3.50c0 0) Idem No. 49 58. 4PaCR (Fig. 3.87b) Pa\\C\\P (Fig. 3.50e0 0) Idem No. 49 59. 4PaCR (Fig. 3.88) Pa||C||R (Fig. 3.50d0 0) Idem No. 49 Table 3.6 Limb topology and connecting conditions of the overactuated solutions with no idle mobilities presented in Figs. 3.89, 3.90, 3.91, 3.92, 3.93, 3.94, 3.95, 3.96, 3.97, 3.98, 3.99, 3.100, 3.101, 3.102, 3.103, 3.104, 3.105, 3.106, 3.107, 3.108, 3.109, 3.110, 3.111, 3.112, 3.113, 3.114, 3.115, 3.116, 3.117, 3.118, 3.119, 3.120, 3.121, 3.122, 3.123, 3.124, 3.125, 3.126, 3.127, 3.128, 3.129, 3.130, 3" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_35_0001419_978-94-007-7401-8-Figure2.108-1.png", + "original_path": "designv11-35/openalex_figure/designv11_35_0001419_978-94-007-7401-8-Figure2.108-1.png", + "caption": "Fig. 2.108 Fully-parallel PMs with coupled Sch\u00f6nflies motions of types 4RPaRRR (a) and 4PaRRPR (b) defined by MF \u00bc SF \u00bc 4; RF\u00f0 \u00de \u00bc \u00f0v1; v2; v3;xd\u00de, TF = 0, NF = 14, limb topology R\\Pa\\kR\\R||R (a) and Pa||R\\R\\P\\kR (b)", + "texts": [], + "surrounding_texts": [ + "162 2 Fully-Parallel Topologies with Coupled Sch\u00f6nflies Motions", + "2.2 Topologies with Complex Limbs 163", + "164 2 Fully-Parallel Topologies with Coupled Sch\u00f6nflies Motions" + ] + }, + { + "image_filename": "designv11_35_0001419_978-94-007-7401-8-Figure3.26-1.png", + "original_path": "designv11-35/openalex_figure/designv11_35_0001419_978-94-007-7401-8-Figure3.26-1.png", + "caption": "Fig. 3.26 Overactuated PMs with coupled Sch\u00f6nflies motions of types 4RPC (a) and 4CRR (b) defined by MF \u00bc SF \u00bc 4; RF\u00f0 \u00de \u00bc v1; v2; v3;xd\u00f0 \u00de, TF = 0, NF = 6, limb topology R\\P\\kC (a) and C||R||R (b)", + "texts": [ + "21b) R||C||R (Fig. 3.2b) Idem No. 1 3. 4PCP (Fig. 3.22a) P\\C\\P (Fig. 3.2c) Idem No. 1 4. 4CPP (Fig. 3.22b) C\\P\\\\P (Fig. 3.2d) Idem No. 1 5. 4PPC (Fig. 3.22c) P\\P\\\\C (Fig. 3.2e) Idem No. 1 6. 4PCR (Fig. 3.23a) P\\C||R (Fig. 3.2f) Idem No. 1 7. 4RCP (Fig. 3.23b) R||C\\P (Fig. 3.2g) Idem No. 1 8. 4RPC (Fig. 3.24a) R\\P\\kC (Fig. 3.2h) Idem No. 1 9. 4PRC (Fig. 3.24b) P\\R||C (Fig. 3.2i) Idem No. 1 10. 4CRP (Fig. 3.25a) C||R\\P (Fig. 3.2j) Idem No. 1 11. 4CPR (Fig. 3.25b) C\\P\\kR (Fig. 3.2k) Idem No. 1 12. 4RPC (Fig. 3.26a) R\\P\\kC (Fig. 3.2i) Idem No. 1 13. 4CRR (Fig. 3.26b) C||R||R (Fig. 3.2m) Idem No. 1 14. 4RRRRR (Fig. 3.27a) R\\R||R\\R||R (Fig. 3.3a) The first and the last revolute joints of the four limbs have parallel axes 15. 4RRRRR (Fig. 3.27b) R||R\\R||R\\R (Fig. 3.3b) Idem No. 14 16. 4RRRRR (Fig. 3.28a) R||R||R\\R||R (Fig. 3.3c) The first revolute joints of the four limbs have parallel axes 17. 4RRRRR (Fig. 3.28b) R||R\\R||R||R (Fig. 3.3d) The last revolute joints of the four limbs have parallel axes 18. 4PRRRR (Fig. 3.29a) P||R||R\\R||R (Fig. 3.3e) Idem No. 17 19" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_35_0001419_978-94-007-7401-8-Figure3.64-1.png", + "original_path": "designv11-35/openalex_figure/designv11_35_0001419_978-94-007-7401-8-Figure3.64-1.png", + "caption": "Fig. 3.64 4PPaRP-type overactuated PM with coupled Sch\u00f6nflies motions defined by MF \u00bc SF \u00bc 4; RF\u00f0 \u00de \u00bc v1; v2; v3;xd\u00f0 \u00de, TF = 0, NF = 18, limb topology P\\Pa||R||P", + "texts": [ + "50k) Idem No. 1 12. 4RPaPP (Fig. 3.60a) R||Pa\\P\\kP (Fig. 3.50l) Idem No. 1 13. 4RPPPa (Fig. 3.60b) R||P\\P\\kPa (Fig. 3.50m) Idem No. 1 14. 4PaRPP (Fig. 3.61a) Pa||R||P\\P (Fig. 3.50n) Idem No. 1 15. 4RPaPP (Fig. 3.61b) R||Pa||P\\P (Fig. 3.50o) Idem No. 1 16. 4RPPaP (Fig. 3.62a) R\\P\\kPa||P (Fig. 3.50p) Idem No. 1 17. 4PPPaR (Fig. 3.62b) P\\P\\kPa||R (Fig. 3.50q) Idem No. 1 18. 4PPPaR (Fig. 3.63a) P\\P||Pa||R (Fig. 3.50r) Idem No. 1 19. 4PPaRR (Fig. 3.63b) P||Pa||R||R (Fig. 3.50s) Idem No. 1 20. 4PPaRP (Fig. 3.64) P\\Pa||R||P (Fig. 3.50t) Idem No. 1 21. 4PPaPR (Fig. 3.65) P\\Pa||P||R (Fig. 3.50u) Idem No. 1 22. 4PPaPR (Fig. 3.66a) P||Pa\\P\\kR (Fig. 3.50v) Idem No. 1 23. 4PPaRP (Fig. 3.66b) P||Pa\\R\\P (Fig. 3.50i0) The before last joints of the four limbs have parallel axes 24. 4RPPaP (Fig. 3.67) R||P||Pa\\P (Fig. 3.50w) Idem No. 1 25. 4RPRPa (Fig. 3.68) R||P||R||Pa (Fig. 3.50x) Idem No. 1 26. 4PaPPR (Fig. 3.69a) Pa||P\\P\\kR (Fig. 3.50y) Idem No. 1 27. 4PaPPR (Fig. 3.69b) Pa\\P\\kP||R (Fig. 3.50z) Idem No. 1 28. 4PaPRP (Fig" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_35_0001419_978-94-007-7401-8-Figure3.74-1.png", + "original_path": "designv11-35/openalex_figure/designv11_35_0001419_978-94-007-7401-8-Figure3.74-1.png", + "caption": "Fig. 3.74 4PRPPa-type overactuated PMs with coupled Sch\u00f6nflies motions defined by MF \u00bc SF \u00bc 4; RF\u00f0 \u00de \u00bc v1; v2; v3;xd\u00f0 \u00de, TF = 0, NF = 18, limb topology P\\R\\P||Pa (a) and P\\R\\P\\Pa (b)", + "texts": [ + "50w) Idem No. 1 25. 4RPRPa (Fig. 3.68) R||P||R||Pa (Fig. 3.50x) Idem No. 1 26. 4PaPPR (Fig. 3.69a) Pa||P\\P\\kR (Fig. 3.50y) Idem No. 1 27. 4PaPPR (Fig. 3.69b) Pa\\P\\kP||R (Fig. 3.50z) Idem No. 1 28. 4PaPRP (Fig. 3.70) Pa\\P\\kR||P (Fig. 3.50z1) Idem No. 1 29. 4PaRRP (Fig. 3.71) Pa||R||R||P (Fig. 3.50a0) Idem No. 1 30. 4PaRPR (Fig. 3.72) Pa||R||P||R (Fig. 3.50b0) Idem No. 1 31. 4PaPRP (Fig. 3.73a) Pa||P||R\\P (Fig. 3.50c0) Idem No. 1 32. 4PaRPP (Fig. 3.73b) Pa||R\\P\\kP (Fig. 3.50d0) Idem No. 1 33. 4PRPPa (Fig. 3.74a) P\\R\\P||Pa (Fig. 3.50e0) The second joints of the four limbs have parallel axes 34. 4PRPPa (Fig. 3.74b) P\\R\\P\\Pa (Fig. 3.50f0) Idem No. 33 35. 4PRPaP (Fig. 3.75a) P\\R\\Pa\\\\P (Fig. 3.50g0) Idem No. 33 36. 4PPRPa (Fig. 3.75b) P\\P\\\\R\\Pa (Fig. 3.50h0) The third joints of the four limbs have parallel axes 37. 4PPPaR (Fig. 3.76a) P\\P||Pa\\\\R (Fig. 3.50j0) The last joints of the four limbs have parallel axes (continued) 3.2 Topologies with Complex Limbs 359 Table 3.5 (continued) No. PM type Limb topology Connecting conditions 38. 4PPPaR (Fig. 3.76b) P\\P\\kPa\\\\R (Fig. 3.50k0) Idem No. 37 39. 4PaPPR (Fig. 3" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_35_0000537_sibcon50419.2021.9438890-Figure1-1.png", + "original_path": "designv11-35/openalex_figure/designv11_35_0000537_sibcon50419.2021.9438890-Figure1-1.png", + "caption": "Fig. 1. TurtleBot Waffle Pi", + "texts": [ + " Thus, it is not considered to be light-weight solution and requires complex setting up process. Ros Control Center is a universal tool for controlling robots running ROS. It runs in the browser using a WebSocket connection and roslibjs from RobotWebTools [26]. Provided application has the following features: nodes, topics and service monitoring, publishing messages, controlling ROS parameters. This solution is Angular-based and focused on control rather than monitoring. The mobile robot TurtleBot3 Waffle Pi (Fig. 1) is a differential drive robot designed as a collaboration project by Open Robotics and ROBOTIS It is a relatively small robot for education and research purposes. TurtleBot3 Waffle is equipped with the following hardware: \u2022 360\u00b0 LiDAR (used to emit laser signal, which reflects of nearby obstacles) \u2022 Raspberry Pi (single-board computer, commonly used for applications that do not require high processing power) \u2022 optical camera \u2022 OpenCR (Connects to Raspberry Pi and allows it to control the motors and sensors) \u2022 11" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_35_0000174_j.matpr.2021.01.640-Figure8-1.png", + "original_path": "designv11-35/openalex_figure/designv11_35_0000174_j.matpr.2021.01.640-Figure8-1.png", + "caption": "Fig. 8. Rotational velocity applied on the drive shaft.", + "texts": [ + " Drive shaft model is divided into finite no of convenient sub elements and each element corners are joined with adjacent element for the purpose of finding bonding strength between sub elements. The fixed support is suitable for the drive shaft .Rotational velocity 650 rad\\s is applied on the drive shaft. The moment is applied 350 N.m on the drive shaft. In this analysis fixed dynamic is suitable are shown in Fig. 7. The rotational velocity 650 rad\\s is applied on the drive shaft. The drive shaft rotations are 650,0, 2.063 rad\\s. The locations are 795.07, 7.3163 mm are shown in Fig. 8. The moment load does not cause rotation, the Moment is applied on the drive shaft is 350 N.m as shown in Fig. 9 in order to omit more bending stress. When composite drive shaft breaks, it is divided into fine fibers that do not have any danger for the driver. They have a less specific modulus and less damping capacity. So that conventional drive shaft is replaced with composite materials Fig. 10. Selection of materials are an important parameter for design of any machine element. Our objective is to select suitable material than conventional steel material and preferred metal matrix composite materials consist of Aluminium, Titanium, Vanadium alloy" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_35_0000066_s10846-021-01335-z-Figure1-1.png", + "original_path": "designv11-35/openalex_figure/designv11_35_0000066_s10846-021-01335-z-Figure1-1.png", + "caption": "Fig. 1 Omnidirectional walker platform with vertical linear drive for hip and trunk support. a CAD design with principal axes. b Planar kinematic model of the omnidirectional platform with two ASOC sets. Inset shows the ASOC model employed in our prototype. Passive casters are omitted for clarity", + "texts": [ + " The first objective of the walker control is to produce a virtual cancellation of the walker mechanism\u2019s dynamics. Specifically, we use admittance control to shape the walker\u2019s mass and moment of inertia, friction and damping, in such a way that the burden of walking in the device is reduced as much as possible. In this section we show that, in spite of the system\u2019s nonholonomic characteristics, it is possible to implement a control capable of tracking a reference trajectory generated by the admittance model with global asymptotic stability. We will use the omnidirectional platform design of Fig. 1a as a motivating example; details of a physical prototype based on this design are provided in Section 5. However, we will keep the control formulation general enough to be applicable to any system in which omnidirectional motion is sought from a nonholonomic mechanism. A list of the key variables employed in the formulation of the control is provided in Table 1, along with their descriptions. Consider a mobile robot constrained to the horizontal plane with an n-dimensional configuration space and generalized coordinates z = [z1 \u00b7 \u00b7 \u00b7 zn]T , subject to m nonholonomic constraints", + " The dynamics of this robot are described by M(z)z\u0308 + C(z, z\u0307)z\u0307 = J(z)T \u03c4 a + A(z)\u03bb (1) where M(z) \u2208 R n\u00d7n is the inertia matrix, C(z, z\u0307) \u2208 R n\u00d7n is the centrifugal and Coriolis matrix, A(z) \u2208 R n\u00d7m is the constraint matrix, \u03bb \u2208 R m\u00d71 is a vector of Lagrange multipliers representing the constraint forces, \u03c4 a \u2208 R p\u00d71 is the actuators\u2019 input and J(z) \u2208 R p\u00d7n is the input transformation matrix. For our walker, J(z) is given by J(z) = [ \u03c9Jc(z) 0 0 ] (2) where \u03c9Jc(z) is the Jacobian relating the end effector velocity v to the wheels\u2019 velocities \u03c9a . An early CAD model of our walker design is shown in Fig. 1a. The corresponding kinematic model, featuring two ASOC sets, is shown in Fig. 1b. Our aim is to control the trajectory of the end effector pose x and velocity v, given respectively by x = [xc yc \u03c6c]T v = x\u0307 (3) The full pose of the walker is z = [xT \u03b11 \u03b12]T (4) where \u03b11 and \u03b12 are the angles of the unactuated steering axes. In Eq. 1, A(z) is the non-slip constraint on the walker wheels; this is a classic nonholonomic constraint obeying A(z)T z\u0307 = 0 (5) If we now define S(z) \u2208 R n\u00d7(n\u2212m) as the annihilator of A(z), i.e. a full-rank matrix satisfying S(z)T A(z) = 0, (6) constraint (5) implies the existence of a smooth function \u03bc such that \u03bc\u0307 = \u03bc\u0307(\u03bc, z) [9] such that z\u0307 = S(z)\u03bc (7) It is possible to find a solution to S(z) that allows choosing \u03bc = v, i", + " The required actuator input is then \u03c4 a = J T \u2020 (MM\u22121 o ua + Cv) (11) The next component of our control is an admittance model for motion on a plane, given by Md v\u0307d + Bdvd = Fm + Fasst (12) where vd is the reference velocity trajectory. The reference pose trajectory, xd , is generated by integrating x\u0307d = vd . The desired admittance of the walker is determined by the inertia matrix Md and the damping matrix Bd , both of which are positive definite diagonal matrices given by Md = diag(md w, md w, Id w) Bd = diag(bd w, bd w, \u03b2d w) (13) In Eq. 13, md w and I d w are, respectively, the desired mass and the desired moment of inertia with respect to the endeffector point c (Fig. 1b); bd w and \u03b2d w are, respectively, the desired linear and rotational coefficients of damping. Fm is the force exerted by the user\u2019s body on the end effector, assumed to depend on the compliance of the human body; Fasst is a computed virtual force designed to assist the user\u2019s motion on the horizontal plane. We assume that Fm is composed of a spring-like force Fh,k and a damper-like force Fh,b such that Fm = Fh,k( xh) + Fh,b( vh) (14) where xh = xh \u2212x and vh = vh \u2212v. Here xh and vh are, respectively, position and velocity coordinates that control the deformation of the human body", + " (In the discussion that follows, a vector with prefix (b) is expressed in body-aligned coordinates {b}; a vector with no prefix is expressed in global coordinates {o}.) For the stability analysis we assume that the walker\u2019s end-effector is fitted with a force and torque (FT) sensor. (The FT sensor in our prototype is shown in Fig. 11a and b.) We define bfs = [ fx fy fz \u03c4x \u03c4y \u03c4z ]T (24) as the vector of force and torque measurements from the FT sensor, with respect to coordinate frame {b}, the axes of which are parallel to the main axes x\u2032 b and y\u2032 b of the walker (Fig. 1b). Since the mass and inertia matrices in Eq. 12 are invariant with the walker\u2019s orientation, we can write Md bv\u0307d + Bd bvd = bFm(bfs) + bFasst bx\u0307d = bvd (25) where bxd and bvd are, respectively, the reference pose and reference velocity with respect to {b}. On the right-hand side of Eq. 25, the function bFm converts the FT sensor measurements to a suitable input for the admittance model. bFm is important because it determines how the walker will interact with the human user. For now, we will assume straightforward use of the sensor readings and define bFm = [fx fy \u03c4z]T (26) The admittance model Eq. 25 can be viewed as three decoupled linear systems, each of which can be specified independently. Here we discuss the design of the admittance control for motion parallel to the xb axis (Fig. 1b). The same method can later be applied to the remaining coordinates, yb and \u03c6c. The first step is to introduce compliance in the linearized model of the walker (10). We can express (10) with respect to frame {b} as Mo bv\u0307 = ua (27) where ua = oR\u22121 b (\u03c6c)ua and oRb(\u03c6c) is the rotation matrix between {b} and {o}. System (27) in expanded form is \u23a1 \u23a3 mo 0 0 0 mo 0 0 0 Io \u23a4 \u23a6 \u23a1 \u23a3 bvc,x bvc,y b\u03c6\u0307c \u23a4 \u23a6 = \u23a1 \u23a3 ua,x ua,y ua,\u03c6 \u23a4 \u23a6 (28) The first row of Eq. 28 represents the subsystem governing motion parallel to the xb axis", + " This is a significant improvement in the walker\u2019s dynamics. 4Mobile Platform Therapeutic Action: Assistance on the Horizontal Plane The next step in our design is to add therapeutic action to the admittance control developed in Section 2. Specifically, we want the walker to provide Supplemental forces to help propel the user\u2019s COM on the horizontal plane. In the general case, this is accomplished by adding a vector of virtual forces bFasst to the right-hand side of Eq. 25. Here we will focus exclusively on assistance in the forward direction, xb (Fig. 1b). The corresponding assistive input is bFasst = [Fasst,x 0 0 0 0 0]T (49) Fasst,x = Lasst,x Fref,x (50) where Fref,x is a reference force value and Lasst,x is a modulating function designed to control the amount of assistance felt by the user. It must be noted that the admittance control does not exert the assistive force Fasst,x directly on the user. Fasst,x can be thought of as the force exerted by somebody pushing the walker in the same direction in which the user is walking. The actual force felt by the user along the xb direction is \u2212Fm,x , i" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_35_0001419_978-94-007-7401-8-Figure2.102-1.png", + "original_path": "designv11-35/openalex_figure/designv11_35_0001419_978-94-007-7401-8-Figure2.102-1.png", + "caption": "Fig. 2.102 4RPaRRR-type fully-parallel PM with coupled Sch\u00f6nflies motions defined by MF \u00bc SF \u00bc 4; RF\u00f0 \u00de \u00bc \u00f0v1; v2; v3;xd\u00de, TF = 0, NF = 14, limb topology R\\Pa\\kR||R\\R", + "texts": [], + "surrounding_texts": [ + "156 2 Fully-Parallel Topologies with Coupled Sch\u00f6nflies Motions", + "2.2 Topologies with Complex Limbs 157", + "158 2 Fully-Parallel Topologies with Coupled Sch\u00f6nflies Motions" + ] + }, + { + "image_filename": "designv11_35_0001385_978-981-15-6095-8-Figure13.15-1.png", + "original_path": "designv11-35/openalex_figure/designv11_35_0001385_978-981-15-6095-8-Figure13.15-1.png", + "caption": "Fig. 13.15 a and b Deformation on the pipe (m): a unreinforced case; b geocell and geogrid reinforced case. Sourced from Hegde and Sitharam (2015c)", + "texts": [ + " However, in case of geocells, the stress was found to distribute in the lateral direction to a shallow depth. Similar observations were also made by Saride et al. (2009) and Hegde and Sitharam (2015d, e) during the numerical simulations of the geocell reinforced soil beds. Since geocell distribute the load in the lateral direction, the intensity of the stresswill reduce the soil existing blow the geocells. Therefore, the pipe will also experience less stress in the presence of reinforcement as compared to the unreinforced beds. Figure 13.15a and b shows the distribution of the vertical displacement contours on the surface of the pipe for unreinforced and the reinforced cases. The reported displacements are acting in the downward direction. From the figure, it is evident that the deformation of the pipe significantly reduces the presence of the geocells and geogrids. From the maximum value of the observed deformation, the strain on the pipe was deduced for both unreinforced and reinforced cases. The strain values thus calculated were 0" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_35_0000395_978-3-030-67750-3_16-Figure30.8-1.png", + "original_path": "designv11-35/openalex_figure/designv11_35_0000395_978-3-030-67750-3_16-Figure30.8-1.png", + "caption": "Fig. 30.8 Side view of loader bucket", + "texts": [], + "surrounding_texts": [ + "bucket, which is 100 mm linear actuator and 150 mm linear actuator depending on the length of the arm needed to move. All these actuators are controlled by a motor located at the wheelbarrow body tray and near to the handle where the user can easily manage to operate. The connection harness attached under the wheelbarrow bucket is for safety reason. All the items can be referred in Figs. 30.8 and 30.9 that shows the detail view of the loader bucket." + ] + }, + { + "image_filename": "designv11_35_0000371_tmag.2021.3076134-Figure6-1.png", + "original_path": "designv11-35/openalex_figure/designv11_35_0000371_tmag.2021.3076134-Figure6-1.png", + "caption": "Fig. 6. Basic dimensions of the proposed ST-PMTFLG. (a) Sectional view. (b) Side view.", + "texts": [ + " Although the amplitude difference between 2-D and 3-D FEA results is large, the waveforms are almost the same, so it is still meaningful to optimize the dimension parameters by 2-D simulation. The cost function of the optimization is to maximize the thrust/power density of the generator. The search space of dimension parameters is determined by the volume of the generator or other restrictions. The optimization constraints are usually magnetic circuit saturation limit, current density limit, slot area limit, machine volume limit, and so on. The dimension parameters of the proposed ST-PMTFLG are given in Fig. 6 and Table I. After the optimization of generator parameters by 2-D simulation, the performance of the generator is evaluated by 3-D simulation. In this section, the 12-stator-pole ST-PMTFLG with 10-, 11-, 13-, and 14-translator-pole translators are designed and compared to maximize the average electromagnetic thrust. The machine performance compared including open-circuit flux linkage, back EMF, and ON-load electromagnetic thrust characteristics, are as follows. The open-circuit flux distributions of 12-stator-pole ST-PMTFLGs having different translator pole numbers (10, 11, 13, and 14) when translator positions in electric \u03b8e = 0 are shown in Fig" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_35_0001419_978-94-007-7401-8-Figure3.70-1.png", + "original_path": "designv11-35/openalex_figure/designv11_35_0001419_978-94-007-7401-8-Figure3.70-1.png", + "caption": "Fig. 3.70 4PaPRP-type overactuated PM with coupled Sch\u00f6nflies motions defined by MF \u00bc SF \u00bc 4; RF\u00f0 \u00de \u00bc v1; v2; v3;xd\u00f0 \u00de, TF = 0, NF = 18, limb topology Pa\\P\\kR||P", + "texts": [ + "64) P\\Pa||R||P (Fig. 3.50t) Idem No. 1 21. 4PPaPR (Fig. 3.65) P\\Pa||P||R (Fig. 3.50u) Idem No. 1 22. 4PPaPR (Fig. 3.66a) P||Pa\\P\\kR (Fig. 3.50v) Idem No. 1 23. 4PPaRP (Fig. 3.66b) P||Pa\\R\\P (Fig. 3.50i0) The before last joints of the four limbs have parallel axes 24. 4RPPaP (Fig. 3.67) R||P||Pa\\P (Fig. 3.50w) Idem No. 1 25. 4RPRPa (Fig. 3.68) R||P||R||Pa (Fig. 3.50x) Idem No. 1 26. 4PaPPR (Fig. 3.69a) Pa||P\\P\\kR (Fig. 3.50y) Idem No. 1 27. 4PaPPR (Fig. 3.69b) Pa\\P\\kP||R (Fig. 3.50z) Idem No. 1 28. 4PaPRP (Fig. 3.70) Pa\\P\\kR||P (Fig. 3.50z1) Idem No. 1 29. 4PaRRP (Fig. 3.71) Pa||R||R||P (Fig. 3.50a0) Idem No. 1 30. 4PaRPR (Fig. 3.72) Pa||R||P||R (Fig. 3.50b0) Idem No. 1 31. 4PaPRP (Fig. 3.73a) Pa||P||R\\P (Fig. 3.50c0) Idem No. 1 32. 4PaRPP (Fig. 3.73b) Pa||R\\P\\kP (Fig. 3.50d0) Idem No. 1 33. 4PRPPa (Fig. 3.74a) P\\R\\P||Pa (Fig. 3.50e0) The second joints of the four limbs have parallel axes 34. 4PRPPa (Fig. 3.74b) P\\R\\P\\Pa (Fig. 3.50f0) Idem No. 33 35. 4PRPaP (Fig. 3.75a) P\\R\\Pa\\\\P (Fig. 3.50g0) Idem No. 33 36" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_35_0000797_9781119526483.ch14-Figure14.24-1.png", + "original_path": "designv11-35/openalex_figure/designv11_35_0000797_9781119526483.ch14-Figure14.24-1.png", + "caption": "Figure 14.24 Is diatom motility a special case of cytoplasmic streaming? (1943). Left: Early model of what we now call a motor protein invoked for cytoplasmic streaming. 1) Protein molecule in extended form. Bond at anterior end of molecule. (2) Protein molecule in contracted form. (3) Shifting of bond from anterior to posterior end. (4) Protein molecule in expanded form\u2026. Let us say that the molecules which form bonds in this way so as to contribute to the motion are in phase and those which may form bonds so as to oppose the motion are out of phase. It is clear that streaming can occur only when the number of molecules in phase exceeds those out of phase. This cannot happen by chance alone and we must therefore postulate that some mechanism is present which sets the majority of stream proteins in phase. This mechanism must of course be intimately related to the mechanism responsible for the reversal of streaming\u201d [14.222] (reprinted with permission of the American Philosophical Society). Right: Motion of an inchworm caterpillar [14.287] (reprinted with permission of Elsevier).", + "texts": [ + "182], might prove effective for future investigations of diatom motility. 12 \u201cJe n\u2019ai malheureusement aucune th\u00e9orie nouvelle \u00e0 proposer et n\u2019aper\u00e7ois aucun motif de recourir \u00e0 celles d\u00e9j\u00e0 abandonn\u00e9es\u201d [14.230]. Before we knew about motor molecules that could walk along cytoskeletal microfilaments (=\u00a0F-actin = filamentous actin) or microtubules, one model for cytoplasmic streaming postulated the proteins could extend and contract, thereby pushing or pulling protoplasm along [14.74] [14.222] [14.388] (Figure 14.24). This was dismissed after William Seifriz\u2019s \u201c\u2026disproof of the caterpillar hypothesis\u201d [14.323] (Figure 14.24), perhaps prematurely, in favor of \u201c\u2026a wave of depolarization giving rise to a field parallel to the axis would carry particles and fluids with it\u201d [14.323], apparently supplanting his \u201cchampioning contraction as the cause of protoplasmic streaming\u201d [14.74] in [14.322]. 13\u201cVue connective d\u2019une valve, montrant la canalisation au niveau du nodule m\u00e9dian et les parcours attribu\u00e9s aux courants propulseurs. Le nodule m\u00e9dian est dessin\u00e9 (d\u2019apr\u00e8s M\u00fcller), suivant une orientation qui ne montre pas le d\u00e9bouch\u00e9 de la \u00ab fente externe \u00bb au dehors; la structure interne des nodules polaires n\u2019est pas figur\u00e9e" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_35_0001419_978-94-007-7401-8-Figure3.43-1.png", + "original_path": "designv11-35/openalex_figure/designv11_35_0001419_978-94-007-7401-8-Figure3.43-1.png", + "caption": "Fig. 3.43 Overactuated PMs with coupled Sch\u00f6nflies motions of types 4RPRPR (a) and 4RRPPR (b) defined by MF \u00bc SF \u00bc 4; RF\u00f0 \u00de \u00bc v1; v2; v3;xd\u00f0 \u00de, TF = 0, NF = 2, limb topology R||P||R\\P\\\\R (a) and R||R||P\\P\\\\R (b)", + "texts": [ + "39b) R\\R||R||P\\P (Fig. 3.3z) Idem No. 16 40. 4PRRRP (Fig. 3.40a) P||R||R\\R\\P (Fig. 3.3z1) The fourth joints of the four limbs have parallel axes 41. 4RPRRP (Fig. 3.40b) R||P||R||R\\P (Fig. 3.3a) Idem No. 40 42. 4RPRRP (Fig. 3.41a) R||P||R\\R\\P (Fig. 3.3b0) Idem No. 40 43. 4RRPRP (Fig. 3.41b) R||R||P\\R\\P (Fig. 3.3c0) Idem No. 40 44. 4RRRPP (Fig. 3.42a) R||R\\R\\P\\\\P (Fig. 3.3d0) The third joints of the four limbs have parallel axes 45. 4PRRPR (Fig. 3.42b) P||R||R\\P\\\\R (Fig. 3.3e0) Idem No. 17 46. 4RPRPR (Fig. 3.43a) R||P||R\\P\\\\R (Fig. 3.3f0) Idem No. 17 47. 4RRPPR (Fig. 3.43b) R||R||P\\P\\\\R (Fig. 3.3g0) Idem No. 17 48. 4RPRPR (Fig. 3.44a) R||P||R\\P\\\\R (Fig. 3.3h0) Idem No. 17 49. 4RPRPR (Fig. 3.44b) P\\P\\kR||R\\\\R (Fig. 3.3i0) Idem No. 17 50. 4PPRRR (Fig. 3.44c) P\\P||R||R\\\\R (Fig. 3.3j0) Idem No. 17 51. 4CRRR (Fig. 3.45a) C||R\\R||R (Fig. 3.3k0) Idem No. 17 52. 4RCRR (Fig. 3.45b) R||C\\R||R (Fig. 3.3l0) Idem No. 17 53. 4RRCR (Fig. 3.46a) R||R\\C||R (Fig. 3.3m0) Idem No. 16 54. 4RRRC (Fig. 3.46b) R||R\\R||C (Fig. 3.3n0) Idem No. 16 55. 4RCRP (Fig. 3.47a) R\\C||R\\\\P (Fig. 3.3o0) Idem No" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_35_0001419_978-94-007-7401-8-Figure4.37-1.png", + "original_path": "designv11-35/openalex_figure/designv11_35_0001419_978-94-007-7401-8-Figure4.37-1.png", + "caption": "Fig. 4.37 1PaPn3-3PaPn3R-type fully-parallel PM with decoupled Sch\u00f6nflies motions defined by MF = SF = 4, (RF) = (v1; v2; v3;xa), TF = 0, NF = 27, limb topology Pa ? Pn3||R and Pa ? Pn3 ?|| R", + "texts": [ + " 4.10k) Pa ? Pn2||R ?? R (Fig. 4.11i) Idem No. 14 (continued) 426 4 Fully-Parallel Topologies with Decoupled Sch\u00f6nflies Motions Table 4.4 (continued) No. PM type Limb topology Connecting conditions 23. Pa ?Pn2||R ??R (Fig. 4.34) Pa ? Pn3||R (Fig. 4.10l) Pa ? Pn3 ?? R (Fig. 4.11j) Idem No. 6 24. 1PaPn3\u20133PaPn3R (Fig. 4.35) Pa ? Pn3||R (Fig. 4.10m) Pa ? Pn3 ??|R (Fig. 4.11k) Idem No. 6 25. 1PaPn3\u20133PaPn3R (Fig. 4.36) Pa ? Pn3||R (Fig. 4.10n) Pa ?Pn3 ?? R (Fig. 4.11l) Idem No. 6 26. 1PaPn3\u20133PaPn3R (Fig. 4.37) Pa ? Pn3||R (Fig. 4.10o) Pa ? Pn3 ??R (Fig. 4.11m) Idem No. 6 27. 1CRbRbR\u20133CRbRbRR (Fig. 4.38) C||Rb||Rb||R (Fig. 4.10p) C||Rb||Rb||R ? R (Fig. 4.11n) Idem No. 6 28. 4PaPaPaR (Fig. 4.39) Pa ? P||Pa||R (Fig. 4.12a) Pa ? Pa||Pa ?? R (Fig. 4.12c) Idem No. 6 29. 4PaPaPaR (Fig. 4.40) Pa ? Pa||Pa||R (Fig. 4.12a) Pa ? Pa||Pa ??R (Fig. 4.12b) Idem No. 6 3PaRRbRbRR (Fig. 4.41) Pa ? R||Rb||Rb||R (Fig. 4.12d) Pa ? R||Rb||Rb||R ??R (Fig. 4.13a) Idem No. 6 3PaRRbRbRR (Fig. 4.42) Pa ? R||Rb||Rb||R (Fig. 4.12e) Pa " + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_35_0000502_tia.2021.3088767-Figure11-1.png", + "original_path": "designv11-35/openalex_figure/designv11_35_0000502_tia.2021.3088767-Figure11-1.png", + "caption": "Fig. 11. Experimental setup for iron-ball magnetic levitation (left) and the electromagnets with a permanent magnet (right).", + "texts": [ + " 9 and 10 show the measured capacitor voltages vc1 and vc2 under the condition of izs* = 0 A without motor drive and magnetic suspension, using voltage FB control and voltage sensorless control, respectively. The capacitor voltages were unbalanced by Ioff when izs* = 0 A was applied at 10 s. Then, the capacitor voltages converged to 80 V as soon as two balance control methods were activated at approximately 15 s. The performance of the sensorless control was equivalent to that of the voltage FB control. Fig. 11 shows the experimental setup for the iron-ball magnetic suspension. A hollow iron ball with a diameter of 50 mm and weight of 160 g and a cylindrical electromagnet with an E-shaped cross section were used. To eliminate the bias current for magnetic suspension regulation, a PM was attached to the bottom of the electromagnet. As shown in Fig. 8, the zeropower control [29] was used because the control could automatically compensate the levitation position z0* for balancing the self-weight of the iron ball and the PM attractive Authorized licensed use limited to: Univ of Calif Santa Barbara" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_35_0001419_978-94-007-7401-8-Figure5.8-1.png", + "original_path": "designv11-35/openalex_figure/designv11_35_0001419_978-94-007-7401-8-Figure5.8-1.png", + "caption": "Fig. 5.8 2PaPRRR-1PaPRR-1RPPP-type fully-parallel PMs with uncoupled Sch\u00f6nflies motions defined by MF = SF = 4, RF\u00f0 \u00de \u00bc \u00f0v1; v2; v3;xb\u00de, TF = 0, NF = 13, limb topology R\\P\\\\P\\\\P and Pa\\P\\\\R||R\\\\R, Pa\\P\\\\R||R (a), Pa||P\\R||R\\\\R, Pa||P\\R||R (b)", + "texts": [], + "surrounding_texts": [ + "5.1 Fully-Parallel Topologies 441", + "442 5 Topologies with Uncoupled Sch\u00f6nflies Motions", + "5.1 Fully-Parallel Topologies 443" + ] + }, + { + "image_filename": "designv11_35_0001731_pime_proc_1939_142_025_02-Figure32-1.png", + "original_path": "designv11-35/openalex_figure/designv11_35_0001731_pime_proc_1939_142_025_02-Figure32-1.png", + "caption": "Fig. 32. Sectional Arrangement of 37,500 kW. Brush-Ljungstrom Turbine at Brighton Power Station", + "texts": [ + " This may perhaps be considered as the modern equivalent of the old vertical \u201cCurtis\u201d turboalternators-many examples of which were installed by the British Thomson-Houston Company, Ltd., in this country in the early days of power station developments. Amongst other interesting and recent additions to the power station plants of this country are the 37,500 kW. Brush-Ljungstrom machines installed at the Southwick Station of the Brighton Corporation. These plants run at a speed of 1,500 r.p.m. and are designed for steam at 650 lb. per sq. in. and 850 deg. F., with a vacuum of 29 inches of mercury. A section through the last of these sets is shown in Fig. 32. It will be seen that the turbine comprises two sets of concentric rings of blading carried on the faces of disks rotating oppositely at a speed of 1,500 r.p.m. Each half of the turbine thus drives its own alternator. In the main, therefore, for the greater part of its expansion the steam is expanded in a radial at UQ Library on June 4, 2016pme.sagepub.comDownloaded from FORTY YEARS\u2019 D E V E L O P M E N T I N P O W E R STATIOIL\u2019 P L A N T \u2018A\u201c STATION SITE COAL STORE O O E O 4 325 Barking \u201cB\u201d Station" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_35_0000174_j.matpr.2021.01.640-Figure12-1.png", + "original_path": "designv11-35/openalex_figure/designv11_35_0000174_j.matpr.2021.01.640-Figure12-1.png", + "caption": "Fig. 12. Stress analysis on Metal Matrix Composites.", + "texts": [ + "In this analysis von-mises elastic strain is suitable that is energy associated with shear strains, Strain energy can be separated into energy associated with volume change and energy associated with distortion of the body. The maximum distortion energy failure theory assumes failure by yielding in a more complicated loading situation to occur when the distortion energy in the material reaches the same value as in a tension test at yield. The maximum strain is obtained as 6.5176e-7 and minimum strain is obtained as 3.3411e-9 Engineering analysis Stress on Metal Matrix Composites as shown in Fig. 12.Stress analysis on metal matrix composite, von mises elastic theory is applied, the stress value obtained as maximum 74.301 MPa and minimum as 0.38088 MPa. etal Matrix Composites. By Modal analysis the characteristic frequencies and mode shapes shown in Fig. 13. In the event that the shat revolves at its characteristic recurrence, it very well may be seriously vibrated. The modal investigation performed to locate the normal frequencies Fig. 14 Fig. 15. Modal analysis in structural mechanics is determined the natural mode shapes and frequencies of an object or structure during free vibration" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_35_0000095_s00170-021-06769-1-Figure2-1.png", + "original_path": "designv11-35/openalex_figure/designv11_35_0000095_s00170-021-06769-1-Figure2-1.png", + "caption": "Fig. 2 The test device sketch Fig. 3 The test device outlook", + "texts": [ + " The primary (cutting) movement is done by the workpiece. This arrangement takes advantage of the relative static knife. Thereby forces can be measured directly beneath the knife, avoiding problems occurring by damping, e.g., due to high mass, or resonance effects. Additionally, by avoiding force measurement beneath the workpiece, modification of the force data by the heterogenetic and anisotropic wood material is prevented. It has to be considered, that this configuration also includes the component of chip acceleration. The rotation arm (Fig. 2) has a diameter of 4 m and is powered by amotor. The radius of this length ismore complex to operate but can simulate linear movement for the longitudinal cut. This length is the highest feasible for keeping the device in the lab-scale (Fig. 3). Dimensioning of the rotor arm (Fig. 4a) was assisted by FEM simulations (Fig. 4b). For the simulation, the rotor arm design was simplified (fix connections, removed screws, etc.). Nevertheless, FEM analysis had to serve as a comparison of different design performance; therefore, it was simulated as a cantilever beam only" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_35_0000331_s12206-021-0408-4-Figure11-1.png", + "original_path": "designv11-35/openalex_figure/designv11_35_0000331_s12206-021-0408-4-Figure11-1.png", + "caption": "Fig. 11. The seven states of axle box bearings: (a) normal; (b) slight outer-race faults; (c) serious outer-race faults; (d) slight inner-race faults; (e) serious inner-race faults; (f) slight ball faults; (g) serious ball faults.", + "texts": [ + " 10(b) shows the position where the acceleration sensor is installed on the UWL. The tested bearing is a double-row cylindrical roller bearing (NJ(P)3226X1). More parameters of the bearing are shown in Table 2. The experiment obtains bearings in the normal state, rolling element faults, inner-race faults and outer-race faults through precision machining. The bearing faults are slightly damaged and severely damaged, with dimensions of 5 mm in width, 1.5 mm in depth and 10 mm in width, 2 mm in depth, as shown in Fig. 11. A total of 7 types of bearings are obtained, and they are marked as Norm, BE1, BE2, IR1, IR2, OR1, and OR2. The detailed information is shown in Table 3. In the experiment, the outer ring of the axle box bearing is fixed, the inner ring rotates, the speed is 45 r/min, the sampling frequency is 25.6 KHz. Under such conditions, 30 groups of vibration signals are collected for each bearing state, the length of each group of vibration signals is 2048, and a total of 210 groups of vibration signal data are collected" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_35_0001419_978-94-007-7401-8-Figure2.68-1.png", + "original_path": "designv11-35/openalex_figure/designv11_35_0001419_978-94-007-7401-8-Figure2.68-1.png", + "caption": "Fig. 2.68 4RPPaPa-type fully-parallel PM with coupled Sch\u00f6nflies motions defined by MF \u00bc SF \u00bc 4; RF\u00f0 \u00de \u00bc \u00f0v1; v2; v3;xd\u00de, TF = 0, NF = 30, limb topology R||P\\Pa\\\\Pa", + "texts": [ + " 15 48. 4PaPPaR (Fig. 2.60) Pa\\P\\\\Pa\\kR (Fig. 2.22u) Idem No. 4 49. 4PaRPaP (Fig. 2.61) Pa\\R\\Pa\\kR (Fig. 2.22v) Idem No. 34 50. 4PaRPPa (Fig. 2.62) Pa\\R||P\\Pa (Fig. 2.22w) Idem No. 34 51. 4PaPRPa (Fig. 2.63) Pa\\P||R\\Pa (Fig. 2.22x) Idem No. 33 52. 4PaPaPR (Fig. 2.64) Pa\\Pa\\\\P||R (Fig. 2.22y) Idem No. 4 53. 4PaPaRP (Fig. 2.65) Pa\\Pa\\\\R||P (Fig. 2.22z) Idem No. 41 54. 4RPaPaP (Fig. 2.66) R\\Pa\\\\Pa\\\\P (Fig. 2.22a0) Idem No. 15 55. 4RPaPPa (Fig. 2.67) R\\Pa\\kP\\\\Pa (Fig. 2.22b0) Idem No. 15 56. 4RPPaPa (Fig. 2.68) R||P\\Pa\\\\Pa (Fig. 2.22c0) Idem No. 15 57. 4PRPaPa (Fig. 2.69) P||R\\Pa\\\\Pa (Fig. 2.22d0) Idem No. 5 58. 4PaCPa (Fig. 2.70) Pa\\C\\Pa (Fig. 2.22e0) The axes of the cylindrical joints of the four limbs are parallel 59. 4CPaPa (Fig. 2.71) C\\Pa\\\\Pa (Fig. 2.22f 0) Idem No. 58 60. 4PaPaC (Fig. 2.72) Pa\\Pa\\\\C (Fig. 2.22g0) Idem No. 58 176 2 Fully-Parallel Topologies with Coupled Sch\u00f6nflies Motions T ab le 2. 4 L im b to po lo gy an d co nn ec ti ng co nd it io ns of th e fu ll y- pa ra ll el so lu ti on s w it h no id le m ob il it ie s pr es en te d in F ig s" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_35_0001419_978-94-007-7401-8-Figure6.13-1.png", + "original_path": "designv11-35/openalex_figure/designv11_35_0001419_978-94-007-7401-8-Figure6.13-1.png", + "caption": "Fig. 6.13 Maximaly regular fully-parallel PMs with Sch\u00f6nflies motions of types 2PRPRR1PRPR-1RPPaPa (a) and 2PRPRR-1PRPR-1RPPaPat (b) defined by MF = SF = 4, (RF) = (v1; v2; v3;xb), TF = 0, NF = 10, limb topology P||R\\P\\||R\\R, P||R\\P\\||R and R||P||Pa||Pa (a), R||P||Pa||Pat (b)", + "texts": [ + "11b) P\\P ??P\\||R (Fig. 4.1a) P\\P ??P ??R (Fig. 4.1b) R||P||Pa||Pat (Fig. 5.4n) Idem no. 9 11. 2PPRRR-1PPRR-1RPPaPa (Fig. 6.12a) P\\P\\||R||R\\R (Fig. 4.2a) P\\P\\||R||R (Fig. 4.1d) R||P||Pa||Pa (Fig. 5.4m) Idem no. 9 (continued) 612 6 Maximally Regular Topologies with Sch\u00f6nflies Motions Table 6.5 (continued) No. PM type Limb topology Connecting conditions 12. 2PPRRR-1PPRR-1RPPaPat (Fig. 6.12b) P\\P\\||R||R\\R (Fig. 4.2a) P\\P\\||R||R (Fig. 4.1d) R||P||Pa||Pat (Fig. 5.4n) Idem no. 9 13. 2PRPRR-1PRPR-1RPPaPa (Fig. 6.13a) P||R\\P\\||R\\R (Fig. 4.2b) P||R\\P\\||R (Fig. 4.1e) R||P||Pa||Pa (Fig. 5.4m) Idem no. 9 14. 2PRPRR-1PRPR-1RPPaPat (Fig. 6.13b) P||R\\P\\||R\\R (Fig. 4.2b) P||R\\P\\||R (Fig. 4.1e) R||P||Pa||Pat (Fig. 5.4n) Idem no. 9 15. 2PRRRR-1PRRR-1RPPaPa (Fig. 6.14a) P||R||R||R\\R (Fig. 4.2d) P||R||R||R (Fig. 4.1g) R||P||Pa||Pa (Fig. 5.4m) Idem no. 9 16. 2PRRRR-1PRRR1RPaPatP (Fig. 6.14b) P||R||R||R\\R (Fig. 4.2d) P||R||R||R (Fig. 4.1g) R||P||Pa||Pat (Fig. 5.4n) Idem no. 9 17. 2CPRR-1CPR-1RPaPaP (Fig. 6.15a) C\\P\\||R\\R (Fig. 4.2e) C ?P ?||R (Fig. 4.1i) R||Pa||Pa||P (Fig. 5.4k) Idem no. 1 18. 2CPRR-1CPR1RPaPatP (Fig. 6.15b) C ?P ?||R ?R (Fig" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_35_0001419_978-94-007-7401-8-Figure6.12-1.png", + "original_path": "designv11-35/openalex_figure/designv11_35_0001419_978-94-007-7401-8-Figure6.12-1.png", + "caption": "Fig. 6.12 Maximaly regular fully-parallel PMs with Sch\u00f6nflies motions of types 2PPRRR1PPRR-1RPPaPa (a) and 2PPRRR-1PPRR-1RPPaPat (b) defined by MF = SF = 4, (RF) = (v1; v2; v3;xb), TF = 0, NF = 10, limb topology P\\P\\||R||R\\R, P\\P\\||R||R and R||P||Pa||Pa (a), R||P||Pa||Pat (b)", + "texts": [ + "2d) P||R||R||R (Fig. 4.1g) R||Pa||Pat||P (Fig. 5.4l) Idem no. 1 9. 3PPPR-1RPPaPa (Fig. 6.11a) P\\P ??P\\||R (Fig. 4.1a) P\\P ??P ??R (Fig. 4.1b) R||P||Pa||Pa (Fig. 5.4m) The last revolute joints of limbs G1, G2 and G3 and the revolute joint between links 7D and 9 have superposed axes. The actuated prismatic joints of limbs G1, G2 and G3 have orthogonal directions 10. 3PPPR-1RPPaPat (Fig. 6.11b) P\\P ??P\\||R (Fig. 4.1a) P\\P ??P ??R (Fig. 4.1b) R||P||Pa||Pat (Fig. 5.4n) Idem no. 9 11. 2PPRRR-1PPRR-1RPPaPa (Fig. 6.12a) P\\P\\||R||R\\R (Fig. 4.2a) P\\P\\||R||R (Fig. 4.1d) R||P||Pa||Pa (Fig. 5.4m) Idem no. 9 (continued) 612 6 Maximally Regular Topologies with Sch\u00f6nflies Motions Table 6.5 (continued) No. PM type Limb topology Connecting conditions 12. 2PPRRR-1PPRR-1RPPaPat (Fig. 6.12b) P\\P\\||R||R\\R (Fig. 4.2a) P\\P\\||R||R (Fig. 4.1d) R||P||Pa||Pat (Fig. 5.4n) Idem no. 9 13. 2PRPRR-1PRPR-1RPPaPa (Fig. 6.13a) P||R\\P\\||R\\R (Fig. 4.2b) P||R\\P\\||R (Fig. 4.1e) R||P||Pa||Pa (Fig. 5.4m) Idem no. 9 14. 2PRPRR-1PRPR-1RPPaPat (Fig. 6.13b) P||R\\P\\||R\\R (Fig. 4.2b) P||R\\P\\||R (Fig. 4.1e) R||P||Pa||Pat (Fig. 5.4n) Idem no. 9 15. 2PRRRR-1PRRR-1RPPaPa (Fig. 6.14a) P||R||R||R\\R (Fig. 4.2d) P||R||R||R (Fig. 4.1g) R||P||Pa||Pa (Fig. 5.4m) Idem no. 9 16. 2PRRRR-1PRRR1RPaPatP (Fig. 6.14b) P||R||R||R\\R (Fig" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_35_0001419_978-94-007-7401-8-Figure3.10-1.png", + "original_path": "designv11-35/openalex_figure/designv11_35_0001419_978-94-007-7401-8-Figure3.10-1.png", + "caption": "Fig. 3.10 4RPPR-type overactuated PMs with coupled Sch\u00f6nflies motions defined by MF \u00bc SF \u00bc 4; RF\u00f0 \u00de \u00bc v1; v2; v3;xd\u00f0 \u00de, TF = 0, NF = 6, limb topology R||P\\P\\kR (a) and R\\P\\kP||R (b)", + "texts": [ + " 2 5. 4RRPR (Fig. 3.6a) R||R||P||R (Fig. 3.1e) Idem No. 2 6. 4PPRR (Fig. 3.6b) P\\P\\kR||R (Fig. 3.1i) Idem No. 1 7. 4RRRP (Fig. 3.7a) R||R||R||P (Fig. 3.1f) Idem No. 2 8. 4RPRP (Fig. 3.7b) R\\P\\kR||P (Fig. 3.1g) Idem No. 2 9. 4PPRR (Fig. 3.8a) P\\P||R||R (Fig. 3.1h) The last revolute joints of the four limbs have parallel axes 10. 4PRPR (Fig. 3.8b) P\\R||P||R (Fig. 3.1p) Idem No. 9 11. 4RRPP (Fig. 3.9a) R||R||P\\P (Fig. 3.1j) Idem No. 2 12. 4RPRP (Fig. 3.9b) R||P||R\\P (Fig. 3.1k) Idem No. 2 13. 4RPPR (Fig. 3.10a) R||P\\P\\kR (Fig. 3.1l) Idem No. 2 14. 4RPPR (Fig. 3.10b) R\\P\\kP||R (Fig. 3.1m) Idem No. 2 15. 4PRRP (Fig. 3.11a) P\\R||R||P (Fig. 3.1n) The last prismatic joints of the four limbs have parallel directions 16. 4PRPR (Fig. 3.11b) P\\R||P||R (Fig. 3.1o) Idem No. 9 17. 4PRRP (Fig. 3.12a) P||R||R\\P (Fig. 3.1v) The first prismatic joints of the four limbs have parallel directions 18. 4RPRP (Fig. 3.12b) R||P||R\\P (Fig. 3.1q) Idem No. 2 19. 4PRPR (Fig. 3.13a) P||R\\P\\kR (Fig. 3.1r) Idem No. 1 20. 4RPPR (Fig. 3.13b) R||P\\P\\kR (Fig. 3.1s) Idem No. 2 21. 4RPPR (Fig" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_35_0000245_ilt-02-2020-0072-Figure1-1.png", + "original_path": "designv11-35/openalex_figure/designv11_35_0000245_ilt-02-2020-0072-Figure1-1.png", + "caption": "Figure 1 Slider-crank-mechanism dynamic model", + "texts": [ + " The width, radius, radial clearance and rough surface of crankpin bearing are then optimized based on the genetic algorithm to improve the LTP. The paper\u2019s innovation is an optimal method of genetic algorithm known as a multiobjective optimization with wide search range and short search time; and a combined model of the slider-crank-mechanism dynamic and crankpin bearing lubrication are successfully developed to optimize the LTP. To calculate slider-crank-mechanism dynamic equations, the center of crankshaft coinciding with the cylinder center is assumed to build themodel in Figure 1. Based on the model and its dynamic motion, the piston\u2019s acceleration is: \u20aczp \u00bc rv2 cosf 1 l cos2f\u00f0 \u00de (1) where l = r/l; r and v are the rotation radius and angular velocity of the crankshaft; l is the connecting rod length. Assuming that the connecting rod\u2019s mass is referred to the lumped masses of the big-rod-end (mbr) and small-rod-end (msr). Thus, the inertial force of the piston (mp) and small-rodend is: Fip \u00bc mzp \u00bc mp 1msr\u00f0 \u00derv2 cosf 1 l cos2f\u00f0 \u00de (2) Under the impact of combustion gas pressure (P = pgpR2) on the piston peak, the impacting forces on the crankpin are: F3 \u00bc P1Fip\u00f0 \u00decos f 1 u\u00f0 \u00de=cosu F4 \u00bc P1Fip\u00f0 \u00desin f 1 u\u00f0 \u00de=cosu ( (3) The big-rod-end of connecting rod is rotated with v ; thus, its centrifugal inertial force is: Fic \u00bc mbrrv2 (4) The connecting rod\u2019s bearing provides a rotating motion of the crankpin within the connecting rod" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_35_0000686_j.mechmachtheory.2021.104433-Figure4-1.png", + "original_path": "designv11-35/openalex_figure/designv11_35_0000686_j.mechmachtheory.2021.104433-Figure4-1.png", + "caption": "Fig. 4. Equivalent mechanism of the epicyclic external gear train.", + "texts": [ + " The position of the instantaneous screw axis reveals that the motion of the epicyclic external gear train can be decomposed into two relatively independent motions: the relative mesh transmission of the sun gear and the planet gear, and the rotation of the gear arm and the planet gear around the axis of the sun gear. To examine the metamorphosis, a link is added to the 4-bar linkage [50] which is the replacement mechanism of the fixed axis external gear train by connecting the center of the sun gear and the center of the planet gear, representing the constant distance between two gears. As such, the open-loop link becomes a closed-loop linkage, giving mobility 2. Then, the equivalent mechanism of the epicyclic external gear train is proposed. As shown in Fig. 4, the links leOB, leCD, and leAC are equivalent to the sun gear, the planet gear, and the gear arm, respectively. Joint A and joint B are the centers of curvature of the tooth profile curve at the mesh point. The higher pair at the mesh point is replaced by link leDB, which is perpendicular to link leOB, and link leCD. The angle between link leDB and link leAC is the pressure angle \u03b1e, which is the same as the angle between link leCD and link leAC. Instantaneously, the equivalent mechanism has the same mobility, instantaneous velocity, and instantaneous acceleration as the epicyclic external gear train" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_35_0001419_978-94-007-7401-8-Figure2.44-1.png", + "original_path": "designv11-35/openalex_figure/designv11_35_0001419_978-94-007-7401-8-Figure2.44-1.png", + "caption": "Fig. 2.44 4PaPPaR-type fully-parallel PM with coupled Sch\u00f6nflies motions defined by MF \u00bc SF \u00bc 4; RF\u00f0 \u00de \u00bc \u00f0v1; v2; v3;xd\u00de, TF = 0, NF = 30, limb topology Pa\\P\\\\Pa\\\\R", + "texts": [ + "39) Pa\\C\\P (Fig. 2.21z1) Idem No. 11 28. 4PPaPaR (Fig. 2.40) P\\Pa\\\\Pa\\kR (Fig. 2.22a) Idem No. 4 29. 4PPaRPa (Fig. 2.41) P\\Pa||R\\Pa (Fig. 2.22b) The axes of the revolute joints connecting the two parallelogram loops of the four limbs are parallel (continued) 2.2 Topologies with Complex Limbs 175 Table 2.3 (continued) No. PM type Limb topology Connecting conditions 30. 4PRPaPa (Fig. 2.42) P\\R||Pa\\Pa (Fig. 2.22c) Idem No. 5 31. 4RPPaPa (Fig. 2.43) R\\P\\kPa\\\\Pa (Fig. 2.22d) Idem No. 15 32. 4PaPPaR (Fig. 2.44) Pa\\P\\\\Pa\\\\R (Fig. 2.22e) Idem No. 4 33. 4PaPRPa (Fig. 2.45) Pa\\P\\kR\\Pa (Fig. 2.22f) The axes of the revolute joints connecting links 5 and 6 of the four limbs are parallel 34. 4PaRPPa (Fig. 2.46) Pa||R\\P\\\\Pa (Fig. 2.22g) The axes of the revolute joints connecting links 4 and 5 of the four limbs are parallel 35. 4RPaPPa (Fig. 2.47) R||Pa\\P\\\\Pa (Fig. 2.22h) Idem No. 15 36. 4PaPaPR (Fig. 2.48) Pa\\Pa\\\\P\\\\R (Fig. 2.22i) Idem No. 4 37. 4RPaPaR (Fig. 2.49) R\\Pa||Pa\\kR (Fig. 2.22j) The first and the last revolute joints of the four limbs are parallel 38" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_35_0000109_s40192-021-00201-y-Figure1-1.png", + "original_path": "designv11-35/openalex_figure/designv11_35_0000109_s40192-021-00201-y-Figure1-1.png", + "caption": "Fig. 1 Schematic of the laser-scanned bare metal plate of IN625 (reproduced from [6])", + "texts": [ + " Both systems use continuous-wave ytterbium fiber (Yb: fiber) lasers with a 1070\u00a0nm wavelength. The melt pool is monitored by a near-infrared (NIR) camera and high-speed shortwave infrared camera (SWIR), for the AMMT, and CBM systems, respectively. The system parameters, including the beam diameters, are shown in Table\u00a01. As specified in Table\u00a02, on each machine, a set of 10 measurements was taken for three different power and scanning speeds. The schematic of the laser-scanned substrate is shown in Fig.\u00a01. Each track was scanned at 5-min intervals to avoid heat buildup, and the nominal temperature of the base plate was 25\u00a0\u00b0C. The experimental data released by the AM-Bench committee includes the melt pool geometry of the laserscanned bare plate of IN625. In this paper, the molten pool depth and width, and cooling rates are considered for comparative assessment of different modeling approaches. The Case Power (W) Scan speed (mm/s) Replications AMMT-A 137.9 400 3 AMMT-B 179.2 800 3 AMMT-C 179.2 1200 4 CBM-A 150 400 3 CBM-B 195 800 3 CBM-C 195 1200 4 1 3 detailed report of the benchmarks cases can be found in the AM-Bench webpage [6]" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_35_0001419_978-94-007-7401-8-Figure5.63-1.png", + "original_path": "designv11-35/openalex_figure/designv11_35_0001419_978-94-007-7401-8-Figure5.63-1.png", + "caption": "Fig. 5.63 3PaPaPaR-1RPaPaP-type fully-parallel PM with uncoupled Sch\u00f6nflies motions defined by MF = SF = 4, RF\u00f0 \u00de \u00bc \u00f0v1; v2; v3;xb\u00de, TF = 0, NF = 39, limb topology Pa\\Pa||Pa\\kR, Pa\\Pa||Pa||R and R||Pa||Pa||P", + "texts": [ + "60) Pa\\Pa||Pa\\\\R (Fig. 5.6c) Idem No. 1 Pa\\Pa||Pa||R (Fig. 5.6b) R\\P\\\\P\\\\P (Fig. 5.1a) 27. 3PaPaPaR-1RUPU (Fig. 5.61) Pa\\Pa||Pa\\kR (Fig. 5.6a) The first revolute joint of G4 limb and the last revolute joints of limbs G1, G2 and G3 have parallel axes. The last joints of of limbs G1, G2 and G3 have superposed axes Pa\\Pa||Pa||R (Fig. 5.6b) R\\R\\R\\P\\kR\\R (Fig. 5.1b) 28. 3PaPaPaR-1RUPU (Fig. 5.62) Pa\\Pa||Pa\\\\R (Fig. 5.6c) Idem No. 27 Pa\\Pa||Pa||R (Fig. 5.6b) R\\R\\R\\P\\kR\\R (Fig. 5.1b) 29. 3PaPaPaR-1RPaPaP (Fig. 5.63) Pa\\Pa||Pa\\kR (Fig. 5.6a) Idem No. 1 Pa\\Pa||Pa||R (Fig. 5.6b) R||Pa||Pa||P (Fig. 5.4k) (continued) 5.1 Fully-Parallel Topologies 533 Table 5.4 (continued) No. PM type Limb topology Connecting conditions 30. 3PaPaPaR-1RPaPaP (Fig. 5.64) Pa\\Pa||Pa\\\\R (Fig. 5.6c) Idem No. 1 Pa\\Pa||Pa||R (Fig. 5.6b) R||Pa||Pa||P (Fig. 5.54k) 31. 3PaPaPaR1RPaPatP (Fig. 5.65) Pa\\Pa||Pa\\kR (Fig. 5.6a) Idem No. 1 Pa\\Pa||Pa||R (Fig. 5.6b) R||Pa||Pat||P (Fig. 5.4l) 32. 3PaPaPaR1RPaPatP (Fig. 5.66) Pa\\Pa||Pa\\\\R (Fig. 5.6c) Idem No" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_35_0000535_tmag.2021.3085750-Figure2-1.png", + "original_path": "designv11-35/openalex_figure/designv11_35_0000535_tmag.2021.3085750-Figure2-1.png", + "caption": "Fig. 2 Open-circuit field distributions under different magnetization states.", + "texts": [ + " The variable-flux characteristics of the proposed HMCVFMM can be illustrated by the open-circuit field V Authorized licensed use limited to: San Francisco State Univ. Downloaded on July 03,2021 at 08:24:04 UTC from IEEE Xplore. Restrictions apply. 0018-9464 (c) 2021 IEEE. Personal use is permitted, but republication/redistribution requires IEEE permission. See http://www.ieee.org/publications_standards/publications/rights/index.html for more information. BJ-13 2 distributions under different magnetization states (MSs) in Fig. 2. It can be observed that the air-gap flux can be flexibly adjusted by applying a transient magnetizing or demagnetizing current pulse. Meanwhile, due to the variable flux capability, the losses at different speeds and loads can be also manipulated to realize efficiency improvement over a wide operating range. The nonlinear hysteresis characteristics of the LCF PMs used in the proposed HMC-VFMM can be characterized by a simple Fourier-series based hysteresis model as shown in Fig. 3(a). The PM working point can track along different recoil lines by applying a specific remagnetizing or demagnetizing current pulse and finally stabilize at the intersect of the recoil line and load line" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_35_0000225_j.addma.2021.101955-Figure8-1.png", + "original_path": "designv11-35/openalex_figure/designv11_35_0000225_j.addma.2021.101955-Figure8-1.png", + "caption": "Fig. 8. Meshing strategy and contribution of the distortion factor; a) model with vertex nodes between layer planes; b) model without vertex nodes between layer planes.", + "texts": [ + " In a similar fashion, the distortion of the part before cut off from the substrate can be obtained through the summation of the distortion contribution of all voxels. As mentioned before, the proposed algorithm only considers the influences of latter layers to earlier layers, so the green voxel in Fig. 7 contributes to the distortion of nodes 1, 2 and 3, but does not contribute any distortion to node 4 because the height of node 4 is higher than the height of the green voxel. The rationale for meshing the CAD model layer by layer is depicted in Fig. 8. The force couple from the green voxel will be distributed on nodes 1, 2, 4, 6, 7, and 8, as shown in Fig. 8a. Since the height of the voxel is higher than nodes 5,7 and 8, it will influence the distortion of nodes 5, 7, and 8. But the sub FEM models created by nodes 1, 2, and 6 do not contribute to nodes 5, 7, and 8, because they are lower than nodes 5, 7, and 8. Since this algorithm focuses on part distortion of the skin and the vertex nodes of surface triangles are capable of representing the outer skin of the part, the layer by layer manner of mesh generation is only confined to the surface of each layer, and an arbitrary mesh is created inside each layer. As shown in Fig. 8b, no matter where the voxel is, its contribution to the nodes on the plane of each mesh layer, such as nodes 1, 2, and 6, or 4, 13, and 3 can be obtained accurately. In the next step, the spring back distortion after part cut off from the substrate is predicted as depicted in Fig. 9. First, the interfacial force between the part and the substrate is calculated through a FEM model that contains the substrate only, as shown in Fig. 9b. The boundary conditions on nodes 1 and 2 are the same as those when calculating the distortion factor, as shown in Fig" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_35_0000686_j.mechmachtheory.2021.104433-Figure22-1.png", + "original_path": "designv11-35/openalex_figure/designv11_35_0000686_j.mechmachtheory.2021.104433-Figure22-1.png", + "caption": "Fig. 22. Metamorphic epicyclic bevel gear train based on the combination constraint.", + "texts": [ + " Based on Section 2.3, a metamorphic epicyclic bevel gear clamping mechanism is designed by using the epicyclic bevel gear train to replace the equivalent mechanism. According to the equivalent geometrical relationship, the key condition, \u03c9A = \u03c9B, of equivalent metamorphic clamping mechanism means that the angular velocity of sun bevel gear is equal to the angular velocity of the bevel gear arm. To realize and adjust this condition, a structure design scheme using combination constraint (inspired from Fig. 22) has been completed as shown in Fig. 27(a). The structure of the metamorphic epicyclic bevel gear clamping mechanism can be described as follow: the sun bevel gear, the driver, drives the planet bevel gear. The lead screw nut is fixed on the planet bevel gear and drives the lead screw claw to move. A limit spring connects the lead screw claw and the lead screw nut. The bevel gear arm holds the bevel gears and limits the rotation of the lead screw claw. In its rotation configuration as shown in Fig" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_35_0000395_978-3-030-67750-3_16-Figure30.1-1.png", + "original_path": "designv11-35/openalex_figure/designv11_35_0000395_978-3-030-67750-3_16-Figure30.1-1.png", + "caption": "Fig. 30.1 Wheelbarrow applying second class lever", + "texts": [ + "my \u00a9 The Author(s), under exclusive license to Springer Nature Switzerland AG 2021 M. H. Abu Bakar et al. (eds.), Progress in Engineering Technology III, Advanced Structured Materials 148, https://doi.org/10.1007/978-3-030-67750-3_30 357 A wheelbarrow is small transporting hardware and usually with one wheel at the front, two handles for users to hold and lift it, and the load will be in between of the handle and front wheel. The wheelbarrow eases the user by having the load in the middle of the fulcrum and the effort, applying the second-class lever as in Fig. 30.1. Nowadays, there have been numerous designs of wheelbarrow developed, aiming to ease the users and to increase productivity. There is a difference between wheelbarrow used at nursery or being used by gardeners if compared to the one that is used at the construction site. Thewheelbarrow used by the gardener has a smaller and shallower tray bucket to lodge the load while the wheelbarrow used at the construction site has a deeper tray bucket. Meanwhile, wheelbarrows at the construction sites are also varying depending on their application, such aswheelbarrowwith onewheel and two wheels" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_35_0000321_s40430-021-02964-z-Figure4-1.png", + "original_path": "designv11-35/openalex_figure/designv11_35_0000321_s40430-021-02964-z-Figure4-1.png", + "caption": "Fig. 4 Proposed experimental setup: a main components, b damaged gear. Source: Author", + "texts": [ + " According to [6], this method allows the extraction and obtaining of compact information, being able to quantify the energy values, relating them the frequency of a fault. Depending on the phenomenon to be studied, there will be a need for scale harmonization through the normalization of the entropy values (WPEm)N. A widely used method consists of linearly normalizing the data between [0, 1] by means of the maximum (WPEmmax) and minimum (WPEmmin) values according to Eq.\u00a020. The experimental setup presents in Fig.\u00a04a consists of a Siemens\u2122 three-phase Induction Motor of 2\u00a0hp (1.5\u00a0kW), 8.6\u00a0Nm, 4 poles, 3.64\u00a0A/380\u00a0V. The motor can be driven with a WEG\u2122 CFW08 vector control frequency inverter (2\u00a0hp/220\u00a0V) or direct start (380\u00a0V/60\u00a0Hz). The gearbox present in Fig.\u00a04b is made up of spur cylindrical gears of module 3 and a gear ratio of 2:1. The drive gear has 23 teeth, tooth height of 675\u00a0mm and pressure angle of 20\u00b0. The motor drives the gearbox which is coupled to a mechanical brake by means of a toothed belt. To obtain surface wear, the load imposed on the electric motor was approximately 60% of the nominal. This is an orientation suggested by [22]. After the formation of the fault, the system was activated in the other loads. (18)WPEm = \u2212pm ln(pm) (19)WPEtot = \u2211 m WPEm (20)(WPEm)N = WPEm \u2212WPEm min WPEm max \u2212WPEm min The frequency inverter does not follow the motor\u2019s nominal characteristics" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_35_0001419_978-94-007-7401-8-Figure2.114-1.png", + "original_path": "designv11-35/openalex_figure/designv11_35_0001419_978-94-007-7401-8-Figure2.114-1.png", + "caption": "Fig. 2.114 4PaRPaRR-type fully-parallel PM with coupled Sch\u00f6nflies motions defined by MF \u00bc SF \u00bc 4; RF\u00f0 \u00de \u00bc \u00f0v1; v2; v3;xd\u00de, TF = 0, NF = 26, limb topology Pa||R\\Pa\\kR\\R", + "texts": [], + "surrounding_texts": [ + "168 2 Fully-Parallel Topologies with Coupled Sch\u00f6nflies Motions", + "2.2 Topologies with Complex Limbs 169", + "170 2 Fully-Parallel Topologies with Coupled Sch\u00f6nflies Motions" + ] + }, + { + "image_filename": "designv11_35_0000395_978-3-030-67750-3_16-Figure11.3-1.png", + "original_path": "designv11-35/openalex_figure/designv11_35_0000395_978-3-030-67750-3_16-Figure11.3-1.png", + "caption": "Fig. 11.3 Proposed layout", + "texts": [ + " In this proposed layout the number of workers is reduced to three persons to operate the production system. This proposed layout has reduced the workers and the material\u2019s trolley has been redesigned and changed to a location which is to near the conveyor A. The function of worker 1 is to transfer the part from conveyor A direct to the material\u2019s trolley. Then, worker 2 is stamping the product\u2019s logo to the part. While worker 3 requires to transfer the part to pallet into the next workstation. The time required and line balancing for this proposed workstation is as follows (Fig. 11.3): Tare weight (s) = 6.862 s Logo stamp (s) = 2.451 s Worker = 3 Line balancing = (6.862+ 2.451)\u00d7 100% (6.862)\u00d7 (3) Line balancing = (931.3) (20.6) Line balancing = 45.2% A comparison of the results before and after the improvement process is shown in Fig. 11.4. It compares from the existing layout of the company and the proposed layout in this research. Figure 11.4 indicates that improvements occur at bottleneck stations which is the tare weight station. By that, the objective of the research has been achieved by minimizing the cycle time" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_35_0000686_j.mechmachtheory.2021.104433-Figure26-1.png", + "original_path": "designv11-35/openalex_figure/designv11_35_0000686_j.mechmachtheory.2021.104433-Figure26-1.png", + "caption": "Fig. 26. The configurations of the equivalent metamorphic clamping mechanism.", + "texts": [ + " Thus, the novel metamorphic clamping equivalent mechanism has mobility 2. For the equivalent metamorphic clamping mechanism, the claw is the output link and its motions are described as follows: (1) The claw rotates about the zO-axis with the angular velocity \u03c9A. (2) The claw with lead screw translates along xO-axis with the velocity vF driven by the helical joint F. According to the constraint condition and configurations of the equivalent mechanism of the epicyclic bevel gear train, the motion modes of the clamping mechanism can be realized in Fig. 26. H. Yuan et al. Mechanism and Machine Theory 166 (2021) 104433 When the constraint condition is Eq. (56), \u03c9A = \u03c9B,\u03c9C1 = \u03c9C2 = \u03c9C3 = \u03c9D1 = \u03c9D2 = \u03c9E = vF = 0 (56) the links lAC, lCD, lDE, lEF, lBE are relatively static, and the equivalent mechanism of the epicyclic bevel gear train is in the single-mobility configuration as shown in Fig. 26(a). Meanwhile, the mobility of the equivalent metamorphic clamping mechanism degenerates from 2 to 1, the motion of it is rotating the axis of joint A(B). The motion-screw system becomes .Eq. (57) { TA = [ 0 0 \u03c9A 0 0 0 ]T (57) When \u03c9A \u2215= \u03c9B, the mobility of the equivalent metamorphic clamping mechanism is 2. And the motion-screw system becomes Eq. (58). { TA = [ 0 0 \u03c9A 0 0 0 ]T TF2 = [ 0 0 0 vF 0 0 ]T (58) Considering the clamping process, the rotation of the link lAE is limited by the target object (\u03c9A= 0) so that the motion of the equivalent metamorphic clamping mechanism translates along with link lOE" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_35_0000395_978-3-030-67750-3_16-Figure20.2-1.png", + "original_path": "designv11-35/openalex_figure/designv11_35_0000395_978-3-030-67750-3_16-Figure20.2-1.png", + "caption": "Fig. 20.2 Electrical braking system (EBS)", + "texts": [ + " A disadvantage age of the current technology is the captured image signal that is influenced by environmental brightness. In 2004 a commercial version of the automatic parking assistance was introduced by Toyota Motor Corporation in the Toyota Prius. Lexus 2007 LS also has an advanced parking guidance system. In automatic parking, there are two requiredmovements, which are the vehiclemovement (in reverse mode) and the steering angle. These movements are not discussed in this paper. The electrohydraulic brake shown in Fig. 20.2 is important for automatic reverse parking because the braking effect helps to decelerate the vehicle. The system has an electric motor and when operated can replace the action of manual brake by using the feet. The speed of car in the reverse mode automatic parking is determined by the propulsion motor and brake system. The brake system is used to decelerate, and stop at a fixed parking before the final hand brake action is applied. The brake force needs to be applied to the disc brake while controlling the motor speed so as to create the required deceleration effect" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_35_0000512_s42417-021-00289-8-Figure2-1.png", + "original_path": "designv11-35/openalex_figure/designv11_35_0000512_s42417-021-00289-8-Figure2-1.png", + "caption": "Fig. 2 Coordinate systems of ball bearing", + "texts": [ + " The outer ring is immobilized and does not have any DOF. Two bearings can be identical or asymmetric ones. The basic parameters of structure, operation and load can be different. This ball bearing\u2013elastic rotor system is a basic model and can be used in widespread conditions. In a ball bearing, the relative position and sliding speed of each part (including ball, ring and cage) are the basis for obtaining their interactive force. Therefore, it is necessary to establish different coordinate systems (Fig.\u00a02) to express multiple motion state and mutual forces for each part conveniently and accurately: (1) Inertial coordinate system: origin O is set at the mass center of outer ring which is stationary when the angular-contact ball bearings are working; axis X coincides with the central axis of bearing; axis Z is the opposite 1 3 direction of gravity; axis Y is determined according to the right hand rule. (2) Body-fixed coordinate systems: BXbYbZb, RXrYrZr and CXcYcZc are the body-fixed coordinate systems for ball, inner ring and cage; B, R and C are the origins which are fixed on the mass center of each part; each axis directions coincide with the coordinate axis of inertial coordinate system, and then they are translated and rotated together with the corresponding part bodies" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_35_0001419_978-94-007-7401-8-Figure3.25-1.png", + "original_path": "designv11-35/openalex_figure/designv11_35_0001419_978-94-007-7401-8-Figure3.25-1.png", + "caption": "Fig. 3.25 Overactuated PMs with coupled Sch\u00f6nflies motions of types 4CRP (a) and 4CPR (b) defined by MF \u00bc SF \u00bc 4; RF\u00f0 \u00de \u00bc v1; v2; v3;xd\u00f0 \u00de, TF = 0, NF = 6, limb topology C||R\\P (a) and C\\P\\kR (b)", + "texts": [ + "21a) R||R||C (Fig. 3.2a) The cylindrical joints of the four limbs have parallel axes 2. 4RCR (Fig. 3.21b) R||C||R (Fig. 3.2b) Idem No. 1 3. 4PCP (Fig. 3.22a) P\\C\\P (Fig. 3.2c) Idem No. 1 4. 4CPP (Fig. 3.22b) C\\P\\\\P (Fig. 3.2d) Idem No. 1 5. 4PPC (Fig. 3.22c) P\\P\\\\C (Fig. 3.2e) Idem No. 1 6. 4PCR (Fig. 3.23a) P\\C||R (Fig. 3.2f) Idem No. 1 7. 4RCP (Fig. 3.23b) R||C\\P (Fig. 3.2g) Idem No. 1 8. 4RPC (Fig. 3.24a) R\\P\\kC (Fig. 3.2h) Idem No. 1 9. 4PRC (Fig. 3.24b) P\\R||C (Fig. 3.2i) Idem No. 1 10. 4CRP (Fig. 3.25a) C||R\\P (Fig. 3.2j) Idem No. 1 11. 4CPR (Fig. 3.25b) C\\P\\kR (Fig. 3.2k) Idem No. 1 12. 4RPC (Fig. 3.26a) R\\P\\kC (Fig. 3.2i) Idem No. 1 13. 4CRR (Fig. 3.26b) C||R||R (Fig. 3.2m) Idem No. 1 14. 4RRRRR (Fig. 3.27a) R\\R||R\\R||R (Fig. 3.3a) The first and the last revolute joints of the four limbs have parallel axes 15. 4RRRRR (Fig. 3.27b) R||R\\R||R\\R (Fig. 3.3b) Idem No. 14 16. 4RRRRR (Fig. 3.28a) R||R||R\\R||R (Fig. 3.3c) The first revolute joints of the four limbs have parallel axes 17. 4RRRRR (Fig. 3.28b) R||R\\R||R||R (Fig. 3.3d) The last revolute joints of the four limbs have parallel axes 18" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_35_0001604_tai.1958.6369020-Figure9-1.png", + "original_path": "designv11-35/openalex_figure/designv11_35_0001604_tai.1958.6369020-Figure9-1.png", + "caption": "Fig. 9. Case 3: e; / \u2014e trajectory", + "texts": [ + " Lewis reports a second-order nonlinear feedback control system with multiplicative term cc\\ where c is the output and c' its derivative.6 The present third-order control system is hence a further extension of the second-order nonlinear feedback control system. Self-Oscillations and Limit Cycle Case 3 represents a peculiar system where self-oscillation occurs. Fig. 8 shows the e'\u2014e projection of a 3-dimensional space trajectory starting at e = 0.5 and e' = 0 and reaching a limit cycle in the form of a circle with emax/=: : t0.35 and emax==b0.35 approximately. Fig. 9 shows the e\"\u2014e projection of the same space trajectory starting at e = 0.5 and e\" = 0 and reaching a limit cycle with e m a x \"=\u00b10.35 and emax=\u00b10.35 approximately. I t is to be noted that the two projections reaching the form of limit cycles are really the projections of a 3- dimensional limit cycle in phase space, as are the projections of a general space trajectory discussed in reference 7. The authors believe that the extension of the limit cycle in a phase plane to the limit cycle in a phase space is a new contribution and hence of interest in the stability study of high-order control systems" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_35_0000535_tmag.2021.3085750-Figure6-1.png", + "original_path": "designv11-35/openalex_figure/designv11_35_0000535_tmag.2021.3085750-Figure6-1.png", + "caption": "Fig. 6 Frozen permeability analysis results. (a) Remagenetization. (b) Demagnetization.", + "texts": [ + " The blueline blocks denote the specification and algorism of the hysteresis characteristics of the LCF PMs, while the yellow blocks refer to the conventional FE process. In order to investigate the reason for balanced bidirectional magnetization effect, the frozen permeability method [10] is adopted in the FE simulation to study the influence of NdFeB PM and id to the MS manipulation process respectively. The HMC-VFMM is firstly demagnetized with sorely PMs, and then id current pulses are added. Fig. 6(a) illustrates the remagnetization field of id. For the initial flux-weakened state, the parallel branch of HCF PM form a magnetic circuit with the negative LCF PM, and the series branch hardly provide magnetic field for the LCF. When id excitement is added, the HCF PM of the series branch helps to magnetize the LCF PM Authorized licensed use limited to: San Francisco State Univ. Downloaded on July 03,2021 at 08:24:04 UTC from IEEE Xplore. Restrictions apply. 0018-9464 (c) 2021 IEEE. Personal use is permitted, but republication/redistribution requires IEEE permission. See http://www.ieee.org/publications_standards/publications/rights/index.html for more information. BJ-13 4 to generate positive magnetic field. Similar phenomenon can be found in the demagnetization process shown in Fig. 6(b). When -id is applied, the HCF PM of the parallel branch contribute to the demagnetization of the LCF PM and the magnetic field of series branch HCF PM is short-circuited. The FE results confirm the balanced magnetization effect. The back-EMF and current pulse waveforms for a bidirectional magnetization process are plotted in Fig. 7(a). In the second and fourth electric period, positive and negative current pulses are applied respectively, i.e., remagnetization and demagnetization processes are conducted successively" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_35_0001419_978-94-007-7401-8-Figure3.86-1.png", + "original_path": "designv11-35/openalex_figure/designv11_35_0001419_978-94-007-7401-8-Figure3.86-1.png", + "caption": "Fig. 3.86 Overactuated PMs with coupled Sch\u00f6nflies motions of types 4CPPa (a) and 4PaPC (b) defined by MF \u00bc SF \u00bc 4; RF\u00f0 \u00de \u00bc v1; v2; v3;xd\u00f0 \u00de, TF = 0, NF = 18, limb topology C\\P\\kPa (a) and Pa\\P\\kC (b)", + "texts": [ + "82a) Pa||P\\R\\P (Fig. 3.50t0) Idem No. 23 48. 4PaPRP (Fig. 3.82b) Pa\\P\\\\R\\P (Fig. 3.50u0) Idem No. 23 49. 4CPaP (Fig. 3.83a) C||Pa\\P (Fig. 3.50v0) The cylindrical joints of the four limbs have parallel directions 50. 4PaCP (Fig. 3.83b) Pa||C\\P (Fig. 3.50w0) Idem No. 49 51. 4CRPa (Fig. 3.84a) C||R||Pa (Fig. 3.50x0) Idem No. 49 52. 4PCPa (Fig. 3.84b) P\\C||Pa (Fig. 3.50y0) Idem No. 49 53. 4PPaC (Fig. 3.85a) P\\Pa||C (Fig. 3.50z0) Idem No. 49 54. 4RCPa (Fig. 3.85b) R||C||Pa (Fig. 3.50z01) Idem No. 49 55. 4CPPa (Fig. 3.86a) C\\P\\kPa (Fig. 3.50a0 0) Idem No. 49 56. 4PaPC (Fig. 3.86b) Pa\\P\\kC (Fig. 3.50b0 0) Idem No. 49 57. 4PaRC (Fig. 3.87a) Pa||R||C (Fig. 3.50c0 0) Idem No. 49 58. 4PaCR (Fig. 3.87b) Pa\\C\\P (Fig. 3.50e0 0) Idem No. 49 59. 4PaCR (Fig. 3.88) Pa||C||R (Fig. 3.50d0 0) Idem No. 49 Table 3.6 Limb topology and connecting conditions of the overactuated solutions with no idle mobilities presented in Figs. 3.89, 3.90, 3.91, 3.92, 3.93, 3.94, 3.95, 3.96, 3.97, 3.98, 3.99, 3.100, 3.101, 3.102, 3.103, 3.104, 3.105, 3.106, 3.107, 3.108, 3.109, 3.110, 3.111, 3.112, 3.113, 3" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_35_0001419_978-94-007-7401-8-Figure3.69-1.png", + "original_path": "designv11-35/openalex_figure/designv11_35_0001419_978-94-007-7401-8-Figure3.69-1.png", + "caption": "Fig. 3.69 Overactuated PMs with coupled Sch\u00f6nflies motions of types 4PaPPR (a) and 4PaPPR (b) defined by MF \u00bc SF \u00bc 4; RF\u00f0 \u00de \u00bc v1; v2; v3;xd\u00f0 \u00de, TF = 0, NF = 18, limb topology Pa||P\\P\\kR (a) and Pa\\P\\kP||R (b)", + "texts": [ + "63a) P\\P||Pa||R (Fig. 3.50r) Idem No. 1 19. 4PPaRR (Fig. 3.63b) P||Pa||R||R (Fig. 3.50s) Idem No. 1 20. 4PPaRP (Fig. 3.64) P\\Pa||R||P (Fig. 3.50t) Idem No. 1 21. 4PPaPR (Fig. 3.65) P\\Pa||P||R (Fig. 3.50u) Idem No. 1 22. 4PPaPR (Fig. 3.66a) P||Pa\\P\\kR (Fig. 3.50v) Idem No. 1 23. 4PPaRP (Fig. 3.66b) P||Pa\\R\\P (Fig. 3.50i0) The before last joints of the four limbs have parallel axes 24. 4RPPaP (Fig. 3.67) R||P||Pa\\P (Fig. 3.50w) Idem No. 1 25. 4RPRPa (Fig. 3.68) R||P||R||Pa (Fig. 3.50x) Idem No. 1 26. 4PaPPR (Fig. 3.69a) Pa||P\\P\\kR (Fig. 3.50y) Idem No. 1 27. 4PaPPR (Fig. 3.69b) Pa\\P\\kP||R (Fig. 3.50z) Idem No. 1 28. 4PaPRP (Fig. 3.70) Pa\\P\\kR||P (Fig. 3.50z1) Idem No. 1 29. 4PaRRP (Fig. 3.71) Pa||R||R||P (Fig. 3.50a0) Idem No. 1 30. 4PaRPR (Fig. 3.72) Pa||R||P||R (Fig. 3.50b0) Idem No. 1 31. 4PaPRP (Fig. 3.73a) Pa||P||R\\P (Fig. 3.50c0) Idem No. 1 32. 4PaRPP (Fig. 3.73b) Pa||R\\P\\kP (Fig. 3.50d0) Idem No. 1 33. 4PRPPa (Fig. 3.74a) P\\R\\P||Pa (Fig. 3.50e0) The second joints of the four limbs have parallel axes 34. 4PRPPa (Fig. 3.74b) P\\R\\P\\Pa (Fig. 3.50f0) Idem No. 33 35" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_35_0001419_978-94-007-7401-8-Figure5.67-1.png", + "original_path": "designv11-35/openalex_figure/designv11_35_0001419_978-94-007-7401-8-Figure5.67-1.png", + "caption": "Fig. 5.67 3PaPaPaR-1RPPaPa-type fully-parallel PM with uncoupled Sch\u00f6nflies motions defined by MF = SF = 4, RF\u00f0 \u00de \u00bc \u00f0v1; v2; v3;xb\u00de, TF = 0, NF = 39, limb topology Pa\\Pa||Pa\\kR, Pa\\Pa||Pa||R and R||P||Pa||Pa", + "texts": [ + "4k) (continued) 5.1 Fully-Parallel Topologies 533 Table 5.4 (continued) No. PM type Limb topology Connecting conditions 30. 3PaPaPaR-1RPaPaP (Fig. 5.64) Pa\\Pa||Pa\\\\R (Fig. 5.6c) Idem No. 1 Pa\\Pa||Pa||R (Fig. 5.6b) R||Pa||Pa||P (Fig. 5.54k) 31. 3PaPaPaR1RPaPatP (Fig. 5.65) Pa\\Pa||Pa\\kR (Fig. 5.6a) Idem No. 1 Pa\\Pa||Pa||R (Fig. 5.6b) R||Pa||Pat||P (Fig. 5.4l) 32. 3PaPaPaR1RPaPatP (Fig. 5.66) Pa\\Pa||Pa\\\\R (Fig. 5.6c) Idem No. 1 Pa\\Pa||Pa||R (Fig. 5.6b) R||Pa||Pat||P (Fig. 5.54l) 33. 3PaPaPaR-1RPPaPa (Fig. 5.67) Pa\\Pa||Pa\\kR (Fig. 5.6a) The revolute joint between links 7D and 11 of G4 limb and the last revolute joints of limbs G1, G2 and G3 have superposed axes Pa\\Pa||Pa||R (Fig. 5.6b) R||P||Pa||Pa (Fig. 5.4m) 34. 3PaPaPaR-1RPPaPa (Fig. 5.68) Pa\\Pa||Pa\\\\R (Fig. 5.6c) Idem No. 33 Pa\\Pa||Pa||R (Fig. 5.6b) R||P||Pa||Pa (Fig. 5.54m) Table 5.5 Limb topology and connecting conditions of the fully-parallel solutions with no idle mobilities presented in Figs. 5.69, 5.70, 5.71, 5.72, 5.73, 5.74, 5.75, 5.76, 5.77, 5" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_35_0001419_978-94-007-7401-8-Figure6.6-1.png", + "original_path": "designv11-35/openalex_figure/designv11_35_0001419_978-94-007-7401-8-Figure6.6-1.png", + "caption": "Fig. 6.6 Maximaly regular fully-parallel PMs with Sch\u00f6nflies motions of types 2CPRR-1CPR1RUPU (a) and 2CRRR-1CRR-1RUPU (b) defined by MF = SF = 4, (RF) = (v1; v2; v3;xb), TF = 0, NF = 2, limb topology R\\R\\R\\P\\||R\\R and C\\P\\||R\\R, C\\P\\||R (a), C||R||R\\R, C||R||R (b)", + "texts": [ + " rG4 0 0 19. MGi (i = 1, 3) 4 5 20. MG2 4 4 21. MG4 6 6 22. (RF) (v1; v2; v3;xb) (v1; v2; v3;xb) 23. SF 4 4 24. rl 0 0 25. rF 14 16 26. MF 4 4 27. NF 4 2 28. TF 0 0 29. Pp1 j\u00bc1 fj 4 5 30. Pp2 j\u00bc1 fj 4 4 31. Pp3 j\u00bc1 fj 4 5 32. Pp4 j\u00bc1 fj 6 6 33. Pp j\u00bc1 fj 18 20 a See footnote of Table 2.2 for the nomenclature of structural parameters 6.1 Fully-Parallel Topologies with Simple Limbs 589 Table 6.4 Structural parametersa of parallel mechanisms in Figs. 6.5 and 6.6 No. Structural parameter Solution Figure 6.5 Figure 6.6 1. m 13 15 2. pi (i = 1,3) 4 4 3. p2 3 3 4. p4 4 6 5. p 15 17 6. q 3 3 7. k1 4 4 8. k2 0 0 9. k 4 4 10. (RG1) (v1; v2; v3;xa;xb) (v1; v2; v3;xa;xb) 11. (RG2) (v1; v2; v3;xb) (v1; v2; v3;xb) 12. (RG3) (v1; v2; v3;xb;xd) (v1; v2; v3;xb;xd) 13. (RG4) (v1; v2; v3;xb) (v1; v2; v3;xa;xb;xd) 14. SGi (i = 1, 3) 5 5 15. SG2 4 4 16. SG4 4 6 17. rGi (i = 1, 2, 3) 0 0 18. rG4 0 0 19. MGi (i = 1, 3) 5 5 20. MG2 4 4 21. MG4 4 6 22. (RF) (v1; v2; v3;xb) (v1; v2; v3;xb) 23. SF 4 4 24. rl 0 0 25. rF 14 16 26. MF 4 4 27" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_35_0000472_s00170-021-07362-2-Figure17-1.png", + "original_path": "designv11-35/openalex_figure/designv11_35_0000472_s00170-021-07362-2-Figure17-1.png", + "caption": "Fig. 17 NIST IN625 cantilever distortion after build plate removal", + "texts": [], + "surrounding_texts": [ + "parameters. An example is the yield strength of additively manufactured IN718 where it was assumed in finite element distortion models as 1172 MPa in [39] and 308 MPa in [40], but experimentally measured as 849 MPa in [24], 723 MPa in [41], and 710 MPa in [25].\nBesides the yield strength, the material plasticity behavior at higher temperature was also investigated. To date, the material behavior of additively manufactured parts at high temperature is not easily available nor well investigated. Thus, the effect of the softening constant (n) on the Ti6Al4V bridge\nstructure distortion was investigated. It was found that distortion predictions are greatly affected by the temperature dependency of material plastic behavior (Fig. 14b). This behavior was expected as the value of softening constant m greatly affects the yield strength temperature dependency (Fig. 4). Therefore, understanding the material behavior at high temperature for additively manufactured parts is crucial for accurate distortion prediction. As mentioned earlier, the values of softening constants m in the proposed models with different materials were taken from wrought material testing data.", + "In this section, the induced residual stresses and strains, and distortions during the build of the NIST IN625 cantilever are investigated. The initial analysis was performed using N=10 and C=1.75, and the same conditions are shown in Fig. 8. The elastic residual strain results before and after cutting cantilever supports are shown in Fig. 16. Before cutting the cantilever supports, the residual elastic strains in x-direction (\u03b5xx) at the overhangs were found tensile where the maximum value is at the top of the cantilever (Fig. 16a). The residual elastic strains in z-direction (\u03b5zz, build direction) were highly compressive close to the build plate\nand moderately compressive at the overhangs, but highly tensile close to the edges (Fig. 16b). By cutting the cantilever supports through deactivating the build plate elements beneath supports, the residual strains were released and a new equilibrium state, at which the cantilever deforms, will be achieved. These predicted results can find a good agreement with the experimentally measured strains by X-ray diffraction in [18], previous finite element predictions in [42].\nThe distorted NIST cantilever is shown in Figs. 17 and 18, where the predicted maximum vertical displacement is 0.94 mm. The predicted maximum displacement is underestimated compared to the reported experimentally measured value of 1.276 mm (Fig. 18). This deviation, which was previously reported in different previous studies in [6], could be due to", + "the assumption of isotropic and homogenous material properties. The different cooling rates during the build of parts with features may result in a heterogenous microstructure, and the mechanical properties at the parts\u2019 features (e.g., overhangs) are likely to be different. Thus, the predicted distortion result could be improved by allowing higher tensile elastic strains by changing the yield strength as discussed in Section 2.4.\nThe predicted residual stresses in x-direction (\u03c3xx) before cutting supports can also find good agreement with the experimentally measured using the contour method at one of the cantilever legs [18] (Fig. 19). The residual stresses are compressive near the build plate and leg center. However, these residual stresses turn tensile at the overhang reaching their maximum value at the top.\nIn this section, the case of the IN718 canonical part is investigated. Unlike the cantilever case where distortions take place after cutting supports, the distortion for this canonical part takes place during the process. The final displacement solution could be obtained by comparing the final deformed mesh to the initial mesh. The predicted distortion for this part with\nand without adaptive remeshing and at two adaptive remeshing configurations is shown in Fig. 20. This predicted distortion can find an agreement with the experimental solution in [7], though the maximum displacement value is underestimated by 25%. As discussed earlier, the main reason for that deviation might be the assumed yield strength given that microstructure and mechanical properties are different at the thin features. To evaluate the accuracy of using adaptive remeshing, firstly, the distortion solutions at different adaptive remeshing configurations were compared, where it was found that solution slightly changed within the considered adaptive mesh configurations. Secondly, the predicted residual elastic strain distributions achieved with and without adaptive remeshing are slightly different due to the less integration points when mesh is coarsened (Fig. 21). Therefore, a compromise between accuracy and computational time should always be achieved.\nThe solution and computational efficiency using the proposed adaptive remeshing framework is compared to the NetFabb solution with adaptive remeshing and hexahedral mesh in [7] (Table 2). Firstly, the proposed approach with tetrahedral mesh could help in reducing the number of nodes and elements by 50 layers when compared to the NetFabb approach even with less layers (16 layers). Secondly, with using 8 core and 50 layers, the average computational time per step using the current approach is 48 s/step,which is two times longer than that reported byNetFabb when using 26 cores and 16 layers. The computational time can be further improved by using more processor cores and changing the mesh and adaptive remeshing configurations (e.g., number of layers, N and C).\nThe effect of the remeshing frequency, controlled by number of layers between remeshing (N), on the computational time" + ] + }, + { + "image_filename": "designv11_35_0001419_978-94-007-7401-8-Figure3.41-1.png", + "original_path": "designv11-35/openalex_figure/designv11_35_0001419_978-94-007-7401-8-Figure3.41-1.png", + "caption": "Fig. 3.41 Overactuated PMs with coupled Sch\u00f6nflies motions of types 4RPRRP (a) and 4RRPRP (b) defined by MF \u00bc SF \u00bc 4; RF\u00f0 \u00de \u00bc v1; v2; v3;xd\u00f0 \u00de, TF = 0, NF = 2, limb topology R||P||R\\R\\P (a) and R||R||P\\R\\P (b)", + "texts": [ + "37a) R||R\\R||R\\P (Fig. 3.3u) Idem No. 16 35. 4PRRRP (Fig. 3.37b) P\\R\\R||R\\P (Fig. 3.3v) Idem No. 29 36. 4RRPRP (Fig. 3.38a) R\\R||P||R\\P (Fig. 3.3w) Idem No. 16 37. 4RPRRP (Fig. 3.38b) R\\P||R||R\\P (Fig. 3.3x) Idem No. 16 38. 4RRPRP (Fig. 3.39a) R\\R||P||R\\P (Fig. 3.3y) Idem No. 16 39. 4RRRPP (Fig. 3.39b) R\\R||R||P\\P (Fig. 3.3z) Idem No. 16 40. 4PRRRP (Fig. 3.40a) P||R||R\\R\\P (Fig. 3.3z1) The fourth joints of the four limbs have parallel axes 41. 4RPRRP (Fig. 3.40b) R||P||R||R\\P (Fig. 3.3a) Idem No. 40 42. 4RPRRP (Fig. 3.41a) R||P||R\\R\\P (Fig. 3.3b0) Idem No. 40 43. 4RRPRP (Fig. 3.41b) R||R||P\\R\\P (Fig. 3.3c0) Idem No. 40 44. 4RRRPP (Fig. 3.42a) R||R\\R\\P\\\\P (Fig. 3.3d0) The third joints of the four limbs have parallel axes 45. 4PRRPR (Fig. 3.42b) P||R||R\\P\\\\R (Fig. 3.3e0) Idem No. 17 46. 4RPRPR (Fig. 3.43a) R||P||R\\P\\\\R (Fig. 3.3f0) Idem No. 17 47. 4RRPPR (Fig. 3.43b) R||R||P\\P\\\\R (Fig. 3.3g0) Idem No. 17 48. 4RPRPR (Fig. 3.44a) R||P||R\\P\\\\R (Fig. 3.3h0) Idem No. 17 49. 4RPRPR (Fig. 3.44b) P\\P\\kR||R\\\\R (Fig. 3.3i0) Idem No. 17 50. 4PPRRR (Fig. 3.44c) P\\P||R||R\\\\R (Fig. 3.3j0) Idem No" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_35_0001419_978-94-007-7401-8-Figure2.47-1.png", + "original_path": "designv11-35/openalex_figure/designv11_35_0001419_978-94-007-7401-8-Figure2.47-1.png", + "caption": "Fig. 2.47 4RPaPPa-type fully-parallel PM with coupled Sch\u00f6nflies motions defined by MF \u00bc SF \u00bc 4; RF\u00f0 \u00de \u00bc \u00f0v1; v2; v3;xd\u00de, TF = 0, NF = 30, limb topology R||Pa\\P\\\\Pa", + "texts": [ + " PM type Limb topology Connecting conditions 30. 4PRPaPa (Fig. 2.42) P\\R||Pa\\Pa (Fig. 2.22c) Idem No. 5 31. 4RPPaPa (Fig. 2.43) R\\P\\kPa\\\\Pa (Fig. 2.22d) Idem No. 15 32. 4PaPPaR (Fig. 2.44) Pa\\P\\\\Pa\\\\R (Fig. 2.22e) Idem No. 4 33. 4PaPRPa (Fig. 2.45) Pa\\P\\kR\\Pa (Fig. 2.22f) The axes of the revolute joints connecting links 5 and 6 of the four limbs are parallel 34. 4PaRPPa (Fig. 2.46) Pa||R\\P\\\\Pa (Fig. 2.22g) The axes of the revolute joints connecting links 4 and 5 of the four limbs are parallel 35. 4RPaPPa (Fig. 2.47) R||Pa\\P\\\\Pa (Fig. 2.22h) Idem No. 15 36. 4PaPaPR (Fig. 2.48) Pa\\Pa\\\\P\\\\R (Fig. 2.22i) Idem No. 4 37. 4RPaPaR (Fig. 2.49) R\\Pa||Pa\\kR (Fig. 2.22j) The first and the last revolute joints of the four limbs are parallel 38. 4RPaPaR (Fig. 2.50) R\\Pa||Pa\\kR (Fig. 2.22k) Idem No. 37 39. 4RPaRPa (Fig. 2.51) R\\Pa\\kR\\Pa (Fig. 2.22l) Idem No. 15 40. 4PaRPPa (Fig. 2.52) Pa\\R\\P||Pa (Fig. 2.22m) Idem No. 34 41. 4PaPaRP (Fig. 2.53) Pa||Pa\\R\\P (Fig. 2.22n) The axes of the revolute joints connecting links 7 and 8 of the four limbs are parallel 42" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_35_0001419_978-94-007-7401-8-Figure5.89-1.png", + "original_path": "designv11-35/openalex_figure/designv11_35_0001419_978-94-007-7401-8-Figure5.89-1.png", + "caption": "Fig. 5.89 3PaPaPaR-1CPaPa-type fully-parallel PM with uncoupled Sch\u00f6nflies motions defined by MF = SF = 4, RF\u00f0 \u00de \u00bc \u00f0v1; v2; v3;xb\u00de, TF = 0, NF = 39, limb topology Pa\\Pa||Pa\\kR, Pa\\Pa||Pa||R and C||Pa||Pa", + "texts": [ + "5a) The revolute joint between links 6D and 10 of G4 limb and the last revolute joints of limbs G1, G2 and G3 have superposed axes Pa\\Pa||R||R (Fig. 5.4i) C||Pa||Pa (Fig. 5.40) 24. 2PaPaRRR1PaPaRR1CPaPa (Fig. 5.86) Pa\\Pa||R||R\\kR (Fig. 5.5b) Idem No. 23 Pa\\Pa||R||R (Fig. 5.4i) C||Pa||Pa (Fig. 5.40) 25. 2PaPaRRR1PaPaRR1CPaPat (Fig. 5.87) Pa\\Pa||R||R\\\\R (Fig. 5.5a) Idem No. 23 Pa\\Pa||R||R (Fig. 5.4i) C||Pa||Pat (Fig. 5.4p) 26. 2PaPaRRR1PaPaRR1CPaPat (Fig. 5.88) Pa\\Pa||R||R\\kR (Fig. 5.5b) Idem No. 23 Pa\\Pa||R||R (Fig. 5.4i) C||Pa||Pat (Fig. 5.4p) 27. 3PaPaPaR1CPaPa (Fig. 5.89) Pa\\Pa||Pa\\kR (Fig. 5.6a) The revolute joint between links 6D and 11 of G4 limb and the last revolute joints of limbs G1, G2 and G3 have superposed axes Pa\\Pa||Pa||R (Fig. 5.6b) C||Pa||Pa (Fig. 5.4o) 28. 3PaPaPaR1CPaPa (Fig. 5.90) Pa\\Pa||Pa\\\\R (Fig. 5.6c) Idem No. 27 Pa\\Pa||Pa||R (Fig. 5.6b) C||Pa||Pa (Fig. 5.54o) 29. 3PaPaPaR1CPaPat (Fig. 5.91) Pa\\Pa||Pa\\kR (Fig. 5.6a) Idem No. 27 Pa\\Pa||Pa||R (Fig. 5.6b) C||Pa||Pat (Fig. 5.4p) 30. 3PaPaPaR1CPaPat (Fig. 5.92) Pa\\Pa||Pa\\\\R (Fig. 5.6c) Idem No. 27 Pa\\Pa||Pa||R (Fig" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_35_0001419_978-94-007-7401-8-Figure5.97-1.png", + "original_path": "designv11-35/openalex_figure/designv11_35_0001419_978-94-007-7401-8-Figure5.97-1.png", + "caption": "Fig. 5.97 Redundantly actuated PMs with uncoupled Sch\u00f6nflies motions of types 1PPPR-2PPC1PPPC (a) and 1CPR-2CPRR-1PCPRR (b) defined by MF = 5, SF = 4, RF\u00f0 \u00de \u00bc \u00f0v1; v2; v3;xa\u00de, TF = 1, NF = 6 (a), NF = 3 (b), limb topology P\\P\\\\P\\\\R, P\\P\\\\C, P||P\\P\\\\C (a) and C\\P\\kR, C\\P\\kR\\R, P|| C\\P\\kR\\R (b)", + "texts": [ + " 1PPRR2PPRRR1PPPRR (Fig. 5.95a) P\\P\\kR||R (Fig. 4.1d) Idem No. 1 P\\P\\kR||R\\R (Fig. 4.2a) P||P\\P\\kR||R\\R (Fig. 5.93d) 4. 1PRPR2PRPRR1PPRPRR (Fig. 5.95b) P||R\\P\\kR (Fig. 4.1e) Idem No. 1 P||R\\P\\kR\\R (Fig. 4.2b) P||P||R\\P\\kR\\R (Fig. 5.93e) 5. 1PRRP2PRRPR1PPRRPR (Fig. 5.96a) P||R||R\\P (Fig. 4.1f) Idem No. 1 P||R||R\\P\\\\R (Fig. 4.2c) P||P||R||R\\P\\\\R (Fig. 5.93c) 6. 1PRRR2PRRRR1PPRRRR (Fig. 5.96b) P||R||R||R (Fig. 4.1 g) Idem No. 1 P||R||R||R\\R (Fig. 4.2d) P||P||R||R||R\\R (Fig. 5.93f) 7. 1PPPR-2PPC1PPPC (Fig. 5.97a) P\\P\\\\P\\\\R (Fig. 4.1b) Idem No. 1 P\\P\\\\C (Fig. 4.1 h) P||P\\P\\\\C (Fig. 5.93g) 8. 1CPR-2CPRR1PCPRR (Fig. 5.97b) C\\P\\kR (Fig. 4.1i) Idem No. 1 C\\P\\kR\\R (Fig. 4.2e) P|| C\\P\\kR\\R (Fig. 5.93i) 9. 1CRP-2CRPR1PCRPR (Fig. 5.98a) C||R\\P (Fig. 4.1j) Idem No. 1 C||R\\P\\\\R (Fig. 4.2f) P||C||R\\P\\\\R (Fig. 5.93h) 10. 1CRR-2CRRR1PCRRR (Fig. 5.98b) C||R||R (Fig. 4.1 k) Idem No. 1 C||R||R\\R (Fig. 4.2 g) P||C||R||R\\R (Fig. 5.93j) 552 5 Topologies with Uncoupled Sch\u00f6nflies Motions Table 5.14 Structural parametersa of parallel mechanisms in Figs. 5.94, 5.95, 5.96 No. Structural parameter Solution Figure 5.94 Figures 5.95 and 5", + "2 Redundantly Actuated Topologies 553 In the redundantly actuated topologies of PMs with uncoupled Sch\u00f6nflies motions F / G1-G2-G3-G4 presented in this section, the moving platform n : nGi (i = 1, 2, 3, 4) is connected to the reference platform 1 : 1Gi : 0 by four spatial complex limbs with four or five degrees of connectivity. Two linear actuators are combined in the two first joints of G4-limb. Limbs G1, G2 and G3 are actuated by one linear motor mounted on the fixed base. Table 5.15 Structural parametersa of parallel mechanisms in Figs. 5.97 and 5.98 No. Structural parameter Solution Figure 5.97a Figure 5.97b and 5.98 1. m 12 14 2. p1 3 3 3. pi (i = 2, 3) 3 4 4. p4 5 5 5. p 14 16 6. q 3 3 7. k1 4 4 8. k2 0 0 9. k 4 4 10. (RG1) (v1; v2; v3;xa) (v1; v2; v3;xa) 11. (RG2) (v1; v2; v3;xa) (v1; v2; v3;xa;xb) 12. (RG3) (v1; v2; v3;xa) (v1; v2; v3;xa;xd) 13. (RG4) (v1; v2; v3;xa) (v1; v2; v3;xa;xd) 14. SG1 4 4 15. SGi (i = 2, 3) 4 5 16. SG4 4 5 17. rGi (i = 1,\u2026,4) 0 0 18. MG1 4 4 19. MGi (i = 2, 3) 4 5 20. MG4 5 6 21. (RF) (v1; v2; v3;xa) (v1; v2; v3;xa) 22. SF 4 4 23. rl 0 0 24. rF 12 15 25. MF 5 5 26. NF 6 3 27" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_35_0001419_978-94-007-7401-8-Figure2.9-1.png", + "original_path": "designv11-35/openalex_figure/designv11_35_0001419_978-94-007-7401-8-Figure2.9-1.png", + "caption": "Fig. 2.9 Fully-parallel PMs with coupled Sch\u00f6nflies motions of types 4RRPRR (a) and 4RRRPR (b) defined by MF \u00bc SF \u00bc 4; RF\u00f0 \u00de \u00bc \u00f0v1; v2; v3;xd\u00de, TF = 0, NF = 2, limb topology R||R||P||R\\R (a) and R||R||R||P\\R (b)", + "texts": [ + "5b) R||R||P\\R||R (Fig. 2.1g) Idem No. 5 9. 4PRRRR (Fig. 2.6a) P\\R\\R||R\\R (Fig. 2.1h) The second and the last joints of the four limbs have parallel axes 10. 4PRRRR (Fig. 2.6b) P||R\\R||R\\R (Fig. 2.1i) Idem No. 9 11. 4RRPRR (Fig. 2.7a) R\\R\\P\\kR\\R (Fig. 2.1j) Idem No. 1 12. 4PRRRR (Fig. 2.7b) P\\R||R\\R||R (Fig. 2.1k) Idem No. 3 13. 4PRRRR (Fig. 2.8a) P||R||R||R||R\\R (Fig. 2.1l) The last revolute joints of the four limbs have parallel axes 14. 4RPRRR (Fig. 2.8b) R||P||R||R\\R (Fig. 2.1m) Idem No. 13 15. 4RRPRR (Fig. 2.9a) R||R||P||R\\R (Fig. 2.1n) Idem No. 13 16. 4RRRPR (Fig. 2.9b) R||R||R||P\\R (Fig. 2.1o) Idem No. 13 17. 4RPRRR (Fig. 2.10a) R||P||R||R\\R (Fig. 2.1p) Idem No. 13 18. 4RRPRR (Fig. 2.10b) R||R||P||R\\R (Fig. 2.1q) Idem No. 13 19. 4RRRRP (Fig. 2.11a) R||R||R\\R\\P (Fig. 2.1r) The before last revolute joints of the four limbs have parallel axes 20. 4RPRRR (Fig. 2.11b) R\\P\\kR\\R||R (Fig. 2.1s) Idem No. 3 21. 4RRRRP (Fig. 2.12a) R\\R||R||R||P (Fig. 2.1t) The first revolute joints of the four limbs have parallel axes 22. 4RRRPR (Fig. 2.12b) R\\R||R||P||R Fig. 2.1u) Idem No" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_35_0001411_978-3-030-55061-5-Figure16-1.png", + "original_path": "designv11-35/openalex_figure/designv11_35_0001411_978-3-030-55061-5-Figure16-1.png", + "caption": "Fig. 16. Translational actuator upon X axis Fig. 17. Translational actuator upon Y axis", + "texts": [ + " 9, whose sides form parallelograms, the calculation of mobility by the segmentation method in detail is presented in works [1] and [3]. For the first independent cycle (Fig. 10) the mobility number is 3. The simulation of the three degrees of freedom of the end-effector in the first cycle is presented in Figs. 11, 12 and 13. Fig. 9. Parallelogram mechanism modeled with ADAMS For the second cycle (Fig. 14) the mobility number is 2. A rotation around the Z axis will be possible (Fig. 15) and the two independent translations, X (Fig. 16) and Y (Fig. 17), will appear as impossible (for example in Fig. 18). the base and an actuator was placed along the X axis, between the cross element and the base (Fig. 19). Following the simulation, it is observed visually (Fig. 20), but also in the superimposed kinematic diagrams (Fig. 21), the existence of a degree of freedom and the appearance of two dependent movements, translation along the X and Y axis. Computer-Assisted Learning Used to Overconstrained Mechanism\u2019s Mobility 525 The authors use computer-assisted learning/teaching based on the ADAMS software and develop a new teaching method for calculating the mobility of overconstrained mechanismswith a structural formula" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_35_0000776_s10854-021-06832-3-Figure7-1.png", + "original_path": "designv11-35/openalex_figure/designv11_35_0000776_s10854-021-06832-3-Figure7-1.png", + "caption": "Fig. 7 Fabrication process of the enzyme-free glucose sensors: a preparation of the silicon wafer and cleaning; b thermal oxidation, LPCVD Si3N4, lithography, sputtering Pt layer, and lift-off; c lithography, sputtering Ni layer and lift-off; d annealing and wet etching", + "texts": [ + " The working electrode is composed of bimetallic PtNi material, and glucose can be oxidized on the surface of the electrode. The active areas of the working electrodes are designed to be 10 9 10 mm2, 5 9 5 mm2, and 1 9 1 mm2, respectively. The counter electrode is made of Pt, with the width of 400 lm. Si3N4 is served as an insulating layer between the working electrode and counter electrode. The enzyme-free glucose sensors based on bimetallic PtNi materials were fabricated by MEMS techniques, as shown in Fig. 7. An antimony-doped 1\u201315 X cm n-type (100)-oriented silicon wafer was used to process enzyme-free glucose sensors. Thermal oxide (100 nm) and LPCVD Si3N4 (160 nm) were deposited on both sides of the wafer and served as insulating layer. Ti (20 nm) layer and Pt layer (100 nm) were sputtered on the substrate and patterned Fig. 5 CV curves of bimetallic PtNi600 at different glucose concentrations in the range of 1\u20137 mM (a) and at different scan rates ranging from 10 to 100 mV s-1 (c) in the presence of 1" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_35_0001419_978-94-007-7401-8-Figure6.25-1.png", + "original_path": "designv11-35/openalex_figure/designv11_35_0001419_978-94-007-7401-8-Figure6.25-1.png", + "caption": "Fig. 6.25 3PPaPaR-1RPPP-type maximaly regular fully-parallel PM with Sch\u00f6nflies motions defined by MF = SF = 4, (RF) = (v1; v2; v3;xb), TF = 0, NF = 24, limb topology P||Pa||Pa\\R, P||Pa||Pa||R and R\\P ??P ??P", + "texts": [ + " The actuated prismatic joints of limbs G1, G2 and G3 have orthogonal directions (continued) 614 6 Maximally Regular Topologies with Sch\u00f6nflies Motions Table 6.6 (continued) No. PM type Limb topology Connecting conditions 14. 3PPPaR-1RUPU (Fig. 6.23b) P ?P ?||Pa\\\\R (Fig. 4.8a) P ?P ?||Pa||R, (Fig. 4.8c) R ?R ?R ?P ?||R ?R (Fig. 5.1b) Idem no. 13 15. 3PPaPR-1RUPU (Fig. 6.24a) P||Pa ?P\\\\R (Fig. 4.8d) P||Pa ?P ?||R (Fig. 4.8e) R ?R ?R ?P ?||R ?R (Fig. 5.1b) Idem no. 13 16. 2PPaRRR-1PPaRR-1RUPU (Fig. 6.24b) P||Pa||R||R ?R (Fig. 4.9a) P||Pa||R||R (Fig. 4.8g) R ?R ?R ?P ?||R ?R (Fig. 5.1b) Idem no. 13 17. 3PPaPaR-1RPPP (Fig. 6.25) P||Pa||Pa ?R (Fig. 4.10a) P||Pa||Pa||R (Fig. 4.10b) R ?P\\\\P\\\\P (Fig. 5.1a) Idem no. 9 18. 3PPaPaR-1RUPU (Fig. 6.26) P||Pa||Pa ?R (Fig. 4.10a) P||Pa||Pa||R (Fig. 4.10b) R ?R ?R ?P ?||R ?R (Fig. 5.1b) Idem no. 13 Table 6.7 Structural parametersa of parallel mechanisms in Figs. 6.7, 6.8, 6.9, 6.10, 6.11, 6.12, 6.13, 6.14 No. Structural parameter Solution Figures 6.7 and 6.11 Figures 6.8, 6.9, 6.10, 6.12, 6.13, 6.14 1. m 18 20 2. pi (i = 1, 3) 4 5 3. p2 4 4 4. p4 10 10 5. p 22 24 6. q 5 5 7. k1 3 3 8", + "4o and p to replace a group of revolute and prismatic pairs with coincident axis/direction. Fully-parallel topologies with at least two identical limbs are illustrated in Figs. 6.27, 6.28, 6.29, 6.30, 6.31, 6.32, 6.33, 6.34, 6.35, 6.36, 6.37, 6.38, 6.39, 6.40, 6.41, 6.42, 6.43, 6.44, 6.45, 6.46. The limb:topology and connecting conditions of these solutions are systematized in Tables 6.13 and 6.14, as are their structural parameters in Tables 6.15 and 6.16. Table 6.12 (continued) No. Structural parameter Solution Figure 6.25 Figure 6.26 Pp1 j\u00bc1 fj 10 10 30. Pp2 j\u00bc1 fj 10 10 31. Pp3 j\u00bc1 fj 10 10 32. Pp4 j\u00bc1 fj 4 6 33. Pp j\u00bc1 fj 34 36 a See footnote of Table 2.2 for the nomenclature of structural parameters 6.3 Fully-Parallel Topologies with Complex Limbs 621 622 6 Maximally Regular Topologies with Sch\u00f6nflies Motions 6.3 Fully-Parallel Topologies with Complex Limbs 623 624 6 Maximally Regular Topologies with Sch\u00f6nflies Motions 6.3 Fully-Parallel Topologies with Complex Limbs 625 626 6 Maximally Regular Topologies with Sch\u00f6nflies Motions 6" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_35_0000128_j.cryogenics.2021.103283-Figure2-1.png", + "original_path": "designv11-35/openalex_figure/designv11_35_0000128_j.cryogenics.2021.103283-Figure2-1.png", + "caption": "Fig. 2. (a) The general structure of the cryogenic induction motor. a: squirrel cage rotor, b: stator core, c: vacuum, d: double-layers Dewar, e: stator bearing columns; (b) The cross section of the motor.", + "texts": [ + " The heat insulation of the Dewar and running of the cryogenic motor were studied, which provides the basis for the development of HTS motors. The structure of the cryogenic induction motor is similar to that of the conventional induction motor with three exceptions. (1) The entire stator core of the cryogenic induction motor is surrounded by a double- layer Dewar, so that the copper coil and the stator core are entirely immersed in liquid nitrogen. (2) The air gap is increased from 0.5 to 5.1 mm due to the existence of Dewar. (3) The bearing structure of the stator core is changed to reduce forces on the Dewar. Fig. 2(a) and (b) shows the general structure and the cross section of the cryogenic induction motor with a squirrel cage rotor, a stator core, a double-layer Dewar and a bearing structure of the stator core. The bearing structure of the stator core consists of the support column welded on the stator core, the end cap, and the other axial force transmission structure. The stator core with copper winding is surrounded by the Dewar. The structure of the Dewar is shown in Fig. 3. The Dewar is a double-layer structure consisting of four cylinders and four covers" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_35_0001419_978-94-007-7401-8-Figure2.54-1.png", + "original_path": "designv11-35/openalex_figure/designv11_35_0001419_978-94-007-7401-8-Figure2.54-1.png", + "caption": "Fig. 2.54 4PaPaPR-type fully-parallel PM with coupled Sch\u00f6nflies motions defined by MF \u00bc SF \u00bc 4; RF\u00f0 \u00de \u00bc \u00f0v1; v2; v3;xd\u00de, TF = 0, NF = 30, limb topology Pa||Pa||P\\R", + "texts": [ + " 4PaPaPR (Fig. 2.48) Pa\\Pa\\\\P\\\\R (Fig. 2.22i) Idem No. 4 37. 4RPaPaR (Fig. 2.49) R\\Pa||Pa\\kR (Fig. 2.22j) The first and the last revolute joints of the four limbs are parallel 38. 4RPaPaR (Fig. 2.50) R\\Pa||Pa\\kR (Fig. 2.22k) Idem No. 37 39. 4RPaRPa (Fig. 2.51) R\\Pa\\kR\\Pa (Fig. 2.22l) Idem No. 15 40. 4PaRPPa (Fig. 2.52) Pa\\R\\P||Pa (Fig. 2.22m) Idem No. 34 41. 4PaPaRP (Fig. 2.53) Pa||Pa\\R\\P (Fig. 2.22n) The axes of the revolute joints connecting links 7 and 8 of the four limbs are parallel 42. 4PaPaPR (Fig. 2.54) Pa||Pa||P\\R (Fig. 2.22o) Idem No. 4 43. 4PRPaPa (Fig. 2.55) P\\R\\Pa||Pa (Fig. 2.22p) Idem No. 5 44. 4PPaRPa (Fig. 2.56) P||Pa\\R\\Pa (Fig. 2.22q) Idem No. 33 45. 4PPaPaR (Fig. 2.57) P||Pa||Pa\\R (Fig. 2.22r) Idem No. 4 46. 4PaPPaR (Fig. 2.58) Pa||P||Pa\\R (Fig. 2.22s) Idem No. 4 47. 4RPPaPa (Fig. 2.59) R\\P||Pa||Pa (Fig. 2.22t) Idem No. 15 48. 4PaPPaR (Fig. 2.60) Pa\\P\\\\Pa\\kR (Fig. 2.22u) Idem No. 4 49. 4PaRPaP (Fig. 2.61) Pa\\R\\Pa\\kR (Fig. 2.22v) Idem No. 34 50. 4PaRPPa (Fig. 2.62) Pa\\R||P\\Pa (Fig. 2" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_35_0001419_978-94-007-7401-8-Figure4.13-1.png", + "original_path": "designv11-35/openalex_figure/designv11_35_0001419_978-94-007-7401-8-Figure4.13-1.png", + "caption": "Fig. 4.13 Complex limbs for fully-parallel PMs with decoupled Sch\u00f6nflies motions, combining three closed loops, defined by MG = SG = 5, RG\u00f0 \u00de \u00bc v1; v2; v3;xa;xb", + "texts": [ + " 4.10n) Pa ?Pn3 ?? R (Fig. 4.11l) Idem No. 6 26. 1PaPn3\u20133PaPn3R (Fig. 4.37) Pa ? Pn3||R (Fig. 4.10o) Pa ? Pn3 ??R (Fig. 4.11m) Idem No. 6 27. 1CRbRbR\u20133CRbRbRR (Fig. 4.38) C||Rb||Rb||R (Fig. 4.10p) C||Rb||Rb||R ? R (Fig. 4.11n) Idem No. 6 28. 4PaPaPaR (Fig. 4.39) Pa ? P||Pa||R (Fig. 4.12a) Pa ? Pa||Pa ?? R (Fig. 4.12c) Idem No. 6 29. 4PaPaPaR (Fig. 4.40) Pa ? Pa||Pa||R (Fig. 4.12a) Pa ? Pa||Pa ??R (Fig. 4.12b) Idem No. 6 3PaRRbRbRR (Fig. 4.41) Pa ? R||Rb||Rb||R (Fig. 4.12d) Pa ? R||Rb||Rb||R ??R (Fig. 4.13a) Idem No. 6 3PaRRbRbRR (Fig. 4.42) Pa ? R||Rb||Rb||R (Fig. 4.12e) Pa ? R||Rb||Rb||R ?? R (Fig. 4.13b) Idem No. 6 4.2 Topologies with Complex Limbs 427 Table 4.5 Structural parametersa of parallel mechanisms in Figures 4.14, 4.15, 4.16, 4.17, 4.18, 4.19, 4.20, 4.21, 4.22, 4.23 No. Structural parameter Solution Figures 4.14 and 4.15 Figures 4.16, 4.17, 4.18, 4.19, 4.20, 4.21, 4.22 Figure 4.23 1. m 22 25 21 2. p1 7 7 6 3. pi (i = 2, 3, 4) 7 8 7 4. p 28 31 27 5. q 7 7 7 6. k1 0 0 0 7. k2 4 4 4 8. k 4 4 4 9. (RG1) (v1; v2; v3;xa) (v1; v2; v3;xa) (v1; v2; v3;xa) 10. (RG2) (v1; v2; v3;xa) (v1; v2; v3;xa;xb) (v1; v2; v3;xa;xb) 11" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_35_0001672_s0368393100116256-Figure3-1.png", + "original_path": "designv11-35/openalex_figure/designv11_35_0001672_s0368393100116256-Figure3-1.png", + "caption": "Fig. 3. Stages in the production of the stator with bridged slots for a hysteresis motor.", + "texts": [ + " Finally, to reduce the eddy current loss in the sleeve and also to provide an easier stator to load, we adopted the device of using externally slotted laminations allowing the coils to be loaded from the outside, the stator iron path being provided by a laminated annular ring enclosing the stator after load ing. The bridges across the slots were then reduced to the minimum thickness, giving satisfactory reduction of the rotor parasitic loss by boring the stator to its final dimension after loading. 413 F. W. MEREDITH The stages in the production of the finished stator are shown in Fig. 3. With this form of construction the disparity in the D.C. and A.C. torques was reduced, as shown in Fig. 4. The dotted curve on Fig. 4 shows the value of Q\u00a5 on the assumption that the parasitic loss is the difference between this curve and either of the others. It will be seen that the parasitic loss is only a few per cent, of the output power at the operating excitation, which is a very satisfactory result. Figure 5 shows output power against input power for rotors of several degrees of magnetic hardness in the same stator, from which it appears that there is an optimum induction for each material and an optimum material for each output torque per unit of rotor volume" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_35_0001419_978-94-007-7401-8-Figure2.46-1.png", + "original_path": "designv11-35/openalex_figure/designv11_35_0001419_978-94-007-7401-8-Figure2.46-1.png", + "caption": "Fig. 2.46 4PaRPPa-type fully-parallel PM with coupled Sch\u00f6nflies motions defined by MF \u00bc SF \u00bc 4; RF\u00f0 \u00de \u00bc \u00f0v1; v2; v3;xd\u00de, TF = 0, NF = 30, limb topology Pa||R\\P\\\\Pa", + "texts": [ + "22b) The axes of the revolute joints connecting the two parallelogram loops of the four limbs are parallel (continued) 2.2 Topologies with Complex Limbs 175 Table 2.3 (continued) No. PM type Limb topology Connecting conditions 30. 4PRPaPa (Fig. 2.42) P\\R||Pa\\Pa (Fig. 2.22c) Idem No. 5 31. 4RPPaPa (Fig. 2.43) R\\P\\kPa\\\\Pa (Fig. 2.22d) Idem No. 15 32. 4PaPPaR (Fig. 2.44) Pa\\P\\\\Pa\\\\R (Fig. 2.22e) Idem No. 4 33. 4PaPRPa (Fig. 2.45) Pa\\P\\kR\\Pa (Fig. 2.22f) The axes of the revolute joints connecting links 5 and 6 of the four limbs are parallel 34. 4PaRPPa (Fig. 2.46) Pa||R\\P\\\\Pa (Fig. 2.22g) The axes of the revolute joints connecting links 4 and 5 of the four limbs are parallel 35. 4RPaPPa (Fig. 2.47) R||Pa\\P\\\\Pa (Fig. 2.22h) Idem No. 15 36. 4PaPaPR (Fig. 2.48) Pa\\Pa\\\\P\\\\R (Fig. 2.22i) Idem No. 4 37. 4RPaPaR (Fig. 2.49) R\\Pa||Pa\\kR (Fig. 2.22j) The first and the last revolute joints of the four limbs are parallel 38. 4RPaPaR (Fig. 2.50) R\\Pa||Pa\\kR (Fig. 2.22k) Idem No. 37 39. 4RPaRPa (Fig. 2.51) R\\Pa\\kR\\Pa (Fig. 2.22l) Idem No. 15 40. 4PaRPPa (Fig. 2.52) Pa\\R\\P||Pa (Fig" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_35_0001780_joaiee.1923.6592178-Figure8-1.png", + "original_path": "designv11-35/openalex_figure/designv11_35_0001780_joaiee.1923.6592178-Figure8-1.png", + "caption": "FIG. 8 FIG. 1 0", + "texts": [ + " Con- and the alternating component is sider these consequences first in the light of the equations, and then from the physical point of view. By equation (25), for \u00f9 ti = \u00f0 / 2 *i = t o sin \u00f9 t - K2 1 - K2 (29) 1 - K2 In (30), the direct component is \u0399 i2' = i0 M/L2 \u2022 for values of t > t h since the equations obviously hold only for conditions after the short circuit. And by (26) 1 - K2 and the alternating component is sin \u00f9 t i/ = - io M/L2 1 - K2 (32) (33) (34) Equations (29) and (30) are plotted in Fig. 8, showing graphically the relation between the alternating and direct components. Now examine the physics of the problem. Consider the case in which the switch sh Fig. 6, is closed when the secondary is linked with maximum flux, that is, when co t = 7T/2. This flux, or linkages rather, must remain constant. But obviously one-half cycle later, the flux in the primary is full value in the opposite direction, since the impressed voltage is maintained. That is, the primary is trying to force flux through the closed secondary, which must not only prevent this opposite flux from entering, but also must maintain constant the flux with which it is linked. This instant corresponds to the maximum current in Fig. 8. Since the reversed primary flux cannot enter the secondary, and the secondary flux cannot enter the primary, it follows that both fluxes must pass through the leakage paths between the windings. I t thus requires twice 1 1 as much maximum current as the case in which the switch Si was closed when the secondary enclosed zero flux. Thus the actual current in both the primary and the secondary comprises a direct and an alternating component of about equal maximum values. In the secondary the direct, or constant flux, with which it is linked, requires a direct current to maintain it, and an alternating current is required to prevent the alternating flux of the primary from linking the secondary. And similarly the primary requires a direct current to prevent the direct flux of the secondary from linking it (since the primary flux must at all times correspond to the impressed voltage), and also requires an alternating component to force the alternating flux through the leakage paths. Thus the total current in either of these windings comprises the two components as shown in Fig. 8. The transformer problem thus illustrates in fundamental respects the facility of calculation and of visualization made possible by the \"theorem of constant magnetic linkages/' Other illustrations will now be given in a much briefer way. Single-Phase Alternator. In Fig. 9, a a represents the armature winding, b b the field winding. Assume that the field winding is excited by the current i0; that the field winding, having zero resistance, is shortcircuited at the collector rings.12 Thus as long as the permeance of the magnetic circuit is not changed, i0 would continue to flow, by the condition that the field linkages must remain constant" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_35_0001419_978-94-007-7401-8-Figure5.21-1.png", + "original_path": "designv11-35/openalex_figure/designv11_35_0001419_978-94-007-7401-8-Figure5.21-1.png", + "caption": "Fig. 5.21 2PaPRRR-1PaPRR-1RPaPatP-type fully-parallel PMs with uncoupled Sch\u00f6nflies motions defined by MF = SF = 4, RF\u00f0 \u00de \u00bc \u00f0v1; v2; v3;xb\u00de), TF = 0, NF = 19, limb topology R||Pa||Pat||P and Pa\\P\\\\R||R\\\\R, Pa\\\\P\\\\R||R (a), Pa||P\\R||R\\\\R, Pa||P\\R||R (b)", + "texts": [ + "19b) Pa||P\\R||R\\\\R (Fig. 5.3b) Idem No. 9 Pa||P\\R||R (Fig. 5.2f) R||Pa||Pa||P (Fig. 5.4k) (continued) 528 5 Topologies with Uncoupled Sch\u00f6nflies Motions Table 5.2 (continued) No. PM type Limb topology Connecting conditions 11. 2PaPRRR-1PaPRR1RPPaPa (Fig. 5.20a) Pa\\P\\\\R||R\\\\R (Fig. 5.3a) Idem No. 9 Pa\\P\\\\R||R (Fig. 5.2e) R||P||Pa||Pa (Fig. 5.4m) 12. 2PaPRRR-1PaPRR1RPPaPa (Fig. 5.20b) Pa||P\\R||R\\\\R (Fig. 5.3b) Idem No. 9 Pa||P\\R||R (Fig. 5.2f) R||P||Pa||Pa (Fig. 5.4m) 13. 2PaPRRR-1PaPRR1RPaPatP (Fig. 5.21a) Pa\\P\\\\R||R\\\\R (Fig. 5.3a) Idem No. 9 Pa\\P\\\\R||R (Fig. 5.2e) R||Pa||Pat||P (Fig. 5.4l) 14. 2PaPRRR-1PaPRR1RPaPatP (Fig. 5.21b) Pa||P\\R||R\\\\R (Fig. 5.3b) Idem No. 9 Pa||P\\R||R (Fig. 5.2f) R||Pa||Pat||P (Fig. 5.4l) 15. 2PaPRRR-1PaPRR1RPPaPat (Fig. 5.22a) Pa\\P\\\\R||R\\\\R (Fig. 5.3a) Idem No. 9 Pa\\P\\\\R||R (Fig. 5.2e) R||P||Pa||Pat (Fig. 5.4n) 16. 2PaPRRR-1PaPRR1RPPaPat (Fig. 5.22b) Pa||P\\R||R\\\\R (Fig. 5.3b) Idem No. 9 Pa||P\\R||R (Fig. 5.2f) R||P||Pa||Pat (Fig. 5.4n) 17. 2PaRPRR-1PaRPR1RPaPaP (Fig. 5.23) Pa\\R\\P\\kR\\R (Fig. 5.3c) The last joints of the four limbs have superposed axes/directions. The revolute joints between links 4 and 5 of limbs G1, G2 and G3 have orthogonal axes Pa\\R\\P\\kR (Fig" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_35_0001419_978-94-007-7401-8-Figure1.1-1.png", + "original_path": "designv11-35/openalex_figure/designv11_35_0001419_978-94-007-7401-8-Figure1.1-1.png", + "caption": "Fig. 1.1 Symbols used to represent the lower kinematic pairs and the kinematic joints: a revolute pair, b prismatic pair, c helical pair, d cylindrical pair, e spherical pair, f planar contact pair, g universal joint, h homokinetic joint, i two superposed revolute joints (1, 2) and (2, 3) with the same axis, j superposed cylindrical (1, 2) and revolute (2, 3) joints with the same axis, k superposed revolute (1, 2) and cylindrical (2, 3) joints with the same axis, and l two superposed cylindrical joints (1, 2) and (2, 3) with the same axis", + "texts": [ + " The word joint is used as a synonym for the kinematic pair and also to define the physical realization of a kinematic pair, including connection via intermediate mechanism elements. Both synonymous terms are used in this text. Usually, in parallel robots, lower pairs are used: revolute R, prismatic P, helical H, cylindrical C, spherical S and planar pair E. The definitions of these kinematic pairs are presented in Table 1.1\u2014Part 1. The graphical representations used in this book for the lower pairs are presented in Fig. 1.1a\u2013f. Universal joints and homokinetic joints are also currently used in the mechanical structure of the parallel 4 1 Introduction robots to transmit the rotational motion between two shafts with intersecting axes. If the instantaneous velocities of the two shafts are always the same, the kinematic joint is homokinetic (from the Greek \u2018\u2018homos\u2019\u2019 and \u2018\u2018kinesis\u2019\u2019 meaning \u2018\u2018same\u2019\u2019 and \u2018\u2018movement\u2019\u2019). We know that the universal joint (Cardan joint or Hooke\u2019s joint) are heterokinetic joints. Various types of homokinetic joints (HJ) are known today: Tracta, Weiss, Bendix, Dunlop, Rzeppa, Birfield, Glaenzer, Thompson, Triplan, Tripode, UF (undercut-free) ball joint, AC (angular contact) ball joint, VL plunge ball joint, DO (double offset) plunge ball joint, AAR (angular adjusted roller), helical flexure U-joints, etc. [6\u20139]. The graphical representations used in this book for the universal homokinetic joints are presented in Fig. 1.1g\u2013h. Joints with idle mobilities are commonly used to reduce the number of overconstraints in a mechanism. The idle mobility is a potential mobility of a joint that is not used by the mechanism and does not influence mechanism\u2019s mobility in the hypothesis of perfect manufacturing and assembling precision. In theoretical conditions, when no errors exist with respect to parallel, perpendicular or intersecting positions of joint axes, motion amplitude associated with an idle mobility is zero. Real life manufacturing and assembling processes introduce errors in the relative positions of the joint axes and, in this case, the idle mobilities become effective mobilities usually with small amplitudes, depending on the precision of the mechanism", + "3 gives an example of the various representations of a Gough-Stewart type parallel robot largely used today in industrial applications. The physical implementation in Fig. 1.3a is a photograph of the parallel robot built by Deltalab (http://www.deltalab.fr/). In a CAD model (Fig. 1.3b) the links and the joints are represented as being as close as possible to the physical implementation (Fig. 1.3a). In a structural diagram (Fig. 1.3c) they are represented by simplified symbols, such as those introduced in Fig. 1.1, respecting the geometric relations defined by the relative positions of joint axes. A structural graph (Fig. 1.3d) is a network of vertices or nodes connected by edges or arcs with no geometric relations. The links are noted in the nodes and the joints on the edges. We can see that the Gough-Stewart type parallel robot has six identical limbs denoted in Fig. 1.3c by A, B, C, D, E and F. The final link is the mobile platform 4 : 4A : 4B: 4C : 4D : 4E: 4F and the reference member is the fixed platform 1A : 1B: 1C : 1D: 1E : 1F : 0" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_35_0001419_978-94-007-7401-8-Figure4.33-1.png", + "original_path": "designv11-35/openalex_figure/designv11_35_0001419_978-94-007-7401-8-Figure4.33-1.png", + "caption": "Fig. 4.33 1PaPn2R-3PaPn2RR-type fully-parallel PM with decoupled Sch\u00f6nflies motions defined by MF = SF = 4, (RF) = (v1; v2; v3;xa), TF = 0, NF = 27, limb topology Pa ? Pn2||R and Pa ? Pn2||R ?? R", + "texts": [ + " 4.28) Pa ? R||Rb||R (Fig. 4.10f) Pa ? R||Rb||R ?? R (Fig. 4.11d) Idem No. 6 18. 1PaRRbR-3PaRRbRR (Fig. 4.29) Pa ? R||Rb||R (Fig. 4.10g) Pa ? R||Rb||R ??R (Fig. 4.11e) Idem No. 1 19. 1PaPn2R\u20133PaPn2RR (Fig. 4.30) Pa ? Pn2||R (Fig. 4.10h) Pa ? Pn2||R ?? R (Fig. 4.11f) Idem No. 14 20. 1PaPn2R\u20133PaPn2RR (Fig. 4.31) Pa ? Pn2||R (Fig. 4.10i) Pa ? Pn2||R ??R (Fig. 4.11g) Idem No. 14 21. 1PaPn2R\u20133PaPn2RR (Fig. 4.32) Pa ? Pn2||R (Fig. 4.10j) Pa ? Pn2||R ??R (Fig. 4.11h) Idem No. 14 22. 1PaPn2R\u20133PaPn2RR (Fig. 4.33) Pa ? Pn2||R (Fig. 4.10k) Pa ? Pn2||R ?? R (Fig. 4.11i) Idem No. 14 (continued) 426 4 Fully-Parallel Topologies with Decoupled Sch\u00f6nflies Motions Table 4.4 (continued) No. PM type Limb topology Connecting conditions 23. Pa ?Pn2||R ??R (Fig. 4.34) Pa ? Pn3||R (Fig. 4.10l) Pa ? Pn3 ?? R (Fig. 4.11j) Idem No. 6 24. 1PaPn3\u20133PaPn3R (Fig. 4.35) Pa ? Pn3||R (Fig. 4.10m) Pa ? Pn3 ??|R (Fig. 4.11k) Idem No. 6 25. 1PaPn3\u20133PaPn3R (Fig. 4.36) Pa ? Pn3||R (Fig. 4.10n) Pa ?Pn3 ?? R (Fig. 4.11l) Idem No. 6 26" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_35_0001419_978-94-007-7401-8-Figure2.40-1.png", + "original_path": "designv11-35/openalex_figure/designv11_35_0001419_978-94-007-7401-8-Figure2.40-1.png", + "caption": "Fig. 2.40 4PPaPaR-type fully-parallel PM with coupled Sch\u00f6nflies motions defined by MF \u00bc SF \u00bc 4; RF\u00f0 \u00de \u00bc \u00f0v1; v2; v3;xd\u00de, TF = 0, NF = 30, limb topology P\\Pa\\\\Pa\\||R", + "texts": [ + "21t) Idem No. 5 21. 4PaPRP (Fig. 2.36a) Pa\\P||R\\P (Fig. 2.21u) Idem No. 11 22. 4PaRPP (Fig. 2.36b) Pa\\R\\P\\kP (Fig. 2.21v) Idem No. 5 23. 4PCPa (Fig. 2.37a) P\\C\\Pa (Fig. 2.21w) Idem No. 5 24. 4PPaC (Fig. 2.37b) P||Pa\\C (Fig. 2.21x) The last joints of the four limbs have parallel axes 25. 4CPPa (Fig. 2.38a) C\\P||Pa (Fig. 2.21y) The first joints of the four limbs have parallel axes 26. 4PaPC (Fig. 2.38b) Pa||P\\C (Fig. 2.21z) Idem No. 24 27. 4PaCP (Fig. 2.39) Pa\\C\\P (Fig. 2.21z1) Idem No. 11 28. 4PPaPaR (Fig. 2.40) P\\Pa\\\\Pa\\kR (Fig. 2.22a) Idem No. 4 29. 4PPaRPa (Fig. 2.41) P\\Pa||R\\Pa (Fig. 2.22b) The axes of the revolute joints connecting the two parallelogram loops of the four limbs are parallel (continued) 2.2 Topologies with Complex Limbs 175 Table 2.3 (continued) No. PM type Limb topology Connecting conditions 30. 4PRPaPa (Fig. 2.42) P\\R||Pa\\Pa (Fig. 2.22c) Idem No. 5 31. 4RPPaPa (Fig. 2.43) R\\P\\kPa\\\\Pa (Fig. 2.22d) Idem No. 15 32. 4PaPPaR (Fig. 2.44) Pa\\P\\\\Pa\\\\R (Fig. 2.22e) Idem No. 4 33. 4PaPRPa (Fig" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_35_0000008_j.autcon.2021.103573-Figure28-1.png", + "original_path": "designv11-35/openalex_figure/designv11_35_0000008_j.autcon.2021.103573-Figure28-1.png", + "caption": "Fig. 28. Gantry crane system with external disturbances.", + "texts": [ + " As the actual block was heavier, the winches hoisted up the wire ropes to maintain the position of the block in the z-direction. Subsequently, the control moments at the terminal time was increased compared to Case 1. The position of the trolleys and the length of the wire ropes are presented in Fig. 27. The length of the wire ropes was increased at the beginning and adjusted in 10 s by the winches. In Case 3, the environmental load was exerted on the block as an external force. An irregular wind was applied to the block with a mean speed of 10 m/s and a direction of 45\u25e6 (see Fig. 28). The wind fluctuation was calculated from the NPD (Norwegian Petroleum Directorate) wind spectrum given by ISO 19901-1, where U(H) is the mean wind speed at height H above the water plane. U(H) = U(10)\u22c5 ( H 10 )\u03b1 (39) In Eq.(39), U(10) is the mean wind speed at 10 m above the water plane. \u03b1 is the height coefficient and is usually set to be 0.11 according to ISO 19901-1. The NPD wind spectrum S(f) with frequency f is then given by the following equations, where n is 0.468 [25]. H.-W. Lee et al" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_35_0000807_s12206-021-0724-8-Figure4-1.png", + "original_path": "designv11-35/openalex_figure/designv11_35_0000807_s12206-021-0724-8-Figure4-1.png", + "caption": "Fig. 4. FE model of the components: (a) crankcase housing; (b) crankshaft.", + "texts": [ + " The FE model scale influences the calculation accuracy, and three different scale FE models can be selected for the calculation of the main bearing lubrication characteristics of a V-type six-cylinder engine, namely, the entire engine block, crankcase housing, and single bearing housings, as shown in Fig. 2. The crankcase housing model has a smaller scale and higher stiffness accuracy than the engine block and single bearing housing models. Simulations are conducted using different FE models, and the results show that the PACP of the main bearing in the first two cases is consistent, as shown in Fig. 3. The maximum relative difference of the average PACP is less than 10 %. Therefore, the crankcase housing model is selected in the simulation. Fig. 4 shows the FE model of the crankcase housing and crankshaft. The main bearing region is the area of interest for the calculation, where the FE mesh is refined. The main bearing shell grid is divided into eight layers, five layers, and 60 parts in the axial, radial, and circumferential directions, respectively. The grid of the mean bearing corresponds to the grid of the bearing shell in the axial and circumferential directions, and the grid size in the radial direction increases with the radius, as shown in Fig. 4(a). The grid of the crankshaft journal also corresponds to the grid of the bearing shell, and the grid at the rounded corners of the crankshaft is axially divided into four layers, as shown in Fig. 4(b). Structural dynamic calculations have lower requirements for mesh refinement than static calculations. Therefore, the grid away from the regions of interest has not been refined specifically. However, a fine oil film grid is required for an accurate calculation of the pressure distribution in fluid calculations. Therefore, a high density oil film grid is developed; it is divided into 24 parts and 120 parts in the axial and circumferential directions, respectively, as shown in Fig. 4(a). An area mapping approach [33, 34] is used for transferring information between the structural and oil film grid; it ensures the accuracy and the efficiency of the calculation of the main bearing lubrication characteristics. The stiffness and mass information of the crankcase housing and crankshaft are condensed to the nodes of the bearing surface and nine nodes on the centerline of each journal. A flexible dynamic model of 30 and 60 normal modes of the crankshaft and crankcase housing is used to reduce the internal DOF" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_35_0001419_978-94-007-7401-8-Figure5.42-1.png", + "original_path": "designv11-35/openalex_figure/designv11_35_0001419_978-94-007-7401-8-Figure5.42-1.png", + "caption": "Fig. 5.42 3PaPPaR-1RPPaPat-type fully-parallel PMs with uncoupled Sch\u00f6nflies motions defined by MF = SF = 4, RF\u00f0 \u00de \u00bc \u00f0v1; v2; v3;xb\u00de, TF = 0, NF = 30 limb topology R||P||Pa||Pat and Pa||P\\Pa\\kR, Pa||P\\Pa||R (a), Pa\\P\\\\Pa\\kR, Pa\\P\\\\Pa||R (b)", + "texts": [ + "4d) R||P||Pa||Pa (Fig. 5.4m) (continued) 5.1 Fully-Parallel Topologies 531 Table 5.4 (continued) No. PM type Limb topology Connecting conditions 4. 3PaPPaR-1RPPaPa (Fig. 5.40b) Pa\\P\\\\Pa\\kR (Fig. 5.4b) Idem No. 3 Pa\\P\\\\Pa||R (Fig. 5.4c) R||P||Pa||Pa (Fig. 5.4m) 5. 3PaPPaR1RPaPatP (Fig. 5.41a) Pa||P\\Pa\\kR (Fig. 5.4a) Idem No. 1 Pa||P\\Pa||R (Fig. 5.4d) R||Pa||Pat||P (Fig. 5.4l) 6. 3PaPPaR1RPaPatP (Fig. 5.41b) Pa\\P\\\\Pa\\kR (Fig. 5.4b) Idem No. 1 Pa\\P\\\\Pa||R (Fig. 5.4c) R||Pa||Pat||P (Fig. 5.4l) 7. 3PaPPaR-1RPPaPat (Fig. 5.42a) Pa||P\\Pa\\kR (Fig. 5.4a) Idem No. 3 Pa||P\\Pa||R (Fig. 5.4d) R||P||Pa||Pat (Fig. 5.4n) 8. 3PaPPaR-1RPPaPat (Fig. 5.42b) Pa\\P\\\\Pa\\kR (Fig. 5.4b) Idem No. 3 Pa\\P\\\\Pa||R (Fig. 5.4c) R||P||Pa||Pat (Fig. 5.4n) 9. 3PaPaPR-1RPaPaP (Fig. 5.43) Pa\\Pa\\kP\\\\R (Fig. 5.4f) Idem No. 1 Pa\\Pa\\kP\\kR (Fig. 5.4h) R||Pa||Pa||P (Fig. 5.4k) 10. 3PaPaPR-1RPaPaP (Fig. 5.44) Pa\\Pa\\\\P\\\\R (Fig. 5.4g) Idem No. 1 Pa\\Pa\\\\P\\kR (Fig. 5.4e) R||Pa||Pa||P (Fig. 5.4k) 11. 3PaPaPR1RPaPatP (Fig. 5.45) Pa\\Pa\\kP\\\\R (Fig. 5.4f) Idem No. 1 Pa\\Pa\\kP\\kR (Fig. 5.4h) R||Pa||Pat||P (Fig. 5.4l) 12. 3PaPaPR1RPaPatP (Fig. 5.46) Pa\\Pa\\\\P\\\\R (Fig. 5.4g) Idem No. 1 Pa\\Pa\\\\P\\kR (Fig" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_35_0001419_978-94-007-7401-8-Figure5.32-1.png", + "original_path": "designv11-35/openalex_figure/designv11_35_0001419_978-94-007-7401-8-Figure5.32-1.png", + "caption": "Fig. 5.32 3PaPaPR-1RPPP-type fully-parallel PMs with uncoupled Sch\u00f6nflies motions defined by MF = SF = 4, RF\u00f0 \u00de \u00bc \u00f0v1; v2; v3;xb\u00de, TF = 0, NF = 24, limb topology R\\P\\\\P\\\\P and Pa\\Pa\\kP\\\\R, Pa\\Pa\\kP\\kR (a), Pa\\Pa\\\\P\\\\R, Pa\\Pa\\\\P\\kR (b)", + "texts": [ + "3 Limb topology and connecting conditions of the fully-parallel solutions with no idle mobilities presented in Figs. 5.31, 5.32, 5.33, 5.34, 5.35, 5.36, 5.37, 5.38 No. PM type Limb topology Connecting conditions 1. 3PaPPaR1RPPP (Fig. 5.31a) Pa||P\\Pa\\kR (Fig. 5.4a) The last joints of the four limbs have superposed axes/directionsPa||P\\Pa||R (Fig. 5.4d) R\\P\\\\P\\\\P (Fig. 5.1a) 2. 3PaPPaR- 1RPPP (Fig. 5.31b) Pa\\P\\\\Pa\\kR (Fig. 5.4b) Idem No. 1 Pa\\P\\\\Pa||R (Fig. 5.4c) R\\P\\\\P\\\\P (Fig. 5.1a) 3. 3PaPaPR1RPPP (Fig. 5.32a) Pa\\Pa\\kP\\\\R (Fig. 5.4f) Idem No. 1 Pa\\Pa\\kP\\kR (Fig. 5.4h) R\\P\\\\P\\\\P (Fig. 5.1a) 4. 3PaPaPR1RPPP (Fig. 5.32b) Pa\\Pa\\\\P\\\\R (Fig. 5.4g) Idem No. 1 Pa\\Pa\\\\P\\kR (Fig. 5.4e) R\\P\\\\P\\\\P (Fig. 5.1a) 5. 2PaPaRRR1PaPaRR1RPPP (Fig. 5.33) Pa\\Pa||R||R\\\\R (Fig. 5.5a) The last joints of the four limbs have superposed axes/directions. The revolute joints between links 7 and 8 of limbs G1, G2 and G3 have orthogonal axes Pa\\Pa||R||R (Fig. 5.4i) R\\P\\\\P\\\\P (Fig. 5.1a) 6. 2PaPaRRR1PaPaRR1RPPP (Fig. 5.34) Pa\\Pa||R||R\\kR (Fig. 5.5b) Idem No. 5 Pa\\Pa||R||R (Fig. 5.4j) R\\P\\\\P\\\\P (Fig. 5.1a) 7. 3PaPPaR1RUPU (Fig. 5.35a) Pa||P\\Pa\\kR (Fig" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_35_0000246_tte.2021.3068819-Figure4-1.png", + "original_path": "designv11-35/openalex_figure/designv11_35_0000246_tte.2021.3068819-Figure4-1.png", + "caption": "Fig. 4. n=3 topology of the equivalent permeability modeling method.", + "texts": [ + " The position error can be expressed as: ( ) 11 0 1 , in n i i n n d n n +\u2212 = = \u2212 (9) So, the maximum error is: ( ) 2 0 1 1, = 2 d = \u2212 (10) We can get other errors, just like: ( ) 11 22 0 2 1 1 2, = 2 2 8 i i i d + = = \u2212 (11) ( ) 12 23 0 3 1 1 3, = 3 3 18 i i i d + = = \u2212 (12) From (10)-(12), it is clear that the position accumulative error is small enough and will not affect the calculation accuracy when number of separated models is 3 (n=3). A. Verification and analysis of 5-parallel-strand model In this section, the 5-parallel-strands model, shown in Fig. 2 and Fig. 3, is taken as an example to verify the correctness of the equivalent permeability modeling method and to analyze the winding AC loss. For analysis and verification, this section takes n=3 as an example for analysis. Fig. 4 is a schematic diagram of an equivalent permeability modeling method. Taking \u03b8=0\u00b0 as an example, Fig. 5(a) shows the physical circuit of 2D-EMM. The axial lengths of Model 1, Model 2, and Model 3 are l1, l2, and l3, respectively. When \u03be1=\u03be2=\u03be3, the strands are evenly displaced in the axial direction; When \u03be1\u2260\u03be2\u2260\u03be3, strands are unevenly displaced in the axial direction. As a result, 2D-EMM can be used to optimize the conductor transposition angle and transposition length by changing the equivalent coefficient" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_35_0000437_acs.langmuir.1c00520-Figure8-1.png", + "original_path": "designv11-35/openalex_figure/designv11_35_0000437_acs.langmuir.1c00520-Figure8-1.png", + "caption": "Figure 8. Schematics of a buckled film clamped around a circular perimeter. The arrows indicate the direction of film displacements needed to develop a Y-shaped fold (left) or a corral-shaped fold (right).", + "texts": [ + " Consider drop sizes that are sufficiently small that only one fold can nucleate within the drop. Near the center of the drop, both Y-shaped folds or corral-shaped folds can release in-plane biaxial compression. However, near the edge, tangential compression cannot be released because the film remains unswollen and fully bonded beyond the drop edge. Thus, drop edge acts as a clamped boundary condition for the film. This clamped boundary affects Y vs perimeter folds very differently, as illustrated in Figure 8. Developing a Y-fold requires radially inwards displacement of the film. Since the near-edge region has no out-of-plane deflection, radially inwards displacement would increase tangential compression. In contrast, for a corral fold, material points in the film move outwards and hence biaxial compression within the corral can be relaxed at least https://doi.org/10.1021/acs.langmuir.1c00520 Langmuir 2021, 37, 6985\u22126994 6992 partially. We speculate therefore that for sufficiently small drops, the corral fold offers a lower energy than a single Y-fold" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_35_0000775_j.engfailanal.2021.105672-Figure15-1.png", + "original_path": "designv11-35/openalex_figure/designv11_35_0000775_j.engfailanal.2021.105672-Figure15-1.png", + "caption": "Fig. 15. Detail \u201cA\u201d: Stress distribution in the fracture zone and at the corresponding cross-section.", + "texts": [ + " Engineering Failure Analysis 129 (2021) 105672 The stress field of the helicopter skid landing gear and its aft cross tube was obtained and the values of the von Mises stress are presented. Visualization of stress field patterns is shown in Figs. 13-14. The static deflections of the helicopter skid landing gear, when a helicopter is parked on the ground, is such that the aft cross tube is exposed to bending load, inducing tension on the underside and compression on the upper side. Resulting stress field of the critical cross-section is shown at Fig. 15. The fractographic analysis revealed features associated with progressive propagation under initial fatigue on approximately 50% cross-section of the wall tube. The fatigue was initiated at two pre-existing cracks on the inner surface of the aft cross tube (Fig. 5). Identified beach, ratchet and chevron marks indicate initiation sites and cracks growth directions. The presence of two initial sites (multiple origins) indicates that cross tube operated at high nominal stress. When the fatigue crack reached the critical size, fast fracture occurred as the last phase", + " Engineering Failure Analysis 129 (2021) 105672 imoprtant role in the crack propagation and failure proccess of the helicopter skid landing gear cross-tube. The technial manual prescribes parking and mooring procedures that were not followed in presented case here, namely aft tube strap was not installed. This induce loads higher than the nominal static loads expected while the helicopter is in the parking position. Consequently this leads to the even higher stress state on the critial cross-section than the calculated one presented in Fig. 15. Frequent omission of installing tube strap during the period while the helicopter is in the parking position has significanly negative contribution on fatigue process of the structure. Based on the presented results it can be concluded that the failure of the helicopter skid landing gear has occurred as the cumulative influence of following factors: \u2022 Negative influence of the extrusion production process. Applied extrusion manufacturing method resulted with numerous surface imperfections at inner surface of aft cross tube" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_35_0001419_978-94-007-7401-8-Figure2.41-1.png", + "original_path": "designv11-35/openalex_figure/designv11_35_0001419_978-94-007-7401-8-Figure2.41-1.png", + "caption": "Fig. 2.41 4PPaRPa-type fully-parallel PM with coupled Sch\u00f6nflies motions defined by MF \u00bc SF \u00bc 4; RF\u00f0 \u00de \u00bc \u00f0v1; v2; v3;xd\u00de, TF = 0, NF = 30, limb topology P\\Pa||R\\Pa", + "texts": [ + "21u) Idem No. 11 22. 4PaRPP (Fig. 2.36b) Pa\\R\\P\\kP (Fig. 2.21v) Idem No. 5 23. 4PCPa (Fig. 2.37a) P\\C\\Pa (Fig. 2.21w) Idem No. 5 24. 4PPaC (Fig. 2.37b) P||Pa\\C (Fig. 2.21x) The last joints of the four limbs have parallel axes 25. 4CPPa (Fig. 2.38a) C\\P||Pa (Fig. 2.21y) The first joints of the four limbs have parallel axes 26. 4PaPC (Fig. 2.38b) Pa||P\\C (Fig. 2.21z) Idem No. 24 27. 4PaCP (Fig. 2.39) Pa\\C\\P (Fig. 2.21z1) Idem No. 11 28. 4PPaPaR (Fig. 2.40) P\\Pa\\\\Pa\\kR (Fig. 2.22a) Idem No. 4 29. 4PPaRPa (Fig. 2.41) P\\Pa||R\\Pa (Fig. 2.22b) The axes of the revolute joints connecting the two parallelogram loops of the four limbs are parallel (continued) 2.2 Topologies with Complex Limbs 175 Table 2.3 (continued) No. PM type Limb topology Connecting conditions 30. 4PRPaPa (Fig. 2.42) P\\R||Pa\\Pa (Fig. 2.22c) Idem No. 5 31. 4RPPaPa (Fig. 2.43) R\\P\\kPa\\\\Pa (Fig. 2.22d) Idem No. 15 32. 4PaPPaR (Fig. 2.44) Pa\\P\\\\Pa\\\\R (Fig. 2.22e) Idem No. 4 33. 4PaPRPa (Fig. 2.45) Pa\\P\\kR\\Pa (Fig. 2.22f) The axes of the revolute joints connecting links 5 and 6 of the four limbs are parallel 34" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_35_0000027_j.procir.2020.05.200-Figure4-1.png", + "original_path": "designv11-35/openalex_figure/designv11_35_0000027_j.procir.2020.05.200-Figure4-1.png", + "caption": "Fig. 4. Arrangement of nozzles and resulting cone projections (a) four nozzles with spray axis perpendicular to workpiece axis, (b) eight nozzles with spray axis inclined to workpiece axis (c) eight nozzles with spray axis perpendicular to workpiece axis.", + "texts": [ + " Investigation of nozzle arrangement To determine the arrangement of the nozzles, an analysis is caried out. This will show whether the setting parameters are suffiient to adjust the nozzles individually. The projection of the spray ones onto the surface of the workpieces before and after formng is simulated using the software Blender. The objective of inestigation is to identify and maximize the area reached with the ooling fluid while simultaneously achieving a maximum homoge- eous cooling surface. Fig. 4 shows an example of possible nozle configurations for the bearing bushing. Surfaces covered by the pray cone are shown in light grey, areas with higher spray denities in lighter grey scales. The heat transfer also decreases raially from the spray center ( Puschmann, 0 0 0 0 ). For a more ho- ogeneous heat transfer, an overlapping of the spray cones in the dge area is advantageous. Several configurations are considered, n which the number, position and orientation of the nozzles were aried. The position and rotation of the nozzles can be adjusted to dapt to different workpiece sizes and shapes.. In addition, differnt arrangement configurations can be compared and the cooling an be optimized. The translatory adjustment l of the nozzles is one manually via screw drives or slotted holes. The angles \u03c6can e adjusted via swivel joints. Both parameters are marked in Fig. 4 . In addition to the already mentioned parameters, the number f nozzles is also varied. Arrangements with four nozzles and eight ozzles are considered. By using four nozzles ( Fig. 4 a), the disance to the workpiece is greatest compared to the other arrange- ents, which results in a lower spray density. In addition, there s a greater overlap of the spray cones.As a result, this arrange- ent is not useful. In Fig. 4 b eight nozzles are rotated around the -axis to reach areas behind the pins that are left untouched in ther arrangements. The disadvantage of this solution is that areas nside the semi-finished product are sprayed directly or there are nsprayed areas in the edge area. On the semi-finished product in ig. 4 c the spray cones overlap in the edge area and on the formed omponent there are unsprayed areas between the spray cones. he variant in Fig. 4 c can be modified in such a way that there re no unsprayed surfaces between the spray cones in the formed omponent, which leads to a larger overlap on the semi-finished roduct. This would resemble Fig. 4 b, but without rotation around he z-direction. The objective is therefore to find a good compro- ise between homogeneous spraying of the semi-finished product nd the formed workpiece. It should be noted that the spray cone is approximated by a ight cone. The interaction with the environment cannot be anticpated in this simplified simulation. Therefore, experimental deter- ination of the ideal nozzle arrangement is needed. Due to turbuences, surfaces not located in the ideal spray cone can be affected y the cooling medium" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_35_0000253_s11661-021-06229-1-Figure15-1.png", + "original_path": "designv11-35/openalex_figure/designv11_35_0000253_s11661-021-06229-1-Figure15-1.png", + "caption": "Fig. 15\u2014Inverse pole figures in the vicinities of the fracture regions of as-built 17-4 PH stainless steel under different strain rate conditions: (a) 0.01 s 1; (b) 10 s 1; (c) 600 s 1.", + "texts": [ + " Since the hard martensite phase and soft austenite phase may have different mechanical behaviors, the phase transformation during the tensile deformation causes the phase composition to always change, thus significantly affecting the stress\u2013strain curve. It is more reasonable to describe the plastic deformation behavior by introducing the additional phase transformation kinetics based on a 2376\u2014VOLUME 52A, JUNE 2021 METALLURGICAL AND MATERIALS TRANSACTIONS A phenomenological-based constitutive model such as the Johnson\u2013Cook model. Much work will be done in the future. Figure 15 depicts the IPFs in the vicinities of the fracture region at the strain rates of 0.01, 10 and 600 s 1. From the figures, when the strain rate equals 0.01 and 10 s 1, the preferred orientations in both martensite and austenite are not along the building direction, but along the tensile direction, indicating that the tensile deformation is large enough to make the texture along the building direction disappear, but when the strain rate reaches 600 s 1, the preferred orientations in both phases are not along the tensile direction, which can be ascribed to the relatively small deformation along the tensile direction" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_35_0001419_978-94-007-7401-8-Figure3.37-1.png", + "original_path": "designv11-35/openalex_figure/designv11_35_0001419_978-94-007-7401-8-Figure3.37-1.png", + "caption": "Fig. 3.37 Overactuated PMs with coupled Sch\u00f6nflies motions of types 4RRRRP (a) and 4PRRRP (b) defined by MF \u00bc SF \u00bc 4; RF\u00f0 \u00de \u00bc v1; v2; v3;xd\u00f0 \u00de, TF = 0, NF = 2, limb topology R||R\\R||R\\P (a) and P\\R\\R||R\\P (b)", + "texts": [ + " 4PRRRR (Fig. 3.34b) P\\R||R\\R||R (Fig. 3.3p) The second joints of the four limbs have parallel axes (continued) 244 3 Overactuated Topologies with Coupled Sch\u00f6nflies Motions Table 3.2 (continued) No. PM type Limb topology Connecting conditions 30. 4RPRRR (Fig. 3.35a) R\\P\\kR\\R||R (Fig. 3.3q) Idem No. 16 31. 4RRRPR (Fig. 3.35b) R||R\\R||P||R (Fig. 3.3r) Idem No. 16 32. 4RRRPR (Fig. 3.36a) R||R\\R||P||R (Fig. 3.3s) Idem No. 16 33. 4RRRRP (Fig. 3.36b) R||R\\R||R||P (Fig. 3.3t) Idem No. 16 34. 4RRRRP (Fig. 3.37a) R||R\\R||R\\P (Fig. 3.3u) Idem No. 16 35. 4PRRRP (Fig. 3.37b) P\\R\\R||R\\P (Fig. 3.3v) Idem No. 29 36. 4RRPRP (Fig. 3.38a) R\\R||P||R\\P (Fig. 3.3w) Idem No. 16 37. 4RPRRP (Fig. 3.38b) R\\P||R||R\\P (Fig. 3.3x) Idem No. 16 38. 4RRPRP (Fig. 3.39a) R\\R||P||R\\P (Fig. 3.3y) Idem No. 16 39. 4RRRPP (Fig. 3.39b) R\\R||R||P\\P (Fig. 3.3z) Idem No. 16 40. 4PRRRP (Fig. 3.40a) P||R||R\\R\\P (Fig. 3.3z1) The fourth joints of the four limbs have parallel axes 41. 4RPRRP (Fig. 3.40b) R||P||R||R\\P (Fig. 3.3a) Idem No. 40 42. 4RPRRP (Fig. 3.41a) R||P||R\\R\\P (Fig. 3.3b0) Idem No. 40 43" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_35_0001419_978-94-007-7401-8-Figure3.73-1.png", + "original_path": "designv11-35/openalex_figure/designv11_35_0001419_978-94-007-7401-8-Figure3.73-1.png", + "caption": "Fig. 3.73 Overactuated PMs with coupled Sch\u00f6nflies motions of types 4PaPRP (a) and 4PaRPP (b) defined by MF \u00bc SF \u00bc 4; RF\u00f0 \u00de \u00bc v1; v2; v3;xd\u00f0 \u00de, TF = 0, NF = 18, limb topology Pa||P||R\\P (a) and Pa||R\\P\\kP (b)", + "texts": [ + "66b) P||Pa\\R\\P (Fig. 3.50i0) The before last joints of the four limbs have parallel axes 24. 4RPPaP (Fig. 3.67) R||P||Pa\\P (Fig. 3.50w) Idem No. 1 25. 4RPRPa (Fig. 3.68) R||P||R||Pa (Fig. 3.50x) Idem No. 1 26. 4PaPPR (Fig. 3.69a) Pa||P\\P\\kR (Fig. 3.50y) Idem No. 1 27. 4PaPPR (Fig. 3.69b) Pa\\P\\kP||R (Fig. 3.50z) Idem No. 1 28. 4PaPRP (Fig. 3.70) Pa\\P\\kR||P (Fig. 3.50z1) Idem No. 1 29. 4PaRRP (Fig. 3.71) Pa||R||R||P (Fig. 3.50a0) Idem No. 1 30. 4PaRPR (Fig. 3.72) Pa||R||P||R (Fig. 3.50b0) Idem No. 1 31. 4PaPRP (Fig. 3.73a) Pa||P||R\\P (Fig. 3.50c0) Idem No. 1 32. 4PaRPP (Fig. 3.73b) Pa||R\\P\\kP (Fig. 3.50d0) Idem No. 1 33. 4PRPPa (Fig. 3.74a) P\\R\\P||Pa (Fig. 3.50e0) The second joints of the four limbs have parallel axes 34. 4PRPPa (Fig. 3.74b) P\\R\\P\\Pa (Fig. 3.50f0) Idem No. 33 35. 4PRPaP (Fig. 3.75a) P\\R\\Pa\\\\P (Fig. 3.50g0) Idem No. 33 36. 4PPRPa (Fig. 3.75b) P\\P\\\\R\\Pa (Fig. 3.50h0) The third joints of the four limbs have parallel axes 37. 4PPPaR (Fig. 3.76a) P\\P||Pa\\\\R (Fig. 3.50j0) The last joints of the four limbs have parallel axes (continued) 3.2 Topologies with Complex Limbs 359 Table 3" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_35_0000686_j.mechmachtheory.2021.104433-Figure19-1.png", + "original_path": "designv11-35/openalex_figure/designv11_35_0000686_j.mechmachtheory.2021.104433-Figure19-1.png", + "caption": "Fig. 19. Metamorphic epicyclic bevel gear train based on force constraint.", + "texts": [ + " Then the corresponding configuration transforms from configuration 2 to configuration 1, and the mobility has degenerated from 2 to 1. The metamorphic epicyclic external gear train is in revolution configuration (Fig. 15(b)). Likewise, a set of metamorphic epicyclic gear trains are constructed by adding force constraints to control the constraint condition. As shown in Fig. 18\u201320, the spring pin which consists of a block and spring provides a certain constraint force. For instance, the metamorphic epicyclic bevel gear train is in revolution configuration (Fig.19 (b)) when the bevel gear arm has been locked to the sun bevel gear by the spring pin. At this moment, the angular velocity \u03c9b2 is equal to \u03c9b1, which means the corresponding motion branch has mobility 1. Until driving force drive the bevel gear arm to overcome the constraint force (Fig. 19(a)), the metamorphic epicyclic bevel gear train transforms to rotation configuration which is similar to the example discussed above and has mobility 2 H. Yuan et al. Mechanism and Machine Theory 166 (2021) 104433 By combining geometrical constraint and force constraint, a set of metamorphic epicyclic gear trains can be obtained based on a combination constraint. As shown in Figs. Figures 21-23, the combination constraint can provide both geometrical limitation and constraint forces. The metamorphic epicyclic gear-rack train in Fig" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_35_0000395_978-3-030-67750-3_16-Figure17.3-1.png", + "original_path": "designv11-35/openalex_figure/designv11_35_0000395_978-3-030-67750-3_16-Figure17.3-1.png", + "caption": "Fig. 17.3 Rolling resistance", + "texts": [ + " The tractive force must fulfill the requirements of vehicle dynamics to overcome forces such as the rolling resistance, gravitational, and aerodynamically which are summed together as the road load force FRL as shown in Fig. 17.2 (Arof et al. 2020a, b, c, d). Thegravitational force Fg dependson the slopeof the roadway, as shown inEq. (17.1). Fg = mg sin \u03b1 (17.1) \u03b1 is the grade angle, m is the total mass of the vehicle, g is the gravity constant. The hysteresis of the tire material causes it at the contact surfaces with the roadway. The centroid of the vertical forces on the wheel moves forward when the tire rolls. Therefore, from beneath the axle toward the direction of motion by the vehicle, as shown in Fig. 17.3. Tractive Force The tractive forcewas used to overcome the Froll force alongwith the gravity force and the aerodynamic drag force. The rolling resistance has been minimized by keeping the tires as inflated as possible by reducing the hysteresis. The ratio of retarding forces due to rolling resistance and the vertical load on the wheel known as the coefficient of rolling resistance C0. The rolling resistance force is given by Eroll = \u23a7 \u23a8 \u23a9 sgn[vxt ]mg(C0 + C1v2xT ) if vxt = 0 FTR \u2212 FgxT if Vxt = 0 and \u2223 \u2223FTR \u2212 FgxT \u2223 \u2223 \u2264 C0mg sgn(FTR \u2212 FgxT )(C0mg) if Vxt = 0 and \u2223 \u2223FTR \u2212 FgxT \u2223 \u2223 > C0mg \u23ab \u23ac \u23ad (17" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_35_0000577_s40313-021-00754-5-Figure10-1.png", + "original_path": "designv11-35/openalex_figure/designv11_35_0000577_s40313-021-00754-5-Figure10-1.png", + "caption": "Fig. 10 Aircraft forces, velocities, moments, and orientations definitions", + "texts": [ + " Before the doublet input was given, the aircraft is conditioned in trim and level conditions (the aircraft is considered in a state of straight flight and the right-left wing of the aircraft is in the same state horizontally). The diagram of the flight data acquisition procedure is shown in Fig.\u00a08. The recorded flight data were then processed and converted to a variable that was used for empirical modeling or system identification as shown in Fig.\u00a09. In the fixed-wing UAV flight, there are many forces, velocities, moments, and orientations that work. All could be summarized and visualized in Fig.\u00a010. In a flight, the aircraft is not always horizontally straight when flying straight ahead (on the X-axis, see Fig.\u00a011) but has an angle-of-attack angle commonly written \u03b1 (alpha). The aircraft angle-of-attack is the deflection angle of the aircraft on the x-axis caused by the airplane\u2019s wing geometry. Similar to angle-of-attack, there is also a sideslip angle (Fig.\u00a011), which is usually written \u03b2 (beta). The difference is that the sideslip angle is the deflection on the y-axis due to the air/wind from the side that concerns the vertical stabilizer of the aircraft" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_35_0000174_j.matpr.2021.01.640-Figure13-1.png", + "original_path": "designv11-35/openalex_figure/designv11_35_0000174_j.matpr.2021.01.640-Figure13-1.png", + "caption": "Fig. 13. Mode shapes 1.", + "texts": [ + " The maximum distortion energy failure theory assumes failure by yielding in a more complicated loading situation to occur when the distortion energy in the material reaches the same value as in a tension test at yield. The maximum strain is obtained as 6.5176e-7 and minimum strain is obtained as 3.3411e-9 Engineering analysis Stress on Metal Matrix Composites as shown in Fig. 12.Stress analysis on metal matrix composite, von mises elastic theory is applied, the stress value obtained as maximum 74.301 MPa and minimum as 0.38088 MPa. etal Matrix Composites. By Modal analysis the characteristic frequencies and mode shapes shown in Fig. 13. In the event that the shat revolves at its characteristic recurrence, it very well may be seriously vibrated. The modal investigation performed to locate the normal frequencies Fig. 14 Fig. 15. Modal analysis in structural mechanics is determined the natural mode shapes and frequencies of an object or structure during free vibration. In this analysis the frequency varies 196.67 Hz, the mode shape level is 1. The mode shape can be extended on level 1, due to the rotations. The mode shapes level can be extend upto level 5" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_35_0001233_978-3-319-09435-9_3-Figure2-1.png", + "original_path": "designv11-35/openalex_figure/designv11_35_0001233_978-3-319-09435-9_3-Figure2-1.png", + "caption": "Fig. 2. Displacements and spring forces. The figure shows an angular displacement (\u03b1) between two adjacent segments (segment j and j + 1 respectively) in the radial direction. A torsion spring couple Mj,j+1 = kt,j \u00b7 \u03b1 is applied between the two. kt,j is fixed for all the hinge joints having the same distance from the body. The figure also shows a displacement (d, vector) between the center of mass of two adjacent segments in the longitudinal direction (segments i and i+ 1). A force Fi,i+1 = kl,i \u00b7 d is applied between the two centers of mass, with kl,i being fixed for all segments at the same distance from the body.", + "texts": [ + " We modeled the wing by means of 10 adjacent radial chains, each being composed by a variable number (10 in the longest chain) of rigid prismatic segments (22 x 24 x 5 mm). This discretization was selected as a trade-off between accuracy and computational cost. The resulting fin shape is approximately semi-elliptic, with root chord of 240 mm and span of 220 mm, following the experimental set-up presented in [4]. The segments are interconnected by means of hinge joints with torsion springs (modeling bending elasticity). Adjacent chains are free to move one respect to the others, with interactions among them being modeled by means of linear springs (Fig. 2). This structure is able to mimic the passive compliance using the minimum number of DOFs. No structural damping was included as it was assumed to be negligible with respect to the one provided by fluid drag forces. Only the joints belonging to the frontal chain are actuated, by means of simple sinusoidal oscillators (the same oscillation is applied at each actuated joint). Although it may seem oversimplified, the resulting global actuation mimics the one observed in batoids, that do not appear to adopt more complex control strategies (traveling waves are not directly actuated, instead they emerge from the interaction with the environment)" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_35_0001419_978-94-007-7401-8-Figure3.113-1.png", + "original_path": "designv11-35/openalex_figure/designv11_35_0001419_978-94-007-7401-8-Figure3.113-1.png", + "caption": "Fig. 3.113 4RRRPaR-type overactuated PM with coupled Sch\u00f6nflies motions defined by MF \u00bc SF \u00bc 4; RF\u00f0 \u00de \u00bc v1; v2; v3;xd\u00f0 \u00de, TF = 0, NF = 14, limb topology R||R\\R\\Pa\\kR", + "texts": [ + "51r) The cylindrical joints of the four limbs have parallel axes 19. 4PaPaC (Fig. 3.107) Pa||Pa||C (Fig. 3.51s) Idem No. 18 20. 4PaCPa (Fig. 3.108) Pa||C||Pa (Fig. 3.51t) Idem No. 18 21. 4RPaRRR (Fig. 3.109) R\\Pa\\kR\\R||R (Fig. 3.52a) Idem No. 13 22. 4RRRRPa (Fig. 3.110a) R||R\\R||R||Pa (Fig. 3.52b) Idem No. 15 23. 4RRRRPa (Fig. 3.110b) R\\R||R\\R||Pa (Fig. 3.52d) Idem No. 15 24. 4PaRRRR (Fig. 3.111) Pa||R\\R||R\\kR (Fig. 3.52c) Idem No. 13 25. 4PaRRRR (Fig. 3.112) Pa||R||R\\R||R (Fig. 3.52e) Idem No. 17 26. 4RRRPaR (Fig. 3.113) R||R\\R\\Pa\\kR (Fig. 3.52f) Idem No. 12 27. 4RRRRPa (Fig. 3.114) R||R\\R||R\\kPa (Fig. 3.52g) Idem No. 12 28. 4PRRRPa (Fig. 3.115) P\\R\\R||R\\kPa (Fig. 3.52h) The second joints of the four limbs have parallel axes 29. 4RRPRPa (Fig. 3.116) R\\R||P||R\\kPa (Fig. 3.52i) Idem No. 12 30. 4RPRRPa (Fig. 3.117) R\\P||R||R\\kPa (Fig. 3.52j) Idem No. 12 31. 4RRPRPa (Fig. 3.118) R\\R||P||R\\kPa (Fig. 3.52k) Idem No. 12 32. 4RRRPPa (Fig. 3.119) R\\R||R||P\\kPa (Fig. 3.52l) Idem No. 12 33. 4PRRRPa (Fig. 3.120) P||R||R\\R||Pa (Fig" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_35_0000477_j.procir.2021.05.085-Figure6-1.png", + "original_path": "designv11-35/openalex_figure/designv11_35_0000477_j.procir.2021.05.085-Figure6-1.png", + "caption": "Figure 6. Non-assembly with normal support placed only at build plate", + "texts": [ + " As most of the support structures lie on the outer side, so removal process of the tree support structure is finer. In this case, the support structure type is \u2018normal\u2019 and placed only at the build plate to support the non-assembly from the base. The purpose of this support type is quite similar to the brim is to keep the non-assembly components in contact with the build plate of the printer. No other support has been placed at the joints and below the overhanging features of the nonassembly as shown in Fig. 6. This support offers lesser time and support material among all cases studied. Due to the absence of the abundant support structure, the least build time (142 min) is observed as compared to other support structures. A slight support structure at the base has a weight of only 1.45 g. Further, it is quite easy to remove the support from the base. 730 WK Leung et al. / Procedia CIRP 100 (2021) 726\u2013731 Leung et / Procedia CIRP 00 (2021) 000\u2013000 5 It has been observed that the circular feature in the structure of the non-assembly that are made without any support has a fine surface and accuracy due to lack of physical contact of non-assembly features with any support structure and reduced post processing" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_35_0001419_978-94-007-7401-8-Figure5.44-1.png", + "original_path": "designv11-35/openalex_figure/designv11_35_0001419_978-94-007-7401-8-Figure5.44-1.png", + "caption": "Fig. 5.44 3PaPaPR-1RPaPaP-type fully-parallel PM with uncoupled Sch\u00f6nflies motions defined by MF = SF = 4, RF\u00f0 \u00de \u00bc \u00f0v1; v2; v3;xb\u00de, TF = 0, NF = 30 limb topology Pa\\Pa\\\\P\\\\R, Pa\\Pa\\\\P\\kR and R||Pa||Pa||P", + "texts": [ + "4d) R||Pa||Pat||P (Fig. 5.4l) 6. 3PaPPaR1RPaPatP (Fig. 5.41b) Pa\\P\\\\Pa\\kR (Fig. 5.4b) Idem No. 1 Pa\\P\\\\Pa||R (Fig. 5.4c) R||Pa||Pat||P (Fig. 5.4l) 7. 3PaPPaR-1RPPaPat (Fig. 5.42a) Pa||P\\Pa\\kR (Fig. 5.4a) Idem No. 3 Pa||P\\Pa||R (Fig. 5.4d) R||P||Pa||Pat (Fig. 5.4n) 8. 3PaPPaR-1RPPaPat (Fig. 5.42b) Pa\\P\\\\Pa\\kR (Fig. 5.4b) Idem No. 3 Pa\\P\\\\Pa||R (Fig. 5.4c) R||P||Pa||Pat (Fig. 5.4n) 9. 3PaPaPR-1RPaPaP (Fig. 5.43) Pa\\Pa\\kP\\\\R (Fig. 5.4f) Idem No. 1 Pa\\Pa\\kP\\kR (Fig. 5.4h) R||Pa||Pa||P (Fig. 5.4k) 10. 3PaPaPR-1RPaPaP (Fig. 5.44) Pa\\Pa\\\\P\\\\R (Fig. 5.4g) Idem No. 1 Pa\\Pa\\\\P\\kR (Fig. 5.4e) R||Pa||Pa||P (Fig. 5.4k) 11. 3PaPaPR1RPaPatP (Fig. 5.45) Pa\\Pa\\kP\\\\R (Fig. 5.4f) Idem No. 1 Pa\\Pa\\kP\\kR (Fig. 5.4h) R||Pa||Pat||P (Fig. 5.4l) 12. 3PaPaPR1RPaPatP (Fig. 5.46) Pa\\Pa\\\\P\\\\R (Fig. 5.4g) Idem No. 1 Pa\\Pa\\\\P\\kR (Fig. 5.4e) R||Pa||Pat||P (Fig. 5.4l) 13. 3PaPaPR-1RPPaPa (Fig. 5.47) Pa\\Pa\\kP\\\\R (Fig. 5.4f) Idem No. 3 Pa\\Pa\\kP\\kR (Fig. 5.4h) R||P||Pa||Pa (Fig. 5.4m) 14. 3PaPaPR-1RPPaPa (Fig. 5.48) Pa\\Pa\\\\P\\\\R (Fig. 5.4g) Idem No. 3 Pa\\Pa\\\\P\\kR (Fig" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_35_0000371_tmag.2021.3076134-Figure2-1.png", + "original_path": "designv11-35/openalex_figure/designv11_35_0000371_tmag.2021.3076134-Figure2-1.png", + "caption": "Fig. 2. Configuration of the proposed stator-PM spiral translator PMTFLG. (a) Sectional view. (b) 3-D full view. (c) Stator part. (d) Translator part.", + "texts": [ + " However, this also causes the area conflict between PMs and armature windings, which restricts the power density [11], [17]. 0018-9464 \u00a9 2021 IEEE. Personal use is permitted, but republication/redistribution requires IEEE permission. See https://www.ieee.org/publications/rights/index.html for more information. Authorized licensed use limited to: California State University Fresno. Downloaded on June 26,2021 at 18:41:18 UTC from IEEE Xplore. Restrictions apply. In this article, a stator-PM spiral translator PMTFLG (ST-PMTFLG) topology is proposed, as shown in Fig. 2. Armature winding and PMs are arranged in outer stator and inner stator, respectively, so the total slot area of the generator is enlarged and the thermal dissipation capability is improved [13]\u2013[19]. As shown in Fig. 2(c), the inner stator consists of three inner stator rings and PMs mounted on them. The outer stator consists of 3 outer stator rings and 12 coils. Each coil passes through three outer stator rings. Therefore, the output power of PMTFLG can be enhanced by increasing the number of outer and inner stator rings. The translator is made of spiral steel blocks and nonmagnetic blocks. In order to facilitate manufacturing, the translator is divided by four blue dotted lines, as shown in Fig. 2(d). Two nonmagnetic ribs are required to assemble the four equal parts of the translator. Because of the existence of nonmagnetic ribs, the air gap will inevitably increase. The armature windings are placed in the outer stator slots and the winding diagrams of the 12-stator-pole ST-PMTFLGs with different translator poles are shown in Fig. 3. Similar to the magnetically geared machine analyzed in [20], ST-PMTFLG is suitable to be used in low-speed application. PMs are mounted on inner stator surface, so the volume of PM is reduced when using in long-stroke application [5]", + " To sum up, the proposed ST-PMTFLG is a good candidate in direct-drive, low-speed, and long-stroke application, like DD-WEC [12]. This article is organized as follows. In Section II, the machine configuration, operation principle, and 2-D simplification model are introduced. In Section III, some design parameters are optimized and the electromagnetic performance of the proposed ST-PMTFLG is evaluated, followed by conclusion in Section IV. The configuration of the proposed ST-PMTFLG is shown in Fig. 2. Armature winding is placed in the outer stator while PMs are alternately mounted on the inner stator surface. The translator is made of spiral steel blocks and nonmagnetic blocks. Similar to the magnetically geared machine and magnetic gear, armature winding pole-pair paw, translator number Nt , and stator tooth number pst match [19], [20] Nt = pst \u00b1 paw. (1) In 12/11 and 12/13 (outer stator slot number/translator number) ST-PMTFLGs, pst = 12, paw = 1, and Nt = 11 and 13, respectively. In 12/10 and 12/14 topologies, pst = 12, paw = 2, and Nt = 10 and 14, respectively [19]", + " The translator in the generator is made up of spiral steel blocks and spiral nonmagnetic blocks. The relationship between the translator position in electric and mechanical degrees is the same as that in the conventional stator-excitation machines [18] \u03b8e = Nt\u03b8m (2) where \u03b8e and \u03b8m are the translator positions in electric and mechanical degrees, respectively, and Nt is the translator pole number, i.e., the number of the moving spiral steel blocks. The proposed ST-PMTFLG is a kind of transverse flux generator. The moving direction of the translator is along the zaxis, as shown in Fig. 2. The 3-D finite-element analysis (FEA) costs too much time; therefore, 2-D simplification model is needed to reduce the optimization time [13]. The spiral translator of the proposed ST-PMTFLG is the only moving part. When the spiral translator moves along the z-axis, it can be considered that it is rotating in the xy plane, like a partitioned stator flux-reversal permanent magnet (FRPM) machine [17]. The linear velocity along the z-axis can be converted to the equivalent angular speed of rotation in the xy plane, as in \u03b8twist \u03c9 = lt v (3) where \u03b8twist is the translator twist arc, lt is the translator length, \u03c9 is the equivalent angular speed of rotation in the xy plane, and v is the moving speed of translator along the z-axis" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_35_0000746_lra.2021.3103641-Figure2-1.png", + "original_path": "designv11-35/openalex_figure/designv11_35_0000746_lra.2021.3103641-Figure2-1.png", + "caption": "Fig. 2. Diagrams of vertical grouser pushing soil backward and blade bulldozing soil: (a) interaction between vertical grouser and soil; and (b) interaction between blade and soil.", + "texts": [ + "{ FDP = (1/r)TD + 0FV (a) TD = rFDP + 0FV (b) (1) When a WPR traverses sandy terrain, uniformly studded grousers can be used to improve its mobility performance by pushing soil particles backward and transferring soil-metal friction into terrain internal friction. To accurately model the contact mechanics between a grouser-wheel and sandy terrain, the interaction between a single grouser and terrain should be considered first. 1) Research on Grouser Pushing Action: A vertical grouser pushing soil particles backward was first considered, as shown in Fig. 2(a) [17], where hG is the grouser height, and rmax is the maximum wheel radius (rmax = r + hG). Interaction between the vertical grouser and terrain can be generalized as a specific case where an inclined blade is pushing soil where the inclined angle \u03b2 0, as shown in Fig. 2(b). The TABLE I VALUES OF PARAMETERS USED IN COMPUTING MAXIMUM PUSHING FORCE Fig. 3. FPmax acting on a vertical grouser versus its penetration depth. soil behind the blade will be brought into a state of passive failure. If the blade is relatively wide compared with its penetration depth, which holds for wheel grousers, the case can be treated as two-dimensional. The reaction force FP exerted on the grouser side face can be determined by integrating the bulldozing pressure Rb generated per unit width along the penetration depth, as shown in Eq" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_35_0001419_978-94-007-7401-8-Figure3.47-1.png", + "original_path": "designv11-35/openalex_figure/designv11_35_0001419_978-94-007-7401-8-Figure3.47-1.png", + "caption": "Fig. 3.47 Overactuated PMs with coupled Sch\u00f6nflies motions of types 4RCRP (a) and 4RRCP (b) defined by MF \u00bc SF \u00bc 4; RF\u00f0 \u00de \u00bc v1; v2; v3;xd\u00f0 \u00de, TF = 0, NF = 2, limb topology R\\C||R\\\\P (a) and R\\R||C\\\\P (b)", + "texts": [ + "3f0) Idem No. 17 47. 4RRPPR (Fig. 3.43b) R||R||P\\P\\\\R (Fig. 3.3g0) Idem No. 17 48. 4RPRPR (Fig. 3.44a) R||P||R\\P\\\\R (Fig. 3.3h0) Idem No. 17 49. 4RPRPR (Fig. 3.44b) P\\P\\kR||R\\\\R (Fig. 3.3i0) Idem No. 17 50. 4PPRRR (Fig. 3.44c) P\\P||R||R\\\\R (Fig. 3.3j0) Idem No. 17 51. 4CRRR (Fig. 3.45a) C||R\\R||R (Fig. 3.3k0) Idem No. 17 52. 4RCRR (Fig. 3.45b) R||C\\R||R (Fig. 3.3l0) Idem No. 17 53. 4RRCR (Fig. 3.46a) R||R\\C||R (Fig. 3.3m0) Idem No. 16 54. 4RRRC (Fig. 3.46b) R||R\\R||C (Fig. 3.3n0) Idem No. 16 55. 4RCRP (Fig. 3.47a) R\\C||R\\\\P (Fig. 3.3o0) Idem No. 16 56. 4RRCP (Fig. 3.47b) R\\R||C\\\\P (Fig. 3.3p0) Idem No. 16 57. 4CRRP (Fig. 3.48a) C||R\\R\\P (Fig. 3.3q0) Idem No. 44 58. 4RCRP (Fig. 3.48b) R||C\\R\\P (Fig. 3.3r0) Idem No. 44 59. 4CRPR (Fig. 3.49a) C||R\\P\\\\R (Fig. 3.3s0) Idem No. 17 60. 4PCRR (Fig. 3.49b) P\\C||R\\\\R (Fig. 3.3t0) Idem No. 17 3.1 Topologies with Simple Limbs 245 Table 3.3 Structural parametersa of parallel mechanisms in Figs. 3.4, 3.5, 3.6, 3.7, 3.8, 3.9, 3.10, 3.11, 3.12, 3.13, 3.14, 3.15, 3.16, 3.17, 3.18, 3.19, 3.20, 3.21, 3.22, 3.23, 3.24, 3.25, 3.26 No" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_35_0000095_s00170-021-06769-1-Figure7-1.png", + "original_path": "designv11-35/openalex_figure/designv11_35_0000095_s00170-021-06769-1-Figure7-1.png", + "caption": "Fig. 7 Knife holder design", + "texts": [ + " For examination, the sample is mounted to the aluminum holder using woodscrews from the bottom side. To avoid initial hit excitation, which can influence measured force values, the sample is squeezed between two polyurethane blocks (homogeneous density of 530 kg/m3). Samples were examined in cutting direction B according to Kivimaa [3]. For cutting examinations, a standardized single knife (LEITZ, Turnblade knife, HW-05: 30 \u00d7 12 \u00d7 1.5) was used. It was placed to the knife holder (Fig. 6) with a constant cutting angle \u03b4 = (\u03b1 + \u03b2) of 60\u00b0. The knife holder (Fig. 7) has been optimized to achieve the highest natural frequency possible. The topology optimization of shape was used to achieving the best stiffness\u2013mass ratio. The knife holder is made of high stiffness steel. It weighs only 146 g; nevertheless, it can resist the stress of approx. 1000 MPa from the force sensor pre-load. Experiments were performed at speeds of 10\u201380 m/s, in steps of 10 m/s. Investigated chip thicknesses were 0.05, 0.1, 0.15, and 0.2 mm. For testing device accuracy, chip thicknesses up to 0" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_35_0000159_j.matpr.2021.02.306-Figure2-1.png", + "original_path": "designv11-35/openalex_figure/designv11_35_0000159_j.matpr.2021.02.306-Figure2-1.png", + "caption": "Fig. 2. Different parts of (a) filament extrusion sy", + "texts": [ + " are parameters which are decided during process planning [9,10]. This printer uses thermoplastic material for printing. Material like Acrylonitrile Butadiene Styrene (ABS), Poly Lactic Acid (PLA), etc. are used [17,18,19,20]. A spool contains the filament is attached in the printer and with the help of geared stepper motor, filament is pushed into the extruder. The hot end of the extruder melts the filament and then it is extruded through nozzle into a platform. The object is created by layer-by-layer deposition of the filament (see Fig. 2(a) and Fig. 3(a)). A typical FFF based 3D printer is consist of mainly two types of components \u2013mechanical components and electronics components (see Fig. 2(a) and Fig. 3(a)) [21]. All the components of FFF 3D printer are mentioned in Table 1. Linear rails are the guiding track for the components. A rail is the riding track over which carriage move and this rail is supported d addi- from underneath. There motion of carriage is restricted, it can\u2019t move in any direction it can only move in the direction of rail. There are different types of rails. Traditional rail is a thin and long rectangular piece of steel with two dovetails that run along either side", + " In pellet extrusion material is in form of small particle instead of wire. There many benefits of pellets over filament. In filament form only those material could be used which have the ability to convert in form of wires [2324]. So, pellet extrusion removes this limitation. Waste could also be reused for fabrication using pellet extrusion [25]. Changing the material and color of material during fabrication is very easy using pellet extrusion. Setup for pellet extrusion also consist of mechanical and electronics parts (see Fig. 2(b) and Fig. 3(b)) [21]. Table 2 shows the different parts used in pellet extrusion setup. Many components are similar to filament extrusion setup, but some different parts are discussed below. Extruder or auger bit is the central part of pellet extrusion system. This is a threaded rod or screw like structure to carry the pellets from hopper to heating block via barrel. This part is made of steel. Design of this extrusion screw plays an important role in material extrusion (see Fig. 4). Barrel is a cylindrical part made of steel" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_35_0000213_j.triboint.2021.106993-Figure9-1.png", + "original_path": "designv11-35/openalex_figure/designv11_35_0000213_j.triboint.2021.106993-Figure9-1.png", + "caption": "Fig. 9. Geometry and mesh for TPJBs; (a) smooth TPJB, (b) pocketed TPJB", + "texts": [ + " The eccentricity ratio can be expressed as \u03b5= e Cl,b (17) The primary purpose of the novel TPJB design is to reduce the drag torque acting opposite to the rotation direction, as seen in Fig. 8. The mathematical expressions of the drag torque (DT) and power loss (PL) are given in Eqs. (18) and (19), respectively. DT = \u222b As ( Rs \u00d7 \u03bcf \u2202ucir \u2202r ) dAs (18) PL=DT \u00d7 \u03c9s (19) J. Yang and A. Palazzolo Tribology International 159 (2021) 106993 where Rs is the shaft radius, As is the journal surface area, ucir and \u03c9s are the circumferential velocity and rotation speed, respectively. As described earlier, the shaft, pad, and fluid domains are considered in the CFD simulation. Fig. 9 depicts the modeled TPJB with the mesh for the smooth and pocketed TPJB. A mesh independence study was conducted on the Thermo-Hydrodynamic (THD) Model, and is summarized in Fig. 10. The number of elements was selected to be 433,436 based on the results in Fig. 10 and the corresponding computation times. The final mesh was identical for the smooth and pocketed bearings as illustrated in Fig. 9. Comparison between the simulation results for the smooth and pocketed TPJBs demonstrates the benefits of the latter. Fig. 11 shows the pressure contours when the pockets are installed on pad1, pad2, and J. Yang and A. Palazzolo Tribology International 159 (2021) 106993 pad3 of the pocketed TPJB. The color contour levels are adjusted to focus on the pressure distribution of the upper pads, since their pressures are much lower than that in the bottom pads. As shown in Fig. 11(b), the peak pressures in the pocketed TPJB are produced at the step and pocket trailing edges, which helps stabilize the pad tilting motion" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_35_0001419_978-94-007-7401-8-Figure2.36-1.png", + "original_path": "designv11-35/openalex_figure/designv11_35_0001419_978-94-007-7401-8-Figure2.36-1.png", + "caption": "Fig. 2.36 Fully-parallel PMs with coupled Sch\u00f6nflies motions of types 4PaPRP (a) and 4PaRPP (b) defined by MF \u00bc SF \u00bc 4; RF\u00f0 \u00de \u00bc \u00f0v1; v2; v3;xd\u00de, TF = 0, NF = 18, limb topology Pa\\P||R\\P (a) and Pa\\R\\P\\kP (b)", + "texts": [ + "32a) P\\Pa||P\\\\R (Fig. 2.21m) Idem No. 4 14. 4PPaRP (Fig. 2.32b) P\\Pa\\kR\\kP (Fig. 2.21n) Idem No. 4 15. 4RPPaP (Fig. 2.33a) R\\P||Pa\\kP (Fig. 2.21o) The first revolute joints of the four limbs have parallel axes 16. 4RPPPa (Fig. 2.33b) R||P\\P||Pa (Fig. 2.21p) Idem No. 15 17. 4PaPPR (Fig. 2.34a) Pa||P\\P||R (Fig. 2.21q) Idem No. 4 18. 4PaPPR (Fig. 2.34b) Pa\\P\\kP\\kR (Fig. 2.21r) Idem No. 4 19. 4PaRPP (Fig. 2.35a) Pa\\R||P\\P (Fig. 2.21s) Idem No. 11 20. 4PaPRP (Fig. 2.35b) Pa||P\\R||P (Fig. 2.21t) Idem No. 5 21. 4PaPRP (Fig. 2.36a) Pa\\P||R\\P (Fig. 2.21u) Idem No. 11 22. 4PaRPP (Fig. 2.36b) Pa\\R\\P\\kP (Fig. 2.21v) Idem No. 5 23. 4PCPa (Fig. 2.37a) P\\C\\Pa (Fig. 2.21w) Idem No. 5 24. 4PPaC (Fig. 2.37b) P||Pa\\C (Fig. 2.21x) The last joints of the four limbs have parallel axes 25. 4CPPa (Fig. 2.38a) C\\P||Pa (Fig. 2.21y) The first joints of the four limbs have parallel axes 26. 4PaPC (Fig. 2.38b) Pa||P\\C (Fig. 2.21z) Idem No. 24 27. 4PaCP (Fig. 2.39) Pa\\C\\P (Fig. 2.21z1) Idem No. 11 28. 4PPaPaR (Fig. 2.40) P\\Pa\\\\Pa\\kR (Fig. 2.22a) Idem No. 4 29. 4PPaRPa (Fig. 2.41) P\\Pa||R\\Pa (Fig. 2.22b) The axes of the revolute joints connecting the two parallelogram loops of the four limbs are parallel (continued) 2" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_35_0001411_978-3-030-55061-5-Figure14-1.png", + "original_path": "designv11-35/openalex_figure/designv11_35_0001411_978-3-030-55061-5-Figure14-1.png", + "caption": "Fig. 14. The second cycle of the mechanism Fig. 15. Rotation around Z axis", + "texts": [ + " The Calculus of Mobility for Parallelogram Mechanism For the mechanism modeled with ADAMS shown in Fig. 9, whose sides form parallelograms, the calculation of mobility by the segmentation method in detail is presented in works [1] and [3]. For the first independent cycle (Fig. 10) the mobility number is 3. The simulation of the three degrees of freedom of the end-effector in the first cycle is presented in Figs. 11, 12 and 13. Fig. 9. Parallelogram mechanism modeled with ADAMS For the second cycle (Fig. 14) the mobility number is 2. A rotation around the Z axis will be possible (Fig. 15) and the two independent translations, X (Fig. 16) and Y (Fig. 17), will appear as impossible (for example in Fig. 18). the base and an actuator was placed along the X axis, between the cross element and the base (Fig. 19). Following the simulation, it is observed visually (Fig. 20), but also in the superimposed kinematic diagrams (Fig. 21), the existence of a degree of freedom and the appearance of two dependent movements, translation along the X and Y axis" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_35_0000435_j.cirp.2021.04.050-Figure2-1.png", + "original_path": "designv11-35/openalex_figure/designv11_35_0000435_j.cirp.2021.04.050-Figure2-1.png", + "caption": "Fig. 2. Mobile manipulator for workpiece exchange automation.", + "texts": [ + " A fiducial marker is set inside a machine tool, and a camera is mounted on the robot manipulator to take images of the marker. This paper also proposes the camera localization process to enhance accuracy of the pose estimation. The effectiveness of the process was verified through experiments. The system overview is described in Section 2. The proposed pose estimation method is explained in Section 3. Experimental procedures and results are shown in Section 4while a conclusion is given in Section 5. The AGV\u2019s structure was newly designed as shown in Fig. 2(a). It has two drive wheels with suspension and four omni wheels supporting its weight. In general, an AGV with two-wheel drive has freely rotatable casters, but the casters need large space to rotate, resulting in enlarging the size of a vehicle. Using omni wheels instead of the casters makes it possible to reduce the vehicle size and lower the center of gravity. This AGV is capable of driving over steps of up to 30mm (e.g. cable ducts, imperfections in flooring etc.) and therefore can work without modifying the preexisting factory environment", + " Also, the transformation from the robot frame to the camera frame robot cameraM can be determined by multiplying robot gripperM and the fixed one from the gripper frame to the camera frame gripper cameraM derived from a CAD model. The superscript character of the transformation means a source frame while the subscript character indicates a destination frame. Also, each transformation is expressed in the form of a homogeneous coordinate system. An appearance of the mobile manipulator is shown in Fig. 2(b). Mapping [11] is the process of making a static map for navigation of AGV. While the operator manually moves the AGV in the factory, Simultaneous Localization And Mapping (SLAM) algorithms are used to build the map of the factory. Navigation is the process to intelligently drive to a destination on the map using both a global path and a local path. The global path is the shortest route from the current location to the destination. On the other hand, the local path needs two steps to determine the route. The first step is to create a waypoint on the global path and several candidate routes from the current location to the waypoint. The second step is to select the optimal way from the candidate routes as the local path, considering collision risk with obstacles, speed, and smoothness. The obstacles are detected by Light Detection and Ranging (LiDAR) mounted on the AGV shown in Fig. 2(b). The LiDARs recognize the obstacle as a human only when the obstacle has two leg-like features. In the case of the human detection, it calculates collision risk for each route based on the location and the walking speed of the human. The AGV driving speed changes dynamically based on factors including a passage width and the distance to obstacles and humans. The positioning error of the AGV is about \u00a730mm after the navigation, so its accuracy enhancement is essential for workpiece exchange automation" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_35_0001419_978-94-007-7401-8-Figure3.22-1.png", + "original_path": "designv11-35/openalex_figure/designv11_35_0001419_978-94-007-7401-8-Figure3.22-1.png", + "caption": "Fig. 3.22 Overactuated PMs with coupled Sch\u00f6nflies motions of types 4PCP (a), 4CPP (b) and 4PPC (c) defined by MF \u00bc SF \u00bc 4; RF\u00f0 \u00de \u00bc v1; v2; v3;xd\u00f0 \u00de, TF = 0, NF = 6, limb topology P\\C\\P (a), C\\P\\\\P (b) and P\\P\\\\C (c)", + "texts": [ + "1 Topologies with Simple Limbs 243 Table 3.2 Limb topology and connecting conditions of the overactuated solutions with no idle mobilities presented in Figs. 3.21, 3.22, 3.23, 3.24, 3.25, 3.26, 3.27, 3.28, 3.29, 3.30, 3.31, 3.32, 3.33, 3.34, 3.35, 3.36, 3.37, 3.38, 3.39, 3.40, 3.41, 3.42, 3.43, 3.44, 3.45, 3.46, 3.47, 3.48, 3.49 No. PM type Limb topology Connecting conditions 1. 4RRC (Fig. 3.21a) R||R||C (Fig. 3.2a) The cylindrical joints of the four limbs have parallel axes 2. 4RCR (Fig. 3.21b) R||C||R (Fig. 3.2b) Idem No. 1 3. 4PCP (Fig. 3.22a) P\\C\\P (Fig. 3.2c) Idem No. 1 4. 4CPP (Fig. 3.22b) C\\P\\\\P (Fig. 3.2d) Idem No. 1 5. 4PPC (Fig. 3.22c) P\\P\\\\C (Fig. 3.2e) Idem No. 1 6. 4PCR (Fig. 3.23a) P\\C||R (Fig. 3.2f) Idem No. 1 7. 4RCP (Fig. 3.23b) R||C\\P (Fig. 3.2g) Idem No. 1 8. 4RPC (Fig. 3.24a) R\\P\\kC (Fig. 3.2h) Idem No. 1 9. 4PRC (Fig. 3.24b) P\\R||C (Fig. 3.2i) Idem No. 1 10. 4CRP (Fig. 3.25a) C||R\\P (Fig. 3.2j) Idem No. 1 11. 4CPR (Fig. 3.25b) C\\P\\kR (Fig. 3.2k) Idem No. 1 12. 4RPC (Fig. 3.26a) R\\P\\kC (Fig. 3.2i) Idem No. 1 13. 4CRR (Fig. 3.26b) C||R||R (Fig. 3.2m) Idem No. 1 14. 4RRRRR (Fig. 3.27a) R\\R||R\\R||R (Fig. 3.3a) The first and the last revolute joints of the four limbs have parallel axes 15" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_35_0001419_978-94-007-7401-8-Figure6.45-1.png", + "original_path": "designv11-35/openalex_figure/designv11_35_0001419_978-94-007-7401-8-Figure6.45-1.png", + "caption": "Fig. 6.45 2PPaRRR-1PPaRR-1CPaPa-type maximaly regular fully-parallel PM with Sch\u00f6nflies motions defined by MF = SF = 4, (RF) = (v1; v2; v3;xb), TF = 0, NF = 19, limb topology P||Pa||R||R\\R, P||Pa||R||R and C||Pa||Pa", + "texts": [ + " PM type Limb topology Connecting conditions 3. 3PPPaR-1CPaPat (Fig. 6.43a) P\\P\\||Pa\\||R (Fig. 4.8b) P\\P\\||Pa||R, (Fig. 4.8c) C||Pa||Pat (Fig. 5.4p) Idem no. 1 4. 3PPPaR-1CPaPat (Fig. 6.43b) P\\P\\||Pa\\R (Fig. 4.8a) P\\P\\||Pa||R, (Fig. 4.8c) C||Pa||Pat (Fig. 5.4p) Idem no. 1 5. 3PPaPR-1CPaPa (Fig. 6.44a) P||Pa\\P\\\\R (Fig. 4.8d) P||Pa\\P\\||R (Fig. 4.8e) C||Pa||Pa (Fig. 5.4o) Idem no. 1 6. 3PPaPR-1CPaPat (Fig. 6.44b) P||Pa\\P\\\\R (Fig. 4.8d) P||Pa\\P\\||R (Fig. 4.8e) C||Pa||Pat (Fig. 5.4p) Idem no. 1 7. 2PPaRRR-1PPaRR1CPaPa (Fig. 6.45) P||Pa||R||R\\R (Fig. 4.9a) P||Pa||R||R (Fig. 4.8g) C||Pa||Pa (Fig. 5.4o) Idem no. 1 8. 2PPaRRR-1PPaRR1CPaPat (Fig. 6.46) P||Pa||R||R\\R (Fig. 4.9a) P||Pa||R||R (Fig. 4.8g) C||Pa||Pat (Fig. 5.4p) Idem no. 1 Table 6.15 Structural parametersa of parallel mechanisms in Figs. 6.27, 6.28, 6.29, 6.30, 6.31, 6.32, 6.33, 6.34, 6.35, 6.36, 6.37 No. Structural parameter Solution Figures 6.27, 6.28, 6.29, 6.32, 6.33, 6.34, 6.35 Figures 6.30, 6.31, 6.36, 6.37 1. m 24 26 2. pi (i = 1, 3) 7 8 3. p2 7 7 4. p4 10 10 5" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_35_0000174_j.matpr.2021.01.640-Figure5-1.png", + "original_path": "designv11-35/openalex_figure/designv11_35_0000174_j.matpr.2021.01.640-Figure5-1.png", + "caption": "Fig. 5. Drive Shaft with Geometric View.", + "texts": [ + " In this type of analysis to determine the time-varying displacements, strains, stresses, and forces in a structure as it responds to any combination of static, transient, and harmonic loads. The time scale of the loading is such that the inertia or damping effects are considered to be important. The model is generated by Pro-E Software package is transferred to ANSYS software and its shown in Fig. 4 for dynamic analysis. Drive shaft have more benefit than two part drive shaft is made by Isotropic materials. It has higher explicit quality, more life, less mass, high basic speed and higher force conveying limit. Drive shaft with Geometric view of shown in Fig. 5. The pro-E model is imported to 3 dimensional view. Isometric view permits somewhat precise view of a three-dimensional object on a twodimensional or computer screen, and are a great way to visualize the shape of an object. Generated model is imported into ANSYS, is Drive Shaft. the element type, real constants and the material properties are fed into analyzing area. Drive Shaft With Mesh perspective is appeared in Fig. 6. The drive shaft is meshed number of elements. Mesh influences the accuracy, convergence, speed of the solution, Here used triangular type of mesh is always quick and easy to create" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_35_0001419_978-94-007-7401-8-Figure3.34-1.png", + "original_path": "designv11-35/openalex_figure/designv11_35_0001419_978-94-007-7401-8-Figure3.34-1.png", + "caption": "Fig. 3.34 Overactuated PMs with coupled Sch\u00f6nflies motions of types 4RRRRP (a) and 4PRRRR (b) defined by MF \u00bc SF \u00bc 4; RF\u00f0 \u00de \u00bc v1; v2; v3;xd\u00f0 \u00de, TF = 0, NF = 2, limb topology R\\R||R\\R\\P (a) and P\\R||R\\R||R (b)", + "texts": [ + "3g) Idem No. 17 21. 4RRPRR (Fig. 3.30b) R||R||P\\R||R (Fig. 3.3h) Idem No. 17 22. 4RRRPR (Fig. 3.31a) R||R\\R\\P\\kR (Fig. 3.3i) Idem No. 17 23. 4RRRRP (Fig. 3.31b) R||R\\R||R\\kP (Fig. 3.3j) Idem No. 17 24. 4PRRRR (Fig. 3.32a) P\\R\\R||R\\R (Fig. 3.3k) The second and the last joints of the four limbs have parallel axes 25. 4RPRRR (Fig. 3.32b) R\\P\\R||R\\R (Fig. 3.3l) Idem No. 14 26. 4RRPRR (Fig. 3.33a) R\\R\\P\\kR\\R (Fig. 3.3m) Idem No. 14 27. 4RRRPR (Fig. 3.33b) R\\R||R\\P\\kR (Fig. 3.3n) Idem No. 14 28. 4RRRRP (Fig. 3.34a) R\\R||R\\R\\P (Fig. 3.3o) The first and the fourth revolute joints of the four limbs have parallel axes 29. 4PRRRR (Fig. 3.34b) P\\R||R\\R||R (Fig. 3.3p) The second joints of the four limbs have parallel axes (continued) 244 3 Overactuated Topologies with Coupled Sch\u00f6nflies Motions Table 3.2 (continued) No. PM type Limb topology Connecting conditions 30. 4RPRRR (Fig. 3.35a) R\\P\\kR\\R||R (Fig. 3.3q) Idem No. 16 31. 4RRRPR (Fig. 3.35b) R||R\\R||P||R (Fig. 3.3r) Idem No. 16 32. 4RRRPR (Fig. 3.36a) R||R\\R||P||R (Fig. 3.3s) Idem No. 16 33. 4RRRRP (Fig. 3.36b) R||R\\R||R||P (Fig. 3.3t) Idem No. 16 34. 4RRRRP (Fig. 3.37a) R||R\\R||R\\P (Fig" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_35_0001419_978-94-007-7401-8-Figure2.104-1.png", + "original_path": "designv11-35/openalex_figure/designv11_35_0001419_978-94-007-7401-8-Figure2.104-1.png", + "caption": "Fig. 2.104 4RRRRPa-type fully-parallel PM with coupled Sch\u00f6nflies motions defined by MF \u00bc SF \u00bc 4; RF\u00f0 \u00de \u00bc \u00f0v1; v2; v3;xd\u00de, TF = 0, NF = 14, limb topology R||R||R\\R||Pa", + "texts": [], + "surrounding_texts": [ + "158 2 Fully-Parallel Topologies with Coupled Sch\u00f6nflies Motions", + "2.2 Topologies with Complex Limbs 159", + "160 2 Fully-Parallel Topologies with Coupled Sch\u00f6nflies Motions" + ] + }, + { + "image_filename": "designv11_35_0001419_978-94-007-7401-8-Figure5.1-1.png", + "original_path": "designv11-35/openalex_figure/designv11_35_0001419_978-94-007-7401-8-Figure5.1-1.png", + "caption": "Fig. 5.1 Simple limbs for fully-parallel PMs with uncoupled Sch\u00f6nflies motions defined by rG = 0, RG\u00f0 \u00de \u00bc v1; v2; v3;xa\u00f0 \u00de, MG = SG = 4 (a) and rG = 0, RG\u00f0 \u00de \u00bc v1; v2; v3;xa;xb;xd , MG = SG = 6 (b)", + "texts": [ + " The limbs G1, G2 and G3 are used for positioning the moving platform and limb G4 for orienting it. There are no idle mobilities in these basic solutions. G. Gogu, Structural Synthesis of Parallel Robots, Solid Mechanics and Its Applications 206, DOI: 10.1007/978-94-007-7401-8_5, Springer Science+Business Media Dordrecht 2014 431 The various types of simple and complex limbs used in the fully-parallel basic solutions illustrated in this section are presented in Figs. 5.1, 5.2, 5.3, 5.4, 5.5, 5.6. The simple limbs combine only revolute and prismatic joints (Fig. 5.1). A cylindrical joint is also used in the complex limbs to replace two consecutive revolute and prismatic joints. The simple limb in Fig. 5.1b combines a homokinetic double universal joint with telescopic intermediary shaft. One (Figs. 5.2 and 5.3), two (Figs. 5.4 and 5.5) or three (Fig. 5.6) planar parallelogram loops are combined in the complex limbs. A planar telescopic parallelogram loop is combined in the limbs in Fig. 5.4l, n and p. Three joint parameters loose their independence in each parallelogram loop. Various topologies of PMs with uncoupled Sch\u00f6nflies motions of the moving platform and no idle mobilities can be obtained by using different topologies presented in Figs", + "1 Fully-Parallel Topologies 521 522 5 Topologies with Uncoupled Sch\u00f6nflies Motions 5.1 Fully-Parallel Topologies 523 524 5 Topologies with Uncoupled Sch\u00f6nflies Motions 5.1 Fully-Parallel Topologies 525 526 5 Topologies with Uncoupled Sch\u00f6nflies Motions 7. 3PaPPR-1RUPU (Fig. 5.11a) Pa||P\\P\\kR (Fig. 5.2a) The first revolute joint of G4 limb and the last revolute joints of limbs G1, G2 and G3 have parallel axes. The last joints of limbs G1, G2 and G3 have superposed axes Pa||P\\P\\\\R (Fig. 5.2d) R\\R\\R\\P\\kR\\R (Fig. 5.1b) 8. 3PaPPR-1RUPU (Fig. 5.11b) Pa\\P\\kP\\kR (Fig. 5.2b) Idem No. 7 Pa\\P\\kP\\\\R (Fig. 5.2c) R\\R\\R\\P\\kR\\R (Fig. 5.1b) 9. 2PaPRRR1PaPRR1RUPU (Fig. 5.12a) Pa\\P\\\\R||R\\\\R (Fig. 5.3a) The first revolute joint of G4 limb and the last revolute joints of limbs G1, G2 and G3 have parallel axes. The last joints of of limbs G1, G2 and G3 have superposed axes. The revolute joints between links 5 and 6 of of limbs G1, G2 and G3 have orthogonal axes Pa\\P\\\\R||R (Fig. 5.2e) R\\R\\R\\P\\kR\\R (Fig. 5.1b) 10. 2PaPRRR1PaPRR1RUPU (Fig. 5.12b) Pa||P\\R||R\\\\R (Fig. 5.3b) Idem No. 9 Pa||P\\R||R (Fig. 5.2f) R\\R\\R\\P\\kR\\R (Fig. 5.1b) 11. 2PaRPRR1PaRPR1RUPU (Fig. 5.13) Pa\\R\\P\\kR\\R (Fig. 5.3c) The first revolute joint of G4 limb and the last revolute joints of limbs G1, G2 and G3 have parallel axes. The last joints of of limbs G1, G2 and G3 have superposed axes, and their revolute joints between links 4 and 5 have orthogonal axes Pa\\R\\P\\kR (Fig. 5.2g) R\\R\\R\\P\\kR\\R (Fig. 5.1b) 12. 2PaRRRR1PaRRR1RUPU (Fig. 5.14) Pa\\R||R||R\\kR (Fig. 5.3d) Idem No. 11 Pa\\R||R||R (Fig. 5.2h) R\\R\\R\\P\\kR\\R (Fig. 5.1b) 5.1 Fully-Parallel Topologies 527 Table 5.2 Limb topology and connecting conditions of the fully-parallel solutions with no idle mobilities presented in Figs. 5.15, 5.16, 5.17, 5.18, 5.19, 5.20, 5.21, 5.22, 5.23, 5.24, 5.25, 5.26, 5.27, 5.28, 5.29, 5.30 No. PM type Limb topology Connecting conditions 1. 3PaPPR-1RPaPaP (Fig. 5.15a) Pa||P\\P\\kR (Fig. 5.2a) The last joints of the four limbs have superposed axes/directions Pa||P\\P\\\\R (Fig. 5.2d) R||Pa||Pa||P (Fig. 5.4k) 2. 3PaPPR-1RPaPaP (Fig. 5.15b) Pa\\P\\kP\\kR (Fig", + "4l) 24. 2PaRRRR-1PaRRR1RPPaPat (Fig. 5.30) Pa\\R||R||R\\kR (Fig. 5.3d) Idem No. 20 Pa\\R||R||R (Fig. 5.2h) R||P||Pa||Pat (Fig. 5.4n) Table 5.3 Limb topology and connecting conditions of the fully-parallel solutions with no idle mobilities presented in Figs. 5.31, 5.32, 5.33, 5.34, 5.35, 5.36, 5.37, 5.38 No. PM type Limb topology Connecting conditions 1. 3PaPPaR1RPPP (Fig. 5.31a) Pa||P\\Pa\\kR (Fig. 5.4a) The last joints of the four limbs have superposed axes/directionsPa||P\\Pa||R (Fig. 5.4d) R\\P\\\\P\\\\P (Fig. 5.1a) 2. 3PaPPaR- 1RPPP (Fig. 5.31b) Pa\\P\\\\Pa\\kR (Fig. 5.4b) Idem No. 1 Pa\\P\\\\Pa||R (Fig. 5.4c) R\\P\\\\P\\\\P (Fig. 5.1a) 3. 3PaPaPR1RPPP (Fig. 5.32a) Pa\\Pa\\kP\\\\R (Fig. 5.4f) Idem No. 1 Pa\\Pa\\kP\\kR (Fig. 5.4h) R\\P\\\\P\\\\P (Fig. 5.1a) 4. 3PaPaPR1RPPP (Fig. 5.32b) Pa\\Pa\\\\P\\\\R (Fig. 5.4g) Idem No. 1 Pa\\Pa\\\\P\\kR (Fig. 5.4e) R\\P\\\\P\\\\P (Fig. 5.1a) 5. 2PaPaRRR1PaPaRR1RPPP (Fig. 5.33) Pa\\Pa||R||R\\\\R (Fig. 5.5a) The last joints of the four limbs have superposed axes/directions. The revolute joints between links 7 and 8 of limbs G1, G2 and G3 have orthogonal axes Pa\\Pa||R||R (Fig. 5.4i) R\\P\\\\P\\\\P (Fig. 5.1a) 6. 2PaPaRRR1PaPaRR1RPPP (Fig. 5.34) Pa\\Pa||R||R\\kR (Fig. 5.5b) Idem No. 5 Pa\\Pa||R||R (Fig. 5.4j) R\\P\\\\P\\\\P (Fig. 5.1a) 7. 3PaPPaR1RUPU (Fig. 5.35a) Pa||P\\Pa\\kR (Fig. 5.4a) The first revolute joint of G4 limb and the last revolute joints of limbs G1, G2 and G3 have parallel axes. The last joints of of limbs G1, G2 and G3 have superposed axes Pa||P\\Pa||R (Fig. 5.4d) R\\R\\R\\P\\kR\\R (Fig. 5.1b) (continued) 530 5 Topologies with Uncoupled Sch\u00f6nflies Motions Table 5.3 (continued) No. PM type Limb topology Connecting conditions 8. 3PaPPaR1RUPU (Fig. 5.35b) Pa\\P\\\\Pa\\kR (Fig. 5.4b) Idem No. 7 Pa\\P\\\\Pa||R (Fig. 5.4c) R\\R\\R\\P\\kR\\R (Fig. 5.1b) 9. 3PaPaPR1RUPU (Fig. 5.36a) Pa\\Pa\\kP\\\\R (Fig. 5.4f) Idem No. 7 Pa\\Pa\\kP\\kR (Fig. 5.4h) R\\R\\R\\P\\kR\\R (Fig. 5.1b) 10. 3PaPaPR1RUPU (Fig. 5.36b) Pa\\Pa\\\\P\\\\R (Fig. 5.4g) Idem No. 7 Pa\\Pa\\\\P\\kR (Fig. 5.4e) R\\R\\R\\P\\kR\\R (Fig. 5.1b) 11. 2PaPaRRR1PaPaRR1RUPU (Fig. 5.37) Pa\\Pa||R||R\\\\R (Fig. 5.5a) The last joints of of limbs G1, G2 and G3 have superposed axes. The revolute joints between links 7 and 8 of limbs G1, G2 and G3 have orthogonal axes Pa\\Pa||R||R (Fig. 5.4i) R\\R\\R\\P\\kR\\R (Fig. 5.1b) 12. 2PaPaRRR1PaPaRR1RUPU (Fig. 5.38) Pa\\Pa||R||R\\kR (Fig. 5.5b) Idem No. 11 Pa\\Pa||R||R (Fig. 5.4j) R\\R\\R\\P\\kR\\R (Fig. 5.1b) Table 5.4 Limb topology and connecting conditions of the fully-parallel solutions with no idle mobilities presented in Figs. 5.39, 5.40, 5.41, 5.42, 5.43, 5.44, 5.45, 5.46, 5.47, 5.48, 5.49, 5.50, 5.51, 5.52, 5.53, 5.54, 5.55, 5.56, 5.57, 5.58, 5.59, 5.60, 5.61, 5.62, 5.63, 5.64, 5.65, 5.66, 5.67, 5.68 No. PM type Limb topology Connecting conditions 1. 3PaPPaR-1RPaPaP (Fig. 5.39a) Pa||P\\Pa\\kR (Fig. 5.4a) The last joints of the four limbs have superposed axes/directions Pa||P\\Pa||R (Fig. 5.4d) R||Pa||Pa||P (Fig", + "4i) R||P||Pa||Pa (Fig. 5.4m) 22. 2PaPaRRR-1PaPaRR1RPPaPa (Fig. 5.56) Pa\\Pa||R||R\\kR (Fig. 5.5b) Idem No. 21 Pa\\Pa||R||R (Fig. 5.4i) R||P||Pa||Pa (Fig. 5.4m) 23. 2PaPaRRR-1PaPaRR1RPPaPat (Fig. 5.57) Pa\\Pa||R||R\\\\R (Fig. 5.5a) Idem No. 21 Pa\\Pa||R||R (Fig. 5.4i) R||P||Pa||Pat (Fig. 5.4n) 24. 2PaPaRRR-1PaPaRR1RPPaPat (Fig. 5.58) Pa\\Pa||R||R\\kR (Fig. 5.5b) Idem No. 21 Pa\\Pa||R||R (Fig. 5.4i) R||P||Pa||Pat (Fig. 5.4n) 25. 3PaPaPaR-1RPPP (Fig. 5.59) Pa\\Pa||Pa\\kR (Fig. 5.6a) Idem No. 1 Pa\\Pa||Pa||R (Fig. 5.6b) R\\P\\\\P\\\\P (Fig. 5.1a) 26. 3PaPaPaR-1RPPP (Fig. 5.60) Pa\\Pa||Pa\\\\R (Fig. 5.6c) Idem No. 1 Pa\\Pa||Pa||R (Fig. 5.6b) R\\P\\\\P\\\\P (Fig. 5.1a) 27. 3PaPaPaR-1RUPU (Fig. 5.61) Pa\\Pa||Pa\\kR (Fig. 5.6a) The first revolute joint of G4 limb and the last revolute joints of limbs G1, G2 and G3 have parallel axes. The last joints of of limbs G1, G2 and G3 have superposed axes Pa\\Pa||Pa||R (Fig. 5.6b) R\\R\\R\\P\\kR\\R (Fig. 5.1b) 28. 3PaPaPaR-1RUPU (Fig. 5.62) Pa\\Pa||Pa\\\\R (Fig. 5.6c) Idem No. 27 Pa\\Pa||Pa||R (Fig. 5.6b) R\\R\\R\\P\\kR\\R (Fig. 5.1b) 29. 3PaPaPaR-1RPaPaP (Fig. 5.63) Pa\\Pa||Pa\\kR (Fig. 5.6a) Idem No. 1 Pa\\Pa||Pa||R (Fig. 5.6b) R||Pa||Pa||P (Fig. 5.4k) (continued) 5.1 Fully-Parallel Topologies 533 Table 5.4 (continued) No. PM type Limb topology Connecting conditions 30. 3PaPaPaR-1RPaPaP (Fig. 5.64) Pa\\Pa||Pa\\\\R (Fig. 5.6c) Idem No. 1 Pa\\Pa||Pa||R (Fig. 5.6b) R||Pa||Pa||P (Fig. 5.54k) 31. 3PaPaPaR1RPaPatP (Fig. 5.65) Pa\\Pa||Pa\\kR (Fig. 5.6a) Idem No. 1 Pa\\Pa||Pa||R (Fig. 5.6b) R||Pa||Pat||P (Fig. 5.4l) 32. 3PaPaPaR1RPaPatP (Fig", + "R (Fig. 4.2e) C ?P ?||R (Fig. 4.1i) C||Pa||Pa (Fig. 5.4o) Idem no. 1 6. 2CPRR-1CPR-1CPaPat (Fig. 6.19b) C ?P ?||R ?R (Fig. 4.2e) C ?P ?||R (Fig. 4.1i) C||Pa||Pat (Fig. 5.4p) Idem no. 1 7. 2CRRR-1CRR-1CPaPa (Fig. 6.20a) C||R||R ?R (Fig. 4.2g) C||R||R (Fig. 4.1k) C||Pa||Pa (Fig. 5.4o) Idem no. 1 8. 2CRRR-1CRR-1CPaPat (Fig. 6.20b) C||R||R ?R (Fig. 4.2g) C||R||R (Fig. 4.1k) C||Pa||Pat (Fig. 5.4p) Idem no. 1 9. 3PPPaR-1RPPP (Fig. 6.21a) P ?P ?||Pa ?||R (Fig. 4.8b) P ?P ?||Pa||R, (Fig. 4.8c) R ?P\\\\P\\\\P (Fig. 5.1a) The last joints of the four limbs have superposed axes/ directions. The actuated prismatic joints of limbs G1, G2 and G3 have orthogonal directions 10. 3PPPaR-1RPPP (Fig. 6.21b) P ?P ?||Pa\\\\R (Fig. 4.8a) P ?P ?||Pa||R, (Fig. 4.8c) R\\\\P\\\\P\\\\P (Fig. 5.1a) Idem no. 9 11. 3PPaPR-1RPPP (Fig. 6.22a) P||Pa ?P\\\\R (Fig. 4.8d) P||Pa ?P ?||R (Fig. 4.8e) R ?P\\\\P\\\\P (Fig. 5.1a) Idem no. 9 12. 2PPaRRR-1PPaRR-1RPPP (Fig. 6.22b) P||Pa||R||R ?R (Fig. 4.9a) P||Pa||R||R (Fig. 4.8g) R ?P\\\\P\\\\P (Fig. 5.1a) Idem no. 9 13. 3PPPaR-1RUPU (Fig. 6.23a) P ?P ?||Pa ?||R (Fig. 4.8b) P ?P ?||Pa||R, (Fig. 4.8c) R ?R ?R ?P ?||R ?R (Fig. 5.1b) The first revolute joint of G4-limb and the last last joints of limbs G1, G2 and G3 have parallel axes. The last revolute joints of limbs G1, G2 and G3 have superposed axes. The actuated prismatic joints of limbs G1, G2 and G3 have orthogonal directions (continued) 614 6 Maximally Regular Topologies with Sch\u00f6nflies Motions Table 6.6 (continued) No. PM type Limb topology Connecting conditions 14. 3PPPaR-1RUPU (Fig. 6.23b) P ?P ?||Pa\\\\R (Fig. 4.8a) P ?P ?||Pa||R, (Fig. 4.8c) R ?R ?R ?P ?||R ?R (Fig. 5.1b) Idem no. 13 15. 3PPaPR-1RUPU (Fig. 6.24a) P||Pa ?P\\\\R (Fig. 4.8d) P||Pa ?P ?||R (Fig. 4.8e) R ?R ?R ?P ?||R ?R (Fig. 5.1b) Idem no. 13 16. 2PPaRRR-1PPaRR-1RUPU (Fig. 6.24b) P||Pa||R||R ?R (Fig. 4.9a) P||Pa||R||R (Fig. 4.8g) R ?R ?R ?P ?||R ?R (Fig. 5.1b) Idem no. 13 17. 3PPaPaR-1RPPP (Fig. 6.25) P||Pa||Pa ?R (Fig. 4.10a) P||Pa||Pa||R (Fig. 4.10b) R ?P\\\\P\\\\P (Fig. 5.1a) Idem no. 9 18. 3PPaPaR-1RUPU (Fig. 6.26) P||Pa||Pa ?R (Fig. 4.10a) P||Pa||Pa||R (Fig. 4.10b) R ?R ?R ?P ?||R ?R (Fig. 5.1b) Idem no. 13 Table 6.7 Structural parametersa of parallel mechanisms in Figs. 6.7, 6.8, 6.9, 6.10, 6.11, 6.12, 6.13, 6.14 No. Structural parameter Solution Figures 6.7 and 6.11 Figures 6.8, 6.9, 6.10, 6.12, 6.13, 6.14 1. m 18 20 2. pi (i = 1, 3) 4 5 3. p2 4 4 4. p4 10 10 5. p 22 24 6. q 5 5 7. k1 3 3 8. k2 1 1 9. k 4 4 10. (RG1) (v1; v2; v3;xb) (v1; v2; v3;xa;xb) 11. (RG2) (v1; v2; v3;xb) (v1; v2; v3;xb) 12. (RG3) (v1; v2; v3;xb) (v1; v2; v3;xb;xd) 13. (RG4) (v1; v2; v3;xb) (v1; v2; v3;xb) 14. SGi (i = 1, 3) 4 5 15" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_35_0000535_tmag.2021.3085750-Figure8-1.png", + "original_path": "designv11-35/openalex_figure/designv11_35_0000535_tmag.2021.3085750-Figure8-1.png", + "caption": "Fig. 8. (a) The prototype HMC-VFMM. (b) Experimental setup.", + "texts": [ + " That is to say, the major and minor hysteresis loops of the hysteresis curve share approximately two same sided branches, and the recoil lines can be regarded as approximately linear lines. From the FE results, it can be deduced that the HMCVFMM design can effectively avoid the mismatch of working point movement between one-way and bidirectional MS manipulation. This magnetization characteristic of the HMCVFMM confirms that the alternate MS manipulation process has weak impact on the practical MS current pulses to be applied, and will further verify the validity of the balanced bidirectional-magnetization effect. As shown in Fig. 8(a), a HMC-VFMM prototype is manufactured and tested. Fig. 8(b) shows the experimental setup. The magnetization test is established based on the MS control circuit described in Fig. 1. The FE-predicted and measured line back-EMF waveforms and harmonic spectra under different MSs are shown in Fig. 9. It can be seen that the HMC-VFMM exhibits a good flux regulation capability. Besides, the FE-predicted and tested torque versus speed curves are compared as shown in Fig. 10. It can be deduced that the speed range of the HMC-VFMM can be effectively extended by online MS manipulation" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_35_0000686_j.mechmachtheory.2021.104433-Figure8-1.png", + "original_path": "designv11-35/openalex_figure/designv11_35_0000686_j.mechmachtheory.2021.104433-Figure8-1.png", + "caption": "Fig. 8. The equivalent mechanism of the epicyclic gear-rack train.", + "texts": [ + " The motion of the epicyclic gear-rack train can be decomposed into two relatively independent motions similar to the epicyclic external gear train: the relative mesh transmission of the sun rack-gear and the planet rack, and the rotation of the rack arm and the planet rack around the axis of the sun rack-gear. Then a close loop 5-bar planar equivalent linkage with two prismatic joints is proposed as the equivalent mechanism of the epicyclic gear-rack train by adding a link connecting the center of the sun rack-gear and the pitch line of the planet rack. Fig. 8 shows that link lrAB, lrCD, and lrOC are equivalent to the sun rack-gear, the planet rack, and the rack arm, respectively. Joint B and joint C are the centers of curvature of the tooth profile curve at the mesh point. The higher pair at the mesh point is replaced by link lrBC, which is perpendicular to link lrAB. Instantaneously, the equivalent mechanism has the same mobility, instantaneous velocity, and instantaneous acceleration as the epicyclic gear-rack train. In addition, Eq. (19) reveals the equivalent geometrical relationships", + " 10 Jb\u03c9b = [ \u2212 SbA SbB SbC1 SbC2 SbC3 SbD1 SbD2 SbE ] \u23a1 \u23a2 \u23a2 \u23a2 \u23a2 \u23a2 \u23a2 \u23a2 \u23a2 \u23a2 \u23a2 \u23a3 \u03c9bA \u03c9bB \u03c9bC1 \u03c9bC2 \u03c9bC3 \u03c9bD1 \u03c9bD2 \u03c9bE \u23a4 \u23a5 \u23a5 \u23a5 \u23a5 \u23a5 \u23a5 \u23a5 \u23a5 \u23a5 \u23a5 \u23a6 = 0 (23) The topological loop Lr of the equivalent mechanism of the epicyclic gear-rack train is presented by the directed graph in Fig.11, where the nodes represent the links lrAB, lrBC, lrCD, lrOD, GrOA and the lines with an arrow represent the screws SrO, SrA, SrB, SrC, SrD added H. Yuan et al. Mechanism and Machine Theory 166 (2021) 104433 on the joints O, A, B, C, D. The link GrOA is the ground link in the equivalent mechanism. As shown in Fig. 8, the screw system of the joints in the epicyclic gear-rack train is given by Eq. (24). \u23a7 \u23aa \u23aa \u23a8 \u23aa \u23aa \u23a9 SrO = [ 0 0 1 0 0 0 ]T SrA = [ 0 0 1 0 0 0 ]T SrB = [ 0 0 1 lrABsin(\u03c6r2 \u2212 \u03b1r) \u2212 lrABcos(\u03c6r2 \u2212 \u03b1r) 0 ]T SrC = [ 0 0 0 \u2212 sin\u03c6r2 cos\u03c6r2 0 ]T SrD = [ 0 0 0 cos(\u03c6r2 \u2212 \u03b1r) sin(\u03c6r2 \u2212 \u03b1r) 0 ]T (24) The closed-loop velocity equation can be obtained as Eq. (25) in matrix form Jr\u03c9r = [SrA SrB SrC SrD \u2212 SrO ] \u23a1 \u23a2 \u23a2 \u23a2 \u23a2 \u23a3 \u03c9rA \u03c9rB vrC vrD \u03c9rO \u23a4 \u23a5 \u23a5 \u23a5 \u23a5 \u23a6 = 0 (25) By taking derivatives of Eq. (21), the closed-loop acceleration equation is obtained as Eq" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_35_0000395_978-3-030-67750-3_16-Figure26.2-1.png", + "original_path": "designv11-35/openalex_figure/designv11_35_0000395_978-3-030-67750-3_16-Figure26.2-1.png", + "caption": "Fig. 26.2 a Design 1. b Design 2", + "texts": [ + " These two designs are using the systems that use a barrel-like principle (Abdullah and Jamil 2015), as in Fig. 26.1a. Basic components in the designs can be termed as in Fig. 26.1b (Abdullah and Jamil 2015). The main advantage of this design is the possibility to separate the housing from the top and bottom cover that makes this concept easy to set the parameter of the device. By disassembling the part of the system, the coil and magnet can be modified, which made the parameter setting simpler. Both conceptual designs can refer to as in Fig. 26.2a,b, where Fig. 26.2a shows Design 1, while Fig. 26.2b shows Design 2. For design 1, a set of two coils is placed outside the magnet. All the linkages are connected to each top and bottom mount. The main shaft is from the bottom linkage connected to themagnet, and themagnet\u2019smotion is controlled by the spring,which is Parameters Design 1 Design 2 Magnet height, mm 200 200 Magnet radius, mm 100 100 Wound coil height, mm 220 220 Number of coil turn 20 10 Coil internal diameter, mm 140 140 Distance between magnet and coil, mm 10 10 Wire diameter, mm 2 2 attachedbelow themagnet" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_35_0000412_asjc.2565-Figure1-1.png", + "original_path": "designv11-35/openalex_figure/designv11_35_0000412_asjc.2565-Figure1-1.png", + "caption": "FIGURE 1 Multirotor vehicle with variable mass used for monitoring and pest combat. (A) Concept of the multirotor vehicle with variable mass. (B) Model of the multirotor vehicle used to conduct Model In the Loop simulations [Color figure can be viewed at wileyonlinelibrary.com]", + "texts": [ + " The releasing task is conducted by using a Smart Releasing Device (SRD) which, in combination with the biological agents, represent two fifths of the Maximum Takeoff Mass (MTOM) of the UAS. The releasing of the payload (flies) represents a big variation in the overall vehicle mass, which can trigger a malfunctioning of the nominal controller's gains. The technical problem addressed in this paper is the development of analytic strategies for mass estimation, attitude stabilization, and trajectory tracking control of the multirotor UAS depicted in Figure 1A. We propose a soft sensor approach to enable online identification of the vehicle mass during the releasing of the sterile insects in crops, which allows online adaptation of the proposed controller's gains. Figure 1B depicts the developed model of the multirotor UAS to conduct Model In the Loop (MIL) simulations to validate the proposed strategies. The proposed soft sensor is composed by (i) the combination of a Terminal SMO with a LSM to perform the identification of the UAS mass and inertial moments, and (ii) an adaptive controller and an attractive ellipsoid method to stabilize the altitude dynamics in combination with classical PD controllers for the remaining vehicle's subsystem. The SMO provides the estimation of the linear accelerations and linear velocities of the UAS in a finite time T [21]", + " The stabilization of the Longitudinal and Lateral dynamics is performed based on PD Controllers in combination with the control signal obtained from the altitude controller, see Figure 3 in Section 4. The PD controllers use the initial known UAS mass m0 before the convergence time T, and the estimated mass m\u0302 once T has been reached. The directional (yaw) dynamics are controlled independently of the changes in the vehicle mass, since the vehicle mass is not directly related to its dynamics, by a PD controller. A simplified dynamic model of the multi-rotor UAS shown in Figure 1 can be written as follows[22]: x\u0308 = 1 m (c\ud835\udf19s\ud835\udf03c\ud835\udf13 + s\ud835\udf19s\ud835\udf13 )u1, ?\u0308? = 1 Ix \ud835\udf0f\ud835\udf19 ?\u0308? = 1 m (c\ud835\udf19s\ud835\udf03s\ud835\udf13 \u2212 s\ud835\udf19c\ud835\udf13 )u1, ?\u0308? = 1 I\ud835\udc66 \ud835\udf0f\ud835\udf03 z\u0308 = 1 m (c\ud835\udf19c\ud835\udf03)u1 \u2212 g, ?\u0308? = 1 Iz \ud835\udf0f\ud835\udf13 (1) where the terms c\u2217 and s\u2217 stand for cos(\u2217) and sin(\u2217), respectively. The inertial frame is given by the coordinate system {} = {x , \ud835\udc66 , z}, while {} = {x, \ud835\udc66, z} represents the multirotor's body-fixed frame, both of them in East-North-Up (ENU) coordinates. The position vector of the vehicle's center of mass with respect to {} is denoted by {x, y, z}" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_35_0000686_j.mechmachtheory.2021.104433-Figure24-1.png", + "original_path": "designv11-35/openalex_figure/designv11_35_0000686_j.mechmachtheory.2021.104433-Figure24-1.png", + "caption": "Fig. 24. Clamping motion modes of the clamping mechanism.", + "texts": [ + " With two mobility states, when the metamorphosis is initiated, either geometrical constraint or force constraint is utilized to restrict one mobility. An example describing the way of controlling the transformation is demonstrated in the following section. A metamorphic epicyclic gear-rack clamping mechanism is introduced as an example for describing the design method using equivalent mechanisms of metamorphic epicyclic gear trains. The clamping process of a 2 mobility clamping mechanism is shown in Fig. 24. At the beginning of the auction period, the claw of H. Yuan et al. Mechanism and Machine Theory 166 (2021) 104433 the clamping mechanism rotates towards the target object until contacts it. Then, the motion of the claw transforms from rotation to translation for clamping the target object with claw and support plate. When clamping is over, the motion process is opposite of the clamping process described above and the claw returns to its initial position. According to the clamping process and demands of the task, the clamping mechanism has two motion modes: H", + " Mechanism and Machine Theory 166 (2021) 104433 H. Yuan et al. Mechanism and Machine Theory 166 (2021) 104433 (1) Rotation motion output to provide the motion of claw to closer or away from the target object. (2) Translation motion output to provide the motion of claw to clamp or release target object. The motion modes of the clamping mechanism include rotation and translation. Therefore, we choose the equivalent mechanism of the epicyclic bevel gear mechanism to replace the rotation link of the clamping mechanism in Fig. 24. An equivalent metamorphic clamping mechanism is constructed in Fig. 25. The motion-screw system Tm of the equivalent metamorphic clamping mechanism is obtained as Eq. (52). H. Yuan et al. Mechanism and Machine Theory 166 (2021) 104433 Tm = \u23a7 \u23aa \u23aa \u23aa \u23aa \u23aa \u23aa \u23aa \u23aa \u23aa \u23aa \u23aa \u23aa \u23aa \u23aa \u23aa \u23aa \u23aa \u23aa \u23aa \u23aa \u23aa \u23aa \u23aa \u23a8 \u23aa \u23aa \u23aa \u23aa \u23aa \u23aa \u23aa \u23aa \u23aa \u23aa \u23aa \u23aa \u23aa \u23aa \u23aa \u23aa \u23aa \u23aa \u23aa \u23aa \u23aa \u23aa \u23aa \u23a9 TA = [ 0 0 \u03c9A 0 0 0 ]T TB = [ 0 0 \u03c9B 0 0 0 ]T TC1 = [ \u2212 sin\u03b1b3\u03c9C1 \u2212 cos\u03b1b\u03c9C1 \u2212 sin\u03b1b1\u03c9C1 0 Rbsin\u03b1b1\u03c9C1 \u2212 Rbcos\u03b1b\u03c9C1 ] T TC2 = [ cos\u03b1b3\u03c9C2 \u2212 cos\u03b1btan\u03b1b3\u03c9C2 \u2212 sin\u03b1b1tan\u03b1b3\u03c9C2 0 0 0 ]T TC3 = [ 0 \u2212 sin\u03b1b1\u03c9C3 cos\u03b1b\u03c9C3 q11\u03c9C3 q12\u03c9C3 q13\u03c9C3 ] T cos\u03b1b3 TD1 = [ \u2212 sin\u03b1b1sin\u03b1b3\u03c9D1 \u2212 cos\u03b1bsin\u03b1b1\u03c9D1 cos2\u03b1b1\u03c9D1 q14\u03c9D1 q15\u03c9D1 q16\u03c9D1 ]T cos\u03b1b1 TD2 = [ cos\u03b1b\u03c9D2 \u2212 sin\u03b1b3\u03c9D2 0 q17\u03c9D2 q18\u03c9D2 q19\u03c9D2 ] T cos\u03b1b1 TE = [\u03c9E 0 0 0 rb\u03c9E 0 ]T TF1 = [ 0 0 0 vF 0 0 ]T TF2 = [\u03c9F2 0 0 0 rb\u03c9F2 0 ]T TG = [ 0 0 0 vG 0 0 ] (52) where the q11 to q19 are given in Eq" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_35_0001478_1.4035286-Figure2-1.png", + "original_path": "designv11-35/openalex_figure/designv11_35_0001478_1.4035286-Figure2-1.png", + "caption": "Fig. 2 Single-DOF model of a cyclic blisk segment for first flapwise bending mode", + "texts": [ + " As soon as all contributions to the residual are known, the solution of the nonlinear algebraic system of equations can be found by means of any root-finding algorithm, such as Newton\u2013Raphson method. An overlying continuation procedure manages then to find neighboring solutions. 3.1 Single-DOF Model of a Blisk 3.1.1 Testcase Setup. In order to give basic insights into the behavior of rotational speed-dependent systems, a single-DOF model, capable of approximating the first flapwise bending mode of a blisk, is utilized, see Fig. 2. The assumptions of small angles sin(d) d, quasistationary rotational speed jdX=dtj 1, planar problem description, and a stiff blisk (nodal diameter of zero) lead to the scalar equation of motion m \u20182\u20acd \u00fe c _d \u00fe k \u00fe m \u20180\u2018X 2 d \u00bc fe (19) with m \u00bc 0:5kg; c\u00bc 2Nms=rad; k\u00bc 5000Nm=rad; \u20180 \u00bc 0:3m, and \u2018\u00bc0.1m as an example. Since the equation is derived analytically, the auxiliary variables k0 \u00bc k; k1 \u00bcm \u20180\u2018 and k2\u00bc0 for the interpolation are readily legible. Except centrifugal forces, all other nonlinearities are neglected" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_35_0001419_978-94-007-7401-8-Figure2.28-1.png", + "original_path": "designv11-35/openalex_figure/designv11_35_0001419_978-94-007-7401-8-Figure2.28-1.png", + "caption": "Fig. 2.28 Fully-parallel PMs with coupled Sch\u00f6nflies motions of types 4PRPPa (a) and 4PPRPa (b) defined by MF \u00bc SF \u00bc 4; RF\u00f0 \u00de \u00bc \u00f0v1; v2; v3;xd\u00de, TF = 0, NF = 18, limb topology P||R\\P||Pa (a) and P\\P||R\\Pa (b)", + "texts": [ + " PM type Limb topology Connecting conditions 1. 4RRPaR (Fig. 2.26a) R||R\\Pa\\kR (Fig. 2.21a) The first two and the last revolute joints of the four limbs have parallel axes 2. 4RRRPa (Fig. 2.26b) R||R||R\\Pa (Fig. 2.21b) The three first revolute joints of the four limbs have parallel axes 3. 4PRPaR (Fig. 2.27a) P\\R\\Pa\\kR (Fig. 2.21c) The second and the last joints of the four limbs have parallel axes 4. 4PPPaR (Fig. 2.27b) P\\P\\kPa\\R (Fig. 2.21d) The last revolute joints of the four limbs have parallel axes 5. 4PRPPa (Fig. 2.28a) P||R\\P||Pa (Fig. 2.21e) The second joints of the four limbs have parallel axes 6. 4PPRPa (Fig. 2.28b) P\\P||R\\Pa (Fig. 2.21f) The third joints of the four limbs have parallel axes 7. 4PRPPa (Fig. 2.29a) P\\R||P\\Pa (Fig. 2.21g) Idem No. 5 8. 4PRPaP (Fig. 2.29b) P\\R\\Pa \\kP (Fig. 2.21h) Idem No. 5 9. 4PPPaR (Fig. 2.30a) P\\P\\kPa\\kR (Fig. 2.21i) Idem No. 4 10. 4PPPaR (Fig. 2.30b) P\\P||Pa\\R (Fig. 2.21j) Idem No. 4 11. 4PPaRP (Fig. 2.31a) P||Pa\\R||P (Fig. 2.21k) The before last joints of the four limbs have parallel axes 12. 4PPaPR (Fig. 2.31b) P||Pa\\P|| R (Fig. 2.21l) Idem No. 4 13. 4PPaPR (Fig. 2.32a) P\\Pa||P\\\\R (Fig" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_35_0000808_s11665-021-06133-0-Figure2-1.png", + "original_path": "designv11-35/openalex_figure/designv11_35_0000808_s11665-021-06133-0-Figure2-1.png", + "caption": "Fig. 2 Schematic diagram of welding groove (mm)", + "texts": [ + " The microstructure, dislocation, Kernel average misorientation, texture and microhardness of joint were investigated, the microstructure characteristics and microhardness distribution of titanium alloy joints by narrow gap fiber laser welding were studied, and the tensile strength and impact toughness of joints in upper and middle were tested, which laid a good foundation for the subsequent narrow gap laser welding of titanium alloy with large thickness. The material used in the test is the annealed TA2 plate with the size of 200 mm 9100 mm 9 40 mm. The microstructure of the base metal is shown in Fig. 1. The microstructure of the base metal is equiaxed a phase. Mechanical properties of TA2pure titanium at room temperature are given in Table 1. Schematic diagram of welding groove geometry is shown in Fig. 2. Before welding, the surface oxide is removed by acid pickling and then dried after cleaning with water. The surface of the dried test plate is required to be clean and free of oxidation color. Before assembly, use stainless steel wire brush to polish the welding area and then wipe clean with acetone. Ventilation protection of welding pool is a key technology in titanium alloy welding process. The quality of protection directly determines the weld formation and the comprehensive mechanical properties of the weld" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_35_0000213_j.triboint.2021.106993-Figure15-1.png", + "original_path": "designv11-35/openalex_figure/designv11_35_0000213_j.triboint.2021.106993-Figure15-1.png", + "caption": "Fig. 15. Temperature Contour by CFD simulation (10,000 N load, 9000 rpm); (a) Journal Surface, (b) Pad Surface; (1) Smooth TPJB, (2) Pocketed TPJB.", + "texts": [ + " In the pockets, the evaporation significantly lessens the dynamic viscosity (\u03bcf ) of the fluid, and the larger film thickness decreases the velocity gradient (\u2202ucir/\u2202r). Therefore, the pockets reduce the shear stress (\u03bcf \u2202ucir/\u2202r). A substantial shear stress reduction is confirmed in Fig. 14, which implies a corresponding significant drop in drag torque and power loss by Eqs. (18) and (19), respectively. Fig. 14(a) shows a significant level of shear stress on the upper pads of the conventional TPJB despite the low pad loading. So, reducing the power loss from the upper pads can yield a significant overall power savings for the TPJB. Fig. 15 shows the temperature distribution on the pad and journal at 9000 rpm and 10,000 N load, for both the smooth and pocketed bearings. Figs. 15(a-1) shows the temperature varying in the axial direction due to the three nozzle oil injection cooling flows. The pad surface temperature increases along the rotation direction, and the bottom pads have relatively higher temperatures because of the larger viscous heat generation in the thinner film. As shown in Fig. 15, the overall temperature at both pad and journal surfaces is decreased in the pocketed TPJB. Fig. 11 shows that the negative gauge pressure in the pockets yields a lower load condition on the upper pads, than for the smooth TPJB. Thus, the journal lifts to the opposite direction of the applied load, and the eccentricity ratio is decreased, with an increase in the minimum film thickness. This indicates the pockets can provide power loss reduction benefits without sacrificing bearing load capacity" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_35_0001596_isscc.1959.1157072-Figure1-1.png", + "original_path": "designv11-35/openalex_figure/designv11_35_0001596_isscc.1959.1157072-Figure1-1.png", + "caption": "Figure 1-Simplest usable array, shown in CLEARED (or ZERO) state. Pulsed f, can change state to those of Figures 2 and 3. Current and voltage arrows show positive refer-", + "texts": [], + "surrounding_texts": [ + "FRIDAY, FEBRUARY 13, 1959 University Muse~m-9:OO A.M.-1 2:OO Noon\n.SESSION VI: Switching Circuits I\n.6.3: A New All-Magnetic Logic System Using Simple Cores\nD. C. Engelbart, Stanford Research Institute, Menlo Park, Calif.\nIT WILL BE SHOWN how a few simple, single-path (e.g., toroidal) magnetic elements can be wired into an array whose properties permit construction of useful, homogenous networks cmnposed only of these all-magnetic arrays. Such networks, when driven by an appropriate sequence of clock-source currents, can transfer binary information states from array to array with stable gain characteristics, and can possess facility for logical manipulation. Two different sizes of cores are utilized, providing an important difference in respective switching currents. For simplicity in the ensuing functional description, single turns will be portrayed throughout, and we shall assume that the cores all contribute the same switchable flux-linkage, regardless of size.\nThe simplest usable array is shown in Figures 1-3 in the three magnetization states of immediate concern. Observe, in Figures 1 and 3, that when looking into an array in either the CLEARED or RESET states, exactly the same conditions are seen from either set of terminals. Further, in both the SET .and the RESET states, essentially identical situations are seen from bb. From this we see that, for current entering the bottom terminal, hb sees a high threshold in the CLEARED state, and low thresholds in the SET and RESET states.\nThe information state of one of these arrays can be transferred to a second array, which must be in the CLEARED state, by applying a current to the coupling loop (Figure 5 ) . The applied current can be divided into two equal, diverginq components, I t , which satisfies Kirchhoff's current law. A loop current, I,, can be superposed, without affecting current-law observarxe, to satisfy Rirchhofs voltage law about the loop. If A were CLEARED (ZERO state), I, ideally would equal zero, since conditions lo,oking both ways from the driven nodes would appear identical. Consequently, current I t going into each of these CLEARED-state arrays causes no switching in either (sub-threshold; see Figure 4) .\nO n the other hand, if A were RESET (ONE state), and we looked left and right from the driven nodes, we would see ,conditions associated respectively with paths ( 4 ) and (2) of\nFigure 4. In this case, I,, would tend to grow toward a value equnl to the difference between the large and the small tl;resholds (Figure 4 ) , but would settle at some lesser value which causes switching (Figure 5) of cores ZA and YB (and %E?) to proceed at equal rates as array B is forced to the SET\nstate.\nI t can be shown that neither during the above transfer-AB .operation nor during the subsequent clear-A operation does a disturbance propagate out of this two-array system which is serious enough to affect the information states of adjoining arrays. No extra decoupling devices are thus needed, and it is found that extension of this transfer-clear procedure to a chain of arrays yields a serial shift register (Figure 6) . Gain characteristics are readily achievable such that the register can be closed upon itself to obtain continuous circulation of stable, arbitrary ONE-ZERO patterns.\nReversing the output connections (Figure 7 ) of an array provides a view, looking left from the driven nodes, of a low threshold in CLEARED state, and a high threshold in SET or RESET states. This results in a complementary output from the array. Furthemore, it is easy to demonstrate that a multi-input array ( e .g . , A or B of Figure 8 ) assumes the state, after a succession of input transfers, which is the logical sum of the variables being transferred into it. Judicious use of the abilities\noutlined above can provide networks capable in principle of realizing any logical function.\nWorking networks of ferrite-toroid arrays have been built and shown to work with drive-current tolerances as high as 20%. The ONE/ZERO flux-switching ratios are usually of the order of ten to one. The basic practicability of this all-magnetic system generally seems quite good.\ni j R\nFigure &SET (or QNE) state, reached from CLEARED state only by positive I , exceeding combined threshold ( I t ) of cores Y and 2. During switching, IR exceeds 2 threshold\nand ( I , - I , ) must exceed Y threshold.\n66 Digesf of Technical Papers", + "1 I\n~\nL\nI c I I I\nclock sequence shown produces rightward shifting, and can cessive arrays of simple magnetic elements. Four-phase provide stable, arbitrary-pattern circulation if register is closed upon itself.\nFigure &Input characteristics at aa: Flux-linkage (voltseconds) change vs. I,. Path (2) from CLEARED to SET, then path (3) t o RESET. Subsequent I , alternations can\nthen follow ( 4 ) - (3) loop.\nstate.\ninformation.\nseveral successive outputs.\nDigest of Technical Papers 67" + ] + }, + { + "image_filename": "designv11_35_0001698_bf02080603-Figure1-1.png", + "original_path": "designv11-35/openalex_figure/designv11_35_0001698_bf02080603-Figure1-1.png", + "caption": "Fig. 1. The growth-rate curve of Neurosporc~ with increasing amounts of biotin, and the values of molasses I and II plotted against this curve.", + "texts": [ + " The growth in the controls with no biotin was inconsiderable, generally 0.2 mg in 6 days. An amount of biotin of 0.05 my could be measured accurately. However, even though the amount of inoculum remained constant, and the conidia were of the same age, the growth of different series varied a little from one experiment to the other. Presumably these fluctuations were due to the varying vigour of the stock-culture. At the beginning of the investigation, for instance, the shortest test-period that was suitable was that of 6 days (see Fig. 1 shows a s tandard biotin curve from which the biotin contents of two types of molasses were calculated. These values and the d ry weights of mycelia are given in Table II . The average biotincontent of molasses I was 124 x 10 -~ mv/mg, tha t of molasses II , 127 \u00d7 10 -3. The exper iment was repeated la ter when the values obta ined were 120 and 128, respectively. Determinat ions of the biot in content of a number of o ther types of molasses are summarized in Table I I I . F rom this table it is obvious tha t it is possible to determine biotin quanti t ies in the range between 0" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_35_0000027_j.procir.2020.05.200-Figure1-1.png", + "original_path": "designv11-35/openalex_figure/designv11_35_0000027_j.procir.2020.05.200-Figure1-1.png", + "caption": "Fig. 1. Demonstrators grouped by first and second period of the Collaborative Research Centre 1153.", + "texts": [ + " Processes such as friction welding, extrusion and uild-up welding are used to produce the semi-finished products. he components are then heated, before being formed. The transort and handling of the components before and after the actual orming process has a special role, which will be explained in the ollowing. In order to investigate the process of Tailored Forming, some emonstrators have been developed. The requirements for hanling are based on manufacturing processes of the demonstrators. he demonstrators are shown in Fig. 1 . For a better overview, the emonstrators are divided according to their funding period. In the rst period, several shafts W1-3, a bearing bush H and a bevel ear K were investigated, which are mainly rotationally symmet- under the CC BY-NC-ND license r m t i a t t c j i t w t b t s t k T 1 r t w p ( t o r a t o t 1 t r f w c s t a T o 2 h p 2 s i u l V m d t t O t t g o p m b g I a a d t g c s t a ( t p m t S r o o g p t a a g f t t l t h s g a H k ical components with diameters between 25mm and 90mm and asses varying between 0.5kg and 1kg. These have already passed hrough the Tailored Forming process. Each of these components s subject to different heating strategies, manufacturing processes nd material pairings. The demonstrators of the current period of he CRC 1153 are under development. However, it can be seen that he complexity in terms of shape and form is increasing. The most omplex demonstrator is the wishbone Q shown in Fig. 1 . A simple aw gripper does not suffice to handle the different shapes. A flexble universal gripper is needed, as it is characterized by its ability o handle a wide range of different objects. A closer look at the manufacturing process of the bevel gear K ill show what conditions the handling must withstand. Two ypes of steel are joined to form a semi-finished product by the uild-up welding process. Heating is carried out by induction and akes place from the outside. The gripper must therefore grip the urface, which is 10 0 0 \u25e6C hot" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_35_0000115_iros45743.2020.9340920-Figure2-1.png", + "original_path": "designv11-35/openalex_figure/designv11_35_0000115_iros45743.2020.9340920-Figure2-1.png", + "caption": "Fig. 2. A new gait pattern for a powered exoskeleton in which body inclination angle is taking account.", + "texts": [ + "(b) shows the early ground contact by inclining forward, which also ruins gait stability by causing a big impact due to continuous extension of the joints. If the ground contact occurs at the desired time as shown in Fig. 1.(c) by adjusting the body inclination angle during walking appropriately, the powered exoskeleton can effectively assist the paraplegic while maintaining stability. Therefore, the gait pattern for the powered exoskeleton need to consider two factors to fully assist the paraplegics: 1) the body inclination angle during walking, and 2) the compensation method for the individual not to have erroneous ground contact time. Figure. 2 shows the configuration of the powered exoskeleton at the moment of the initial contact. Notice that the shank of the front leg is positioned to be perpendicular to the ground depending on the body inclination angle. Each joint angle of the proposed gait pattern can be defined as, \u03b8 f hip = 1 lth arccos\u03bb + v, (1) \u03b8 f knee = \u03b8 f hip\u2212 v, (2) \u03b8 f ankle = 0, (3) \u03b8 i ankle = v\u2212\u03b8 i hip +\u03b8 i knee (4) where \u03bb = lth cos(v\u2212\u03b8 i hip)+ lsh cos(v\u2212\u03b8 i hip +\u03b8 i knee)\u2212 lsh. (5) lsh, lth, \u03b8 i hip and \u03b8 i hip represent the length of shank, the length of thigh, and target values of the hip and knee joints based on the range of motion (RoM) of the paraplegics data, respectively" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_35_0001419_978-94-007-7401-8-Figure3.30-1.png", + "original_path": "designv11-35/openalex_figure/designv11_35_0001419_978-94-007-7401-8-Figure3.30-1.png", + "caption": "Fig. 3.30 Overactuated PMs with coupled Sch\u00f6nflies motions of types 4RPRRR (a) and 4RRPRR (b) defined by MF \u00bc SF \u00bc 4; RF\u00f0 \u00de \u00bc v1; v2; v3;xd\u00f0 \u00de, TF = 0, NF = 2, limb topology R||P||R\\R||R (a) and R||R||P\\R||R (b)", + "texts": [ + " 3.3a) The first and the last revolute joints of the four limbs have parallel axes 15. 4RRRRR (Fig. 3.27b) R||R\\R||R\\R (Fig. 3.3b) Idem No. 14 16. 4RRRRR (Fig. 3.28a) R||R||R\\R||R (Fig. 3.3c) The first revolute joints of the four limbs have parallel axes 17. 4RRRRR (Fig. 3.28b) R||R\\R||R||R (Fig. 3.3d) The last revolute joints of the four limbs have parallel axes 18. 4PRRRR (Fig. 3.29a) P||R||R\\R||R (Fig. 3.3e) Idem No. 17 19. 4RPRRR (Fig. 3.29b) R||P||R\\R||R (Fig. 3.3f) Idem No. 17 20. 4RPRRR (Fig. 3.30a) R||P||R\\R||R (Fig. 3.3g) Idem No. 17 21. 4RRPRR (Fig. 3.30b) R||R||P\\R||R (Fig. 3.3h) Idem No. 17 22. 4RRRPR (Fig. 3.31a) R||R\\R\\P\\kR (Fig. 3.3i) Idem No. 17 23. 4RRRRP (Fig. 3.31b) R||R\\R||R\\kP (Fig. 3.3j) Idem No. 17 24. 4PRRRR (Fig. 3.32a) P\\R\\R||R\\R (Fig. 3.3k) The second and the last joints of the four limbs have parallel axes 25. 4RPRRR (Fig. 3.32b) R\\P\\R||R\\R (Fig. 3.3l) Idem No. 14 26. 4RRPRR (Fig. 3.33a) R\\R\\P\\kR\\R (Fig. 3.3m) Idem No. 14 27. 4RRRPR (Fig. 3.33b) R\\R||R\\P\\kR (Fig. 3.3n) Idem No. 14 28. 4RRRRP (Fig. 3.34a) R\\R||R\\R\\P (Fig. 3.3o) The first and the fourth revolute joints of the four limbs have parallel axes 29" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_35_0001419_978-94-007-7401-8-Figure5.37-1.png", + "original_path": "designv11-35/openalex_figure/designv11_35_0001419_978-94-007-7401-8-Figure5.37-1.png", + "caption": "Fig. 5.37 2PaPaRRR-1PaPaRR-1RUPU-type fully-parallel PM with uncoupled Sch\u00f6nflies motions defined by MF = SF = 4, RF\u00f0 \u00de \u00bc \u00f0v1; v2; v3;xb\u00de, TF = 0, NF = 20, limb topology Pa\\Pa||R||R\\\\R, Pa\\Pa||R||R and R\\R\\R\\P\\kR\\R", + "texts": [ + "1b) (continued) 530 5 Topologies with Uncoupled Sch\u00f6nflies Motions Table 5.3 (continued) No. PM type Limb topology Connecting conditions 8. 3PaPPaR1RUPU (Fig. 5.35b) Pa\\P\\\\Pa\\kR (Fig. 5.4b) Idem No. 7 Pa\\P\\\\Pa||R (Fig. 5.4c) R\\R\\R\\P\\kR\\R (Fig. 5.1b) 9. 3PaPaPR1RUPU (Fig. 5.36a) Pa\\Pa\\kP\\\\R (Fig. 5.4f) Idem No. 7 Pa\\Pa\\kP\\kR (Fig. 5.4h) R\\R\\R\\P\\kR\\R (Fig. 5.1b) 10. 3PaPaPR1RUPU (Fig. 5.36b) Pa\\Pa\\\\P\\\\R (Fig. 5.4g) Idem No. 7 Pa\\Pa\\\\P\\kR (Fig. 5.4e) R\\R\\R\\P\\kR\\R (Fig. 5.1b) 11. 2PaPaRRR1PaPaRR1RUPU (Fig. 5.37) Pa\\Pa||R||R\\\\R (Fig. 5.5a) The last joints of of limbs G1, G2 and G3 have superposed axes. The revolute joints between links 7 and 8 of limbs G1, G2 and G3 have orthogonal axes Pa\\Pa||R||R (Fig. 5.4i) R\\R\\R\\P\\kR\\R (Fig. 5.1b) 12. 2PaPaRRR1PaPaRR1RUPU (Fig. 5.38) Pa\\Pa||R||R\\kR (Fig. 5.5b) Idem No. 11 Pa\\Pa||R||R (Fig. 5.4j) R\\R\\R\\P\\kR\\R (Fig. 5.1b) Table 5.4 Limb topology and connecting conditions of the fully-parallel solutions with no idle mobilities presented in Figs. 5.39, 5.40, 5.41, 5.42, 5" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_35_0001419_978-94-007-7401-8-Figure2.8-1.png", + "original_path": "designv11-35/openalex_figure/designv11_35_0001419_978-94-007-7401-8-Figure2.8-1.png", + "caption": "Fig. 2.8 Fully-parallel PMs with coupled Sch\u00f6nflies motions of types 4PRRRR (a) and 4RPRRR (b) defined by MF \u00bc SF \u00bc 4; RF\u00f0 \u00de \u00bc \u00f0v1; v2; v3;xd\u00de, TF = 0, NF = 2, limb topology P||R||R||R||R\\R (a) and R||P||R||R\\R (b)", + "texts": [ + " 2.4b) P\\R\\R||R||R (Fig. 2.1e) The second joints of the four limbs have parallel axes 7. 4RRRPR (Fig. 2.5a) R||R\\R\\P\\kR (Fig. 2.1f) Idem No. 5 8. 4RRPRR (Fig. 2.5b) R||R||P\\R||R (Fig. 2.1g) Idem No. 5 9. 4PRRRR (Fig. 2.6a) P\\R\\R||R\\R (Fig. 2.1h) The second and the last joints of the four limbs have parallel axes 10. 4PRRRR (Fig. 2.6b) P||R\\R||R\\R (Fig. 2.1i) Idem No. 9 11. 4RRPRR (Fig. 2.7a) R\\R\\P\\kR\\R (Fig. 2.1j) Idem No. 1 12. 4PRRRR (Fig. 2.7b) P\\R||R\\R||R (Fig. 2.1k) Idem No. 3 13. 4PRRRR (Fig. 2.8a) P||R||R||R||R\\R (Fig. 2.1l) The last revolute joints of the four limbs have parallel axes 14. 4RPRRR (Fig. 2.8b) R||P||R||R\\R (Fig. 2.1m) Idem No. 13 15. 4RRPRR (Fig. 2.9a) R||R||P||R\\R (Fig. 2.1n) Idem No. 13 16. 4RRRPR (Fig. 2.9b) R||R||R||P\\R (Fig. 2.1o) Idem No. 13 17. 4RPRRR (Fig. 2.10a) R||P||R||R\\R (Fig. 2.1p) Idem No. 13 18. 4RRPRR (Fig. 2.10b) R||R||P||R\\R (Fig. 2.1q) Idem No. 13 19. 4RRRRP (Fig. 2.11a) R||R||R\\R\\P (Fig. 2.1r) The before last revolute joints of the four limbs have parallel axes 20. 4RPRRR (Fig. 2.11b) R\\P\\kR\\R||R (Fig. 2.1s) Idem No. 3 21. 4RRRRP (Fig. 2.12a) R\\R||R||R||P (Fig. 2" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_35_0001419_978-94-007-7401-8-Figure3.55-1.png", + "original_path": "designv11-35/openalex_figure/designv11_35_0001419_978-94-007-7401-8-Figure3.55-1.png", + "caption": "Fig. 3.55 Overactuated PMs with coupled Sch\u00f6nflies motions of types 4PPaRP (a) and 4PRRPa (b) defined by MF \u00bc SF \u00bc 4; RF\u00f0 \u00de \u00bc v1; v2; v3;xd\u00f0 \u00de, TF = 0, NF = 18, limb topology P||Pa||R\\P (a) and P||R||R||Pa (b)", + "texts": [ + "5 Limb topology and connecting conditions of the overactuated solutions with no idle mobilities presented in Figs. 3.54, 3.55, 3.56, 3.57, 3.58, 3.59, 3.60, 3.61, 3.62, 3.63, 3.64, 3.65, 3.66, 3.67, 3.68, 3.69, 3.70, 3.71, 3.72, 3.73, 3.74, 3.75, 3.76, 3.77, 3.78, 3.79, 3.80, 3.81, 3.82, 3.83, 3.84, 3.85, 3.86, 3.87, 3.88 No. PM type Limb topology Connecting conditions 1. 4PRPPa (Fig. 3.54a) P||R\\P\\kPa (Fig. 3.50a) The revolute joints of the four limbs have parallel axes 2. 4PRPaP (Fig. 3.54b) P||R||Pa\\P (Fig. 3.50b) Idem No. 1 3. 4PPaRP (Fig. 3.55a) P||Pa||R\\P (Fig. 3.50c) Idem No. 1 4. 4PRRPa (Fig. 3.55b) P||R||R||Pa (Fig. 3.50d) Idem No. 1 5. 4PPRPa (Fig. 3.56a) P\\P||R||Pa (Fig. 3.50e) Idem No. 1 6. 4PPRPa (Fig. 3.56b) P\\P\\kR||Pa (Fig. 3.50f) Idem No. 1 7. 4PRPPa (Fig. 3.57a) P\\R||P||Pa (Fig. 3.50g) Idem No. 1 8. 4PRPaP (Fig. 3.57b) P\\R||Pa||P (Fig. 3.50h) Idem No. 1 9. 4RPRPa (Fig. 3.58a) R||P||R||Pa (Fig. 3.50i) Idem No. 1 10. 4RRPPa (Fig. 3.58b) R||R||P||Pa (Fig. 3.50j) Idem No. 1 11. RRPaP (Fig. 3.59) R||R||Pa||P (Fig. 3.50k) Idem No. 1 12. 4RPaPP (Fig. 3.60a) R||Pa\\P\\kP (Fig. 3.50l) Idem No" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_35_0001419_978-94-007-7401-8-Figure2.16-1.png", + "original_path": "designv11-35/openalex_figure/designv11_35_0001419_978-94-007-7401-8-Figure2.16-1.png", + "caption": "Fig. 2.16 4PRRPR-type fully-parallel PM with coupled Sch\u00f6nflies motions defined by MF \u00bc SF \u00bc 4; RF\u00f0 \u00de \u00bc \u00f0v1; v2; v3;xd\u00de, TF = 0, NF = 2, limb topology P\\R\\R\\P\\kR", + "texts": [ + "1t) The first revolute joints of the four limbs have parallel axes 22. 4RRRPR (Fig. 2.12b) R\\R||R||P||R Fig. 2.1u) Idem No. 21 23. 4RRPRR (Fig. 2.13a) R\\R||P||R||R (Fig. 2.1v) Idem No. 21 24. 4RPRRR (Fig. 2.13b) R\\P||R||R||R (Fig. 2.1w) Idem No. 21 25. 4RRRPR (Fig. 2.14a) R\\R||R||P||R (Fig. 2.1x) Idem No. 21 26. 4RRPRR (Fig. 2.14b) R\\R||P||R||R||R (Fig. 2.1y) Idem No. 21 27. 4RPRRP (Fig. 2.15a) R\\P\\kR\\R\\P (Fig. 2.1z) Idem No. 19 28. 4PRRRP (Fig. 2.15b) P\\R||R\\R\\P (Fig. 2.1z1) Idem No. 19 29. 4PRRPR (Fig. 2.16) P\\R\\R\\P\\kR (Fig. 2.1a0) Idem No. 5 30. 4CRRR (Fig. 2.17a) C\\R||R\\R (Fig. 2.1d0) Idem No. 1 (continued) 2.1 Topologies with Simple Limbs 59 Table 2.1 (continued) No. PM type Limb topology Connecting conditions 31. 4CRRR (Fig. 2.17b) C||R||R\\R (Fig. 2.1e0) Idem No. 13 32. 4RRCR (Fig. 2.18a) R||R||C\\R (Fig. 2.1f0) Idem No. 13 33. 4RCRR (Fig. 2.18b) R||C||R\\R (Fig. 2.1g0) Idem No. 13 34. 4RRRC (Fig. 2.19a) R\\R||R||C (Fig. 2.1h0) Idem No. 21 35. 4RCRR (Fig. 2.19b) R\\C||R||R (Fig. 2.1l0) Idem No. 21 36" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_35_0001478_1.4035286-Figure6-1.png", + "original_path": "designv11-35/openalex_figure/designv11_35_0001478_1.4035286-Figure6-1.png", + "caption": "Fig. 6 Multi-DOF model of a blisk with shroud interfaces", + "texts": [ + " 5, where the peak is best approximated by constant-kept forced responses if their associated stiffness is evaluated near that resonance frequency. Apparently, the widened passage could mislead one to attribute a greater damping to the system, whereas this is clearly an effect of variable rotational speed. 3.2 Multi-DOF Model of a Blisk 3.2.1 Testcase Setup. The examined system is a principle cyclic symmetric model of a blisk with a Young\u2019s modulus E\u00bc 243,000 MPa, density . \u00bc 8900 kg=m3, and a Poisson\u2019s ratio of \u00bc 0.3, see Fig. 6. It exhibits cyclic shroud interface contact, pretwist of the blades, and cyclic boundary conditions at the disk and showcases the applicability of the proposed method to large-scale industrial test cases. This model will be analyzed exemplarily within a frequency range of 0\u2013800 Hz rotational speed; therefore, selected sample points for second-order interpolation are chosen to 0 Hz, 400 Hz, and 800 Hz. Truly, this is a quite large covered range, which in real cases should be selected more thoroughly and adapted to the problem depending on 062503-4 / Vol", + " Within the framework of MMR, the classical Craig\u2013Bampton approach is utilized containing 50 fixed interface normal modes and 36 DOF for describing the relative contact motion per sample model. In fact, not considering variable rotational speed (CMS-only) leads to 86 complex coordinates in total, whereas performing MMR returns 258 total coordinates (136 after SVD). If not otherwise specified, a hysteretic damping of g\u00bc 0.0032 is applied to the system. This equals about 1% logarithmic decrement as long as Rayleigh parameters a and b are kept zero. The excitation stimulus is fixed to a unit loading and acts along the axial direction at the location depicted in Fig. 6. The response is evaluated in circumferential direction. Since the harmonic order of the Fourier ansatz Nh does not influence the principle of variable speed, the analyses were mostly performed monoharmonically. Nevertheless, the extension to higher order is straightforward and results are also shown in the following. 3.2.2 Principle Dynamic Behavior. To show the main effects when facing variable rotational speed in conjunction with largescale nonlinear blisks, subsequently several Campbell and forced response diagrams are presented" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_35_0000816_j.mechmachtheory.2021.104477-Figure3-1.png", + "original_path": "designv11-35/openalex_figure/designv11_35_0000816_j.mechmachtheory.2021.104477-Figure3-1.png", + "caption": "Fig. 3. (a) High-speed indexing cam mechanism, and (b) its dwell-rise-dwell working cycle [16].", + "texts": [ + " (7) can be further improved, which will be detailed in Subsection 4.2.3. In this section, three numerical cases are presented to validate the effectiveness and versatility of the proposed framework in cam curve design and optimization, which are: (i) data-driven reconstruction of cam curves through PHOI, (ii) kinematics optimization based on pointwise scaling and piecewise modulation, and (iii) dynamics optimization based on PHOI and pointwise scaling. H. Luo et al. Mechanism and Machine Theory 167 (2022) 104477 Fig. 3(a) shows a typical indexing cam mechanism (ICM) used in a high-speed automatic production line, wherein the indexing plate is driven by the working profile of the globoidal cam whose rotational speed is up to 1800 r/min. The working cycle of this cam mechanism conforms to a dwell-rise-dwell (DRD) type (Fig. 3(b)), and the motion curve at the rising stage is a standard high-speed cam curve\u2013MS curve [16]. To validate the interpolation accuracy and convergence of the proposed PHOI method, the reconstruction of this DRD high-speed cam curve is performed in this subsection. First, by sampling the dimensionless MS curve, the multiorder motion parameters (Si, Vi, Ai and Ji) on multiple time nodes Ti are extracted, which are taken as the raw motion constraints for interpolation. Since classic Hermite, uniform B-spline and NURBS interpolation methods mainly focus on low-order motion constraints (Si, Vi and/or Ai) [17], two high-order derivative involved interpolation methods (CHOI given by Eq", + " While the original MS cam curve is generally considered as a standard high-speed motion law, two jerk jumps are observed at the beginning and end of the rising stage (Fig. 5(b)), which may degrade its dynamic performance at high speeds [6,20]. It is also noted that the absolute jerk |Jimp| at the crossover point (where the acceleration goes from positive to negative) is 23.16, indicating an appreciable impact velocity (Vimp = (4.5|Jimp|\u03b42/h2)1/3, where \u03b4 denotes the clearance between the roller and cam working profile [6]) at the rising stage, which is undesired in shape-locked intermittent mechanisms (such as the ICM in Fig. 3(a)). To reduce the crossover impact and realize the global jerk continuity, five interpolation nodes are specified, and the jerk scaling factors at nodes T = 0, 0.5 and 1 are assigned as \u03bbJ,1 = \u03bbJ,3 = \u03bbJ,5 = 0 (while other scaling factors are kept as 1). Then, the scaled motion constraints are determined, listed in Table 1. H. Luo et al. Mechanism and Machine Theory 167 (2022) 104477 By interpolating the scaled motion parameters (Table 1) with PHOI, a modified cam curve is obtained, with its derivative curves shown in Figs" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_35_0001419_978-94-007-7401-8-Figure5.25-1.png", + "original_path": "designv11-35/openalex_figure/designv11_35_0001419_978-94-007-7401-8-Figure5.25-1.png", + "caption": "Fig. 5.25 2PaRPRR-1PaRPR-1RPaPatP-type fully-parallel PMs with uncoupled Sch\u00f6nflies motions defined by MF = SF = 4, RF\u00f0 \u00de \u00bc \u00f0v1; v2; v3;xb\u00de, TF = 0, NF = 19, limb topology Pa\\R\\P\\kR\\R, Pa\\R\\P\\kR and R||Pa||Pat||P", + "texts": [ + "22b) Pa||P\\R||R\\\\R (Fig. 5.3b) Idem No. 9 Pa||P\\R||R (Fig. 5.2f) R||P||Pa||Pat (Fig. 5.4n) 17. 2PaRPRR-1PaRPR1RPaPaP (Fig. 5.23) Pa\\R\\P\\kR\\R (Fig. 5.3c) The last joints of the four limbs have superposed axes/directions. The revolute joints between links 4 and 5 of limbs G1, G2 and G3 have orthogonal axes Pa\\R\\P\\kR (Fig. 5.2g) R||Pa||Pa||P (Fig. 5.4k) 18. 2PaRPRR-1PaRPR1RPPaPa (Fig. 5.24) Pa\\R\\P\\kR\\R (Fig. 5.3c) Idem No. 17 Pa\\R\\P\\kR (Fig. 5.2g) R||P||Pa||Pa (Fig. 5.4m) 19. 2PaRPRR-1PaRPR1RPaPatP (Fig. 5.25) Pa\\R\\P\\kR\\R (Fig. 5.3c) Idem No. 17 Pa\\R\\P\\kR (Fig. 5.2g) R||Pa||Pat||P (Fig. 5.4l) 20. 2PaRPRR-1PaRPR1RPPaPat (Fig. 5.26) Pa\\R\\P\\kR\\R (Fig. 5.3c) The revolute joint between links 7D and 9 of G4 limb and the last revolute joints of limbs G1, G2 and G3 have superposed axes. The revolute joints between links 4 and 5 of limbs G1, G2 and G3 have orthogonal axes Pa\\R\\P\\kR (Fig. 5.2g) R||P||Pa||Pat (Fig. 5.4n) (continued) 5.1 Fully-Parallel Topologies 529 Table 5.2 (continued) No. PM type Limb topology Connecting conditions 21" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_35_0001419_978-94-007-7401-8-Figure2.65-1.png", + "original_path": "designv11-35/openalex_figure/designv11_35_0001419_978-94-007-7401-8-Figure2.65-1.png", + "caption": "Fig. 2.65 4PaPaRP-type fully-parallel PM with coupled Sch\u00f6nflies motions defined by MF \u00bc SF \u00bc 4; RF\u00f0 \u00de \u00bc \u00f0v1; v2; v3;xd\u00de, TF = 0, NF = 30, limb topology Pa\\Pa\\\\R||P", + "texts": [ + "22q) Idem No. 33 45. 4PPaPaR (Fig. 2.57) P||Pa||Pa\\R (Fig. 2.22r) Idem No. 4 46. 4PaPPaR (Fig. 2.58) Pa||P||Pa\\R (Fig. 2.22s) Idem No. 4 47. 4RPPaPa (Fig. 2.59) R\\P||Pa||Pa (Fig. 2.22t) Idem No. 15 48. 4PaPPaR (Fig. 2.60) Pa\\P\\\\Pa\\kR (Fig. 2.22u) Idem No. 4 49. 4PaRPaP (Fig. 2.61) Pa\\R\\Pa\\kR (Fig. 2.22v) Idem No. 34 50. 4PaRPPa (Fig. 2.62) Pa\\R||P\\Pa (Fig. 2.22w) Idem No. 34 51. 4PaPRPa (Fig. 2.63) Pa\\P||R\\Pa (Fig. 2.22x) Idem No. 33 52. 4PaPaPR (Fig. 2.64) Pa\\Pa\\\\P||R (Fig. 2.22y) Idem No. 4 53. 4PaPaRP (Fig. 2.65) Pa\\Pa\\\\R||P (Fig. 2.22z) Idem No. 41 54. 4RPaPaP (Fig. 2.66) R\\Pa\\\\Pa\\\\P (Fig. 2.22a0) Idem No. 15 55. 4RPaPPa (Fig. 2.67) R\\Pa\\kP\\\\Pa (Fig. 2.22b0) Idem No. 15 56. 4RPPaPa (Fig. 2.68) R||P\\Pa\\\\Pa (Fig. 2.22c0) Idem No. 15 57. 4PRPaPa (Fig. 2.69) P||R\\Pa\\\\Pa (Fig. 2.22d0) Idem No. 5 58. 4PaCPa (Fig. 2.70) Pa\\C\\Pa (Fig. 2.22e0) The axes of the cylindrical joints of the four limbs are parallel 59. 4CPaPa (Fig. 2.71) C\\Pa\\\\Pa (Fig. 2.22f 0) Idem No. 58 60. 4PaPaC (Fig. 2.72) Pa\\Pa\\\\C (Fig. 2.22g0) Idem No" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_35_0001419_978-94-007-7401-8-Figure2.63-1.png", + "original_path": "designv11-35/openalex_figure/designv11_35_0001419_978-94-007-7401-8-Figure2.63-1.png", + "caption": "Fig. 2.63 4PaPRPa-type fully-parallel PM with coupled Sch\u00f6nflies motions defined by MF \u00bc SF \u00bc 4; RF\u00f0 \u00de \u00bc \u00f0v1; v2; v3;xd\u00de, TF = 0, NF = 30, limb topology Pa\\P||R\\Pa", + "texts": [ + "22o) Idem No. 4 43. 4PRPaPa (Fig. 2.55) P\\R\\Pa||Pa (Fig. 2.22p) Idem No. 5 44. 4PPaRPa (Fig. 2.56) P||Pa\\R\\Pa (Fig. 2.22q) Idem No. 33 45. 4PPaPaR (Fig. 2.57) P||Pa||Pa\\R (Fig. 2.22r) Idem No. 4 46. 4PaPPaR (Fig. 2.58) Pa||P||Pa\\R (Fig. 2.22s) Idem No. 4 47. 4RPPaPa (Fig. 2.59) R\\P||Pa||Pa (Fig. 2.22t) Idem No. 15 48. 4PaPPaR (Fig. 2.60) Pa\\P\\\\Pa\\kR (Fig. 2.22u) Idem No. 4 49. 4PaRPaP (Fig. 2.61) Pa\\R\\Pa\\kR (Fig. 2.22v) Idem No. 34 50. 4PaRPPa (Fig. 2.62) Pa\\R||P\\Pa (Fig. 2.22w) Idem No. 34 51. 4PaPRPa (Fig. 2.63) Pa\\P||R\\Pa (Fig. 2.22x) Idem No. 33 52. 4PaPaPR (Fig. 2.64) Pa\\Pa\\\\P||R (Fig. 2.22y) Idem No. 4 53. 4PaPaRP (Fig. 2.65) Pa\\Pa\\\\R||P (Fig. 2.22z) Idem No. 41 54. 4RPaPaP (Fig. 2.66) R\\Pa\\\\Pa\\\\P (Fig. 2.22a0) Idem No. 15 55. 4RPaPPa (Fig. 2.67) R\\Pa\\kP\\\\Pa (Fig. 2.22b0) Idem No. 15 56. 4RPPaPa (Fig. 2.68) R||P\\Pa\\\\Pa (Fig. 2.22c0) Idem No. 15 57. 4PRPaPa (Fig. 2.69) P||R\\Pa\\\\Pa (Fig. 2.22d0) Idem No. 5 58. 4PaCPa (Fig. 2.70) Pa\\C\\Pa (Fig. 2.22e0) The axes of the cylindrical joints of the four limbs are parallel 59" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_35_0001419_978-94-007-7401-8-Figure6.44-1.png", + "original_path": "designv11-35/openalex_figure/designv11_35_0001419_978-94-007-7401-8-Figure6.44-1.png", + "caption": "Fig. 6.44 Maximaly regular fully-parallel PMs with Sch\u00f6nflies motions of types 3PPaPR1CPaPa (a) and 3PPaPR-1CPaPat (b) defined by MF = SF = 4, (RF) = (v1; v2; v3;xb), TF = 0, NF = 21, limb topology P||Pa\\P ??R, P||Pa\\P\\||R and C||Pa||Pa (a), C||Pa||Pat (b)", + "texts": [ + " The actuated prismatic joints of limbs G1, G2 and G3 have orthogonal directions 2. 3PPPaR-1CPaPa (Fig. 6.42b) P\\P\\||Pa\\\\R (Fig. 4.8a) P\\P\\||Pa||R, (Fig. 4.8c) C||Pa||Pa (Fig. 5.4o) Idem no. 1 (continued) 6.3 Fully-Parallel Topologies with Complex Limbs 643 Table 6.14 (continued) No. PM type Limb topology Connecting conditions 3. 3PPPaR-1CPaPat (Fig. 6.43a) P\\P\\||Pa\\||R (Fig. 4.8b) P\\P\\||Pa||R, (Fig. 4.8c) C||Pa||Pat (Fig. 5.4p) Idem no. 1 4. 3PPPaR-1CPaPat (Fig. 6.43b) P\\P\\||Pa\\R (Fig. 4.8a) P\\P\\||Pa||R, (Fig. 4.8c) C||Pa||Pat (Fig. 5.4p) Idem no. 1 5. 3PPaPR-1CPaPa (Fig. 6.44a) P||Pa\\P\\\\R (Fig. 4.8d) P||Pa\\P\\||R (Fig. 4.8e) C||Pa||Pa (Fig. 5.4o) Idem no. 1 6. 3PPaPR-1CPaPat (Fig. 6.44b) P||Pa\\P\\\\R (Fig. 4.8d) P||Pa\\P\\||R (Fig. 4.8e) C||Pa||Pat (Fig. 5.4p) Idem no. 1 7. 2PPaRRR-1PPaRR1CPaPa (Fig. 6.45) P||Pa||R||R\\R (Fig. 4.9a) P||Pa||R||R (Fig. 4.8g) C||Pa||Pa (Fig. 5.4o) Idem no. 1 8. 2PPaRRR-1PPaRR1CPaPat (Fig. 6.46) P||Pa||R||R\\R (Fig. 4.9a) P||Pa||R||R (Fig. 4.8g) C||Pa||Pat (Fig. 5.4p) Idem no. 1 Table 6.15 Structural parametersa of parallel mechanisms in Figs. 6.27, 6.28, 6.29, 6.30, 6.31, 6.32, 6.33, 6.34, 6.35, 6.36, 6.37 No. Structural parameter Solution Figures 6.27, 6" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_35_0001419_978-94-007-7401-8-Figure5.108-1.png", + "original_path": "designv11-35/openalex_figure/designv11_35_0001419_978-94-007-7401-8-Figure5.108-1.png", + "caption": "Fig. 5.108 1PPn2R-2PPn2RR-1PPPn2RR type redundantly actuated PM with uncoupled Sch\u00f6nflies motions defined by MF = 5, SF = 4, RF\u00f0 \u00de \u00bc \u00f0v1; v2; v3;xa\u00de, TF = 1, NF = 15, limb topology P||Pn2||R, P||Pn2||R\\R, P||P||Pn2||R\\R", + "texts": [ + "104) P||Pa\\P\\kR (Fig. 4.8e) Idem No. 1 P||Pa\\P||R (Fig. 4.8f) P||P||Pa\\P||R (Fig. 5.99d) 5. 1PPaRR-2PPaRRR-1PPPaRRR (Fig. 5.105) P||Pa||R||R (Fig. 4.8g) Idem No. 1 P||Pa||R||R\\R (Fig. 4.9a) P||P||Pa||R||R\\R (Fig. 5.99e) 6. 1PRRbR-2PRRbRR-1PPRRbRR (Fig. 5.106) P||R||Rb||R (Fig. 4.1i) Idem No. 1 P||R||Rb||R\\R (Fig. 4.9c) P||P||R||Rb||R\\R (Fig. 5.99f) 7. 1PPn2R-2PPn2RR-1PPPn2RR (Fig. 5.107) P||Pn2||R (Fig. 4.8j) Idem No. 1 P||Pn2||R\\R (Fig. 4.9d) P||P||Pn2||R\\R (Fig. 5.99g) 8. 1PPn2R-2PPn2RR-1PPPn2RR (Fig. 5.108) P||Pn2||R (Fig. 4.8k) Idem No. 1 P||Pn2||R\\R (Fig. 4.9e) P||P||Pn2||R\\R (Fig. 5.99h) 9. 1PPn3-2PPn3R-1PPPn3R (Fig. 5.109) P||Pn3 (Fig. 4.8l) Idem No. 1 P||Pn3\\R (Fig. 4.9f) P||P||Pn3\\R (Fig. 5.99i) 10. 1PPn3-2PPn3R-1PPPn3R (Fig. 5.110) P||Pn3 (Fig. 4.8m) Idem No. 1 P||Pn3\\R (Fig. 4.9g) P||P||Pn3\\R (Fig. 5.99j) 11. 1PPaPR-2PPaC-1PPPaC (Fig. 5.111) P||Pa\\P\\kR (Fig. 4.8e) Idem No. 1 P||Pa\\C (Fig. 4.8o) P||P||Pa\\C (Fig. 5.99k) 12. 1CRbR-2CRbRR-1PCRbRR (Fig. 5.112) C||Rb||R (Fig. 4.8n) Idem No. 1 C||Rb||R\\R (Fig" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_35_0000395_978-3-030-67750-3_16-Figure17.5-1.png", + "original_path": "designv11-35/openalex_figure/designv11_35_0000395_978-3-030-67750-3_16-Figure17.5-1.png", + "caption": "Fig. 17.5 Ideal gearbox: steady-state model", + "texts": [ + " The vehicle is not accelerating, dv/dt is 0, and FTR is the maximum tractive force delivered by the motor at or near zero speed. With the assumptions, at near stall conditions, \u2211 F = 0 \u2192 FTR \u2212 FgxT = 0 \u2192 FTR = mg sin \u03b2 (17.8) If required mass (m) = 1300 kg, (full load) FTR = 1300 \u00d7 9.81 sin 18 = 3940N the maximum percent grade is max % grade = 100 tan \u03b2, max% grade = 100FTR \u221a (mg)2 \u2212 FTR2 = 100 \u00d7 3940 \u221a (1300 \u00d7 9.81)2 \u2212 (3940)2 = 32.48% TheEV transmission equation has established by assuming an ideal gearbox as shown in Fig. 17.5 with Plosses is 0, and the efficiency is 100%, perfectly rigid gears, and no gear backlash. For a tire wheel with radius r, the tangential and the angular velocity are related by: \u03c9r = v, \u03c9 = v r (17.9) The tangential velocity at the gear teeth contact point is the same for the two gears with different radius. rin\u03c9in = v = rout\u03c9out (17.10) The gear ratio has defined in terms of speed transformation between the input shaft and the output shaft. GR = \u03c9in \u03c9out = rout rin (17.11) Assuming 100% efficiency of the gear train: Pout = Pin,\u21d2 Tout\u03c9out = Tin\u03c9in (17" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_35_0000101_iccr51572.2020.9344394-Figure1-1.png", + "original_path": "designv11-35/openalex_figure/designv11_35_0000101_iccr51572.2020.9344394-Figure1-1.png", + "caption": "Fig. 1. Earth and Body fixed reference System.", + "texts": [ + " Section V presents the robustness properties, and finally section VI draws the conclusions. 154 978-1-7281-7562-1/20/$31.00 \u00a92020 IEEE 20 20 3 rd In te rn at io na l C on fe re nc e on C on tr ol a nd R ob ot s ( IC CR ) | 9 78 -1 -7 28 1- 75 62 -1 /2 0/ $3 1. 00 \u00a9 20 20 IE EE | D O I: 10 .1 10 9/ IC CR 51 57 2. 20 20 .9 34 Authorized licensed use limited to: Rutgers University. Downloaded on May 20,2021 at 12:52:35 UTC from IEEE Xplore. Restrictions apply. While investigating the motion of AUV in 6 DOF it is essential to define two coordinate systems as shown in Fig. 1. The moving coordinate system (O-xyz) is fixed to the vehicle and is called the body-fixed reference system, [4]. The motion of body-fixed reference system is depicted to earth fixed reference system (\u03c6,\u03b8,\u03a8). The six DOF AUV\u2019s equation of motion follows rigid body dynamics. Let us consider, Q=[N ,B], where N=[x,y,z] and B=(\u03c6,\u03b8,\u03a8) are the position and orientation of body and earth fixed reference system. Body fixed system B is transformed into earth fixed reference system N. Linear and angular velocities are denoted as V1=[u,v,w] and V2=[p,q,r] respectively" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_35_0001419_978-94-007-7401-8-Figure2.11-1.png", + "original_path": "designv11-35/openalex_figure/designv11_35_0001419_978-94-007-7401-8-Figure2.11-1.png", + "caption": "Fig. 2.11 Fully-parallel PMs with coupled Sch\u00f6nflies motions of types 4RRRRP (a) and 4RPRRR (b) defined by MF \u00bc SF \u00bc 4; RF\u00f0 \u00de \u00bc \u00f0v1; v2; v3;xd\u00de, TF = 0, NF = 2, limb topology R||R||R\\R\\P (a) and R\\P\\kR\\R||R (b)", + "texts": [ + "1j) Idem No. 1 12. 4PRRRR (Fig. 2.7b) P\\R||R\\R||R (Fig. 2.1k) Idem No. 3 13. 4PRRRR (Fig. 2.8a) P||R||R||R||R\\R (Fig. 2.1l) The last revolute joints of the four limbs have parallel axes 14. 4RPRRR (Fig. 2.8b) R||P||R||R\\R (Fig. 2.1m) Idem No. 13 15. 4RRPRR (Fig. 2.9a) R||R||P||R\\R (Fig. 2.1n) Idem No. 13 16. 4RRRPR (Fig. 2.9b) R||R||R||P\\R (Fig. 2.1o) Idem No. 13 17. 4RPRRR (Fig. 2.10a) R||P||R||R\\R (Fig. 2.1p) Idem No. 13 18. 4RRPRR (Fig. 2.10b) R||R||P||R\\R (Fig. 2.1q) Idem No. 13 19. 4RRRRP (Fig. 2.11a) R||R||R\\R\\P (Fig. 2.1r) The before last revolute joints of the four limbs have parallel axes 20. 4RPRRR (Fig. 2.11b) R\\P\\kR\\R||R (Fig. 2.1s) Idem No. 3 21. 4RRRRP (Fig. 2.12a) R\\R||R||R||P (Fig. 2.1t) The first revolute joints of the four limbs have parallel axes 22. 4RRRPR (Fig. 2.12b) R\\R||R||P||R Fig. 2.1u) Idem No. 21 23. 4RRPRR (Fig. 2.13a) R\\R||P||R||R (Fig. 2.1v) Idem No. 21 24. 4RPRRR (Fig. 2.13b) R\\P||R||R||R (Fig. 2.1w) Idem No. 21 25. 4RRRPR (Fig. 2.14a) R\\R||R||P||R (Fig. 2.1x) Idem No. 21 26. 4RRPRR (Fig. 2.14b) R\\R||P||R||R||R (Fig. 2.1y) Idem No. 21 27. 4RPRRP (Fig. 2.15a) R\\P\\kR\\R\\P (Fig. 2.1z) Idem No" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_35_0001419_978-94-007-7401-8-Figure2.42-1.png", + "original_path": "designv11-35/openalex_figure/designv11_35_0001419_978-94-007-7401-8-Figure2.42-1.png", + "caption": "Fig. 2.42 4PRPaPa-type fully-parallel PM with coupled Sch\u00f6nflies motions defined by MF \u00bc SF \u00bc 4; RF\u00f0 \u00de \u00bc \u00f0v1; v2; v3;xd\u00de, TF = 0, NF = 30, limb topology P\\R||Pa\\Pa", + "texts": [ + "21y) The first joints of the four limbs have parallel axes 26. 4PaPC (Fig. 2.38b) Pa||P\\C (Fig. 2.21z) Idem No. 24 27. 4PaCP (Fig. 2.39) Pa\\C\\P (Fig. 2.21z1) Idem No. 11 28. 4PPaPaR (Fig. 2.40) P\\Pa\\\\Pa\\kR (Fig. 2.22a) Idem No. 4 29. 4PPaRPa (Fig. 2.41) P\\Pa||R\\Pa (Fig. 2.22b) The axes of the revolute joints connecting the two parallelogram loops of the four limbs are parallel (continued) 2.2 Topologies with Complex Limbs 175 Table 2.3 (continued) No. PM type Limb topology Connecting conditions 30. 4PRPaPa (Fig. 2.42) P\\R||Pa\\Pa (Fig. 2.22c) Idem No. 5 31. 4RPPaPa (Fig. 2.43) R\\P\\kPa\\\\Pa (Fig. 2.22d) Idem No. 15 32. 4PaPPaR (Fig. 2.44) Pa\\P\\\\Pa\\\\R (Fig. 2.22e) Idem No. 4 33. 4PaPRPa (Fig. 2.45) Pa\\P\\kR\\Pa (Fig. 2.22f) The axes of the revolute joints connecting links 5 and 6 of the four limbs are parallel 34. 4PaRPPa (Fig. 2.46) Pa||R\\P\\\\Pa (Fig. 2.22g) The axes of the revolute joints connecting links 4 and 5 of the four limbs are parallel 35. 4RPaPPa (Fig. 2.47) R||Pa\\P\\\\Pa (Fig. 2.22h) Idem No. 15 36. 4PaPaPR (Fig" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_35_0001419_978-94-007-7401-8-Figure5.26-1.png", + "original_path": "designv11-35/openalex_figure/designv11_35_0001419_978-94-007-7401-8-Figure5.26-1.png", + "caption": "Fig. 5.26 2PaRPRR-1PaRPR-1RPPaPat-type fully-parallel PMs with uncoupled Sch\u00f6nflies motions defined by MF = SF = 4, RF\u00f0 \u00de \u00bc \u00f0v1; v2; v3;xb\u00de, TF = 0, NF = 19, limb topology Pa\\R\\P\\kR\\R, Pa\\R\\P\\kR and R||P||Pa||Pat", + "texts": [ + "23) Pa\\R\\P\\kR\\R (Fig. 5.3c) The last joints of the four limbs have superposed axes/directions. The revolute joints between links 4 and 5 of limbs G1, G2 and G3 have orthogonal axes Pa\\R\\P\\kR (Fig. 5.2g) R||Pa||Pa||P (Fig. 5.4k) 18. 2PaRPRR-1PaRPR1RPPaPa (Fig. 5.24) Pa\\R\\P\\kR\\R (Fig. 5.3c) Idem No. 17 Pa\\R\\P\\kR (Fig. 5.2g) R||P||Pa||Pa (Fig. 5.4m) 19. 2PaRPRR-1PaRPR1RPaPatP (Fig. 5.25) Pa\\R\\P\\kR\\R (Fig. 5.3c) Idem No. 17 Pa\\R\\P\\kR (Fig. 5.2g) R||Pa||Pat||P (Fig. 5.4l) 20. 2PaRPRR-1PaRPR1RPPaPat (Fig. 5.26) Pa\\R\\P\\kR\\R (Fig. 5.3c) The revolute joint between links 7D and 9 of G4 limb and the last revolute joints of limbs G1, G2 and G3 have superposed axes. The revolute joints between links 4 and 5 of limbs G1, G2 and G3 have orthogonal axes Pa\\R\\P\\kR (Fig. 5.2g) R||P||Pa||Pat (Fig. 5.4n) (continued) 5.1 Fully-Parallel Topologies 529 Table 5.2 (continued) No. PM type Limb topology Connecting conditions 21. 2PaRRRR-1PaRRR1RPaPaP (Fig. 5.27) Pa\\R||R||R\\kR (Fig. 5.3d) Idem No. 17 Pa\\R||R||R (Fig. 5.2h) R||Pa||Pa||P (Fig" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_35_0000174_j.matpr.2021.01.640-Figure7-1.png", + "original_path": "designv11-35/openalex_figure/designv11_35_0000174_j.matpr.2021.01.640-Figure7-1.png", + "caption": "Fig. 7. Input values applied on the drive shaft.", + "texts": [ + " Mesh influences the accuracy, convergence, speed of the solution, Here used triangular type of mesh is always quick and easy to create. Drive shaft model is divided into finite no of convenient sub elements and each element corners are joined with adjacent element for the purpose of finding bonding strength between sub elements. The fixed support is suitable for the drive shaft .Rotational velocity 650 rad\\s is applied on the drive shaft. The moment is applied 350 N.m on the drive shaft. In this analysis fixed dynamic is suitable are shown in Fig. 7. The rotational velocity 650 rad\\s is applied on the drive shaft. The drive shaft rotations are 650,0, 2.063 rad\\s. The locations are 795.07, 7.3163 mm are shown in Fig. 8. The moment load does not cause rotation, the Moment is applied on the drive shaft is 350 N.m as shown in Fig. 9 in order to omit more bending stress. When composite drive shaft breaks, it is divided into fine fibers that do not have any danger for the driver. They have a less specific modulus and less damping capacity. So that conventional drive shaft is replaced with composite materials Fig" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_35_0001584_978-1-4614-8544-5_15-Figure15.19-1.png", + "original_path": "designv11-35/openalex_figure/designv11_35_0001584_978-1-4614-8544-5_15-Figure15.19-1.png", + "caption": "FIGURE 15.19. Behavior of 2 as a function of , .", + "texts": [ + " The fourth invariant frequency 4 is shown in Figure 15.15. 4 increases when decreases. However, 4 settles when & 0 6. lim 0 2 = (15.76) To have a better picture about the behavior of invariant frequencies, Figures 15.17 and 15.18 depict the relative frequency ratio 4 3 and 3 2. Example 578 Frequency response at invariant frequencies. The frequency response is a function of , , and . Damping always diminishes the amplitude of vibration, so at rst we set = 0 and plot the behavior of as a function of , . Figure 15.19 illustrates the behavior of at the second invariant frequency 2. Because lim 0 2 = 1 (15.77) 2 starts at one, regardless of the value of . The value of 2 is always greater than one. Figure 15.20 shows that 3 is not a function of and is a decreasing function of . Figure 15.21 shows that 4 1 regardless of the value of and . The relative behavior of 2, 3, and 4 is shown in Figures 15.22 and 15.23. Example 579 Natural frequencies and vibration isolation of a quarter car. For a modern typical passenger car, the values of natural frequencies are around 1Hz and 10Hz respectively" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_35_0000156_s12666-021-02208-7-Figure2-1.png", + "original_path": "designv11-35/openalex_figure/designv11_35_0000156_s12666-021-02208-7-Figure2-1.png", + "caption": "Fig. 2 Sampling schematic of tensile and microhardness tests (unit: mm)", + "texts": [ + " Due to the high thermal conductivity of the copper plate, the filler material was needed to be biased toward the copper plate side, and the offset distance was about 3 mm from the weld center. That was to obtain more heat on the copper side to melt the copper plate, so that it could achieve good metallurgical bond with the stainless steel plate. The welding conditions and parameters are listed in Table 2. After welding, all samples used for the performance characterization were extracted from the welded joint using wire electrical discharge machining, as shown in Fig. 2. With the purpose of understanding the feature of microstructure, the metallographic specimens were etched by 25 g FeCl3 ? 25 ml HCl ? 100 ml H2O, and the etched specimen was placed under an optic microscope (OM) to observe the microstructure of the joint. In addition, the scanning electron microscope (SEM) equipped with energy-dispersive X-ray spectroscopy (EDS) was used to carry out more detailed observation of microstructure and study the variation of chemical composition across the weld interface" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_35_0001395_s00170-021-07842-5-Figure4-1.png", + "original_path": "designv11-35/openalex_figure/designv11_35_0001395_s00170-021-07842-5-Figure4-1.png", + "caption": "Fig. 4 Scheme of a device for assessing wear resistance", + "texts": [ + " In these studies, the contribution to the diffraction pattern was made not only by the surface layers of the sample, but by the entire volume of the substance through which the radiation flux passes. The phase composition was studied using a beam of monochromatic synchrotron radiation with a cross section of 100 \u00d7 400\u03bcm. The resulting diffraction rings are integrated over the radius and recalculated for \u03bb = 1.5406 \u00c5. Wear resistance was evaluated using tests carried out by interaction of the test sample with abrasive paper (Fig. 4). The use of a spiral trajectory on abrasive paper ensures a stable wear process of the sample, because it provides a constant contact of the sample surface with new abrasive particles. The tests were carried out at a load of 10 N for 30 min in air, with P80 abrasive paper (particle size of 200 \u03bcm); the rotation speed was 150 rpm. On average, three measurements were performed on each of the composites. The roughness of the working surface before testing was not more than Ra 2.5. It is necessary to optimize the parameters of the laser exposure of the powder mixture during the formation of single tracks to form a high-quality multilayer cermet structure" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_35_0000577_s40313-021-00754-5-Figure1-1.png", + "original_path": "designv11-35/openalex_figure/designv11_35_0000577_s40313-021-00754-5-Figure1-1.png", + "caption": "Fig. 1 UX-6 fixed-wing UAV", + "texts": [ + "ol.:(0123456789) Keywords UAV\u00a0\u00b7 Fixed-wing aircraft modeling\u00a0\u00b7 Datcom + pro\u00a0\u00b7 System Identification\u00a0\u00b7 Analytical-empirical comparison The UX-6 UAV shown in Fig.\u00a01 is a fixed-wing UAV developed by The Department of Computer Science and Electronics, Universitas Gadjah Mada (Priyambodo et\u00a0al., 2016a). It was designed as an aerial mapping platform with flying wing configuration. The UX-6 is classified as a small UAV with less than 2\u00a0m wingspan. The UX-6 UAV autonomous control system is being developed, using a model-based design approach (Aarenstrup, 2015). This approach needs a mathematical or flight model, as the base for system development. Through the flight model, the aircraft attitude and motion could be learned, as well as the aircraft controller design could be simulated like a linear quadratic regulator (LQR) controller (Dhewa et\u00a0al" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_35_0000522_j.biosystems.2021.104451-Figure5-1.png", + "original_path": "designv11-35/openalex_figure/designv11_35_0000522_j.biosystems.2021.104451-Figure5-1.png", + "caption": "Fig. 5. Swarm robot motion platform in simulation experiment.", + "texts": [ + " If the detected obstacle angle is negative, the swarm robot rotates right; if it is positive, it turns left. Angular velocity of the swarm robot is determined as( \u2212 \u03c0) for the left turn and a right turn \u03c0). Linear velocity of the swarm robot is determined based on the coefficient k1 subject to the distance of the obstacle (OMr,n ). The experiments applied in this study were performed in the Mobile Robotic Toolbox in MATLAB environment. Mobile Robotic Toolbox is a simulation developed for swarm robot applications. Fig. 5 shows the motion platform for the swarm robots used in the simulation experiments. The dimensions of this platform can be changed and different numbers of swarm robots can be placed inside. The swarm robots used in the experiments performed were nonholonomic differential drive 2-wheel model with similar features. The swarm robots were circular with a radius of 0.40 units. The maximum velocity of the swarm robot is vmax = 3 unit step . The k1coefficient specified in Equations (5)\u2013(7) is calculated according to vmaxand LRF. k1is as expressed in equation (8). k1 = vmax LRF (8) As shown in Fig. 5, the detection range of the swarm robot for the obstacle detector has a circular area with a radius of 1unit with red dashes, and the detection range for the robot detector has a circular area with a radius of LRF. A zero average, 0.2 deviation value of noise in N (0, 0.2) form in Gaussian distribution was applied to the sensor inputs of the swarm robot used in the simulation environment. The simulation experiment parameters used to examine the flexibility and scalability of the DISA method in the study are given in Table 2" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_35_0001419_978-94-007-7401-8-Figure4.6-1.png", + "original_path": "designv11-35/openalex_figure/designv11_35_0001419_978-94-007-7401-8-Figure4.6-1.png", + "caption": "Fig. 4.6 Fully-parallel PMs with decoupled Sch\u00f6nflies motions of types 1PPPR-3PPC (a) and 1CPR-3CPRR (b) defined by MF = SF = 4, RF\u00f0 \u00de \u00bc v1; v2; v3;xa\u00f0 \u00de, TF = 0, NF = 6 (a), NF = 3 (b), limb topology P ? P ?? P ?? R and P ? P ??C (a), C ? P ?||R and C ? P ?||R ? R (b)", + "texts": [ + " The actuated prismatic joints of limbs G3 and G4 have parallel directions 2. 4PPPR (Fig. 4.3b) P ?P ??P ??R (Fig. 2.1b) P ?P ??P||R (Fig. 4.1c) Idem No. 1 3. 1PPRR-3PPRRR (Fig. 4.4a) P ?P ?|R||R (Fig. 4.1d) P ?P ?|R||R ?R (Fig. 4.2a) Idem No. 1 4. 1PRPR-3PRPRR (Fig. 4.4b) P||R ?P ?|R (Fig. 4.1e) P||R ?P ?|R ?R (Fig. 4.2b) Idem No. 1 5. 1PRRP-3PRRPR (Fig. 4.5a) P||R||R ?P (Fig. 4.1f) P||R||R ?P ??R (Fig. 4.2c) Idem No. 1 6. 1PRRR-3PRRRR (Fig. 4.5b) P||R||R||R (Fig. 4.1g) P||R||R||R ?R (Fig. 4.2d) Idem No. 1 7. 1PPPR-3PPC (Fig. 4.6a) P ?P ??P ??R (Fig. 4.1b) P ?P ??C (Fig. 4.1h) The last joints of the four limbs have parallel axes. The actuated prismatic joints of limbs G1, G2 and G3 hvave orthogonal directions. The actuated prismatic joints of limbs G3 and G4 have parallel directions 8. 1CPR-3CPRR (Fig. 4.6b) C ?P ?|R (Fig. 4.1i) C ?P ?|R ?R (Fig. 4.2e) Idem No. 1 9. 1CRP-3CRPR (Fig. 4.7a) C||R ?P (Fig. 4.1j) C||R ?P ??R (Fig. 4.2f) Idem No. 1 10. 1CRR-3CRRR (Fig. 4.7b) C||R||R (Fig. 4.1k) C||R||R ?R (Fig. 4.2g) Idem No. 1 376 4 Fully-Parallel Topologies with Decoupled Sch\u00f6nflies Motions Table 4.2 Structural parametersa of parallel mechanisms in Figs. 4.3, 4.4, 4.5 No. Structural parameter Solution Figure 4.3 Figures 4.4 and 4.5 1. m 14 17 2. p1 4 4 3. pi (i = 2, 3, 4) 4 5 4. p 16 19 5. q 3 3 6. k1 4 4 7", + " SGi (i = 2, 3, 4) 4 5 15. rGi (i = 1,\u2026,4) 0 0 16. MG1 4 4 17. MGi (i = 2, 3, 4) 4 5 18. (RF) (v1; v2; v3;xa) (v1; v2; v3;xa) 19. SF 4 4 20. rl 0 0 21. rF 12 15 22. MF 4 4 23. NF 6 3 24. TF 0 0 25. Pp1 j\u00bc1 fj 4 4 26. Pp2 j\u00bc1 fj 4 5 27. Pp3 j\u00bc1 fj 4 5 28. Pp4 j\u00bc1 fj 4 5 29. Pp j\u00bc1 fj 16 19 a See footnote of Table 2.2 for the nomenclature of structural parameters 4.1 Topologies with Simple Limbs 377 Table 4.3 Structural parametersa of parallel mechanisms in Figs. 4.6 and 4.7 No. Structural parameter Solution Figure 4.6a Figures 4.6b and 4.7 1. m 11 13 2. p1 4 3 3. pi (i = 2, 3, 4) 3 4 4. p 13 15 5. q 3 3 6. k1 4 4 7. k2 0 0 8. k 4 4 9. (RG1) (v1; v2; v3;xa) (v1; v2; v3;xa) 10. (RG2) (v1; v2; v3;xa) (v1; v2; v3;xa;xb) 11. (RG3) (v1; v2; v3;xa) (v1; v2; v3;xa;xd) 12. (RG4) (v1; v2; v3;xa) (v1; v2; v3;xa;xd) 13. SG1 4 4 14. SGi (i = 2, 3, 4) 4 5 15. rGi (i = 1,\u2026,4) 0 0 16. MG1 4 4 17. MGi (i = 2,3,4) 4 5 18. (RF) (v1; v2; v3;xa) (v1; v2; v3;xa) 19. SF 4 4 20. rl 0 0 21. rF 12 15 22. MF 4 4 23. NF 6 3 24. TF 0 0 25" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_35_0001380_b978-0-12-818983-2.00004-9-Figure4.4-1.png", + "original_path": "designv11-35/openalex_figure/designv11_35_0001380_b978-0-12-818983-2.00004-9-Figure4.4-1.png", + "caption": "Figure 4.4 Load displacement curves of untreated and alkali-treated jute composites: (A) vinylester resin, 23 wt.% fiber; (B) untreated; (C) 4 h NaOH-treated, 35 wt.% fiber; (D) untreated; (E) 4 h NaOH-treated [13].", + "texts": [ + " Similar to silane treatment, alkali (NaOH) treatment improves the crystallinity in the jute fibers and increases its modulus. The strength of the fibers improves while elongation to break reduces after treatment (Table 4.10). The flexural strength of the alkali-treated jute fiber-reinforced vinylester composites improves from 199.1 to 238.9 MPa, the modulus improves from 11.89 to 14.69 GPa, and the interlaminar shear strength increases from 0.238 to 0.2834 MPa. The load displacement curves of the composites containing untreated and alkali-treated jute fibers are shown in Fig. 4.4 [13]. Jute composites based on acetylated jute fibers possess improved mechanical properties, the degree of improvement depending on the jute fiber content and the duration of the treatment [14]. The tensile and flexural strengths of composites consisting of untreated and acetylated jute fibers are shown in Fig. 4.5 [14]. Several studies have been conducted to study the effectiveness of maleic anhydride polypropylene (MAH PP) copolymer as a coupling agent. For instance Mieck et al. [15] have reported an increase in shear and tensile strengths of about 100% and 25%, respectively, for flax PP composites, when the coupling agent was applied to the flax fibers before the composite was processed" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_35_0000174_j.matpr.2021.01.640-Figure17-1.png", + "original_path": "designv11-35/openalex_figure/designv11_35_0000174_j.matpr.2021.01.640-Figure17-1.png", + "caption": "Fig. 17. Mode shapes 5.", + "texts": [ + " The shaft rotates on its natural frequency, the mode shapes obtained their speeds. In this analysis the mode shape level 4 investigated and is shown in Fig. 16. In this step the natural frequency varies due to rotations. The results are plotted. While analysis the frequency var- ies 399.43 Hz, the mode shape level is 4. The review results are obtained by general post processor. The mode shapes can be determined by natural frequencies in this analysis the frequency varies 490.63 Hz, the mode shape level is 5 and its shown in Fig. 17. Fig. 18 shows that the mode shapes can be changed with regular frequencies. It shows that natural frequency gradually increases from first mode shape to second and continuously increases up to fifth mode shape and vibration of the system eliminated efficiently by suitable element design. Through assessment of the assortment of the repeat with their modes it was contemplated that the Aluminum \u2013 Titanium-Vanadium composite material shows the implausible material properties for the arrangement of composite drive shaft" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_35_0001419_978-94-007-7401-8-Figure3.20-1.png", + "original_path": "designv11-35/openalex_figure/designv11_35_0001419_978-94-007-7401-8-Figure3.20-1.png", + "caption": "Fig. 3.20 4RPPP-type overactuated PMs with coupled Sch\u00f6nflies motions defined by MF \u00bc SF \u00bc 4; RF\u00f0 \u00de \u00bc v1; v2; v3;xd\u00f0 \u00de, TF = 0, NF = 6, limb topology R\\P\\kP\\\\P (a) and R||P\\P\\\\P (b)", + "texts": [ + " 4PPPR (Fig. 3.16a) P\\P\\\\P||R (Fig. 3.1x) Idem No. 9 26. 4PPRP (Fig. 3.16b) P\\P||R\\P (Fig. 3.1y) The revolute joints of the four limbs have parallel axes 27. 4PPRP (Fig. 3.17a) P\\P\\\\R||P (Fig. 3.1z) Idem No. 26 28. 4PPRP (Fig. 3.17b) P\\P\\kR\\P (Fig. 3.1a0) Idem No. 26 29. 4PRPP (Fig. 3.18a) P\\R||P\\P (Fig. 3.1b0) Idem No. 26 30. 4PRPP (Fig. 3.18b) P\\R\\P\\kP (Fig. 3.1c0) Idem No. 26 31. 4PRPP (Fig. 3.19a) P||R\\P\\\\P (Fig. 3.1d0) Idem No. 26 32. 4RPPP (Fig. 3.19b) R\\P\\\\P\\\\P (Fig. 3.1e0) Idem No. 26 33. 4RPPP (Fig. 3.20a) R\\P\\kP\\\\P (Fig. 3.1f0) Idem No. 26 34. 4RPPP (Fig. 3.20b) R||P\\P\\\\P (Fig. 3.1g0) Idem No. 26 3.1 Topologies with Simple Limbs 243 Table 3.2 Limb topology and connecting conditions of the overactuated solutions with no idle mobilities presented in Figs. 3.21, 3.22, 3.23, 3.24, 3.25, 3.26, 3.27, 3.28, 3.29, 3.30, 3.31, 3.32, 3.33, 3.34, 3.35, 3.36, 3.37, 3.38, 3.39, 3.40, 3.41, 3.42, 3.43, 3.44, 3.45, 3.46, 3.47, 3.48, 3.49 No. PM type Limb topology Connecting conditions 1. 4RRC (Fig. 3.21a) R||R||C (Fig. 3.2a) The cylindrical joints of the four limbs have parallel axes 2" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_35_0001419_978-94-007-7401-8-Figure3.14-1.png", + "original_path": "designv11-35/openalex_figure/designv11_35_0001419_978-94-007-7401-8-Figure3.14-1.png", + "caption": "Fig. 3.14 Overactuated PMs with coupled Sch\u00f6nflies motions of types 4RPPR (a) and 4RPRP (b) defined by MF \u00bc SF \u00bc 4; RF\u00f0 \u00de \u00bc v1; v2; v3;xd\u00f0 \u00de, TF = 0, NF = 6, limb topology R\\P\\kP||R (a) and R\\P\\kR||P (b)", + "texts": [ + "10b) R\\P\\kP||R (Fig. 3.1m) Idem No. 2 15. 4PRRP (Fig. 3.11a) P\\R||R||P (Fig. 3.1n) The last prismatic joints of the four limbs have parallel directions 16. 4PRPR (Fig. 3.11b) P\\R||P||R (Fig. 3.1o) Idem No. 9 17. 4PRRP (Fig. 3.12a) P||R||R\\P (Fig. 3.1v) The first prismatic joints of the four limbs have parallel directions 18. 4RPRP (Fig. 3.12b) R||P||R\\P (Fig. 3.1q) Idem No. 2 19. 4PRPR (Fig. 3.13a) P||R\\P\\kR (Fig. 3.1r) Idem No. 1 20. 4RPPR (Fig. 3.13b) R||P\\P\\kR (Fig. 3.1s) Idem No. 2 21. 4RPPR (Fig. 3.14a) R\\P\\kP||R (Fig. 3.1t) Idem No. 2 22. 4RPRP (Fig. 3.14b) R\\P\\kR||P (Fig. 3.1u) Idem No. 2 23. 4PPPR (Fig. 3.15a) P\\P\\\\P\\kR (Fig. 3.1h0) Idem No. 9 24. 4PPPR (Fig. 3.15b) P\\P\\\\P\\\\R (Fig. 3.1w) Idem No. 9 25. 4PPPR (Fig. 3.16a) P\\P\\\\P||R (Fig. 3.1x) Idem No. 9 26. 4PPRP (Fig. 3.16b) P\\P||R\\P (Fig. 3.1y) The revolute joints of the four limbs have parallel axes 27. 4PPRP (Fig. 3.17a) P\\P\\\\R||P (Fig. 3.1z) Idem No. 26 28. 4PPRP (Fig. 3.17b) P\\P\\kR\\P (Fig. 3.1a0) Idem No. 26 29. 4PRPP (Fig. 3.18a) P\\R||P\\P (Fig. 3.1b0) Idem No. 26 30. 4PRPP (Fig. 3.18b) P\\R\\P\\kP (Fig" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_35_0000686_j.mechmachtheory.2021.104433-Figure20-1.png", + "original_path": "designv11-35/openalex_figure/designv11_35_0000686_j.mechmachtheory.2021.104433-Figure20-1.png", + "caption": "Fig. 20. Metamorphic epicyclic gear-rack train based on force constraint.", + "texts": [], + "surrounding_texts": [ + "By combining geometrical constraint and force constraint, a set of metamorphic epicyclic gear trains can be obtained based on a combination constraint. As shown in Figs. Figures 21-23, the combination constraint can provide both geometrical limitation and constraint forces. The metamorphic epicyclic gear-rack train in Fig. 23 is used to demonstrate the configuration transformation. In the rotation configuration, the driving force drives the planet rack overcoming the constraint force, and the spring in combination constraint is compressed or released. The angular velocity \u03c9r2 is not equal to \u03c9r1 the corresponding motion branch has mobility 2 as discussed in Fig. 14(b). When the driving force is equal to the constraint force or the planet rack is locked by geometrical limitation, the metamorphic epicyclic gear-rack train transforms to revolution configuration. The planet rack is rigidified with the sun rack-gear, and the angular velocity \u03c9r2 is equal to \u03c9r1, which means the corresponding motion branch has degenerated from mobility 2 to mobility 1. In practical design, we can adjust the constraint force to determine the constraint condition and control the configuration transformation. By inserting three types of constraints including geometrical constraint, force constraint, and combination constraint separately on the mechanism, one mobility is restricted that the metamorphic epicyclic gear train and initiates one active mobility. In practical design, we use external forces, such as gravity, magnetic force, constraint force from the target object, to overcome the inner constraint to release the mobility. These proposed metamorphic mechanisms usually operate with one active mobility and change their configuration and state of motion when the intervention of external force. With two mobility states, when the metamorphosis is initiated, either geometrical constraint or force constraint is utilized to restrict one mobility. An example describing the way of controlling the transformation is demonstrated in the following section." + ] + }, + { + "image_filename": "designv11_35_0001419_978-94-007-7401-8-Figure3.9-1.png", + "original_path": "designv11-35/openalex_figure/designv11_35_0001419_978-94-007-7401-8-Figure3.9-1.png", + "caption": "Fig. 3.9 Overactuated PMs with coupled Sch\u00f6nflies motions of types 4RRPP (a) and 4RPRP (b) defined by MF \u00bc SF \u00bc 4; RF\u00f0 \u00de \u00bc v1; v2; v3;xd\u00f0 \u00de, TF = 0, NF = 6, limb topology R||R||P\\P (a) and R||P||R\\P (b)", + "texts": [ + "1b) The first revolute joints of the four limbs have parallel axes 3. 4RPRR (Fig. 3.5a) R||P||R||R (Fig. 3.1c) Idem No. 2 4. 4RRPR (Fig. 3.5b) R||R||P||R (Fig. 3.1d) Idem No. 2 5. 4RRPR (Fig. 3.6a) R||R||P||R (Fig. 3.1e) Idem No. 2 6. 4PPRR (Fig. 3.6b) P\\P\\kR||R (Fig. 3.1i) Idem No. 1 7. 4RRRP (Fig. 3.7a) R||R||R||P (Fig. 3.1f) Idem No. 2 8. 4RPRP (Fig. 3.7b) R\\P\\kR||P (Fig. 3.1g) Idem No. 2 9. 4PPRR (Fig. 3.8a) P\\P||R||R (Fig. 3.1h) The last revolute joints of the four limbs have parallel axes 10. 4PRPR (Fig. 3.8b) P\\R||P||R (Fig. 3.1p) Idem No. 9 11. 4RRPP (Fig. 3.9a) R||R||P\\P (Fig. 3.1j) Idem No. 2 12. 4RPRP (Fig. 3.9b) R||P||R\\P (Fig. 3.1k) Idem No. 2 13. 4RPPR (Fig. 3.10a) R||P\\P\\kR (Fig. 3.1l) Idem No. 2 14. 4RPPR (Fig. 3.10b) R\\P\\kP||R (Fig. 3.1m) Idem No. 2 15. 4PRRP (Fig. 3.11a) P\\R||R||P (Fig. 3.1n) The last prismatic joints of the four limbs have parallel directions 16. 4PRPR (Fig. 3.11b) P\\R||P||R (Fig. 3.1o) Idem No. 9 17. 4PRRP (Fig. 3.12a) P||R||R\\P (Fig. 3.1v) The first prismatic joints of the four limbs have parallel directions 18. 4RPRP (Fig. 3.12b) R||P||R\\P (Fig. 3.1q) Idem No. 2 19. 4PRPR (Fig" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_35_0001563_9781119682684.ch1-Figure1.26-1.png", + "original_path": "designv11-35/openalex_figure/designv11_35_0001563_9781119682684.ch1-Figure1.26-1.png", + "caption": "Figure 1.26 Cross-sectional view of different machines.", + "texts": [ + " The stator carries AC armature winding (three-phase distributed) and the rotor has DC field winding. The rotor is supplied by an external DC source via the slip ring and brush arrangement. The rotor can have two shapes; cylindrical and salient pole. Cylindrical rotor machines are used for high-speed (thermal and nuclear power plants) and salient pole is used for low-speed (hydroelectric power plant) applications. The cross-sectional view of the DC, induction and synchronous machines are given in Figure 1.26. The stator and rotor structures are clearly seen [5]. If the rotor winding is replaced by Permanent Magnets, the machine is called a Permanent Magnet Synchronous machine (PMSM) or sinusoidal PMSM. Use of PM offers several advantages, for example, no rotor losses (hence higher efficiency), faster response time as the electromechanical time constant is reduced. The PM material used for the rotor are; ferrite, aluminium-nickel-cobalt (Al-Ni-Co), samarium-cobalt (Sm-Co) and neodymium-iron-born (Nd-Fe-B)" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_35_0000403_s10527-020-10002-w-Figure4-1.png", + "original_path": "designv11-35/openalex_figure/designv11_35_0000403_s10527-020-10002-w-Figure4-1.png", + "caption": "Fig. 4. Exploded view of MFD (explanation in text).", + "texts": [ + " For ease of working with the data processing system, the PC runs a program with a user interface. The program displays the data in real time and stores them in the PC memory. Having selected the sensor coordinate system as the basal coordinate system, the forces and torques applied to the instrument are determined by solving the static prob lem under the conditions of equilibrium of the system of forces in space: However, the obtained values for the resultant vector F include device component positioning errors. Figure 4 shows the MFD in exploded view. Errors will have two components. The first consists of errors in the attachment of adapter flange 2 to multicomponent sensor 1 and the second consists of play (backlash) in the mount between surgical instrument 3 and adapter flange 2. The resultant load is computed as: where F is the force applied to the point of contact, equal to the modulus of the vector of forces measured by the sensor; \u0394F = f(\u03b1device, \u03b2device, \u03b3device, \u03b1play, \u03b2play, \u03b3play) is the error in the force arising from the three maximal errors in flange fitting angles and the three maximum permissible angles associated with play in the attachment of the surgi cal instrument to the probe" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_35_0001393_ijhm.2020.109916-Figure3-1.png", + "original_path": "designv11-35/openalex_figure/designv11_35_0001393_ijhm.2020.109916-Figure3-1.png", + "caption": "Figure 3 Meshed view of spur gear (see online version for colours)", + "texts": [ + " 3 Modal analysis method is used to calculate the suitable material for the design of spur gears. The geometrical model of spur gear was constructed using SOLIDWORKS software for the designing process (Figures 1 and 2). After creating the model in SOLIDWORKS software, the design is imported into ANSYS workbench software for further analysis of the key factors such as stresses, deflections and modal analysis of the spur gear. The dimensions and the material properties of the spur gear are given Tables 1 and 2. Figure 3 shows the meshed view of the spur gear. Accuracy and efficiency of finite element method depends upon the meshing size of the model. A very fine mesh is used to design the model so that the results would be more accurate. In the spur gear, the shear stress (\u03c4), equivalent stress (\u03c3v) and maximum shear stress (\u03c3m) directly depends on the amount of load and type of load acting on the spur gear. Performance and life span depend upon stress and deformation. To get the analysis results, the SOLIDWORKS design model is imported to ANSYS Workbench and boundary conditions are applied (Figure 4)" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_35_0000174_j.matpr.2021.01.640-Figure10-1.png", + "original_path": "designv11-35/openalex_figure/designv11_35_0000174_j.matpr.2021.01.640-Figure10-1.png", + "caption": "Fig. 10. Elongation analysis on Metal Matrix Composite.", + "texts": [ + " The drive shaft rotations are 650,0, 2.063 rad\\s. The locations are 795.07, 7.3163 mm are shown in Fig. 8. The moment load does not cause rotation, the Moment is applied on the drive shaft is 350 N.m as shown in Fig. 9 in order to omit more bending stress. When composite drive shaft breaks, it is divided into fine fibers that do not have any danger for the driver. They have a less specific modulus and less damping capacity. So that conventional drive shaft is replaced with composite materials Fig. 10. Selection of materials are an important parameter for design of any machine element. Our objective is to select suitable material than conventional steel material and preferred metal matrix composite materials consist of Aluminium, Titanium, Vanadium alloy. Material Properties of Metal Matrix Composite Ti-Al-V alloy are listed below in the table 1. Steel (SM45C) The conventional two piece steel drive shaft is made in two sections connected by a support structure, bearings and U-joints hence overall weight of assembly will be more" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_35_0000512_s42417-021-00289-8-Figure4-1.png", + "original_path": "designv11-35/openalex_figure/designv11_35_0000512_s42417-021-00289-8-Figure4-1.png", + "caption": "Fig. 4 Geometric relationship between ball and cage pocket", + "texts": [ + " (12) \u23a7 \u23aa\u23aa\u23a8\u23aa\u23aa\u23a9 i b_r = nb\ufffd i=1 \ufffd T\u22121 i_a T\u22121 a_p \ufffd \u222b d p xc_r \ufffd\ufffd r b_r = nb\ufffd i=1 \ufffd Ti_rT \u22121 i_a T\u22121 a_c \ufffd \u222b d p xc_r \ufffd\ufffd , Table 1 System parameters of ball bearing\u2013rotor system Item Parameter Value Ball bearing Number of balls 9 Contact angle 25\u00b0 Pitch diameter 20.45\u00a0mm Ball diameter 4.74\u00a0mm Inner race clearance 0.001\u00a0mm Outer race clearance 0.002\u00a0mm Inner diameter of cage 21\u00a0mm Outer diameter of cage 17\u00a0mm Pocket clearance 0.2\u00a0mm Rotor L 177.5\u00a0mm D 10\u00a0mm Outer diameter 72\u00a0mm Length of disk 25.7\u00a0mm Rotor density 7950 kg\u00a0 m\u22123 Modulus of elasticity 211 GPa Disk eccentricity 32\u00a0\u03bcm 1 3 The relative geometric position between the ball and cage pocket is shown in Fig.\u00a04. In the pocket coordinate system, the position vector of the ball center relative to pocket center is expressed as Then, the minimum clearance between ball and pocket wall is shown as (13) d b_d = Tc_d ( Ti_c i b_c \u2212 c d_c ) . where Bd is the width of pocket. In the contact (between ball and pocket) coordinate system, the position vector of cage center is The relative velocity vector between ball and cage is Taking the comprehensive roughness \u03c3bd between the ball and pocket surface as a reference, when hbd\u00a0\u2265\u00a0\u03c3bd, there is (14)hbd = 0" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_35_0001419_978-94-007-7401-8-Figure5.41-1.png", + "original_path": "designv11-35/openalex_figure/designv11_35_0001419_978-94-007-7401-8-Figure5.41-1.png", + "caption": "Fig. 5.41 3PaPPaR-1RPaPatP-type fully-parallel PMs with uncoupled Sch\u00f6nflies motions defined by MF = SF = 4, RF\u00f0 \u00de \u00bc \u00f0v1; v2; v3;xb\u00de, TF = 0, NF = 30 limb topology R||Pa||Pat||P and Pa||P\\Pa\\kR, Pa||P\\Pa||R (a), Pa\\P\\\\Pa\\kR, Pa\\P\\\\Pa||R (b)", + "texts": [ + "4c) R||Pa||Pa||P (Fig. 5.4k) 3. 3PaPPaR-1RPPaPa (Fig. 5.40a) Pa||P\\Pa\\kR (Fig. 5.4a) The revolute joint between links 7D and 9 of G4 limb and the last revolute joints of limbs G1, G2 and G3 have superposed axes Pa||P\\Pa||R (Fig. 5.4d) R||P||Pa||Pa (Fig. 5.4m) (continued) 5.1 Fully-Parallel Topologies 531 Table 5.4 (continued) No. PM type Limb topology Connecting conditions 4. 3PaPPaR-1RPPaPa (Fig. 5.40b) Pa\\P\\\\Pa\\kR (Fig. 5.4b) Idem No. 3 Pa\\P\\\\Pa||R (Fig. 5.4c) R||P||Pa||Pa (Fig. 5.4m) 5. 3PaPPaR1RPaPatP (Fig. 5.41a) Pa||P\\Pa\\kR (Fig. 5.4a) Idem No. 1 Pa||P\\Pa||R (Fig. 5.4d) R||Pa||Pat||P (Fig. 5.4l) 6. 3PaPPaR1RPaPatP (Fig. 5.41b) Pa\\P\\\\Pa\\kR (Fig. 5.4b) Idem No. 1 Pa\\P\\\\Pa||R (Fig. 5.4c) R||Pa||Pat||P (Fig. 5.4l) 7. 3PaPPaR-1RPPaPat (Fig. 5.42a) Pa||P\\Pa\\kR (Fig. 5.4a) Idem No. 3 Pa||P\\Pa||R (Fig. 5.4d) R||P||Pa||Pat (Fig. 5.4n) 8. 3PaPPaR-1RPPaPat (Fig. 5.42b) Pa\\P\\\\Pa\\kR (Fig. 5.4b) Idem No. 3 Pa\\P\\\\Pa||R (Fig. 5.4c) R||P||Pa||Pat (Fig. 5.4n) 9. 3PaPaPR-1RPaPaP (Fig. 5.43) Pa\\Pa\\kP\\\\R (Fig. 5.4f) Idem No. 1 Pa\\Pa\\kP\\kR (Fig. 5.4h) R||Pa||Pa||P (Fig. 5.4k) 10. 3PaPaPR-1RPaPaP (Fig. 5.44) Pa\\Pa\\\\P\\\\R (Fig. 5.4g) Idem No" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_35_0001596_isscc.1959.1157072-Figure5-1.png", + "original_path": "designv11-35/openalex_figure/designv11_35_0001596_isscc.1959.1157072-Figure5-1.png", + "caption": "Figure 5-Two-step, transfer-clear operation will transfer information state of A to B, and prepare A for receiving new information. Assume solid arrows indicate ZERO (or CLEARED) state; dotted arrows indicate ONE (or SET)", + "texts": [ + " Observe, in Figures 1 and 3, that when looking into an array in either the CLEARED or RESET states, exactly the same conditions are seen from either set of terminals. Further, in both the SET .and the RESET states, essentially identical situations are seen from bb. From this we see that, for current entering the bottom terminal, hb sees a high threshold in the CLEARED state, and low thresholds in the SET and RESET states. The information state of one of these arrays can be transferred to a second array, which must be in the CLEARED state, by applying a current to the coupling loop (Figure 5 ) . The applied current can be divided into two equal, diverginq components, I t , which satisfies Kirchhoff's current law. A loop current, I,, can be superposed, without affecting current-law observarxe, to satisfy Rirchhofs voltage law about the loop. If A were CLEARED (ZERO state), I, ideally would equal zero, since conditions lo,oking both ways from the driven nodes would appear identical. Consequently, current I t going into each of these CLEARED-state arrays causes no switching in either (sub-threshold; see Figure 4) . O n the other hand, if A were RESET (ONE state), and we looked left and right from the driven nodes, we would see ,conditions associated respectively with paths ( 4 ) and (2) of Figure 4. In this case, I,, would tend to grow toward a value equnl to the difference between the large and the small tl;resholds (Figure 4 ) , but would settle at some lesser value which causes switching (Figure 5) of cores ZA and YB (and %E?) to proceed at equal rates as array B is forced to the SET state. I t can be shown that neither during the above transfer-AB .operation nor during the subsequent clear-A operation does a disturbance propagate out of this two-array system which is serious enough to affect the information states of adjoining arrays. No extra decoupling devices are thus needed, and it is found that extension of this transfer-clear procedure to a chain of arrays yields a serial shift register (Figure 6) " + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_35_0000718_j.scitotenv.2021.149314-Figure1-1.png", + "original_path": "designv11-35/openalex_figure/designv11_35_0000718_j.scitotenv.2021.149314-Figure1-1.png", + "caption": "Fig. 1. The schematic of the PSCP and PBCP. (a) the structure of PSCP, (b) top view of nozzles cross-section, (c) the structure of PBCP, (d) and (e) coordinate system of PSCP. Notes: 1. gas outlet, 2. microalgal solution outlet, 3. microalgal solution, 4. the circulating pump, 5. nozzles, 6. CO2 bubbles, 7. porous gas aerators, 8. inlet tubes of pump, 9. outlet tubes of pump.", + "texts": [ + " In the previous work, the author proposed a single column TSCP with high radial velocity, goodmixing andmass transfer ofmicroalgal solution. The microalgal solution was pumped from the top of the column, and then injected into the column through four nozzles in lower part of the column, thus forming a spiral-up flow field to promote the light/dark cycle of microalgal cells (Ye et al., 2019). In order tomeet the needs of industrial application for microalgal biomass yield and CO2 fixation rate, eight TSCPs are designed in parallel to form PSCP. The structure of PSCP is shown in Fig. 1 (a). The total height of PSCP is 1.6 m and the total width is 1.7 m. It is composed of 8 vertical transparent hollow PET columns, 32 stainless steel nozzles, 8 gas aerators, upper and bottom headers, the circulating pump, circulating pipes and other connecting parts. The diameter and the length of the top and bottom header are 0.1 m and 1.7 m, respectively. The diameter of the vertical column is 0.1 m, and the distance between adjacent columns is also 0.1 m. Some MSOs were set in the upper of partial columns, which were connected to the inlet of the circulating pump", + " The number of MSO is set as 2,4,6, and 8. A maximum of one MSO can be set on a single column. The MSOs distribution in PSCP with different number of MSOs was shown in the Fig. A.1 in the Appendices. The height ofMSO is set as 0.65m, 0.85m, 1.05m, and 1.15 m. The outlet of the circulating pump is connected to four nozzles at the lower part of each column through pipes, and the height of the nozzles cross-section is 0.25 m. Each column is equipped with four nozzles. The arrangement of nozzles is shown in Fig. 1 (b). Four nozzles are symmetrically arranged on the horizontal cross-section of each column. The axes of these nozzles do not coincide with the cross-section diameter directly, but form an angle of 10\u00b0 with the cross-section diameter, so that the design streamline of microalgal solution is tangent to the central imaginary circle. Gas aerators are arranged at the bottom center of each column to introduce CO2 bubbles into the microalgal solution. CO2 aeration rate is controlled by rotameter. The tangential flowon the nozzle cross-section coupled with the upward flow driven by rising CO2 bubbles to forma spiral-upflowfield. The PSCPwith a total volumeof about 96 L is supported by a stainless steel bracket. Parallel bubble column photobioreactor (PBCP) has the same structure, size and material as PSCP (Fig. 1 (c)), but without circulating pump, hose and nozzles. The total volume of PBCP is also about 96 L. Creo parametric 6.0 is used to build the three-dimensional model of PSCP, and ANSYS 2020 is used for meshing and CFD calculation. Considering the calculation time, after the grid independence test, the mesh with 117,000 cells was used for study. In order to analyze theflowfield ofmicroalgal solution containing CO2 bubbles, the Eulerian two-phasemodelwas used. Themicroalgal solution (liquid) was the main phase, and its density was set as 1000 kg/m3" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_35_0001419_978-94-007-7401-8-Figure2.43-1.png", + "original_path": "designv11-35/openalex_figure/designv11_35_0001419_978-94-007-7401-8-Figure2.43-1.png", + "caption": "Fig. 2.43 4RPPaPa-type fully-parallel PM with coupled Sch\u00f6nflies motions defined by MF \u00bc SF \u00bc 4; RF\u00f0 \u00de \u00bc \u00f0v1; v2; v3;xd\u00de, TF = 0, NF = 30, limb topology R\\P\\kPa\\\\Pa", + "texts": [ + " 4PaPC (Fig. 2.38b) Pa||P\\C (Fig. 2.21z) Idem No. 24 27. 4PaCP (Fig. 2.39) Pa\\C\\P (Fig. 2.21z1) Idem No. 11 28. 4PPaPaR (Fig. 2.40) P\\Pa\\\\Pa\\kR (Fig. 2.22a) Idem No. 4 29. 4PPaRPa (Fig. 2.41) P\\Pa||R\\Pa (Fig. 2.22b) The axes of the revolute joints connecting the two parallelogram loops of the four limbs are parallel (continued) 2.2 Topologies with Complex Limbs 175 Table 2.3 (continued) No. PM type Limb topology Connecting conditions 30. 4PRPaPa (Fig. 2.42) P\\R||Pa\\Pa (Fig. 2.22c) Idem No. 5 31. 4RPPaPa (Fig. 2.43) R\\P\\kPa\\\\Pa (Fig. 2.22d) Idem No. 15 32. 4PaPPaR (Fig. 2.44) Pa\\P\\\\Pa\\\\R (Fig. 2.22e) Idem No. 4 33. 4PaPRPa (Fig. 2.45) Pa\\P\\kR\\Pa (Fig. 2.22f) The axes of the revolute joints connecting links 5 and 6 of the four limbs are parallel 34. 4PaRPPa (Fig. 2.46) Pa||R\\P\\\\Pa (Fig. 2.22g) The axes of the revolute joints connecting links 4 and 5 of the four limbs are parallel 35. 4RPaPPa (Fig. 2.47) R||Pa\\P\\\\Pa (Fig. 2.22h) Idem No. 15 36. 4PaPaPR (Fig. 2.48) Pa\\Pa\\\\P\\\\R (Fig. 2.22i) Idem No. 4 37. 4RPaPaR (Fig" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_35_0000172_012019-Figure2-1.png", + "original_path": "designv11-35/openalex_figure/designv11_35_0000172_012019-Figure2-1.png", + "caption": "Figure 2. Simplified illustration of an angular misalignment at cardan shaft.", + "texts": [ + " According to [14], when a torque transmission shaft has three or more bearing supports, for example when front and rear axles of a wheel loader are coupled together, there is a potential for misalignment, which can be parallel misalignment, meaning that one of the two shafts is displaced laterally, but still parallel to the other, or angular misalignment, where the axis of one is at an angle to that of the other. The angular misalignment of cardan shafts can introduce bending deflections to both input and output shafts while rotating with respect to the shafts. The excited bending modes of shafts show themselves then as harmonics of the shaft\u2019s rotational speed in the frequency spectrum. This is considered as the basis for cardan shaft angular misalignment detection using vibration signals. Figure 2 shows a typical cardan shaft with an angular misalignment. The angular misalignment can be studied mathematically using the equation (1), which shows the relative instantaneous angular velocity of output to input [17] as \u03c92 \u03c91 = cos \u03b1 1 \u2212 sin2 \u03b1 sin2 \u03b8i (1) where \u03c92 \u03c91 , \u03b1, and \u03b8i are output-to-input ratio of instantaneous angular velocity, the misalignment angle, and the instantaneous shaft rotating angle (depending on the vehicle steering condition), respectively. For small misalignment angles, say \u03b1 up to about 10\u00ba, cos \u03b1 and sin2 \u03b1 can be considered as 1 and \u03b12, respectively" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_35_0000555_j.engfailanal.2021.105525-Figure1-1.png", + "original_path": "designv11-35/openalex_figure/designv11_35_0000555_j.engfailanal.2021.105525-Figure1-1.png", + "caption": "Fig. 1. Schematic view of the involute spur gear drive.", + "texts": [ + " In the text, the influences of defects on the lubrication of gear meshing, as well as the changes in normal stiffness and normal contact height with the meshing path, are discussed. A detailed comparison and analysis of the influence of defect sizes and depths on gear contact and stiffness are analyzed, and the impacts of the pit X. Pei et al. Engineering Failure Analysis 127 (2021) 105525 under different speeds are also investigated. The contact curvature radius, velocity, and load force change constantly during the meshing of a single pair of tooth surfaces. A schematic diagram of the transmission is shown in Fig. 1. For involute gear pairs, the geometric parameters and speed parameters of the meshing region can be obtained according to the following formula [25]. \u23a7 \u23aa \u23aa \u23aa \u23aa \u23aa \u23aa\u23a8 \u23aa \u23aa \u23aa \u23aa \u23aa \u23aa\u23a9 gyc = \u2213 1 2 d1Sin\u03b1 \u00b1 \u0305\u0305\u0305\u0305\u0305\u0305\u0305\u0305\u0305\u0305\u0305\u0305\u0305\u0305\u0305\u0305\u0305\u0305\u0305\u0305\u0305\u0305\u0305\u0305\u0305\u0305\u0305\u0305\u0305\u0305\u0305\u0305\u0305\u0305\u0305\u0305\u0305\u0305\u0305\u0305\u0305\u0305\u0305\u0305\u0305\u0305\u0305\u0305\u0305\u0305\u0305\u0305\u0305\u0305\u0305\u0305\u0305 ( 1 2 d1Sin\u03b1 )2 \u2212 ( 1 2 d1 )2 + rc 2 \u221a r1 = 1 2 d1Sin\u03b1 \u00b1 gyc r2 = 1 2 d1Sin\u03b1 \u2213 gyc u1,2 = \u03c91,2 \u00d7 r1,2 (1) The subscripts 1 and 2 represent the driving gear and the driven gear. d denotes the pitch circle diameter and it is a constant when the center distance of the two gears is fixed" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_35_0000395_978-3-030-67750-3_16-Figure11.6-1.png", + "original_path": "designv11-35/openalex_figure/designv11_35_0000395_978-3-030-67750-3_16-Figure11.6-1.png", + "caption": "Fig. 11.6 Proposed design of material trolley", + "texts": [ + " So, the new material\u2019s trolley has been designed and fabricated to ensure comfort for workers. The current design of the material\u2019s trolley was designed based on four multi-direction wheels. That is the reason that the current design is heavy once it is loaded by the semi-finished good part. Furthermore, it will affect the body of workers and at the same time reduce the productivity of the production process. The newly designedmaterial\u2019s trolley applied the \u201cPipe& Joint\u201d concept to ensure more comfort for the transferring process as shown in Fig. 11.6. Furthermore, the cost of this newly designedmaterial\u2019s trolley is cheaper and it is easilymaintained because it just used a single Allen-key size for the maintenance process (shown in Fig. 11.7). Moreover, this new design of this material\u2019s trolley is lighter. This pipe and joint system is considered as a smart system, dynamic, and flexible modular assembly system consisting of plastic-coated steel pipes and metal joints. The flexibility of this system to tailor with several requirements such as racks, trolleys, workstations, gravity flow racks, and even light-duty machine structures was considered during the selection of this system" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_35_0000132_lra.2021.3062583-Figure4-1.png", + "original_path": "designv11-35/openalex_figure/designv11_35_0000132_lra.2021.3062583-Figure4-1.png", + "caption": "Fig. 4. (a) 3D model of the robot wrist applied with cable pretension FT . (b) 3D schematic showing the wrist being subject to the external load F that can be resolved into Fz and FL in the reference system {O}. (c) Exaggerated illustration of the small angle deflection of the wrist due to the lateral force FL in the bending plane.", + "texts": [ + " The actuation cables were passed through the channels and terminated at the DF. The silicone sheath was stretched and slid onto the wrist before the needle was secured on the needle hub of the DF using glue. The wristed percutaneous robot must be sufficiently stiff in its straight configuration to puncture the various layers of biological tissues to enter the pericardial space. The stiffness of the robot (wrist) could be influenced by the temperature of the SMA spring, T , and the pretension of the actuation cables, FT , as shown in Fig. 4(a). Higher T and larger FT would increase the robot stiffness. A stiffness model is therefore desired to understand theoretically the capability of the robot to resist lateral deflection during the robot insertion process. It helps to determine the optimal T , FT , and other geometric parameters of the SMA spring in the robot that allows smooth percutaneous operation, especially across soft tissue layers with different mechanical properties. In this preliminary work, we intend to verify the model accuracy and determine the stiffness required for gelatin phantoms of two stiffnesses", + " While bending stiffness is the focus of the work, spring compression must be taken into consideration in the stiffness model. The spring with a mean radius r, spring wire diameter d, and n number of coils has a free (uncompressed) length of ls. The compressive, flexural, and shearing rigidities of the spring in its free state can be expressed as [29]: ks = Gd4ls 64r3n ; \u03b2s = 2lsEIG n\u03c0r(2G+ E) ; \u03b3s = lsEI \u03c0nr3 (1) whereE,G, and I are the Young\u2019s modulus, shear modulus, and second moment of area of the spring wire, respectively. When the wrist is pretensioned with four actuation cables of each FT (see Fig. 4(a)) and an arbitrary force F is applied at the central point of the distal end of the wrist (see Fig. 4(b)), the total axial load FA can be written as: FA = Fz + 4FT (2) where Fz is the component of F in the axial direction. The compressed spring length l becomes l = ls \u2212 FA ks . k, \u03b2 and \u03b3 that represent the three rigidities of the spring in the compressed state can be expressed as: k = k0 l ls ; \u03b2 = \u03b2s l ls ; \u03b3 = \u03b3s l ls (3) It should be noted that E of the SMA spring is a function of its martensite volume fraction, \u03be, which changes according to T . It Authorized licensed use limited to: San Francisco State Univ", + " When the SMA is below the martensite finish temperature of around 30 \u25e6C, \u03be = 1 and E = EM . When it exceeds the austenite finish temperature of around 45 \u25e6C, \u03be = 0 and E = EA. With all mechanical properties of spring in its compressed state known, the 2-D bending stiffness of the robot wrist considering the flexural and shear rigidities can be developed. The lateral force, FL, resolved from the external load F , can be further resolved into Fx and Fy as follows: Fx = FL cos\u03c3; Fy = FL sin\u03c3 (5) where \u03c3 is the angle of FL relative to the x-axis, as shown in Fig. 4(b). These lateral forces cause an internal moment and shear force at the distal end of the spring. The spring is then modeled as a cantilever beam with an equivalent flexural rigidity \u03b2 and equivalent shearing rigidity \u03b3 of the spring during the bending process. Based on the classical Euler-Bernoulli beam theory under the assumption of small-angle bending and considering both the internal moment and shear force, the deflection of the spring \u03b4s in the x axis and y axis can be expressed as [18], [29]:[ \u03b4xs \u03b4ys ] = [ 1 3 l 3+ 1 2 lrl 2 \u03b2 + (l+lr) \u03b3 0 0 1 3 l 3+ 1 2 lrl 2 \u03b2 + (l+lr) \u03b3 ][ Fx Fy ] (6) where lr is the length of the rigid part (see Fig. 4). \u03b4 is the sum of the spring deflection, \u03b4s and the rigid part deflection \u03b4r as a result of the bending slope of the spring \u03b8. Its two components in x axis and y axis can be expressed as:[ \u03b4x \u03b4y ] = [ 1 3Fxl 3+ 1 2Fxlrl 2 \u03b2 + Fx(l+lr) \u03b3 + lr sin \u03b8y 1 3Fyl 3+ 1 2Fylrl 2 \u03b2 + Fy(l+lr) \u03b3 + lr sin \u03b8x ] (7) where the spring bending angle about the y-axis, \u03b8y = 1 2Fxl 2+Fxlrl \u03b2 and the spring bending angle about the x-axis \u03b8x = 1 2Fyl 2+Fylrl \u03b2 . Under the axial load FA, there also exists a deflec- tion magnification, given approximately by C = 1 1\u2212FA/Fc [29], where Fc is the critical load before the spring buckles and is computed to be 30 N in this case" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_35_0001419_978-94-007-7401-8-Figure3.127-1.png", + "original_path": "designv11-35/openalex_figure/designv11_35_0001419_978-94-007-7401-8-Figure3.127-1.png", + "caption": "Fig. 3.127 4RPRPaR-type overactuated PM with coupled Sch\u00f6nflies motions defined by MF \u00bc SF \u00bc 4; RF\u00f0 \u00de \u00bc v1; v2; v3;xd\u00f0 \u00de, TF = 0, NF = 14, limb topology R||P||R\\Pa||R", + "texts": [ + " 3.121) R||P||R\\R||Pa (Fig. 3.52n) Idem No. 15 35. 4RPRRPa (Fig. 3.122) R||P||R\\R||Pa (Fig. 3.52o) Idem No. 15 36. 4RRPRPa (Fig. 3.123) R||R||P\\R||Pa (Fig. 3.52p) Idem No. 15 37. 4RRRPPa (Fig. 3.124) R||R\\R\\P\\kPa (Fig. 3.52q) Idem No. 15 (continued) 3.2 Topologies with Complex Limbs 361 Table 3.6 (continued) No. PM type Limb topology Connecting conditions 38. 4PRRPaR (Fig. 3.125) P||R||R\\Pa||R (Fig. 3.52r) Idem No. 13 39. 4RPRPaR (Fig. 3.126) R||P||R\\Pa||R (Fig. 3.52s) Idem No. 13 40. 4RPRPaR (Fig. 3.127) R||P||R\\Pa||R (Fig. 3.52t) Idem No. 13 41. 4RRPPaR (Fig. 3.128) R||R||P\\Pa||R (Fig. 3.52u) Idem No. 13 42. 4PPaRRR (Fig. 3.129a) P\\Pa\\kR||R||\\kR (Fig. 3.52v) Idem No. 13 43. 4PaPRRR (Fig. 3.129b) Pa\\P\\\\R||R\\\\R (Fig. 3.52w) Idem No. 13 44. 4PPaRRR (Fig. 3.130) P\\Pa\\\\R||R\\kR (Fig. 3.52x) Idem No. 13 45. 4PaPRRR (Fig. 3.131) Pa\\P||R||R\\kR (Fig. 3.52y) Idem No. 13 46. 4PaRPRR (Fig. 3.132) Pa\\R||P||R\\kR (Fig. 3.52z) Idem No. 13 47. 4PaRPRR (Fig. 3.133) Pa\\R||P||R\\kR (Fig. 3.52a0) Idem No. 13 48. 4PaRRPR (Fig" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_35_0001419_978-94-007-7401-8-Figure3.48-1.png", + "original_path": "designv11-35/openalex_figure/designv11_35_0001419_978-94-007-7401-8-Figure3.48-1.png", + "caption": "Fig. 3.48 Overactuated PMs with coupled Sch\u00f6nflies motions of types 4CRRP (a) and 4RCRP (b) defined by MF \u00bc SF \u00bc 4; RF\u00f0 \u00de \u00bc v1; v2; v3;xd\u00f0 \u00de, TF = 0, NF = 2, limb topology C||R\\R\\P (a) and R||C\\R\\P (b)", + "texts": [ + "3h0) Idem No. 17 49. 4RPRPR (Fig. 3.44b) P\\P\\kR||R\\\\R (Fig. 3.3i0) Idem No. 17 50. 4PPRRR (Fig. 3.44c) P\\P||R||R\\\\R (Fig. 3.3j0) Idem No. 17 51. 4CRRR (Fig. 3.45a) C||R\\R||R (Fig. 3.3k0) Idem No. 17 52. 4RCRR (Fig. 3.45b) R||C\\R||R (Fig. 3.3l0) Idem No. 17 53. 4RRCR (Fig. 3.46a) R||R\\C||R (Fig. 3.3m0) Idem No. 16 54. 4RRRC (Fig. 3.46b) R||R\\R||C (Fig. 3.3n0) Idem No. 16 55. 4RCRP (Fig. 3.47a) R\\C||R\\\\P (Fig. 3.3o0) Idem No. 16 56. 4RRCP (Fig. 3.47b) R\\R||C\\\\P (Fig. 3.3p0) Idem No. 16 57. 4CRRP (Fig. 3.48a) C||R\\R\\P (Fig. 3.3q0) Idem No. 44 58. 4RCRP (Fig. 3.48b) R||C\\R\\P (Fig. 3.3r0) Idem No. 44 59. 4CRPR (Fig. 3.49a) C||R\\P\\\\R (Fig. 3.3s0) Idem No. 17 60. 4PCRR (Fig. 3.49b) P\\C||R\\\\R (Fig. 3.3t0) Idem No. 17 3.1 Topologies with Simple Limbs 245 Table 3.3 Structural parametersa of parallel mechanisms in Figs. 3.4, 3.5, 3.6, 3.7, 3.8, 3.9, 3.10, 3.11, 3.12, 3.13, 3.14, 3.15, 3.16, 3.17, 3.18, 3.19, 3.20, 3.21, 3.22, 3.23, 3.24, 3.25, 3.26 No. Structural parameter Solution Figures 3.4, 3.5, 3.6, 3.7, 3.8, 3.9, 3.10, 3.11, 3.12, 3.13, 3.14, 3.15, 3.16, 3" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_35_0001579_2015-24-2526-Figure4-1.png", + "original_path": "designv11-35/openalex_figure/designv11_35_0001579_2015-24-2526-Figure4-1.png", + "caption": "Figure 4. Normal force distribution in a helical gear pair predicted by gear", + "texts": [ + " Available validations are focusing on gear rattle and gear whine comparison of transmission gears with experimental results. Specific validation for MBU or scissors gears are missing. Main difference for rattle is that we have loaded condition in case of the MBU (similar to a gear hammering, but we face rather low load in balancer drive gearing as long as no scissors gears are present), but this approach is general and therefore no limitations in application are valid. A so-called 2.5D sliced approach is used (see Figure 4, left picture). For a better performance, the resolution of gear contact is utilizing the special geometrical properties of involute shaped gears, rather than solving a general 3D-contact problem. However, the model reflects deviations from the ideal geometry, caused by shape modifications and manufacturing tolerances, as well as edge loading effects due to angular misalignments. In addition, a direct evaluation of root stresses is possible. A more simple, but for MBD rather typical, engagement-line approach, would not completely fulfill those requirements due to its known limitations", + " In order to meet the requirements regarding computational efficiency and having in mind that the user requests to analyze various load and speed cases, parameter and model variants the gear contact is resolved in the following steps: \u2022 Discretization into a series of slices along the gear's face width (\u201csliced approach\u201d) \u2022 Detection of contact by intersecting candidate flanks with Plane of Action \u2022 Determination of displacement field by considering local connection node positions and deviations from the ideal involute, enabling the investigation in shape modifications and correction, which are essential for gear profile optimization as well as investigation in manufacturing tolerances \u2022 Computation of normal contact forces (as shown by Figure 4) considering deflections induced by flank contact, tooth bending and tooth tilting in wheel body \u2022 Determination of damping forces (contact and backlash) and friction forces \u2022 Calculation of gear root stresses This model is capable to predict gear noise in respect to rattle and whine on a high accuracy level, which was validated by different comparison with measurements during numerical model development (see [12]) as well as durability investigation, based on analytical approach following ISO standard" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_35_0000067_jestpe.2021.3061120-Figure15-1.png", + "original_path": "designv11-35/openalex_figure/designv11_35_0000067_jestpe.2021.3061120-Figure15-1.png", + "caption": "Fig. 15. Flux distribution of different LSPMVMs under power angle of 70\u00b0. (a) Type I: V-shaped LSPMVM. (b) Type II: Consequent-pole U-shaped LSPMVM. (c) Type III: Dual-stator spoke LSPMVM. (d) Type IV: Consequent-pole SPM LSPMVM. (e) Type V: LSPMSM.", + "texts": [ + " Once the load is over the breakdown torque, the machine cannot keep in the steady state and the speeds are sharply decreased. Moreover, the power angles corresponding to the maximum torque are different, where the consequent-pole U-shaped LSPMVM has the minimum value of 60\u00b0, which means that the contribution of reluctance torque in which is maximum compared to other topologies. As shown in (8), the first item is PM torque components, and the second item represents the reluctance torque components due to the different reactance of d-axis and q-axis. 2 1 1 sin sin 2 2 e d q d mpEU mpU T X X X (8) Fig. 15 shows the flux distributions with power angle of 70\u00b0. It is worth noting that the maximum flux density points of all the topologies occur at the rotor teeth. And the Type II and Type IV suffer from the serious stator yoke saturation, which would restrict the overload capability. The steady state currents are given in Fig. 16, where the RMS of LSPMVMs are almost the same, and the regular LSPMSM has the higher current amplitude. The current THD of which are 3.3%, 5.0%, 1.0%, 10.0% and 65.1%, respectively, where the 3rd harmonic of LSPMSM is more serious than other topologies" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_35_0000142_012076-Figure1-1.png", + "original_path": "designv11-35/openalex_figure/designv11_35_0000142_012076-Figure1-1.png", + "caption": "Figure 1. Belt transmission schematic and brush model concept.", + "texts": [ + " In addition, the results are also useful during the design phase of a transmission since they allow to select and, eventually, optimize the set of parameters, which determine the transmission kinematics and efficiency. The mathematical model implemented to perform the sensitivity analysis is the one developed in [10] which is recalled in this section. The belt is assumed as composed of the tension member, made of the reinforcement fibers which are much stiffer than the rubber matrix, composed of a bed of elastically deformable bristles in contact with the pulley. The model is planar and the following equations are valid per unit of width of the belt. Figure 1 shows a schematic of the problem where the main geometry, static and kinematic parameters are given: Rdg and Rdn are the driving and the driven pulleys\u2019 radii, \u03c9dg and \u03c9dn is the driving and driven pulley angular velocity andMdg andMdn are the torques acting on the driving and driven pulley, respectively. Considering the continuity equation, for a given angular coordinate \u03b1, the belt velocity Vb and the belt tension T are linked by the following relationship Vb(\u03b1) = v1 EA+ T (\u03b1) EA+ T1 = v2 EA+ T (\u03b1) EA+ T2 (1) where T1 and T2 are the belt tension on the tight and slack side, respectively, EA represents the longitudinal stiffness per unit width of the belt, v1 [v2] is the peripheral speed of the belt in correspondence of the belt tension T1 [T2]" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_35_0000225_j.addma.2021.101955-Figure15-1.png", + "original_path": "designv11-35/openalex_figure/designv11_35_0000225_j.addma.2021.101955-Figure15-1.png", + "caption": "Fig. 15. FEM model of the part with the substrate.", + "texts": [ + " The inherent strain of this sample part was calculated as the average residual plastic strain inside the interest area of a heat affected zone model simulated by ANSYS [46] with temperature history generated by a thermal simulation model developed by the authors [52]. The temperature-dependent thermal properties and mechanical properties used in the simulation were based on the properties of SUS316 stainless steel [53]. The FEM model of the part with the substrate includes 7501 nodes and 4842 quadratic tetrahedron elements, as shown in Fig. 15. The FEM model was discretized into 0.8 million voxels (part only excluding the substrate) with 0.5 \u00d7 0.5 \u00d7 0.1 mm dimension. Because the gaps between the unit force couple were 0.1 mm, 0.1 mm and 0.02 mm along the X, Y, and Z axis, respectively, the distortion factors were linearly scaled up five times for this voxel model. The inherent strain applied on each voxel was the same because the 90-degree rotation hatch pattern would theoretically form plastic strains with the same amplitude on all layers" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_35_0000149_s11548-021-02338-9-Figure5-1.png", + "original_path": "designv11-35/openalex_figure/designv11_35_0000149_s11548-021-02338-9-Figure5-1.png", + "caption": "Fig. 5 Guiding LODEM (Follower 1): a photograph of the prototype device composed of a circular ring-guided rail mechanism driven by stepping motors with three DOFs, and b schematic diagram of the device and the polar coordinates", + "texts": [ + "9\u00b0 for the pitch axis and 0.8\u00b0 for the yaw axis. The position of the leader handle (xm, ym, zm) in Cartesian coordinates is calculated using the cross product of two-unit vectors in the pitch and yaw axes, the pitch angle \u03b81, the yaw angle \u03b82, and the insertion position lm, as follows: (1) \u23a1\u23a2\u23a2\u23a3 xm ym zm \u23a4 \u23a5\u23a5\u23a6 = lm\ufffd 1 \u2212 sin 2 1 sin 2 2 \u23a1 \u23a2\u23a2\u23a3 sin 1 cos 2 \u2212 cos 1 sin 2 cos 1 cos 2 \u23a4\u23a5\u23a5\u23a6 1 3 The guiding LODEM, which has three stepping-motorcontrolled wire-tube-driven DOFs in polar coordinates (Follower 1), is shown in Fig.\u00a05. The mechanism of the manipulator, which can be autoclave sterilized, is constructed by a circular ring slider and a circular ring-guided rail for the yaw axis, a pair of slide blocks and a pair of half circular ring-guided rails for the pitch axis, and linear guide rollers with a linear telescopic slider for the insertion axis with 5-mm-diameter forceps. In the present study, the insertion axis is modified to provide a laparoscope of 12\u00a0mm in diameter. Each axis is driven by a super polyethylene (PE) wire (outer diameter: 0" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_35_0000168_j.mechatronics.2021.102514-Figure2-1.png", + "original_path": "designv11-35/openalex_figure/designv11_35_0000168_j.mechatronics.2021.102514-Figure2-1.png", + "caption": "Fig. 2. Example of the cutting process (a) 5R manipulator (b) The open-chain system of 5R.", + "texts": [ + " It should be noted that the bias of the hidden neurons and output neurons are replaced by using augmented inputs for simplicity. The multi-layer neural network possesses the following property. Property 1 ([21]). An arbitrary continuous function is denoted by \ud835\udc53 (\ud835\udc67) that transforms a compact set from R\ud835\udc56 to R\ud835\udc5c, for any positive constant \ud835\udf0c, there exists hidden neuron number \ud835\udc41\u2217, and the ideal weights matrices \ud835\udc4a \u2217 and \ud835\udc49 \u2217, such that \u2016\ud835\udc53 (\ud835\udc67) \u2212\ud835\udc4a \u2217?\u0304?(\ud835\udc49 \u2217?\u0304?)\u2016 \u2264 \ud835\udf0c. According to [10], the PM dynamics can be modeled by three steps. Firstly, cut the PM to several open-chain systems, and an example is given in Fig. 2 to show the cutting process. Secondly, derive \ud835\udc65 the dynamics equation of each open-chain system through the Euler\u2013 Lagrange approach. Finally, combine these equations through matrix manipulation methods to derive the PM system dynamics. Assume the PM is cut to \ud835\udc41 open-chain systems. Since each openchain system can be modeled as a serial manipulator, the \ud835\udc56\ud835\udc61\u210e, \ud835\udc56 = 1, 2,\u2026 , \ud835\udc41 , open-chain system dynamics can be written in the following manner [16]. \ud835\udc40\ud835\udc56(\ud835\udc5e\ud835\udc56)\ud835\udc5e\ud835\udc56 + \ud835\udc36\ud835\udc56(\ud835\udc5e\ud835\udc56, ?\u0307?\ud835\udc56)?\u0307?\ud835\udc56 + \ud835\udc3a\ud835\udc56(\ud835\udc5e\ud835\udc56) = \ud835\udf0f\ud835\udc56, (1) where \ud835\udc5e\ud835\udc56, \ud835\udc40\ud835\udc56, \ud835\udc36\ud835\udc56, \ud835\udc3a\ud835\udc56 and \ud835\udf0f\ud835\udc56 are the generalized coordinate, the inertia matrix, the Coriolis matrix, the gravitational force vector and the control input vector of the \ud835\udc56\ud835\udc61\u210e open-chain system", + " Programmable controller, National Instruments cRIO-9047, is chosen as the control unit. The sampling frequency of the joint angle and joint velocity is 1000(Hz). The implementation of the ESC to the system is separated into two parts. On the one hand, an interface is built to send the highlevel requests such as enabling the servos, starting the experiment and receiving the recorded data on a computer. On the other hand, the main program of the ESC is deployed to the controller using LabView. The mechanical model is given in Fig. 2. The position vector of point \ud835\udc5d is denoted by \ud835\udc65 = [\ud835\udc65\ud835\udc5d, \ud835\udc66\ud835\udc5d]\ud835\udc47 . According to the experiment setup, we have \ud835\udc5e = [\ud835\udc5e1, \ud835\udc5e2, \ud835\udc5e3, \ud835\udc5e4]\ud835\udc47 and \ud835\udc5e\ud835\udc4e = [\ud835\udc5e1, \ud835\udc5e3]\ud835\udc47 . The kinematic parameters of the 5R are given as follows |\ud835\udc341\ud835\udc342| = 0.08 (m), |\ud835\udc341\ud835\udc351| = |\ud835\udc342\ud835\udc352| = 0.08 (m) and |\ud835\udc351\ud835\udc5d| = |\ud835\udc352\ud835\udc5d| = 0.2 (m). The estimation of the system parameters is given in Table 2, where \ud835\udc59\ud835\udc501, \ud835\udc59\ud835\udc502, \ud835\udc591, \ud835\udc592, \ud835\udc5a1, \ud835\udc5a2, \ud835\udc3d\ud835\udc501 and \ud835\udc3d\ud835\udc502 represent the centroid location length, the mass and the moment of inertia with respect to the centroid of the link \ud835\udc34\ud835\udc56\ud835\udc35\ud835\udc56 and link \ud835\udc35\ud835\udc56\ud835\udc5d, \ud835\udc56 = 1, 2, and \ud835\udc3d\ud835\udc5012 = \ud835\udc3d\ud835\udc501 + \ud835\udc3d\ud835\udc502" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_35_0001419_978-94-007-7401-8-Figure2.6-1.png", + "original_path": "designv11-35/openalex_figure/designv11_35_0001419_978-94-007-7401-8-Figure2.6-1.png", + "caption": "Fig. 2.6 4PRRRR-type fully-parallel PMs with coupled Sch\u00f6nflies motions defined by MF \u00bc SF \u00bc 4; RF\u00f0 \u00de \u00bc \u00f0v1; v2; v3;xd\u00de, TF = 0, NF = 2, limb topology P\\R\\R||R\\R (a) and P||R\\R||R\\R (b)", + "texts": [ + "1c) The two last revolute joints of the four limbs have parallel axes 4. 4RRRRR (Fig. 2.3b) R||R||R\\R||R (Fig. 2.1c) The three first revolute joints of the four limbs have parallel axes 5. 4RRRRR (Fig. 2.4a) R||R\\R||R||R (Fig. 2.1d) The two first revolute joints of the four limbs have parallel axes. 6. 4PRRRR (Fig. 2.4b) P\\R\\R||R||R (Fig. 2.1e) The second joints of the four limbs have parallel axes 7. 4RRRPR (Fig. 2.5a) R||R\\R\\P\\kR (Fig. 2.1f) Idem No. 5 8. 4RRPRR (Fig. 2.5b) R||R||P\\R||R (Fig. 2.1g) Idem No. 5 9. 4PRRRR (Fig. 2.6a) P\\R\\R||R\\R (Fig. 2.1h) The second and the last joints of the four limbs have parallel axes 10. 4PRRRR (Fig. 2.6b) P||R\\R||R\\R (Fig. 2.1i) Idem No. 9 11. 4RRPRR (Fig. 2.7a) R\\R\\P\\kR\\R (Fig. 2.1j) Idem No. 1 12. 4PRRRR (Fig. 2.7b) P\\R||R\\R||R (Fig. 2.1k) Idem No. 3 13. 4PRRRR (Fig. 2.8a) P||R||R||R||R\\R (Fig. 2.1l) The last revolute joints of the four limbs have parallel axes 14. 4RPRRR (Fig. 2.8b) R||P||R||R\\R (Fig. 2.1m) Idem No. 13 15. 4RRPRR (Fig. 2.9a) R||R||P||R\\R (Fig. 2.1n) Idem No. 13 16. 4RRRPR (Fig. 2.9b) R||R||R||P\\R (Fig. 2.1o) Idem No. 13 17. 4RPRRR (Fig. 2.10a) R||P||R||R\\R (Fig. 2.1p) Idem No" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_35_0001419_978-94-007-7401-8-Figure2.111-1.png", + "original_path": "designv11-35/openalex_figure/designv11_35_0001419_978-94-007-7401-8-Figure2.111-1.png", + "caption": "Fig. 2.111 4PRRRPa-type fully-parallel PM with coupled Sch\u00f6nflies motions defined by MF \u00bc SF \u00bc 4; RF\u00f0 \u00de \u00bc \u00f0v1; v2; v3;xd\u00de, TF = 0, NF = 14, limb topology P\\R||R\\R||Pa", + "texts": [], + "surrounding_texts": [ + "2.2 Topologies with Complex Limbs 165", + "166 2 Fully-Parallel Topologies with Coupled Sch\u00f6nflies Motions", + "2.2 Topologies with Complex Limbs 167" + ] + }, + { + "image_filename": "designv11_35_0000440_mias.2021.3065325-Figure2-1.png", + "original_path": "designv11-35/openalex_figure/designv11_35_0000440_mias.2021.3065325-Figure2-1.png", + "caption": "FIGURE 2. PM constructions. (a) A surface PM. (b) An interior PM.", + "texts": [ + " A PMM is a synchronous machine. This motor type may also be referred to as a PMac, PM synchronous, and brushless ac synchronous. The stator is similar in design to that of any conventional induction or synchronous machine in which electrical power is supplied through this winding. What is unique about this motor is that the rotor employs permanent rare-earth magnets that are either surface mounted on the rotor lamination stack or embedded within the rotor laminations, producing what is known as an interior PM. Figure 2 describes these two options. Regardless of how the Motor Efficiency (%) Motor Rating (hp) Standard Efficiency High Efficiency 25-Year Energy Cost Savings 1 76.5 83.8 US$716.75 5 84.1 89.3 US$2,179.11 7.5 85.9 91 US$3,080 10 86.9 91.3 US$3,490.72 100 91 95.5 US$32,592.7 500 93.3 96.2 US$101,686.60 Table 1. The 25-year projected energy savings from high-efficiency motors Authorized licensed use limited to: San Francisco State Univ. Downloaded on June 16,2021 at 01:11:30 UTC from IEEE Xplore. Restrictions apply" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_35_0000556_iemdc47953.2021.9449493-Figure4-1.png", + "original_path": "designv11-35/openalex_figure/designv11_35_0000556_iemdc47953.2021.9449493-Figure4-1.png", + "caption": "Fig. 4. Analysis model of the target motor (a) Stator (b) Rotor. (a) (b)", + "texts": [ + " 3 (c) and (d) are the S-HPMM with the bar type Fe-PM in [12] and the proposed S-HPMM. The stator structure of all 4 models is the same as that of the target motor, and have the same usage of Fe-PM and Nd-PM. Also, the thickness of Nd-PM is 6.5mm, which is the same for all. The material of the rotor core is applied high-tensile electrical steel sheet with a yield strength of 700 MPa to maintain rotational stiffness at the maximum speed. III. TARGET MOTOR The target motor is the traction motor of TOYOTA PRIUS 4th-generation HEV commercialized in 2015. Fig. 4 is the analysis model of the target motor, and Table I is the detailed specifications of the target motor and the proposed HPMM. The number of stator slots of the target motor is 48, and it is a distributed winding. The rotor has an 8-pole structure. Each model has the same stator structure and overall volume. As each magnet material for analysis, Nd-PM is N42SH-R (21kOe, ShinEtsu Co., Ltd.), and Fe-PM is NMF-15J (5.5kOe, Hitachi metals Co., Ltd.). The remanence of the target motor was estimated to be 1" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_35_0000228_j.matpr.2021.02.620-Figure1-1.png", + "original_path": "designv11-35/openalex_figure/designv11_35_0000228_j.matpr.2021.02.620-Figure1-1.png", + "caption": "Fig. 1. 3-D Model of Connecting Rod.", + "texts": [ + " The results are compared at the end pertaining to static stress analysis, fatigue stress analysis and cost of the connecting rod. The different materials chosen for connecting rod are namely AISI 4140 Cr-Mo High Tensile Steel, 42crMo4, Aluminum 7075 T6 and Forged steel. The different properties of the material chosen are listed in Table 1. To begin with, the analytical model of connecting rod was built in NX11 Unigraphics and the numerical analysis has been carried out assigning the all chosen materials with corresponding properties. The 3D model of connecting rod built in Unigraphics is shown in Fig. 1.To obtain the numerical analysis, the build model of NX11 Unigraphics has been imported to ANSYS work bench environment Select simulation for analysis of connecting rod Made the connection between the cup of the connecting rod and the whole parts Select the meshing method of tetrahedral Apply the boundary condition Apply the option of simulation Select the result of simulation result (see Figs. 2 and 3). The analytical computations associated with the design of connecting rod and as well numerical outcomes of the same are discussed using the 42crMo4 material" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_35_0000146_012064-Figure3-1.png", + "original_path": "designv11-35/openalex_figure/designv11_35_0000146_012064-Figure3-1.png", + "caption": "Figure 3. Contact description.", + "texts": [ + " Series: Materials Science and Engineering 1038 (2021) 012064 IOP Publishing doi:10.1088/1757-899X/1038/1/012064 \ud835\udc3e = [ \ud835\udc3e\ud835\udc4e\ud835\udc65 0 0 \u2212\ud835\udc3e\ud835\udc4e\ud835\udc65 0 0 0 12\ud835\udc3e\ud835\udc53\ud835\udc59 \ud835\udc3f2(1 + \u03a6) 6\ud835\udc3e\ud835\udc53\ud835\udc59 \ud835\udc3f(1 + \u03a6) 0 \u221212\ud835\udc3e\ud835\udc53\ud835\udc59 \ud835\udc3f2(1 + \u03a6) 6\ud835\udc3e\ud835\udc53\ud835\udc59 \ud835\udc3f(1 + \u03a6) \ud835\udc3e\ud835\udc53\ud835\udc59(4 + \u03a6) (1 + \u03a6) 0 \u22126\ud835\udc3e\ud835\udc53\ud835\udc59 \ud835\udc3f(1 + \u03a6) \ud835\udc3e\ud835\udc53\ud835\udc59(2 \u2212 \u03a6) (1 + \u03a6) \ud835\udc3e\ud835\udc4e\ud835\udc65 0 0 \ud835\udc46\ud835\udc66\ud835\udc5a 12\ud835\udc3e\ud835\udc53\ud835\udc59 \ud835\udc3f2(1 + \u03a6) \u22126\ud835\udc3e\ud835\udc53\ud835\udc59 \ud835\udc3f(1 + \u03a6) \ud835\udc3e\ud835\udc53\ud835\udc59(4 + \u03a6) (1 + \u03a6) ] where shearing factor \u03a6 can be computed as: \u03a6 = \ud835\udc3e\ud835\udc53\ud835\udc59 \ud835\udc3e\ud835\udc60\u210e\ud835\udc3f2 A penalty-based contact algorithm has been developed to account for coils contact in a realistic, yet efficient manner. The contact is evaluated on working plane using contact circles A and B in figure 3. Those circles are obtained as the intersection of the working plane with the spring wire (neglecting a secondary effect due to helix angle). When the distance between centers is less than the mean wire diameter, a contact force generates. The direction of the force is that of the line connecting the center of the two circles. Of course, the forces are applied to nodes 1 and 3 involved in contact and thus a moment arise. Thus, the contact is computed in the real location where it develops and then transferred to the equivalent beam nodes. With reference to figure 3, the coordinates of contact circles centers are: { \ud835\udc65\ud835\udc34 \ud835\udc66\ud835\udc34 } = { \ud835\udc651 \ud835\udc661 } + \ud835\udc371 2 { cos\ud835\udf031 sin\ud835\udf031 } (10) { \ud835\udc65\ud835\udc35 \ud835\udc66\ud835\udc35 } = { \ud835\udc653 \ud835\udc663 } + \ud835\udc373 2 { cos\ud835\udf033 sin\ud835\udf033 } (11) The 49th AIAS Conference (AIAS 2020) IOP Conf. Series: Materials Science and Engineering 1038 (2021) 012064 IOP Publishing doi:10.1088/1757-899X/1038/1/012064 Thus, it is possible to compute the actual distance and penetration: \u210e = \u221a(\ud835\udc65\ud835\udc34 \u2212 \ud835\udc65\ud835\udc35)2 + (\ud835\udc66\ud835\udc34 \u2212 \ud835\udc66\ud835\udc35)2 \ud835\udc51 = 1 2 (\ud835\udc511 + \ud835\udc513) Introducing a convenient contact force expression [2] with appropriate stiffness and exponent: { \ud835\udc39 = \ud835\udc3e\ud835\udc36(\ud835\udc51 \u2212 \u210e)\ud835\udc58 (\ud835\udc51 \u2212 \u210e) > 0 (12) Angle \ud835\udf19 gives the direction of the force and allows to calculate the coordinates of contact point C: cos\ud835\udf19 = \ud835\udc65\ud835\udc35 \u2212 \ud835\udc65\ud835\udc34 \u210e sin\ud835\udf19 = \ud835\udc66\ud835\udc35 \u2212 \ud835\udc66\ud835\udc34 \u210e { \ud835\udc65\ud835\udc36 \ud835\udc66\ud835\udc36 } = { \ud835\udc65\ud835\udc34 \u2212 \ud835\udc513 2 cos\ud835\udf19 \ud835\udc66\ud835\udc34 \u2212 \ud835\udc513 2 sin\ud835\udf19 } = { \ud835\udc65\ud835\udc35 + \ud835\udc511 2 cos\ud835\udf19 \ud835\udc66\ud835\udc35 + \ud835\udc511 2 sin\ud835\udf19 } (13) Combining equations (10) and (13), and magnitude \ud835\udc39 from equation (12), it is possible to compute the generalized nodal forces, as function of nodal coordinates only: \ud835\udc6d\ud835\udfcf = { \ud835\udc391\ud835\udc65 \ud835\udc391\ud835\udc66 \ud835\udc401 } = \ud835\udc39 { \u2212 cos\ud835\udf19 \u2212sin\ud835\udf19 sin\ud835\udf19 ( \ud835\udc371 2 cos \ud835\udf031 + \ud835\udc511 2 cos\ud835\udf19) \u2212 cos\ud835\udf19 ( \ud835\udc371 2 sin \ud835\udf031 + \ud835\udc511 2 sin\ud835\udf19) } \ud835\udc6d\ud835\udfd0 = { \ud835\udc393\ud835\udc65 \ud835\udc393\ud835\udc66 \ud835\udc403 } = \ud835\udc39 { cos\ud835\udf19 sin\ud835\udf19 sin\ud835\udf19 ( \ud835\udc373 2 cos \ud835\udf033 \u2212 \ud835\udc513 2 cos\ud835\udf19) \u2212 cos\ud835\udf19 ( \ud835\udc373 2 sin \ud835\udf032 \u2212 \ud835\udc513 2 sin\ud835\udf19) } Contact stiffness matrix is given by contact forces derivatives with respect to nodal degrees of freedom" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_35_0001419_978-94-007-7401-8-Figure3.39-1.png", + "original_path": "designv11-35/openalex_figure/designv11_35_0001419_978-94-007-7401-8-Figure3.39-1.png", + "caption": "Fig. 3.39 Overactuated PMs with coupled Sch\u00f6nflies motions of types 4RRPRP (a) and 4RRRPP (b) defined by MF \u00bc SF \u00bc 4; RF\u00f0 \u00de \u00bc v1; v2; v3;xd\u00f0 \u00de, TF = 0, NF = 2, limb topology R\\R||P||R\\P (a) and R\\R||R||P\\P (b)", + "texts": [ + " PM type Limb topology Connecting conditions 30. 4RPRRR (Fig. 3.35a) R\\P\\kR\\R||R (Fig. 3.3q) Idem No. 16 31. 4RRRPR (Fig. 3.35b) R||R\\R||P||R (Fig. 3.3r) Idem No. 16 32. 4RRRPR (Fig. 3.36a) R||R\\R||P||R (Fig. 3.3s) Idem No. 16 33. 4RRRRP (Fig. 3.36b) R||R\\R||R||P (Fig. 3.3t) Idem No. 16 34. 4RRRRP (Fig. 3.37a) R||R\\R||R\\P (Fig. 3.3u) Idem No. 16 35. 4PRRRP (Fig. 3.37b) P\\R\\R||R\\P (Fig. 3.3v) Idem No. 29 36. 4RRPRP (Fig. 3.38a) R\\R||P||R\\P (Fig. 3.3w) Idem No. 16 37. 4RPRRP (Fig. 3.38b) R\\P||R||R\\P (Fig. 3.3x) Idem No. 16 38. 4RRPRP (Fig. 3.39a) R\\R||P||R\\P (Fig. 3.3y) Idem No. 16 39. 4RRRPP (Fig. 3.39b) R\\R||R||P\\P (Fig. 3.3z) Idem No. 16 40. 4PRRRP (Fig. 3.40a) P||R||R\\R\\P (Fig. 3.3z1) The fourth joints of the four limbs have parallel axes 41. 4RPRRP (Fig. 3.40b) R||P||R||R\\P (Fig. 3.3a) Idem No. 40 42. 4RPRRP (Fig. 3.41a) R||P||R\\R\\P (Fig. 3.3b0) Idem No. 40 43. 4RRPRP (Fig. 3.41b) R||R||P\\R\\P (Fig. 3.3c0) Idem No. 40 44. 4RRRPP (Fig. 3.42a) R||R\\R\\P\\\\P (Fig. 3.3d0) The third joints of the four limbs have parallel axes 45. 4PRRPR (Fig. 3.42b) P||R||R\\P\\\\R (Fig. 3.3e0) Idem No. 17 46. 4RPRPR (Fig" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_35_0000686_j.mechmachtheory.2021.104433-Figure1-1.png", + "original_path": "designv11-35/openalex_figure/designv11_35_0000686_j.mechmachtheory.2021.104433-Figure1-1.png", + "caption": "Fig. 1. Structures of the planar epicyclic gear trains.", + "texts": [], + "surrounding_texts": [ + "The epicyclic gear train is one of the most basic gear transmission mechanisms and is widely used in various mechanisms with rotary motion [1] due to its advantages including a high transmission ratio, low weight, and good adaptability to speed and direction changes [2]. However, there are no study and application of epicyclic gear trains with metamorphosis. Therefore, the revelation of the metamorphic phenomenon with configuration analysis in epicyclic gear mechanism becomes an urgent issue to further expand the research and application fields of the metamorphic mechanisms. Metamorphic mechanisms were proposed by Dai and Rees Jones [3] in the 1990s and were inspired by the abstraction and evolution of complex origami art. The mechanisms change their topologies and mobility with link-annexing and with joint property * Corresponding author. Contents lists available at ScienceDirect Mechanism and Machine Theory journal homepage: www.elsevier.com/locate/mechmt https://doi.org/10.1016/j.mechmachtheory.2021.104433 Received 23 April 2021; Received in revised form 9 June 2021; Accepted 16 June 2021 Mechanism and Machine Theory 166 (2021) 104433 change resulting from geometrical constraints and with reconfiguration joints. Therefore, metamorphic mechanisms meet various demands and adapt to various working environments due to their multiconfiguration, multi-topological structure with multi-mobility characteristics, becoming an important research direction, in special robotics, in the aerospace industry, and the manufacturing industry. At present, extensive research has been carried out on metamorphic mechanism design and configuration analysis [4]. Li et al. [5-8] proposed a synthesis method of metamorphic mechanisms by matrix operation of configuration transformations. Carroll et al. presented metamorphic mechanisms based on compliant links [9,10]. A metamorphic mechanism design method was developed in [11] by utilizing the coupled links, the link relationship, and the characteristics of the kinematic pairs based on metamorphic research results. Reference [12] defined five metamorphic joints due to constraint conditions and proposed a structure synthesis method by combinations of metamorphic joints according to mobility changes. Li and Dai [13,14] proposed a structural theory of metamorphic mechanisms based on the augmented Assur group and extended it to a spatial metamorphic mechanism. Rodriguez et al. [15-18] designed a variety of mechanisms inspired by origami progress. Gan and Dai [19] introduced an invented reconfigurable Hooke (rT) joint and synthesized metamorphic parallel mechanisms. Additionally, Zhang et al. [21] created variable-axis (vA) joints that can realize changes in mobility and investigated one type of metamorphic parallel mechanism. Gan [22] proposed a reconfigurable revolute joint (rR) which qualified the new 3rRPS metamorphic parallel mechanism to be able to provide both 3R and 1T2R motion independently. Wei [23] proposed a variable revolute (vR) joint and verified its application by a reconfigurable generic 4R linkage. Ma, Song et al. [24-25] introduced a kind of metamorphic linkage that is capable of changing its motion branches for reconfigurable robots. For configuration transforming and working sequence problems of practical constrained metamorphic processes, Li and Yang et al. [26,27] proposed a task-based design method based on the relationship of equivalent resistances, constraint characteristics and forms, and the structures of metamorphic joints in the working stages, using the equivalent resistance gradient of metamorphic joints. A method for configuration synthesis of metamorphic mechanisms was presented based on functional analysis [28]. Meanwhile, studies have been conducted on metamorphic higher pair mechanisms. Guo [29] used the method of replacing higher pairs with lower pairs to analyse kinematic and mobility variations of the three configurations of planetary wheeled metamorphic inspection robots. The concept of higher pair Assur groups [30] based on the augmented Assur group was proposed but needed further study in terms of design and synthesis. Here, the metamorphic phenomenon of epicyclic gear trains is essential in terms of using the metamorphosis technique in higher pair gear trains. Several graph-theoretic methods [31-34] have been developed to simplify the kinematic analysis of gear trains. In reference [35,36], an ePaddle-EGM robot was designed and optimized based on the relative position of the sun gear and planetary gear in an epicyclic gear train. Laus et al. [37] analysed the efficiency of epicyclic gear trains using graph and screw theories. A novel tricolour topology expression method was proposed in [38] for the topology synthesis of a multirow planetary gear. The current design methods for metamorphic mechanisms mainly focus on the configuration synthesis of lower pair mechanisms and the application of reconfiguration joints, while only graph theory is used. Graph theory mainly focuses on the relation changing of the components in metamorphic mechanisms, while the higher pair Assur groups method develops basic higher pair Assur groups for metamorphic structure synthesis. For the higher pair study, a linkage that is kinematically equivalent to a mechanism with a higher H. Yuan et al. Mechanism and Machine Theory 166 (2021) 104433 pair was presented based on tori in [39,40] but not specified any higher pair [41] reviewed analytical methods for cylindrical gear pairs. For the replacement method of lower pair and higher pair, [42,43] introduced a way to use a link mechanism to replace fixed an axis external gear train and a gear-rack train with 1 mobility. For linkage synthesis, Liu et al [44-46] presented motion synthesis, function synthesis, and path synthesis for linkages. However, little literature is reported in the linkages that are equivalent to the epicyclic gear train. Hence, the epicyclic gear train, which is a higher pair mechanism, still needs more study to establish an appropriate kinematic model for using the metamorphosis principle for configurations, motion branches, and constraint condition analysis. In this paper, three 5-bar mechanisms are presented for the first time to be utilized as equivalent mechanisms for the epicyclic gear trains. Motion branches and constraint conditions of epicyclic gear trains are analysed by equivalent mechanism methods using firstand second-order kinematics. This generated a set of novel metamorphic epicyclic gear trains based on geometrical and force constraints. The layout of this paper is as follows. The equivalent mechanisms and equivalent geometrical relationships of three typical epicyclic gear trains are established in Section 2 based on the instantaneous screw axes of gears. The first- and second-order kinematics based on Lie brackets of the equivalent mechanisms are presented and solved in Section 3 to reveal the motion branches and obtain constraint conditions. Following the equivalent geometrical relationships, Section 4 describes the metamorphic configurations, motion branches, and metamorphic constraint conditions of epicyclic gear trains. This leads to a set of novel metamorphic epicyclic gear trains designed based on constrained metamorphic mechanisms in Section 5. The study further leads to a metamorphic clamping mechanism invented as an example to demonstrate the application of metamorphic epicyclic gear trains in both Sections 6 and 7. Finally, conclusions are drawn in Section 8." + ] + }, + { + "image_filename": "designv11_35_0000149_s11548-021-02338-9-Figure6-1.png", + "original_path": "designv11-35/openalex_figure/designv11_35_0000149_s11548-021-02338-9-Figure6-1.png", + "caption": "Fig. 6 Portable LODEM (Follower 2): a photograph of the prototype device composed of a gimbal-mounted parallel linkage mechanism driven by DC servo motors with three DOFs, and b schematic diagram of the device and the gimbal coordinates", + "texts": [ + " The positional accuracy when loaded by 5\u00a0N at the tip of the forceps is 0.1\u00a0mm for the pitch axis, 0.1\u00a0mm for the yaw axis, and 2.0\u00a0mm for the insertion axis. The mechanical deflection is 0.2\u00a0mm for the pitch axis and 0.1\u00a0mm for the yaw axis. The position of the forceps handle [xsyszs]t in Cartesian coordinates is calculated using a homogeneous transformation matrix, yaw angle \u03b1, pitch angle \u03b2, and insertion position ls, as follows: The portable LODEM, which has three DC servo motor-controlled DOFs in gimbal coordinates (Follower 2), is shown in Fig.\u00a06. The LODEM consists of a separable pivot restraint part that can be sterilized and an actuator unit draped with (2) \u23a1\u23a2\u23a2\u23a3 xs ys zs \u23a4 \u23a5\u23a5\u23a6 = \u23a1 \u23a2\u23a2\u23a3 ls cos sin ls sin sin ls cos \u23a4 \u23a5\u23a5\u23a6 . 1 3 1 3 a sterilized cover. This manipulator uses a gimbal-mounted parallel linkage mechanism of the closed loop structure for the pitch and yaw axes, and a wire-driven linear slider mechanism for the insertion axis attached to the 5-mm-diameter forceps. Each axis is driven by a DC motor (maximum continuous torque: 30" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_35_0000417_012020-Figure6-1.png", + "original_path": "designv11-35/openalex_figure/designv11_35_0000417_012020-Figure6-1.png", + "caption": "Figure 6. (a) Tire testing machine, Ektron PL-2003, (b) FE model of vertical stiffness testing of NPT, and (c) boundary conditions.", + "texts": [ + " In addition, each rebar element composed of mathematically sublayers of belt. The outer layers, middle layers, and inner layers composed of 3, 1, and 2 sublayers, respectively. The bead wire diameter was estimated by measuring to be 1 mm, while the number of wires per unit length was estimated to 0.3582 mm-1. The simulation of vertical stiffness testing was performed using FE software, MSC. Marc. The FE mesh of NPT is combined with rigid plate, which represented tire testing machine\u2019s moving plate as shown in Figure 6(a). The contact algorithm with Coulomb\u2019s friction of 0.8 was used to calculate interactive force between NPT and rigid plate. The plate was assigned to press upward against the NPT in vertical direction with force of 20 kN or until the displacement of 26 mm was obtained. The NPT was assigned to be fixed at the inner spoke portion. This represented the locking adapter at the rim in the actual experiment. The analysis was then performed and the results were collected. The11th International Conference on Mechanical Engineering (TSME-ICOME 2020) IOP Conf", + " On the other hand, the bending and compression were observed on the lower portion of spoke on both models. This indicate the feature to uniformly distribute the load along all portion of spoke, not only on the lower portion which was directly contact the plate. The vertical force and displacement were then collected. In addition, the force and displacement data were also collected from the vertical stiffness testing experiment of actual NPT. The experiment is performed by tire testing machine, Ektron PL-2003 (Figure 6). The vertical force and displacement relationship of the experiment and both FE model are plotted as shown in Figure 8. The vertical stiffness could be derived from ratio between vertical force and displacement. The estimated vertical stiffness values of FE models and experiment are shown in Table 5. The FEM\u2019s analysis error could be estimated by comparing between the waterjet FE model and the experiment because they were based on the same material and conditions. The comparison showed that the estimated vertical stiffness has error of 9" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_35_0001419_978-94-007-7401-8-Figure6.24-1.png", + "original_path": "designv11-35/openalex_figure/designv11_35_0001419_978-94-007-7401-8-Figure6.24-1.png", + "caption": "Fig. 6.24 Maximaly regular fully-parallel PMs with Sch\u00f6nflies motions of types 3PPaPR1RUPU (a) and 2PPaRRR-1PPaRR-1RUPU (b) defined by MF = SF = 4, (RF) = (v1; v2; v3;xb), TF = 0, NF = 13 (a), NF = 11 (b), limb topology R\\R\\R\\P\\||R\\R and P||Pa\\P ??R, P||Pa\\P\\||R (a), P||Pa||R||R\\R, P||Pa||R||R (b)", + "texts": [ + "1b) The first revolute joint of G4-limb and the last last joints of limbs G1, G2 and G3 have parallel axes. The last revolute joints of limbs G1, G2 and G3 have superposed axes. The actuated prismatic joints of limbs G1, G2 and G3 have orthogonal directions (continued) 614 6 Maximally Regular Topologies with Sch\u00f6nflies Motions Table 6.6 (continued) No. PM type Limb topology Connecting conditions 14. 3PPPaR-1RUPU (Fig. 6.23b) P ?P ?||Pa\\\\R (Fig. 4.8a) P ?P ?||Pa||R, (Fig. 4.8c) R ?R ?R ?P ?||R ?R (Fig. 5.1b) Idem no. 13 15. 3PPaPR-1RUPU (Fig. 6.24a) P||Pa ?P\\\\R (Fig. 4.8d) P||Pa ?P ?||R (Fig. 4.8e) R ?R ?R ?P ?||R ?R (Fig. 5.1b) Idem no. 13 16. 2PPaRRR-1PPaRR-1RUPU (Fig. 6.24b) P||Pa||R||R ?R (Fig. 4.9a) P||Pa||R||R (Fig. 4.8g) R ?R ?R ?P ?||R ?R (Fig. 5.1b) Idem no. 13 17. 3PPaPaR-1RPPP (Fig. 6.25) P||Pa||Pa ?R (Fig. 4.10a) P||Pa||Pa||R (Fig. 4.10b) R ?P\\\\P\\\\P (Fig. 5.1a) Idem no. 9 18. 3PPaPaR-1RUPU (Fig. 6.26) P||Pa||Pa ?R (Fig. 4.10a) P||Pa||Pa||R (Fig. 4.10b) R ?R ?R ?P ?||R ?R (Fig. 5.1b) Idem no. 13 Table 6.7 Structural parametersa of parallel mechanisms in Figs. 6.7, 6.8, 6.9, 6.10, 6.11, 6.12, 6.13, 6.14 No. Structural parameter Solution Figures 6.7 and 6.11 Figures 6", + "2 for the nomenclature of structural parameters 618 6 Maximally Regular Topologies with Sch\u00f6nflies Motions Table 6.10 (continued) No. Structural parameter Solution Figures 6.21 and 6.22a Figure 6.22b 24. rl 9 9 25. rF 21 23 26. MF 4 4 27. NF 15 13 28. TF 0 0 29. Pp1 j\u00bc1 fj 7 8 30. Pp2 j\u00bc1 fj 7 7 31. Pp3 j\u00bc1 fj 7 8 32. Pp4 j\u00bc1 fj 4 4 33. Pp j\u00bc1 fj 25 27 a See footnote of Table 2.2 for the nomenclature of structural parameters 6.2 Fully-Parallel Topologies with Simple and Complex Limbs 619 Table 6.11 (continued) No. Structural parameter Solution Figures 6.23 and 6.24a Figure 6.24b 25. rF 23 25 26. MF 4 4 27. NF 13 11 28. TF 0 0 29. Pp1 j\u00bc1 fj 7 8 30. Pp2 j\u00bc1 fj 7 7 31. Pp3 j\u00bc1 fj 7 8 32. Pp4 j\u00bc1 fj 6 6 33. Pp j\u00bc1 fj 27 29 a See footnote of Table 2.2 for the nomenclature of structural parameters 620 6 Maximally Regular Topologies with Sch\u00f6nflies Motions In the fully-parallel and maximally regular topologies of PMs with Sch\u00f6nflies motions F G1-G2-G3-G4 presented in this section, the moving platform n : nGi (i = 1, 2, 3, 4) is connected to the reference platform 1 : 1Gi : 0 by four complex limbs with four or five degrees of connectivity" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_35_0000585_00405175211026533-Figure8-1.png", + "original_path": "designv11-35/openalex_figure/designv11_35_0000585_00405175211026533-Figure8-1.png", + "caption": "Figure 8. Some technical possibilities for implementation of delays.", + "texts": [ + " Something that is not well specified in this workflow is \u201chow long is the certain time\u201d in which the carrier has to look around. This task requires some more systematic analysis. Figure 7 demonstrates the position of the carriers with triangles in the initial (old) and the end (new) configurations. The empty slots are represented by thin rectangles. Ultimately, the carrier at the wrong position has to move one slot back, meaning one slot is delayed in its movement. The technical solution for implementation of a one slot delay can be done in several ways; 1. Use of horn gear with one slot only (Figure 8(a)). This additional gear takes the carrier and moves it around and gives it back to another gear after one slot duration time. An advantage of this solution is that it supports the good dynamics of the process: the horn gear can rotate with the same speed as the others. There may be problems with this approach if the size of the gear is too small to accommodate the complete slot geometry. 2. Use an additional gear with any number of slots, but with controllable velocity. This is calculated so that the received carrier is moved around and given back after the time of one slot. This solution is technically possible with modern motors, but causes impact forces, accelerations, and requires very precise electronics. It is not clear if this performs in production state and remains stable over long periods of time. 3. Use a gear with five slots (Figure 8(b)) or (x * 4\u00fe 1) slots, which rotates synchronous in the system and delivers the carrier in some of the positions in the next groups. (Here x is any number larger than one, but in practice gears with larger number slots for x\u00bc 2, 3 etc. require a lot of space, and are not used.) Here the equation for the delays is: Di \u00bc i NSlots \u00fe d (5) where Di is the required carrier delay in number of slots, i is an integer number representing how many groups the carrier will be delayed, Nslots are the number of slots on each horn gear (assuming equal numbers) and d is the required minimal delay or slot shift in order to reach the new arrangement. This equation is derived by deduction from Figure 8, based on the numbers 1, 5, and 9, using the knowledge that, for the current case machine with horn gears, four slots are used and the repeat length of the pattern is four. So, any position of later groups with four slots will be useful. In reality, if only a machine with 4 4 horn gears is available, none of these solutions can be applied directly. The only solution is the application of the theory of the extended horn gears,32 where the path of the carrier is generated as a composite curve around several gears" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_35_0000296_s12206-021-0435-1-Figure3-1.png", + "original_path": "designv11-35/openalex_figure/designv11_35_0000296_s12206-021-0435-1-Figure3-1.png", + "caption": "Fig. 3. Schematic overview of the joints 5 and 6 assembly [36].", + "texts": [ + " Geometrical modeling of robots can be achieved by the Denavit-Hartenberg (DH) formulation based on kinematic frames located on the joint axes. The St\u00e4ubli RX160 is equipped with 6 revolute joints. The first four joints of the robot are directly driven by servo motors via helical gear transmissions as shown in Fig. 2. For the remaining two joints, the servos are mounted inside the fourth link of the robot and a gear transmission transfers the motion from servo 6 through the 5th joint to the 6th joint [37]. This mechanical design causes a kinematic coupling between the servo 5 and joint 6 as shown in Fig. 3. The angular velocity of servo 6 is related to the difference of joint velocities 5q and 6q . Consequently, one needs to define additional coordinate axes in order to take into account this extra friction model. The velocity 7q associated with the additional coordinate axes is defined as follows: 7 5 6 q q q= + . (1) The frame definition of the robot is shown in Fig. 1, and the corresponding DH parameters in Table 1. For each link in the Table 1, the transformation matrix 1 i iT \u2212 of the frame attached to the link i with respect to the frame attached to link 1i \u2212 can be given as follows: 1 0 0 0 0 1 i i i i i i i i i i i i i ii i i i i i c s c s s a c s c c c s a s T s c d \u03b8 \u03b8 \u03b1 \u03b8 \u03b1 \u03b8 \u03b8 \u03b8 \u03b1 \u03b8 \u03b1 \u03b8 \u03b1 \u03b1\u2212 \u2212\u23a1 \u23a4 \u23a2 \u23a5\u2212\u23a2 \u23a5= \u23a2 \u23a5 \u23a2 \u23a5 \u23a2 \u23a5\u23a3 \u23a6 (2) where ic\u03b8 and is\u03b8 are the standard abbreviations of cos( )i\u03b8 and ( )sin " + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_35_0001419_978-94-007-7401-8-Figure3.82-1.png", + "original_path": "designv11-35/openalex_figure/designv11_35_0001419_978-94-007-7401-8-Figure3.82-1.png", + "caption": "Fig. 3.82 4PaPRP-type overactuated PMs with coupled Sch\u00f6nflies motions defined by MF \u00bc SF \u00bc 4; RF\u00f0 \u00de \u00bc v1; v2; v3;xd\u00f0 \u00de, TF = 0, NF = 18, limb topology Pa||P\\R\\P (a) and Pa\\P\\\\R\\P (b)", + "texts": [ + " 37 39. 4PaPPR (Fig. 3.77a) Pa||P\\P\\\\R (Fig. 3.50l0) Idem No. 37 40. 4PaPPR (Fig. 3.77b) Pa\\P\\kP\\\\R (Fig. 3.50m0) Idem No. 37 41. 4PaRPR (Fig. 3.78) Pa\\R\\P\\kR (Fig. 3.50n0) Idem No. 37 42. 4PaPRR (Fig. 3.79a) Pa||P\\R||R (Fig. 3.50o0) Idem No. 37 43. 4PaPRR (Fig. 3.79b) Pa\\P\\\\R||R (Fig. 3.50p0) Idem No. 37 44. 4PaRRP (Fig. 3.80a) Pa\\R||R\\P (Fig. 3.50q0) Idem No. 23 45. 4RPaRR (Fig. 3.80b) R\\Pa\\kR||R (Fig. 3.50s0) Idem No. 37 46. 4PaRRR (Fig. 3.81) Pa\\R||R||R (Fig. 3.50r0) Idem No. 37 47. 4PaPRP (Fig. 3.82a) Pa||P\\R\\P (Fig. 3.50t0) Idem No. 23 48. 4PaPRP (Fig. 3.82b) Pa\\P\\\\R\\P (Fig. 3.50u0) Idem No. 23 49. 4CPaP (Fig. 3.83a) C||Pa\\P (Fig. 3.50v0) The cylindrical joints of the four limbs have parallel directions 50. 4PaCP (Fig. 3.83b) Pa||C\\P (Fig. 3.50w0) Idem No. 49 51. 4CRPa (Fig. 3.84a) C||R||Pa (Fig. 3.50x0) Idem No. 49 52. 4PCPa (Fig. 3.84b) P\\C||Pa (Fig. 3.50y0) Idem No. 49 53. 4PPaC (Fig. 3.85a) P\\Pa||C (Fig. 3.50z0) Idem No. 49 54. 4RCPa (Fig. 3.85b) R||C||Pa (Fig. 3.50z01) Idem No. 49 55. 4CPPa (Fig. 3.86a) C\\P\\kPa (Fig. 3.50a0 0) Idem No. 49 56. 4PaPC (Fig" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_35_0000360_j.matpr.2021.04.147-Figure6-1.png", + "original_path": "designv11-35/openalex_figure/designv11_35_0000360_j.matpr.2021.04.147-Figure6-1.png", + "caption": "Fig. 6. Equivalent (von mises) stress for Rear Impact Test.", + "texts": [], + "surrounding_texts": [ + "Off road vehicle is shaped to race and steer on different terrains. ORV is designed in such a way that it can endure off-roading terrains. In off-terrain circumstances, the vehicle bears dynamic loads and all that is sustained through the chassis frame. Chassis frame bears every mountings and assembly, so it is expected from an ORV chassis frame to sustain both static and dynamic loads. The selection of materials for chassis greatly depends on the high tensile strength and material light weight. The majority of manufacturers favour lightweight, cost-effective, safe, and recyclable materials." + ] + }, + { + "image_filename": "designv11_35_0000802_s41315-021-00190-3-Figure6-1.png", + "original_path": "designv11-35/openalex_figure/designv11_35_0000802_s41315-021-00190-3-Figure6-1.png", + "caption": "Fig. 6 Trajectory planning method with integration of MPI", + "texts": [ + " 1 3 \u2022 The \u201cMinimization\u201d procedure is called for calculating the minimum time trajectory P(t)MPI Optim for the whole sys- tem and with minimized cost function CostMPI Optim . The Simulated Annealing algorithm is used for finding the adequate trajectory minimizing the execution time of the task. The positions of control points related to both the mobile platform and the manipulator are modified during the minimization process. The positions of the control points of the mobile robot are modified (Pajak and Pajak 2017; Yu et\u00a0al. 2017) in the square spaces. The trajectories of the mobile robot are, thus, generated around the planned path. The figure (Fig.\u00a06)\u00a0illustrates the representation of the trajectory planning method. We note that the orientation is variable. The trajectories are generated in pose space P, are 1 3 then transformed to be defined in generalized space (Pajak and Pajak 2017). The final position [ Xend p , Yend p ] is maintained when MPI index is introduced as additional restriction. However, the orientation end p of the mobile robot and the configuration of the manipulator qend m are variable. The positions of the control points CP are variable" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_35_0000606_j.measurement.2021.109637-Figure1-1.png", + "original_path": "designv11-35/openalex_figure/designv11_35_0000606_j.measurement.2021.109637-Figure1-1.png", + "caption": "Fig. 1. Illustration of the ratchet mechanism with the visible gear and teeth: (a) ratchet ring, (b) pawl, (c) spring.", + "texts": [ + " [24] Alternative AI-based approaches used to diagnose mechanical systems include unsupervised methods, also with fuzzy sets theory (kmeans [17], fuzzy c-means [17], medoids of fuzzy sets [12] etc.). Also, k nearest neighbors [17] and tree-based approaches were applied. Diagnostics of ratchet mechanisms was not yet investigated, though their mathematical models were proposed [25]. Also, they are considered as important elements of larger structures [25]. This justifies experiments presented in the paper. The ratchet mechanisms are widely used in various mechanical systems, such as clocks, bridges, The mechanism consists of the gear or the toothed ring (Fig. 1). Inside there are non-symmetrical teeth and ratchets (or their assemblies). The latter must be constantly pushed towards the teeth surface, for instance by using the spring. The mechanism requires at least one ratchet to operate. Though there are versions revolving in both directions (like in socket wrenches), the type examined in this research allows only for the forward rotation. Mechanisms vary (depending on applications) in the diameter size (ranging from single mm in handheld watches to 80\u201390 cm in agricultural machines)" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_35_0001419_978-94-007-7401-8-Figure5.80-1.png", + "original_path": "designv11-35/openalex_figure/designv11_35_0001419_978-94-007-7401-8-Figure5.80-1.png", + "caption": "Fig. 5.80 3PaPPaR-1CPaPat-type fully-parallel PMs with uncoupled Sch\u00f6nflies motions defined by MF = SF = 4, RF\u00f0 \u00de \u00bc \u00f0v1; v2; v3;xb\u00de, TF = 0, NF = 30 limb topology C||Pa||Pat and Pa||P\\Pa\\kR, Pa||P\\Pa||R (a), Pa\\P\\\\Pa\\kR, Pa\\P\\\\Pa||R (b)", + "texts": [ + "2h) C||Pa||Pa (Fig. 5.40) 14. 2PaRRRR1PaRRR1CPaPat (Fig. 5.78) Pa\\R||R||R\\kR (Fig. 5.3d) Idem No. 3 Pa\\R||R||R (Fig. 5.2h) C||Pa||Pat (Fig. 5.4p) 15. 3PaPPaR1CPaPa (Fig. 5.79a) Pa||P\\Pa\\kR (Fig. 5.4a) The revolute joint between links 6D and 9 of G4 limb and the last revolute joints of limbs G1, G2 and G3 have superposed axes Pa||P\\Pa||R (Fig. 5.4d) C||Pa||Pa (Fig. 5.40) 16. 3PaPPaR1CPaPa (Fig. 5.79b) Pa\\P\\\\Pa\\kR (Fig. 5.4b) Idem No. 15 Pa\\P\\\\Pa||R (Fig. 5.4c) C||Pa||Pa (Fig. 5.40) 17. 3PaPPaR1CPaPat (Fig. 5.80a) Pa||P\\Pa\\kR (Fig. 5.4a) Idem No. 15 Pa||P\\Pa||R (Fig. 5.4d) C||Pa||Pat (Fig. 5.4p) 18. 3PaPPaR1CPaPat (Fig. 5.80b) Pa\\P\\\\Pa\\kR (Fig. 5.4b) Idem No. 15 Pa\\P\\\\Pa||R (Fig. 5.4c) C||Pa||Pat (Fig. 5.4p) (continued) 5.1 Fully-Parallel Topologies 535 Table 5.5 (continued) No. PM type Limb topology Connecting conditions 19. 3PaPaPR1CPaPa (Fig. 5.81) Pa\\Pa\\kP\\\\R (Fig. 5.4f) Idem No. 15 Pa\\Pa\\kP\\kR (Fig. 5.4h) C||Pa||Pa (Fig. 5.4o) 20. 3PaPaPR1CPaPa (Fig. 5.82) Pa\\Pa\\\\P\\\\R (Fig. 5.4g) Idem No. 15 Pa\\Pa\\\\P\\kR (Fig. 5.4e) C||Pa||Pa (Fig. 5.40) 21. 3PaPaPR1CPaPat (Fig. 5.83) Pa\\Pa\\kP\\\\R (Fig. 5.4f) Idem No. 15 Pa\\Pa\\kP\\kR (Fig" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_35_0001419_978-94-007-7401-8-Figure5.28-1.png", + "original_path": "designv11-35/openalex_figure/designv11_35_0001419_978-94-007-7401-8-Figure5.28-1.png", + "caption": "Fig. 5.28 2PaRRRR-1PaRRR-1RPPaPa-type fully-parallel PMs with uncoupled Sch\u00f6nflies motions defined by MF = SF = 4, RF\u00f0 \u00de \u00bc \u00f0v1; v2; v3;xb\u00de, TF = 0, NF = 19, limb topology Pa\\R||R||R\\kR, Pa\\R||R||R and R||P||Pa||Pa", + "texts": [ + "3c) The revolute joint between links 7D and 9 of G4 limb and the last revolute joints of limbs G1, G2 and G3 have superposed axes. The revolute joints between links 4 and 5 of limbs G1, G2 and G3 have orthogonal axes Pa\\R\\P\\kR (Fig. 5.2g) R||P||Pa||Pat (Fig. 5.4n) (continued) 5.1 Fully-Parallel Topologies 529 Table 5.2 (continued) No. PM type Limb topology Connecting conditions 21. 2PaRRRR-1PaRRR1RPaPaP (Fig. 5.27) Pa\\R||R||R\\kR (Fig. 5.3d) Idem No. 17 Pa\\R||R||R (Fig. 5.2h) R||Pa||Pa||P (Fig. 5.4k) 22. 2PaRRRR-1PaRRR1RPPaPa (Fig. 5.28) Pa\\R||R||R\\kR (Fig. 5.3d) Idem No. 20 Pa\\R||R||R (Fig. 5.2h) R||P||Pa||Pa (Fig. 5.4m) 23. 2PaRRRR-1PaRRR1RPaPatP (Fig. 5.29) Pa\\R||R||R\\kR (Fig. 5.3d) Idem No. 17 Pa\\R||R||R (Fig. 5.2h) R||Pa||Pat||P (Fig. 5.4l) 24. 2PaRRRR-1PaRRR1RPPaPat (Fig. 5.30) Pa\\R||R||R\\kR (Fig. 5.3d) Idem No. 20 Pa\\R||R||R (Fig. 5.2h) R||P||Pa||Pat (Fig. 5.4n) Table 5.3 Limb topology and connecting conditions of the fully-parallel solutions with no idle mobilities presented in Figs. 5.31, 5.32, 5.33, 5.34, 5.35, 5.36, 5" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_35_0000174_j.matpr.2021.01.640-Figure14-1.png", + "original_path": "designv11-35/openalex_figure/designv11_35_0000174_j.matpr.2021.01.640-Figure14-1.png", + "caption": "Fig. 14. Mode shapes 2.", + "texts": [ + "3411e-9 Engineering analysis Stress on Metal Matrix Composites as shown in Fig. 12.Stress analysis on metal matrix composite, von mises elastic theory is applied, the stress value obtained as maximum 74.301 MPa and minimum as 0.38088 MPa. etal Matrix Composites. By Modal analysis the characteristic frequencies and mode shapes shown in Fig. 13. In the event that the shat revolves at its characteristic recurrence, it very well may be seriously vibrated. The modal investigation performed to locate the normal frequencies Fig. 14 Fig. 15. Modal analysis in structural mechanics is determined the natural mode shapes and frequencies of an object or structure during free vibration. In this analysis the frequency varies 196.67 Hz, the mode shape level is 1. The mode shape can be extended on level 1, due to the rotations. The mode shapes level can be extend upto level 5. After applying boundary conditions, the modal analysis is to be perform. The free vibrations can be oscillated due to the drive shaft, the natural frequencies are obtained" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_35_0001563_9781119682684.ch1-Figure1.31-1.png", + "original_path": "designv11-35/openalex_figure/designv11_35_0001563_9781119682684.ch1-Figure1.31-1.png", + "caption": "Figure 1.31 SRM, (a) unaligned, (b) aligned position.", + "texts": [ + " Thus, by appropriately switching the phases, the machine rotates continuously in one direction. The major advantages of the SRM are their simple construction, high-power density and great efficiency. Since the rotor has no winding and no magnet, it is light in weight and there are no copper losses. Nowadays, SRM is mostly used in Electric Vehicle applications. Cross-sectional view of a 6/4 SRM machine is shown in Figure 1.30. The aligned and non-aligned positions in an 8/6 SRM machine is shown in Figure 1.31. SRM machines are rugged, reliable, have high-power density and require a simple unidirectional power converter. SRM requires rotor position sensor, has higher torque ripple with noisy machine operation. The current shape is non-sinusoidal. The machine construction is simple; however, the control is complex. A stepper motor is not a continuously rotating machine but rather it rotates in steps. It rotates by a specific number of degrees. A train of input current pulses are supplied and Figure 1.30 A 6/4 SRM motor" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_35_0000636_j.mechmachtheory.2021.104432-Figure4-1.png", + "original_path": "designv11-35/openalex_figure/designv11_35_0000636_j.mechmachtheory.2021.104432-Figure4-1.png", + "caption": "Fig. 4. Kinematic representation of the MLTA using body-fixed inline joints. (a) Overall kinematic model and hardpoint notation, (b) Twist-beam equivalent body isolated from the surroundings and with reference coordinate system R, (c) Isolated and rotated right twist-beam equivalent with parallel and perpendicular velocities at point Ur and (d) Transfer of the translational and rotational motion of the right twist-beam equivalent body to the right knuckle.", + "texts": [ + "2 these are iteratively integrated so that the trajectories of each hardpoint can be obtained and thus used in the optimisation process shown in Section 4. The given calculation approach is based on the works of T. Niessing and X. Fang Mechanism and Machine Theory 165 (2021) 104432 T. Niessing and X. Fang Mechanism and Machine Theory 165 (2021) 104432 Matschinsky and Schnelle carried out in [20,23,24] and applied to the mechanism of the MLTA. In contrast to the kinematic model shown in Fig. 3, Schnelle proposes a revolute joint instead of a cylindrical one located in the SC (see, also, Fig. 4(a)) [23]. With that, the former spherical joints used for both body mounts are replaced by two joints with a translational DOF in the lateral direction. In the further course of this work, these types of joints will be referred to as inline joints. This approach can be considered a more realistic representation of the elasto-kinematic deformations in the bushings and the cross beam during body roll [20]. Due to the assumptions made above, both joints always move symmetrically [23]. To prevent a lateral translation of the entire suspension, an additional in-plane joint is added to the SC as shown in Fig. 4(a). The goal of this approach is to compute the translational v k and rotational velocities \u03a9 k of the knuckle as well as the rigid twist-beam equivalent bodies. In the first step, the rigid equivalent bodies are isolated from the vehicle body and the knuckles (see Fig. 4(b)). Similar to [23], a reference coordinate system R is introduced. Its origin is located at the intersection point of the left trailing arm and cross beam. The corresponding yR axis is directed along the longitudinal axis of the cross beam and is coincident with the axis of the revolute joint. The inertial coordinate frame I is considered to be positioned in front of the vehicle with an x I axis that points opposite to the driving direction; the z I axis points upwards in the vertical direction", + " The drives of the mechanism are the rotations of the left and right halves of the revolute joint and can be defined as: I \u03c8\u0307 l/r = cl/r \u22c5 \u03c8\u0307 \u22c5 TIR RyR (2) where: I\u03c8\u0307 l/r Angular velocity for the left/right equivalent body with reference to the inertial frame I TIR Transformation matrix from reference to inertial frame I cl/r Constant for the left/right body to define the type of wheel travel (jounce, rebound, no travel values: \u2212 1,1,0) \u03c8\u0307 Magnitude of the rotational drive in rad time step The constants cl and cr can be considered as the DOFs of the mechanism. In the following, a single rebound travel of the right wheel will be discussed, therefore cl = 0 and cr = 1. To compute the motion of the left wheel, the index \"r\" and \"l\" in the upcoming equations can be exchanged with each other. Additionally, the vectors will be given in inertial frame I, and the index of the frame will be neglected for better readability. Fig. 4(c) shows the isolated mechanism with a torsionally rotated right twist-beam equivalent. The velocity of the isolated right body mount can be written as [23]: v Ur = \u03c8\u0307 r \u00d7 U r (3) It is obvious that the distance between the body mounts increases due to torsional rotation (see Fig. 4(c)). The velocity of the changing distance between the two body mounts can be found with projection of v Ur onto the translational axis of the body mounts which, in this case, also coincides with the connection vector from the left to the right body mounts r lr. The velocity can be written as [23]: v Ur,\u2016 = ( v Ur \u22c5 elr ) elr = d\u0307Ure lr (4) where: v Ur, \u2016 Velocity due to torsional rotation of the isolated right body mount projected onto the translational DOF e lr Unit vector from the left-to-right body mount d\u0307Ur Magnitude of the changing distance at the right body mount In the following, closing conditions will be formulated that allow a reintegration of the isolated bodies into their original boundaries. Since in the given scenario, the mounts are modelled as spherical joints with a translational DOF in the lateral direction, it can be found that the velocities at these inline joints cannot have proportions perpendicular to their translational DOFs (see Fig. 4(c)). To compensate for this, Schnelle [23] introduces an additional rotational vector \u03c9 B that represents the motion of the vehicle body relative to the suspension system. The body rotation is then superimposed on the torsional rotation of the isolated twist-beam equivalents so that the velocity of the body mount perpendicular to the translational DOF is eliminated. This leads to the following closing condition for the right side: 0 = v Ur, \u22a5 \u2212 \u03c9 B,r \u00d7 r lr = ( v Ur \u2212 v Ur, \u2016 ) \u2212 \u03c9 B,r \u00d7 r lr (5) where: v Ur, \u22a5 Velocity due to torsional rotation of the isolated right body mount perpendicular to the translational DOF of the body mounts This condition can also be interpreted as the result of two consecutive rotations", + " 4(b), (c) and (d)) can be computed as: v RU,r = ( \u03c8\u0307 r \u2212 \u03c9 B ) \u00d7 ( RU r \u2212 U r ) + v Ur, \u2016 = \u03a9r \u00d7 ( RU r \u2212 U r ) + v Ur, \u2016 (7) in which the velocity of the right body mount has been chosen as the translational reference velocity of the rigid body motion. In addition, the torsional rotation \u03c8\u0307 r as well as the relative rotation of the vehicle body \u03c9B are combined into the overall angular velocity of the right twist-beam equivalent body \u03a9r. This motion state can now be transferred to the knuckle (see Fig. 4(d)). Since the knuckle and the equivalent body are connected via a revolute joint RU, only the translational velocity vRU,r and the angular velocity \u03a9r,\u22a5 that is perpendicular to the axis of the revolute joint can be directly transferred to the knuckle. To compute the overall angular velocity of the right knuckle \u03a9K,r, the overall angular velocity \u03a9r will be projected onto the axis of the revolute joint: \u03a9r,\u2016 = ( \u03a9r \u22c5 eRU,r ) eRU,r \u03a9r,\u22a5 = \u03a9r \u2212 \u03a9r,\u2016 \u03c9RU, r = \u03c9Ru,r eRU,r \u03a9K,r = \u03c9RU, r + \u03a9r,\u22a5 (8) where: \u03a9 K,r Angular velocity of the right knuckle eRU,r Unit vector of the rotational axis of the right revolute joint RU\u03c9RU, r Angular velocity of the right knuckle about the revolute joint RU\u03a9r,\u2016 Overall angular velocity of the right twist-beam equivalent body parallel to the revolute joint \u03a9r,\u22a5 Overall angular velocity of the right twist-beam equivalent body perpendicular to the revolute joint The only unknown remains the magnitude \u03c9Ru,r which is the angular velocity about the revolute joint" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_35_0001411_978-3-030-55061-5-Figure12-1.png", + "original_path": "designv11-35/openalex_figure/designv11_35_0001411_978-3-030-55061-5-Figure12-1.png", + "caption": "Fig. 12. The semilogarithmic plot of the correspondence interval versus the accuracy of initial parameter", + "texts": [ + " Comparison between approximate and exact solutions for two different values of modified initial parameter For a set of values of the accuracy for the initial conditions there were found the values of the corresponding interval, considering a certain parameter (\u03b82) as objective. There were found the times for which the next condition is fulfilled: |\u03b8 \u20322 \u2212 \u03b82| \u2264 0.01 (9) The notation Bs is used in Fig. 11 for the order number of the subinterval up to which the condition (9) is satisfied; obviously, the corresponding interval is proportional to Bs. The values of Bs have been found for the range q = 1\u00f7 14 and in Fig. 12 it is presented the dependency of the corresponding interval on the accuracy of the initial data, using semilogarithmic plot. Chaos Illustrations in Dynamics of Mechanisms 303 From Fig. 12 it is suggested a quasi-linear variation in semilogarithmic plot and the possibility of approximating the points from the graph with a straight line. To be recalled that the graph fromFig. 12was traced for a dynamical systemwithwell specified structure, properties and initial conditions. From the plot, a first important conclusion emerges: while the corresponding time presents a linear variation, the accuracy for the stipulation of initial parameters increases exponentially. A second conclusion developing from the plot in Fig. 12 consists in the fact that with known accuracy of the initial experimental data measurements, the dimension of the corresponding interval can be found. Reciprocate, when the dimension of the corresponding interval is known the required accuracy for evaluating the initial data can be found, fact that represents valuable information both technically and economically. The paper aims to draw attention to the results that can be provided by the numerical modellingmethods used in system dynamics. Themain concern of thework is to identify a dynamic system for which an analytical model exists" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_35_0001603_301-Figure8-1.png", + "original_path": "designv11-35/openalex_figure/designv11_35_0001603_301-Figure8-1.png", + "caption": "Fig. 8. Disk machine", + "texts": [], + "surrounding_texts": [ + "The mechanical wear of metals\nshows clearly that there is a thin layer near the surface in which the crystalline structure remains almost unaffected.\nIn sliding, the friction superposes a tangential component on to the stress distribution and in these circumstances the shearing stress in the surface may become very close to the maximum shear It is then to be expected that the surface films will become disrupted and extensive intermetallic contact will occur. Experiment shows(10) that when the normal load is sufficient to cause extensive bulk flow, large scale, continuous rupture of the surfaces results. When the load is within the elastic limit, the stresses at some surface irregularity may, nevertheless, locally exceed the yield stress, and a small weld may form and then grow as sliding continues. From this it follows that the frictional forces at the onset of sliding are typical of contaminated surfaces, and that as sliding proceeds the coefficient of friction in general changes.\nDifferent frictional phenomena and different degrees of surface damage are therefore to be expected in mechanical systems in which the sliding distance is short, as with a pair of spur gears, from those in which it is long, as with a piston in a cylinder. The effect of sliding distance is conveniently studied in a friction apparatus in which the specimens are cylinders, one inclined at 45\u2019 to the direction of sliding, the other being mounted on a turntable (Fig. 3). The sliding distance is varied by rotating the turntable. Such an apparatus has been designed by Dr. Archard and results obtained with\nFriction records corresponding to the specimen arrangements shown at A , B and C. With A and C the sliding distance is long, wth,B it is short. Upper specimen 0.4% carbon steel; lower specimen 1 \u2018 5 ;4 nickel, 1 % chromium steel unlubricatcd. Speed 0.05 cmis. Loads: A , 10 kg; B, 7 . 5 to 25 kg: C. 10 kg.\nit show that the influence of the sliding distance can indeed be very great (Fig. 4). This, therefore, is a factor which must be borne in mind when attempting to design experiments intended to show the probable behaviour of a material in a particular type of machine. The dependence of the extent and nature of the surface damage on the distance of sliding is probably the main reason for the well-known difference between the behaviour of materials in rolling and in sliding. In pure sliding very effective lubrication must be provided, whereas rollers will often operate satisfactorily with hardly any lubrication at all. Moreover, the stresses which can safely be imposed on sliding elements are very small in comparison with those which can be carried by rollers.\nC O N T I N U O U S R U B B I N G I N C O N D I T I O N S O F P O I N T C O N T A C T\nTo maintain nominal point contact conditions indefinitely a fairly complicated apparatus is required. Dr. Archard and Mr. Tillen have designed a machine in which the test specimens are cylinders both rotating with their axes crossed at SO\u201d. Simultaneously, one cylinder reciprocates in a direction at 45\u201c to its axis (Fig. 5), the reciprocating motion being deliberately\nAA\u2019, upper specimen axis; BB\u2019, lower specimen axis; CC\u2019, fixed pivot axis, W, load.\nunrelated to the speed of rotation. Each point of the surface of each cylinder therefore rubs against the other, so that although the specimens wear they retain their cylindrical form. In this way point contact conditions are maintained and, since wear changes the diameter of the cylinders very little, the stresses are also effectively constant. The sliding distance can be varied by varying the speed of rotation of one cylinder with respect to the other.\nExploratory experiments with this apparatus show that at light loads, after careful running-in, the electrical contact resistance between lubricated specimens can be kery high, some tens of thousands of ohms. As the load is increased the oil film begins to break down locally giving temporary excursions of the contact resistance to low valnes, one ohm or less. Eventually at heavy loads low resistances below one ohm are observed all the time. At sufficiently heavy loads, the lubrication fails and the surfaces become heavily damaged; this is commonly termed \u201cscuffing.\u201d This general pattern of events is influenced by the speed; at low speeds the low resistance region extends over a considerable range of loads, but at higher speeds scuffing may follow almost immediately after the electrical resistance begins to fluctuate to low values. The variation of the scuffing load with speed is shown in\nBRITISH JOURKAL OF APPLIED PHYSICS - , . - 12: VOL. 9, APRIL 1958", + "running without the surfaces seizing. Examples of the flow produced are shown in Fig. 7 . The second point is that the very high electrical contact resistances which can be observed clearly denote some form of hydrodynamic lubrication. Simple hydrodynamic theory, assuming rigid bodies and an\nA, fixed cylinder; B, cylinder suspended from pivot C; D, gears determining slidelroll ratio; E, cardon shaft; F, fixed shaft,\nU1, U>, disks.\nMachines of this kind are often used for simulating the conditions at a chosen instant during the engagement of a pair of spur or helical gear teeth. In these gears pure rolling occurs at the pitch line, but above and below this line sliding also occurs, the direction of sliding reversing at the pitch line.\nFor many pears the mechanism of lubrication of gears has been a subject for argument. The practical fact that gears often show the original machining marks after many years of running suggests that they operate in conditions of hydrodynamic lubrication, and electrical contact resistance measurements using a disk machine(13) support this suggestion. Nevertheless, elementary hydrodynamic theory(I4) assuming rigid disks gives the following expression,\nwhere F is the load per unit face width, 17 is the viscosity of the oil, Y is the relative radius of curvature of the disks, u1 and I(? are their peripheral speeds and h, is the minimum film thickness. If values corresponding to practical operation are inserted. the calculated film thickness is usually about in.\n(b) Plastic flow in track. (e) Surface finish of track. (d) Original surface finish.\nor even less. The inaccuracies in manufackre exceed this value so that one might expect the surface irregularities to penetrate the films and the gears to operate under conditions of boundary lubrication.\nHowever, the inadequacy of the simple theory is shown by\noil Of constant viscosity, predicts that the load which can be supported in point contact conditions is given by(11)\nW = ~ . ~ T V R ( ~ ) 2R ' 0 the fact that it predicts the development of pressures in the\nBRITISH JOURNAL OF APPLIED PHYSICS 128 Voi. 9, APRIL 1958", + "oil film of the same order of magnitude as the elastic limit of the disk materials. Under such stresses, the disks would deform and the hydrodynamic forces therefore ought to be calculated for the deformed shape. Moreover, pressure changes the viscosity of oils and a complete mathematical analysis of the problem needs to allow for this also; it should, moreover, allow for the viscosity change due to the heat generated by the shearing of the oil film and for the conduction of some of this heat into the material of the disks. Very serious attempts have been made to give consideration to these factors,(\u20195) but no complete solution of this extremely complicated problem has yet been obtained.\nCrook(\u20196) has found recently that it is possible to measure the oil film thickness between the disks by a method which does not rest upon the uncertainties in these theories. If the capacitance between the disks is measured, the film thickness can be derived once the dielectric constant of the oil and the shape of the gap is known. The dielectric constant of a non-polar oil is not likely to be changed by the temperatures and pressures by more than say 20%; the main problem is that the shape of the gap is unknown. Fortunately, after passing between the conjunction of the disks, the oil film divides to form thin sheets of oil on the surface of each disk. Very lightly loaded plane pads can be made to float upon these films and the capacitance between these pads and the disks can be measured. Because there is no appreciable load: the pads remain sensibly undeformed; the shape is therefore known and the film thickness can be derived. Therefore, from the capacitance between the disks and the pads, a measure of the oil film thickness at the conjunction of the loaded disks may be derived. A check on the method is provided by measuring the capacitance between the disks themselves when they are loaded so lightly as to be sensibly undeformed. This also provides a means of determining for the first time the range of validity of Martin\u2019s elementary theory.\nDr. Crook\u2019s experiments show that as the load is increased the first departure from elementary theory arises from the increase in viscosity of the oil due to the pressures developing. Later the pressures rise sufficiently to cause the disks to deform noticeably and, finally, the oil film thickness becomes independent of load. In the last stage an increase in load increases the width of the flattened conjunction of the disks sufficiently to compensate for the tendency to diminish the oil film thickness. Fig. 9 shows the variation of film thickness with load for rolling conditions. A particularly interesiing feature of Dr. Crook\u2018s work is that he shows that the\u2019film thickness is\nThe rizechanical weay of metals\n9, APRIL 1958 I29 BRITISH JOURNAL OF APPLIED PHYSICS\nDisk diameter, 3 in.; peripheral speed 597 cm s-1; oil to Admiralty specification OMi00: inlet temperature of oil 50\u201c C;\nviscosity of inlet oil 0.4 P.\ninfluenced far more by changes in the overall temperature of the disks, than by changes in the temperature in the heavily loaded zone where the properties of the oil are.so abnormal. This suggests that the oil film thickness is primarily determined by the hydrodynamic phenomena at the entry side of the disks. If this were to be confirmed it might speed the development of an adequate theory of line lubrication because it is a t the entry side, before the oil drastically changes its properties, that the conditions are definable with least uncertainty.\nT H E L A W S O F U N L U B R I C A T E D W E A R A N D T H E I N F L U E N C E O F M A T E R I A L P R O P E R T I E S\nThe above experiments have also shown that hydrodynamic support up to heavy loads can only be obtained on carefully prepared or run-in specimens. If a heavy load is applied to a rough surface immediately after the onset of running, the subsequent events depend on the manner in which the materials wear. If the surface irregularities reduce in size as the experiment proceeds hydrodynamic conditions will ultimately be achieved. If this does not happen the usual consequence is early catastrophic failure. It is clear, therefore, that the achievement of adequate lubrication depends on the properties of the solid materials as well as on the properties of the oil.\nWear experiments in lubricated conditions are very sensitive to the conditions of lubrication, and it is easier to obtain a picture of the way in which material properties influence wear from experiments made in unlubricated conditions. Also, of course, the study of the wear of unlubricated materials has its own intrinsic importance. The apparatus needed is simpler and many useful conclusions can be drawn from experiments using a pin and ring apparatus such as is shown in Fig. 10. Two general types of wear can be observed, a\nA , rotating shaft; B, ring, s i n . diametei; C, flat-ended f in . diameter pin, pressed under load against ring.\nmild type and a severe type. In severe wear the surfaces remain metallic in appearance; there is extensive tearing, the wear fragments are of the order of tenths of a millimetre in size, and the structure of the rubbing metals is disrupted to a comparable depth. In mild wear, the surfaces often become discoloured, the wear fragments are usually smaller than one micron and, in a normal section, there is no distortion of the underlying structure visible in the optical microscope. Fuller descriptions of these two forms of wear have been given by Archard and Hirst.(\u201d)\nIn general, the change in the nature of the surfaces produced by rubbing causes the wear rate to change with time, until an equilibrium surface condition has been achieved,(\u2019*) but thereafter it becomes constant and independent of the apparent area of contact. This behaviour has now been" + ] + }, + { + "image_filename": "designv11_35_0000231_00368504211003383-Figure2-1.png", + "original_path": "designv11-35/openalex_figure/designv11_35_0000231_00368504211003383-Figure2-1.png", + "caption": "Figure 2. Structure diagram of picking manipulator.", + "texts": [ + " Kv, Kp represent the feedback gain vector of velocity and displacement respectively; D0(q) represent estimated symmetric definite inertia matrix; C0(q, _q) represent the vector of centripetal force torque and Coriolis force torque. G0(q) represent estimated gravity vector; t1, t2 represent the output torques of compute torque controller and fuzzy logic compensation controller respectively, e and _e are error vector and error rate vector respectively. _q and q are the actual velocity vector and displacement vector of the joints respectively. Simulation and analysis of trajectory tracking control Virtual prototype of picking manipulator The structure diagram of picking manipulator is shown in Figure 2, the structural parameters are shown in Table 1. In order to realize the effective combination of mechanical system analysis and control design simulation, the electromechanical joint simulation of Adams and MATLAB is adopted to form a closed data cycle between Adams and MATLAB. The joint simulation system platform is shown in Figure 3. Dof 7_input is the system function name of joint variable input, Dof 7_control is the system function name of control module. Adams sub is virtual prototype module imported from ADAMS to Matlab" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_35_0001419_978-94-007-7401-8-Figure3.13-1.png", + "original_path": "designv11-35/openalex_figure/designv11_35_0001419_978-94-007-7401-8-Figure3.13-1.png", + "caption": "Fig. 3.13 Overactuated PMs with coupled Sch\u00f6nflies motions of types 4PRPR (a) and 4RPPR (b) defined by MF \u00bc SF \u00bc 4; RF\u00f0 \u00de \u00bc v1; v2; v3;xd\u00f0 \u00de, TF = 0, NF = 6, limb topology P||R\\P\\kR (a) and R||P\\P\\kR (b)", + "texts": [ + "9b) R||P||R\\P (Fig. 3.1k) Idem No. 2 13. 4RPPR (Fig. 3.10a) R||P\\P\\kR (Fig. 3.1l) Idem No. 2 14. 4RPPR (Fig. 3.10b) R\\P\\kP||R (Fig. 3.1m) Idem No. 2 15. 4PRRP (Fig. 3.11a) P\\R||R||P (Fig. 3.1n) The last prismatic joints of the four limbs have parallel directions 16. 4PRPR (Fig. 3.11b) P\\R||P||R (Fig. 3.1o) Idem No. 9 17. 4PRRP (Fig. 3.12a) P||R||R\\P (Fig. 3.1v) The first prismatic joints of the four limbs have parallel directions 18. 4RPRP (Fig. 3.12b) R||P||R\\P (Fig. 3.1q) Idem No. 2 19. 4PRPR (Fig. 3.13a) P||R\\P\\kR (Fig. 3.1r) Idem No. 1 20. 4RPPR (Fig. 3.13b) R||P\\P\\kR (Fig. 3.1s) Idem No. 2 21. 4RPPR (Fig. 3.14a) R\\P\\kP||R (Fig. 3.1t) Idem No. 2 22. 4RPRP (Fig. 3.14b) R\\P\\kR||P (Fig. 3.1u) Idem No. 2 23. 4PPPR (Fig. 3.15a) P\\P\\\\P\\kR (Fig. 3.1h0) Idem No. 9 24. 4PPPR (Fig. 3.15b) P\\P\\\\P\\\\R (Fig. 3.1w) Idem No. 9 25. 4PPPR (Fig. 3.16a) P\\P\\\\P||R (Fig. 3.1x) Idem No. 9 26. 4PPRP (Fig. 3.16b) P\\P||R\\P (Fig. 3.1y) The revolute joints of the four limbs have parallel axes 27. 4PPRP (Fig. 3.17a) P\\P\\\\R||P (Fig. 3.1z) Idem No. 26 28. 4PPRP (Fig. 3.17b) P\\P\\kR\\P (Fig" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_35_0001563_9781119682684.ch1-Figure1.14-1.png", + "original_path": "designv11-35/openalex_figure/designv11_35_0001563_9781119682684.ch1-Figure1.14-1.png", + "caption": "Figure 1.14 Illustration of Faraday\u2019s law.", + "texts": [ + "53) where qoE is the magnitude of the force acting on the charge and 2\u03c0r is the distance over which that force acts. Or the Eq. (1.46) can be written as W = ( qoE ) (2\ud835\udf0br) = qo or = 2\ud835\udf0brE (1.54) Rewriting Eq. (1.53) W = \u222b F \u22c5 dl = qo \u222e E \u22c5 dl = qo or = \u222e E \u22c5 dl (1.55) The integral in Eq. (1.54) becomes Eq. (1.53) when Figure 1.13b is considered. According to Faraday\u2019s law or rewriting Eq. (1.37a) = \u2212d\ud835\udf19 dt Therefore, reformulated Faraday\u2019s law equation shall be \u222e C E \u22c5 dl = \u2212d\ud835\udf11 dt = \u2212 d dt\u222b S B \u22c5 dS (1.56) where S is the surface area enclosed by C as shown in Figure 1.14 Faraday\u2019s law can be stated as a time-varying magnetic field that gives rise to an electric field. Specifically, the electromotive force around a closed path C is equal to the negative of the time rate of increase of the magnetic flux enclosed by that path. Hence, Faraday\u2019s law can be applied to any closed path. In Figure 1.13d, it is shown that when there is a change in a magnetic field, then EMFs are induced: the inner most region \u20181\u2019 experiences inducement of emf = \u03b5, outside magnetic field region \u20183\u2019 emf = 0, and region \u20182\u2019 near the outer edge of the changing magnetic field emf< \u03b5" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_35_0000277_14644207211003321-Figure1-1.png", + "original_path": "designv11-35/openalex_figure/designv11_35_0000277_14644207211003321-Figure1-1.png", + "caption": "Figure 1. Experimental arrangement: (a) metal and polymer driver and driven gears; (b) in-house test setup; and (c) schematic illustration.", + "texts": [ + " Following the completion of tests, the teeth were subjected to failure analysis under optical and fieldemission scanning electron microscopes. In this study, involute spur gears were used as master (driver) and test (driven) gears. The polymer master and test gears were made by injection molding PA 66. The metal master gear was manufactured using the wire cut electric discharge process, and AISI 316 was the material. The fabricated master and test gears of symmetric and asymmetric configurations are shown in Figure 1(a). The symmetric gear was designed with a baseline pressure angle of 20 . In asymmetric gear teeth, the pressure angles of the flanks were 34 and 20 . The pressure angles of asymmetric gear were optimized based on the criteria of minimum allowable thickness at tooth tip and minimum bending stress. The rationale behind adopting these pressure angles is described in detail elsewhere.17 The gear configurations are described by referring to the drive side pressure angle and coast side pressure angle (left to right)", + " The measurement system comprised of two sensors: a torque transducer (HBM, T20WN), which determined the applied torque, and a rotary potentiometer (MCB, PR27M), which calculated the deflection of the tooth pair by measuring the angular displacement of the drive shaft. During the operation, the output signals emerging from the transducers were acquired by a data acquisition system (HBM, QuantumX\u2014MX840A) at a steady rate of 300 Hz. In addition, the rise in the gear surface temperature was measured by an infrared thermal imaging camera (Testo, 870-1). Figure 1(b) and (c) show the bending fatigue test rig. The servo motor and the master gear were interconnected, with the torque transducer in the middle. The rotary potentiometer was mounted on the drive gear shaft. The master gear acted as the driver gear, whereas the test gear was firmly attached to the stationary driven gear shaft. The alternate rotary motion of the driver gear exerted cyclic load on the test gear tooth. A locking plate was employed to arrest the rotation of the test gear shaft. Consequently, the sliding action was eliminated, and the test gear tooth was subjected to bending fatigue only" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_35_0001706_pime_proc_1950_163_014_02-Figure14-1.png", + "original_path": "designv11-35/openalex_figure/designv11_35_0001706_pime_proc_1950_163_014_02-Figure14-1.png", + "caption": "Fig. 14. Split Assembly", + "texts": [ + " That method had been used \u20acor rotating joints on aircraft t CLAYTON, D. 1945 Proc. I.Mech.E., vol. 153, p. 332, \u201cA Short Review of Surface Finish in Relation to Friction and Lubrication\u201d. at UNIV CALIFORNIA SANTA BARBARA on June 8, 2016pme.sagepub.comDownloaded from COMMUNICATIONS O N THE F R I C T I O N OF FLEXIBLE PACKINGS 109 turrets where it was essential to have smooth performance at low rotation speeds. Those joints were used to conduct the oil flow to and from the turret control valves, and a split assembly was indicated in Fig. 14. The increase in static friction with time of contact had been _ _ _ _ - known for some time; the relationship given in equation (5) I t was probable that the interstices of the seal surfaces became filled with correctly orientated graphite thus preventing direct adhesion, cohesion, or deformation of the points of contact, as had been suggested by Hunter (1944)$. Certain control valves incorporating formed seals had been tested after being stationary for four weeks under atmospheric conditions; no measurable increase in static friction was recorded when employing D " + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_35_0001419_978-94-007-7401-8-Figure6.18-1.png", + "original_path": "designv11-35/openalex_figure/designv11_35_0001419_978-94-007-7401-8-Figure6.18-1.png", + "caption": "Fig. 6.18 Maximaly regular fully-parallel PMs with Sch\u00f6nflies motions of types 2PPRRR1PPRR-1CPaPa (a) and 2PPRRR-1PPRR-1CPaPat (b) defined by MF = SF = 4, (RF) = (v1; v2; v3;xb), TF = 0, NF = 10, limb topology P\\P\\||R||R\\R, P\\P\\||R||R and C||Pa||Pa (a), C||Pa||Pat (b)", + "texts": [ + "23, 6.24, 6.25, 6.26 No. PM type Limb topology Connecting conditions 1. 3PPPR-1CPaPaP (Fig. 6.17a) P ?P\\\\P ?||R (Fig. 4.1a) P ?P\\\\P\\\\R (Fig. 4.1b) C||Pa||Pa (Fig. 5.4o) The last revolute joints of limbs G1, G2 and G3 and the revolute joint between links 6D and 8 have superposed axes. The actuated prismatic joints of limbs G1, G2 and G3 have orthogonal directions 2. 3PPPR-1CPaPatP (Fig. 6.17b) P ?P\\\\P ?||R (Fig. 4.1a) P ?P\\\\P\\\\R (Fig. 4.1b) CPa||Pat (Fig. 5.4p) Idem no. 1 3. 2PPRRR-1PPRR-1CPaPa (Fig. 6.18a) P ?P ?||R||R ?R (Fig. 4.2a) P ?P ?||R||R (Fig. 4.1d) C||Pa||Pa (Fig. 5.4o) Idem no. 1 (continued) 6.2 Fully-Parallel Topologies with Simple and Complex Limbs 613 Table 6.6 (continued) No. PM type Limb topology Connecting conditions 4. 2PPRRR-1PPRR-1CPaPat (Fig. 6.18b) P ?P ?||R||R ?R (Fig. 4.2a) P ?P ?||R||R (Fig. 4.1d) C||Pa||Pat (Fig. 5.4p) Idem no. 1 5. 2CPRR-1CPR-1CPaPa (Fig. 6.19a) C ?P ?||R ?R (Fig. 4.2e) C ?P ?||R (Fig. 4.1i) C||Pa||Pa (Fig. 5.4o) Idem no. 1 6. 2CPRR-1CPR-1CPaPat (Fig. 6.19b) C ?P ?||R ?R (Fig. 4.2e) C ?P ?||R (Fig. 4.1i) C||Pa||Pat (Fig. 5.4p) Idem no. 1 7. 2CRRR-1CRR-1CPaPa (Fig. 6.20a) C||R||R ?R (Fig. 4.2g) C||R||R (Fig. 4.1k) C||Pa||Pa (Fig. 5.4o) Idem no. 1 8. 2CRRR-1CRR-1CPaPat (Fig. 6.20b) C||R||R ?R (Fig. 4.2g) C||R||R (Fig. 4", + "2 for the nomenclature of structural parameters 616 6 Maximally Regular Topologies with Sch\u00f6nflies Motions Table 6.8 (continued) No. Structural parameter Solution Figures 6.15 and 6.16 Figure 6.17 21. MG4 4 4 22. (RF) (v1; v2; v3;xb) (v1; v2; v3;xb) 23. SF 4 4 24. rl 6 6 25. rF 20 18 26. MF 4 4 27. NF 10 12 28. TF 0 0 29. Pp1 j\u00bc1 fj 5 4 30. Pp2 j\u00bc1 fj 4 4 31. Pp3 j\u00bc1 fj 5 4 32. Pp4 j\u00bc1 fj 10 10 33. Pp j\u00bc1 fj 24 22 a See footnote of Table 2.2 for the nomenclature of structural parameters 6.2 Fully-Parallel Topologies with Simple and Complex Limbs 617 Table 6.9 (continued) No. Structural parameter Solution Figure 6.18 Figures 6.19 and 6.20 23. SF 4 4 24. rl 6 6 25. rF 20 20 26. MF 4 4 27. NF 10 10 28. TF 0 0 29. Pp1 j\u00bc1 fj 5 5 30. Pp2 j\u00bc1 fj 4 4 31. Pp3 j\u00bc1 fj 5 5 32. Pp4 j\u00bc1 fj 10 10 33. Pp j\u00bc1 fj 24 24 a See footnote of Table 2.2 for the nomenclature of structural parameters 618 6 Maximally Regular Topologies with Sch\u00f6nflies Motions Table 6.10 (continued) No. Structural parameter Solution Figures 6.21 and 6.22a Figure 6.22b 24. rl 9 9 25. rF 21 23 26. MF 4 4 27. NF 15 13 28. TF 0 0 29. Pp1 j\u00bc1 fj 7 8 30. Pp2 j\u00bc1 fj 7 7 31" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_35_0001419_978-94-007-7401-8-Figure5.5-1.png", + "original_path": "designv11-35/openalex_figure/designv11_35_0001419_978-94-007-7401-8-Figure5.5-1.png", + "caption": "Fig. 5.5 Complex limbs combining two parallelogram loops for fully-parallel PMs with uncoupled Sch\u00f6nflies motions defined by MG = SG = 5, RG\u00f0 \u00de \u00bc v1; v2; v3;xa;xd\u00f0 \u00de, rG = 6", + "texts": [ + "4a) The last joints of the four limbs have superposed axes/directionsPa||P\\Pa||R (Fig. 5.4d) R\\P\\\\P\\\\P (Fig. 5.1a) 2. 3PaPPaR- 1RPPP (Fig. 5.31b) Pa\\P\\\\Pa\\kR (Fig. 5.4b) Idem No. 1 Pa\\P\\\\Pa||R (Fig. 5.4c) R\\P\\\\P\\\\P (Fig. 5.1a) 3. 3PaPaPR1RPPP (Fig. 5.32a) Pa\\Pa\\kP\\\\R (Fig. 5.4f) Idem No. 1 Pa\\Pa\\kP\\kR (Fig. 5.4h) R\\P\\\\P\\\\P (Fig. 5.1a) 4. 3PaPaPR1RPPP (Fig. 5.32b) Pa\\Pa\\\\P\\\\R (Fig. 5.4g) Idem No. 1 Pa\\Pa\\\\P\\kR (Fig. 5.4e) R\\P\\\\P\\\\P (Fig. 5.1a) 5. 2PaPaRRR1PaPaRR1RPPP (Fig. 5.33) Pa\\Pa||R||R\\\\R (Fig. 5.5a) The last joints of the four limbs have superposed axes/directions. The revolute joints between links 7 and 8 of limbs G1, G2 and G3 have orthogonal axes Pa\\Pa||R||R (Fig. 5.4i) R\\P\\\\P\\\\P (Fig. 5.1a) 6. 2PaPaRRR1PaPaRR1RPPP (Fig. 5.34) Pa\\Pa||R||R\\kR (Fig. 5.5b) Idem No. 5 Pa\\Pa||R||R (Fig. 5.4j) R\\P\\\\P\\\\P (Fig. 5.1a) 7. 3PaPPaR1RUPU (Fig. 5.35a) Pa||P\\Pa\\kR (Fig. 5.4a) The first revolute joint of G4 limb and the last revolute joints of limbs G1, G2 and G3 have parallel axes. The last joints of of limbs G1, G2 and G3 have superposed axes Pa||P\\Pa||R (Fig. 5.4d) R\\R\\R\\P\\kR\\R (Fig. 5.1b) (continued) 530 5 Topologies with Uncoupled Sch\u00f6nflies Motions Table 5.3 (continued) No. PM type Limb topology Connecting conditions 8. 3PaPPaR1RUPU (Fig. 5.35b) Pa\\P\\\\Pa\\kR (Fig. 5.4b) Idem No. 7 Pa\\P\\\\Pa||R (Fig. 5.4c) R\\R\\R\\P\\kR\\R (Fig. 5.1b) 9. 3PaPaPR1RUPU (Fig. 5.36a) Pa\\Pa\\kP\\\\R (Fig. 5.4f) Idem No. 7 Pa\\Pa\\kP\\kR (Fig. 5.4h) R\\R\\R\\P\\kR\\R (Fig. 5.1b) 10. 3PaPaPR1RUPU (Fig. 5.36b) Pa\\Pa\\\\P\\\\R (Fig. 5.4g) Idem No. 7 Pa\\Pa\\\\P\\kR (Fig. 5.4e) R\\R\\R\\P\\kR\\R (Fig. 5.1b) 11. 2PaPaRRR1PaPaRR1RUPU (Fig. 5.37) Pa\\Pa||R||R\\\\R (Fig. 5.5a) The last joints of of limbs G1, G2 and G3 have superposed axes. The revolute joints between links 7 and 8 of limbs G1, G2 and G3 have orthogonal axes Pa\\Pa||R||R (Fig. 5.4i) R\\R\\R\\P\\kR\\R (Fig. 5.1b) 12. 2PaPaRRR1PaPaRR1RUPU (Fig. 5.38) Pa\\Pa||R||R\\kR (Fig. 5.5b) Idem No. 11 Pa\\Pa||R||R (Fig. 5.4j) R\\R\\R\\P\\kR\\R (Fig. 5.1b) Table 5.4 Limb topology and connecting conditions of the fully-parallel solutions with no idle mobilities presented in Figs. 5.39, 5.40, 5.41, 5.42, 5.43, 5.44, 5.45, 5.46, 5.47, 5.48, 5.49, 5.50, 5.51, 5.52, 5.53, 5.54, 5.55, 5.56, 5.57, 5.58, 5.59, 5.60, 5.61, 5.62, 5.63, 5.64, 5.65, 5.66, 5.67, 5.68 No. PM type Limb topology Connecting conditions 1. 3PaPPaR-1RPaPaP (Fig. 5.39a) Pa||P\\Pa\\kR (Fig. 5.4a) The last joints of the four limbs have superposed axes/directions Pa||P\\Pa||R (Fig", + " 3PaPaPR-1RPPaPa (Fig. 5.47) Pa\\Pa\\kP\\\\R (Fig. 5.4f) Idem No. 3 Pa\\Pa\\kP\\kR (Fig. 5.4h) R||P||Pa||Pa (Fig. 5.4m) 14. 3PaPaPR-1RPPaPa (Fig. 5.48) Pa\\Pa\\\\P\\\\R (Fig. 5.4g) Idem No. 3 Pa\\Pa\\\\P\\kR (Fig. 5.4e) R||P||Pa||Pa (Fig. 5.4m) 15. 3PaPaPR-1RPPaPat (Fig. 5.49) Pa\\PakP\\\\R (Fig. 5.4f) Idem No. 3 Pa\\Pa\\kP\\kR (Fig. 5.4h) R||P||Pa||Pat (Fig. 5.4n) 16. 3PaPaPR-1RPPaPat (Fig. 5.50) Pa\\Pa\\\\P\\\\R (Fig. 5.4g) Idem No. 3 Pa\\Pa\\\\P\\kR (Fig. 5.4e) R||P||Pa||Pat (Fig. 5.4n) 17. 2PaPaRRR-1PaPaRR1RPaPaP (Fig. 5.51) Pa\\Pa||R||R\\\\R (Fig. 5.5a) Idem No. 1 Pa\\Pa||R||R (Fig. 5.4i) R||Pa||Pa||P (Fig. 5.4k) (continued) 532 5 Topologies with Uncoupled Sch\u00f6nflies Motions Table 5.4 (continued) No. PM type Limb topology Connecting conditions 18. 2PaPaRRR-1PaPaRR1RPaPaP (Fig. 5.52) Pa\\Pa||R||R\\kR (Fig. 5.5b) Idem No. 1 Pa\\Pa||R||R (Fig. 5.4i) R||Pa||Pa||P (Fig. 5.4k) 19. 2PaPaRRR-1PaPaRR1RPaPatP (Fig. 5.53) Pa\\Pa||R||R\\\\R (Fig. 5.5a) Idem No. 1 Pa\\Pa||R||R (Fig. 5.4i) R||Pa||Pat||P (Fig. 5.4l) 20. 2PaPaRRR-1PaPaRR1RPaPatP (Fig. 5.54) Pa\\Pa||R||R\\kR (Fig. 5.5b) Idem No. 1 Pa\\Pa||R||R (Fig. 5.4i) R||Pa||Pat||P (Fig. 5.4l) 21. 2PaPaRRR-1PaPaRR1RPPaPa (Fig. 5.55) Pa\\Pa||R||R\\\\R (Fig. 5.5a) The revolute joint between links 7D and 10 of G4 limb and the last revolute joints of limbs G1, G2 and G3 have superposed axes Pa\\Pa||R||R (Fig. 5.4i) R||P||Pa||Pa (Fig. 5.4m) 22. 2PaPaRRR-1PaPaRR1RPPaPa (Fig. 5.56) Pa\\Pa||R||R\\kR (Fig. 5.5b) Idem No. 21 Pa\\Pa||R||R (Fig. 5.4i) R||P||Pa||Pa (Fig. 5.4m) 23. 2PaPaRRR-1PaPaRR1RPPaPat (Fig. 5.57) Pa\\Pa||R||R\\\\R (Fig. 5.5a) Idem No. 21 Pa\\Pa||R||R (Fig. 5.4i) R||P||Pa||Pat (Fig. 5.4n) 24. 2PaPaRRR-1PaPaRR1RPPaPat (Fig. 5.58) Pa\\Pa||R||R\\kR (Fig. 5.5b) Idem No. 21 Pa\\Pa||R||R (Fig. 5.4i) R||P||Pa||Pat (Fig. 5.4n) 25. 3PaPaPaR-1RPPP (Fig. 5.59) Pa\\Pa||Pa\\kR (Fig. 5.6a) Idem No. 1 Pa\\Pa||Pa||R (Fig. 5.6b) R\\P\\\\P\\\\P (Fig. 5.1a) 26. 3PaPaPaR-1RPPP (Fig. 5.60) Pa\\Pa||Pa\\\\R (Fig. 5.6c) Idem No. 1 Pa\\Pa||Pa||R (Fig. 5.6b) R\\P\\\\P\\\\P (Fig. 5.1a) 27. 3PaPaPaR-1RUPU (Fig. 5.61) Pa\\Pa||Pa\\kR (Fig. 5.6a) The first revolute joint of G4 limb and the last revolute joints of limbs G1, G2 and G3 have parallel axes. The last joints of of limbs G1, G2 and G3 have superposed axes Pa\\Pa||Pa||R (Fig", + " 3PaPaPR1CPaPa (Fig. 5.81) Pa\\Pa\\kP\\\\R (Fig. 5.4f) Idem No. 15 Pa\\Pa\\kP\\kR (Fig. 5.4h) C||Pa||Pa (Fig. 5.4o) 20. 3PaPaPR1CPaPa (Fig. 5.82) Pa\\Pa\\\\P\\\\R (Fig. 5.4g) Idem No. 15 Pa\\Pa\\\\P\\kR (Fig. 5.4e) C||Pa||Pa (Fig. 5.40) 21. 3PaPaPR1CPaPat (Fig. 5.83) Pa\\Pa\\kP\\\\R (Fig. 5.4f) Idem No. 15 Pa\\Pa\\kP\\kR (Fig. 5.4h) C||Pa||Pat (Fig. 5.4p) 22. 3PaPaPR1CPaPat (Fig. 5.84) Pa\\Pa\\\\P\\\\R (Fig. 5.4g) Idem No. 15 Pa\\Pa\\\\P\\kR (Fig. 5.4e) C||Pa||Pat (Fig. 5.4p) 23. 2PaPaRRR1PaPaRR1CPaPa (Fig. 5.85) Pa\\Pa||R||R\\\\R (Fig. 5.5a) The revolute joint between links 6D and 10 of G4 limb and the last revolute joints of limbs G1, G2 and G3 have superposed axes Pa\\Pa||R||R (Fig. 5.4i) C||Pa||Pa (Fig. 5.40) 24. 2PaPaRRR1PaPaRR1CPaPa (Fig. 5.86) Pa\\Pa||R||R\\kR (Fig. 5.5b) Idem No. 23 Pa\\Pa||R||R (Fig. 5.4i) C||Pa||Pa (Fig. 5.40) 25. 2PaPaRRR1PaPaRR1CPaPat (Fig. 5.87) Pa\\Pa||R||R\\\\R (Fig. 5.5a) Idem No. 23 Pa\\Pa||R||R (Fig. 5.4i) C||Pa||Pat (Fig. 5.4p) 26. 2PaPaRRR1PaPaRR1CPaPat (Fig. 5.88) Pa\\Pa||R||R\\kR (Fig. 5.5b) Idem No. 23 Pa\\Pa||R||R (Fig. 5.4i) C||Pa||Pat (Fig. 5.4p) 27. 3PaPaPaR1CPaPa (Fig. 5.89) Pa\\Pa||Pa\\kR (Fig. 5.6a) The revolute joint between links 6D and 11 of G4 limb and the last revolute joints of limbs G1, G2 and G3 have superposed axes Pa\\Pa||Pa||R (Fig. 5.6b) C||Pa||Pa (Fig. 5.4o) 28. 3PaPaPaR1CPaPa (Fig. 5.90) Pa\\Pa||Pa\\\\R (Fig. 5.6c) Idem No. 27 Pa\\Pa||Pa||R (Fig. 5.6b) C||Pa||Pa (Fig. 5.54o) 29. 3PaPaPaR1CPaPat (Fig. 5.91) Pa\\Pa||Pa\\kR (Fig. 5.6a) Idem No. 27 Pa\\Pa||Pa||R (Fig. 5.6b) C||Pa||Pat (Fig" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_35_0000239_j.dt.2021.03.026-Figure6-1.png", + "original_path": "designv11-35/openalex_figure/designv11_35_0000239_j.dt.2021.03.026-Figure6-1.png", + "caption": "Fig. 6. Rubber tracked assembly.", + "texts": [ + " Besides, to reduce the impact load when the steering cylinder reaches the limit position, elastic limit devices are installed at both ends of the steering cylinder. Fig. 4. Multifunctional special vehicle. Fig. 5. Main steering system and auxiliary steering system. This rubber track assembly provides the vehicle with highspeed driving ability. It is mainly used as a rubber track drive system for utility vehicles. The system can improve the maneuverability of the utility vehicle so that the utility vehicle can quickly reach the workspace to carry out operations. As shown in Fig. 6, the rubber track assembly includes a driving wheel, tension wheel, roller wheels, tension arm, and tension spring. The driving wheels are located on top of the triangle track. The driving wheel is connected to the wheel-side reducer driving the axle. The driving wheel drives the rubber track through driving teeth. The tension wheel is located at the front of the rubber track assembly, and the tension arm and tension elastic device is used to tension the rubber track. The lower part of the rubber track assembly includes two sets of equal load wheels", + " Due to the small steering angle, the steering angle error caused by Ackerman steering is ignored. The value s0 \u00bc Mm=Dq is obtained as 5393.7 Nm/ . Table 2 Friction coefficient of pavements. Type of pavement Values Initial Static friction coefficient mS Initial Coulomb friction coefficient mC Concrete or asphalt (dry) 0.85 0.75 Asphalt (wet) 0.6 0.525 Concrete (wet) 0.8 0.7 Gravel 0.6 0.55 Dirt road (dry) 0.68 0.65 The input of this paper is the real-time curve of the left and right rubber track steering angles. During the in-situ steering experiment, the test system in Fig. 6 was used to test and record the steering angle of the left and right rubber track. The recorded data is shown in Fig. 8(a). The solid black line indicates the steering angle of the left track. The red dotted line indicates the steering angle of the right track. The turning process can be divided into 15 stages, namely ta-tb, tb-tc, \u2026to-tp. In the interval of tc-td-te, tf-tgth, ti-tj-tk and tl-tm-tn in Fig. 9, the steering angle gradually approaches the limit angle and remains stable. At this time, the stroke of the steering cylinder is close to the displacement limit of the cylinder" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_35_0001419_978-94-007-7401-8-Figure2.30-1.png", + "original_path": "designv11-35/openalex_figure/designv11_35_0001419_978-94-007-7401-8-Figure2.30-1.png", + "caption": "Fig. 2.30 4PPPaR-type fully-parallel PMs with coupled Sch\u00f6nflies motions defined by MF \u00bc SF \u00bc 4; RF\u00f0 \u00de \u00bc \u00f0v1; v2; v3;xd\u00de, TF = 0, NF = 18, limb topology P\\P\\kPa\\kR (a) and P\\P||Pa\\\\R (b)", + "texts": [ + "21c) The second and the last joints of the four limbs have parallel axes 4. 4PPPaR (Fig. 2.27b) P\\P\\kPa\\R (Fig. 2.21d) The last revolute joints of the four limbs have parallel axes 5. 4PRPPa (Fig. 2.28a) P||R\\P||Pa (Fig. 2.21e) The second joints of the four limbs have parallel axes 6. 4PPRPa (Fig. 2.28b) P\\P||R\\Pa (Fig. 2.21f) The third joints of the four limbs have parallel axes 7. 4PRPPa (Fig. 2.29a) P\\R||P\\Pa (Fig. 2.21g) Idem No. 5 8. 4PRPaP (Fig. 2.29b) P\\R\\Pa \\kP (Fig. 2.21h) Idem No. 5 9. 4PPPaR (Fig. 2.30a) P\\P\\kPa\\kR (Fig. 2.21i) Idem No. 4 10. 4PPPaR (Fig. 2.30b) P\\P||Pa\\R (Fig. 2.21j) Idem No. 4 11. 4PPaRP (Fig. 2.31a) P||Pa\\R||P (Fig. 2.21k) The before last joints of the four limbs have parallel axes 12. 4PPaPR (Fig. 2.31b) P||Pa\\P|| R (Fig. 2.21l) Idem No. 4 13. 4PPaPR (Fig. 2.32a) P\\Pa||P\\\\R (Fig. 2.21m) Idem No. 4 14. 4PPaRP (Fig. 2.32b) P\\Pa\\kR\\kP (Fig. 2.21n) Idem No. 4 15. 4RPPaP (Fig. 2.33a) R\\P||Pa\\kP (Fig. 2.21o) The first revolute joints of the four limbs have parallel axes 16. 4RPPPa (Fig. 2.33b) R||P\\P||Pa (Fig. 2.21p) Idem No. 15 17. 4PaPPR (Fig" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_35_0000498_2050-7038.12961-Figure9-1.png", + "original_path": "designv11-35/openalex_figure/designv11_35_0000498_2050-7038.12961-Figure9-1.png", + "caption": "FIGURE 9 Magnetic field density distribution plots of (A) 4p24s, (B) 4p36s, and (C) 4p42s machine models under the saturation limit of 0.65T", + "texts": [ + " The fundamental component generates the main stator field, which is rotating in nature, while the third-harmonic component generates the harmonic field which is pulsating in nature. The investigated machine models are operated at a constant speed of 1800 rpm and the simulations are performed for 1.4 seconds. The armature voltages of the investigated machine models when they are supplied with a three-phase fundamental current of 1.3 A (peak) and a single-phase third-harmonic current of 5 A (peak) through an arrangement discussed above are shown in Figure 8. The magnetic field density distribution plots of 4p24s, 4p36s, and 4p42s machine models are shown in Figure 9. These plots indicate that the operation of machine models employing the proposed simplified harmonic field-excitation scheme is under the saturation level of 0.65T. The flux linkages of the investigated machine models are shown in Figure 10. The magnitude of flux linkages for the 4p24s machine model under the steady-state operation is 0.153 Wb (peak), whereas the peak values of the flux linkages for 4p36s and 4p42s machine models are 0.351 and 0.441 Wb, respectively. The pulsating magnetic field caused by the harmonic current of the stator winding induces the EMF in the harmonic winding of the specially designed rotor which is later rectified using the full-bridge diode rectifier to supply the direct field-excitation current to the field winding of the rotor" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_35_0001419_978-94-007-7401-8-Figure3.18-1.png", + "original_path": "designv11-35/openalex_figure/designv11_35_0001419_978-94-007-7401-8-Figure3.18-1.png", + "caption": "Fig. 3.18 4PRPP-type overactuated PMs with coupled Sch\u00f6nflies motions defined by MF \u00bc SF \u00bc 4; RF\u00f0 \u00de \u00bc v1; v2; v3;xd\u00f0 \u00de, TF = 0, NF = 6, limb topology P\\R||P\\P (a) and P\\R\\P\\kP (b)", + "texts": [ + " 4RPPR (Fig. 3.14a) R\\P\\kP||R (Fig. 3.1t) Idem No. 2 22. 4RPRP (Fig. 3.14b) R\\P\\kR||P (Fig. 3.1u) Idem No. 2 23. 4PPPR (Fig. 3.15a) P\\P\\\\P\\kR (Fig. 3.1h0) Idem No. 9 24. 4PPPR (Fig. 3.15b) P\\P\\\\P\\\\R (Fig. 3.1w) Idem No. 9 25. 4PPPR (Fig. 3.16a) P\\P\\\\P||R (Fig. 3.1x) Idem No. 9 26. 4PPRP (Fig. 3.16b) P\\P||R\\P (Fig. 3.1y) The revolute joints of the four limbs have parallel axes 27. 4PPRP (Fig. 3.17a) P\\P\\\\R||P (Fig. 3.1z) Idem No. 26 28. 4PPRP (Fig. 3.17b) P\\P\\kR\\P (Fig. 3.1a0) Idem No. 26 29. 4PRPP (Fig. 3.18a) P\\R||P\\P (Fig. 3.1b0) Idem No. 26 30. 4PRPP (Fig. 3.18b) P\\R\\P\\kP (Fig. 3.1c0) Idem No. 26 31. 4PRPP (Fig. 3.19a) P||R\\P\\\\P (Fig. 3.1d0) Idem No. 26 32. 4RPPP (Fig. 3.19b) R\\P\\\\P\\\\P (Fig. 3.1e0) Idem No. 26 33. 4RPPP (Fig. 3.20a) R\\P\\kP\\\\P (Fig. 3.1f0) Idem No. 26 34. 4RPPP (Fig. 3.20b) R||P\\P\\\\P (Fig. 3.1g0) Idem No. 26 3.1 Topologies with Simple Limbs 243 Table 3.2 Limb topology and connecting conditions of the overactuated solutions with no idle mobilities presented in Figs. 3.21, 3.22, 3.23, 3.24, 3.25, 3.26, 3.27, 3.28, 3.29, 3.30, 3.31, 3.32, 3" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_35_0000034_detc2014-35384-Figure2-1.png", + "original_path": "designv11-35/openalex_figure/designv11_35_0000034_detc2014-35384-Figure2-1.png", + "caption": "Figure 2: Designed configuration of the proposed hybrid system", + "texts": [ + " The offset contour will be used for mask image generation which is presented in section 4. An efficient and robust contour to image conversion algorithm is proposed. In section 5, we presented the physical experiments for proof-of-concept verification. The conclusion and future work was discussed in section 6. The proposed hybrid system is a combination of vector scanning and mask projection as shown in Figure 1. Thus, it includes all the components used in the two subsystems. A simplified schematic drawing is presented in Figure 2. The developed system consists of laser source, projection system, optics, Z-axis elevator, and recoating system. In order to reduce the total cost of the system, an off-the-shelf LED portable projector (< $1000) with native resolution of 1024x768 is used for generating dynamic mask image. The original RGB LED is not strong enough to cure the photopolymer, thus a high power near-UV LED (wavelength = 405nm) is used to replace the original visible-light LED. Instead of using the expensive UV laser as the commercial SLA system, an inexpensive (<$100) but much more powerful (2Watt) Blu-ray laser diode is used as laser source" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_35_0000565_978-3-030-73882-2_165-Figure15-1.png", + "original_path": "designv11-35/openalex_figure/designv11_35_0000565_978-3-030-73882-2_165-Figure15-1.png", + "caption": "Fig. 15 The schematic of an H-bridge in EasyEDA software", + "texts": [ + " Figure 13 shows the schematic of our proposed Fire Safety System in EasyEDA software. And Fig. 14 shows the schematic of the power Supplies system, Voltage Sensor and input pins in same software. We have used the LM317T regulator in schematic to adapt the input power supply voltage that equal 24 V to an adapted voltage at 9 Vwhich used to power the Arduino Uno Card. And also the voltage sensor used to measure the voltage at the resistor terminal. In the power part, the siren and the polarity resistor located in the middle of the H-bridge as illustrated in Fig. 15. In the control part, the Arduino Uno Card with the connection connectors and all the components as well as the output pin and also the part who declares to the proposed Fire Safety System if there is a fire and the status of the siren, are illustrated in Fig. 16. Design and Realization of Fire Safety System \u2026 441 Fig. 16 The schematic of the control part in EasyEDA software We have used a 24 V DC power supply for the siren which powered during a fire. When there is no fire the voltage sensor will measure a zero voltage which means a disconnection for the proposed Fare Safety system" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_35_0001419_978-94-007-7401-8-Figure2.5-1.png", + "original_path": "designv11-35/openalex_figure/designv11_35_0001419_978-94-007-7401-8-Figure2.5-1.png", + "caption": "Fig. 2.5 Fully-parallel PMs with coupled Sch\u00f6nflies motions of types 4RRRPR (a) and 4RRPRR (b) defined by MF \u00bc SF \u00bc 4; RF\u00f0 \u00de \u00bc \u00f0v1; v2; v3;xd\u00de, TF = 0, NF = 2, limb topology R||R\\R\\P\\||R (a) and R||R||P\\R||R (b)", + "texts": [ + "1b) The first, the second and the last revolute joints of the four limbs have parallel axes 3. 4RRRRR (Fig. 2.3a) R||R||R\\R||R (Fig. 2.1c) The two last revolute joints of the four limbs have parallel axes 4. 4RRRRR (Fig. 2.3b) R||R||R\\R||R (Fig. 2.1c) The three first revolute joints of the four limbs have parallel axes 5. 4RRRRR (Fig. 2.4a) R||R\\R||R||R (Fig. 2.1d) The two first revolute joints of the four limbs have parallel axes. 6. 4PRRRR (Fig. 2.4b) P\\R\\R||R||R (Fig. 2.1e) The second joints of the four limbs have parallel axes 7. 4RRRPR (Fig. 2.5a) R||R\\R\\P\\kR (Fig. 2.1f) Idem No. 5 8. 4RRPRR (Fig. 2.5b) R||R||P\\R||R (Fig. 2.1g) Idem No. 5 9. 4PRRRR (Fig. 2.6a) P\\R\\R||R\\R (Fig. 2.1h) The second and the last joints of the four limbs have parallel axes 10. 4PRRRR (Fig. 2.6b) P||R\\R||R\\R (Fig. 2.1i) Idem No. 9 11. 4RRPRR (Fig. 2.7a) R\\R\\P\\kR\\R (Fig. 2.1j) Idem No. 1 12. 4PRRRR (Fig. 2.7b) P\\R||R\\R||R (Fig. 2.1k) Idem No. 3 13. 4PRRRR (Fig. 2.8a) P||R||R||R||R\\R (Fig. 2.1l) The last revolute joints of the four limbs have parallel axes 14. 4RPRRR (Fig. 2.8b) R||P||R||R\\R (Fig. 2.1m) Idem No. 13 15. 4RRPRR (Fig" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_35_0001385_978-981-15-6095-8-Figure4.8-1.png", + "original_path": "designv11-35/openalex_figure/designv11_35_0001385_978-981-15-6095-8-Figure4.8-1.png", + "caption": "Fig. 4.8 Sketch of effective parameters on the shear characteristics of geocell\u2013grains interface (Tavakoli Mehrjardi and Motarjemi 2018)", + "texts": [ + " \u2013 geocells mobilize an apparent cohesion on the shear interface owing to the provision of some confinement for the aggregates located in the neighbor of the shear plane. For geocell-reinforced samples with Dr = 50%, the apparent cohesion has substantially increased by about 1.9\u201323 kPa. \u2013 The results clarify that among the studied variables, geocell with cell aspect ratio [the ratio of the geocell\u2019s cells size (b) to the medium grains size (D50)] 4 has the best performance in the improvement of interface\u2019s shear strength. Moreover, to observe the effective parameters on the shear characteristics of geocell\u2013grains interface, Fig. 4.8 is illustrated. From this figure, during shearing, interlocking effect which mobilizes apparent cohesion and friction at the interface, 84 G. Tavakoli Mehrjardi and S. N. Moghaddas Tafreshi besides the confinement effect on grains within the geocell\u2019s cells, producing interface\u2019s shear strength. Based on the acquired results, it was found out that shear strength of the interface encountered weakness in the aftermath of grain sliding alongside the geocell\u2019s walls and also, geocell\u2019s walls distortion" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_35_0001419_978-94-007-7401-8-Figure6.23-1.png", + "original_path": "designv11-35/openalex_figure/designv11_35_0001419_978-94-007-7401-8-Figure6.23-1.png", + "caption": "Fig. 6.23 3PPPaR-1RUPU-type maximaly regular fully-parallel PMs with Sch\u00f6nflies motions defined by MF = SF = 4, (RF) = (v1; v2; v3;xb), TF = 0, NF = 13, limb topology P\\P\\||Pa||R, R\\R\\R\\P\\||R\\R and P\\P\\||Pa\\||R (a), P\\P\\||Pa ??R (b)", + "texts": [ + "1a) The last joints of the four limbs have superposed axes/ directions. The actuated prismatic joints of limbs G1, G2 and G3 have orthogonal directions 10. 3PPPaR-1RPPP (Fig. 6.21b) P ?P ?||Pa\\\\R (Fig. 4.8a) P ?P ?||Pa||R, (Fig. 4.8c) R\\\\P\\\\P\\\\P (Fig. 5.1a) Idem no. 9 11. 3PPaPR-1RPPP (Fig. 6.22a) P||Pa ?P\\\\R (Fig. 4.8d) P||Pa ?P ?||R (Fig. 4.8e) R ?P\\\\P\\\\P (Fig. 5.1a) Idem no. 9 12. 2PPaRRR-1PPaRR-1RPPP (Fig. 6.22b) P||Pa||R||R ?R (Fig. 4.9a) P||Pa||R||R (Fig. 4.8g) R ?P\\\\P\\\\P (Fig. 5.1a) Idem no. 9 13. 3PPPaR-1RUPU (Fig. 6.23a) P ?P ?||Pa ?||R (Fig. 4.8b) P ?P ?||Pa||R, (Fig. 4.8c) R ?R ?R ?P ?||R ?R (Fig. 5.1b) The first revolute joint of G4-limb and the last last joints of limbs G1, G2 and G3 have parallel axes. The last revolute joints of limbs G1, G2 and G3 have superposed axes. The actuated prismatic joints of limbs G1, G2 and G3 have orthogonal directions (continued) 614 6 Maximally Regular Topologies with Sch\u00f6nflies Motions Table 6.6 (continued) No. PM type Limb topology Connecting conditions 14. 3PPPaR-1RUPU (Fig. 6.23b) P ?P ?||Pa\\\\R (Fig. 4.8a) P ?P ?||Pa||R, (Fig. 4.8c) R ?R ?R ?P ?||R ?R (Fig. 5.1b) Idem no. 13 15. 3PPaPR-1RUPU (Fig. 6.24a) P||Pa ?P\\\\R (Fig. 4.8d) P||Pa ?P ?||R (Fig. 4.8e) R ?R ?R ?P ?||R ?R (Fig. 5.1b) Idem no. 13 16. 2PPaRRR-1PPaRR-1RUPU (Fig. 6.24b) P||Pa||R||R ?R (Fig. 4.9a) P||Pa||R||R (Fig. 4.8g) R ?R ?R ?P ?||R ?R (Fig. 5.1b) Idem no. 13 17. 3PPaPaR-1RPPP (Fig. 6.25) P||Pa||Pa ?R (Fig. 4.10a) P||Pa||Pa||R (Fig. 4.10b) R ?P\\\\P\\\\P (Fig. 5.1a) Idem no. 9 18. 3PPaPaR-1RUPU (Fig. 6.26) P||Pa||Pa " + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_35_0001419_978-94-007-7401-8-Figure2.69-1.png", + "original_path": "designv11-35/openalex_figure/designv11_35_0001419_978-94-007-7401-8-Figure2.69-1.png", + "caption": "Fig. 2.69 4PRPaPa-type fully-parallel PM with coupled Sch\u00f6nflies motions defined by MF \u00bc SF \u00bc 4; RF\u00f0 \u00de \u00bc \u00f0v1; v2; v3;xd\u00de, TF = 0, NF = 30, limb topology P||R\\Pa\\\\Pa", + "texts": [ + " 4 49. 4PaRPaP (Fig. 2.61) Pa\\R\\Pa\\kR (Fig. 2.22v) Idem No. 34 50. 4PaRPPa (Fig. 2.62) Pa\\R||P\\Pa (Fig. 2.22w) Idem No. 34 51. 4PaPRPa (Fig. 2.63) Pa\\P||R\\Pa (Fig. 2.22x) Idem No. 33 52. 4PaPaPR (Fig. 2.64) Pa\\Pa\\\\P||R (Fig. 2.22y) Idem No. 4 53. 4PaPaRP (Fig. 2.65) Pa\\Pa\\\\R||P (Fig. 2.22z) Idem No. 41 54. 4RPaPaP (Fig. 2.66) R\\Pa\\\\Pa\\\\P (Fig. 2.22a0) Idem No. 15 55. 4RPaPPa (Fig. 2.67) R\\Pa\\kP\\\\Pa (Fig. 2.22b0) Idem No. 15 56. 4RPPaPa (Fig. 2.68) R||P\\Pa\\\\Pa (Fig. 2.22c0) Idem No. 15 57. 4PRPaPa (Fig. 2.69) P||R\\Pa\\\\Pa (Fig. 2.22d0) Idem No. 5 58. 4PaCPa (Fig. 2.70) Pa\\C\\Pa (Fig. 2.22e0) The axes of the cylindrical joints of the four limbs are parallel 59. 4CPaPa (Fig. 2.71) C\\Pa\\\\Pa (Fig. 2.22f 0) Idem No. 58 60. 4PaPaC (Fig. 2.72) Pa\\Pa\\\\C (Fig. 2.22g0) Idem No. 58 176 2 Fully-Parallel Topologies with Coupled Sch\u00f6nflies Motions T ab le 2. 4 L im b to po lo gy an d co nn ec ti ng co nd it io ns of th e fu ll y- pa ra ll el so lu ti on s w it h no id le m ob il it ie s pr es en te d in F ig s. 2. 73 , 2" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_35_0000344_j.cie.2021.107366-Figure2-1.png", + "original_path": "designv11-35/openalex_figure/designv11_35_0000344_j.cie.2021.107366-Figure2-1.png", + "caption": "Fig. 2. Graphical illustration of the information obtained from the machine vision system.", + "texts": [ + ", the lowest priority level, is utilized to conduct the surface treatment on a particular part of the workpiece by means of a manual or an automatic mode of operation. The following input information is considered for these levels: the robot pose p and the robot configuration and its derivative {q, q\u0307}, which are obtained from the robot controller; the force vector F obtained from the guidance sensor, which is located at the robot tool; the data {d, n} obtained using machine vision, where d represents the length of the vector from the robot tool, see Fig. 2, to the nearest point of the workpiece, whereas n denotes the unit vector of the mentioned vector (note that n is normal to the workpiece surface as long as it is smooth at the nearest point to the tool); and the reference pref for the tool position p = [ x y z ]T. The equation Aix = bi (4) for each priority level is obtained below, where x corresponds to the commanded acceleration q\u0308c for the robot system. The errors of these equations are minimized using (5) and (6), as shown in Fig. 1. Thus, the acceleration command q\u0308c,3 is double integrated to get the robot configuration command qc", + " (9), (10) and (12), the control equation for Level 1 results in: [ pos(\u03d5d) 0 0 pos(\u03d5b) ] Lg\u03d51 q\u0308c = \u2212 [ pos(\u03d5d) pos(\u03d5b) ] u+ 1 , \u2192A1q\u0308c = b1, (17) where u+ 1 represents switching gain of the SMC, b1 and A1 denote the vector and matrix for the control equation of Level 1 and, according to (10)\u2013(16), matrix Lg\u03d51 is given by: Lg\u03d51 = [ (\u2202\u03d5d/\u2202x)T (\u2202\u03d5b/\u2202x)T ] g = [ (\u2202\u03d5d/\u2202q\u0307)T (\u2202\u03d5b/\u2202q\u0307)T ] = [ \u2212 Kd1 ( \u2202d\u0307 / \u2202q\u0307 )T Kb1(\u2202\u03c3\u0307b/\u2202q\u0307)T ] = [ Kd1 nT Kb1 CT RT w ] Jv. (18) A key requirement for surface treatment operations is that the robot tool has to be orthogonal to the workpiece surface, that is, the Z-axis of the robot tool (see Fig. 2) must point in the direction of n. Thus, the reference for the tool orientation is vector n, which can be easily transformed (Siciliano and Khatib, 2008) to roll and pitch reference values, i.e., \u03b1ref and \u03b2ref . It is worth noting that there is no requirement for the yaw angle and, hence, it can be used, for instance, for tool guidance, see Section 3.4.1. Thus, the control equation for Level 2 results in: M2Jq\u0308c = o\u0308ref + Kd2e\u0307o + Kp2eo + sign ( e\u0307o + ( Kp2 / Kd2 ) eo ) u+ 2 \u2192A2q\u0308c = b2, (19) where matrix M2 = [ 0 0 0 1 0 0 0 0 0 0 1 0 ] is used to affect only \u03b1 and \u03b2 angles (i" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_35_0000686_j.mechmachtheory.2021.104433-Figure28-1.png", + "original_path": "designv11-35/openalex_figure/designv11_35_0000686_j.mechmachtheory.2021.104433-Figure28-1.png", + "caption": "Fig. 28. Prototype and experiment.", + "texts": [ + " In this configuration, the driving force provided by the sun bevel gear overcomes the spring force, compresses the limit spring, and drives the lead screw claw to move. Meanwhile, the lead screw claw is constrained by the target object and moves along the lead screw to clamp the target object. To demonstrate the process of configuration transformation better, a metamorphic epicyclic bevel gear clamping mechanism is designed and the corresponding mechanical parts based on the principle model in Fig. 27 are shown in Fig. 28(a). An insulator inspection robot that consists of two metamorphic epicyclic bevel gear clamping mechanisms and a control box is designed as an experimental prototype. The metamorphic epicyclic bevel gear clamping mechanism is used to clamp the insulator to fix the robot. The prototype experiments focused on the configuration transformation are carried out by clamping the insulator as displayed in Fig. 28(b)-(e). When the robot starts to clamp the insulator, the lead screw claw rotates to the insulator by the driving motor and doesn\u2019t move along the lead screw. So metamorphic epicyclic bevel gear clamping mechanism of the robot is in the rotation configuration of the claw as shown in Fig. 28(b). Till the claw contacts the insulator, the limit spring is compressed, and the lead screw claw translates along the lead screw to clamp the insulator as shown in Fig. 28(c). From this, the clamping mechanism transforms from the rotation configuration of the claw to the translation configuration of the claw. On the contrary, when the robot releases the insulator, the spring force makes the lead screw claw translates away from the insulator firstly as shown in Fig. 28(d). As displayed in Fig. 28(e), the lead screw claw starts rotating away from the insulator when the lead screw claw reaches the maximum length limited by combination constraint. From this, the clamping mechanism inversely transforms from the translation configuration of the claw to the rotation configuration of the claw. According to the configuration, motion mode, and constraint condition analysis of the example mentioned above, the equivalent mechanisms are proved to be equivalent to the corresponding metamorphic epicyclic gear trains and effective for designing metamorphic epicyclic gear mechanisms" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_35_0001419_978-94-007-7401-8-Figure5.51-1.png", + "original_path": "designv11-35/openalex_figure/designv11_35_0001419_978-94-007-7401-8-Figure5.51-1.png", + "caption": "Fig. 5.51 2PaPaRRR-1PaPaRR-1RPaPaP-type fully-parallel PM with uncoupled Sch\u00f6nflies motions defined by MF = SF = 4, RF\u00f0 \u00de \u00bc \u00f0v1; v2; v3;xb\u00de, TF = 0, NF = 28, limb topology Pa\\Pa||R||R\\\\R, Pa\\Pa||R||R and R||Pa||Pa||P", + "texts": [ + "4l) 13. 3PaPaPR-1RPPaPa (Fig. 5.47) Pa\\Pa\\kP\\\\R (Fig. 5.4f) Idem No. 3 Pa\\Pa\\kP\\kR (Fig. 5.4h) R||P||Pa||Pa (Fig. 5.4m) 14. 3PaPaPR-1RPPaPa (Fig. 5.48) Pa\\Pa\\\\P\\\\R (Fig. 5.4g) Idem No. 3 Pa\\Pa\\\\P\\kR (Fig. 5.4e) R||P||Pa||Pa (Fig. 5.4m) 15. 3PaPaPR-1RPPaPat (Fig. 5.49) Pa\\PakP\\\\R (Fig. 5.4f) Idem No. 3 Pa\\Pa\\kP\\kR (Fig. 5.4h) R||P||Pa||Pat (Fig. 5.4n) 16. 3PaPaPR-1RPPaPat (Fig. 5.50) Pa\\Pa\\\\P\\\\R (Fig. 5.4g) Idem No. 3 Pa\\Pa\\\\P\\kR (Fig. 5.4e) R||P||Pa||Pat (Fig. 5.4n) 17. 2PaPaRRR-1PaPaRR1RPaPaP (Fig. 5.51) Pa\\Pa||R||R\\\\R (Fig. 5.5a) Idem No. 1 Pa\\Pa||R||R (Fig. 5.4i) R||Pa||Pa||P (Fig. 5.4k) (continued) 532 5 Topologies with Uncoupled Sch\u00f6nflies Motions Table 5.4 (continued) No. PM type Limb topology Connecting conditions 18. 2PaPaRRR-1PaPaRR1RPaPaP (Fig. 5.52) Pa\\Pa||R||R\\kR (Fig. 5.5b) Idem No. 1 Pa\\Pa||R||R (Fig. 5.4i) R||Pa||Pa||P (Fig. 5.4k) 19. 2PaPaRRR-1PaPaRR1RPaPatP (Fig. 5.53) Pa\\Pa||R||R\\\\R (Fig. 5.5a) Idem No. 1 Pa\\Pa||R||R (Fig. 5.4i) R||Pa||Pat||P (Fig. 5.4l) 20. 2PaPaRRR-1PaPaRR1RPaPatP (Fig" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_35_0001563_9781119682684.ch1-Figure1.28-1.png", + "original_path": "designv11-35/openalex_figure/designv11_35_0001563_9781119682684.ch1-Figure1.28-1.png", + "caption": "Figure 1.28 BLDC machine.", + "texts": [ + " For motoring operations, the supplied current is made positive for positive emf and negative for negative emf. Hence, their product is always positive. BLDC are ac synchronous machines with permanent magnets on rotor and trapezoidal back emf shape. The stator poles (projected type) are supplied by current to produce magnetic poles that attract/repel the rotor permanent magnets. The stator poles are energized in a proper sequence to produce a continuous motion. The cross-sectional view of a three-phase BLDC machine is shown in Figure 1.28. The synchronous reluctance machine operates on the principle of reluctance torque. The torque is produced due to saliency in the rotor. The stator has a slotted structure with three-phase distributed winding placed inside the slots. The rotor has no windings and no permanent magnets. The number of poles on stator and rotors are same. The cross-sectional view of a synchronous reluctance machine is shown in Figure 1.29. The switched reluctance machine has salient structure on both stator and rotor" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_35_0001737_8.10872-Figure11-1.png", + "original_path": "designv11-35/openalex_figure/designv11_35_0001737_8.10872-Figure11-1.png", + "caption": "FIG. 11.", + "texts": [ + " It is possible to get a better agreement between the Bo and Bres curves if B0 is calculated for a movable sys tem obtained by the introduction of four imaginary joints. If an equilibrium with the external loads can be established in this movable system, by Castigliano's theorem the actual moment distribution can be found as that combination of the so-obtained B0 moment dia gram with the Bm, Bp, and Bv moment diagrams of Fig. 3 that corresponds to the least strain energy. The movable system is shown in Fig. 11. RPQ is the resultant shear force acting on the portion of the ring extending from point P to point Q. Its magnitude, direction, and line of action have been determined by making use of the reduced moment areas, as explained before. The magnitude and direction of the resultant shear force acting on parts O-P and C-Q have been similarly determined. The three form a closed polygon with the vertical external force 100 lbs. = 9.62 in. 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 6, 2 01 5 | h ttp :// ar c", + "62 in. and of the vertical component of ROP- A similar condition exists for the vertical component of the reaction at Q. Hence, the ratio of the vertical distances of points P and Qy respectively, from the intersection point of the two reactions with RPQ is (RP)v/(RQ)v = 8.16/0.73. This requirement determines the location of the inter section point and, consequently, the directions of the reactions. These being known, the magnitudes of the reactions follow from the graphic construction shown in Fig. 11. The equilibrium of part O-P is established if the necessary horizontal reaction and bending mo ment are assumed to act at point 0. This is permis sible, since by symmetry an equal and opposite moment and horizontal force are required for the equilibrium of the other (not shown) half of the ring. Similar considerations hold true for part C-Q. From this point the procedure is the same as before. Fig. 12 shows the curves B0 and B0Bp/2. The area below these curves was determined with a planimeter and was 2" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_35_0001613_05698195908972366-Figure2-1.png", + "original_path": "designv11-35/openalex_figure/designv11_35_0001613_05698195908972366-Figure2-1.png", + "caption": "FIG . 2. Types of loading.", + "texts": [ + " These two categories consist of (a) offset bearings and (b) centrally loaded bearings. In the centrally loaded partial bearing the line of action of the load is located so as to bisect the bearing arc whereas in the offset bearing (or eccentrically loaded bearing as it is sometimes called) the load line can occur at any angular position other than that yielding central loading. Offset bearings may be further subdivided depending upon the direction of offset relative to the direction of rotation (Fig. 2). This paper is mainly concerned with the properties of offset bearings. In the past, only a few workers have concerned them selves with offset bearings. During the years 1928-29, Howarth (1) and Boswall (2) described respectively the theoretical behavior of offset bearings with arcs of 120\u00b0 and 90\u00b0. These analyses, however, were limited to bearings of infinite length and did not rigorously account for the effects of rupture of the lubricant film. In 1932 Boswall and Brierly (3) conducted some tests which were qualita tively compared with the limited theoretical data then available", + " film pressure, deg P = film pressure, psig Pm.x= max. pr essure developed in lubricant film, psig S = (R/C)2 I-'-N/P = bearing characteristic number, dimensionless t1 = lub ricant inlet temp., deg F t2 = lubricant outlet temp., deg F /:).t = temp. rise = t2-t1, deg F c = specific heat of lubricant, Btu /lb deg F y = weight per unit volume of lubricant, Ib/in3 ] = mechanical equivalent of heat = 9336 in lb/Btu Type of loading The various types of loading that can be experienced with a partial bearing are illustrated in Fig. 2 for a bearing having the arc fl. Figure 2(b) shows the \"centrally loaded\" case where the load line bisects the bearing arc giving IJ. /fl = 0.500; Figs. 2(a, c) show the \"offset\" cases . In Fig. 2(a) the load is offset counter to the direction of rotation and the load angle IJ. is less than one-half the bearing are, i.e. , IX/fl < 0.500 . However, in Fig. 2(c) the load is offset in the direction of rotation and the ratio IX/fl exceeds 0.500. Figure 2, drawn for a fixed value of the eccentricity ratio \u20ac , shows that as the load line shifts, the position of the journal which is characterized by the \"entrance angle\" (JA will vary thus altering the shape of the lubricant film. A change in the film shape will in turn rearrange the distribution of the film pressures (Fig. 1) so that the resultant film pressure will coincide with the line of action of the load in order that equilibrium be restored. As a consequence of the altered film shape, th e performance characteristics, to be discussed subsequently, are affected", + " The lubricant film shape can fall into one of three classifications. These classifications are illustrated respec tively in Figs. 2(a, b, c) as: (a) converging-diverging, (b) wholly converging, (c) diverging-converging. It can be seen from F ig. 2(c) that negative valu es of (JA 2 will cause the film to be diverging at the leading edge while positive values of (JA will result in converging films at the bearing entrance. It should be mentioned that the film shape is not always related to the type of load ing in the manner shown in Fig. 2. The exact relationship may be determined from either Fig. 4 or Fig. 14 which will be discussed subse quently. Analysis The results given in this paper were obtained by solving numerically the fundamental lubrication equations by means of a digital computer. The method of attack , the mesh sizes employed , and other pertinent information have been reported elsewhere (12) and need not be repeated. However, mention should be made of the two main as sumptions used: (a) The viscosity of the lubricant was assumed to be constant as was done before in (10, 11, 12) in order that the present paper be a logical and consistent extension of previous work, thus permitting comparisons and inferences to be made readily. (b) The effects of rupture of the lubricant film in the diverging regions were accounted for by using the boundary conditions that were discussed in detail in (12)3. Results Pressure Distributions With converging-diverging film shapes such as that illustrated by Fig. 2(a) cavitation or film rupture was found to occur in the diverging portion of the film, that i~ near the trailing edge. A typical pressure distribution taken from the results for such an operating condition is drawn in Fig. 3(a) for \u20ac = 0.8, (JA = 100\u00b0. With wholly converg ing film shapes (Fig. 2b) no film rupture was experienced; a typical pressure distribution is shown in Fig. 3(b). On the other hand, with the diverging-converging film shape 2l1 A is taken po sitively in the direction of rotation (see Nomen clature). 3 Namely, that dp /dl1 = P = 0 at point of film rupture. D ow nl oa de d by [ U ni ve rs ity o f C al if or ni a, S an D ie go ] at 1 3: 19 2 4 Ju ne 2 01 6 Effect of Offset L oads on P erformance of 1200 Partial J ournal Bearing 149 FIG. 3(a). Typical p ressure distribution for con verging-diverging film shape (. = 0.8, 8A = 100 0 , a.IP = 0.450). such as that shown by Fig. 2(c), film rupture would tend to occur near the leading edge of the bearing. A typical pressure distribution for such a case is given in Fig . 3(c), drawn for \u20ac = 0.8 but 8A = - 40\u00b0. An interesting result of this study was the discovery that for all the eccentricity ratios investigated (up to and including \u20ac = 0.97), no film rupture was experienced near the entrance of the bearing as long as the entrance angle 8A was greater than - 30\u00b0. 0 .0 E' 'Cc, 3(h). T ypical pressure distribution for wholl y converging film shape (", + " -20 f-- '----,=-- - -----'f---- +-++ -\u00a5----:>-'\"\"\"\"-- - --j -40 Entrance Angle 8A The manner in which the position of the line of centers varies with the bearing characteristic number S and the load position rx/fl is given in Fig. 4. As S increases from zero D ow nl oa de d by [ U ni ve rs ity o f C al if or ni a, S an D ie go ] at 1 3: 19 2 4 Ju ne 2 01 6 D ow nl oa de d by [ U ni ve rs ity o f C al if or ni a, S an D ie go ] at 1 3: 19 2 4 Ju ne 2 01 6 152 A. A. RAIMONDI film thickness prevailing in the bearing arc. If the line of centers, when projected backwards, intersects the bearing arc as shown in Fig. 2(a), then hn must necessarily occur along the line of centers. This condition will always exist as long as the film shape is convergent-divergenrt . However, when the journal assumes a position such as that shown in Fig. 2(c) and the insert in Fig. 5 where the line of centers, when projected backwards, does not intersect the bearing arc, th en the least film thickness prevailing in the bearing arc does not occur along the line of centers but at the trailing edge of the bearing, i.e., hn equals the outlet film thickness. Th is condition will exist for all film shapes other than converging-divergings. 'When the minimum film thickness is defined in this manner, no limit need be placed on its maximum value in contrast to the case of the full journal bearing where the maximum value cannot exceed the radial clearance because of physical limitations", + " 'Tvpical bearing design. with small lands \"I\" extending along the full periphery of the bearing, which design is typical of many turbine bearings and ring-oiled bearings, the data will be applicable as long as the eccentricity ratio \u20ac is equal to or less than unity for it is evident that the radial movement of the journal is limited by the lands to a value equal to the radial clearance of the bearing. Moreover, the least film thickness in the bearing will now necessarily always occur along the line of centers (Fig. 2a) because the lands encircle the journal; its value can be calculated from the equation hn = C(l- \u20ac), where \u20ac is given in Fig. 6. Since the hydrodynamic forces developed over the lands will be negligible the journal can contact the cap . As contact will take place when \u20ac = 1.0, the range of S where contact is likely to occur can be determined from either Fig. 6 or Fig. 7. These figures show that operation with rx jfJ exceeding, say, 0.665, should be avoided with bearings similar to Fig. 20 in order to prevent rubbing on the lands in the top half of the bearing" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_35_0001391_s00170-021-07282-1-Figure5-1.png", + "original_path": "designv11-35/openalex_figure/designv11_35_0001391_s00170-021-07282-1-Figure5-1.png", + "caption": "Fig. 5 Radial force Fr vs deformation displacement \u03b4r of the front/rear bearings measured by EMM and IMM (n = 24,000 rpm). a Rear bearings & EMM, b Front bearings & EMM, c Rear bearings & IMM, d Front bearings & IMM", + "texts": [ + " The original experimental data of the HSMS at a speed of 24,000 rpm is shown in Fig. 4. It shows that under the same loading force, the deformation displacement of the measuring points in the EMM was much larger than that in the IMM, which shows that the bending deformation displacement of the loading rod due to the cantilever beam cannot be ignored. Using the EMM and IMM described in Section 2.2 to analyse the experimental results, the relationship between the loading force and deformation displacement of the front/rear bearings at 24,000 rpm was derived (Fig. 5). The original measured data (blue curves) was corrected by low-pass filtering (green curves); then, the one-to-one correspondence between loading force and deformation displacement (red curve) was obtained by polynomial fitting. According to Fig. 5, the radial deformation displacement \u03b4r of the front/rear bearings increases with increases in the radial force Fr. According to Eqs. (11) and (12), we can verify that the DSSB can be obtained by deducing the radial force Fr to the deformation displacement \u03b4r. That is, the polynomial model (red curve) is derived to obtain the DSSB. Figure 6 shows the experimental results regarding the effect of the radial force on the DSSB, including analyses by EMM and IMM. It can be seen from Fig. 6 that, first, as the radial force Fr increases, the DSSB Kr increases nonlinearly" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_35_0000213_j.triboint.2021.106993-Figure5-1.png", + "original_path": "designv11-35/openalex_figure/designv11_35_0000213_j.triboint.2021.106993-Figure5-1.png", + "caption": "Fig. 5. Prescribed boundary conditions for the CFD solver.", + "texts": [ + " Likewise, the displacement solutions in the structure solver modify the interface boundary conditions to include effects of the solid thermal-elastic deformations, and shaft centrifugal force deformations. The mesh deformation equation is expressed in Eq. (16), where the dependent variable is the mesh displacement (\u03b4i) relative to the initial node location. The mesh stiffness (\u0393dis) is chosen with any constant value to maintain the mesh orthogonal quality and induce uniform mesh deformations in the film thickness direction. Mesh Deformation Equation (\u03b4i): \u2202 \u2202xi \u0393dis \u2202\u03b4i \u2202xi = 0 (16) Fig. 5 shows the prescribed boundary conditions for the CFD solver. The interface boundaries between the journal and the fluid-film and between the pad and the fluid-film are also defined in the CFD solver. For stable convergence, the supply oil inlet is prescribed with total pressures in the continuity and momentum equations, where the total pressure is the sum of the static and dynamic pressures. Ambient pressure is imposed at the side oil outlet. The interface boundaries between the pad and fluid-film (including the step and pocket surfaces) are prescribed with a no-slip wall condition, and a moving wall boundary condition is applied to the interface boundary between the journal and fluid-film, due to the spinning journal" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_35_0000395_978-3-030-67750-3_16-Figure30.9-1.png", + "original_path": "designv11-35/openalex_figure/designv11_35_0000395_978-3-030-67750-3_16-Figure30.9-1.png", + "caption": "Fig. 30.9 Top view of loader bucket", + "texts": [], + "surrounding_texts": [ + "bucket, which is 100 mm linear actuator and 150 mm linear actuator depending on the length of the arm needed to move. All these actuators are controlled by a motor located at the wheelbarrow body tray and near to the handle where the user can easily manage to operate. The connection harness attached under the wheelbarrow bucket is for safety reason. All the items can be referred in Figs. 30.8 and 30.9 that shows the detail view of the loader bucket." + ] + }, + { + "image_filename": "designv11_35_0001419_978-94-007-7401-8-Figure5.39-1.png", + "original_path": "designv11-35/openalex_figure/designv11_35_0001419_978-94-007-7401-8-Figure5.39-1.png", + "caption": "Fig. 5.39 3PaPPaR-1RPaPaP-type fully-parallel PMs with uncoupled Sch\u00f6nflies motions defined by MF = SF = 4, RF\u00f0 \u00de \u00bc \u00f0v1; v2; v3;xb\u00de, TF = 0, NF = 30 limb topology R||Pa||Pa||P and Pa||P\\Pa\\kR, Pa||P\\Pa||R (a), Pa\\P\\\\Pa\\kR, Pa\\P\\\\Pa||R (b)", + "texts": [ + "1b) 12. 2PaPaRRR1PaPaRR1RUPU (Fig. 5.38) Pa\\Pa||R||R\\kR (Fig. 5.5b) Idem No. 11 Pa\\Pa||R||R (Fig. 5.4j) R\\R\\R\\P\\kR\\R (Fig. 5.1b) Table 5.4 Limb topology and connecting conditions of the fully-parallel solutions with no idle mobilities presented in Figs. 5.39, 5.40, 5.41, 5.42, 5.43, 5.44, 5.45, 5.46, 5.47, 5.48, 5.49, 5.50, 5.51, 5.52, 5.53, 5.54, 5.55, 5.56, 5.57, 5.58, 5.59, 5.60, 5.61, 5.62, 5.63, 5.64, 5.65, 5.66, 5.67, 5.68 No. PM type Limb topology Connecting conditions 1. 3PaPPaR-1RPaPaP (Fig. 5.39a) Pa||P\\Pa\\kR (Fig. 5.4a) The last joints of the four limbs have superposed axes/directions Pa||P\\Pa||R (Fig. 5.4d) R||Pa||Pa||P (Fig. 5.4k) 2. 3PaPPaR-1RPaPaP (Fig. 5.39b) Pa\\P\\\\Pa\\kR (Fig. 5.4b) Idem No. 1 Pa\\P\\\\Pa||R (Fig. 5.4c) R||Pa||Pa||P (Fig. 5.4k) 3. 3PaPPaR-1RPPaPa (Fig. 5.40a) Pa||P\\Pa\\kR (Fig. 5.4a) The revolute joint between links 7D and 9 of G4 limb and the last revolute joints of limbs G1, G2 and G3 have superposed axes Pa||P\\Pa||R (Fig. 5.4d) R||P||Pa||Pa (Fig. 5.4m) (continued) 5.1 Fully-Parallel Topologies 531 Table 5.4 (continued) No. PM type Limb topology Connecting conditions 4. 3PaPPaR-1RPPaPa (Fig. 5.40b) Pa\\P\\\\Pa\\kR (Fig. 5.4b) Idem No. 3 Pa\\P\\\\Pa||R (Fig" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_35_0001419_978-94-007-7401-8-Figure2.64-1.png", + "original_path": "designv11-35/openalex_figure/designv11_35_0001419_978-94-007-7401-8-Figure2.64-1.png", + "caption": "Fig. 2.64 4PaPaPR-type fully-parallel PM with coupled Sch\u00f6nflies motions defined by MF \u00bc SF \u00bc 4; RF\u00f0 \u00de \u00bc \u00f0v1; v2; v3;xd\u00de, TF = 0, NF = 30, limb topology Pa\\Pa\\\\P||R", + "texts": [ + "22p) Idem No. 5 44. 4PPaRPa (Fig. 2.56) P||Pa\\R\\Pa (Fig. 2.22q) Idem No. 33 45. 4PPaPaR (Fig. 2.57) P||Pa||Pa\\R (Fig. 2.22r) Idem No. 4 46. 4PaPPaR (Fig. 2.58) Pa||P||Pa\\R (Fig. 2.22s) Idem No. 4 47. 4RPPaPa (Fig. 2.59) R\\P||Pa||Pa (Fig. 2.22t) Idem No. 15 48. 4PaPPaR (Fig. 2.60) Pa\\P\\\\Pa\\kR (Fig. 2.22u) Idem No. 4 49. 4PaRPaP (Fig. 2.61) Pa\\R\\Pa\\kR (Fig. 2.22v) Idem No. 34 50. 4PaRPPa (Fig. 2.62) Pa\\R||P\\Pa (Fig. 2.22w) Idem No. 34 51. 4PaPRPa (Fig. 2.63) Pa\\P||R\\Pa (Fig. 2.22x) Idem No. 33 52. 4PaPaPR (Fig. 2.64) Pa\\Pa\\\\P||R (Fig. 2.22y) Idem No. 4 53. 4PaPaRP (Fig. 2.65) Pa\\Pa\\\\R||P (Fig. 2.22z) Idem No. 41 54. 4RPaPaP (Fig. 2.66) R\\Pa\\\\Pa\\\\P (Fig. 2.22a0) Idem No. 15 55. 4RPaPPa (Fig. 2.67) R\\Pa\\kP\\\\Pa (Fig. 2.22b0) Idem No. 15 56. 4RPPaPa (Fig. 2.68) R||P\\Pa\\\\Pa (Fig. 2.22c0) Idem No. 15 57. 4PRPaPa (Fig. 2.69) P||R\\Pa\\\\Pa (Fig. 2.22d0) Idem No. 5 58. 4PaCPa (Fig. 2.70) Pa\\C\\Pa (Fig. 2.22e0) The axes of the cylindrical joints of the four limbs are parallel 59. 4CPaPa (Fig. 2.71) C\\Pa\\\\Pa (Fig. 2.22f 0) Idem No" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_35_0000512_s42417-021-00289-8-Figure3-1.png", + "original_path": "designv11-35/openalex_figure/designv11_35_0000512_s42417-021-00289-8-Figure3-1.png", + "caption": "Fig. 3 Geometric relationship between ball and raceway", + "texts": [ + " In a certain coordinate system \u2018m\u2019, parameters m d_e and m d_e are the position and velocity vector from part \u2018d\u2019 to part \u2018e\u2019 (subscript \u2018d\u2019 and \u2018e\u2019 can be \u2018b\u2019, \u2018r\u2019, \u2018c\u2019 for each part). If a part is in the inertial coordinate system, m d_e and m d_e can be simply described as rd i and vd i. Parameters \u03c9b b, \u03c9c c, \u03c9r r represent the rotation speed vector of each part in the corresponding body-fixed coordinate system. The relative geometric position between the ball and ring raceway is shown in Fig.\u00a03. In the body-fixed coordinate system of ring, the azimuth angle of ball is \u03b8br and the position vector of the raceway curvature center in the direction of ball is expressed as where rf is the radius of curvature center locus circle of raceway. In the azimuth coordinate system of ball, the position vector of ball center to curvature center of raceway is written as Then, the actual contact angle is written as with subscripts \u20181\u2019, \u20182\u2019, \u20183\u2019 meaning the projection of a b_cv in the direction of X, Y and Z axes", + " Then, the transformation matrix from the azimuth coordinate system of ball to the contact coordinate system is Tap = T(\u03b1c1, \u03b1c2, 0), and the contact deformation between ball and ring raceway can be expressed as: where fr is the radius coefficient of groove curvature; db is the ball diameter. (1) r cv_r = ( 0 \u2212rf sin br rf cos br )T , (2) a b_cv = Ti_aT \u22121 i_r ( r b_r \u2212 r cv_r ) . (3) \u23a7\u23aa\u23aa\u23a8\u23aa\u23aa\u23a9 c1 = arctan \ufffd ra b_r1 \u2215ra b_r3 \ufffd c2 = arctan \ufffd \u2212ra b_r2 \u2215 \ufffd\ufffd ra b_r1 \ufffd2 + \ufffd ra b_r3 \ufffd2 \ufffd , (4) br = ||| a b_cv ||| \u2212 ( fr \u2212 0.5 ) db, 1 3 After the contact deformation is obtained, the elliptical contact area will appear at the contact between ball and raceway which is shown in Fig.\u00a03. Hertz point contact theory is used to calculate the long half axis ap, short half axis bp and contact stress ph. When the ellipse contact area is divided into several narrow strips, the position vector of contact point (on the narrow strip) relative to ball center is obtained as where R is the curvature radius of contact compression surface between ball and raceway; xc is the coordinate value of contact point in x direction. In the inertial coordinate system, the position vector of contact point relative to ring center is described as In the contact coordinate system, the sliding velocity vector of raceway relative to ball is After introducing p r_b into the formula of traction coefficients [19], the vector of traction coefficients is (5) p p_b = ( xc, 0, \u221a R 2 \u2212 x2 c \u2212 \u221a R 2 \u2212 a2 p + \u221a 0" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_35_0001419_978-94-007-7401-8-Figure3.133-1.png", + "original_path": "designv11-35/openalex_figure/designv11_35_0001419_978-94-007-7401-8-Figure3.133-1.png", + "caption": "Fig. 3.133 4PaRPRR-type overactuated PM with coupled Sch\u00f6nflies motions defined by MF \u00bc SF \u00bc 4; RF\u00f0 \u00de \u00bc v1; v2; v3;xd\u00f0 \u00de, TF = 0, NF = 14, limb topology Pa\\R||P||R\\kR", + "texts": [ + "126) R||P||R\\Pa||R (Fig. 3.52s) Idem No. 13 40. 4RPRPaR (Fig. 3.127) R||P||R\\Pa||R (Fig. 3.52t) Idem No. 13 41. 4RRPPaR (Fig. 3.128) R||R||P\\Pa||R (Fig. 3.52u) Idem No. 13 42. 4PPaRRR (Fig. 3.129a) P\\Pa\\kR||R||\\kR (Fig. 3.52v) Idem No. 13 43. 4PaPRRR (Fig. 3.129b) Pa\\P\\\\R||R\\\\R (Fig. 3.52w) Idem No. 13 44. 4PPaRRR (Fig. 3.130) P\\Pa\\\\R||R\\kR (Fig. 3.52x) Idem No. 13 45. 4PaPRRR (Fig. 3.131) Pa\\P||R||R\\kR (Fig. 3.52y) Idem No. 13 46. 4PaRPRR (Fig. 3.132) Pa\\R||P||R\\kR (Fig. 3.52z) Idem No. 13 47. 4PaRPRR (Fig. 3.133) Pa\\R||P||R\\kR (Fig. 3.52a0) Idem No. 13 48. 4PaRRPR (Fig. 3.134) Pa\\R||R||P\\kR (Fig. 3.52b0) Idem No. 13 49. 4RCRPa (Fig. 3.135) R\\C||R\\kPa (Fig. 3.52c0) Idem No. 12 50. 4RRCPa (Fig. 3.136) R\\R||C\\kPa (Fig. 3.52d0) Idem No. 12 51. 4CRRPa (Fig. 3.137) C||R\\R||Pa (Fig. 3.52e0) Idem No. 15 52. 4RCRPa (Fig. 3.138) R||C\\R||Pa (Fig. 3.52f0) Idem No. 15 53. 4CRPaR (Fig. 3.139) C||R\\Pa||R (Fig. 3.52g0) Idem No. 13 54. 4RCPaR (Fig. 3.140) R||C\\Pa||R (Fig. 3.52h0) Idem No. 13 55. 4PaCRR (Fig. 3.141a) Pa\\C||R\\kR (Fig" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_35_0000417_012020-Figure7-1.png", + "original_path": "designv11-35/openalex_figure/designv11_35_0000417_012020-Figure7-1.png", + "caption": "Figure 7. Maximum stress at spoke and deformation of: (a) 3D printed NPT and (b) waterjet cut NPT.", + "texts": [ + " This represented the locking adapter at the rim in the actual experiment. The analysis was then performed and the results were collected. The11th International Conference on Mechanical Engineering (TSME-ICOME 2020) IOP Conf. Series: Materials Science and Engineering 1137 (2021) 012020 IOP Publishing doi:10.1088/1757-899X/1137/1/012020 FEA of vertical stiffness testing of NPT was performed and the results were achieved. The stress and deformation results of FEA of both 3D printed and waterjet material are shown on Figure 7. The maximum stress value at spoke and deformation was found to be high on the 3D printed model due to The11th International Conference on Mechanical Engineering (TSME-ICOME 2020) IOP Conf. Series: Materials Science and Engineering 1137 (2021) 012020 IOP Publishing doi:10.1088/1757-899X/1137/1/012020 its lower modulus of elasticity. The maximum stress value at spoke of both models are summarized in table 4. The maximum of stress that occurred on 3D printed NPT was estimated to be 12.21 % higher than the waterjet cut one" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_35_0001419_978-94-007-7401-8-Figure3.45-1.png", + "original_path": "designv11-35/openalex_figure/designv11_35_0001419_978-94-007-7401-8-Figure3.45-1.png", + "caption": "Fig. 3.45 Overactuated PMs with coupled Sch\u00f6nflies motions of types 4CRRR (a) and 4RCRR (b) defined by MF \u00bc SF \u00bc 4; RF\u00f0 \u00de \u00bc v1; v2; v3;xd\u00f0 \u00de, TF = 0, NF = 2, limb topology C||R\\R||R (a) and R||C\\R||R (b)", + "texts": [ + "3c0) Idem No. 40 44. 4RRRPP (Fig. 3.42a) R||R\\R\\P\\\\P (Fig. 3.3d0) The third joints of the four limbs have parallel axes 45. 4PRRPR (Fig. 3.42b) P||R||R\\P\\\\R (Fig. 3.3e0) Idem No. 17 46. 4RPRPR (Fig. 3.43a) R||P||R\\P\\\\R (Fig. 3.3f0) Idem No. 17 47. 4RRPPR (Fig. 3.43b) R||R||P\\P\\\\R (Fig. 3.3g0) Idem No. 17 48. 4RPRPR (Fig. 3.44a) R||P||R\\P\\\\R (Fig. 3.3h0) Idem No. 17 49. 4RPRPR (Fig. 3.44b) P\\P\\kR||R\\\\R (Fig. 3.3i0) Idem No. 17 50. 4PPRRR (Fig. 3.44c) P\\P||R||R\\\\R (Fig. 3.3j0) Idem No. 17 51. 4CRRR (Fig. 3.45a) C||R\\R||R (Fig. 3.3k0) Idem No. 17 52. 4RCRR (Fig. 3.45b) R||C\\R||R (Fig. 3.3l0) Idem No. 17 53. 4RRCR (Fig. 3.46a) R||R\\C||R (Fig. 3.3m0) Idem No. 16 54. 4RRRC (Fig. 3.46b) R||R\\R||C (Fig. 3.3n0) Idem No. 16 55. 4RCRP (Fig. 3.47a) R\\C||R\\\\P (Fig. 3.3o0) Idem No. 16 56. 4RRCP (Fig. 3.47b) R\\R||C\\\\P (Fig. 3.3p0) Idem No. 16 57. 4CRRP (Fig. 3.48a) C||R\\R\\P (Fig. 3.3q0) Idem No. 44 58. 4RCRP (Fig. 3.48b) R||C\\R\\P (Fig. 3.3r0) Idem No. 44 59. 4CRPR (Fig. 3.49a) C||R\\P\\\\R (Fig. 3.3s0) Idem No. 17 60. 4PCRR (Fig. 3.49b) P\\C||R\\\\R (Fig. 3.3t0) Idem No. 17 3" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_35_0001419_978-94-007-7401-8-Figure6.40-1.png", + "original_path": "designv11-35/openalex_figure/designv11_35_0001419_978-94-007-7401-8-Figure6.40-1.png", + "caption": "Fig. 6.40 3PPaPaR-1RPPaPa-type maximaly regular fully-parallel PM with Sch\u00f6nflies motions defined by MF = SF = 4, (RF) = (v1; v2; v3;xb), TF = 0, NF = 30, limb topology P||Pa||Pa\\R, P||Pa||Pa||R and R||P||Pa||Pa", + "texts": [ + " 2PPaRRR-1PPaRR-1RPPaPa (Fig. 6.36) P||Pa||R||R\\R (Fig. 4.9a) P||Pa||R||R (Fig. 4.8g) R||P||Pa||Pa (Fig. 5.4m) Idem no. 9 16. 2PPaRRR-1PPaRR-1RPPaPat (Fig. 6.37) P||Pa||R||R\\R (Fig. 4.9a) P||Pa||R||R (Fig. 4.8g) R||P||Pa||Pat (Fig. 5.4n) Idem no. 9 17. 3PPaPaR-1RPaPaP (Fig. 6.38) P||Pa||Pa\\R (Fig. 4.10a) P||Pa||Pa||R (Fig. 4.10b) R||Pa||Pa||P (Fig. 5.4k) Idem no. 1 18. 3PPaPaR-1RPaPatP (Fig. 6.39) P||Pa||Pa\\R (Fig. 4.10a) P||Pa||Pa||R (Fig. 4.10b) R||Pa||Pat||P (Fig. 5.4l) Idem no. 1 19. 3PPaPaR-1RPPaPa (Fig. 6.40) P||Pa||Pa\\R (Fig. 4.10a) P||Pa||Pa||R (Fig. 4.10b) R||P||Pa||Pa (Fig. 5.4m) Idem no. 9 20. 3PPaPaR-1RPaPatP (Fig. 6.41) P||Pa||Pa\\R (Fig. 4.10a) P||Pa||Pa||R (Fig. 4.10b) R||P||Pa||Pat (Fig. 5.4n) Idem no. 9 Table 6.14 Limb topology and connecting conditions of the fully-parallel solutions with no idle mobilities presented in Figs. 6.42, 6.43, 6.44, 6.45, 6.46 No. PM type Limb topology Connecting conditions 1. 3PPPaR-1CPaPa (Fig. 6.42a) P\\P\\||Pa\\||R (Fig. 4.8b) P\\P\\||Pa||R, (Fig. 4.8c) C||Pa||Pa (Fig" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_35_0000091_iros45743.2020.9341321-Figure1-1.png", + "original_path": "designv11-35/openalex_figure/designv11_35_0000091_iros45743.2020.9341321-Figure1-1.png", + "caption": "Fig. 1. (Left) Quadrupedal robot, Vision 60, whose full-order and 18- DOF model is considered in this paper for numerical simulations. (Right) Conceptual illustration of the proposed decomposition schemes to synthesize decentralized controllers (Top Right: front-hind decomposition) (Bottom Right: left-right decomposition).", + "texts": [ + " The dynamics of each subsystem are addressed to set up a feedback linearization based Quadratic Programming (QP) problem that synthesizes local controllers. The proposed approach estimates the nonlinear interactions amongst subsystems by assuming that the other subsystem is on the periodic orbit. This reduces the need for information sharing between subsystems. We validate our HZD-based decentralization control scheme through simulations on forward and inplace walking gaits, using the full-order model of Vision 60, manufactured by GhostRobotics1 (see Fig. 1). In our previous work [33], [34], we demonstrated the feasibility of time-invariant decentralized algorithms for bipedal locomotion albeit under the assumption that each subsystem has access to position and orientation data from strategically placed additional inertial measurement units (IMUs). The current paper extends 1https://www.ghostrobotics.io/ the work in [33], [34] to quadrupedal locomotion by 1) proposing decentralized control laws that eliminate the need for communication among subsystems, and 2) extending the decentralized feedback laws to be time-varying", + " We remark that all local subsystems in the aforementioned decomposition schemes include two legs of the robot and we will design local controllers for all individual subsystems. Here, we present the control synthesis for the left-right decentralization scheme. The decentralized controllers for the other decompositions can be obtained similarly and will be discussed in Section V. This section eliminates the need for full-order model knowledge for the purposes of control synthesis by the construction of two subsystems under the left-right decomposition scheme (see Fig. 1). We assume that both subsystems have access to the absolute positions and orientations of the body frame B that are encoded in vectors \u03d1 and \u03d1\u0307. This ensures each subsystem is aware of its own contribution to the cumulative goal of locomotion. However, there is no other variable sharing/communication between the local controllers. Consequently, the configuration variables for the left subsystem can be represented by ql := col(\u03d1, q1, q2, q3, q4, q5, q6), where, (q1, ..., q6) are the six actuated DOFs that parameterize the two legs on the left half of the robot" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_35_0000775_j.engfailanal.2021.105672-Figure2-1.png", + "original_path": "designv11-35/openalex_figure/designv11_35_0000775_j.engfailanal.2021.105672-Figure2-1.png", + "caption": "Fig. 2. Schematic of the helicopter skid landing gear assembly: 1 - forward cross tube, 2 - aft cross tube, 3 - skid tube, 4 - step, 5 - saddle, 6 - shoe, 7 - support.", + "texts": [ + " According to the [14], the helicopter is powered with two turbo engines of 1342 kW and with two-bladed rotor. The crew consists of two members and the capacity is 9 to 14 passengers. The cruising speed is 186 km/h, while the maximum flight speed is 240 km/h. The operating altitude of the flight is 3810 m, the maximum flight height is 6096 m, while the range is 439 km. The weight of the empty helicopter is 3097 kg, while the maximum takeoff weight is 5080 kg [15]. The helicopter is equipped with a landing gear with two parallel skids and two cross tubes, Fig. 2. Two cross tubes (the front and aft cross tube) are used for connecting the skids to the helicopter fuselage on four supports. Comparing it with other types of landing gear, the landing gear with skids has the following advantages: simple and light construction, low costs and reduced maintenance time. For normal landing, landing energy is accumulated in tubular elements of the helicopter landing gear. However, for more difficult landing, this energy can be absorbed by deformation of the landing gear structure" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_35_0001419_978-94-007-7401-8-Figure3.105-1.png", + "original_path": "designv11-35/openalex_figure/designv11_35_0001419_978-94-007-7401-8-Figure3.105-1.png", + "caption": "Fig. 3.105 4PaRPaP-type overactuated PM with coupled Sch\u00f6nflies motions defined by MF \u00bc SF \u00bc 4; RF\u00f0 \u00de \u00bc v1; v2; v3;xd\u00f0 \u00de, TF = 0, NF = 30, limb topology Pa||R\\Pa\\\\P", + "texts": [ + "100) R||Pa\\Pa\\\\P (Fig. 3.51l) The first revolute joints of the four limbs have parallel axes 13. 4PaRPaR (Fig. 3.101) Pa||R\\Pa\\kR (Fig. 3.51m) The last revolute joints of the four limbs have parallel axes 14. 4PaPaRR (Fig. 3.102) Pa\\Pa||R||R (Fig. 3.51n) Idem No. 13 15. 4PaRRPa (Fig. 3.103) Pa\\R||R||Pa (Fig. 3.51o) The revolute joints of the parallelogram loops connecting the four limbs to the moving plateform have parallel axes 16. 4PaPPaR (Fig. 3.104) Pa\\P\\\\Pa||R (Fig. 3.51p) Idem No. 13 17. 4PaRPaP (Fig. 3.105) Pa||R\\Pa\\\\P (Fig. 3.51q) The revolute joints of the parallelogram loops connecting the four limbs to the fixed base have parallel axes 18. 4CPaPa (Fig. 3.106) C||Pa||Pa (Fig. 3.51r) The cylindrical joints of the four limbs have parallel axes 19. 4PaPaC (Fig. 3.107) Pa||Pa||C (Fig. 3.51s) Idem No. 18 20. 4PaCPa (Fig. 3.108) Pa||C||Pa (Fig. 3.51t) Idem No. 18 21. 4RPaRRR (Fig. 3.109) R\\Pa\\kR\\R||R (Fig. 3.52a) Idem No. 13 22. 4RRRRPa (Fig. 3.110a) R||R\\R||R||Pa (Fig. 3.52b) Idem No. 15 23. 4RRRRPa (Fig" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_35_0001419_978-94-007-7401-8-Figure5.47-1.png", + "original_path": "designv11-35/openalex_figure/designv11_35_0001419_978-94-007-7401-8-Figure5.47-1.png", + "caption": "Fig. 5.47 3PaPaPR-1RPPaPa-type fully-parallel PM with uncoupled Sch\u00f6nflies motions defined by MF = SF = 4, RF\u00f0 \u00de \u00bc \u00f0v1; v2; v3;xb\u00de, TF = 0, NF = 30 limb topology Pa\\Pa\\kP\\\\R, Pa\\Pa\\kP\\kR and R||P||Pa||Pa", + "texts": [ + "4c) R||P||Pa||Pat (Fig. 5.4n) 9. 3PaPaPR-1RPaPaP (Fig. 5.43) Pa\\Pa\\kP\\\\R (Fig. 5.4f) Idem No. 1 Pa\\Pa\\kP\\kR (Fig. 5.4h) R||Pa||Pa||P (Fig. 5.4k) 10. 3PaPaPR-1RPaPaP (Fig. 5.44) Pa\\Pa\\\\P\\\\R (Fig. 5.4g) Idem No. 1 Pa\\Pa\\\\P\\kR (Fig. 5.4e) R||Pa||Pa||P (Fig. 5.4k) 11. 3PaPaPR1RPaPatP (Fig. 5.45) Pa\\Pa\\kP\\\\R (Fig. 5.4f) Idem No. 1 Pa\\Pa\\kP\\kR (Fig. 5.4h) R||Pa||Pat||P (Fig. 5.4l) 12. 3PaPaPR1RPaPatP (Fig. 5.46) Pa\\Pa\\\\P\\\\R (Fig. 5.4g) Idem No. 1 Pa\\Pa\\\\P\\kR (Fig. 5.4e) R||Pa||Pat||P (Fig. 5.4l) 13. 3PaPaPR-1RPPaPa (Fig. 5.47) Pa\\Pa\\kP\\\\R (Fig. 5.4f) Idem No. 3 Pa\\Pa\\kP\\kR (Fig. 5.4h) R||P||Pa||Pa (Fig. 5.4m) 14. 3PaPaPR-1RPPaPa (Fig. 5.48) Pa\\Pa\\\\P\\\\R (Fig. 5.4g) Idem No. 3 Pa\\Pa\\\\P\\kR (Fig. 5.4e) R||P||Pa||Pa (Fig. 5.4m) 15. 3PaPaPR-1RPPaPat (Fig. 5.49) Pa\\PakP\\\\R (Fig. 5.4f) Idem No. 3 Pa\\Pa\\kP\\kR (Fig. 5.4h) R||P||Pa||Pat (Fig. 5.4n) 16. 3PaPaPR-1RPPaPat (Fig. 5.50) Pa\\Pa\\\\P\\\\R (Fig. 5.4g) Idem No. 3 Pa\\Pa\\\\P\\kR (Fig. 5.4e) R||P||Pa||Pat (Fig. 5.4n) 17. 2PaPaRRR-1PaPaRR1RPaPaP (Fig. 5.51) Pa\\Pa||R||R\\\\R (Fig. 5.5a) Idem No" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_35_0001419_978-94-007-7401-8-Figure5.92-1.png", + "original_path": "designv11-35/openalex_figure/designv11_35_0001419_978-94-007-7401-8-Figure5.92-1.png", + "caption": "Fig. 5.92 3PaPaPaR-1CPaPat-type fully-parallel PM with uncoupled Sch\u00f6nflies motions defined by MF = SF = 4, RF\u00f0 \u00de \u00bc \u00f0v1; v2; v3;xb\u00de, TF = 0, NF = 39, limb topology Pa\\Pa||Pa\\\\R, Pa\\Pa||Pa||R and C||Pa||Pat", + "texts": [ + "4i) C||Pa||Pat (Fig. 5.4p) 27. 3PaPaPaR1CPaPa (Fig. 5.89) Pa\\Pa||Pa\\kR (Fig. 5.6a) The revolute joint between links 6D and 11 of G4 limb and the last revolute joints of limbs G1, G2 and G3 have superposed axes Pa\\Pa||Pa||R (Fig. 5.6b) C||Pa||Pa (Fig. 5.4o) 28. 3PaPaPaR1CPaPa (Fig. 5.90) Pa\\Pa||Pa\\\\R (Fig. 5.6c) Idem No. 27 Pa\\Pa||Pa||R (Fig. 5.6b) C||Pa||Pa (Fig. 5.54o) 29. 3PaPaPaR1CPaPat (Fig. 5.91) Pa\\Pa||Pa\\kR (Fig. 5.6a) Idem No. 27 Pa\\Pa||Pa||R (Fig. 5.6b) C||Pa||Pat (Fig. 5.4p) 30. 3PaPaPaR1CPaPat (Fig. 5.92) Pa\\Pa||Pa\\\\R (Fig. 5.6c) Idem No. 27 Pa\\Pa||Pa||R (Fig. 5.6b) C||Pa||Pat (Fig. 5.54p) 536 5 Topologies with Uncoupled Sch\u00f6nflies Motions No. Structural parameter Solution Figure 5.7 Figures 5.8, 5.9, 5.10 Figures 5.11 1. m 20 22 22 2. pi (i = 1, 3) 7 8 7 3. p2 7 7 7 4. p4 4 4 6 5. p 25 27 27 6. q 6 6 6 7. k1 1 1 1 8. k2 3 3 3 9. k 4 4 4 10. (RG1) (v1; v2; v3;xb) (v1; v2; v3;xa;xb) (v1; v2; v3;xb) 11. (RG2) (v1; v2; v3;xb) (v1; v2; v3;xb) (v1; v2; v3;xb) 12. (RG3) (v1; v2; v3;xb) (v1; v2; v3;xb;xd) (v1; v2; v3;xb) 13" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_35_0000373_17415977.2021.1910683-Figure12-1.png", + "original_path": "designv11-35/openalex_figure/designv11_35_0000373_17415977.2021.1910683-Figure12-1.png", + "caption": "Figure 12. Schematic representation of the SHM for a three-axle vehicle.", + "texts": [ + " It is important to highlight some assumptions made for the SHM application: \u2022 Constant vehicle velocity; \u2022 No lateral or longitudinal load transfer, consequently the vehicle roll and pitch move- ments are not considered; \u2022 Planar model; \u2022 Use of a linear tyre model, whose reliability range occurs for small slip angles, and at low lateral acceleration and velocity; \u2022 Cornering stiffness considered constant. In order to analyse the lateral dynamics of a 6x6 vehicle, the SHM of a 3-axle, 1-rigid body ground vehicle is proposed in Figure 12, considering that the front and themiddle axles are steerable with angles \u03b4f and \u03b4m, respectively, and disregarding the roll motion and other parameters such as ride motion (e.g. pitch and bounce), which affect the real behaviour of a vehicle. The steer angle of the front axle \u03b4f is associated with the driver\u2019s steering wheel input. For low speed turning, the geometry of a manoeuvre during curves may be designed Figure 13. Equivalent SHM for a front andmiddle wheel-steering 6x6 vehicle. Ackermann condition for a three axle model", + " (26) Assuming small and equal tyre slip angles on the left and the right wheels [23] for each axle, the lateral force in the \u2018equivalent\u2019 wheel is obtained byEquation (13), applying the resultant cornering stiffness parameters for an SHM (as expressed by Equations (15) and (16)): Fyf = \u2212C\u03b1f (\u03b2f \u2212 \u03b4f ), (27) Fym = \u2212C\u03b1m(\u03b2m \u2212 \u03b4m), (28) Fyr = \u2212C\u03b1r\u03b2r, (29) where: \u03b2f = tan\u22121 ( vy + a1r vx ) , (30) \u03b2m = tan\u22121 ( vy + a2r vx ) , (31) \u03b2r = tan\u22121 ( vy \u2212 a3r vx ) . (32) C\u03b1f , C\u03b1m and C\u03b1r are the cornering stiffness of the front, middle and rear \u2018equivalent\u2019 tyres (or axles), respectively, for the SHM proposed in Figure 12. Since the measurement of the cornering stiffness is a complex task, the values of these parameters for a 6x6 vehicle were estimated so that it would be possible to compare a simulatedmodel with a real experiment test on the vehicle during a double-lane change manoeuvre. Given the formulation expressed by Equation 24 for the forward problem of a 6x6 vehicle, Simulink R\u00a9 was used to model and simulate the vehicle behaviour using a blockdiagram language given in Figure 14, while the parameters values were initialized using MATLAB R\u00a9" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_35_0001419_978-94-007-7401-8-Figure5.24-1.png", + "original_path": "designv11-35/openalex_figure/designv11_35_0001419_978-94-007-7401-8-Figure5.24-1.png", + "caption": "Fig. 5.24 2PaRPRR-1PaRPR-1RPPaPa-type fully-parallel PMs with uncoupled Sch\u00f6nflies motions defined by MF = SF = 4, RF\u00f0 \u00de \u00bc \u00f0v1; v2; v3;xb\u00de), TF = 0, NF = 19, limb topology Pa\\R\\P\\kR\\R, Pa\\R\\P\\kR and R||P||Pa||Pa", + "texts": [ + "22a) Pa\\P\\\\R||R\\\\R (Fig. 5.3a) Idem No. 9 Pa\\P\\\\R||R (Fig. 5.2e) R||P||Pa||Pat (Fig. 5.4n) 16. 2PaPRRR-1PaPRR1RPPaPat (Fig. 5.22b) Pa||P\\R||R\\\\R (Fig. 5.3b) Idem No. 9 Pa||P\\R||R (Fig. 5.2f) R||P||Pa||Pat (Fig. 5.4n) 17. 2PaRPRR-1PaRPR1RPaPaP (Fig. 5.23) Pa\\R\\P\\kR\\R (Fig. 5.3c) The last joints of the four limbs have superposed axes/directions. The revolute joints between links 4 and 5 of limbs G1, G2 and G3 have orthogonal axes Pa\\R\\P\\kR (Fig. 5.2g) R||Pa||Pa||P (Fig. 5.4k) 18. 2PaRPRR-1PaRPR1RPPaPa (Fig. 5.24) Pa\\R\\P\\kR\\R (Fig. 5.3c) Idem No. 17 Pa\\R\\P\\kR (Fig. 5.2g) R||P||Pa||Pa (Fig. 5.4m) 19. 2PaRPRR-1PaRPR1RPaPatP (Fig. 5.25) Pa\\R\\P\\kR\\R (Fig. 5.3c) Idem No. 17 Pa\\R\\P\\kR (Fig. 5.2g) R||Pa||Pat||P (Fig. 5.4l) 20. 2PaRPRR-1PaRPR1RPPaPat (Fig. 5.26) Pa\\R\\P\\kR\\R (Fig. 5.3c) The revolute joint between links 7D and 9 of G4 limb and the last revolute joints of limbs G1, G2 and G3 have superposed axes. The revolute joints between links 4 and 5 of limbs G1, G2 and G3 have orthogonal axes Pa\\R\\P\\kR (Fig. 5" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_35_0001419_978-94-007-7401-8-Figure2.106-1.png", + "original_path": "designv11-35/openalex_figure/designv11_35_0001419_978-94-007-7401-8-Figure2.106-1.png", + "caption": "Fig. 2.106 Fully-parallel PMs with coupled Sch\u00f6nflies motions of types 4RRRPaR (a) and 4RRPaRR (b) defined by MF \u00bc SF \u00bc 4; RF\u00f0 \u00de \u00bc \u00f0v1; v2; v3;xd\u00de, TF = 0, NF = 14, limb topology R\\R||R\\Pa\\kR (a) and R\\R\\Pa\\kR||R (b)", + "texts": [], + "surrounding_texts": [ + "160 2 Fully-Parallel Topologies with Coupled Sch\u00f6nflies Motions", + "2.2 Topologies with Complex Limbs 161", + "162 2 Fully-Parallel Topologies with Coupled Sch\u00f6nflies Motions" + ] + }, + { + "image_filename": "designv11_35_0001419_978-94-007-7401-8-Figure5.84-1.png", + "original_path": "designv11-35/openalex_figure/designv11_35_0001419_978-94-007-7401-8-Figure5.84-1.png", + "caption": "Fig. 5.84 3PaPaPR-1CPaPat-type fully-parallel PM with uncoupled Sch\u00f6nflies motions defined by MF = SF = 4, RF\u00f0 \u00de \u00bc \u00f0v1; v2; v3;xb\u00de, TF = 0, NF = 30 limb topology Pa\\Pa\\\\P\\\\R, Pa\\Pa\\\\P\\kR and C||Pa||Pat", + "texts": [ + "4c) C||Pa||Pat (Fig. 5.4p) (continued) 5.1 Fully-Parallel Topologies 535 Table 5.5 (continued) No. PM type Limb topology Connecting conditions 19. 3PaPaPR1CPaPa (Fig. 5.81) Pa\\Pa\\kP\\\\R (Fig. 5.4f) Idem No. 15 Pa\\Pa\\kP\\kR (Fig. 5.4h) C||Pa||Pa (Fig. 5.4o) 20. 3PaPaPR1CPaPa (Fig. 5.82) Pa\\Pa\\\\P\\\\R (Fig. 5.4g) Idem No. 15 Pa\\Pa\\\\P\\kR (Fig. 5.4e) C||Pa||Pa (Fig. 5.40) 21. 3PaPaPR1CPaPat (Fig. 5.83) Pa\\Pa\\kP\\\\R (Fig. 5.4f) Idem No. 15 Pa\\Pa\\kP\\kR (Fig. 5.4h) C||Pa||Pat (Fig. 5.4p) 22. 3PaPaPR1CPaPat (Fig. 5.84) Pa\\Pa\\\\P\\\\R (Fig. 5.4g) Idem No. 15 Pa\\Pa\\\\P\\kR (Fig. 5.4e) C||Pa||Pat (Fig. 5.4p) 23. 2PaPaRRR1PaPaRR1CPaPa (Fig. 5.85) Pa\\Pa||R||R\\\\R (Fig. 5.5a) The revolute joint between links 6D and 10 of G4 limb and the last revolute joints of limbs G1, G2 and G3 have superposed axes Pa\\Pa||R||R (Fig. 5.4i) C||Pa||Pa (Fig. 5.40) 24. 2PaPaRRR1PaPaRR1CPaPa (Fig. 5.86) Pa\\Pa||R||R\\kR (Fig. 5.5b) Idem No. 23 Pa\\Pa||R||R (Fig. 5.4i) C||Pa||Pa (Fig. 5.40) 25. 2PaPaRRR1PaPaRR1CPaPat (Fig. 5.87) Pa\\Pa||R||R\\\\R (Fig" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_35_0000174_j.matpr.2021.01.640-Figure3-1.png", + "original_path": "designv11-35/openalex_figure/designv11_35_0000174_j.matpr.2021.01.640-Figure3-1.png", + "caption": "Fig. 3. Extrude Model Of The Drive Shaft.", + "texts": [], + "surrounding_texts": [ + "Drive Shaft With Mesh perspective is appeared in Fig. 6. The drive shaft is meshed number of elements. Mesh influences the accuracy, convergence, speed of the solution, Here used triangular type of mesh is always quick and easy to create. Drive shaft model is divided into finite no of convenient sub elements and each element corners are joined with adjacent element for the purpose of finding bonding strength between sub elements." + ] + }, + { + "image_filename": "designv11_35_0001419_978-94-007-7401-8-Figure3.49-1.png", + "original_path": "designv11-35/openalex_figure/designv11_35_0001419_978-94-007-7401-8-Figure3.49-1.png", + "caption": "Fig. 3.49 Overactuated PMs with coupled Sch\u00f6nflies motions of types 4CRPR (a) and 4PCRR (b) defined by MF \u00bc SF \u00bc 4; RF\u00f0 \u00de \u00bc v1; v2; v3;xd\u00f0 \u00de, TF = 0, NF = 2, limb topology C||R\\P\\\\R (a) and P\\C||R\\\\R (b)", + "texts": [ + "44c) P\\P||R||R\\\\R (Fig. 3.3j0) Idem No. 17 51. 4CRRR (Fig. 3.45a) C||R\\R||R (Fig. 3.3k0) Idem No. 17 52. 4RCRR (Fig. 3.45b) R||C\\R||R (Fig. 3.3l0) Idem No. 17 53. 4RRCR (Fig. 3.46a) R||R\\C||R (Fig. 3.3m0) Idem No. 16 54. 4RRRC (Fig. 3.46b) R||R\\R||C (Fig. 3.3n0) Idem No. 16 55. 4RCRP (Fig. 3.47a) R\\C||R\\\\P (Fig. 3.3o0) Idem No. 16 56. 4RRCP (Fig. 3.47b) R\\R||C\\\\P (Fig. 3.3p0) Idem No. 16 57. 4CRRP (Fig. 3.48a) C||R\\R\\P (Fig. 3.3q0) Idem No. 44 58. 4RCRP (Fig. 3.48b) R||C\\R\\P (Fig. 3.3r0) Idem No. 44 59. 4CRPR (Fig. 3.49a) C||R\\P\\\\R (Fig. 3.3s0) Idem No. 17 60. 4PCRR (Fig. 3.49b) P\\C||R\\\\R (Fig. 3.3t0) Idem No. 17 3.1 Topologies with Simple Limbs 245 Table 3.3 Structural parametersa of parallel mechanisms in Figs. 3.4, 3.5, 3.6, 3.7, 3.8, 3.9, 3.10, 3.11, 3.12, 3.13, 3.14, 3.15, 3.16, 3.17, 3.18, 3.19, 3.20, 3.21, 3.22, 3.23, 3.24, 3.25, 3.26 No. Structural parameter Solution Figures 3.4, 3.5, 3.6, 3.7, 3.8, 3.9, 3.10, 3.11, 3.12, 3.13, 3.14, 3.15, 3.16, 3.17, 3.18, 3.19, 3.20 Figures 3.21, 3.22, 3.23, 3.24, 3.25, 3.26 1. m 14 10 2. pi (i = 1,\u2026,4) 4 3 3. p 16 12 4. q 3 3 5" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_35_0001419_978-94-007-7401-8-Figure2.48-1.png", + "original_path": "designv11-35/openalex_figure/designv11_35_0001419_978-94-007-7401-8-Figure2.48-1.png", + "caption": "Fig. 2.48 4PaPaPR-type fully-parallel PM with coupled Sch\u00f6nflies motions defined by MF \u00bc SF \u00bc 4; RF\u00f0 \u00de \u00bc \u00f0v1; v2; v3;xd\u00de, TF = 0, NF = 30, limb topology Pa\\Pa\\\\P\\\\R", + "texts": [ + "42) P\\R||Pa\\Pa (Fig. 2.22c) Idem No. 5 31. 4RPPaPa (Fig. 2.43) R\\P\\kPa\\\\Pa (Fig. 2.22d) Idem No. 15 32. 4PaPPaR (Fig. 2.44) Pa\\P\\\\Pa\\\\R (Fig. 2.22e) Idem No. 4 33. 4PaPRPa (Fig. 2.45) Pa\\P\\kR\\Pa (Fig. 2.22f) The axes of the revolute joints connecting links 5 and 6 of the four limbs are parallel 34. 4PaRPPa (Fig. 2.46) Pa||R\\P\\\\Pa (Fig. 2.22g) The axes of the revolute joints connecting links 4 and 5 of the four limbs are parallel 35. 4RPaPPa (Fig. 2.47) R||Pa\\P\\\\Pa (Fig. 2.22h) Idem No. 15 36. 4PaPaPR (Fig. 2.48) Pa\\Pa\\\\P\\\\R (Fig. 2.22i) Idem No. 4 37. 4RPaPaR (Fig. 2.49) R\\Pa||Pa\\kR (Fig. 2.22j) The first and the last revolute joints of the four limbs are parallel 38. 4RPaPaR (Fig. 2.50) R\\Pa||Pa\\kR (Fig. 2.22k) Idem No. 37 39. 4RPaRPa (Fig. 2.51) R\\Pa\\kR\\Pa (Fig. 2.22l) Idem No. 15 40. 4PaRPPa (Fig. 2.52) Pa\\R\\P||Pa (Fig. 2.22m) Idem No. 34 41. 4PaPaRP (Fig. 2.53) Pa||Pa\\R\\P (Fig. 2.22n) The axes of the revolute joints connecting links 7 and 8 of the four limbs are parallel 42. 4PaPaPR (Fig. 2.54) Pa||Pa||P\\R (Fig" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_35_0001393_ijhm.2020.109916-Figure4-1.png", + "original_path": "designv11-35/openalex_figure/designv11_35_0001393_ijhm.2020.109916-Figure4-1.png", + "caption": "Figure 4 Boundary condition of spur gear (see online version for colours)", + "texts": [ + " Accuracy and efficiency of finite element method depends upon the meshing size of the model. A very fine mesh is used to design the model so that the results would be more accurate. In the spur gear, the shear stress (\u03c4), equivalent stress (\u03c3v) and maximum shear stress (\u03c3m) directly depends on the amount of load and type of load acting on the spur gear. Performance and life span depend upon stress and deformation. To get the analysis results, the SOLIDWORKS design model is imported to ANSYS Workbench and boundary conditions are applied (Figure 4). For the design of the spur gear, 31 number of teeth with a square keyway type are used. In this validation study, material type is structural steel (Manickaraj, 201)). From Table 3, it is concluded that the generated value for a single square keyway with pressure angle 14.5\u00b0 through simulation is approximately the same as the reference value. Reference value and current analysis error value for equivalent stress is 10% and for maximum shear stress value error is 2% and for shear stress is 7%. From Table 4, it is concluded that the generated value for square keyway with pressure angle 20\u00b0 through simulation is approximately same to the reference value" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_35_0000025_j.talanta.2021.122185-Figure1-1.png", + "original_path": "designv11-35/openalex_figure/designv11_35_0000025_j.talanta.2021.122185-Figure1-1.png", + "caption": "Fig. 1. Schematic experimental setup for fully automated measurements.", + "texts": [ + " The pH meter was calibrated using NIST traceable buffer solutions from SI Analytics. The Arduino\u00ae Nano prototyping platform was used for controlling the stepper motor driver. For running the peristaltic pump, a stepper motor driver DRV8825 from Pololu was used and connected according to the manufacturer\u2019s recommendations [10]. A precision peristaltic pump of type KCS Plus \u2013 SM8B04 A from Kamoer\u00ae Fluid Tech was used. A Samsung Galaxy Grand Prime SM-G530FZ mobile phone was used to record the pH meter display (see Fig. 1). Software. The WinASPECT\u00ae software version 2.3.9.0 was used to run the spectrophotometer. This automated system was created with using the Arduino\u00ae Open-Source prototyping platform [11]. The peristaltic pump and the pH value readings were controlled by a user friendly Windows Form Application created in Microsoft\u00ae Visual Studio 2017\u00ae software [12] in the C# programming language. The DroidCam application version 6.8.3 from Dev47Apps was used for sharing the video recording in real time to the computer [13]; however, any other kind of USB or web camera can be used" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_35_0000213_j.triboint.2021.106993-Figure6-1.png", + "original_path": "designv11-35/openalex_figure/designv11_35_0000213_j.triboint.2021.106993-Figure6-1.png", + "caption": "Fig. 6. Illustration of prescribed boundary conditions for the FEA solver; (a) shaft domain, (b) pad domain.", + "texts": [ + " Palazzolo Tribology International 159 (2021) 106993 set to zero at all inlet and outlet boundary conditions. A 40 degC oil temperature is applied at the supply oil inlet and the side outlet for the energy equations. Heat convection boundary conditions, with heat convection coefficients and surrounding temperatures, are applied on the shaft, pad, and groove outer surfaces. All interface boundaries between the journal and fluid-film and between the pad and fluid film take the displacement values for all shaft and pad dofs and the thermal-elastic deformations. Fig. 6 illustrates the prescribed boundary conditions in the FEA solver. Symmetry boundary conditions are imposed by applying zero displacements in the z direction. As shown in Fig. 6(b), the additional y\u2019-constraints in the middle of the inner pad surface are taken to make a solvable problem. The pad pivot boundaries are fixed in the x\u2019, y\u2019, and z\u2019 directions. Displacements from the nonlinear pivot stiffness are calculated separately in a Python-based code [16]. The original TPJB is referred to as \u201cSmooth TPJB,\u201d and the proposed TPJB with the pockets and steps is referred to as \u201cPocketed TPJB\u201d. The benefits of the latter are demonstrated by comparing the predicted performances of the Smooth TPJB and Pocketed TPJB" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_35_0001385_978-981-15-6095-8-Figure5.6-1.png", + "original_path": "designv11-35/openalex_figure/designv11_35_0001385_978-981-15-6095-8-Figure5.6-1.png", + "caption": "Fig. 5.6 Strain variation at mid-height of geocell mattress (d/B = 1.2)\u2014Test series B", + "texts": [ + " Even at a settlement equal to about 50% of the footing width, clear signs of failure were not evident in many cases with geocell reinforcement. At this stage, the load on the footing was as high as 8 times the ultimate capacity of the unreinforced sand. In many of these tests, the failure was not apparent in the geocell-reinforced sand beds even at settlements as high as 5 times the failure settlement of the unreinforced footings. Hence, a five-fold increase in permissible settlement may be allowed for the geocell-reinforced sand beds without much serious consequence in the superstructure. Figure 5.6 shows strain variations alongwidth of geocellmattress at itsmid-height, for a typical case with d/B = 1.2, h/B = 0.8, b/B = 12, u/B = 0.1. Similar pattern of strain variation was observed in other cases as well. It can be seen that the strain in the geocell mattress is maximum in the central region underneath the footing. This is because in the region underneath the footing, the geocell reinforcement actively restrains the stress concentration induced yield in the soil mass and thereby its strength is significantly mobilized leading to enhanced load-carrying capacity" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_35_0000004_j.ijpvp.2021.104314-Figure2-1.png", + "original_path": "designv11-35/openalex_figure/designv11_35_0000004_j.ijpvp.2021.104314-Figure2-1.png", + "caption": "Fig. 2. The diagram of the bi-directional pig applied in the experiment.", + "texts": [ + " An infrared sensor was installed near the outlet of the pipeline for ensuring the security of the measurement. Once the pig arrives at the detecting area, the frequency of the main winch drops sharply and the pig will stop gradually due to the friction force. In addition, to ensure the stability of the measurement, the centering device is set between the main winch and the receiver, which restraint the vertical displacement of the drag rope by using two fixed pulleys and maintain the drag rope always along the axis of the pipeline. The bi-directional pig adopted in the experiment is shown in Fig. 2. The sealing discs are made of polyurethane rubber which has typical viscoelastic characteristics; the spacers and the mandrel are made of structural steel. The structure sizes and the property parameters are illustrated in Table 1. By changing the frequency of the main winch, the bi-directional pig was driven with the velocity of 0.02 m/s, 0.5 m/s, and 1 m/s respectively in this experiment. The procedure of the measurement is as follows. First, the front end and the rear end of the pig were respectively connected with the main winch and the auxiliary winch through the drag rope" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_35_0001419_978-94-007-7401-8-Figure5.30-1.png", + "original_path": "designv11-35/openalex_figure/designv11_35_0001419_978-94-007-7401-8-Figure5.30-1.png", + "caption": "Fig. 5.30 2PaRRRR-1PaRRR-1RPPaPat-type fully-parallel PMs with uncoupled Sch\u00f6nflies motions defined by MF = SF = 4, RF\u00f0 \u00de \u00bc \u00f0v1; v2; v3;xb\u00de, TF = 0, NF = 19, limb topology Pa\\R||R||R\\kR, Pa\\R||R||R and R||P||Pa||Pat", + "texts": [ + "1 Fully-Parallel Topologies 529 Table 5.2 (continued) No. PM type Limb topology Connecting conditions 21. 2PaRRRR-1PaRRR1RPaPaP (Fig. 5.27) Pa\\R||R||R\\kR (Fig. 5.3d) Idem No. 17 Pa\\R||R||R (Fig. 5.2h) R||Pa||Pa||P (Fig. 5.4k) 22. 2PaRRRR-1PaRRR1RPPaPa (Fig. 5.28) Pa\\R||R||R\\kR (Fig. 5.3d) Idem No. 20 Pa\\R||R||R (Fig. 5.2h) R||P||Pa||Pa (Fig. 5.4m) 23. 2PaRRRR-1PaRRR1RPaPatP (Fig. 5.29) Pa\\R||R||R\\kR (Fig. 5.3d) Idem No. 17 Pa\\R||R||R (Fig. 5.2h) R||Pa||Pat||P (Fig. 5.4l) 24. 2PaRRRR-1PaRRR1RPPaPat (Fig. 5.30) Pa\\R||R||R\\kR (Fig. 5.3d) Idem No. 20 Pa\\R||R||R (Fig. 5.2h) R||P||Pa||Pat (Fig. 5.4n) Table 5.3 Limb topology and connecting conditions of the fully-parallel solutions with no idle mobilities presented in Figs. 5.31, 5.32, 5.33, 5.34, 5.35, 5.36, 5.37, 5.38 No. PM type Limb topology Connecting conditions 1. 3PaPPaR1RPPP (Fig. 5.31a) Pa||P\\Pa\\kR (Fig. 5.4a) The last joints of the four limbs have superposed axes/directionsPa||P\\Pa||R (Fig. 5.4d) R\\P\\\\P\\\\P (Fig. 5.1a) 2. 3PaPPaR- 1RPPP (Fig. 5.31b) Pa\\P\\\\Pa\\kR (Fig" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_35_0001419_978-94-007-7401-8-Figure5.23-1.png", + "original_path": "designv11-35/openalex_figure/designv11_35_0001419_978-94-007-7401-8-Figure5.23-1.png", + "caption": "Fig. 5.23 2PaRPRR-1PaRPR-1RPaPaP-type fully-parallel PMs with uncoupled Sch\u00f6nflies motions defined by MF = SF = 4, RF\u00f0 \u00de \u00bc \u00f0v1; v2; v3;xb\u00de, TF = 0, NF = 19, limb topology Pa\\R\\P\\kR\\R, Pa\\R\\P\\kR and R||Pa||Pa||P", + "texts": [ + " 5.21a) Pa\\P\\\\R||R\\\\R (Fig. 5.3a) Idem No. 9 Pa\\P\\\\R||R (Fig. 5.2e) R||Pa||Pat||P (Fig. 5.4l) 14. 2PaPRRR-1PaPRR1RPaPatP (Fig. 5.21b) Pa||P\\R||R\\\\R (Fig. 5.3b) Idem No. 9 Pa||P\\R||R (Fig. 5.2f) R||Pa||Pat||P (Fig. 5.4l) 15. 2PaPRRR-1PaPRR1RPPaPat (Fig. 5.22a) Pa\\P\\\\R||R\\\\R (Fig. 5.3a) Idem No. 9 Pa\\P\\\\R||R (Fig. 5.2e) R||P||Pa||Pat (Fig. 5.4n) 16. 2PaPRRR-1PaPRR1RPPaPat (Fig. 5.22b) Pa||P\\R||R\\\\R (Fig. 5.3b) Idem No. 9 Pa||P\\R||R (Fig. 5.2f) R||P||Pa||Pat (Fig. 5.4n) 17. 2PaRPRR-1PaRPR1RPaPaP (Fig. 5.23) Pa\\R\\P\\kR\\R (Fig. 5.3c) The last joints of the four limbs have superposed axes/directions. The revolute joints between links 4 and 5 of limbs G1, G2 and G3 have orthogonal axes Pa\\R\\P\\kR (Fig. 5.2g) R||Pa||Pa||P (Fig. 5.4k) 18. 2PaRPRR-1PaRPR1RPPaPa (Fig. 5.24) Pa\\R\\P\\kR\\R (Fig. 5.3c) Idem No. 17 Pa\\R\\P\\kR (Fig. 5.2g) R||P||Pa||Pa (Fig. 5.4m) 19. 2PaRPRR-1PaRPR1RPaPatP (Fig. 5.25) Pa\\R\\P\\kR\\R (Fig. 5.3c) Idem No. 17 Pa\\R\\P\\kR (Fig. 5.2g) R||Pa||Pat||P (Fig. 5.4l) 20. 2PaRPRR-1PaRPR1RPPaPat (Fig" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_35_0000387_tmag.2021.3079978-Figure1-1.png", + "original_path": "designv11-35/openalex_figure/designv11_35_0000387_tmag.2021.3079978-Figure1-1.png", + "caption": "Fig. 1 Structure of three-phase 12/8-pole DSBLM.", + "texts": [ + " See http://www.ieee.org/publications_standards/publications/rights/index.html for more information. > BW-17 < 2 iron and copper loss of the motor, the converter loss is also an important part of the motor system. In this paper, the loss of a DSBLM with rectangular wire armature winding driven by square and sinusoidal wave currents are comparatively studied. The effect of PWM on motor loss for actual operation is considered, and the converter loss under different driving modes is also studied. Fig. 1 shows the structure of the 12/8-pole DSBLM. The stator and rotor are all salient pole structure. The stator is wound with armature winding and DC field winding, while there is no permanent magnet and winding on the rotor. The stator is divided into large slot and small slot according to slot shape. There are armature winding and field winding in the large slot, while there is only armature winding in the small slot. The armature winding adopts rectangular wire and the DC field winding adopts round wire" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_35_0000146_012064-Figure6-1.png", + "original_path": "designv11-35/openalex_figure/designv11_35_0000146_012064-Figure6-1.png", + "caption": "Figure 6. Spring B deformed configuration at maximum load. Result of proposed model in blue, FE in red.", + "texts": [ + " Figure 4 shows undeformed Spring A (wireframe) and deformed configurations from both proposed model (blue) and FE (red) results. The two shapes are practically superimposed and is hard to distinguish between them. Load deflections curves at loaded node are plotted in figure 5. It is possible to notice the effect of contacts, especially on the vertical displacement of the tip. Furthermore, it is noticed that the proposed model is always showing a softer behavior than FE, in both with and without contact solutions. Deformation of Spring B at maximum load is shown in figure 6 and the load-displacement characteristic of the spring in figure 7. Again, it is hard to distinguish between the rendering based on the proposed model and that based on finite elements. Coils in contact Coil compenetration Figure 4. Spring A deformed at maximum load, without coil contact (left) and with coil contact. Proposed model result is blue and FE is red. The 49th AIAS Conference (AIAS 2020) IOP Conf. Series: Materials Science and Engineering 1038 (2021) 012064 IOP Publishing doi:10.1088/1757-899X/1038/1/012064 Figure 7" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_35_0001419_978-94-007-7401-8-Figure3.134-1.png", + "original_path": "designv11-35/openalex_figure/designv11_35_0001419_978-94-007-7401-8-Figure3.134-1.png", + "caption": "Fig. 3.134 4PaRRPR-type overactuated PM with coupled Sch\u00f6nflies motions defined by MF \u00bc SF \u00bc 4; RF\u00f0 \u00de \u00bc v1; v2; v3;xd\u00f0 \u00de, TF = 0, NF = 14, limb topology Pa\\R||R||P\\kR", + "texts": [ + "127) R||P||R\\Pa||R (Fig. 3.52t) Idem No. 13 41. 4RRPPaR (Fig. 3.128) R||R||P\\Pa||R (Fig. 3.52u) Idem No. 13 42. 4PPaRRR (Fig. 3.129a) P\\Pa\\kR||R||\\kR (Fig. 3.52v) Idem No. 13 43. 4PaPRRR (Fig. 3.129b) Pa\\P\\\\R||R\\\\R (Fig. 3.52w) Idem No. 13 44. 4PPaRRR (Fig. 3.130) P\\Pa\\\\R||R\\kR (Fig. 3.52x) Idem No. 13 45. 4PaPRRR (Fig. 3.131) Pa\\P||R||R\\kR (Fig. 3.52y) Idem No. 13 46. 4PaRPRR (Fig. 3.132) Pa\\R||P||R\\kR (Fig. 3.52z) Idem No. 13 47. 4PaRPRR (Fig. 3.133) Pa\\R||P||R\\kR (Fig. 3.52a0) Idem No. 13 48. 4PaRRPR (Fig. 3.134) Pa\\R||R||P\\kR (Fig. 3.52b0) Idem No. 13 49. 4RCRPa (Fig. 3.135) R\\C||R\\kPa (Fig. 3.52c0) Idem No. 12 50. 4RRCPa (Fig. 3.136) R\\R||C\\kPa (Fig. 3.52d0) Idem No. 12 51. 4CRRPa (Fig. 3.137) C||R\\R||Pa (Fig. 3.52e0) Idem No. 15 52. 4RCRPa (Fig. 3.138) R||C\\R||Pa (Fig. 3.52f0) Idem No. 15 53. 4CRPaR (Fig. 3.139) C||R\\Pa||R (Fig. 3.52g0) Idem No. 13 54. 4RCPaR (Fig. 3.140) R||C\\Pa||R (Fig. 3.52h0) Idem No. 13 55. 4PaCRR (Fig. 3.141a) Pa\\C||R\\kR (Fig. 3.52i0) Idem No. 13 56. 4PaRCR (Fig. 3.141b) Pa\\R||C\\kR (Fig" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_35_0001419_978-94-007-7401-8-Figure5.50-1.png", + "original_path": "designv11-35/openalex_figure/designv11_35_0001419_978-94-007-7401-8-Figure5.50-1.png", + "caption": "Fig. 5.50 3PaPaPR-1RPPaPat-type fully-parallel PM with uncoupled Sch\u00f6nflies motions defined by MF = SF = 4, RF\u00f0 \u00de \u00bc \u00f0v1; v2; v3;xb\u00de, TF = 0, NF = 30 limb topology Pa\\Pa\\\\P\\\\R, Pa\\Pa\\\\P\\kR and R||P||Pa||Pat", + "texts": [ + "4h) R||Pa||Pat||P (Fig. 5.4l) 12. 3PaPaPR1RPaPatP (Fig. 5.46) Pa\\Pa\\\\P\\\\R (Fig. 5.4g) Idem No. 1 Pa\\Pa\\\\P\\kR (Fig. 5.4e) R||Pa||Pat||P (Fig. 5.4l) 13. 3PaPaPR-1RPPaPa (Fig. 5.47) Pa\\Pa\\kP\\\\R (Fig. 5.4f) Idem No. 3 Pa\\Pa\\kP\\kR (Fig. 5.4h) R||P||Pa||Pa (Fig. 5.4m) 14. 3PaPaPR-1RPPaPa (Fig. 5.48) Pa\\Pa\\\\P\\\\R (Fig. 5.4g) Idem No. 3 Pa\\Pa\\\\P\\kR (Fig. 5.4e) R||P||Pa||Pa (Fig. 5.4m) 15. 3PaPaPR-1RPPaPat (Fig. 5.49) Pa\\PakP\\\\R (Fig. 5.4f) Idem No. 3 Pa\\Pa\\kP\\kR (Fig. 5.4h) R||P||Pa||Pat (Fig. 5.4n) 16. 3PaPaPR-1RPPaPat (Fig. 5.50) Pa\\Pa\\\\P\\\\R (Fig. 5.4g) Idem No. 3 Pa\\Pa\\\\P\\kR (Fig. 5.4e) R||P||Pa||Pat (Fig. 5.4n) 17. 2PaPaRRR-1PaPaRR1RPaPaP (Fig. 5.51) Pa\\Pa||R||R\\\\R (Fig. 5.5a) Idem No. 1 Pa\\Pa||R||R (Fig. 5.4i) R||Pa||Pa||P (Fig. 5.4k) (continued) 532 5 Topologies with Uncoupled Sch\u00f6nflies Motions Table 5.4 (continued) No. PM type Limb topology Connecting conditions 18. 2PaPaRRR-1PaPaRR1RPaPaP (Fig. 5.52) Pa\\Pa||R||R\\kR (Fig. 5.5b) Idem No. 1 Pa\\Pa||R||R (Fig. 5.4i) R||Pa||Pa||P (Fig. 5.4k) 19. 2PaPaRRR-1PaPaRR1RPaPatP (Fig" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_35_0000403_s10527-020-10002-w-Figure1-1.png", + "original_path": "designv11-35/openalex_figure/designv11_35_0000403_s10527-020-10002-w-Figure1-1.png", + "caption": "Fig. 1. Diagram of instruments for accessing soft tissues: straight (a) and angled (b) structures.", + "texts": [ + " The surgeon addresses the spinal muscles using various retractors and chisels with a working zone 2 3 cm in width. The estimated range of forces in this type of procedure is 150 200 N and the range of torques is 10 20 N\u22c5m. Instruments for gaining access to the soft tissues should be designed on the basis of such tools as blades for working on tissues in the surgical field. Two designs for interchangeable instruments have been developed: straight and angled. These are diagrammed, along with possible sensors, in Fig. 1. Pressure on the patient\u2019s tissues will create a nonuniform distribution of the load perpendicular to the z axis of the sensor in the working zone of the instrument. The resulting torque depends directly on the shoulder created by the vector of the resultant force applied at a point depending on the distribution of the load. Following the surgeon\u2019s actions, the point of application of the force will move from the center to a distance of up to half the working zone. Knowing the range of measure ments of the torque for the sensor and the length of the working zone, the maximum possible load can be deter mined by locating the vector of the resultant force on the boundaries of the movement of the point at which force is applied", + " This relationship and the design features of the tool gave the optimum dimensions L1, L2, and L3, of 84, 60, and 30 mm, respectively. For these lengths, the limiting loads F1 and F2 are 60.97 and 45.45 N, respectively. Thus, the maximum measurable force for an instrument with straight geometry is 45.45 N. A second design of the instrument is provided to cover the whole range of force values measured by the sensor. In this instrument, an unevenly distributed load is directed parallel to the z axis of the sensor (Fig. 1b) to decrease the force shoulder for the resultant force vector, thus decreasing the torque acting on the sensor. The lim iting measured loads for the two boundary cases of the displacement of the resultant force application point are determined by the relationships: where L1 is the length of the straight part, L2 is the length of the angled part of the instrument, L3 is the length of the working zone, and L4 is the length of the displacement of the point of application of the distributed force. Optimum values of L1, L2, L3, and L4 for this instru ment were found to be 104, 5, 30, and 15 mm, respective ly" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_35_0000321_s40430-021-02964-z-Figure5-1.png", + "original_path": "designv11-35/openalex_figure/designv11_35_0000321_s40430-021-02964-z-Figure5-1.png", + "caption": "Fig. 5 Experimental tests of broken teeth in the detection stage. Source: Author", + "texts": [ + " In the 60\u00a0Hz rotation (supply line and frequency inverter), the wavelet details \u03bb7.0 and \u03bb7.5 were used. The \u03bb7.0 wavelet detail is the SLR, and the \u03bb7.5 detail is the GMR. For the 20\u00a0Hz rotation, the \u03bb8.0 wavelet detail is the SLR and the \u03bb8.3 detail is the GMR. The failure of a broken tooth was obtained by cutting a tooth at its base. Were broken two gears (driving wheel), to observe the repeatability of the results. For each failed gear, five samples of each operating condition were obtained. Figure\u00a05 presents the transmission system with a broken tooth failure. Figure\u00a06 presents a schematic of the procedure used to detect failure. Figure\u00a07 presents the percentage changes in the entropy of each wavelet packet detail (first fifty) with broken tooth failure. In Fig.\u00a07, it is possible to verify that the literature references on the location of the characteristic frequencies of the broken tooth present significant changes in the distribution of the set of entropies. Figure\u00a08 presents the wavelet entropy values for each sample in the broken tooth tests" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_35_0000008_j.autcon.2021.103573-Figure11-1.png", + "original_path": "designv11-35/openalex_figure/designv11_35_0000008_j.autcon.2021.103573-Figure11-1.png", + "caption": "Fig. 11. Wire rope vector between boom and block.", + "texts": [ + " The coordinates of the floating crane and the drive system and mass matrix are defined as follows: pcrane = [ zcrane \u03b8crane ] T ,p = [ pb pc ] pb = [ pb1 pb2 ] T ,pc = [ pc1 pc2 pc3 pc4 ] T , (19) Mp=diag ( Iboom1+(l/2)2mboom1 ,Iboom2+(l/2)2mboom2 ,J1/r1,J2/r2,J3/r3,J4/r4 ) . (20) As the floating crane has six control inputs, the motion of the block is controllable as it has 6-DOF. The block coordinates are defined in the same way as for the gantry crane. In this system, we ignored the motion of the hooks and assumed the block is connected to the boom directly with wire ropes. The wire rope constraint between the boom and block is modeled as restraining the length of wire rope between the connection point of the boom and the block (see Fig. 11). The constraint equation is defined as. \u03a6ij(p, x) = 1 2 ( cij T(p, x)cij ( p , x ) \u2212 pci 2), (21) where cij is the wire rope vector, i is the index of the wire rope from 1 to 4, and j is the index of the boom, which is 1 or 2. The wire rope vector and its derivatives can be written as follows: cij = rlug \u2212 rboom = ( rB + CRBXi ) \u2212 ( rboomj + ERJjYi ) = (rB + di) \u2212 ( rcrane + ERcraneZj + ERJjYi ) , ( Yboomi = [0, 0, li] T ) (22) c\u0307ij = ( r\u0307B +\u03c9B \u00d7di ) \u2212 ( r\u0307crane \u2212 \u0303ERcraneZjey\u03b8\u0307crane \u2212 \u0303ERJjYiey ( p\u0307bj + \u03b8\u0307crane )) , (23) where Zj, Yi, and Xi are the local position vectors of the joint of the boom, j, from the crane, connection point, i, of the boom, and the lug point, i, of the block. li is the distance between the wire connection point of ith boom from the local frame of the boom (see Fig. 11). Finally, the derivative of the constraint is calculated from the equations above. From these equations, the Jacobian of the constraint is obtained for the block, crane barge, and the drive system. Lastly, the equations of motion of the crane barge, drive system, and \u03a6\u0307ij = cij T c\u0307ij \u2212 pcip\u0307ci = 0 = cij T { ( r\u0307B + \u03c9B \u00d7 di ) + ( \u2212 r\u0307crane + \u0303ERcraneZjey\u03b8\u0307crane + \u0303ERJjYiey ( p\u0307bj + \u03b8\u0307crane ))} \u2212 pcip\u0307ci = [ cij T \u2212 cij T d\u0303i ][ r\u0307B \u03c9B ] + [ \u2212 cij T cij T \u0303ERcraneZjey + cij T \u0303ERJjYiey ] [ r\u0307crane \u03b8\u0307crane ] + [ cij T \u0303ERJjYiey \u2212 pci ] \u23a1 \u23a3 p\u0307b p\u0307c \u23a4 \u23a6 = Gs(p, x) [ r\u0307B \u03c9B ] + Gc(pcrane, p, x) [ r\u0307crane \u03b8\u0307crane ] + Gp(pcrane, p, x) \u23a1 \u23a3 p\u0307b p\u0307c \u23a4 \u23a6 (24) H" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_35_0001419_978-94-007-7401-8-Figure3.66-1.png", + "original_path": "designv11-35/openalex_figure/designv11_35_0001419_978-94-007-7401-8-Figure3.66-1.png", + "caption": "Fig. 3.66 Overactuated PMs with coupled Sch\u00f6nflies motions of types 4PPaPR (a) and 4PPaRP (b) defined by MF \u00bc SF \u00bc 4; RF\u00f0 \u00de \u00bc v1; v2; v3;xd\u00f0 \u00de, TF = 0, NF = 18, limb topology P||Pa\\P\\kR (a) and P||Pa\\R\\P (b)", + "texts": [ + "50m) Idem No. 1 14. 4PaRPP (Fig. 3.61a) Pa||R||P\\P (Fig. 3.50n) Idem No. 1 15. 4RPaPP (Fig. 3.61b) R||Pa||P\\P (Fig. 3.50o) Idem No. 1 16. 4RPPaP (Fig. 3.62a) R\\P\\kPa||P (Fig. 3.50p) Idem No. 1 17. 4PPPaR (Fig. 3.62b) P\\P\\kPa||R (Fig. 3.50q) Idem No. 1 18. 4PPPaR (Fig. 3.63a) P\\P||Pa||R (Fig. 3.50r) Idem No. 1 19. 4PPaRR (Fig. 3.63b) P||Pa||R||R (Fig. 3.50s) Idem No. 1 20. 4PPaRP (Fig. 3.64) P\\Pa||R||P (Fig. 3.50t) Idem No. 1 21. 4PPaPR (Fig. 3.65) P\\Pa||P||R (Fig. 3.50u) Idem No. 1 22. 4PPaPR (Fig. 3.66a) P||Pa\\P\\kR (Fig. 3.50v) Idem No. 1 23. 4PPaRP (Fig. 3.66b) P||Pa\\R\\P (Fig. 3.50i0) The before last joints of the four limbs have parallel axes 24. 4RPPaP (Fig. 3.67) R||P||Pa\\P (Fig. 3.50w) Idem No. 1 25. 4RPRPa (Fig. 3.68) R||P||R||Pa (Fig. 3.50x) Idem No. 1 26. 4PaPPR (Fig. 3.69a) Pa||P\\P\\kR (Fig. 3.50y) Idem No. 1 27. 4PaPPR (Fig. 3.69b) Pa\\P\\kP||R (Fig. 3.50z) Idem No. 1 28. 4PaPRP (Fig. 3.70) Pa\\P\\kR||P (Fig. 3.50z1) Idem No. 1 29. 4PaRRP (Fig. 3.71) Pa||R||R||P (Fig. 3.50a0) Idem No. 1 30. 4PaRPR (Fig. 3.72) Pa||R||P||R (Fig. 3.50b0) Idem No. 1 31" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_35_0000805_j.ejcon.2021.08.003-Figure1-1.png", + "original_path": "designv11-35/openalex_figure/designv11_35_0000805_j.ejcon.2021.08.003-Figure1-1.png", + "caption": "Fig. 1. Two-axis gimbal azimuth \u03b11 and elevation \u03b12 correction angles.", + "texts": [ + " Application to inertially stabilized platforms The main control objective in an Inertially Stabilized Platform is o maintain the line-of-sight (LOS) invariant, that is, the direction to hich the object of interest is pointed at, in the presence of intenional or unintentional maneuvers of the host vehicle. This is the ase of cameras in mobile robots such as UAV (unmanned aerial ehicle) and ROV (remotely operated vehicle), where mechanical ompensation of the LOS is usually employed. This task involves estimating the platform attitude, usually usng a set of sensors in an inertial measurement unit (IMU) [12] or ven visual odometry [8] , and compensating this attitude using mechanical assembly. A common choice is the two-axis gimbal 9,25] depicted in Fig. 1 which perform rotations in two angles, amely the azimuth and the elevation. The object to be stabilized s represented by a red rectangle positioned in the inner gimbal nd the correction angles are denoted by \u03b11 (azimuth) and \u03b12 (elvation). In the absence of mechanical unbalances, azimuth and elevaion angles can be considered as two decoupled SISO (single-input, ingle-output) systems for control purposes. However, in the presnce of such unbalances, dynamic interaction between inputs and utputs will cause azimuth commands to affect the elevation and ice-versa [1] " + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_35_0000193_tmag.2021.3067323-Figure3-1.png", + "original_path": "designv11-35/openalex_figure/designv11_35_0000193_tmag.2021.3067323-Figure3-1.png", + "caption": "Fig. 3. Finite-element subdivisions except to air. (a) Simple pole model (b) Full model", + "texts": [ + " Each phase current to output the target torque in an arbitrary posture is calculated by multiplying both sides of (1) by the inverse matrix of Km. When the rotor is rotated, each phase current that outputs the target torque is calculated in each posture. This current is input to the full model in Fig. 1, and a magnetic field analysis calculates the output torque. The analysis accuracy of this analysis method is verified by comparing the output torque with the target torque. In this paper, the difference between the target value and the analysis value is called the error. Fig. 3 shows the finite-element subdivisions in the magnetic field analysis of the simple pole model for a torque constant map analysis and the full model for a torque analysis, and Table II shows the discretization data for the case to analyze the simple pole model and full model. In both models, the element size of the magnetic pole part was set to 1 mm with a function of the mesh generator in JMAG [8]. The torque constant map resolution was 2.3 degrees in the latitude direction and 3 degrees in the longitude direction" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_35_0000168_j.mechatronics.2021.102514-Figure3-1.png", + "original_path": "designv11-35/openalex_figure/designv11_35_0000168_j.mechatronics.2021.102514-Figure3-1.png", + "caption": "Fig. 3. The structure of a typical Delta manipulator. The blue platform is a fixed platform. The three black cylinders indicate three motors. The red platform is a moving platform which has three translational degrees of freedom. (For interpretation of the references to color in this figure legend, the reader is referred to the web version of this article.)", + "texts": [ + " (37) yields \ud835\udc49\ud835\udc56(\ud835\udc61) \u2264 \ud835\udc49 (0)e\u2212\ud835\udefe\ud835\udc61 + \ud835\udf12\u2215\ud835\udefe(1 \u2212 e\u2212\ud835\udefe\ud835\udc61) + \ud835\udeefe\ud835\udefe\ud835\udee5\ud835\udc61 e\ud835\udefe\ud835\udee5\ud835\udc61 \u2212 1 \u2212 \ud835\udeefe\u2212\ud835\udefe\ud835\udc61, (39) where \ud835\udc61 \u2208 (\ud835\udc61\ud835\udc60(\ud835\udc56), \ud835\udc61\ud835\udc60(\ud835\udc56 + 1)). Therefore, we have \ud835\udc52\ud835\udc47 \ud835\udc52 \u22642\ud835\udc49 (0)e\u2212\ud835\udefe\ud835\udc61 + 2\ud835\udf12\u2215\ud835\udefe(1 \u2212 e\u2212\ud835\udefe\ud835\udc61) + 2\ud835\udeefe\ud835\udefe\ud835\udee5\ud835\udc61 e\ud835\udefe\ud835\udee5\ud835\udc61 \u2212 1 . (40) It can be seen from Eq. (40) that \ud835\udc52\ud835\udc47 \ud835\udc52 can be upper bounded such that arbitrary small tracking error can be achieved by choosing \ud835\udefe sufficiently large and \ud835\udf12 and \ud835\udeef sufficiently small. This ends the proof of Theorem 1. A simulation example based on a Delta manipulator is given in this section. The structure of a typical Delta manipulator is shown in Fig. 3. The mechanical model of the Delta manipulator and the definition of the generalized coordinates are given in Fig. 4. The position vector of the end-effector is denoted by [\ud835\udc65\ud835\udc5d, \ud835\udc66\ud835\udc5d, \ud835\udc67\ud835\udc5d]\ud835\udc47 . To build the dynamics equation of the Delta manipulator, the PM is cut into three open-chain systems \ud835\udc5c \u2212 \ud835\udc34\ud835\udc56 \u2212 \ud835\udc35\ud835\udc56 \u2212 \ud835\udc36\ud835\udc56 where \ud835\udc56 \u2208 1, 2, 3 and a moving platform. Each open-chain system has three degrees of freedom, i.e. \ud835\udc5e\ud835\udc561, \ud835\udc5e\ud835\udc562 and \ud835\udc5e . Thus, \ud835\udc5e = [\ud835\udc5e , \ud835\udc5e , \ud835\udc5e , \ud835\udc5e , \ud835\udc5e , \ud835\udc5e , \ud835\udc5e , \ud835\udc5e , \ud835\udc5e , \ud835\udc65 , \ud835\udc66 , \ud835\udc67 ]\ud835\udc47 \u2208 R12. It \ud835\udc563 11 12 13 21 22 23 31 32 33 \ud835\udc5d \ud835\udc5d \ud835\udc5d should be noted that motors are only placed at \ud835\udc341, \ud835\udc342 and \ud835\udc343 such that \ud835\udc5e\ud835\udc4e = [\ud835\udc5e11, \ud835\udc5e21, \ud835\udc5e31]\ud835\udc47 " + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_35_0001419_978-94-007-7401-8-Figure6.5-1.png", + "original_path": "designv11-35/openalex_figure/designv11_35_0001419_978-94-007-7401-8-Figure6.5-1.png", + "caption": "Fig. 6.5 Maximaly regular fully-parallel PMs with Sch\u00f6nflies motions of types 2CPRR-1CPR1RPPP (a) and 2CRRR-1CRR-1RPPP (b) defined by MF = SF = 4, (RF) = (v1; v2; v3;xb), TF = 0, NF = 4, limb topology R\\P\\\\P ??P and C\\P\\||R\\R, C\\P\\||R (a), C||R||R\\R, C||R||R (b)", + "texts": [ + " rGi (i = 1, 2, 3) 0 0 18. rG4 0 0 19. MGi (i = 1, 3) 4 5 20. MG2 4 4 21. MG4 6 6 22. (RF) (v1; v2; v3;xb) (v1; v2; v3;xb) 23. SF 4 4 24. rl 0 0 25. rF 14 16 26. MF 4 4 27. NF 4 2 28. TF 0 0 29. Pp1 j\u00bc1 fj 4 5 30. Pp2 j\u00bc1 fj 4 4 31. Pp3 j\u00bc1 fj 4 5 32. Pp4 j\u00bc1 fj 6 6 33. Pp j\u00bc1 fj 18 20 a See footnote of Table 2.2 for the nomenclature of structural parameters 6.1 Fully-Parallel Topologies with Simple Limbs 589 Table 6.4 Structural parametersa of parallel mechanisms in Figs. 6.5 and 6.6 No. Structural parameter Solution Figure 6.5 Figure 6.6 1. m 13 15 2. pi (i = 1,3) 4 4 3. p2 3 3 4. p4 4 6 5. p 15 17 6. q 3 3 7. k1 4 4 8. k2 0 0 9. k 4 4 10. (RG1) (v1; v2; v3;xa;xb) (v1; v2; v3;xa;xb) 11. (RG2) (v1; v2; v3;xb) (v1; v2; v3;xb) 12. (RG3) (v1; v2; v3;xb;xd) (v1; v2; v3;xb;xd) 13. (RG4) (v1; v2; v3;xb) (v1; v2; v3;xa;xb;xd) 14. SGi (i = 1, 3) 5 5 15. SG2 4 4 16. SG4 4 6 17. rGi (i = 1, 2, 3) 0 0 18. rG4 0 0 19. MGi (i = 1, 3) 5 5 20. MG2 4 4 21. MG4 4 6 22. (RF) (v1; v2; v3;xb) (v1; v2; v3;xb) 23. SF 4 4 24. rl 0 0 25. rF 14 16 26" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_35_0001627_600010-Figure3-1.png", + "original_path": "designv11-35/openalex_figure/designv11_35_0001627_600010-Figure3-1.png", + "caption": "Fig. 3 \u2014 Analogy \u2014 contacting gear teeth to contacting rolls", + "texts": [], + "surrounding_texts": [ + "Many designers use speed factors, or dynamic loading factors to modify gear loading formulas. Compressive stress Eqs. 2 and 3 do not have such factors; as a result, the plotted points and final curve (Fig. 5) have no stress load corrections for VOLUME 68, 1960 55 SSION GEAR LIFE- SURFACE FATIGUE STRESS REPETITIONS VS. um COMPRESSIVE STRESS 11111 FOR DRIVING GEARS 1111111 2200 R.M. TRANSMISSION knikliTik 2 111:1UTISIPIED111111111111 0 1111 :1 11:10112 Iiin:.i 111411111441 100,000 50,000 20,00 \u2022 10,000 5,000 02poo rn IS 1,000 H ttd wa- 500 (r) I-. 200 RANSMI TRANSMISSION GEAR LIFE-SURFACE FATIGUE STRESS REPETITIONS VS. COMPRESSIVE STRESS 00. 1000 100 180 200 250 300 350 400 450 COMPRESSIVE STRESS\"Scil-P.S.1.X103 Fig. 4\u2014 Transmission gear life \u2014 surface fatigue, stress repetitions versus compressive stress for driving gears (2200 rpm transmission input speed) the speed of gears. Points, which determined the higher cycle portion of the curve at the least compressive stresses, are transmission input drive gears. Points which determined the lowest cycle gear life and highest compressive loading are first or low speed gearsets. Between these two extremes are second-, third-, and fourth-speed gearsets. With transmission input speeds of 2200 rpm, pitch line velocities vary from 1825 fpm for the input drive gear to 570 fpm for the first-speed gear. If the curve is used in transmission gear design based on an input shaft speed of 2200 rpm, we believe it has a \"built-in\" speed factor. Future input drive gears will fall on the same part of the curve as for those tested. Future first-, second-, third-, and fourth-speed gears will fall on a portion of the curve which was obtained by gearsets operating at similar speeds. If gears that will operate at speeds considerably different than those tested are to be designed, a speed factor should be used. One of our current projects is the determination of such a speed factor by dynamometer testing. In cases of speed-up ratios where the driving gear member is turning at slower rpm than the output gear, our testing indicated the driving gear failed prior to the driven gear even though the driver had fewer gear tooth stress repetition than the driven. This has been observed by others. D. W. Dudley of General Electric Co. indicates he has performed tests on gear speed-up sets of 4/1 ratio in which the driving gear pitted before the driven pinion. 4 Due to these results, all gear stress repetitions are calculated and plotted for the driving gear member. In many fatigue curves, where stress cycles are plotted against stress, the curve will bend toward 20,000 opoo 5,000 te) 0 x2,000 F.: 1,000 500 cn in ta cc i- 200 cn 100 180 200 250 300 350 400 450 COMPRESSIVE STRESS li So n P.S.I. X 103 Fig. 5\u2014Transmission gear life \u2014 surface fatigue, stress repetitions versus compressive stress for driving gears (2200 rpm transmission input speed) infinite repetitions as the stress is reduced beyond a certain point. We had suspected this might happen near the high cycle end of our curve, as many authorities have indicated pitting will not be encountered at compressive stresses of 200,000 psi and less. Our plotted gear failures did not give an indication of infinite life within the range of compressive stresses tested. Earle Buckingham has described tests made with steel rolls to define the surface endurance limit of case hardened steel. 5 Tests of up to 400,000,000 stress cycles indicated no definite endurance limit. The load limit followed the same line as for tests of heavier loads and fewer cycles. This would indicate our curve could be extended toward high repetitions and low stress with no bend toward infinite life being encountered. We would like to explore this high repetitions low stress area, but have not for the following reasons: 1. It is not necessary as most of our design work falls in the area of the defined curve. 2. A great amount of dynamometer time would be necessary to obtain a sufficient number of points as each point in this area of the curve would require many hours of testing. Some of our plotted points represent 500 hr of dynamometer testing. As loading equations for bearings and gears are both results of Hertz' work on elastic bodies, it is of interest to compare results of testing each. From the literature on bearing life published by the various companies, we find life is inversely proportional 4 \"Practical Gear Design,\" by D. W. Dudley, pp. 288-289. Pub. by McGraw-Hill Book Co., Inc., New York, 1954. 5 \"Analytical Mechanics of Gears,\" by Earle Buckingham, pp. 521-522. Pub. by McGraw-Hill Book Co., Inc., New York, 1949. 50P FOR DRIVING GEARS -L 2200 R. P. M. TRANS. INPUT SPEED 56 SAE TRANSACTIONS = 0.59 x 0.8339 6 x 10 7 x 4732 ( 1 0.940 x 0.5558 3.3333 + 6.1666 ) to the load to a power varying from the third to fourth. Expressed as an equation: Life = K (Lad) (4) where: K = Constant of proportionality x varies in range from 3-4 Some bearing companies use the third power, others the fourth, and the 3.33 power is often used for roller bearings. Setting our results in similar form, we find in terms of stress and cycles, the cycles vary inversely as the compressive stress to the 6.75 power, or: Cycles = K 1 Cycles = K (Load) 38 This 3.38 power is in the range of results found by the various bearing companies by test, and is very close to the 3.33 power which is often used for roller bearings. We believe this tends to prove the analogy of gear teeth to contacting rolls. Upon examining Eq. 3, it will be noticed the compressive stress may be varied by changing the following gear dimension or items: (1) face width, (2) pressure angle, and (3) helix angle. In the past, gear designers regulated compressive stress only by changing the face width. Pressure angles were somewhat standardized and helix angles were determined for desired overlap or permissible end thrusts. The value of careful consideration of pressure angles, and helix angles for compact high capacity gearsets, is now known and utilized in gear design. This paper represents work which is only beginning. Dynamometer testing is now under way to establish speed or dynamic loading factors which will increase the flexibility of the present curve for design use." + ] + }, + { + "image_filename": "designv11_35_0000027_j.procir.2020.05.200-Figure2-1.png", + "original_path": "designv11-35/openalex_figure/designv11_35_0000027_j.procir.2020.05.200-Figure2-1.png", + "caption": "Fig. 2. Overall system consisting of: 2-jaw gripper with formflexible gripper jaws and spray nozzles for the active cooling.", + "texts": [], + "surrounding_texts": [ + "r m t i a t t c j i t\nw t b t s t\nk T 1 r t w p ( t o r a\nt o t\n1\nt r f w c\ns t a T o\n2\nh p\n2\ns i u l V m d t t O t t g o p m b g I a a d t g c s t a ( t\np m t\nS r o o g p t a a g f t t l t h s g a H k\nical components with diameters between 25mm and 90mm and\nasses varying between 0.5kg and 1kg. These have already passed hrough the Tailored Forming process. Each of these components s subject to different heating strategies, manufacturing processes nd material pairings. The demonstrators of the current period of he CRC 1153 are under development. However, it can be seen that he complexity in terms of shape and form is increasing. The most omplex demonstrator is the wishbone Q shown in Fig. 1 . A simple aw gripper does not suffice to handle the different shapes. A flexble universal gripper is needed, as it is characterized by its ability o handle a wide range of different objects.\nA closer look at the manufacturing process of the bevel gear K ill show what conditions the handling must withstand. Two ypes of steel are joined to form a semi-finished product by the uild-up welding process. Heating is carried out by induction and akes place from the outside. The gripper must therefore grip the urface, which is 10 0 0 \u25e6C hot. For such an application special high emperature grippers are necessary.\nFurthermore, the distribution of heat in the component is a ey factor for the quality, as identified by Behrens et al. (2018b) . he temperature of the steel should be in the range from 900 \u25e6C to 200 \u25e6C and the temperature of the aluminium in the range from oom temperature to 300 \u25e6C ( Behrens et al., 2018a ). In this case he flow behaviour of both materials is adapted to each other,\nhich is necessary for the forming process. The necessary tem-\nerature gradient could not yet be achieved with inductive heating Behrens et al., 2018b ). The aim of the cooling system integration is o increase the temperature gradient. To achieve the aim, flexibility f both the cooling performance and the cooling system design is equired in order to ensure adaptation to different workpiece sizes nd shapes.\nMoreover, investigations have shown that due to heat conducion in the material, the transport time has a significant influence n the temperature distribution. Thus, the aim is to minimize the ransport time or adaptive cooling ( Behrens et al., 2016 ).\n.2. Objectives\nFrom the explanations on the process of Tailored Forming and he associated boundary conditions (section 1.1), the following equirements and objectives can be deduced: Development of a orm-flexible handling system for Tailored Forming components\nith surface temperatures up to 1250 C with an integrated active\nooling of the components.\nThis paper will show how such a system can be designed. The ub-systems form-flexible gripper jaw, cooling system and the auomated process of gripping are developed for this purpose. After\ngeneral introduction, the sub-systems are presented successively.\nhey complement each other to form a system which fulfils the bjectives.\n. State of the art\nThis chapter provides an overview of variable shape gripping, eat-resistant gripping and cooling techniques. The information resented serves as a base for the development of a new system.\n.1. Universal grippers\nUniversal grippers are characterized by the fact that they are uitable for several assembly and handling tasks due to their flexbility. The flexibility of universal grippers is realized through the\nsing of grippers that are adaptable in shape. The adaptability al-\nows the grippers to adjust automatically to different geometries.\narious effects are used for this purpose. Since the entire the-\natic breadth of universal grippers would be far too broad to be\niscussed in detail in this paper, some grippers that are imporant for the paper are mentioned. A good overview of the enire field of universal grippers is given by Fantoni et al. (2014) .\nn the one hand, there are grippers which consist of highly elas-\nic materials and adapt to the shape of the object to be gripped, he so called soft robotic grippers. An overview of this type of ripper is given by Shintake et al. (2018) . There are two types f soft grippers, those that combine elastic materials and comressed air ( Ilievski et al., 2011 ) and those that use Granular Jam-\ning ( Brown et al., 2010 ). When compressed air is used, the elastic\nody of the gripper is deformed in such a way that the object is ripped or embraced, resulting in a form and force closure grip. n Granular Jamming, there is an elastic skin which is filled with\ngranulate. In the initial state, the granulate is permeated with ir and deformable. The skin is applied to the object to be hanled and thereby assumes the contour of the object. The skin is hen evacuated. The air is extracted from the granulate and the ranulate is stiffened. In this way, the skin retains its assumed ontour and a form-locking grip is performed. The materials used, uch as silicones or other polymers, are limited in their application emperature. For elastic variable shape grippers and their materils, a maximum operating temperature of 300 \u25e6C has been observed Mosadegh et al., 2014 ). For this reason, it is not possible to use hem in the high-temperature range of Tailored Forming.\nA further type of variable shape grippers are those which reroduce the contour of the object to be gripped by moving ele-\nents and thereby enable a form closure. The advantage here is\nhat metal materials are used instead of silicone.\nOne example of such grippers is the Omnigripper by cott (1985) . In the Omnigripper, pins are arranged in two sepaate matrices. The pins can be moved axially independently of each ther. At first the pins are extended by compressed air. Then, withut compressed air, the gripper is applied to the object and the finers that come into contact are retracted again. This creates a griper that is specially adapted to the contour. To perform the grip, he two matrices are moved together and there is a force-locking nd form-locking connection. Another gripper that uses pins in\nmatrix arrangement to grip with variable shapes is the matrix ripper by Matrix GmbH ( Meinstrup, 2013 ). Here, two of the jaws ace each other. The pins are pushed out by a spring. Upon conact with the object to be gripped they retract again. As soon as he negative is ready, the pins are locked and thus the image is ocked. Another approach that is being researched is the reproducion of the human hand as a universal gripping tool. Such technical\nands have 3 to 5 joint fingers, with one of the fingers being deigned as an independently movable thumb to achieve a high deree of flexibility. Examples of highly developed technical hands re the DLR-Hand II ( Butterfass et al., 2001 ) and the Robonaut-\nand 2 ( Bridgwater et al., 2012 ). These are very complex in their inematics and meet requirements that are not set by the process.", + "c p K t p e o i w l\nm ( p T i t w h u p i\n2\np s c o\nc t a a I e w\nd v a v h ( l t g A s g t t d fi g p c a\n3\nm F\ni a t m l\n3\na p M a w d r o g t s e a p m fi t c c\nw\nt m i l e i a\nt p c c t t\nThe grippers shown so far are variable in their shape. In ontrast, the most important design-elements of a high temerature resistance gripper are characterized by Cutkosky and urokawa (1983) . Their work shows that the fingers or jaws are he thermally most stressed elements. The choice of material is articularly important. Depending on the material, there are differnt mechanical and thermal properties, such as thermal resistance r thermal conduction. The gripper developed by Cutkosky et al. s adaptable to uneven surfaces. For this purpose, fingers are used\nhich are independently beared. However, the shape variability is\nimited compared to the elastic grippers.\nSummarizing, the highest form variability is achieved by elastic aterials, because they can reversibly perform high deformations Mosadegh et al., 2014 ). These materials belong to the group of olymers that have a maximum working temperature of 300 \u25e6C . In ailored Forming, temperatures of up to 1250 \u25e6C are reached, makng the use of polymer based materials impossible. Grippers for his high-temperature range are made from metals, in order to\nithstand temperatures above 10 0 0 \u25e6C . These materials, in turn,\nave no form flexibility and therefore metal grippers must make s of rigid jaws or fingers. Overall, the metal grippers with sliding ins are more suitable for the temperature range in tailored form-\nng due to their simple kinematics and the material used.\n.2. Cooling techniques\nAs outlined before, the cooling is necessary to adjust the temerature gradient in the object. The temperature distribution is esential for the quality of the formed parts. Therefore, the different ooling techniques are explained to provide a basis for the choice f method.\nBasic cooling techniques are immersion cooling, spray cooling, ooling with gas and cooling by conduction in solids of different emperatures. In immersion cooling, the workpieces are cooled in\nbath of a liquid. Spray cooling is the spraying of a workpiece with cooling medium, using single or dual fluid nozzles for spraying. f several solids of different temperatures come into contact with ach other, conduction occurs and the temperatures equalize, the\norkpiece with the higher temperature is cooled.\nThe immersion cooling achieves the highest cooling effect. The isadvantage is that the fluid must be in a reservoir. This preents permanent cooling because it is difficult to move. Therefore n integration into the gripper is not possible. The outstanding adantage of spray cooling is the flexible and simple change of the eat transfer in a wide spectrum by varying the media pressures Golovko et al., 2014; Herbst et al., 2015 ) as well as the easy instalation and low weight of the nozzles. This allows for an adaptation o different applications. Selected areas can be sprayed with the enerated spray cone, which increases the geometric adaptability.\ndisadvantage is the need for a liquid coolant flow and the necesity to remove the coolant after the cooling process. Cooling with as has the advantage that harmless gases such as air do not have o be removed from the environment. However the achievable heat ransfer coefficients are relatively low ( Stark et al., 2011 ). In conuction with contact of solids, the workpiece can be cooled at de-\nned areas on the workpiece, but solid structures are disadvantaeous due to the limited geometric flexibility. Still, spray cooling is articularly suitable as a cooling technology for the forming proess due to the adaptability of the cooling effect over a wide range nd the geometric flexibility.\n. Development of a form-flexible gripper with active cooling\nThis section introduces the solution concepts for the gripper echanism and the cooling system. A brief look at the system in ig. 2 shows that the system consists of the gripping unit, which\nn turn consists of the two formflexible gripper jaws and a parllel jaw gripper. Furthermore, the cooling system, which includes he nozzles and their holder, is shown. The design of the gripper\nechanics and the active cooling system are presented in the fol-\nowing.\n.1. Gripper mechanics\nIn order to meet the demands of the Tailored Forming process, novel gripper was designed. As shown in Fig. 3 , the gripping rinciple of the Omnigripper, pins actuated by air pressure and the\natrixgripper, opposing gripper jaws, is combined. The jaws have matrix with 7x5 pins each with a stroke of 24mm, which are ell suited for the demonstrators in Tailored Forming ( Fig. 3 ). The imension of one jaw are 86x66x60 mm and can be extended as equired. The independently movable pins allow differently shaped bjects to be handled. Fig. 3 a shows the workpiece of the bevel ear clamped in the gripper. The other pictures of the figure show he gripped demonstrators of the first period of CRC 1153. The hape variability is given by the gripping principle. The pins are xtended by compressed air. The jaws are then closed by the parllel gripper without compressed air being applied to the pins. The ins, which will face resistance, are retracted during further move-\nent of the jaws. When the jaws have reached their position de-\nned by the user, compressed air is applied again. The pins act on\nhe object with a force regulated by the compressed air and seure the grip. To release the grip, a vacuum is applied instead of ompressed air. The pins retract and the object is released.\nWith this, the functionality to grip differently shaped objects\nith one flexible gripper can be provided.\nThe next requirement to the system is the temperature resisance. To withstand the high temperatures acting on the pins, a\naterial research was carried out. A special stainless steel has been dentified which is designed for high temperature use. This stainess steel, called 1.4148, is used for the pins and has long term oprating temperatures of up to 1150 \u25e6C and a low thermal conductivty coefficient ( Brnic et al., 2015 ). Higher temperatures are possible t short term.\nThe poor thermal conductivity of the material has the effect hat the heat input into the rest of the system is low. For this urpose, simulations have been carried out with Ansys Mechanial which take into account the process times of the handling and ompare the associated heating of the system with the resulting hermal expansion. To simplify the simulation, only heat conducion was taken into account. Due to the fact that there were no", + "s n t a\n3\nw s u v u w i i\n3\nr c c i v c n z s s d m e i v a e c d b\no n t m\ni m z o i u F c T a c p t m a\nl i m l b\n3\ne s t s c i t m p f 7 a i c f f i F 6 t\nignificant aspects besides the expected expansion, the results are ot presented in detail in this work. Following from these results, he fits of the pin bearings are adapted to the thermal expansion t operating temperature.\n.2. Active cooling\nAfter the gripper is introduced, the cooling system follows, hich is mainly based on spray cooling. The choice fell on a twotage nozzle, as this allows for cooling with either a gas or a liqid or both simultaneously. The heat transfer can be adjusted by arying the pressure of the media or the ratio of gas to liquid. By sing several nozzles and a flexible arrangement of these, different\norkpiece areas can be cooled. The combination of air and water s used as a coolant for the two-substance nozzle. The advantage s the good availability of the media as well as the low costs.\n.2.1. Investigation of nozzle arrangement\nTo determine the arrangement of the nozzles, an analysis is caried out. This will show whether the setting parameters are suffiient to adjust the nozzles individually. The projection of the spray ones onto the surface of the workpieces before and after formng is simulated using the software Blender. The objective of inestigation is to identify and maximize the area reached with the ooling fluid while simultaneously achieving a maximum homoge-\neous cooling surface. Fig. 4 shows an example of possible nozle configurations for the bearing bushing. Surfaces covered by the pray cone are shown in light grey, areas with higher spray denities in lighter grey scales. The heat transfer also decreases raially from the spray center ( Puschmann, 0 0 0 0 ). For a more ho-\nogeneous heat transfer, an overlapping of the spray cones in the dge area is advantageous. Several configurations are considered, n which the number, position and orientation of the nozzles were aried. The position and rotation of the nozzles can be adjusted to dapt to different workpiece sizes and shapes.. In addition, differnt arrangement configurations can be compared and the cooling an be optimized. The translatory adjustment l of the nozzles is\none manually via screw drives or slotted holes. The angles \u03c6can e adjusted via swivel joints. Both parameters are marked in Fig. 4 .\nIn addition to the already mentioned parameters, the number f nozzles is also varied. Arrangements with four nozzles and eight ozzles are considered. By using four nozzles ( Fig. 4 a), the disance to the workpiece is greatest compared to the other arrange-\nents, which results in a lower spray density. In addition, there\ns a greater overlap of the spray cones.As a result, this arrange-\nent is not useful. In Fig. 4 b eight nozzles are rotated around the -axis to reach areas behind the pins that are left untouched in ther arrangements. The disadvantage of this solution is that areas nside the semi-finished product are sprayed directly or there are\nnsprayed areas in the edge area. On the semi-finished product in ig. 4 c the spray cones overlap in the edge area and on the formed omponent there are unsprayed areas between the spray cones. he variant in Fig. 4 c can be modified in such a way that there re no unsprayed surfaces between the spray cones in the formed omponent, which leads to a larger overlap on the semi-finished roduct. This would resemble Fig. 4 b, but without rotation around he z-direction. The objective is therefore to find a good compro-\nise between homogeneous spraying of the semi-finished product\nnd the formed workpiece.\nIt should be noted that the spray cone is approximated by a ight cone. The interaction with the environment cannot be anticpated in this simplified simulation. Therefore, experimental deter-\nination of the ideal nozzle arrangement is needed. Due to turbuences, surfaces not located in the ideal spray cone can be affected\ny the cooling medium.\n.2.2. Simulation ofcooling process\nAfter investigating the arrangement of the nozzles, the influnce of the spray cooling must be analyzed. For this purpose a imulation in Matlab was set up. For the simulation of the heat ransfer, material properties such as the thermal conductivity \u03bb, the pecific heat capacity c, the material density \u03c1and the heat transfer oefficient \u03b1must be considered. The following heat transfer model s to be considered as the energy balance for heat conduction hrough an infinitesimal non-moving volume. To describe the ther-\nal effects of the system, the model ( Eq. (1) ) is solved. For sim-\nlification, temperature independent material properties are used\nor the simulation. The used average heat transfer coefficient of\nkWm \u22122 K \u22121 is achieved according to Golovko et al. (2014) at an ir pressure of 0.2MPa and water pressure of 0.1MPa and can be ncreased by increasing the pressure, which in turn increases the ooling effect. The temperature distribution is important, especially or the bearing bushing blank, because it is inductively heated rom inside. The problem is that the aluminium on the outside s heated strongly by the thermal conduction, as can be seen in ig. 5 a. The aluminium comes close to its melting temperature of\n50 \u25e6C . By heating with activated cooling, however, the temperaure of the aluminium remains within a normal range of tempera-" + ] + }, + { + "image_filename": "designv11_35_0000213_j.triboint.2021.106993-Figure2-1.png", + "original_path": "designv11-35/openalex_figure/designv11_35_0000213_j.triboint.2021.106993-Figure2-1.png", + "caption": "Fig. 2. Journal-Bearing system (a) geometric parameters depicted with hydrodynamic pressure and shear stress, and (b) degrees of freedom.", + "texts": [ + " Direct lubrication lowers power loss significantly, and it is included to show that the proposed TPJB changes can even further reduce power loss. The proposed method is simply to include a pocket and a step in each upper J. Yang and A. Palazzolo Tribology International 159 (2021) 106993 pad, as illustrated in Fig. 1(b). The TPJB is used to support the shaft, and to provide desired stiffness and damping for vibration control. Oil flows from the three nozzles to both side seals and to the thin fluid-film between the journal and the downstream pad. The pressure and shear stress acting on the journal and pad are illustrated in Fig. 2(a). These cause the lifting force and drag force on the shaft, respectively. Important geometric parameters include the bearing clearance (Cl,b), pad clearance (Cl,p), and preload (mpr), as shown in Fig. 2(a). The pressure and shear stress induced forces, and the applied loads, determine the equilibrium values of the journal and pad degrees of freedom (xs,ys,\u03b4 j tilt ,x j pvt), represented in Fig. 2(b). Fig. 3 shows a TPJB with a pocket and step designed to lower drag power loss. The pocket and step are inserted in the upper pad, with the step at the leading edge having a much shorter circumferential length than the downstream pocket. The pocket and step have the dual functions of stabilizing the pad motion at an equilibrium tilt angle, and reducing the drag loss by increasing the vapor volume fraction in the cavitation region. Cavitation is the phase change from liquid to gas, and occurs when the pressure becomes lower than the ambient or saturation pressure", + " (14) and (15). Note that the convection term in the shaft energy equation is included for the shaft rotation effects in the fixed frame. Shaft Energy Equation (Ts): \u2202 \u2202xi ( \u03c1sushtot,s ) = \u2202 \u2202xi \u03bbs \u2202Ts \u2202xi (14) Pad Energy Equation (Ts): 0= \u2202 \u2202xi \u03bbp \u2202Tp \u2202xi (15) us is the shaft rotating speed at each element, and it is calculated by multiplying the element radial location (Re) and angular velocity (ws). The displacements of the shaft translational x and y motions, pad tilting and pad pivot motions shown in Fig. 2(b) are periodically updated during the Newton-Raphson driven search for the equilibrium states. These displacements are applied to the interface boundaries of the mesh deformation equation between the journal and fluid-film and between the pad and fluid-film [16]. Likewise, the displacement solutions in the structure solver modify the interface boundary conditions to include effects of the solid thermal-elastic deformations, and shaft centrifugal force deformations. The mesh deformation equation is expressed in Eq" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_35_0001419_978-94-007-7401-8-Figure3.16-1.png", + "original_path": "designv11-35/openalex_figure/designv11_35_0001419_978-94-007-7401-8-Figure3.16-1.png", + "caption": "Fig. 3.16 Overactuated PMs with coupled Sch\u00f6nflies motions of types 4PPPR (a) and 4PPRP (b) defined by MF \u00bc SF \u00bc 4; RF\u00f0 \u00de \u00bc v1; v2; v3;xd\u00f0 \u00de, TF = 0, NF = 6, limb topology P\\P\\\\P||R (a) and P\\P||R\\P (b)", + "texts": [ + "12a) P||R||R\\P (Fig. 3.1v) The first prismatic joints of the four limbs have parallel directions 18. 4RPRP (Fig. 3.12b) R||P||R\\P (Fig. 3.1q) Idem No. 2 19. 4PRPR (Fig. 3.13a) P||R\\P\\kR (Fig. 3.1r) Idem No. 1 20. 4RPPR (Fig. 3.13b) R||P\\P\\kR (Fig. 3.1s) Idem No. 2 21. 4RPPR (Fig. 3.14a) R\\P\\kP||R (Fig. 3.1t) Idem No. 2 22. 4RPRP (Fig. 3.14b) R\\P\\kR||P (Fig. 3.1u) Idem No. 2 23. 4PPPR (Fig. 3.15a) P\\P\\\\P\\kR (Fig. 3.1h0) Idem No. 9 24. 4PPPR (Fig. 3.15b) P\\P\\\\P\\\\R (Fig. 3.1w) Idem No. 9 25. 4PPPR (Fig. 3.16a) P\\P\\\\P||R (Fig. 3.1x) Idem No. 9 26. 4PPRP (Fig. 3.16b) P\\P||R\\P (Fig. 3.1y) The revolute joints of the four limbs have parallel axes 27. 4PPRP (Fig. 3.17a) P\\P\\\\R||P (Fig. 3.1z) Idem No. 26 28. 4PPRP (Fig. 3.17b) P\\P\\kR\\P (Fig. 3.1a0) Idem No. 26 29. 4PRPP (Fig. 3.18a) P\\R||P\\P (Fig. 3.1b0) Idem No. 26 30. 4PRPP (Fig. 3.18b) P\\R\\P\\kP (Fig. 3.1c0) Idem No. 26 31. 4PRPP (Fig. 3.19a) P||R\\P\\\\P (Fig. 3.1d0) Idem No. 26 32. 4RPPP (Fig. 3.19b) R\\P\\\\P\\\\P (Fig. 3.1e0) Idem No. 26 33. 4RPPP (Fig. 3.20a) R\\P\\kP\\\\P (Fig. 3.1f0) Idem No. 26 34. 4RPPP (Fig. 3.20b) R||P\\P\\\\P (Fig" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_35_0001419_978-94-007-7401-8-Figure2.17-1.png", + "original_path": "designv11-35/openalex_figure/designv11_35_0001419_978-94-007-7401-8-Figure2.17-1.png", + "caption": "Fig. 2.17 4CRRR-type fully-parallel PMs with coupled Sch\u00f6nflies motions defined by MF \u00bc SF \u00bc 4; RF\u00f0 \u00de \u00bc \u00f0v1; v2; v3;xd\u00de, TF = 0, NF = 2, limb topology C\\R||R\\R (a) and C||R||R\\R (a)", + "texts": [ + "1t) The first revolute joints of the four limbs have parallel axes 22. 4RRRPR (Fig. 2.12b) R\\R||R||P||R Fig. 2.1u) Idem No. 21 23. 4RRPRR (Fig. 2.13a) R\\R||P||R||R (Fig. 2.1v) Idem No. 21 24. 4RPRRR (Fig. 2.13b) R\\P||R||R||R (Fig. 2.1w) Idem No. 21 25. 4RRRPR (Fig. 2.14a) R\\R||R||P||R (Fig. 2.1x) Idem No. 21 26. 4RRPRR (Fig. 2.14b) R\\R||P||R||R||R (Fig. 2.1y) Idem No. 21 27. 4RPRRP (Fig. 2.15a) R\\P\\kR\\R\\P (Fig. 2.1z) Idem No. 19 28. 4PRRRP (Fig. 2.15b) P\\R||R\\R\\P (Fig. 2.1z1) Idem No. 19 29. 4PRRPR (Fig. 2.16) P\\R\\R\\P\\kR (Fig. 2.1a0) Idem No. 5 30. 4CRRR (Fig. 2.17a) C\\R||R\\R (Fig. 2.1d0) Idem No. 1 (continued) 2.1 Topologies with Simple Limbs 59 Table 2.1 (continued) No. PM type Limb topology Connecting conditions 31. 4CRRR (Fig. 2.17b) C||R||R\\R (Fig. 2.1e0) Idem No. 13 32. 4RRCR (Fig. 2.18a) R||R||C\\R (Fig. 2.1f0) Idem No. 13 33. 4RCRR (Fig. 2.18b) R||C||R\\R (Fig. 2.1g0) Idem No. 13 34. 4RRRC (Fig. 2.19a) R\\R||R||C (Fig. 2.1h0) Idem No. 21 35. 4RCRR (Fig. 2.19b) R\\C||R||R (Fig. 2.1l0) Idem No. 21 36. 4RRCR (Fig. 2.20a) R\\R||C||R (Fig. 2.1b0) Idem No. 21 37. 4RCRR (Fig. 2.20b) R||C\\R||R (Fig. 2.1c0) Idem No. 21 60 2 Fully-Parallel Topologies with Coupled Sch\u00f6nflies Motions In the fully-parallel topologies of PMs with coupled Sch\u00f6nflies motions F / G1G2-G3-G4 presented in this section, the moving platform n : nGi (i = 1, 2, 3, 4) is connected to the reference platform 1 : 1Gi : 0 by four spatial complex limbs with four or five degrees of connectivity" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_35_0001563_9781119682684.ch1-Figure1.13-1.png", + "original_path": "designv11-35/openalex_figure/designv11_35_0001563_9781119682684.ch1-Figure1.13-1.png", + "caption": "Figure 1.13 (a) When the magnetic field is changing at a uniform rate, then there will be inducement of current which flows in the copper ring of radius r. (b) Even if no ring is in the changing magnetic field, there is an induced electric field E. The four points are shown. (c) The electric fields are induced in every circular imaginary ring. (d) Three similar closed paths that enclose identical areas. Equal EMFs are induced around paths of region 1, which lie entirely within the region of changing the magnetic field. A smaller emf is induced around path 2, which only partially lies in that region. No net emf is induced around path 3, which lies entirely outside the magnetic field [4].", + "texts": [ + " The work done is given by W = \u2212\u222b Fe \u22c5 dl = \u2212qo \u222b E \u22c5 dl (1.50) The potential difference between point A and point B is \u0394V = VB \u2212 VA = W qo = \u2212\u222b B A E \u22c5 dl (1.51) But when the test charge moves from a particular point and comes back to the same point, then we say there is no work is done by Electric force (Fe). Therefore, the closed integral will be equal to zero. \u222e E \u22c5 dl = 0 (1.52) Now, consider a copper ring of radius r is immersed in a gradually increasing magnetic field of radius R whose direction is into the page as shown in Figure 1.13a, an emf (\u03b5) is induced which makes current i to flow in the ring in a counter clockwise direction as per Lenz\u2019s law. The current set up an electric field in the same direction. As stated earlier, the electric field is induced when there is a change of flux even if there is no current flowing. This has been shown in Figures 1.13b,c in that no copper ring is placed in the magnetic field which is changing with time. This induced electric field originates from a point and returns to the same point, or in other words, it has a closed path whereas the electric field, due to static charges, is radial i.e. it originates from a positive charge and ends up a negative charge. Hence, the close line integral of the Electric field due to static charges is zero (Eq. (1.50)). The electric field E, induced due to changing magnetic field B with time, induces EMFs. The imaginary ring outside B induces zero ems, rings near outer edge induces non zero emf but less than the inner rings. This is shown in Figure 1.13d. In Figure 1.13b, a work W has done on a particle qo in moving around the circular path is equal to \u03b5qo, where \u03b5 is the induced emf. Another way of expressing work W done is W = \u222b F \u22c5 dl = ( qoE ) (2\ud835\udf0br) (1.53) where qoE is the magnitude of the force acting on the charge and 2\u03c0r is the distance over which that force acts. Or the Eq. (1.46) can be written as W = ( qoE ) (2\ud835\udf0br) = qo or = 2\ud835\udf0brE (1.54) Rewriting Eq. (1.53) W = \u222b F \u22c5 dl = qo \u222e E \u22c5 dl = qo or = \u222e E \u22c5 dl (1.55) The integral in Eq. (1.54) becomes Eq. (1.53) when Figure 1.13b is considered. According to Faraday\u2019s law or rewriting Eq. (1.37a) = \u2212d\ud835\udf19 dt Therefore, reformulated Faraday\u2019s law equation shall be \u222e C E \u22c5 dl = \u2212d\ud835\udf11 dt = \u2212 d dt\u222b S B \u22c5 dS (1.56) where S is the surface area enclosed by C as shown in Figure 1.14 Faraday\u2019s law can be stated as a time-varying magnetic field that gives rise to an electric field. Specifically, the electromotive force around a closed path C is equal to the negative of the time rate of increase of the magnetic flux enclosed by that path. Hence, Faraday\u2019s law can be applied to any closed path. In Figure 1.13d, it is shown that when there is a change in a magnetic field, then EMFs are induced: the inner most region \u20181\u2019 experiences inducement of emf = \u03b5, outside magnetic field region \u20183\u2019 emf = 0, and region \u20182\u2019 near the outer edge of the changing magnetic field emf< \u03b5. Example 1.26 A conducting wire moves in a plane perpendicular to magnetic field of 40 mT. The length of the wire is 50 cm and the speed of the wire is 10 m/s. Determine (a) the force exerted on an electron in the wire, (b) the electrostatic field E\u20d7 in the wire, and (c) the potential difference produced between the ends of the wire" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_35_0000165_s40430-021-02894-w-Figure7-1.png", + "original_path": "designv11-35/openalex_figure/designv11_35_0000165_s40430-021-02894-w-Figure7-1.png", + "caption": "Fig. 7 Finite element model of the gear drives", + "texts": [ + " The FE models of the gear sets are obtained by the coordinates of the point cloud (for more detailed steps, see literature [35]). For external cylindrical gear sets, the number of teeth involved in meshing usually cannot exceed three. To keep boundary conditions far enough from the tooth loaded areas, the number of teeth in the FE models is greater than or equal to three [9]. Here, five-teeth models are applied for the FE analysis of the gear sets. In this paper, the torsion and bending of the gear set are neglected. Figure\u00a07 shows the FE models of the five-teeth model of the curvilinear gear sets. The element type is the hexahedral first-order (enhanced by incompatible modes to improve their bending behaviour [36]) three-dimensional solid element C3D8I. The nodes on the pinion rim and two sides as \u2018Node set 1\u2032 are rigidly fixed to a reference point (RP-1) on the Z-axis of the pinion by rigid body constraint (the influences of shaft deformation on the gear set are ignored), and the rotation degree of freedom of the reference point (RP-1) around the Z-axis of the pinion is only released" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_35_0001419_978-94-007-7401-8-Figure6.43-1.png", + "original_path": "designv11-35/openalex_figure/designv11_35_0001419_978-94-007-7401-8-Figure6.43-1.png", + "caption": "Fig. 6.43 3PPPaR-1CPaPat-type maximaly regular fully-parallel PMs with Sch\u00f6nflies motions defined by MF = SF = 4, (RF) = (v1; v2; v3;xb), TF = 0, NF = 21, limb topology P\\P\\||Pa||R, C||Pa||Pat and P\\P\\||Pa\\||R (a), P\\P\\||Pa ??R (b)", + "texts": [ + "8b) P\\P\\||Pa||R, (Fig. 4.8c) C||Pa||Pa (Fig. 5.4o) The last revolute joints of limbs G1, G2 and G3 and the revolute joint between links 6D and 8 have superposed axes. The actuated prismatic joints of limbs G1, G2 and G3 have orthogonal directions 2. 3PPPaR-1CPaPa (Fig. 6.42b) P\\P\\||Pa\\\\R (Fig. 4.8a) P\\P\\||Pa||R, (Fig. 4.8c) C||Pa||Pa (Fig. 5.4o) Idem no. 1 (continued) 6.3 Fully-Parallel Topologies with Complex Limbs 643 Table 6.14 (continued) No. PM type Limb topology Connecting conditions 3. 3PPPaR-1CPaPat (Fig. 6.43a) P\\P\\||Pa\\||R (Fig. 4.8b) P\\P\\||Pa||R, (Fig. 4.8c) C||Pa||Pat (Fig. 5.4p) Idem no. 1 4. 3PPPaR-1CPaPat (Fig. 6.43b) P\\P\\||Pa\\R (Fig. 4.8a) P\\P\\||Pa||R, (Fig. 4.8c) C||Pa||Pat (Fig. 5.4p) Idem no. 1 5. 3PPaPR-1CPaPa (Fig. 6.44a) P||Pa\\P\\\\R (Fig. 4.8d) P||Pa\\P\\||R (Fig. 4.8e) C||Pa||Pa (Fig. 5.4o) Idem no. 1 6. 3PPaPR-1CPaPat (Fig. 6.44b) P||Pa\\P\\\\R (Fig. 4.8d) P||Pa\\P\\||R (Fig. 4.8e) C||Pa||Pat (Fig. 5.4p) Idem no. 1 7. 2PPaRRR-1PPaRR1CPaPa (Fig. 6.45) P||Pa||R||R\\R (Fig. 4.9a) P||Pa||R||R (Fig. 4.8g) C||Pa||Pa (Fig. 5.4o) Idem no. 1 8. 2PPaRRR-1PPaRR1CPaPat (Fig. 6.46) P||Pa||R||R\\R (Fig. 4.9a) P||Pa||R||R (Fig" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_35_0000504_tmag.2021.3087267-Figure4-1.png", + "original_path": "designv11-35/openalex_figure/designv11_35_0000504_tmag.2021.3087267-Figure4-1.png", + "caption": "Fig. 4. Field calculation without (left) and with (right) the flux sink. Dotted flux lines are those that would have taken 3D paths.", + "texts": [ + " Therefore, the flux sink concept is introduced. The idea behind this concept is to add an additional edge to the 2D geometry that absorbs those flux lines that would have taken 3D paths. Authorized licensed use limited to: China Jiliang University. Downloaded on June 30,2021 at 09:40:22 UTC from IEEE Xplore. Restrictions apply. 0018-9464 (c) 2021 IEEE. Personal use is permitted, but republication/redistribution requires IEEE permission. See http://www.ieee.org/publications_standards/publications/rights/index.html for more information. Fig. 4 shows how the flux sink is added to the 2D geometry. The flux sink has the same magnetic potential as the stator core in order to sink flux lines originating from the rotor. It is placed at a distance of half a pole pitch \u03c4p/2 from the rotor. The rational behind this placement as well as validation of this field calculation can be found in the previously mentioned paper [1]. D. Using 2D FEA Instead of Conformal Mappings Relying on Schwarz\u2013Christoffel mappings for the field calculation is computationally inexpensive and provides independence from 2D FEA software, which simplifies integration of the proposed pressure finger loss calculation approach into existing dimensioning tool-chains" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_35_0001419_978-94-007-7401-8-Figure5.66-1.png", + "original_path": "designv11-35/openalex_figure/designv11_35_0001419_978-94-007-7401-8-Figure5.66-1.png", + "caption": "Fig. 5.66 3PaPaPaR-1RPaPatP-type fully-parallel PM with uncoupled Sch\u00f6nflies motions defined by MF = SF = 4, RF\u00f0 \u00de \u00bc \u00f0v1; v2; v3;xb\u00de, TF = 0, NF = 39, limb topology Pa\\Pa||Pa\\\\R, Pa\\Pa||Pa||R and R||Pa||Pat||P", + "texts": [ + "1b) 29. 3PaPaPaR-1RPaPaP (Fig. 5.63) Pa\\Pa||Pa\\kR (Fig. 5.6a) Idem No. 1 Pa\\Pa||Pa||R (Fig. 5.6b) R||Pa||Pa||P (Fig. 5.4k) (continued) 5.1 Fully-Parallel Topologies 533 Table 5.4 (continued) No. PM type Limb topology Connecting conditions 30. 3PaPaPaR-1RPaPaP (Fig. 5.64) Pa\\Pa||Pa\\\\R (Fig. 5.6c) Idem No. 1 Pa\\Pa||Pa||R (Fig. 5.6b) R||Pa||Pa||P (Fig. 5.54k) 31. 3PaPaPaR1RPaPatP (Fig. 5.65) Pa\\Pa||Pa\\kR (Fig. 5.6a) Idem No. 1 Pa\\Pa||Pa||R (Fig. 5.6b) R||Pa||Pat||P (Fig. 5.4l) 32. 3PaPaPaR1RPaPatP (Fig. 5.66) Pa\\Pa||Pa\\\\R (Fig. 5.6c) Idem No. 1 Pa\\Pa||Pa||R (Fig. 5.6b) R||Pa||Pat||P (Fig. 5.54l) 33. 3PaPaPaR-1RPPaPa (Fig. 5.67) Pa\\Pa||Pa\\kR (Fig. 5.6a) The revolute joint between links 7D and 11 of G4 limb and the last revolute joints of limbs G1, G2 and G3 have superposed axes Pa\\Pa||Pa||R (Fig. 5.6b) R||P||Pa||Pa (Fig. 5.4m) 34. 3PaPaPaR-1RPPaPa (Fig. 5.68) Pa\\Pa||Pa\\\\R (Fig. 5.6c) Idem No. 33 Pa\\Pa||Pa||R (Fig. 5.6b) R||P||Pa||Pa (Fig. 5.54m) Table 5.5 Limb topology and connecting conditions of the fully-parallel solutions with no idle mobilities presented in Figs" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_35_0000126_icict50816.2021.9358674-Figure2-1.png", + "original_path": "designv11-35/openalex_figure/designv11_35_0000126_icict50816.2021.9358674-Figure2-1.png", + "caption": "Fig. 2. Structure of PUMA 560 manipulator", + "texts": [ + " To overcome the initial high jerk value, a sinusoidal jerk profile provides a smooth motion and increases life of the manipulator in [21]. Therefore this paper uses synchronized S-curve approach to generate a path for PUMA-560 manipulator such that it can min imize the jerk while it reaches to its destination. II. PROBLEM DESCRIPTION This paper considers a 6-DOF PUMA-560 which is an industrial robot having six axis joints such as waist, shoulder, elbow, wrist pitch, wrist roll, wrist yaw which is shown in figure 2. Brushed DC servo motors are used for activating all the six joints in this robot and 2.5 kg of nominal load with 0.1mm positional repeatability can be operated by using end effector of PUMA-560. The aim of the robot manipulator is to serve pick and place/assembly operations. PUMA-560 robot work space is defined around 0.92m from the Centre to wrist and the end effector can reach maximum velocity of 1m/s. To develop an algorithm for smooth trajectory with minimum jerk and optimum t ime using synchronized Scurve trajectory and intelligent computation of inverse kinematics using Artificial Neural Networks for a PUMA560 robot manipulator for performing pick and place/assembly operations" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_35_0001419_978-94-007-7401-8-Figure3.36-1.png", + "original_path": "designv11-35/openalex_figure/designv11_35_0001419_978-94-007-7401-8-Figure3.36-1.png", + "caption": "Fig. 3.36 Overactuated PMs with coupled Sch\u00f6nflies motions of types 4RRRPR (a) and 4RRRRP (b) defined by MF \u00bc SF \u00bc 4; RF\u00f0 \u00de \u00bc v1; v2; v3;xd\u00f0 \u00de, TF = 0, NF = 2, limb topology R||R\\R||P||R (a) and R||R\\R||R||P (b)", + "texts": [ + " 4RRRRP (Fig. 3.34a) R\\R||R\\R\\P (Fig. 3.3o) The first and the fourth revolute joints of the four limbs have parallel axes 29. 4PRRRR (Fig. 3.34b) P\\R||R\\R||R (Fig. 3.3p) The second joints of the four limbs have parallel axes (continued) 244 3 Overactuated Topologies with Coupled Sch\u00f6nflies Motions Table 3.2 (continued) No. PM type Limb topology Connecting conditions 30. 4RPRRR (Fig. 3.35a) R\\P\\kR\\R||R (Fig. 3.3q) Idem No. 16 31. 4RRRPR (Fig. 3.35b) R||R\\R||P||R (Fig. 3.3r) Idem No. 16 32. 4RRRPR (Fig. 3.36a) R||R\\R||P||R (Fig. 3.3s) Idem No. 16 33. 4RRRRP (Fig. 3.36b) R||R\\R||R||P (Fig. 3.3t) Idem No. 16 34. 4RRRRP (Fig. 3.37a) R||R\\R||R\\P (Fig. 3.3u) Idem No. 16 35. 4PRRRP (Fig. 3.37b) P\\R\\R||R\\P (Fig. 3.3v) Idem No. 29 36. 4RRPRP (Fig. 3.38a) R\\R||P||R\\P (Fig. 3.3w) Idem No. 16 37. 4RPRRP (Fig. 3.38b) R\\P||R||R\\P (Fig. 3.3x) Idem No. 16 38. 4RRPRP (Fig. 3.39a) R\\R||P||R\\P (Fig. 3.3y) Idem No. 16 39. 4RRRPP (Fig. 3.39b) R\\R||R||P\\P (Fig. 3.3z) Idem No. 16 40. 4PRRRP (Fig. 3.40a) P||R||R\\R\\P (Fig. 3.3z1) The fourth joints of the four limbs have parallel axes 41" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_35_0000596_s43452-021-00255-x-Figure12-1.png", + "original_path": "designv11-35/openalex_figure/designv11_35_0000596_s43452-021-00255-x-Figure12-1.png", + "caption": "Fig. 12 von Mises stress contour for the specimen of Sr 9", + "texts": [], + "surrounding_texts": [ + "1 3\nhigher than 4-rivets. Approximately between 4 up to 40% of the hybrid joint\u2019s strength increases. Also, according to Fig.\u00a08b, the rivet-adhesive hybrid connection has the highest strength compared to the two rivet and adhesive joints. When the nanocomposite reaches its maximum shear strength, it does not reach the shear stress yield point of the rivet.\nIn the present research, numerical simulations have been performed using ABAQUS 6.14\u20132 software, which is one of the most powerful software in performing connection simulations. Schematic of the double-lap joints geometry demonstrated in Fig.\u00a09. The overall geometry is designed, meshed, and tensile-loaded as shown in Fig.\u00a010 on a metric scale as a simulation of a built-in connection. Also, material properties for modeling in software as follow: Elastic modulus (GPa) for Polypropylene nanocomposite (Plates), Al (Rivet), and Adhesive (Hot Melt 571) are 1.9, 70, and 1.3 GPa, respectively. Also, Poison\u2019s ratio for Plates, Rivet, and Adhesive\nis 0.36, 0.33, and 0.34, respectively. For the adhesive layer plane, the COH3D8P element of the 3D adhesive layer element and 8-node with 3 degrees of transfer freedom in each node is used. For the adhesive layer, the sweep technique is used in the mesh and the optimal size of the adhesive elements is selected 0.0135\u00a0mm.\nIt is very important to ensure the accuracy of the results of mesh analysis. In this study, we used 8-nodes and 3D brick elements (rectangular cubes) for segmentation, also, hexagonal or prismatic elements for meshing area around rivet holes on the plates, and tetrahedron 10-node elements or tetrahedron element for rivet head. The convergence check of the analysis results is provided for the Sr911, and the optimal mesh elements are similarly selected for all test samples. Here, 174,214 elements have been used because of the convergence of the solution to the problem as well as the appropriate number of meshes that speed up the analysis (see Fig.\u00a011).\nNumerical analysis for the elastic region shows that the greatest stress concentration occurs around the rivet holes. As the number of rivets increases, the joint stiffness increases, and as a result, the maximum stress in the structure decreases. The von Mises stress contour and stress distribution in the internal layer of the joint, 9-rivet layout, are shown in Figs.\u00a012 and 13.\nAccording to Fig.\u00a0 14a observed that the maximum von Mises stress on the rivet equals 74.25\u00a0MPa. Also, in Fig.\u00a014b, the highest amount of the elastic stress distribution occurs in a lateral surface of holes (53.66\u00a0MPa). Results of simulation for the double-lap adhesive joint using ABAQUS software was showed that the highest amount of stress in the adhesive layer equal 17.61\u00a0MPa and occurs around holes and external edges (Fig.\u00a015a, b).", + "1 3", + "1 3\nIn the simulation of an adhesive joint, what is seen is the uniformity of tension that occurs at the surface of the joint. Also, creating tension concentration around the rivet holes in the structure is a feature of these rivet joints. In the composite bonding of the adhesive layer between the nanocomposite plates, it has withstood a large amount of stress due to loading. Due to its flexibility in the direction of tension, increasing the length of the adhesive layer is expected and, in the end, the damage that occurs is not from the adhesive area. The value of maximum stress (MPa) in the simulated models for the rivet and adhesive joints is illustrated in Fig.\u00a016. To compare the experimental and numerical results, the arrangements of 31 and 32 experimental rivets with the results of numerical simulations are shown in Fig.\u00a017. Due to these curves, good compatibility is observed between the two methods. In both simulated and experimental modes\nin the elastic region and before failure occurs, the greatest stress is on the nanocomposite and the onset of damage occurs from the adherend.\nIn the process of modeling the joint structure, after stretching, the maximum stress values are extracted, which generally decreases with increasing the number of rivets, which makes the joint more reliable; because the stress concentration decreases with the strength of the connection. The highest stress on the rivet joint was shown in the Sr32 specimen and the lowest amount of von Mises stress was observed in the Sr9 specimen.\nIn Table\u00a02, internal plates of rivet joints after deformation for the elastic zone (Sr9) for experimental and numerical methods with the value of notches are shown. The value of the normal max stress of adhesive-rivet joints, max shear stresses for internal plate and upper and lower plates are shown in Table\u00a03. Table\u00a04 observed that the created" + ] + }, + { + "image_filename": "designv11_35_0000381_tro.2021.3070102-Figure1-1.png", + "original_path": "designv11-35/openalex_figure/designv11_35_0000381_tro.2021.3070102-Figure1-1.png", + "caption": "Fig. 1. (a) Representative nonplanar surface contact is created when a soft parallel-jaw gripper deforms to a cup while grasping. Bottom: The extracted nonplanar contact surface and the pressure distribution, where redder colors represent higher pressure. (b) Three-dimensional projection of the proposed elliptical (upper) and quartic (bottom) 6-D limit surface model, which approximates the 6DFW limit of a nonplanar surface contact.", + "texts": [ + "com/martinajingyixu/non- planar-surface-contact typically used to plan grasps [1]\u2013[5] or combined with learning techniques to detect grasps [6] and to predict the success of manipulation tasks [7]. Deformable jaws or grippers covered with compliant materials [8] are widely deployed in grasping applications as they deform to the local geometry of the object and can better resist external disturbances. Such soft-finger grasps result in nonplanar surface contacts if the local geometry of the object is nonplanar, as shown in Fig. 1(a), and the frictional wrench, a vector that is composed of the frictional force and torque, is 6-D as the frictional force and torque are in three dimensions, respectively. Existing work in grasp planning [9]\u2013[12] typically assumes a planar contact area and uses a so-called limit surface (LS) [13], which describes all possible 3-D frictional wrenches that can be transmitted through a planar contact. Such planar contact models neglect the nonplanar surface geometry and only consider a 3-D subset of the 6-D frictional wrench (6DFW), potentially lead to an overly conservative friction estimation", + " As friction depends on the relative motion between two bodies in contact, we derive the 6DFW for a given instantaneous motion of the grasped object. However, for many robot grasping applications, the relative motion caused by external disturbances during the manipulation is unknown at the time of grasp planning. Therefore, we propose the 6-D limit surface (6DLS), generalized from the 3DLS, to represent the 6DFW limit for the nonplanar surface contact. We further present an ellipsoid and a quartic model as a low-dimensional representation to approximate a 6DLS. Fig. 1(b) illustrates a 3-D projection of each of the above 6DLS models. We apply the 6DLS models to predict physical grasp success by building a grasp wrench space (GWS) and compare the prediction results with two 3-D planar contact models [14], [15] and two 3-D nonplanar surface contact models from our previous work [16]. This article makes the following main contributions. 1) A concept of the 6DLS, generalized from the 3DLS, to represent the 6DFW limit for a nonplanar deformable contact. 2) Two models to approximate a 6DLS and a pipeline to compute the models, based on the given contact surface and pressure distribution" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_35_0000074_iros45743.2020.9341423-Figure2-1.png", + "original_path": "designv11-35/openalex_figure/designv11_35_0000074_iros45743.2020.9341423-Figure2-1.png", + "caption": "Fig. 2: Hardware design. (a) Overview of the manipulator. Fingers are driven by tendons controlled by three servomotors. The manipulator is attached to a UR3 robot arm. The fingertips are covered with blue silicon films. (b) Design of the thumb and index finger. Tendons are fixed to the holes using screws. The black parts are made of nylon, whereas the white parts (fingertips) are made of another plastic. Bellows-like structures function as flexure joints. (c) Three types of behaviors that compose the in-finger manipulation. The flexure of the index finger leads to parallel sliding motion, side motion of the index leads to lateral sliding, and rotation of the servomotor at the thumb base leads to twisting of the fingertip surface. These behaviors are combined during manipulation.", + "texts": [ + ", enabling sliding motion orthogonal to a manipulator is difficult). Restriction in finger motion is disadvantageous for dexterous in-finger manipulation. Zhou et al. proposed a robotic hand made of soft material for dexterous manipulation [27]. The material compliance simplified the tasks of moving an object in the fingers without dropping. Inspired by this idea, we propose a simple tendon-driven finger design that allows finger motion in multiple directions to achieve the screw-picking task. As shown in Fig. 2(a), the designed manipulator has two fingers: a thumb and an index finger. The fingers are driven by tendons attached to servomotors, as shown. The tactile sensor arrays described later are attached to the fingertips. The details of the finger design are presented in Fig. 2(b). The fingers are made of nylon and fabricated using a 3D printer (Markforged, Inc.). There are no joint assemblies in the fingers. Instead, bellows-like compliant structures function as \u201cflexure joints.\u201d Because of this design, fingers can be batch molded, thus simplifying the fabrication process. The white parts in Fig. 2(b) are the removable fingertips made of plastic. Tactile sensor arrays, explained in the following section, are attached to these fingertips. The index finger has two DoFs, namely, flexure motion and side motion. A tendon fixed to the \u201cflexure hole\u201d in Fig. 2(b) generates flexure motion (i.e., bending around the \u201cflexure bellows\u201d). Two tendons fixed to the \u201cside holes\u201d generate side motion (i.e., bending around the \u201cside bellows\u201d). The thumb has two DoFs, namely, twisting generated by a servomotor right under the thumb in Fig. 2(a) and warping around the \u201cwarp bellows.\u201d The latter DoF is entirely passive, i.e., warping occurs only when the index finger firmly presses the thumb. This passive compliance provides continuous pressure between the fingertips, which ensures that screws do not easily drop during manipulation. Furthermore, it is the key structure to enable parallel surface sliding motion, which is a vital manipulation behavior for screw separation. Note that the warp bellows do not allow a motion in the lateral direction", + " On the other hand, it should be soft enough to realize the sliding motion with the adopted servo motor. The stiffness affects the picking performance, and this relationship should be investigated in future work. A flexible stretch sensor (Images Scientific Instruments, Inc.) was attached to the thumb to obtain the magnitude of the warping. The warp magnitude, together with the angle of the flexure servomotor, implicitly tells a controller the gap between the fingers, which conveys information about the screw configuration in the fingers. Figure 2(c) shows the following three manipulation behaviors generated by the compliant robotic fingers presented. \u2022 Parallel sliding: The sliding motion of the fingertip surface in the parallel direction as the index finger presses the thumb to warp. \u2022 Lateral sliding: The sliding motion of the fingertip surface in the lateral direction as the index finger generates side motion. \u2022 Twisting: Generated by the twisting motion of the thumb. This motion leads to changes in the pressure distribution on the surface" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_35_0001419_978-94-007-7401-8-Figure6.9-1.png", + "original_path": "designv11-35/openalex_figure/designv11_35_0001419_978-94-007-7401-8-Figure6.9-1.png", + "caption": "Fig. 6.9 Maximaly regular fully-parallel PMs with Sch\u00f6nflies motions of types 2PRPRR1PRPR-1RPaPaP (a) and 2PRPRR-1PRPR-1RPaPatP (b) defined by MF = SF = 4, (RF) = (v1; v2; v3;xb), TF = 0, NF = 10, limb topology P||R\\P\\||R\\R, P||R\\P\\||R and R||Pa||Pa||P (a), R||Pa||Pat||P (b)", + "texts": [ + "4k) The last joints of the four limbs have superposed axes/ directions. The actuated prismatic joints of limbs G1, G2 and G3 have orthogonal directions 2. 3PPPR-1RPaPatP (Fig. 6.7b) P\\P ??P\\||R (Fig. 4.1a) P\\P ??P ??R (Fig. 4.1b) R||Pa||Pat||P (Fig. 5.4l) Idem no. 1 3. 2PPRRR-1PPRR-1RPaPaP (Fig. 6.8a) P ?P\\||R||R\\R (Fig. 4.2a) P\\P\\||R||R (Fig. 4.1d) R||Pa||Pa||P (Fig. 5.4k) Idem no. 1 4. 2PPRRR-1PPRR1RPaPatP (Fig. 6.8b) P\\P\\||R||R\\R (Fig. 4.2a) P\\P\\||R||R (Fig. 4.1d) R||Pa||Pat||P (Fig. 5.4l) Idem no. 1 5. 2PRPRR-1PRPR-1RPaPaP (Fig. 6.9a) P||R\\P\\||R\\R (Fig. 4.1b) P||R\\P\\||R (Fig. 4.1e) R||Pa||Pa||P (Fig. 5.4k) Idem no. 1 6. 2PRPRR-1PRPR1RPaPatP (Fig. 6.9b) P||R\\P\\||R\\R (Fig. 4.1b) P||R\\P\\||R (Fig. 4.1e) R||Pa||Pat||P (Fig. 5.4l) Idem no. 1 7. 2PRRRR-1PRRR-1RPaPaP (Fig. 6.10a) P||R||R||R\\R (Fig. 4.2d) P||R||R||R (Fig. 4.1g) R||Pa||Pa||P (Fig. 5.4k) Idem no. 1 8. 2PRRRR-1PRRR1RPaPatP (Fig. 6.10b) P||R||R||R\\R (Fig. 4.2d) P||R||R||R (Fig. 4.1g) R||Pa||Pat||P (Fig. 5.4l) Idem no. 1 9. 3PPPR-1RPPaPa (Fig. 6.11a) P\\P ??P\\||R (Fig. 4.1a) P\\P ??P ??R (Fig. 4.1b) R||P||Pa||Pa (Fig. 5.4m) The last revolute joints of limbs G1, G2 and G3 and the revolute joint between links 7D and 9 have superposed axes" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_35_0001419_978-94-007-7401-8-Figure3.17-1.png", + "original_path": "designv11-35/openalex_figure/designv11_35_0001419_978-94-007-7401-8-Figure3.17-1.png", + "caption": "Fig. 3.17 4PPRP-type overactuated PMs with coupled Sch\u00f6nflies motions defined by MF \u00bc SF \u00bc 4; RF\u00f0 \u00de \u00bc v1; v2; v3;xd\u00f0 \u00de, TF = 0, NF = 6, limb topology P\\P\\\\R||P (a) and P\\P\\kR\\P (b)", + "texts": [ + " 4PRPR (Fig. 3.13a) P||R\\P\\kR (Fig. 3.1r) Idem No. 1 20. 4RPPR (Fig. 3.13b) R||P\\P\\kR (Fig. 3.1s) Idem No. 2 21. 4RPPR (Fig. 3.14a) R\\P\\kP||R (Fig. 3.1t) Idem No. 2 22. 4RPRP (Fig. 3.14b) R\\P\\kR||P (Fig. 3.1u) Idem No. 2 23. 4PPPR (Fig. 3.15a) P\\P\\\\P\\kR (Fig. 3.1h0) Idem No. 9 24. 4PPPR (Fig. 3.15b) P\\P\\\\P\\\\R (Fig. 3.1w) Idem No. 9 25. 4PPPR (Fig. 3.16a) P\\P\\\\P||R (Fig. 3.1x) Idem No. 9 26. 4PPRP (Fig. 3.16b) P\\P||R\\P (Fig. 3.1y) The revolute joints of the four limbs have parallel axes 27. 4PPRP (Fig. 3.17a) P\\P\\\\R||P (Fig. 3.1z) Idem No. 26 28. 4PPRP (Fig. 3.17b) P\\P\\kR\\P (Fig. 3.1a0) Idem No. 26 29. 4PRPP (Fig. 3.18a) P\\R||P\\P (Fig. 3.1b0) Idem No. 26 30. 4PRPP (Fig. 3.18b) P\\R\\P\\kP (Fig. 3.1c0) Idem No. 26 31. 4PRPP (Fig. 3.19a) P||R\\P\\\\P (Fig. 3.1d0) Idem No. 26 32. 4RPPP (Fig. 3.19b) R\\P\\\\P\\\\P (Fig. 3.1e0) Idem No. 26 33. 4RPPP (Fig. 3.20a) R\\P\\kP\\\\P (Fig. 3.1f0) Idem No. 26 34. 4RPPP (Fig. 3.20b) R||P\\P\\\\P (Fig. 3.1g0) Idem No. 26 3.1 Topologies with Simple Limbs 243 Table 3.2 Limb topology and connecting conditions of the overactuated solutions with no idle mobilities presented in Figs" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_35_0000155_012015-Figure6-1.png", + "original_path": "designv11-35/openalex_figure/designv11_35_0000155_012015-Figure6-1.png", + "caption": "Fig. 6 rplidar A3", + "texts": [ + "4 TurtleBot's odometry The odometry is the data comes from motion sensors like wheel encoders, to calculate the estimated change in robots position and heading over time, realative to launching point in terms of an x and y position and an orientation around the z.The odometry data becomes less accurate as the robot moves. this inaccuracy can be due incorrect wheel diameters or due to uneven driving surfaces causing the wheel encoders to output inaccurate data. So odometry data should corrected by using extra source of information, in our case we will use rplidar, figure 6, for this task, which gives the absolute distance and orientation taken from robot pose, now we have all prerequisite to implement the equations (11 to15). Befor implement SALM algorithm , letus see how this approuch effects the robot pose, and as result effects the mapping process. Its sambley can be done depeding on ROS built system which enable the user to choose which sensor allowed to publish data, in this case we will accsess to odomety data, which represent the motion model, equation (11-12)" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_35_0000577_s40313-021-00754-5-Figure11-1.png", + "original_path": "designv11-35/openalex_figure/designv11_35_0000577_s40313-021-00754-5-Figure11-1.png", + "caption": "Fig. 11 Euler angle \u03a6, \u03b8, and \u03c8 to determine the orientation of a flying plane. a Yaw rotation on the Z-axis b Pitch rotation on Y-axis. c Roll rotation on X-axis", + "texts": [ + " Before the doublet input was given, the aircraft is conditioned in trim and level conditions (the aircraft is considered in a state of straight flight and the right-left wing of the aircraft is in the same state horizontally). The diagram of the flight data acquisition procedure is shown in Fig.\u00a08. The recorded flight data were then processed and converted to a variable that was used for empirical modeling or system identification as shown in Fig.\u00a09. In the fixed-wing UAV flight, there are many forces, velocities, moments, and orientations that work. All could be summarized and visualized in Fig.\u00a010. In a flight, the aircraft is not always horizontally straight when flying straight ahead (on the X-axis, see Fig.\u00a011) but has an angle-of-attack angle commonly written \u03b1 (alpha). The aircraft angle-of-attack is the deflection angle of the aircraft on the x-axis caused by the airplane\u2019s wing geometry. Similar to angle-of-attack, there is also a sideslip angle (Fig.\u00a011), which is usually written \u03b2 (beta). The difference is that the sideslip angle is the deflection on the y-axis due to the air/wind from the side that concerns the vertical stabilizer of the aircraft. Equations\u00a01 and 2 are for finding \u03b1 and \u03b2. The Euler angle was used for the aircraft orientation as shown in Fig.\u00a011, which has the following reference: a. Z-axis leads downward (same as the direction of the gravity vector). The angle rotation on this axis is called yaw (\u03c8). b. X-axis leads forward. The angle rotation on this axis is called roll (\u03c6). c. Y-axis leads rightward. The angle rotation on this axis is called pitch (\u03b8). (1) = tan\u22121 w u (2) = tan\u22121 \ufffd v\u221a u2 + w2 \ufffd 1 3 In an aircraft motion modeling, one of the parameters used is the angular velocity on the x, y, and z axes. The angular velocity is called the variables p, q, and r" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_35_0001419_978-94-007-7401-8-Figure2.113-1.png", + "original_path": "designv11-35/openalex_figure/designv11_35_0001419_978-94-007-7401-8-Figure2.113-1.png", + "caption": "Fig. 2.113 4PaRPaRR-type fully-parallel PM with coupled Sch\u00f6nflies motions defined by MF \u00bc SF \u00bc 4; RF\u00f0 \u00de \u00bc \u00f0v1; v2; v3;xd\u00de, TF = 0, NF = 26, limb topology Pa\\R\\Pa\\kR\\R", + "texts": [], + "surrounding_texts": [ + "2.2 Topologies with Complex Limbs 167", + "168 2 Fully-Parallel Topologies with Coupled Sch\u00f6nflies Motions", + "2.2 Topologies with Complex Limbs 169" + ] + }, + { + "image_filename": "designv11_35_0001553_978-3-319-14705-5_14-Figure2-1.png", + "original_path": "designv11-35/openalex_figure/designv11_35_0001553_978-3-319-14705-5_14-Figure2-1.png", + "caption": "Fig. 2 Support polygon and composed force", + "texts": [ + " This method has the advantage that it does not require a dynamic filter because errors between the model and a real robot are small thanks to using the multi-body robot model. This chapter describes a novel online walking pattern generation using Fast Fourier Transform (FFT). This chapter is organized as follows. Sections2 and3describes aFFT-based offline and online walking pattern generation, respectively, and Sect. 4 describes simulation results. In Sect. 5, experimental results are shown. Section6 provides conclusions and future work. Zero moment Point (ZMP) is defined as a point where the total forces and moments acting on a robot are zero [23] (see Fig. 2). If the ZMP exists inside the support polygon formed by the support points between the feet and the ground, a biped robot can walk stably without falling down. One complete walking cycle is divided into two phases: single support phase and double support phase. During the single support phase, one foot is on the ground and the other foot is in the air. The biped robot is in the double support phase as soon as the swing foot reaches the ground. The ZMP should be changed smoothly according to two support phases for dynamic stability" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_35_0000373_17415977.2021.1910683-Figure18-1.png", + "original_path": "designv11-35/openalex_figure/designv11_35_0000373_17415977.2021.1910683-Figure18-1.png", + "caption": "Figure 18. Information extracted from Carsim TM .", + "texts": [ + " Due to its widely application in automotive researches, including validation for dynamics mathematical models, a prior estimation of the cornering stiffness of the front and rear axle of a C-class vehicle included in this software database was performed. Figures 18(a) and 18(b) show the mentioned vehicle and the front wheel steering angle extracted from CarsimTM simulation, which is applied as an input to the SHM model of a two-axle vehicle during a double-lane change manoeuvre at 10 kmh\u22121. The lateral forces database for the 215/55R17 tyre used in this model is presented in Figure 18(c), showing that they varied according to the vertical load and the slip angle. In agreement with [25], it can be visualized in Figure 18(d) the phase lag between the slip angles and lateral forces in each tyre, attributed to the its relaxation length, that difficult the direct identification of the cornering stiffness. Simulation parameters used in the estimation process are presented in Table 6. Figure 19 shows the estimation process and the comparative yaw rate values, while Table 7 presents the estimated cornering stiffness values. The numerical solution is considered satisfactory to represent the yaw rate measurement given the root-mean-square error (RMSE) and multiple correlation coefficient (R2 ) values (greater than 0" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_35_0001419_978-94-007-7401-8-Figure5.88-1.png", + "original_path": "designv11-35/openalex_figure/designv11_35_0001419_978-94-007-7401-8-Figure5.88-1.png", + "caption": "Fig. 5.88 2PaPaRRR-1PaPaRR-1CPaPat-type fully-parallel PM with uncoupled Sch\u00f6nflies motions defined by MF = SF = 4, RF\u00f0 \u00de \u00bc \u00f0v1; v2; v3;xb\u00de, TF = 0, NF = 28, limb topology Pa\\Pa||R||R\\kR, Pa\\Pa||R||R and C||Pa||Pat", + "texts": [ + " 2PaPaRRR1PaPaRR1CPaPa (Fig. 5.85) Pa\\Pa||R||R\\\\R (Fig. 5.5a) The revolute joint between links 6D and 10 of G4 limb and the last revolute joints of limbs G1, G2 and G3 have superposed axes Pa\\Pa||R||R (Fig. 5.4i) C||Pa||Pa (Fig. 5.40) 24. 2PaPaRRR1PaPaRR1CPaPa (Fig. 5.86) Pa\\Pa||R||R\\kR (Fig. 5.5b) Idem No. 23 Pa\\Pa||R||R (Fig. 5.4i) C||Pa||Pa (Fig. 5.40) 25. 2PaPaRRR1PaPaRR1CPaPat (Fig. 5.87) Pa\\Pa||R||R\\\\R (Fig. 5.5a) Idem No. 23 Pa\\Pa||R||R (Fig. 5.4i) C||Pa||Pat (Fig. 5.4p) 26. 2PaPaRRR1PaPaRR1CPaPat (Fig. 5.88) Pa\\Pa||R||R\\kR (Fig. 5.5b) Idem No. 23 Pa\\Pa||R||R (Fig. 5.4i) C||Pa||Pat (Fig. 5.4p) 27. 3PaPaPaR1CPaPa (Fig. 5.89) Pa\\Pa||Pa\\kR (Fig. 5.6a) The revolute joint between links 6D and 11 of G4 limb and the last revolute joints of limbs G1, G2 and G3 have superposed axes Pa\\Pa||Pa||R (Fig. 5.6b) C||Pa||Pa (Fig. 5.4o) 28. 3PaPaPaR1CPaPa (Fig. 5.90) Pa\\Pa||Pa\\\\R (Fig. 5.6c) Idem No. 27 Pa\\Pa||Pa||R (Fig. 5.6b) C||Pa||Pa (Fig. 5.54o) 29. 3PaPaPaR1CPaPat (Fig. 5.91) Pa\\Pa||Pa\\kR (Fig. 5.6a) Idem No. 27 Pa\\Pa||Pa||R (Fig" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_35_0001419_978-94-007-7401-8-Figure5.35-1.png", + "original_path": "designv11-35/openalex_figure/designv11_35_0001419_978-94-007-7401-8-Figure5.35-1.png", + "caption": "Fig. 5.35 3PaPPaR-1RUPU-type fully-parallel PMs with uncoupled Sch\u00f6nflies motions defined by MF = SF = 4, RF\u00f0 \u00de \u00bc \u00f0v1; v2; v3;xb\u00de, TF = 0, NF = 22 limb topology R\\R\\R\\P\\kR\\R and Pa||P\\Pa\\kR, Pa||P\\Pa||R (a), Pa\\P\\\\Pa\\kR, Pa\\P\\\\Pa||R (b)", + "texts": [ + " 3PaPaPR1RPPP (Fig. 5.32b) Pa\\Pa\\\\P\\\\R (Fig. 5.4g) Idem No. 1 Pa\\Pa\\\\P\\kR (Fig. 5.4e) R\\P\\\\P\\\\P (Fig. 5.1a) 5. 2PaPaRRR1PaPaRR1RPPP (Fig. 5.33) Pa\\Pa||R||R\\\\R (Fig. 5.5a) The last joints of the four limbs have superposed axes/directions. The revolute joints between links 7 and 8 of limbs G1, G2 and G3 have orthogonal axes Pa\\Pa||R||R (Fig. 5.4i) R\\P\\\\P\\\\P (Fig. 5.1a) 6. 2PaPaRRR1PaPaRR1RPPP (Fig. 5.34) Pa\\Pa||R||R\\kR (Fig. 5.5b) Idem No. 5 Pa\\Pa||R||R (Fig. 5.4j) R\\P\\\\P\\\\P (Fig. 5.1a) 7. 3PaPPaR1RUPU (Fig. 5.35a) Pa||P\\Pa\\kR (Fig. 5.4a) The first revolute joint of G4 limb and the last revolute joints of limbs G1, G2 and G3 have parallel axes. The last joints of of limbs G1, G2 and G3 have superposed axes Pa||P\\Pa||R (Fig. 5.4d) R\\R\\R\\P\\kR\\R (Fig. 5.1b) (continued) 530 5 Topologies with Uncoupled Sch\u00f6nflies Motions Table 5.3 (continued) No. PM type Limb topology Connecting conditions 8. 3PaPPaR1RUPU (Fig. 5.35b) Pa\\P\\\\Pa\\kR (Fig. 5.4b) Idem No. 7 Pa\\P\\\\Pa||R (Fig. 5.4c) R\\R\\R\\P\\kR\\R (Fig. 5.1b) 9. 3PaPaPR1RUPU (Fig. 5.36a) Pa\\Pa\\kP\\\\R (Fig. 5.4f) Idem No. 7 Pa\\Pa\\kP\\kR (Fig. 5.4h) R\\R\\R\\P\\kR\\R (Fig. 5.1b) 10. 3PaPaPR1RUPU (Fig. 5.36b) Pa\\Pa\\\\P\\\\R (Fig. 5.4g) Idem No. 7 Pa\\Pa\\\\P\\kR (Fig. 5.4e) R\\R\\R\\P\\kR\\R (Fig. 5.1b) 11. 2PaPaRRR1PaPaRR1RUPU (Fig. 5.37) Pa\\Pa||R||R\\\\R (Fig. 5.5a) The last joints of of limbs G1, G2 and G3 have superposed axes. The revolute joints between links 7 and 8 of limbs G1, G2 and G3 have orthogonal axes Pa\\Pa||R||R (Fig" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_35_0000174_j.matpr.2021.01.640-Figure4-1.png", + "original_path": "designv11-35/openalex_figure/designv11_35_0000174_j.matpr.2021.01.640-Figure4-1.png", + "caption": "Fig. 4. Pro-E imported model.", + "texts": [ + " To determine the dynamic response of a structure under the action of any general time-dependent loads. In this type of analysis to determine the time-varying displacements, strains, stresses, and forces in a structure as it responds to any combination of static, transient, and harmonic loads. The time scale of the loading is such that the inertia or damping effects are considered to be important. The model is generated by Pro-E Software package is transferred to ANSYS software and its shown in Fig. 4 for dynamic analysis. Drive shaft have more benefit than two part drive shaft is made by Isotropic materials. It has higher explicit quality, more life, less mass, high basic speed and higher force conveying limit. Drive shaft with Geometric view of shown in Fig. 5. The pro-E model is imported to 3 dimensional view. Isometric view permits somewhat precise view of a three-dimensional object on a twodimensional or computer screen, and are a great way to visualize the shape of an object. Generated model is imported into ANSYS, is Drive Shaft" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_35_0000535_tmag.2021.3085750-Figure1-1.png", + "original_path": "designv11-35/openalex_figure/designv11_35_0000535_tmag.2021.3085750-Figure1-1.png", + "caption": "Fig. 1 Topology of (a) proposed HMC-VFMM, ( b) MS control scheme.", + "texts": [ + ", the required demagnetizing and remagnetizing current levels are basically similar, which will benefit the inverter rating reduction and online flux control. The balanced bidirectional-magnetization effect is analytically revealed by using a simplified Fourier-series based hysteresis model. In addition, the numerical hysteresis model is programmed with FE model so as to accurately investigate the magnetization characteristics of HMC-VFMM. Finally, some experiments on a prototype are carried out to verify the theoretical analyses. The 21-slot/4-pole HMC-VFMM is investigated in this paper, and the cross-section is illustrated in Fig. 1(a). The proposed HMC design is characterized by a dual-layer PM structure. For the parallel branch, the \u201cU\u201d-shaped hybrid PM arrangement is employed, HCF PMs are placed on two sides while the LCF is arranged in the middle. For the series branch, HCF PMs form a spoke-type flux concentration structure. The proposed HMC-VFMM can combine the advantages of wide flux regulation range in parallel type and excellent on-load demagnetization withstand capability in series type. The onload magnetization circuit is shown in Fig. 1(b). The magnetization state controller generates MS control command signals to the inverter in reference to the machine operation properties. The design parameters of the HMC-VFMM are shown in Table I. The variable-flux characteristics of the proposed HMCVFMM can be illustrated by the open-circuit field V Authorized licensed use limited to: San Francisco State Univ. Downloaded on July 03,2021 at 08:24:04 UTC from IEEE Xplore. Restrictions apply. 0018-9464 (c) 2021 IEEE. Personal use is permitted, but republication/redistribution requires IEEE permission", + " From the FE results, it can be deduced that the HMCVFMM design can effectively avoid the mismatch of working point movement between one-way and bidirectional MS manipulation. This magnetization characteristic of the HMCVFMM confirms that the alternate MS manipulation process has weak impact on the practical MS current pulses to be applied, and will further verify the validity of the balanced bidirectional-magnetization effect. As shown in Fig. 8(a), a HMC-VFMM prototype is manufactured and tested. Fig. 8(b) shows the experimental setup. The magnetization test is established based on the MS control circuit described in Fig. 1. The FE-predicted and measured line back-EMF waveforms and harmonic spectra under different MSs are shown in Fig. 9. It can be seen that the HMC-VFMM exhibits a good flux regulation capability. Besides, the FE-predicted and tested torque versus speed curves are compared as shown in Fig. 10. It can be deduced that the speed range of the HMC-VFMM can be effectively extended by online MS manipulation. The FE and measured open-circuit phase back-EMF magnitudes versus various peak current pulse curves subject to one-way continuous and bidirectional demagnetization processes are shown in Fig" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_35_0001419_978-94-007-7401-8-Figure4.3-1.png", + "original_path": "designv11-35/openalex_figure/designv11_35_0001419_978-94-007-7401-8-Figure4.3-1.png", + "caption": "Fig. 4.3 4PPPR-type fully-parallel PMs with decoupled Sch\u00f6nflies motions defined by MF = SF = 4, RF\u00f0 \u00de \u00bc v1; v2; v3;xa\u00f0 \u00de, TF = 0, NF = 6, limb topology P ? P ?? P ??R and P ?P ?? P ?||R (a), P ? P ??P||R (b)", + "texts": [ + " 370 4 Fully-Parallel Topologies with Decoupled Sch\u00f6nflies Motions 4.1 Topologies with Simple Limbs 371 372 4 Fully-Parallel Topologies with Decoupled Sch\u00f6nflies Motions 4.1 Topologies with Simple Limbs 373 374 4 Fully-Parallel Topologies with Decoupled Sch\u00f6nflies Motions 4.1 Topologies with Simple Limbs 375 Table 4.1 Limb topology and connecting conditions of the fully-parallel solutions with no idle mobilities presented in Figs. 4.3, 4.4, 4.5, 4.6, 4.7 No. PM type Limb topology Connecting conditions 1. 4PPPR (Fig. 4.3a) P ?P ??P ??R (Fig. 4.1b) P ?P ??P ?|R (Fig. 4.1a) The last revolute joints of the four limbs have parallel axes. The actuated prismatic joints of limbs G1, G2 and G3 hvave orthogonal directions. The actuated prismatic joints of limbs G3 and G4 have parallel directions 2. 4PPPR (Fig. 4.3b) P ?P ??P ??R (Fig. 2.1b) P ?P ??P||R (Fig. 4.1c) Idem No. 1 3. 1PPRR-3PPRRR (Fig. 4.4a) P ?P ?|R||R (Fig. 4.1d) P ?P ?|R||R ?R (Fig. 4.2a) Idem No. 1 4. 1PRPR-3PRPRR (Fig. 4.4b) P||R ?P ?|R (Fig. 4.1e) P||R ?P ?|R ?R (Fig. 4.2b) Idem No. 1 5. 1PRRP-3PRRPR (Fig. 4.5a) P||R||R ?P (Fig. 4.1f) P||R||R ?P ??R (Fig. 4.2c) Idem No. 1 6. 1PRRR-3PRRRR (Fig. 4.5b) P||R||R||R (Fig. 4.1g) P||R||R||R ?R (Fig. 4.2d) Idem No. 1 7. 1PPPR-3PPC (Fig. 4.6a) P ?P ??P ??R (Fig. 4.1b) P ?P ??C (Fig. 4.1h) The last joints of the four limbs have parallel axes", + " The actuated prismatic joints of limbs G1, G2 and G3 hvave orthogonal directions. The actuated prismatic joints of limbs G3 and G4 have parallel directions 8. 1CPR-3CPRR (Fig. 4.6b) C ?P ?|R (Fig. 4.1i) C ?P ?|R ?R (Fig. 4.2e) Idem No. 1 9. 1CRP-3CRPR (Fig. 4.7a) C||R ?P (Fig. 4.1j) C||R ?P ??R (Fig. 4.2f) Idem No. 1 10. 1CRR-3CRRR (Fig. 4.7b) C||R||R (Fig. 4.1k) C||R||R ?R (Fig. 4.2g) Idem No. 1 376 4 Fully-Parallel Topologies with Decoupled Sch\u00f6nflies Motions Table 4.2 Structural parametersa of parallel mechanisms in Figs. 4.3, 4.4, 4.5 No. Structural parameter Solution Figure 4.3 Figures 4.4 and 4.5 1. m 14 17 2. p1 4 4 3. pi (i = 2, 3, 4) 4 5 4. p 16 19 5. q 3 3 6. k1 4 4 7. k2 0 0 8. k 4 4 9. (RG1) (v1; v2; v3;xa) (v1; v2; v3;xa) 10. (RG2) (v1; v2; v3;xa) (v1; v2; v3;xa;xb) 11. (RG3) (v1; v2; v3;xa) (v1; v2; v3;xa;xd) 12. (RG4) (v1; v2; v3;xa) (v1; v2; v3;xa;xd) 13. SG1 4 4 14. SGi (i = 2, 3, 4) 4 5 15. rGi (i = 1,\u2026,4) 0 0 16. MG1 4 4 17. MGi (i = 2, 3, 4) 4 5 18. (RF) (v1; v2; v3;xa) (v1; v2; v3;xa) 19. SF 4 4 20. rl 0 0 21. rF 12 15 22. MF 4 4 23. NF 6 3 24. TF 0 0 25" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_35_0000417_012020-Figure5-1.png", + "original_path": "designv11-35/openalex_figure/designv11_35_0000417_012020-Figure5-1.png", + "caption": "Figure 5. (a) FEM of NPT, (b) components of NPT, and (c) comparison of embed rebar element and actual belts\u2019 position.", + "texts": [ + " It was found that waterjet specimens show nearly identical stress-strain profile for radial and width cutting direction. However, the ultimate stress and breaking stress were found to be significantly different. The11th International Conference on Mechanical Engineering (TSME-ICOME 2020) IOP Conf. Series: Materials Science and Engineering 1137 (2021) 012020 IOP Publishing doi:10.1088/1757-899X/1137/1/012020 The FE model of NPT was created and the conditions were applied using FE software, MSC. Patran. The FE mesh of NPT along with its dimension are shown in Figure 5(a) The tread and shear band were modelled using continuum element, while the spoke was modelled using shell element. The linear elastic or Hookean material model was used to model the mechanical behaviour of NPT\u2019s components. The details of elements used in modelling the NPT are summarized as shown in table 3. Each component is assembled together using glued contact condition just like the actual NPT assemble process as shown in Figure 5(b). The belt layers were modelled using reinforce bar or rebar element. They were embed into shear band element using tying equation. The comparison of embed rebar element and actual belt\u2019s position is shown as schematic diagram in Figure 5(c). In addition, each rebar element composed of mathematically sublayers of belt. The outer layers, middle layers, and inner layers composed of 3, 1, and 2 sublayers, respectively. The bead wire diameter was estimated by measuring to be 1 mm, while the number of wires per unit length was estimated to 0.3582 mm-1. The simulation of vertical stiffness testing was performed using FE software, MSC. Marc. The FE mesh of NPT is combined with rigid plate, which represented tire testing machine\u2019s moving plate as shown in Figure 6(a)" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_35_0000216_s00202-021-01264-y-Figure7-1.png", + "original_path": "designv11-35/openalex_figure/designv11_35_0000216_s00202-021-01264-y-Figure7-1.png", + "caption": "Fig. 7 Membership functions for a m and I b", + "texts": [ + " The input membership functions are triangular-shaped symmetrically distributed around zero while singleton membership functions are preferred for the output variable owing to their low computational burden during defuzzification. The inputs are normalized to [\u2212 1, 1] and the output is in the range of [ 180 , 30 180 ] rad, from which it is elucidated that can be varied within [1\u00b0, 30\u00b0]. Mamdani\u2019s min\u2013max fuzzy inference method is employed and for defuzzification we adopt the weighted average technique. The membership functions are shown in Fig.\u00a07, where the linguistic labels from left to right are NB = Negative Big, N = Negative, Z = Zero, P = Positive, PB = Positive Big, VS = Very Small, S = Small, M = Medium, B = Big, and VB = Very Big. As five membership functions are chosen for two inputs, 5 \u00d7 5 = 25 fuzzy rules are to be determined. This means a 25D combinatorial optimization problem causing an exhausting work and may need excessive time to perform optimization by GA. To simplify optimization, the rule table is considered as a grid with 25 cells as shown in Fig" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_35_0001572_978-3-319-05371-4_7-Figure7.4-1.png", + "original_path": "designv11-35/openalex_figure/designv11_35_0001572_978-3-319-05371-4_7-Figure7.4-1.png", + "caption": "Fig. 7.4 The friction circle concept (from [19])", + "texts": [ + "8) where h is the distance of the vehicle center of mass from the ground (see Fig. 7.3), and where \u03bc j = D sin(C arctan(Bs j )), \u03bci j = \u2212(si j/s j )\u03bc j , i = x, y; j = f, r, (7.9) for some constants C, B and D. Expression (7.9) is a simplified version of the well-known Pacejka \u201cMagic Formula\u201d (MF) [24] for the tire friction modeling, and combines the longitudinal and lateral motion, thus intrinsically incorporating the non-linear effect of the lateral/longitudinal coupling also known as the \u201cfriction circle\u201d (see Fig. 7.4), according to which, the constraint F2 x, j + F2 y, j \u2264 F2 max, j = (\u03bc j Fz, j ) 2 ( j = f, r) couples the allowable values of longitudinal and lateral tire friction forces. Incorporating the friction circle constraint is necessary for the correct modeling of the dynamics occurring during the aggressive maneuvers we consider in this work. In Eq. (7.9) si j denote the tire longitudinal and lateral slip ratios, given by sx j = Vx j \u2212 \u03c9 j R \u03c9 j R = Vx j \u03c9 j R \u2212 1, sy j = (1 + sx j ) Vyj Vx j , j = f, r, (7" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_35_0000395_978-3-030-67750-3_16-Figure17.2-1.png", + "original_path": "designv11-35/openalex_figure/designv11_35_0000395_978-3-030-67750-3_16-Figure17.2-1.png", + "caption": "Fig. 17.2 Vehicle dynamic of a car", + "texts": [ + " The design requirements and the expected parameters to be determined are needed to allow the DC drive EV using a series motor and FQDC that have been designing to pass the New European Driving Cycle (NEDC) test. The propulsion unit of the vehicle needs a tractive force FTR to propel the vehicle. The tractive force must fulfill the requirements of vehicle dynamics to overcome forces such as the rolling resistance, gravitational, and aerodynamically which are summed together as the road load force FRL as shown in Fig. 17.2 (Arof et al. 2020a, b, c, d). Thegravitational force Fg dependson the slopeof the roadway, as shown inEq. (17.1). Fg = mg sin \u03b1 (17.1) \u03b1 is the grade angle, m is the total mass of the vehicle, g is the gravity constant. The hysteresis of the tire material causes it at the contact surfaces with the roadway. The centroid of the vertical forces on the wheel moves forward when the tire rolls. Therefore, from beneath the axle toward the direction of motion by the vehicle, as shown in Fig. 17.3. Tractive Force The tractive forcewas used to overcome the Froll force alongwith the gravity force and the aerodynamic drag force" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_35_0001419_978-94-007-7401-8-Figure5.72-1.png", + "original_path": "designv11-35/openalex_figure/designv11_35_0001419_978-94-007-7401-8-Figure5.72-1.png", + "caption": "Fig. 5.72 3PaPPR-1CPaPat-type fully-parallel PMs with uncoupled Sch\u00f6nflies motions defined by MF = SF = 4, RF\u00f0 \u00de \u00bc \u00f0v1; v2; v3;xb\u00de, TF = 0, NF = 21, limb topology C||Pa||Pat and Pa||P\\P\\kR, Pa||P\\P\\\\R (a), Pa\\P\\kP\\kR, Pa\\P\\kP\\\\R (b)", + "texts": [ + "6b) R||P||Pa||Pat (Fig. 5.4n) 2. 3PaPaPaR1RPPaPat (Fig. 5.70) Pa\\Pa||Pa\\\\R (Fig. 5.6c) Idem No. 1 Pa\\Pa||Pa||R (Fig. 5.6b) R||P||Pa||Pat (Fig. 5.54n) 3. 3PaPPR1CPaPa (Fig. 5.71a) Pa||P\\P\\kR (Fig. 5.2a) The revolute joint between links 6D and 8 of G4 limb and the last revolute joints of limbs G1, G2 and G3 have superposed axes Pa||P\\P\\\\R (Fig. 5.2d) C||Pa||Pa (Fig. 5.4o) 4. 3PaPPR1CPaPa (Fig. 5.71b) Pa\\P\\kP\\kR (Fig. 5.2b) Idem No. 3 Pa\\P\\kP\\\\R (Fig. 5.2c) C||Pa||Pa (Fig. 5.40) 5. 3PaPPR1CPaPat (Fig. 5.72a) Pa||P\\P\\kR (Fig. 5.2a) Idem No. 3 Pa||P\\P\\\\R (Fig. 5.2d) C||Pa||Pat (Fig. 5.4p) 6. 3PaPPR1CPaPat (Fig. 5.72b) Pa\\P\\kP\\kR (Fig. 5.2b) Idem No. 3 Pa\\P\\kP\\\\R (Fig. 5.2c) C||Pa||Pat (Fig. 5.4p) (continued) 534 5 Topologies with Uncoupled Sch\u00f6nflies Motions Table 5.5 (continued) No. PM type Limb topology Connecting conditions 7. 2PaPRRR1PaPRR1CPaPa (Fig. 5.73a) Pa\\P\\\\R||R\\\\R (Fig. 5.3a) Idem No. 3 Pa\\P\\\\R||R (Fig. 5.2e) C||Pa||Pa (Fig. 5.40) 8. 2PaPRRR1PaPRR1CPaPa (Fig. 5.73b) Pa||P\\R||R\\\\R (Fig. 5.3b) Idem No. 3 Pa||P\\R||R (Fig. 5.2f) C||Pa||Pa (Fig. 5.40) 9. 2PaPRRR1PaPRR1CPaPat (Fig. 5.74a) Pa\\P\\\\R||R\\\\R (Fig" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_35_0001584_978-1-4614-8544-5_15-Figure15.22-1.png", + "original_path": "designv11-35/openalex_figure/designv11_35_0001584_978-1-4614-8544-5_15-Figure15.22-1.png", + "caption": "FIGURE 15.22. Behavior of 4 3 as a function of , .", + "texts": [], + "surrounding_texts": [ + "The most employed and useful model of a vehicle suspension system is a quarter car model, shown in Figure 15.1. We introduce, examine, and optimize the quarter car model in this chapter. 15.1 Mathematical Model We may represent the vertical vibration of a vehicle using a quarter car model made of two solid masses and denoted as sprung and unsprung masses, respectively. The sprung mass represents 1 4 of the body of the vehicle, and the unsprung mass represents one wheel of the vehicle. A spring of sti ness , and a shock absorber with viscous damping coe cient , support the sprung mass and are called the main suspension. The unsprung mass is in direct contact with the ground through a spring , representing the tire sti ness. The governing di erential equations of motion for the quarter car model shown in Figure 15.1, are: \u00a8 + ( ) + ( ) = 0 (15.1) \u00a8 + ( ) + ( + ) = (15.2) R.N. Jazar, Vehicle Dynamics: Theory and Application, DOI 10.1007/978-1-4614-8544-5_15 \u00a9 Springer Science+Business Media New York 2014 985 Proof. The kinetic energy, potential energy, and dissipation function of the system are: = 1 2 2 + 1 2 2 (15.3) = 1 2 ( ) 2 + 1 2 ( ) 2 (15.4) = 1 2 ( ) 2 (15.5) Employing the Lagrange method,\u03bc \u00b6 + + = 0 (15.6)\u03bc \u00b6 + + = 0 (15.7) we nd the equations of motion \u00a8 = ( ) ( ) (15.8) \u00a8 = ( ) + ( ) ( ) (15.9) which can be expressed in a matrix form [ ] x\u0308+ [ ]x+ [ ]x = F (15.10) 0 0 \u00b8 \u00a8 \u00a8 \u00b8 + \u00b8 \u00b8 + + \u00b8 \u00b8 = 0 \u00b8 (15.11) Example 570 Tire damping. We may add a damper in parallel to , as shown in Figure 15.1, to model any damping in tires. However, the value of for tires, compared to , are very small and hence, we may ignore to simplify the model. Having the damper in parallel to makes the equation of motion the same as Equations (12.44) and (12.45) with a matrix form as Equation (12.47). Example 571 Mathematical model\u2019s limitations. The quarter car model contains no representation of the geometric e ects of the full car and o ers no possibility of studying longitudinal and lateral interconnections. However, it contains the most basic features of the real problem and includes a proper representation of the problem of controlling wheel and wheel-body load variations. In the quarter car model, we assume that the tire is always in contact with the ground, which is true at low frequency but might not be true at high frequency. A better model must be able to include the possibility of separation between the tire and ground. Optimal design of two-DOF vibration systems, including a quarter car model, is the subject of numerous investigations since the invention of the vibration absorber theory by Frahm in 1909. It seems that the rst analytical investigation on the damping properties of two-DOF systems is due to Den Hartog (1901 1989). 15.2 Frequency Response To nd the frequency response, we consider a harmonic excitation, = cos (15.12) and look for a harmonic solution in the form = 1 sin + 1 cos = sin ( ) (15.13) = 2 sin + 2 cos = sin ( ) (15.14) = = 3 sin + 3 cos = sin ( ) (15.15) where , , and are complex amplitudes. By introducing the following dimensionless characters: = (15.16) = r (15.17) = r (15.18) = (15.19) = (15.20) = 2 (15.21) we search for the absolute and relative frequency responses: = \u00af\u0304\u0304\u0304 \u00af\u0304\u0304\u0304 (15.22) = \u00af\u0304\u0304\u0304 \u00af\u0304\u0304\u0304 (15.23) = \u00af\u0304\u0304\u0304 \u00af\u0304\u0304\u0304 (15.24) and obtain the following functions: 2 = 4 2 2 + 1 2 1 + 2 2 (15.25) 2 = 4 2 2 + 1 + 2 \u00a1 2 2 \u00a2 2 1 + 2 2 (15.26) 2 = 4 2 1 + 2 2 (15.27) 1 = \u00a3 2 \u00a1 2 2 1 \u00a2 + \u00a1 1 (1 + ) 2 2 \u00a2\u00a4 (15.28) 2 = 2 \u00a1 1 (1 + ) 2 2 \u00a2 (15.29) The absolute acceleration of sprung mass and unsprung mass may be de ned by the following equations: = \u00af\u0304\u0304\u0304 \u00af \u00a8 2 \u00af\u0304\u0304\u0304 \u00af = 2 2 (15.30) = \u00af\u0304\u0304\u0304 \u00af \u00a8 2 \u00af\u0304\u0304\u0304 \u00af = 2 2 (15.31) Proof. To nd the frequency responses, let us apply a harmonic excitation = cos (15.32) and assume that the solutions are harmonic functions with unknown coefcients. = 1 sin + 1 cos (15.33) = 2 sin + 2 cos (15.34) Substituting the solutions in the equations of motion (15.1)-(15.2) and collecting the coe cients of sin and cos in both equations provides the following set of algebraic equations for 1, 1, 2, 2: [ ] 1 2 1 2 = 0 0 0 (15.35) where [ ] is the coe cient matrix. [ ] = 2 2 + 2 + 2 (15.36) The unknowns may be found by matrix inversion 1 2 1 2 = [ ] 1 0 0 0 (15.37) and therefore, the amplitudes and can be found. 2 = 2 1 + 2 1 = \u00a1 2 2 + 2 \u00a2 2 3 + 2 4 2 (15.38) 2 = 2 2 + 2 2 = \u00a1 4 2 + 2 2 2 2 + 2 \u00a2 2 3 + 2 4 2 (15.39) 3 = \u00a1 2 ( + + ) 4 \u00a2 (15.40) 4 = \u00a1 3 ( + ) \u00a2 (15.41) Having and helps us to calculate and its amplitude . = = ( 1 2) sin + ( 1 2) cos = 3 sin + 3 cos = sin ( ) (15.42) 2 = 2 3 + 2 3 = 4 2 2 3 + 2 4 2 (15.43) Taking derivative from and provides the acceleration frequency responses and for the unsprung and sprung masses. Equations (15.30)- (15.31) express and . Using the de nitions (15.16)-(15.21), we may transform Equations (15.38), (15.39), (15.43) to (15.25), (15.26), (15.27). Figures 15.2, 15.3, 15.4, are samples of the frequency responses , , and for = 3, and = 0 2. Example 572 Average value of parameters for street cars. Equations (15.25)-(15.27) indicate that the frequency responses , , and are functions of four parameters: mass ratio , damping ratio , natural frequency ratio , and excitation frequency ratio . The average, minimum, and maximum of practical values of the parameters are indicated in Table 14 1. For a quarter car model, it is known that , and therefore, 1. Typical mass ratio, , for vehicles lies in the range 3 to 8, with small cars closer to 8 and large cars near 3. The excitation frequency is equal to , when = 1 , and equal to , when = 1. For a real model, the order of magnitude of the sti ness is , so , and 1. Therefore, 1 at = . So, we expect to have two resonant frequencies greater than = 1. Example 573 F Three-dimensional visualization for frequency responses. To get a sense about the behavior of di erent frequency responses of a quarter car model, Figures 15.5 to 15.8 are plotted for = 375kg = 35000N m = 75kg = 193000N m (15.44) 15.3 F Natural and Invariant Frequencies The quarter car system is a two-DOF system and therefore it has two natural frequencies 1 , 2 : 1 = s 1 2 2 \u03bc 1 + (1 + ) 2 q (1 + (1 + ) 2) 2 4 2 \u00b6 (15.45) 2 = s 1 2 2 \u03bc 1 + (1 + ) 2 + q (1 + (1 + ) 2) 2 4 2 \u00b6 (15.46) The family of response curves for the displacement frequency response of the sprung mass, , are obtained by keeping and constant, and varying . This family has several points in common, which are at frequencies 1, 2, 3, 4, and 1, 2, 3, 4, 1 = 0 1 = 1 3 = 1 3 = 1 2 2 = 1 1 (1 + ) 2 2 2 4 4 = 1 1 (1 + ) 2 2 2 (15.47) 2 = s 1 2 2 \u03bc 1 + 2 (1 + ) 2 q (1 + 2 (1 + ) 2) 2 8 2 \u00b6 (15.48) 4 = s 1 2 2 \u03bc 1 + 2 (1 + ) 2 + q (1 + 2 (1 + ) 2) 2 8 2 \u00b6 (15.49) where 1 (= 0) 2 1 1 + 3 \u03bc = 1 \u00b6 4 (15.50) The corresponding transmissivities at 2 and 4 are 2 = 1 1 (1 + ) 2 2 2 (15.51) 4 = 1 1 (1 + ) 2 2 2 (15.52) The frequencies 1, 2, 3, and 4 are called invariant frequencies, and their corresponding amplitudes are called invariant amplitudes because they are not dependent on . However, they are dependent on the values of and . The order of magnitude of the natural and invariant frequencies are: 1 (= 0) 1 2 1 1 + 3 \u03bc = 1 2 \u00b6 2 4 (15.53) The curves for have no other common points except 1, 2, 3, 4. The order of frequencies along with the order of corresponding amplitudes can be used to predict the shape of the frequency response curves of the sprung mass . Figure 15.9 shows schematically the shape of the amplitude versus excitation frequency ratio . Proof. The natural and resonant frequencies of a system are at positions where the amplitude goes to in nity when damping is zero. Hence, the natural frequencies would be the roots of the denominator of the function. \u00a1 2 \u00a2 = 2 \u00a1 2 2 1 \u00a2 + \u00a1 1 (1 + ) 2 2 \u00a2 = 2 2 \u00a1 1 + (1 + ) 2 \u00a2 + 1 = 0 (15.54) The solution of this equation are the natural frequencies given in Equations (15.45) and (15.46). The invariant frequencies are independent of , so they can be found by intersecting the curves for = 0 and = . lim 0 2 = \u00b1 1 ( 2 ( 2 2 1) 2 2 ( + 1) + 1) 2 (15.55) lim 2 = \u00b1 1 ( 2 2 ( + 1) 1) 2 (15.56) Therefore, the invariant frequencies, , can be determined by solving the following equation: 2 \u00a1 2 2 1 \u00a2 + \u00a1 1 (1 + ) 2 2 \u00a2 = \u00b1 \u00a11 (1 + ) 2 2 \u00a2 (15.57) Using the (+) sign, we nd 1 and 3 with their corresponding transmissivities 1, and 3, 1 = 0 1 = 1 (15.58) 3 = 1 3 = 1 (15.59) and, with the ( ) sign, we nd the following equation for 2, and 4: 2 4 \u00a1 1 + 2 (1 + ) 2 \u00a2 2 + 2 = 0 (15.60) Equation (15.60) has two real positive roots, 2 and 4, 2 = s 1 2 2 \u03bc 1 + 2 (1 + ) 2 q (1 + 2 (1 + ) 2) 2 8 2 \u00b6 (15.61) 4 = s 1 2 2 \u03bc 1 + 2 (1 + ) 2 + q (1 + 2 (1 + ) 2) 2 8 2 \u00b6 (15.62) with the following relative order of magnitude: 1 (= 0) 2 1 1 + 3 \u03bc = 1 \u00b6 4 (15.63) The corresponding amplitudes at 2, and 4 can be found by substituting Equations (15.61) and (15.62) in (15.25). 2 = 1 1 (1 + ) 2 2 2 (15.64) 4 = 1 1 (1 + ) 2 4 2 (15.65) It can be checked that (1 + ) 2 2 4 1 1 (15.66) and hence, | 4| 1 (= 3) 1 | 2| (15.67) and therefore, 2 1 (15.68) 4 1 (15.69) Using Equation (15.54), we can evaluate \u00a1 2 2 \u00a2 , \u00a1 2 4 \u00a2 , and \u00a1 2 3 \u00a2 as\u00a1 2 2 \u00a2 = (1 + ) 2 2 2 1 0 (15.70)\u00a1 2 4 \u00a2 = (1 + ) 2 2 4 1 0 (15.71)\u00a1 2 3 \u00a2 = \u03bc 1 2 \u00b6 (15.72) therefore, the two positive roots of Equation (15.54), 1 and 2 \u00a1 2 2 \u00a2 , have the order of magnitudes as: 1 (= 0) 1 2 1 1 + 3 \u03bc = 1 2 \u00b6 2 4 (15.73) Example 574 F Nodes of the absolute frequency response . There are four nodes in the absolute displacement frequency response of a quarter car. The rst node is at a trivial point ( 1 = 0 1 = 1), which shows that = when the excitation frequency is zero. The fourth node is at ( 4 4 1). There are also two middle nodes at ( 2 2 1) and \u00a1 3 = 1 3 = 1 \u00a2 . Because 1 1 and 4 1, the middle nodes are important in optimization. To have a better view at the middle nodes, Figure 15.10 illustrates a magni cation of the sprung mass displacement frequency response, = \u00af\u0304 \u00af\u0304 around the middle nodes. Example 575 F There is no Frahm optimal quarter car. Reduction in absolute amplitude is the rst attempt for optimization. If the amplitude frequency response = ( ) contains xed points with respect to some parameters, then using the Frahm method, the optimization process is carried out in two steps: 1 We select the parameters that control the position of the invariant points to equalize the corresponding height at the invariant frequencies, and minimize the height of the xed points as much as possible. 2 We nd the remaining parameters such that the maximum amplitude coincides precisely at the invariant points. For a real problem, the values of mass ratio , and wheel frequency are xed and we are trying to nd the optimum values of and . The parameters and include the unknown sti ness of the main spring and the unknown damping of the main shock absorber, respectively. The amplitude at invariant frequencies , show that the rst invariant point ( 1 = 0 1 = 1) is always xed, and the fourth one ( 4 4 1) happens after the natural frequencies. Therefore, the second and third nodes are the suitable nodes for applying the above optimization steps. However, 2 1 3 1 (15.74) and hence, we cannot apply the above optimization method. It is because 2 and 3 can never be equated by varying . Ever so, we can still nd the optimum value of by evaluating based on other constraints. Example 576 Natural frequency variation. The natural frequencies 1 and 2 , as given in Equations (15.45) and (15.46), are functions of and . Figures 15.11 and 15.12 illustrate the e ect of these two parameters on the variation of the natural frequencies. The rst natural frequency 1 1 decreases by increasing the mass ratio . 1 is close to the natural frequency of a 1 8 car model and indicates the principal natural frequency of a car. Hence, it is called the body bounce natural frequency. The second natural frequency, 2 , approaches in nity when decreases. However, 2 10Hz for street cars with acceptable ride comfort. 1 relates to the unsprung mass, and is called the wheel hop natural frequency. Figure 15.13 that plots the natural frequency ratio 1 2 shows their relative behavior. Example 577 Invariant frequencies variation. The invariant frequencies 2, 3, and 4, as given in Equation (15.47), are functions of and . Figures 15.14 to 15.18 illustrate the e ect of these two parameters on the invariant frequencies. The second invariant frequency 2, as shown in Figure 15.14, is always less than 2 because lim 0 2 = 2 (15.75) So, whatever the value of the mass ratio is, 2 cannot be greater than 2. Such a behavior does not let us control the position of second node freely. The third invariant frequency 3 as shown in Figure 15.15 is not a function of the mass ratio and may have any value depending on . The fourth invariant frequency 4 is shown in Figure 15.15. 4 increases when decreases. However, 4 settles when & 0 6. lim 0 2 = (15.76) To have a better picture about the behavior of invariant frequencies, Figures 15.17 and 15.18 depict the relative frequency ratio 4 3 and 3 2. Example 578 Frequency response at invariant frequencies. The frequency response is a function of , , and . Damping always diminishes the amplitude of vibration, so at rst we set = 0 and plot the behavior of as a function of , . Figure 15.19 illustrates the behavior of at the second invariant frequency 2. Because lim 0 2 = 1 (15.77) 2 starts at one, regardless of the value of . The value of 2 is always greater than one. Figure 15.20 shows that 3 is not a function of and is a decreasing function of . Figure 15.21 shows that 4 1 regardless of the value of and . The relative behavior of 2, 3, and 4 is shown in Figures 15.22 and 15.23. Example 579 Natural frequencies and vibration isolation of a quarter car. For a modern typical passenger car, the values of natural frequencies are around 1Hz and 10Hz respectively. The former is due to the bounce of sprung mass and the latter belongs to the unsprung mass. At average speed, bumps with wavelengths much greater than the wheelbase of the vehicle, will excite bounce motion of the body. At at higher speed, wavelength of the bumps become shorter than a wheelbase length and cause heavy vibrations of the unsprung. Therefore, when the wheels hit a single bump on the road, the impulse will set the wheels into oscillation at the natural frequency of the unsprung mass around 10Hz. In turn, for the sprung mass, the excitation will be the frequency of vibration of the unsprung around 10Hz. Because the natural frequency of the sprung is approximately 1Hz, the excellent isolation for sprung mass occurs and the frequency range around 10Hz has no essential in uence on the sprung discomfort. When the wheel runs over a rough undulating surface, the excitation will consists of a wide range of frequencies. Again, high excitation frequency at 5Hz to 20Hz means high frequency input to the sprung mass, which can e ectively be isolated. Low frequency excitation, however, will cause resonance in the sprung mass. 15.4 F RMS Optimization Figure 15.24 is a design chart for optimal suspension parameters of a base excited two-DOF system such as a quarter car model. The horizontal axis is the root mean square of relative displacement, = ( ), and the vertical axis is the root mean square of absolute acceleration, = ( ). There are two sets of curves that make a mesh. The rst set, which is almost parallel at the right end, are constant damping ratio , and the second set is constant natural frequency ratio . There is a curve, called the optimal design curve, which indicates the optimal main suspension parameters: The optimal design curve is the result of the RMS optimization strategy \u00a8 (15.78) which states that the minimum absolute acceleration with respect to the relative displacement, if there is any, makes the suspension of a quarter car optimal. Mathematically, it is equivalent to the following minimization problem: = 0 (15.79) 2 2 0 (15.80) To use the design curve and determine optimal sti ness and damping for the main suspension of the system, we start from an estimate value for on the horizontal axis and draw a vertical line to hit the optimal curve. The intersection point indicates the optimal and for the . Figure 15.25 illustrates a sample application for = 0 75, which indicates 0 3 and 0 35 for optimal suspension. Having and , determines the optimal value of and . = 2 (15.81) = 2 p (15.82) Proof. The RMS of a continues function ( ) is de ned by ( ) = s 1 2 1 Z 2 1 2 ( ) (15.83) where 2 1 is called the working frequency range. Let us consider a working range for the excitation frequency 0 \u00a1 = 2 \u00a2 20Hz to include almost all ground vehicles, especially road vehicles, and show the RMS of and by = ( ) (15.84) = ( ) (15.85) In applied vehicle dynamics, we usually measure frequencies in [ Hz], instead of [ rad s], we perform design calculations based on cyclic frequencies and in [ Hz], and we do analytic calculation based on angular frequencies and in [ rad s]. To calculate and over the working frequency range = s 1 40 Z 40 0 2 (15.86) = s 1 40 Z 40 0 2 = 2 s 1 40 Z 40 0 2 2 (15.87) we rst nd integrals of 2 and 2.Z 2 = 1 2 6 \u03bc 1 1 + 1 5 \u00b6 ln \u03bc 1 + 1 \u00b6 + 1 2 7 \u03bc 1 2 + 2 5 \u00b6 ln \u03bc 2 + 2 \u00b6 + 1 2 8 \u03bc 1 3 + 3 5 \u00b6 ln \u03bc 3 + 3 \u00b6 + 1 2 9 \u03bc 1 4 + 4 5 \u00b6 ln \u03bc 4 + 4 \u00b6 (15.88) Z 2 = 3 1 2 6 ln \u03bc 1 + 1 \u00b6 ++ 3 2 2 7 ln \u03bc 2 + 2 \u00b6 + 3 3 2 8 ln \u03bc 3 + 3 \u00b6 + 3 4 2 9 ln \u03bc 4 + 4 \u00b6 (15.89) The parameters 1 through 9 are: 1 = 1 2 19 + 23 19 1 4 15 14 (15.90) 2 = 1 2 19 23 19 1 4 15 14 (15.91) 3 = 1 2 19 + 24 19 1 4 15 14 (15.92) 4 = 1 2 19 24 19 1 4 15 14 (15.93) 5 = 4 2 (15.94) 6 = \u00a1 2 1 2 2 \u00a2 \u00a1 2 1 2 3 \u00a2 \u00a1 2 1 2 4 \u00a2 (15.95) 7 = \u00a1 2 2 2 3 \u00a2 \u00a1 2 2 2 3 \u00a2 \u00a1 2 2 2 1 \u00a2 (15.96) 8 = \u00a1 2 3 2 4 \u00a2 \u00a1 2 3 2 1 \u00a2 \u00a1 2 3 2 2 \u00a2 (15.97) 9 = \u00a1 2 4 2 1 \u00a2 \u00a1 2 4 2 2 \u00a2 \u00a1 2 4 2 3 \u00a2 (15.98) 10 = 1 6 3 p 20 + 8 13 + 2 3 2 11 3 20 + 1 3 11 (15.99) 11 = 8 16 14 3 3 15 8 3 14 (15.100) 12 = 4 16 14 15 3 15 8 2 14 17 8 3 14 (15.101) 13 = 64 2 14 17 15 + 256 3 14 18 + 16 14 2 15 16 3 4 15 256 4 14 (15.102) 14 = 4 (15.103) 15 = 2 4 (1 + ) 2 2 + 4 (1 + ) 2 4 2 (15.104) 16 = 8 2 2 (1 + ) + (1 + ) 2 4 2 2 (2 + ) + 1 (15.105) 17 = 4 2 2 2 (1 + ) 2 (15.106) 18 = 1 (15.107) 19 = 10 11 (15.108) 20 = 21 + 12 p 22 (15.109) 21 = 288 11 13 + 108 2 12 + 8 3 11 (15.110) 22 = 768 3 13 + 384 2 11 2 13 48 13 4 11 432 11 2 12 13 + 81 4 12 + 12 3 11 2 12 (15.111) 23 = 19 ( 11 10) 2 12 3 2 19 (15.112) 24 = 19 ( 11 + 10) + 2 12 3 2 19 (15.113) Now the required RMS, , and , over the frequency range 0 20Hz, can be calculated analytically from Equations (15.86) and (15.87). Equations (15.86) and (15.87) show that both and are functions of only three variables: , , and . = ( ) (15.114) = ( ) (15.115) In applied vehicle dynamics, is usually a xed parameter, so, any pair of design parameters ( ) determines and uniquely. Let us set = 3 (15.116) Using Equations (15.86) and (15.87), we may draw Figure 15.26 to illustrate how behaves with respect to when and vary. Keeping constant and varying , it is possible to minimize with respect to . The minimum points make the optimal curve and determine the best and . The way to use the optimal design curve is to estimate a value for or and nd the associated point on the design curve. A magni ed picture is shown in Figure 15.24. The horizontal axis is the root mean square of relative displacement, = ( ), and the vertical axis is the root mean square of absolute acceleration, = ( ). The optimal curve indicates that softening a suspension decreases the body acceleration, however, it requires a large room for relative displacement. Due to physical constraints, the wheel travel is limited, and hence, we must design the suspension such that to use the available suspension travel, and decrease the body acceleration as low as possible. Mathematically it is equivalent to (15.79) and (15.80). Example 580 Examination of the optimal quarter car model. To examine the optimal design curve and compare practical ways to make a suspension optimal, we assume that there is a quarter car with an o - optimal suspension, indicated by point 1 in Figure 15.27. = 3 = 0 35 = 0 4 (15.117) To optimize the suspension practically, we may keep the sti ness constant and change the damper to a corresponding optimal value, or keep the damping constant and change the sti ness to a corresponding optimal value. However, if it is possible, we may change both, sti ness and damping to a point on the optimal curve depending on the physical constraints and requirements. Point 2 in Figure 15.27 has the same as point 1 with an optimal damping ratio 0 3. Point 3 in Figure 15.27 has the same as point 1 with an optimal natural frequency ratio 0 452. Hence, points 2 and 3 are two alternative optimal designs for the o -optimal point 1. Figure 15.28 compares the acceleration frequency response log for the three points 1, 2, and 3. Point 3 has the minimum acceleration frequency response. Figure 15.29 depicts the absolute displacement frequency response log and Figure 15.30 compares the relative displacement frequency response log for the there points 1, 2, 3. These gures show that both points 2 and 3 introduce better suspension than point 1. Suspension 2 has a higher level of acceleration but needs less relative suspension travel than suspension 3. Suspension 3 has a lower level of acceleration, but it needs more room for suspension travel than suspension 2. Example 581 Comparison of an o -optimal quarter car with two optimals. An alternative method to optimize an o -optimal suspension is to keep the RMS of relative displacement or absolute acceleration constant and nd the associated point on the optimal design curve. Figure 15.31 illustrates two alternative optimal designs, points 2 and 3, for an o - optimal design at point 1. The mass ratio is assumed to be = 3 (15.118) and the suspension characteristics at 1 are = 0 0465 = 0 265 = 2 = 0 15 (15.119) The optimal point corresponding to 1 with the same is at 2 with the characteristics = 0 23 = 0 45 = 0 543 = 0 15 (15.120) and the optimal point with the same as point 1 is a point at 2 with the characteristics: = 0 0949 = 0 1858 = 2 = 0 0982 (15.121) Figure 15.32 depicts the sprung mass vibration amplitude , which shows that both points 2 and 3 have lower overall amplitude specially at second resonance. Figure 15.33 shows the amplitude of relative displacement between sprung and unsprung masses. The amplitude of absolute acceleration of the sprung mass is shown in Figure 15.34. Example 582 F Natural frequencies and vibration isolation requirements. Road irregularities are the most common source of excitation for passenger cars. Therefore, the natural frequencies of vehicle system are the primary factors in determining design requirements for conventional isolators. The natural frequency of the vehicle body supported by the primary suspension is usually between 0 2Hz and 2Hz, and the natural frequency of the unsprung mass, called wheel hop frequency, usually is between 2Hz and 20Hz. The higher values generally apply to military vehicles. The isolation of sprung mass from the uneven road can be improved by using a soft spring, which reduces the primary natural frequency. Lowering the natural frequency always improves the ride comfort, however it causes a design problem due to the large relative motion between the sprung and unsprung masses. One of the most important constraints that suspension system designers have to consider is the rattle-space constraint, the maximum allowable relative displacement. Additional factors are imposed by the overall stability, reliability, and economic or cost factors. Example 583 Optimal characteristics variation. We may collect the optimal and and plot them as shown in Figures 15.35 and 15.36. These gures illustrate the trend of their variation. The optimal value of both and are decreasing functions of relative displacement RMS . So, when more room is available, we may reduce and and have a softer suspension for better ride comfort. Figure 15.37 shows how the optimal and change with each other. 15.5 F Optimization Based on Natural Frequency and Wheel Travel Assume a xed value for the mass ratio and natural frequency ratio to x the position of the nodes in the frequency response plot. Then, an optimal value for damping ratio is F = 35 36 sq 2 37 8 2 + 37 8 2 35 (15.122) where 35 = 2 (1 + ) + 1 (15.123) 36 = 4 1 + (15.124) 37 = \u00a1 2 2 (1 + ) + 1 \u00a2 (15.125) The optimal damping ratio F causes the second resonant amplitude 2 to occur at the second invariant frequency 2. The value of relative displacement at = 2 for = F is, 2 = vuuut \u00b3p 2 37 8 2 35 \u00b4 1 + 2 2 \u00b3 28 p 2 37 29 \u00b4 (15.126) where, 28 = 4 4 (1 + ) 4 4 2 (1 + ) 2 (1 ) + \u00a1 1 + 2 \u00a2 (15.127) 29 = 8 6 (1 + ) 5 + 12 4 (1 + ) 3 (1 ) 2 4 (1 + ) \u00a1 1 + 3 2 2 \u00a2 + \u00a1 1 + 2 \u00a2 (15.128) Proof. Natural frequencies of the sprung and unsprung masses, as given in Equations (15.45) and (15.46), are related to and . When is given, we can evaluate by considering the maximum permissible static de ection, which in turn adjusts the value of natural frequencies. If the values of and are determined and kept xed, then the value of damping ratio which cause the rst resonant amplitude to occur at the second node, can be determined as optimum damping. For a damping ratio less or greater than the optimum, the resonant amplitude would be greater. The frequencies related to the maximum of are obtained by di erentiating with respect to and setting the result equal to zero = 1 2 2 = 1 2 25 \u00a1 8 2 25 26 27 \u00a2 = 0 (15.129) where 25 = \u00a3 2 \u00a1 2 2 1 \u00a2 + \u00a1 1 (1 + ) 2 2 \u00a2\u00a42 +4 2 2 \u00a1 1 (1 + ) 2 2 \u00a22 (15.130) 26 = 8 2 \u00a1 4 2 2 + 1 \u00a2 \u00a1 3 2 2 (1 + ) 1 \u00a2 \u00d7 \u00a1 2 2 (1 + ) 1 \u00a2 (15.131) 27 = 4 \u00a1 4 2 2 + 1 \u00a2 \u00a3 2 2 (1 + ) + 2 \u00a1 1 2 2 \u00a2 1 \u00a4 \u00d7 \u00a3 2 (1 + ) 2 2 2 + 1 \u00a4 (15.132) Now, the optimal value F in Equation (15.122) is obtained if the frequency ratio in Equation (15.129) is replaced with 2 given by Equation (15.61). The optimal damping ratio F makes have a maximum at the second invariant frequency 2. Figure 15.38 illustrates an example of frequency response for di erent including = F. Figure 15.39 shows the sensitivity of F to and . Substituting F in the general expression of , the absolute maximum value of would be equal to 2 given by equation (15.51). Substituting = 2 and = F in Equation (15.25) gives us Equation (15.126) for 2. The lower the natural frequency of the suspension, the more e ective the isolation from road irregularities. So, the sti ness of the main spring must be as low as possible. Figure 15.40 shows the behavior of 2 for = F. Example 584 Nodes in 2 for = F. The relative displacement at second node, 2, is a monotonically increasing function of and has two invariant points. The invariant points of may be found from \u00b1 2 \u00a3\u00a1 2 2 1 \u00a2 + \u00a1 1 (1 + ) 2 \u00a2\u00a4 = \u00a3 2 \u00a1 2 2 1 \u00a2 + \u00a1 1 (1 + ) 2 2 \u00a2\u00a42 + 2 \u00a1 1 (1 + ) 2 2 \u00a2 (15.133) that are, 1 = 0 1 = 0 (15.134) 0 = 1 1 + 0 = 1 + 1 (15.135) The value of at 0 is, 0 = 2 2 (1 + ) 3 \u00a3 4 2 + 2 (1 + ) \u00a4 (15.136) Example 585 F Maximum value of . Figure 15.4 shows that has a node at the intersection of the curves for = 0 and = . There might be a speci c damping ratio to make have a maximum at the node. To nd the maximum value of , we have to solve the following equation for : = 1 2 2 = 1 2 23 \u00a1 4 3 23 30 31 \u00a2 (15.137) where 23 = \u00a3 2 \u00a1 2 2 1 \u00a2 + \u00a1 1 (1 + ) 2 2 \u00a2\u00a42 +4 2 2 \u00a1 1 (1 + ) 2 2 \u00a22 (15.138) 30 = 8 2 5 \u00a3 3 2 2 (1 + ) 1 \u00a4 \u00a1 2 2 (1 + ) 1 \u00a2 (15.139) 31 = 4 5 \u00a3 2 2 (1 + ) + 2 \u00a1 1 2 2 \u00a2 1 \u00a4 \u00d7 \u00a3 2 (1 + ) 2 2 2 + 1 \u00a4 (15.140) Therefore, the maximum occurs at the roots of the equation: 32 8 + 33 6 + 34 2 1 = 0 (15.141) where 32 = 4 (15.142) 33 = 2 4 2 (1 + ) 2 + 4 (1 + ) 2 (15.143) 34 = 2 (1 + ) + 1 2 2 (15.144) Equation (15.141) has two positive roots when is less than a speci c value of damping ratio, , and one positive root when is greater than , where, = ( ) (15.145) The positive roots of Equation (15.141) are 5 and 6, and the corresponding relative displacements are denoted by 5 and 6, where 5 6. The invariant frequencies 5 and 6 would be equal when , and they approach 0 when goes to in nity. The invariant frequency 6 is greater than 5 as long as , and they are equal when . The relative displacements 5 and 6 are monotonically decreasing functions of and they approach 0 when goes to in nity. It is seen from (15.135) that the invariant point at 0 depends on and but the value of 0 depends only on . If is given, then 0 is xed. Therefore, the maximum value of the relative displacement, , cannot be less than 0 and we cannot nd any real value for that causes the maximum of to occur at 0 . The optimum value of could be found when we adjust the maximum value of 6, to be equal to the allowed wheel travel. 15.6 Summary The vertical vibration of vehicles may be modeled by a two-DOF linear system called quarter car model. One-fourth of the body mass, known as sprung mass, is suspended by the main suspension of the vehicle and . The main suspension and are mounted on a wheel of the vehicle, known as unsprung mass. The wheel is sitting on the road by a tire with sti ness . Assuming the vehicle is running on a harmonically bumped road we are able to nd the frequency responses of the sprung and unsprung masses, and relative displacement can be found analytically by taking advantage of the linearity of the system. The frequency response of the sprung mass has four nodes. The rst and fourth nodes are usually out of resonance or out of working frequency range. The middle nodes sit at di erent sides of = 1, and therefore, they cannot be equated and Frahm optimization cannot be applied. The root mean square of the absolute acceleration and relative displacement can be found analytically by applying the RMS optimization method. The RMS optimization method is based on minimizing the absolute acceleration RMS with respect to the relative displacement RMS. The result of RMS optimization introduces an optimal design curve for a xed mass ratio. 15.7 Key Symbols \u00a8 acceleration damping main suspension damper [ ] damping matrix 1 road wave length 2 road wave amplitude dissipation function F force = 1 cyclic frequency [ Hz] damper force spring force cyclic natural frequency [ Hz]\u00a1 2 \u00a2 characteristic equation sti ness main suspension spring sti ness tire sti ness equivalent sti ness [ ] sti ness matrix kinetic energy L Lagrangean mass sprung mass unsprung mass [ ] mass matrix = excitation frequency ratio nodal frequency ratio = natural frequency ratio = ( ) RMS of = ( ) RMS of time period = 2 2 sprung mass acceleration frequency response = 2 2 unsprung mass acceleration frequency response potential energy absolute displacement sprung mass displacement unsprung mass displacement steady-state amplitude of steady-state amplitude of steady-state amplitude of base excitation displacement steady-state amplitude of relative displacement steady-state amplitude of short notation parameter = sprung mass ratio = sprung mass ratio = | | sprung mass relative frequency response = | | sprung mass frequency response = \u00a1 2 \u00a2 damping ratio F optimal damping ratio = | | unsprung mass frequency response = 2 angular frequency [ rad s] = p sprung mass frequency = p unsprung mass frequency natural frequency Subscript node number natural sprung unsprung Exercises 1. Quarter car natural frequencies. Determine the natural frequencies of a quarter car with the following characteristics: = 275kg = 45 kg = 200000N m = 10000N m 2. Equations of motion. Derive the equations of motion for the quarter car model that is shown in Figure 15.1, using the relative coordinates: (a) = = (b) = = (c) = = 3. F Natural frequencies for di erent coordinates. Determine and compare the natural frequencies of the three cases in Exercise 2 and check their equality by employing the numerical data of Exercise 1. 4. Quarter car nodal frequencies. Determine the nodal frequencies of a quarter car with the following characteristics: = 275kg = 45 kg = 200000N m = 10000N m Check the order of the nodal frequencies with the natural frequencies found in Exercise 1. 5. Frequency responses of a quarter car. A car is moving on a wavy road with a wave length 1 = 20m and wave amplitude 2 = 0 08m. = 200 kg = 40 kg = 220000N m = 8000N m = 1000N s m Determine the steady-state amplitude , , and if the car is moving at: (a) = 30km h (b) = 60km h (c) = 120km h 6. Quarter car suspension optimization. Consider a car with = 200 kg = 40 kg = 220000N m = 0 75 and determine the optimal suspension parameters. 7. A quarter car has = 0 45 and = 0 4. What is the required wheel travel if the road excitation has an amplitude = 1cm? 8. F Quarter car and time response. Find the optimal suspension of a quarter car with the following characteristics: = 220 kg = 42 kg = 150000N m = 0 75 and determine the response of the optimal quarter car to a unit step excitation. 9. F Quarter car mathematical model. In the mathematical model of the quarter car, we assumed the tire is always sticking to the road. Determine the condition at which the tire leaves the surface of the road. 10. Optimal damping. Consider a quarter car with = 0 45 and = 0 4. Determine the optimal damping ratio F." + ] + }, + { + "image_filename": "designv11_35_0000174_j.matpr.2021.01.640-Figure6-1.png", + "original_path": "designv11-35/openalex_figure/designv11_35_0000174_j.matpr.2021.01.640-Figure6-1.png", + "caption": "Fig. 6. Drive Shaft With Mesh View.", + "texts": [ + " It has higher explicit quality, more life, less mass, high basic speed and higher force conveying limit. Drive shaft with Geometric view of shown in Fig. 5. The pro-E model is imported to 3 dimensional view. Isometric view permits somewhat precise view of a three-dimensional object on a twodimensional or computer screen, and are a great way to visualize the shape of an object. Generated model is imported into ANSYS, is Drive Shaft. the element type, real constants and the material properties are fed into analyzing area. Drive Shaft With Mesh perspective is appeared in Fig. 6. The drive shaft is meshed number of elements. Mesh influences the accuracy, convergence, speed of the solution, Here used triangular type of mesh is always quick and easy to create. Drive shaft model is divided into finite no of convenient sub elements and each element corners are joined with adjacent element for the purpose of finding bonding strength between sub elements. The fixed support is suitable for the drive shaft .Rotational velocity 650 rad\\s is applied on the drive shaft. The moment is applied 350 N" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_35_0000132_lra.2021.3062583-Figure5-1.png", + "original_path": "designv11-35/openalex_figure/designv11_35_0000132_lra.2021.3062583-Figure5-1.png", + "caption": "Fig. 5. Schematic showing geometrical relationships between cable displacement and various shape parameters of the robot for planar bending at an arbitary angle.", + "texts": [ + " Under the axial load FA, there also exists a deflec- tion magnification, given approximately by C = 1 1\u2212FA/Fc [29], where Fc is the critical load before the spring buckles and is computed to be 30 N in this case. The final deflection of the spring can then be expressed as: \u03b4 = C( \u221a (\u03b4x)2 + (\u03b4y)2) (8) The kinematics model of the robot relates the cable displacement with the tip position (at the end of the DF). There are four cables equally spaced around the SMA spring (only three cables are shown in Fig. 5). Since only one DoF is implemented for active bending during pericardiocentesis, we focus on the kinematics for planar motion of the robot. When one cable is pulled, the distal end bends in one direction as shown in Fig. 5. It should be noted that the cable elongation is negligible throughout the required range of motion of the robot, and is thus not considered in the model. l1 and l3 are two cables symmetrically placed about the center axis of the SMA spring (the same for l2 and l4), and \u03c6 is the bending plane angle. In this case, \u03c6 = \u03c0/2. According to the sine theorem, we can obtain: r sin(\u03b7) = x1 sin(\u03b8) ; lr sin(\u03c02 \u2212 \u03d5) = x2 sin(\u03c02 ) (9) The length of the cable l1 is: l1 = x1 + x2 = sin(\u03b8) sin(\u03b7) r + lr cos(\u03d5) = rsin(\u03b8) + lr sin(\u03b1+ \u03b8) (10) while the length of the antagonistic cable is l3 = (r +D)\u03b8 + lr" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_35_0001419_978-94-007-7401-8-Figure4.36-1.png", + "original_path": "designv11-35/openalex_figure/designv11_35_0001419_978-94-007-7401-8-Figure4.36-1.png", + "caption": "Fig. 4.36 1PaPn3-3PaPn3R-type fully-parallel PM with decoupled Sch\u00f6nflies motions defined by MF = SF = 4, (RF) = (v1; v2; v3;xa), TF = 0, NF = 27, limb topology Pa ? Pn3||R and Pa ? Pn3 ?? R", + "texts": [ + "10j) Pa ? Pn2||R ??R (Fig. 4.11h) Idem No. 14 22. 1PaPn2R\u20133PaPn2RR (Fig. 4.33) Pa ? Pn2||R (Fig. 4.10k) Pa ? Pn2||R ?? R (Fig. 4.11i) Idem No. 14 (continued) 426 4 Fully-Parallel Topologies with Decoupled Sch\u00f6nflies Motions Table 4.4 (continued) No. PM type Limb topology Connecting conditions 23. Pa ?Pn2||R ??R (Fig. 4.34) Pa ? Pn3||R (Fig. 4.10l) Pa ? Pn3 ?? R (Fig. 4.11j) Idem No. 6 24. 1PaPn3\u20133PaPn3R (Fig. 4.35) Pa ? Pn3||R (Fig. 4.10m) Pa ? Pn3 ??|R (Fig. 4.11k) Idem No. 6 25. 1PaPn3\u20133PaPn3R (Fig. 4.36) Pa ? Pn3||R (Fig. 4.10n) Pa ?Pn3 ?? R (Fig. 4.11l) Idem No. 6 26. 1PaPn3\u20133PaPn3R (Fig. 4.37) Pa ? Pn3||R (Fig. 4.10o) Pa ? Pn3 ??R (Fig. 4.11m) Idem No. 6 27. 1CRbRbR\u20133CRbRbRR (Fig. 4.38) C||Rb||Rb||R (Fig. 4.10p) C||Rb||Rb||R ? R (Fig. 4.11n) Idem No. 6 28. 4PaPaPaR (Fig. 4.39) Pa ? P||Pa||R (Fig. 4.12a) Pa ? Pa||Pa ?? R (Fig. 4.12c) Idem No. 6 29. 4PaPaPaR (Fig. 4.40) Pa ? Pa||Pa||R (Fig. 4.12a) Pa ? Pa||Pa ??R (Fig. 4.12b) Idem No. 6 3PaRRbRbRR (Fig. 4.41) Pa ? R||Rb||Rb||R (Fig. 4.12d) Pa " + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_35_0000090_j.physa.2021.125842-Figure1-1.png", + "original_path": "designv11-35/openalex_figure/designv11_35_0000090_j.physa.2021.125842-Figure1-1.png", + "caption": "Fig. 1. Distance between the thermocouple tip and the contact point between tool and workpiece during roughing and finishing operations.", + "texts": [ + " Various techniques are available for measuring the temperature during machining, although they vary in sensitivity, repeatability, accuracy, size, cost and simplicity. However, in a machining process such as milling, the embedded thermocouple remains the mostly used measurement instrument. In this work, the workpiece subsurface temperature was measured using K-type embedded thermocouple technique available in our laboratory. The characteristics of the used thermocouple are shown in Table 1. The thermocouples were placed at a distance of 1.27 mm and 0.508 mm beneath the machined surface for the roughing and finishing tests respectively (Fig. 1). To avoid heat flow alteration between thermocouples and the machined part, hot junctions were welded without metal addition. The holes were then filled with thermal cement (Omega OB-400 cement with 62.46 W/\u25e6C m2 of thermal conductivity) to ensure the securely maintain the thermocouples in position and allow proper heat transfer from the aluminum plate to the hot junction (Fig. 2). As shown above, the aluminum alloy 2024-T3 (typical sheet for fuselage and wing skins) was used as workpiece material in this test" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_35_0001437_012004-Figure4-1.png", + "original_path": "designv11-35/openalex_figure/designv11_35_0001437_012004-Figure4-1.png", + "caption": "Figure 4. Prepaid water meter prototype design.", + "texts": [ + " During water usage, if the number of water pulses has reached the minimum pulse limit, the system will activate the bell and generate a sound signal to inform the customer to top up the water pulse. If the water pulse has run out, the system will automatically move the solenoid valve to close the water flow. As long as the customer is using water, the water pulse refilling process can be done via the remote control to update the number of water IConVET 2020 Journal of Physics: Conference Series 1810 (2021) 012004 IOP Publishing doi:10.1088/1742-6596/1810/1/012004 pulses. In this process, the solenoid valve will remain open so that it does not affect the ongoing water use process. Figure 4 shows the results of the prepaid water meter packaging design. The water flow sensor and the solenoid valve components are packed in one place. The placement of the two components is made like there is curve in the pipe to obtain more minimalist dimensional shape when compared to a connection with a straight pipe. Figure 4 shows the placement of the IR sensor components, EEPROM memory, microcontroller, and solenoid driver as well as the buzzer circuit on the printed circuit board. Meanwhile, the LCD is not installed in a printed circuit board, so it required a cable connection to be placed on the panel box. The entire series of hardware can be put in the packaging box along with the water flow sensor components and the solenoid valve. Based on the testing of the prepaid water meter system, the results of testing the infrared sensor performance data are in the form of remote control button code identification as shown in Figure 5" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_35_0001419_978-94-007-7401-8-Figure5.73-1.png", + "original_path": "designv11-35/openalex_figure/designv11_35_0001419_978-94-007-7401-8-Figure5.73-1.png", + "caption": "Fig. 5.73 2PaPRRR-1PaPRR-1CPaPa-type fully-parallel PMs with uncoupled Sch\u00f6nflies motions defined by MF = SF = 4, RF\u00f0 \u00de \u00bc \u00f0v1; v2; v3;xb\u00de, TF = 0, NF = 19, limb topology C||Pa||Pa and Pa\\P\\\\R||R\\\\R, Pa\\P\\\\R||R (a), Pa||P\\R||R\\\\R, Pa||P\\R||R (b)", + "texts": [ + "2d) C||Pa||Pa (Fig. 5.4o) 4. 3PaPPR1CPaPa (Fig. 5.71b) Pa\\P\\kP\\kR (Fig. 5.2b) Idem No. 3 Pa\\P\\kP\\\\R (Fig. 5.2c) C||Pa||Pa (Fig. 5.40) 5. 3PaPPR1CPaPat (Fig. 5.72a) Pa||P\\P\\kR (Fig. 5.2a) Idem No. 3 Pa||P\\P\\\\R (Fig. 5.2d) C||Pa||Pat (Fig. 5.4p) 6. 3PaPPR1CPaPat (Fig. 5.72b) Pa\\P\\kP\\kR (Fig. 5.2b) Idem No. 3 Pa\\P\\kP\\\\R (Fig. 5.2c) C||Pa||Pat (Fig. 5.4p) (continued) 534 5 Topologies with Uncoupled Sch\u00f6nflies Motions Table 5.5 (continued) No. PM type Limb topology Connecting conditions 7. 2PaPRRR1PaPRR1CPaPa (Fig. 5.73a) Pa\\P\\\\R||R\\\\R (Fig. 5.3a) Idem No. 3 Pa\\P\\\\R||R (Fig. 5.2e) C||Pa||Pa (Fig. 5.40) 8. 2PaPRRR1PaPRR1CPaPa (Fig. 5.73b) Pa||P\\R||R\\\\R (Fig. 5.3b) Idem No. 3 Pa||P\\R||R (Fig. 5.2f) C||Pa||Pa (Fig. 5.40) 9. 2PaPRRR1PaPRR1CPaPat (Fig. 5.74a) Pa\\P\\\\R||R\\\\R (Fig. 5.3a) Idem No. 3 Pa\\P\\\\R||R (Fig. 5.2e) C||Pa||Pat(Fig. 5.4p) 10. 2PaPRRR1PaPRR1CPaPat (Fig. 5.74b) Pa||P\\R||R\\\\R (Fig. 5.3b) Idem No. 3 Pa||P\\R||R (Fig. 5.2f) C||Pa||Pat (Fig. 5.4p) 11. 2PaRPRR1PaRPR1CPaPa (Fig. 5.75) Pa\\R\\P\\kR\\R (Fig. 5.3c) Idem No. 3 Pa\\R\\P\\kR (Fig. 5.2g) C||Pa||Pa (Fig. 5.40) 12. 2PaRPRR1PaRPR1CPaPat (Fig. 5.76) Pa\\R\\P\\kR\\R (Fig. 5.3c) Idem No" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_35_0000395_978-3-030-67750-3_16-Figure28.2-1.png", + "original_path": "designv11-35/openalex_figure/designv11_35_0000395_978-3-030-67750-3_16-Figure28.2-1.png", + "caption": "Fig. 28.2 (a) and (b) Mode 1 Fig. 28.2(a) and Mode 2 Fig. 28.2(b) working principles", + "texts": [ + " The fire extinguisher nozzle experiment was conducted in a closed huge environment to ensure the data obtained covered the whole space and avoid wind, temperature and pressure disturbance. The large space allows the researchers to closely monitor the dust particles coverage. It also allows the researchers to measure the distance should be because the particles can be seen and covered the large area. The safety and health issues is also important because the fine dry powder particles can be hazardous to human respiratory system. These factors can give a significant impact on the experiment result. The bi-nozzle system operates in dual mode. Mode one, as shown in Fig. 28.2(a), works with only one outlet, which is the center outlet with 12 mm of diameter. The working principle behind this mode is when the nozzle is fully closed, the pathway for the dry powder to travel through is only limited to one. The aim is to create a high pressure and velocity stream of the fine particles flow out through the outlet, thus creating a narrow spray distance. As for mode two, as presented in Fig. 28.2. Figure 28.2(b), the dry powder particles for this mode flow two direction through outlet 1 and outlet 2. For the current nozzle, outlet 2 has a bigger diameter with 52 mm. The working principle of these outlets requires the nozzle to turn counter clockwise to allow the particles travel through two pathways. This is to allow the fluid to pass through a more significant diameter outlet (outlet 2) that create a wider spray angle with a slow velocity. Figure 28.3 illustrate the evolution of the bi-nozzle design to the existing one" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_35_0001419_978-94-007-7401-8-Figure3.24-1.png", + "original_path": "designv11-35/openalex_figure/designv11_35_0001419_978-94-007-7401-8-Figure3.24-1.png", + "caption": "Fig. 3.24 Overactuated PMs with coupled Sch\u00f6nflies motions of types 4RPC (a) and 4PRC (b) defined by MF \u00bc SF \u00bc 4; RF\u00f0 \u00de \u00bc v1; v2; v3;xd\u00f0 \u00de, TF = 0, NF = 6, limb topology R\\P\\kC (a) and P\\R||C (b)", + "texts": [ + "43, 3.44, 3.45, 3.46, 3.47, 3.48, 3.49 No. PM type Limb topology Connecting conditions 1. 4RRC (Fig. 3.21a) R||R||C (Fig. 3.2a) The cylindrical joints of the four limbs have parallel axes 2. 4RCR (Fig. 3.21b) R||C||R (Fig. 3.2b) Idem No. 1 3. 4PCP (Fig. 3.22a) P\\C\\P (Fig. 3.2c) Idem No. 1 4. 4CPP (Fig. 3.22b) C\\P\\\\P (Fig. 3.2d) Idem No. 1 5. 4PPC (Fig. 3.22c) P\\P\\\\C (Fig. 3.2e) Idem No. 1 6. 4PCR (Fig. 3.23a) P\\C||R (Fig. 3.2f) Idem No. 1 7. 4RCP (Fig. 3.23b) R||C\\P (Fig. 3.2g) Idem No. 1 8. 4RPC (Fig. 3.24a) R\\P\\kC (Fig. 3.2h) Idem No. 1 9. 4PRC (Fig. 3.24b) P\\R||C (Fig. 3.2i) Idem No. 1 10. 4CRP (Fig. 3.25a) C||R\\P (Fig. 3.2j) Idem No. 1 11. 4CPR (Fig. 3.25b) C\\P\\kR (Fig. 3.2k) Idem No. 1 12. 4RPC (Fig. 3.26a) R\\P\\kC (Fig. 3.2i) Idem No. 1 13. 4CRR (Fig. 3.26b) C||R||R (Fig. 3.2m) Idem No. 1 14. 4RRRRR (Fig. 3.27a) R\\R||R\\R||R (Fig. 3.3a) The first and the last revolute joints of the four limbs have parallel axes 15. 4RRRRR (Fig. 3.27b) R||R\\R||R\\R (Fig. 3.3b) Idem No. 14 16. 4RRRRR (Fig. 3.28a) R||R||R\\R||R (Fig. 3.3c) The first revolute joints of the four limbs have parallel axes 17" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_35_0001419_978-94-007-7401-8-Figure3.80-1.png", + "original_path": "designv11-35/openalex_figure/designv11_35_0001419_978-94-007-7401-8-Figure3.80-1.png", + "caption": "Fig. 3.80 Overactuated PMs with coupled Sch\u00f6nflies motions of types 4PaRRP (a) and 4RPaRR (b) defined by MF \u00bc SF \u00bc 4; RF\u00f0 \u00de \u00bc v1; v2; v3;xd\u00f0 \u00de, TF = 0, NF = 18, limb topology Pa\\R||R\\P (a) and R\\Pa\\kR||R (b)", + "texts": [ + "50j0) The last joints of the four limbs have parallel axes (continued) 3.2 Topologies with Complex Limbs 359 Table 3.5 (continued) No. PM type Limb topology Connecting conditions 38. 4PPPaR (Fig. 3.76b) P\\P\\kPa\\\\R (Fig. 3.50k0) Idem No. 37 39. 4PaPPR (Fig. 3.77a) Pa||P\\P\\\\R (Fig. 3.50l0) Idem No. 37 40. 4PaPPR (Fig. 3.77b) Pa\\P\\kP\\\\R (Fig. 3.50m0) Idem No. 37 41. 4PaRPR (Fig. 3.78) Pa\\R\\P\\kR (Fig. 3.50n0) Idem No. 37 42. 4PaPRR (Fig. 3.79a) Pa||P\\R||R (Fig. 3.50o0) Idem No. 37 43. 4PaPRR (Fig. 3.79b) Pa\\P\\\\R||R (Fig. 3.50p0) Idem No. 37 44. 4PaRRP (Fig. 3.80a) Pa\\R||R\\P (Fig. 3.50q0) Idem No. 23 45. 4RPaRR (Fig. 3.80b) R\\Pa\\kR||R (Fig. 3.50s0) Idem No. 37 46. 4PaRRR (Fig. 3.81) Pa\\R||R||R (Fig. 3.50r0) Idem No. 37 47. 4PaPRP (Fig. 3.82a) Pa||P\\R\\P (Fig. 3.50t0) Idem No. 23 48. 4PaPRP (Fig. 3.82b) Pa\\P\\\\R\\P (Fig. 3.50u0) Idem No. 23 49. 4CPaP (Fig. 3.83a) C||Pa\\P (Fig. 3.50v0) The cylindrical joints of the four limbs have parallel directions 50. 4PaCP (Fig. 3.83b) Pa||C\\P (Fig. 3.50w0) Idem No. 49 51. 4CRPa (Fig. 3.84a) C||R||Pa (Fig. 3.50x0) Idem No. 49 52. 4PCPa (Fig. 3.84b) P\\C||Pa (Fig. 3.50y0) Idem No. 49 53" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_35_0001745_jada.archive.1940.0343-Figure20-1.png", + "original_path": "designv11-35/openalex_figure/designv11_35_0001745_jada.archive.1940.0343-Figure20-1.png", + "caption": "Fig. 20.\u2014 Line of force of vertical casting", + "texts": [ + " Eventually, centrifugal force will operate to neutralize inertia and carry the metal in the direction desired. While this is not o f great consequence with comparatively low fusing metals of wide melting range, such as casting gold, for the reason given in the previous dis cussion of crucibles it is very important in casting platinum. While it is possible to partially overcome this by means of a properly designed crucible, it does present an obstacle which is very likely to cause cold shuts unless the mold temperature is undesirably high. Figure 20 shows the lines o f force of the vertical type casting machine. Here again, the same principles of physics hold true. The relation of the metal to the center of rotation determines the lines of force exerted upon it. The direction of rotation is clockwise. The force of inertia, therefore, tends to carry the metal against the floor of the cru cible ; but since the metal is already lying in that spot, it cannot splash or spin out. Since the metal cannot move by inertia, and it will not remain stationary, cen trifugal force is exerted immediately to carry it in the desired direction" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_35_0001419_978-94-007-7401-8-Figure3.21-1.png", + "original_path": "designv11-35/openalex_figure/designv11_35_0001419_978-94-007-7401-8-Figure3.21-1.png", + "caption": "Fig. 3.21 Overactuated PMs with coupled Sch\u00f6nflies motions of types 4RRC (a) and 4RCR (b) defined by MF \u00bc SF \u00bc 4; RF\u00f0 \u00de \u00bc v1; v2; v3;xd\u00f0 \u00de, TF = 0, NF = 6, limb topology R||R||C (a) and R||C||R (b)", + "texts": [ + " 4RPPP (Fig. 3.20a) R\\P\\kP\\\\P (Fig. 3.1f0) Idem No. 26 34. 4RPPP (Fig. 3.20b) R||P\\P\\\\P (Fig. 3.1g0) Idem No. 26 3.1 Topologies with Simple Limbs 243 Table 3.2 Limb topology and connecting conditions of the overactuated solutions with no idle mobilities presented in Figs. 3.21, 3.22, 3.23, 3.24, 3.25, 3.26, 3.27, 3.28, 3.29, 3.30, 3.31, 3.32, 3.33, 3.34, 3.35, 3.36, 3.37, 3.38, 3.39, 3.40, 3.41, 3.42, 3.43, 3.44, 3.45, 3.46, 3.47, 3.48, 3.49 No. PM type Limb topology Connecting conditions 1. 4RRC (Fig. 3.21a) R||R||C (Fig. 3.2a) The cylindrical joints of the four limbs have parallel axes 2. 4RCR (Fig. 3.21b) R||C||R (Fig. 3.2b) Idem No. 1 3. 4PCP (Fig. 3.22a) P\\C\\P (Fig. 3.2c) Idem No. 1 4. 4CPP (Fig. 3.22b) C\\P\\\\P (Fig. 3.2d) Idem No. 1 5. 4PPC (Fig. 3.22c) P\\P\\\\C (Fig. 3.2e) Idem No. 1 6. 4PCR (Fig. 3.23a) P\\C||R (Fig. 3.2f) Idem No. 1 7. 4RCP (Fig. 3.23b) R||C\\P (Fig. 3.2g) Idem No. 1 8. 4RPC (Fig. 3.24a) R\\P\\kC (Fig. 3.2h) Idem No. 1 9. 4PRC (Fig. 3.24b) P\\R||C (Fig. 3.2i) Idem No. 1 10. 4CRP (Fig. 3.25a) C||R\\P (Fig. 3.2j) Idem No. 1 11. 4CPR (Fig. 3.25b) C\\P\\kR (Fig. 3.2k) Idem No. 1 12. 4RPC (Fig" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_35_0000557_tmech.2021.3085761-Figure1-1.png", + "original_path": "designv11-35/openalex_figure/designv11_35_0000557_tmech.2021.3085761-Figure1-1.png", + "caption": "Fig. 1 Traditional active yaw system", + "texts": [ + " Index terms-Wind energy, Maglev, Active yaw system, Pitching suppression, Sliding mode controller. I. INTRODUCTION ind energy has attracted worldwide attention due to its non-pollution and renewability. The yaw system is one of the key parts of the horizontal-axis wind turbines (HAWT) for the maximum energy [1,2]. But at present, the active yaw system adopts multi-motor driving and gear coupling technology, which has many defects such as complicated structure and higher friction [3-5], as shown in Fig.1. Therefore, we put forward a maglev wind yaw system (MWYS) as shown in Fig.2 [6]. A disc motor is introduced to suspend nacelle and yaw for wind, rotor coils generate the axial suction to suspend the nacelle, and then the yaw torque derived from the stator coils to drive the nacelle for wind. As a result, the yaw driving power reduces hugely because of the suspension. This paper focuses on the nacelle suspension for the MWYS. This work was supported by the national natural science foundation of china under Grants 61473170" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_35_0001419_978-94-007-7401-8-Figure3.132-1.png", + "original_path": "designv11-35/openalex_figure/designv11_35_0001419_978-94-007-7401-8-Figure3.132-1.png", + "caption": "Fig. 3.132 4PaRPRR-type overactuated PM with coupled Sch\u00f6nflies motions defined by MF \u00bc SF \u00bc 4; RF\u00f0 \u00de \u00bc v1; v2; v3;xd\u00f0 \u00de, TF = 0, NF = 14, limb topology Pa\\R||P||R\\kR", + "texts": [ + "125) P||R||R\\Pa||R (Fig. 3.52r) Idem No. 13 39. 4RPRPaR (Fig. 3.126) R||P||R\\Pa||R (Fig. 3.52s) Idem No. 13 40. 4RPRPaR (Fig. 3.127) R||P||R\\Pa||R (Fig. 3.52t) Idem No. 13 41. 4RRPPaR (Fig. 3.128) R||R||P\\Pa||R (Fig. 3.52u) Idem No. 13 42. 4PPaRRR (Fig. 3.129a) P\\Pa\\kR||R||\\kR (Fig. 3.52v) Idem No. 13 43. 4PaPRRR (Fig. 3.129b) Pa\\P\\\\R||R\\\\R (Fig. 3.52w) Idem No. 13 44. 4PPaRRR (Fig. 3.130) P\\Pa\\\\R||R\\kR (Fig. 3.52x) Idem No. 13 45. 4PaPRRR (Fig. 3.131) Pa\\P||R||R\\kR (Fig. 3.52y) Idem No. 13 46. 4PaRPRR (Fig. 3.132) Pa\\R||P||R\\kR (Fig. 3.52z) Idem No. 13 47. 4PaRPRR (Fig. 3.133) Pa\\R||P||R\\kR (Fig. 3.52a0) Idem No. 13 48. 4PaRRPR (Fig. 3.134) Pa\\R||R||P\\kR (Fig. 3.52b0) Idem No. 13 49. 4RCRPa (Fig. 3.135) R\\C||R\\kPa (Fig. 3.52c0) Idem No. 12 50. 4RRCPa (Fig. 3.136) R\\R||C\\kPa (Fig. 3.52d0) Idem No. 12 51. 4CRRPa (Fig. 3.137) C||R\\R||Pa (Fig. 3.52e0) Idem No. 15 52. 4RCRPa (Fig. 3.138) R||C\\R||Pa (Fig. 3.52f0) Idem No. 15 53. 4CRPaR (Fig. 3.139) C||R\\Pa||R (Fig. 3.52g0) Idem No. 13 54. 4RCPaR (Fig. 3.140) R||C\\Pa||R (Fig" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_35_0001419_978-94-007-7401-8-Figure6.29-1.png", + "original_path": "designv11-35/openalex_figure/designv11_35_0001419_978-94-007-7401-8-Figure6.29-1.png", + "caption": "Fig. 6.29 Maximaly regular fully-parallel PMs with Sch\u00f6nflies motions of types 3PPaPR1RPaPaP (a) and 3PPaPR-1RPaPatP (b) defined by MF = SF = 4, (RF) = (v1; v2; v3;xb), TF = 0, NF = 15, limb topology P||Pa\\P ??R, P||Pa\\P\\||R and R||Pa||Pa||P (a), R||Pa||Pat||P (b)", + "texts": [ + "4k) The last joints of the four limbs have superposed axes/directions. The actuated prismatic joints of limbs G1, G2 and G3 have orthogonal directions 2. 3PPPaR-1RPaPaP (Fig. 6.27b) P\\P\\||Pa\\R (Fig. 4.8a) P\\P\\||Pa||R, (Fig. 4.8c) R||Pa||Pa||P (Fig. 5.4k) Idem no. 1 3. 3PPPaR-1RPaPatP (Fig. 6.28a) P\\P\\||Pa\\||R (Fig. 4.8b) P\\P\\||Pa||R, (Fig. 4.8c) R||Pa||Pat||P (Fig. 5.4l) Idem no. 1 4. 3PPPaR-1RPaPatP (Fig. 6.28b) P\\P\\||Pa\\R (Fig. 4.8a) P\\P\\||Pa||R, (Fig. 4.8c) R||Pa||Pat||P (Fig. 5.4l) Idem no. 1 5. 3PPaPR-1RPaPaP (Fig. 6.29a) P||Pa\\P\\\\R (Fig. 4.8d) P||Pa\\P\\||R (Fig. 4.8e) R||Pa||Pa||P (Fig. 5.4k) Idem no. 1 6. 3PPaPR-1RPaPatP (Fig. 6.29b) P||Pa\\P\\\\R (Fig. 4.8d) P||Pa\\P\\||R (Fig. 4.8e) R||Pa||Pat||P (Fig. 5.4l) Idem no. 1 7. 2PPaRRR-1PPaRR-1RPaPaP (Fig. 6.30) P||Pa||R||R\\R (Fig. 4.9a) P||Pa||R||R (Fig. 4.8g) R||Pa||Pa||P (Fig. 5.4k) Idem no. 1 8. 2PPaRRR-1PPaRR1RPaPatP (Fig. 6.31) P||Pa||R||R\\R (Fig. 4.9a) P||Pa||R||R (Fig. 4.8g) R||Pa||Pat||P (Fig. 5.4l) Idem no. 1 9. 3PPPaR-1RPPaPa (Fig. 6.32a) P\\P\\||Pa\\||R (Fig. 4.8b) P\\P\\||Pa||R, (Fig. 4.8c) R||P||Pa||Pa (Fig. 5.4m) The last revolute joints of limbs G1, G2 and G3 and the revolute joint between links 7D and 9 have superposed axes" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_35_0000213_j.triboint.2021.106993-Figure11-1.png", + "original_path": "designv11-35/openalex_figure/designv11_35_0000213_j.triboint.2021.106993-Figure11-1.png", + "caption": "Fig. 11. Pressure Contour (10,000 N load, 9000 rpm) by CFD simulation; (a) Smooth TPJB, (b) Pocketed TPJB.", + "texts": [ + " 9 depicts the modeled TPJB with the mesh for the smooth and pocketed TPJB. A mesh independence study was conducted on the Thermo-Hydrodynamic (THD) Model, and is summarized in Fig. 10. The number of elements was selected to be 433,436 based on the results in Fig. 10 and the corresponding computation times. The final mesh was identical for the smooth and pocketed bearings as illustrated in Fig. 9. Comparison between the simulation results for the smooth and pocketed TPJBs demonstrates the benefits of the latter. Fig. 11 shows the pressure contours when the pockets are installed on pad1, pad2, and J. Yang and A. Palazzolo Tribology International 159 (2021) 106993 pad3 of the pocketed TPJB. The color contour levels are adjusted to focus on the pressure distribution of the upper pads, since their pressures are much lower than that in the bottom pads. As shown in Fig. 11(b), the peak pressures in the pocketed TPJB are produced at the step and pocket trailing edges, which helps stabilize the pad tilting motion. The pressure is suddenly decreased in the circumferential direction where the film thickness is suddenly increased, to conserve fluid momentum. Thus, the pressure suddenly drops at the leading edge of the pockets, generating cavitation in the pocket. The negative gauge pressure inside the pockets is recovered at the trailing edge of the pocket. In addition, Fig", + " 15 shows the temperature distribution on the pad and journal at 9000 rpm and 10,000 N load, for both the smooth and pocketed bearings. Figs. 15(a-1) shows the temperature varying in the axial direction due to the three nozzle oil injection cooling flows. The pad surface temperature increases along the rotation direction, and the bottom pads have relatively higher temperatures because of the larger viscous heat generation in the thinner film. As shown in Fig. 15, the overall temperature at both pad and journal surfaces is decreased in the pocketed TPJB. Fig. 11 shows that the negative gauge pressure in the pockets yields a lower load condition on the upper pads, than for the smooth TPJB. Thus, the journal lifts to the opposite direction of the applied load, and the eccentricity ratio is decreased, with an increase in the minimum film thickness. This indicates the pockets can provide power loss reduction benefits without sacrificing bearing load capacity. The eccentricity ratio is increased with increasing applied load, and decreased with increased operating speed, as shown in Fig" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_35_0000743_s11029-021-09953-2-Figure1-1.png", + "original_path": "designv11-35/openalex_figure/designv11_35_0000743_s11029-021-09953-2-Figure1-1.png", + "caption": "Fig. 1. Periodicity cell (RVE) of an \u201cuncut\u201d rubber-cord layer.", + "texts": [ + " It is shown that it is possible to construct a potential for a transversely isotropic medium based on the Treloar potential and, in some cases, use it to approximate the properties of a rubber-cord layer. Examples of determination of its parameters are given, and the stability of this procedure is proved by numerical experiments. The rubber-cord layers in a radial tire are employed as a carcass and breaker. It is assumed that the structure of the material is close to periodic. Its periodicity cell is shown in Fig. 1. We will consider an \u201cuntruncated\u201d cell of the frame with an aspect ratio close to one ( h l , 1 ). In the case of an inhomogeneous layer considered as a 3D body, the very method for determination of averaged elastic properties needs changing. The well-known classical method presupposes the presence of a representative region on whose boundary displacements of a special type (task I) or surface forces (task II) are specified. In this work, the elastic moduli are defined and displacements are specified on the boundary" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_35_0000717_s11665-021-06062-y-Figure2-1.png", + "original_path": "designv11-35/openalex_figure/designv11_35_0000717_s11665-021-06062-y-Figure2-1.png", + "caption": "Fig. 2 Schematic of tensile specimens machined from notional part", + "texts": [ + " Contact e-mails: pcollins@iastate.edu and dgh0@lehigh.edu. *According to the seminal publications on the concept of Materials State by E. Lindgren and J. Aldrin, they define Materials State Awareness (MSA) as \u2018\u2018Digitally Enabled Reliable Nondestructive Quantitative Materials / Damage Characterization Regardless of Scale\u2019\u2019 JMEPEG ASM International https://doi.org/10.1007/s11665-021-06062-y 1059-9495/$19.00 Journal of Materials Engineering and Performance uniaxial tensile specimens have been machined (see Fig. 2). The orientations, shown in Fig. 2, are as follows: \u2022 x\u2014along the build path of the deposited beads, \u2022 y\u2014normal to the deposited beads in the same plane as deposition, \u2022 z\u2014normal to both the deposited beads and the plane con- taining x and y, \u2022 xc\u2014in the plane of deposition where deposited beads crisscross, \u2022 zc\u2014normal to the plane of deposition where deposited beads crisscross, \u2022 I\u2014normal to the plane of deposition where the specimen contains the interface between the plate and the deposit, \u2022 k\u201445 in the y\u2013z plane, and \u2022 kc\u201445 in the y\u2013z plane where deposited beads criss- cross" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_35_0001419_978-94-007-7401-8-Figure5.27-1.png", + "original_path": "designv11-35/openalex_figure/designv11_35_0001419_978-94-007-7401-8-Figure5.27-1.png", + "caption": "Fig. 5.27 2PaRRRR-1PaRRR-1RPaPaP-type fully-parallel PMs with uncoupled Sch\u00f6nflies motions defined by MF = SF = 4, RF\u00f0 \u00de \u00bc \u00f0v1; v2; v3;xb\u00de, TF = 0, NF = 19, limb topology Pa\\R||R||R\\kR, Pa\\R||R||R and R||Pa||Pa||P", + "texts": [ + " 5.2g) R||Pa||Pat||P (Fig. 5.4l) 20. 2PaRPRR-1PaRPR1RPPaPat (Fig. 5.26) Pa\\R\\P\\kR\\R (Fig. 5.3c) The revolute joint between links 7D and 9 of G4 limb and the last revolute joints of limbs G1, G2 and G3 have superposed axes. The revolute joints between links 4 and 5 of limbs G1, G2 and G3 have orthogonal axes Pa\\R\\P\\kR (Fig. 5.2g) R||P||Pa||Pat (Fig. 5.4n) (continued) 5.1 Fully-Parallel Topologies 529 Table 5.2 (continued) No. PM type Limb topology Connecting conditions 21. 2PaRRRR-1PaRRR1RPaPaP (Fig. 5.27) Pa\\R||R||R\\kR (Fig. 5.3d) Idem No. 17 Pa\\R||R||R (Fig. 5.2h) R||Pa||Pa||P (Fig. 5.4k) 22. 2PaRRRR-1PaRRR1RPPaPa (Fig. 5.28) Pa\\R||R||R\\kR (Fig. 5.3d) Idem No. 20 Pa\\R||R||R (Fig. 5.2h) R||P||Pa||Pa (Fig. 5.4m) 23. 2PaRRRR-1PaRRR1RPaPatP (Fig. 5.29) Pa\\R||R||R\\kR (Fig. 5.3d) Idem No. 17 Pa\\R||R||R (Fig. 5.2h) R||Pa||Pat||P (Fig. 5.4l) 24. 2PaRRRR-1PaRRR1RPPaPat (Fig. 5.30) Pa\\R||R||R\\kR (Fig. 5.3d) Idem No. 20 Pa\\R||R||R (Fig. 5.2h) R||P||Pa||Pat (Fig. 5.4n) Table 5.3 Limb topology and connecting conditions of the fully-parallel solutions with no idle mobilities presented in Figs" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_35_0001419_978-94-007-7401-8-Figure1.3-1.png", + "original_path": "designv11-35/openalex_figure/designv11_35_0001419_978-94-007-7401-8-Figure1.3-1.png", + "caption": "Fig. 1.3 Various representations of a Gough-Stewart type parallel robot: physical implementation (a) CAD model (b), structural diagram (c) and its associated graph (d), A-limb (e) and its associated graph (f)", + "texts": [ + " The \u2018\u2018independence\u2019\u2019 of the kinematic chains associated with the limbs of a parallel mechanism must be limited to the structural point of view. This \u2018\u2018independence\u2019\u2019 is not valid for the kinematic and static point of view. 1.1 Terminology 5 A parallel robot can be illustrated by a physical implementation or by an abstract representation. The physical implementation is usually illustrated by robot photography and the abstract representation by a CAD model, structural diagram and structural graph. Figure 1.3 gives an example of the various representations of a Gough-Stewart type parallel robot largely used today in industrial applications. The physical implementation in Fig. 1.3a is a photograph of the parallel robot built by Deltalab (http://www.deltalab.fr/). In a CAD model (Fig. 1.3b) the links and the joints are represented as being as close as possible to the physical implementation (Fig. 1.3a). In a structural diagram (Fig. 1.3c) they are represented by simplified symbols, such as those introduced in Fig. 1.1, respecting the geometric relations defined by the relative positions of joint axes. A structural graph (Fig. 1.3d) is a network of vertices or nodes connected by edges or arcs with no geometric relations. The links are noted in the nodes and the joints on the edges. We can see that the Gough-Stewart type parallel robot has six identical limbs denoted in Fig. 1.3c by A, B, C, D, E and F. The final link is the mobile platform 4 : 4A : 4B: 4C : 4D : 4E: 4F and the reference member is the fixed platform 1A : 1B: 1C : 1D: 1E : 1F : 0. Each limb is connected to both platforms by spherical pairs. A prismatic pair is actuated in each limb. The spherical pairs are not actuated and are called passive pairs. The two platforms are polinary links, the other two links of each limb are binary links. The parallel mechanism 6-SPS-type associated with the Gough-Stewart type parallel robot is a complex mechanism with a multi-loop associated graph (Fig. 1.3d). It has six simple limbs of type SPS. The actuated pair is underlined. The simple open kinematic chain associated with A-limb is denoted by A (1A : 0\u20132A\u20133A\u2013 4A : 4)\u2014Fig. 1.3e and its associated graph is tree-type (Fig. 1.3f). 6 1 Introduction 1.1 Terminology 7 We consider the general case of a robot in which the end-effector is connected to the reference link by k C 1 kinematic chains. The end-effector is a binary or polynary link called a mobile platform in the case of parallel robots, and a monary link for serial robots. The reference link may either be the fixed base or may be deemed to be fixed. The kinematic chains connecting the end-effector to the reference link can be simple or complex. They are called limbs or legs in the case of parallel robots" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_35_0001419_978-94-007-7401-8-Figure5.16-1.png", + "original_path": "designv11-35/openalex_figure/designv11_35_0001419_978-94-007-7401-8-Figure5.16-1.png", + "caption": "Fig. 5.16 3PaPPR-1RPPaPa-type fully-parallel PMs with uncoupled Sch\u00f6nflies motions defined by MF = SF = 4, RF\u00f0 \u00de \u00bc \u00f0v1; v2; v3;xb\u00de, TF = 0, NF = 21, limb topology R||P||Pa||Pa and Pa||P\\P\\kR, Pa||P\\P\\\\R (a), Pa\\P\\kP\\kR, Pa\\P\\kP\\\\R (b)", + "texts": [ + "2 Limb topology and connecting conditions of the fully-parallel solutions with no idle mobilities presented in Figs. 5.15, 5.16, 5.17, 5.18, 5.19, 5.20, 5.21, 5.22, 5.23, 5.24, 5.25, 5.26, 5.27, 5.28, 5.29, 5.30 No. PM type Limb topology Connecting conditions 1. 3PaPPR-1RPaPaP (Fig. 5.15a) Pa||P\\P\\kR (Fig. 5.2a) The last joints of the four limbs have superposed axes/directions Pa||P\\P\\\\R (Fig. 5.2d) R||Pa||Pa||P (Fig. 5.4k) 2. 3PaPPR-1RPaPaP (Fig. 5.15b) Pa\\P\\kP\\kR (Fig. 5.2b) Idem No. 1 Pa\\P\\kP\\\\R (Fig. 5.2c) R||Pa||Pa||P (Fig. 5.4k) 3. 3PaPPR-1RPPaPa (Fig. 5.16a) Pa||P\\P\\kR (Fig. 5.2a) The revolute joint between links 7D and 9 of G4 limb and the last revolute joints of limbs G1, G2 and G3 have superposed axes Pa||P\\P\\\\R (Fig. 5.2d) R||P||Pa||Pa (Fig. 5.4m) 4. 3PaPPR-1RPPaPa (Fig. 5.16b) Pa\\P\\kP\\kR (Fig. 5.2b) Idem No. 3 Pa\\P\\kP\\\\R (Fig. 5.2c) R||P||Pa||Pa (Fig. 5.4m) 5. 3PaPPR1RPaPatP (Fig. 5.17a) Pa||P\\P\\kR (Fig. 5.2a) Idem No. 1 Pa||P\\P\\\\R (Fig. 5.2d) R||Pa||Pat||P (Fig. 5.4l) 6. 3PaPPR1RPaPatP (Fig. 5.17b) Pa\\P\\kP\\kR (Fig. 5.2b) Idem No. 1 Pa\\P\\kP\\\\R (Fig. 5.2c) R||Pa||Pat||P (Fig. 5.4l) 7. 3PaPPR-1RPPaPat (Fig. 5.18a) Pa||P\\P\\kR (Fig. 5.2a) Idem No. 3 Pa||P\\P\\\\R (Fig. 5.2d) R||P||Pa||Pat (Fig. 5.4n) 8. 3PaPPR-1RPPaPat (Fig. 5.18b) Pa\\P\\kP\\kR (Fig. 5.2b) Idem No. 3 Pa\\P\\kP\\\\R (Fig" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_35_0001419_978-94-007-7401-8-Figure4.20-1.png", + "original_path": "designv11-35/openalex_figure/designv11_35_0001419_978-94-007-7401-8-Figure4.20-1.png", + "caption": "Fig. 4.20 1PPn2R-3PPn2RR-type fully-parallel PM with decoupled Sch\u00f6nflies motions defined by MF = SF = 4, RF\u00f0 \u00de \u00bc v1; v2; v3;xa\u00f0 \u00de, TF = 0, NF = 15, limb topology P||Pn2||R and P||Pn2||R ? R", + "texts": [ + " 1PaRPR\u20133PaRPRR (Fig. 4.17) Pa ? R ? P ??R (Fig. 4.8h) Pa ? R ? P ??R ? R (Fig. 4.9b) The last revolute joints of the four limbs have parallel axes. The revolute joints between links 4 and 5 of limbs G1, G2 and G3 hvave orthogonal axes. The revolute joints between links 4 and 5 of limbs G3 and G4 have parallel axes 7. 1PRRbR\u20133PRRbRR (Fig. 4.18) P||R||Rb||R (Fig. 4.8i) P||R||Rb||R ? R (Fig. 4.9c) Idem No. 1 8. 1PPn2R\u20133PPn2RR (Fig. 4.19) P||Pn2||R (Fig. 4.8j) P||Pn2||R ? R (Fig. 4.9d) 9. 1PPn2R\u20133PPn2RR (Fig. 4.20) P||Pn2||R (Fig. 4.8k) P||Pn2||R ?R (Fig. 4.9e) Idem No. 1 10. 1PPn3\u20133PPn3R (Fig. 4.21) P||Pn3 (Fig. 4.8l) P||Pn3 ? R (Fig. 4.9f) Idem No. 1 11. 1PPn3\u20133PPn3R (Fig. 4.22) P||Pn3 (Fig. 4.8m) P||Pn3 ? R (Fig. 4.9g) Idem No. 1 (continued) 4.2 Topologies with Complex Limbs 425 Table 4.4 (continued) No. PM type Limb topology Connecting conditions 12. 1CRbR\u20133CRbRR (Fig. 4.23) C||Rb||R (Fig. 4.8n) C||Rb||R ? R (Fig. 4.9h) Idem No. 1 13. 4PPaPaR (Fig. 4.24) P||Pa||Pa||R (Fig. 4.10b) P||Pa||Pa ? R (Fig. 4" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_35_0000174_j.matpr.2021.01.640-Figure16-1.png", + "original_path": "designv11-35/openalex_figure/designv11_35_0000174_j.matpr.2021.01.640-Figure16-1.png", + "caption": "Fig. 16. Mode shapes 4.", + "texts": [ + "53 Hz, the mode shape is 2. The free vibrations can be obtained, the modal analysis is performed to find natural frequency In this analysis the angular velocity can be applied 3750 rpm on the drive shaft. The drive shaft can be rotated. The frequency varies 277.19 Hz, the mode shape level is 3. The vibration can be oscillated on the drive shaft. The shaft rotates on its natural frequency, the mode shapes obtained their speeds. In this analysis the mode shape level 4 investigated and is shown in Fig. 16. In this step the natural frequency varies due to rotations. The results are plotted. While analysis the frequency var- ies 399.43 Hz, the mode shape level is 4. The review results are obtained by general post processor. The mode shapes can be determined by natural frequencies in this analysis the frequency varies 490.63 Hz, the mode shape level is 5 and its shown in Fig. 17. Fig. 18 shows that the mode shapes can be changed with regular frequencies. It shows that natural frequency gradually increases from first mode shape to second and continuously increases up to fifth mode shape and vibration of the system eliminated efficiently by suitable element design" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_35_0001419_978-94-007-7401-8-Figure2.72-1.png", + "original_path": "designv11-35/openalex_figure/designv11_35_0001419_978-94-007-7401-8-Figure2.72-1.png", + "caption": "Fig. 2.72 4PaPaC-type fully-parallel PM with coupled Sch\u00f6nflies motions defined by MF \u00bc SF \u00bc 4; RF\u00f0 \u00de \u00bc \u00f0v1; v2; v3;xd\u00de, TF = 0, NF = 30, limb topology Pa\\Pa\\\\C", + "texts": [ + "22y) Idem No. 4 53. 4PaPaRP (Fig. 2.65) Pa\\Pa\\\\R||P (Fig. 2.22z) Idem No. 41 54. 4RPaPaP (Fig. 2.66) R\\Pa\\\\Pa\\\\P (Fig. 2.22a0) Idem No. 15 55. 4RPaPPa (Fig. 2.67) R\\Pa\\kP\\\\Pa (Fig. 2.22b0) Idem No. 15 56. 4RPPaPa (Fig. 2.68) R||P\\Pa\\\\Pa (Fig. 2.22c0) Idem No. 15 57. 4PRPaPa (Fig. 2.69) P||R\\Pa\\\\Pa (Fig. 2.22d0) Idem No. 5 58. 4PaCPa (Fig. 2.70) Pa\\C\\Pa (Fig. 2.22e0) The axes of the cylindrical joints of the four limbs are parallel 59. 4CPaPa (Fig. 2.71) C\\Pa\\\\Pa (Fig. 2.22f 0) Idem No. 58 60. 4PaPaC (Fig. 2.72) Pa\\Pa\\\\C (Fig. 2.22g0) Idem No. 58 176 2 Fully-Parallel Topologies with Coupled Sch\u00f6nflies Motions T ab le 2. 4 L im b to po lo gy an d co nn ec ti ng co nd it io ns of th e fu ll y- pa ra ll el so lu ti on s w it h no id le m ob il it ie s pr es en te d in F ig s. 2. 73 , 2. 74 , 2. 75 , 2. 76 , 2. 77 , 2. 78 ,2 .7 9, 2. 80 ,2 .8 1, 2. 82 ,2 .8 3, 2. 84 ,2 .8 5, 2. 86 ,2 .8 7, 2. 88 ,2 .8 9, 2. 90 ,2 .9 1, 2. 92 ,2 .9 3, 2. 94 ,2 .9 5, 2. 96 ,2 .9 7, 2. 98 ,2 .9 9, 2. 10 0, 2. 10 1, 2. 10 2, 2" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_35_0000504_tmag.2021.3087267-Figure9-1.png", + "original_path": "designv11-35/openalex_figure/designv11_35_0000504_tmag.2021.3087267-Figure9-1.png", + "caption": "Fig. 9. Transient 3D FEA model of the test generator used for validation. Pressure fingers highlighted in blue.", + "texts": [], + "surrounding_texts": [ + "A transient time-step 3D FEA model was set up in Opera [16] to validate the new pressure finger loss calculation. The FEA geometry was modeled after a 1.5 MVA, 16-pole test generator. The test generator shares many geometric endregion features (such as the compression system) with much larger hydro generators. The time-step simulation was run for 3 electric periods with a total of 60 time steps. All results were averaged over the last electrical period. One pressure finger was modeled at an increased mesh density. This pressure finger contains approx. 72500 elements with a maximum element edge length of 2 mm. All other pressure fingers were modeled more coarsely at approx. 10500 elements to reduce the calculation time. The pressure fingers are modeled as nonmagnetic steel with a conductivity of \u03c3 = 1.4 MS/m and a relative permeability of \u00b5r = 1. The clamping plate was modeled as magnetic steel with eddy currents enabled. Eddy currents in the stator core were also enabled in radial and in tangential direction. The pressure fingers were electrically insulated against the stator core and the clamping plate to prevent eddy currents from passing between the components. In reality, these currents exist and might cause additional losses, however modeling these effects exceeds the scope of this paper." + ] + }, + { + "image_filename": "designv11_35_0001419_978-94-007-7401-8-Figure2.19-1.png", + "original_path": "designv11-35/openalex_figure/designv11_35_0001419_978-94-007-7401-8-Figure2.19-1.png", + "caption": "Fig. 2.19 Fully-parallel PMs with coupled Sch\u00f6nflies motions of types 4RRRC (a) and 4RCRR (b) defined by MF \u00bc SF \u00bc 4; RF\u00f0 \u00de \u00bc \u00f0v1; v2; v3;xd\u00de, TF = 0, NF = 2, limb topology R\\R||R||C (a) and R\\C||R||R (b)", + "texts": [ + "15a) R\\P\\kR\\R\\P (Fig. 2.1z) Idem No. 19 28. 4PRRRP (Fig. 2.15b) P\\R||R\\R\\P (Fig. 2.1z1) Idem No. 19 29. 4PRRPR (Fig. 2.16) P\\R\\R\\P\\kR (Fig. 2.1a0) Idem No. 5 30. 4CRRR (Fig. 2.17a) C\\R||R\\R (Fig. 2.1d0) Idem No. 1 (continued) 2.1 Topologies with Simple Limbs 59 Table 2.1 (continued) No. PM type Limb topology Connecting conditions 31. 4CRRR (Fig. 2.17b) C||R||R\\R (Fig. 2.1e0) Idem No. 13 32. 4RRCR (Fig. 2.18a) R||R||C\\R (Fig. 2.1f0) Idem No. 13 33. 4RCRR (Fig. 2.18b) R||C||R\\R (Fig. 2.1g0) Idem No. 13 34. 4RRRC (Fig. 2.19a) R\\R||R||C (Fig. 2.1h0) Idem No. 21 35. 4RCRR (Fig. 2.19b) R\\C||R||R (Fig. 2.1l0) Idem No. 21 36. 4RRCR (Fig. 2.20a) R\\R||C||R (Fig. 2.1b0) Idem No. 21 37. 4RCRR (Fig. 2.20b) R||C\\R||R (Fig. 2.1c0) Idem No. 21 60 2 Fully-Parallel Topologies with Coupled Sch\u00f6nflies Motions In the fully-parallel topologies of PMs with coupled Sch\u00f6nflies motions F / G1G2-G3-G4 presented in this section, the moving platform n : nGi (i = 1, 2, 3, 4) is connected to the reference platform 1 : 1Gi : 0 by four spatial complex limbs with four or five degrees of connectivity. The complex limbs combine only revolute, prismatic and cylindrical joints" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_35_0001419_978-94-007-7401-8-Figure6.33-1.png", + "original_path": "designv11-35/openalex_figure/designv11_35_0001419_978-94-007-7401-8-Figure6.33-1.png", + "caption": "Fig. 6.33 3PPPaR-1RPPaPat-type maximaly regular fully-parallel PMs with Sch\u00f6nflies motions defined by MF = SF = 4, (RF) = (v1; v2; v3;xb), TF = 0, NF = 21, limb topology P\\P\\||Pa||R, R||P||Pa||Pat and P\\P\\||Pa\\||R (a), P\\P\\||Pa ??R (b)", + "texts": [ + "9a) P||Pa||R||R (Fig. 4.8g) R||Pa||Pat||P (Fig. 5.4l) Idem no. 1 9. 3PPPaR-1RPPaPa (Fig. 6.32a) P\\P\\||Pa\\||R (Fig. 4.8b) P\\P\\||Pa||R, (Fig. 4.8c) R||P||Pa||Pa (Fig. 5.4m) The last revolute joints of limbs G1, G2 and G3 and the revolute joint between links 7D and 9 have superposed axes. The actuated prismatic joints of limbs G1, G2 and G3 have orthogonal directions 10. 3PPPaR-1PRPaPa (Fig. 6.32b) P\\P\\||Pa\\R (Fig. 4.8a) P\\P\\||Pa||R, (Fig. 4.8c) R||P||Pa||Pa (Fig. 5.4m) Idem no. 9 11. 3PPPaR-1RPPaPat (Fig. 6.33a) P\\P\\||Pa\\||R (Fig. 4.8b) P\\P\\||Pa||R, (Fig. 4.8c) R||P||Pa||Pat (Fig. 5.4n) Idem no. 9 (continued) 642 6 Maximally Regular Topologies with Sch\u00f6nflies Motions Table 6.13 (continued) No. PM type Limb topology Connecting conditions 12. 3PPPaR-1RPPaPat (Fig. 6.33b) P\\P\\||Pa\\\\R (Fig. 4.8a) P\\P\\||Pa||R, (Fig. 4.8c) R||P||Pa||Pat (Fig. 5.4n) Idem no. 9 13. 3PPaPR-1RPPaPa (Fig. 6.34) P||Pa\\P\\R (Fig. 4.8d) P||Pa\\P\\||R (Fig. 4.8e) R||P||Pa||Pa (Fig. 5.4m) Idem no. 9 14. 3PPaPR-1RPPatPa (Fig. 6.35) P||Pa\\P\\\\R (Fig. 4.8d) P||Pa\\P\\||R (Fig. 4.8e) R||P||Pa||Pat (Fig. 5.4n) Idem no. 9 15. 2PPaRRR-1PPaRR-1RPPaPa (Fig. 6.36) P||Pa||R||R\\R (Fig. 4.9a) P||Pa||R||R (Fig. 4.8g) R||P||Pa||Pa (Fig. 5.4m) Idem no. 9 16. 2PPaRRR-1PPaRR-1RPPaPat (Fig. 6.37) P||Pa||R||R\\R (Fig" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_35_0000261_s12239-021-0050-2-Figure10-1.png", + "original_path": "designv11-35/openalex_figure/designv11_35_0000261_s12239-021-0050-2-Figure10-1.png", + "caption": "Figure 10. Driving force limitation according to acting lateral force.", + "texts": [ + " In the model that calculates the inner wheel\u2019s friction limit in real time, the lateral acceleration sensor signal value is used as the input to calculate the load transfer through which the vertical load of the inner wheel can be estimated in turn (Figure 9). The inner wheel\u2019s friction limit is then calculated as the product of the friction coefficient of the road surface and the inner wheel vertical load. However, according to the tire friction circle concept, the driving force limit can be reduced through the extent of the lateral force even about the same resulting friction limit. Therefore, an allowable driving force prediction model was created according to the lateral force (Figure 10 and the following equations). (6) (7) Hence, the logic is designed for the ELSD clutch to be engaged in proportion to the amount that the real driving force from the powertrain exceeds the allowable driving force at inner wheel. Therefore, ELSD is activated when the engaging value is greater than \u03b5, which means that inner wheel spin occurs because the engine driving force of engine is greater than the allowable driving force (see Figure 11 and the following equations). (8) (9) (10) (11) To calibrate the prediction model, two gains are applied" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_35_0000213_j.triboint.2021.106993-Figure12-1.png", + "original_path": "designv11-35/openalex_figure/designv11_35_0000213_j.triboint.2021.106993-Figure12-1.png", + "caption": "Fig. 12. Vapor Volume Fraction Contour by CFD simulation (10,000 N load, 9000 rpm); (a) Smooth TPJB, (b) Pocketed TPJB.", + "texts": [ + " 11(b), the peak pressures in the pocketed TPJB are produced at the step and pocket trailing edges, which helps stabilize the pad tilting motion. The pressure is suddenly decreased in the circumferential direction where the film thickness is suddenly increased, to conserve fluid momentum. Thus, the pressure suddenly drops at the leading edge of the pockets, generating cavitation in the pocket. The negative gauge pressure inside the pockets is recovered at the trailing edge of the pocket. In addition, Fig. 12(b) shows that the negative pressure inside the pockets causes the significant phase change (high vapor volume fraction) from the liquid phase to the gas phase, while cavitation is not observed in the smooth TPJB result, as seen in Fig. 12 (a). The vaporization is activated more at the higher speeds, as demonstrated in Fig. 13, because the pressure drop inside the pockets becomes larger at the higher speeds. Fig. 13 also shows that the degree of cavitation is nearly insensitive to the applied load. Further simulations have demonstrated that a high level of cavitation occurs for a wide range of operating conditions. This is good since it suggests that the proposed novel features for the TPJB may be beneficial for a wide range of applications" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_35_0000174_j.matpr.2021.01.640-Figure9-1.png", + "original_path": "designv11-35/openalex_figure/designv11_35_0000174_j.matpr.2021.01.640-Figure9-1.png", + "caption": "Fig. 9. Moment applied on the drive shaft.", + "texts": [ + " The fixed support is suitable for the drive shaft .Rotational velocity 650 rad\\s is applied on the drive shaft. The moment is applied 350 N.m on the drive shaft. In this analysis fixed dynamic is suitable are shown in Fig. 7. The rotational velocity 650 rad\\s is applied on the drive shaft. The drive shaft rotations are 650,0, 2.063 rad\\s. The locations are 795.07, 7.3163 mm are shown in Fig. 8. The moment load does not cause rotation, the Moment is applied on the drive shaft is 350 N.m as shown in Fig. 9 in order to omit more bending stress. When composite drive shaft breaks, it is divided into fine fibers that do not have any danger for the driver. They have a less specific modulus and less damping capacity. So that conventional drive shaft is replaced with composite materials Fig. 10. Selection of materials are an important parameter for design of any machine element. Our objective is to select suitable material than conventional steel material and preferred metal matrix composite materials consist of Aluminium, Titanium, Vanadium alloy", + " The accompanying suppositions were prepared in our estimations The shaft whirling at a reliable torque with respect to longitudinal alignment. The shaft have an indistinguishable and in circular crosssection. The shaft is superbly adjusted, at each cross area, the mass focus concurs among the geometric focus. All clammy and nonlinear impacts have prohibited The stress - strain correlation is linear and flexible for composite material and its obey Hook\u2019s law. Since lamina is thin and no out-of - plane burdens are associated and considered under plane stress conditions, its details are specified in Fig. 9. The drive shaft can be solid, round section chosen. Material selection (Aluminium, Titanium, Vanadium alloy) The orthotropic properties of metal matrix composites (Al-Ti-V) are listed below in table 3. Selection of materials are Aluminium, Titanium, Vanadium alloy. Ti-Al-V alloy containing 82% Titanium, 6% Aluminium, 12% Vanadium. Here Aluminium and Titanium are matrix, Vanadium is the fiber respectively. This Titanium alloy has good tensile properties. The metal matrix composites are high strength to density ratio and better fatigue resistance Table 4" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_35_0000115_iros45743.2020.9340920-Figure4-1.png", + "original_path": "designv11-35/openalex_figure/designv11_35_0000115_iros45743.2020.9340920-Figure4-1.png", + "caption": "Fig. 4. WalkON Suit, the powered exoskeleton for the paraplegics; (a) the WalkON Suit and (b) the mechanical components of the lower limb", + "texts": [ + " The iterative learning controller in the learning domain generates an updated body inclination array by the following control law: Vi+1 = Vi + krP\u22121 learn[R\u2212Ti] (14) where R represents the desired output array, which is exactly 50% of the gait cycle and kr represents the update gain. As the number of iteration of learning domain goes to infinity, the ground contact error converges to zero since the error array is expressed as, Ei+1 = R\u2212Ti+1 = (1\u2212 kr)Ei. (15) Hence the error array is ensured to be zero as the iteration goes by if, 0 < kr < 2. (16) In order to verify the performance of the proposed gait pattern, an experimental setup shown in Fig. 4, the WalkON Suit manufactured by ANGEL ROBOTICS CO. is utilized. Figure 4(a) represents the overall structure of the WalkON Suit, and Fig. 4(b) represents the detailed mechanical components. Note that the WalkOn Suit has active joints for the hip, knee and ankle. The hip and knee joints are actuated by the brushless motors (MF0127020 manufactured by Alliedmotion Co.) equipped with a set of planetary gears. The overall gear ratio is about 20:1, and the maximum joint torque for each joint is 70 N\u00b7m. The ankle joint is controlled by a linear actuator of which the thrust force is 1500 N. Before the gait adaptation by iterative learning is implemented, the problem of ground contact error appears as shown in Fig" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_35_0001563_9781119682684.ch1-Figure1.11-1.png", + "original_path": "designv11-35/openalex_figure/designv11_35_0001563_9781119682684.ch1-Figure1.11-1.png", + "caption": "Figure 1.11 Magnetic coupling between two conducting loops.", + "texts": [ + " \u222b xmax 0 dx = \u222b \u221e 0 ve\u2212 B2 l2 mR tdt [x]xmax 0 = v\u222b \u221e 0 e\u2212 B2 l2 mR tdt = v \u23a1\u23a2\u23a2\u23a2\u23a3 1( \u2212 B2l2 mR ) e\u2212 B2 l2 mR t \u23a4\u23a5\u23a5\u23a5\u23a6 \u221e 0 Or xmax \u2212 0 = v ( \u2212 mR B2l2 ) ( e\u2212\u221e \u2212 e0) Or xmax = ( \u2212vmR B2l2 ) (0 \u2212 1) = vmR B2l2 Alternative Method: applying Newton\u2019s second law \u2212B2l2v R = ( \u2212B2l2 R ) dx dt = m dv dt or ( \u2212B2l2 R ) dx = mdv Integrating both sides from initial to the stopping point gives \u222b xmax 0 ( \u2212B2l2 R ) dx = \u222b 0 v mdv Or ( \u2212B2l2 R ) [x]xmax 0 = m [v]0 v Or ( \u2212B2l2 R )( xmax \u2212 0 ) = m (0 \u2212 v) Therefore xmax = vmR B2l2 Flux linkage is an extension of magnetic flux. Flux linkage is equivalent to the total flux passing through the surface formed by a coil and it is determined by knowing the value of flux (\ud835\udf19) and the number of turns as the following relation expresses. \ud835\udf06 = \u222b B \u22c5 dS = \u222b dt = N\ud835\udf19 (1.44) The closed loops C1 and C2 are placed near to each other whose surface area S1 and S2 respectively. C1 carries current i1 as shown in Figure 1.11. C1 produces flux B1 because of the flow of current i1. Part of the total flux passes through the surface S2 bounded by C2. Let that flux be \u03d512. The magnetic flux \u03d512 is produced because of current i1 flowing in C1 is given by \ud835\udf1912 = \u222b S2 B \u22c5 dS (1.45) Consider Figure 1.12 which has two turns. The magnetic flux \u03d512 links surface S21 and S22. The total surface area will be S2 = S21 + S22. We see that the same magnetic flux passes through S2 twice in the same direction. So, the magnetic flux linking C2 must be 2\u03d512" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_35_0000128_j.cryogenics.2021.103283-Figure6-1.png", + "original_path": "designv11-35/openalex_figure/designv11_35_0000128_j.cryogenics.2021.103283-Figure6-1.png", + "caption": "Fig. 6. (a) The location of eight PT-100 platinum temperature sensors; (b)\u2013(d) The cross section of the stator core at three different positions.", + "texts": [ + " The vacuum degree was measured by the Pirani gauge with a measuring range of 105\u201310\u2212 1 Pa. The measuring accuracy of the Pirani gauge was \u00b110%. The rotation speed of the motor was measured by a Hall sensor with a measurement accuracy of \u00b10.5%. The temperature of the stator core directly affected the overcurrent capacity of the copper coil and the performance of the motor. To monitor the working temperature of the stator core, eight PT-100 platinum temperature sensors were placed at different positions of the stator core, as shown in Fig. 6. There were three sensors located near the inlet of liquid nitrogen marked X-1, 2 and 3 and three sensors located in the middle of the core marked X-4, 5, and 6. X-7 and 8 were placed near the outlet of the liquid nitrogen. Due to the space limitation, no sensor was placed in the bottom of the core. These sensors were glued to the surface of the stator core with STYCAST 2850 FT. The resistance values of these sensors were measured with a multimeter, and then converted into temperature. Three-wire connection method was used by the PT-100 sensors" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_35_0001419_978-94-007-7401-8-Figure3.141-1.png", + "original_path": "designv11-35/openalex_figure/designv11_35_0001419_978-94-007-7401-8-Figure3.141-1.png", + "caption": "Fig. 3.141 Overactuated PMs with coupled Sch\u00f6nflies motions of types 4PaCRR (a) and 4PaRCR (b) defined by MF \u00bc SF \u00bc 4; RF\u00f0 \u00de \u00bc v1; v2; v3;xd\u00f0 \u00de, TF = 0, NF = 14, limb topology Pa\\C||R\\kR (a) and Pa\\R||C\\kR (b)", + "texts": [ + " 4PaRPRR (Fig. 3.133) Pa\\R||P||R\\kR (Fig. 3.52a0) Idem No. 13 48. 4PaRRPR (Fig. 3.134) Pa\\R||R||P\\kR (Fig. 3.52b0) Idem No. 13 49. 4RCRPa (Fig. 3.135) R\\C||R\\kPa (Fig. 3.52c0) Idem No. 12 50. 4RRCPa (Fig. 3.136) R\\R||C\\kPa (Fig. 3.52d0) Idem No. 12 51. 4CRRPa (Fig. 3.137) C||R\\R||Pa (Fig. 3.52e0) Idem No. 15 52. 4RCRPa (Fig. 3.138) R||C\\R||Pa (Fig. 3.52f0) Idem No. 15 53. 4CRPaR (Fig. 3.139) C||R\\Pa||R (Fig. 3.52g0) Idem No. 13 54. 4RCPaR (Fig. 3.140) R||C\\Pa||R (Fig. 3.52h0) Idem No. 13 55. 4PaCRR (Fig. 3.141a) Pa\\C||R\\kR (Fig. 3.52i0) Idem No. 13 56. 4PaRCR (Fig. 3.141b) Pa\\R||C\\kR (Fig. 3.52j0) Idem No. 13 57. 4PaRRRPa (Fig. 3.142) Pa\\R||R\\kR||Pa (Fig. 3.53a) Idem No. 15 58. 4RPaRRPa (Fig. 3.143) R\\Pa\\kR\\kR||Pa (Fig. 3.53b) Idem No. 15 59. 4RRRPaPa (Fig. 3.144) R||R\\R||Pa||Pa (Fig. 3.53c) Idem No. 15 60. 4RPaRPaR (Fig. 3.145) R\\Pa\\kR\\kPa||R(Fig. 3.53d) Idem No. 13 61. 4PaPaRRR (Fig. 3.146) Pa||Pa\\R||R\\kR (Fig. 3.53e) Idem No. 13 62. 4PaPaRRR (Fig. 3.147) Pa||Pa\\R||R\\kR (Fig. 3.53f) Idem No. 13 Table 3.7 Structural parametersa of parallel mechanisms in Figs. 3" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_35_0001419_978-94-007-7401-8-Figure6.27-1.png", + "original_path": "designv11-35/openalex_figure/designv11_35_0001419_978-94-007-7401-8-Figure6.27-1.png", + "caption": "Fig. 6.27 3PPPaR-1RPaPaP-type maximaly regular fully-parallel PMs with Sch\u00f6nflies motions defined by MF = SF = 4, (RF) = (v1; v2; v3;xb), TF = 0, NF = 21, limb topology P\\P\\||Pa||R, R||Pa||Pa||P and P\\P\\||Pa\\||R (a), P\\P\\||Pa ??R (b)", + "texts": [ + "3 Fully-Parallel Topologies with Complex Limbs 637 638 6 Maximally Regular Topologies with Sch\u00f6nflies Motions 6.3 Fully-Parallel Topologies with Complex Limbs 639 640 6 Maximally Regular Topologies with Sch\u00f6nflies Motions 6.3 Fully-Parallel Topologies with Complex Limbs 641 Table 6.13 Limb topology and connecting conditions of the fully-parallel solutions with no idle mobilities presented in Figs. 6.27, 6.28, 6.29, 6.30, 6.31, 6.32, 6.33, 6.34, 6.35, 6.36, 6.37, 6.38, 6.39, 6.40, 6.41 No. PM type Limb topology Connecting conditions 1. 3PPPaR-1RPaPaP (Fig. 6.27a) P\\P\\||Pa\\||R (Fig. 4.8b) P\\P\\||Pa||R, (Fig. 4.8c) R||Pa||Pa||P (Fig. 5.4k) The last joints of the four limbs have superposed axes/directions. The actuated prismatic joints of limbs G1, G2 and G3 have orthogonal directions 2. 3PPPaR-1RPaPaP (Fig. 6.27b) P\\P\\||Pa\\R (Fig. 4.8a) P\\P\\||Pa||R, (Fig. 4.8c) R||Pa||Pa||P (Fig. 5.4k) Idem no. 1 3. 3PPPaR-1RPaPatP (Fig. 6.28a) P\\P\\||Pa\\||R (Fig. 4.8b) P\\P\\||Pa||R, (Fig. 4.8c) R||Pa||Pat||P (Fig. 5.4l) Idem no. 1 4. 3PPPaR-1RPaPatP (Fig. 6.28b) P\\P\\||Pa\\R (Fig. 4.8a) P\\P\\||Pa||R, (Fig. 4.8c) R||Pa||Pat||P (Fig. 5.4l) Idem no. 1 5. 3PPaPR-1RPaPaP (Fig. 6.29a) P||Pa\\P\\\\R (Fig. 4.8d) P||Pa\\P\\||R (Fig. 4.8e) R||Pa||Pa||P (Fig. 5.4k) Idem no. 1 6. 3PPaPR-1RPaPatP (Fig. 6.29b) P||Pa\\P\\\\R (Fig. 4.8d) P||Pa\\P\\||R (Fig" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_35_0000246_tte.2021.3068819-Figure11-1.png", + "original_path": "designv11-35/openalex_figure/designv11_35_0000246_tte.2021.3068819-Figure11-1.png", + "caption": "Fig. 11. 3D mesh model of different transposition angles.", + "texts": [ + " In order to optimize the axial distribution of the conductor furtherly, the effect of the \u03be2 (the equivalent length coefficient) on the AC loss at different \u03b8 is analyzed in the Fig. 9(b). For the different \u03b8, AC loss will be reduced to varying degrees with different \u03be2. Besides the AC loss, the current density of each strand will be affected by the frequency, including the phase and the magnitude, see the Fig. 10(a)-(c). And the specific transposition angle can effectively reduce the magnitude, see the Fig. 10(d)-(i). Fig. 11 shows the mesh models of 3D-Model at two transposition angles and Fig. 12 shows the mesh model of EMM. When there is no transposition angle (\u03b8=0\u00b0), the accuracy of EMM with n=1 is high enough. However, for transposition angle, n=2 or 3 will be chosen and the number of mesh nodes will also increase. For 3D-Model and 2D-EMM, the number of mesh nodes has been compared in Table II. From the comparison result, it can be seen that the number of mesh nodes in 3D-Model is more than that in EMM. For different transposition angle, the ratio of Mesh-nodes-number is 10" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_35_0001393_ijhm.2020.109916-Figure1-1.png", + "original_path": "designv11-35/openalex_figure/designv11_35_0001393_ijhm.2020.109916-Figure1-1.png", + "caption": "Figure 1 Dimensions of spur gear (see online version for colours)", + "texts": [], + "surrounding_texts": [ + "The geometrical model of spur gear was constructed using SOLIDWORKS software for the designing process (Figures 1 and 2). After creating the model in SOLIDWORKS software, the design is imported into ANSYS workbench software for further analysis of the key factors such as stresses, deflections and modal analysis of the spur gear. The dimensions and the material properties of the spur gear are given Tables 1 and 2. Figure 3 shows the meshed view of the spur gear. Accuracy and efficiency of finite element method depends upon the meshing size of the model. A very fine mesh is used to design the model so that the results would be more accurate." + ] + }, + { + "image_filename": "designv11_35_0000686_j.mechmachtheory.2021.104433-Figure27-1.png", + "original_path": "designv11-35/openalex_figure/designv11_35_0000686_j.mechmachtheory.2021.104433-Figure27-1.png", + "caption": "Fig. 27. Metamorphic epicyclic bevel gear clamping mechanism.", + "texts": [ + "3, a metamorphic epicyclic bevel gear clamping mechanism is designed by using the epicyclic bevel gear train to replace the equivalent mechanism. According to the equivalent geometrical relationship, the key condition, \u03c9A = \u03c9B, of equivalent metamorphic clamping mechanism means that the angular velocity of sun bevel gear is equal to the angular velocity of the bevel gear arm. To realize and adjust this condition, a structure design scheme using combination constraint (inspired from Fig. 22) has been completed as shown in Fig. 27(a). The structure of the metamorphic epicyclic bevel gear clamping mechanism can be described as follow: the sun bevel gear, the driver, drives the planet bevel gear. The lead screw nut is fixed on the planet bevel gear and drives the lead screw claw to move. A limit spring connects the lead screw claw and the lead screw nut. The bevel gear arm holds the bevel gears and limits the rotation of the lead screw claw. In its rotation configuration as shown in Fig. 27(b), the sun bevel gear driven by the motor rotates the planet bevel gear. Due to limitation by the preload force of the limit spring, the lead screw claw and the lead screw nut are relatively static, the lead screw nut H. Yuan et al. Mechanism and Machine Theory 166 (2021) 104433 and the planet bevel gear stop rotating about their rotation axis. Under this constraint condition, the metamorphic epicyclic bevel gear train is in the rotation configuration, the angular velocities of the sun bevel gear and bevel gear arm are the same, and the lead screw claw rotates about the rotation axis of the sun bevel gear towards the target object. The rotation configuration transforms to translation configuration as shown in Fig. 27(c) when the lead screw claw contacts the target object. In this configuration, the driving force provided by the sun bevel gear overcomes the spring force, compresses the limit spring, and drives the lead screw claw to move. Meanwhile, the lead screw claw is constrained by the target object and moves along the lead screw to clamp the target object. To demonstrate the process of configuration transformation better, a metamorphic epicyclic bevel gear clamping mechanism is designed and the corresponding mechanical parts based on the principle model in Fig. 27 are shown in Fig. 28(a). An insulator inspection robot that consists of two metamorphic epicyclic bevel gear clamping mechanisms and a control box is designed as an experimental prototype. The metamorphic epicyclic bevel gear clamping mechanism is used to clamp the insulator to fix the robot. The prototype experiments focused on the configuration transformation are carried out by clamping the insulator as displayed in Fig. 28(b)-(e). When the robot starts to clamp the insulator, the lead screw claw rotates to the insulator by the driving motor and doesn\u2019t move along the lead screw" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_35_0001419_978-94-007-7401-8-Figure6.46-1.png", + "original_path": "designv11-35/openalex_figure/designv11_35_0001419_978-94-007-7401-8-Figure6.46-1.png", + "caption": "Fig. 6.46 2PPaRRR-1PPaRR-1CPaPat-type maximaly regular fully-parallel PM with Sch\u00f6nflies motions defined by MF = SF = 4, (RF) = (v1; v2; v3;xb), TF = 0, NF = 19, limb topology P||Pa||R||R\\R, P||Pa||R||R and C||Pa||Pat", + "texts": [ + "8c) C||Pa||Pat (Fig. 5.4p) Idem no. 1 4. 3PPPaR-1CPaPat (Fig. 6.43b) P\\P\\||Pa\\R (Fig. 4.8a) P\\P\\||Pa||R, (Fig. 4.8c) C||Pa||Pat (Fig. 5.4p) Idem no. 1 5. 3PPaPR-1CPaPa (Fig. 6.44a) P||Pa\\P\\\\R (Fig. 4.8d) P||Pa\\P\\||R (Fig. 4.8e) C||Pa||Pa (Fig. 5.4o) Idem no. 1 6. 3PPaPR-1CPaPat (Fig. 6.44b) P||Pa\\P\\\\R (Fig. 4.8d) P||Pa\\P\\||R (Fig. 4.8e) C||Pa||Pat (Fig. 5.4p) Idem no. 1 7. 2PPaRRR-1PPaRR1CPaPa (Fig. 6.45) P||Pa||R||R\\R (Fig. 4.9a) P||Pa||R||R (Fig. 4.8g) C||Pa||Pa (Fig. 5.4o) Idem no. 1 8. 2PPaRRR-1PPaRR1CPaPat (Fig. 6.46) P||Pa||R||R\\R (Fig. 4.9a) P||Pa||R||R (Fig. 4.8g) C||Pa||Pat (Fig. 5.4p) Idem no. 1 Table 6.15 Structural parametersa of parallel mechanisms in Figs. 6.27, 6.28, 6.29, 6.30, 6.31, 6.32, 6.33, 6.34, 6.35, 6.36, 6.37 No. Structural parameter Solution Figures 6.27, 6.28, 6.29, 6.32, 6.33, 6.34, 6.35 Figures 6.30, 6.31, 6.36, 6.37 1. m 24 26 2. pi (i = 1, 3) 7 8 3. p2 7 7 4. p4 10 10 5. p 31 33 6. q 8 8 7. k1 0 0 8. k2 4 4 9. k 4 4 10. (RG1) (v1; v2; v3;xb) (v1; v2; v3;xa;xb) 11. (RG2) (v1; v2; v3;xb) (v1; v2; v3;xb) 12" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_35_0000512_s42417-021-00289-8-Figure6-1.png", + "original_path": "designv11-35/openalex_figure/designv11_35_0000512_s42417-021-00289-8-Figure6-1.png", + "caption": "Fig. 6 Timoshenko beam element", + "texts": [ + " Under normal conditions, the mass of ring is much smaller than that of rotor; as a result, the motion of inner ring is determined by the journal position of rotor which is controlled by rotor dynamic equations; at the same time, the force and torque from other bearing parts (balls and cage) acting on the inner ring are transferred to the corresponding journal position to provide supporting force and moment which are determined by the dynamic equations of ball bearings. According to the above coordination relationship, the following motion constraint equation can be established after transforming the DOFs of rotor in the bearing coordinate system. (33) ( \ud835\udc0cS +\ud835\udc0cD ) ?\u0308?R + ( \ud835\udc06S +\ud835\udc06D ) ?\u0307?R +\ud835\udc0as\ud835\udc2eR = \ud835\udc05B + \ud835\udc05U + \ud835\udc05G, The rotor in Fig.\u00a01 is composed of a shaft and a disk. The shaft is discretized by several Timoshenko beam element (Fig.\u00a06) which considers the moment of inertia, shear effect and axial force effect [21]. This element includes 3 translation DOFs (x, y, z) and 2 rotational DOFs (\u03c8, \u03c6). Each element matrix is assembled according to the relative geometric position to form an integral matrix of the shaft. (34) \u23a7\u23aa\u23a8\u23aa\u23a9 Bi = (xBi, yBi, zBi, x\u0307Bi, y\u0307Bi, z\u0307Bi, \ud835\udf02Bi, \ud835\udf09Bi, \ud835\udf06Bi, ?\u0307?Bi, ?\u0307?Bi, ?\u0307?Bi) T Rj = (zRj,\u2212xRj,\u2212yRj, z\u0307Rj,\u2212x\u0307Rj,\u2212y\u0307Rj,\ud835\udf14Rt,\u2212\ud835\udf13Rj,\u2212\ud835\udf11Rj,\ud835\udf14R,\u2212?\u0307?Rj,\u2212?\u0307?Rj) T Bi = Rj (i = 1, 2; j = 1,\u2026 , ns) , where uBi is the combined translation and rotation vector of ith bearing; uRj is the combined translation and rotation vector of jth node on the rotor corresponding to the position of ith bearing; \u03c9R is the angular velocity of rotor; t is time" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_35_0001419_978-94-007-7401-8-Figure3.54-1.png", + "original_path": "designv11-35/openalex_figure/designv11_35_0001419_978-94-007-7401-8-Figure3.54-1.png", + "caption": "Fig. 3.54 Overactuated PMs with coupled Sch\u00f6nflies motions of types 4PRPPa (a) and 4PRPaP (b) defined by MF \u00bc SF \u00bc 4; RF\u00f0 \u00de \u00bc v1; v2; v3;xd\u00f0 \u00de, TF = 0, NF = 18, limb topology P||R\\P\\kPa (a) and P||R||Pa\\P (b)", + "texts": [ + "2 Topologies with Complex Limbs 355 356 3 Overactuated Topologies with Coupled Sch\u00f6nflies Motions 3.2 Topologies with Complex Limbs 357 358 3 Overactuated Topologies with Coupled Sch\u00f6nflies Motions Table 3.5 Limb topology and connecting conditions of the overactuated solutions with no idle mobilities presented in Figs. 3.54, 3.55, 3.56, 3.57, 3.58, 3.59, 3.60, 3.61, 3.62, 3.63, 3.64, 3.65, 3.66, 3.67, 3.68, 3.69, 3.70, 3.71, 3.72, 3.73, 3.74, 3.75, 3.76, 3.77, 3.78, 3.79, 3.80, 3.81, 3.82, 3.83, 3.84, 3.85, 3.86, 3.87, 3.88 No. PM type Limb topology Connecting conditions 1. 4PRPPa (Fig. 3.54a) P||R\\P\\kPa (Fig. 3.50a) The revolute joints of the four limbs have parallel axes 2. 4PRPaP (Fig. 3.54b) P||R||Pa\\P (Fig. 3.50b) Idem No. 1 3. 4PPaRP (Fig. 3.55a) P||Pa||R\\P (Fig. 3.50c) Idem No. 1 4. 4PRRPa (Fig. 3.55b) P||R||R||Pa (Fig. 3.50d) Idem No. 1 5. 4PPRPa (Fig. 3.56a) P\\P||R||Pa (Fig. 3.50e) Idem No. 1 6. 4PPRPa (Fig. 3.56b) P\\P\\kR||Pa (Fig. 3.50f) Idem No. 1 7. 4PRPPa (Fig. 3.57a) P\\R||P||Pa (Fig. 3.50g) Idem No. 1 8. 4PRPaP (Fig. 3.57b) P\\R||Pa||P (Fig. 3.50h) Idem No. 1 9. 4RPRPa (Fig. 3.58a) R||P||R||Pa (Fig. 3.50i) Idem No. 1 10. 4RRPPa (Fig. 3.58b) R||R||P||Pa (Fig. 3.50j) Idem No" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_35_0001419_978-94-007-7401-8-Figure5.10-1.png", + "original_path": "designv11-35/openalex_figure/designv11_35_0001419_978-94-007-7401-8-Figure5.10-1.png", + "caption": "Fig. 5.10 2PaRRRR-1PaRRR-1RPPP-type fully-parallel PMs with uncoupled Sch\u00f6nflies motions defined by MF = SF = 4, RF\u00f0 \u00de \u00bc \u00f0v1; v2; v3;xb\u00de, TF = 0, NF = 13, limb topology Pa\\R||R||R\\kR, Pa\\R||R||R and R\\P\\\\P\\\\P", + "texts": [], + "surrounding_texts": [ + "5.1 Fully-Parallel Topologies 443", + "444 5 Topologies with Uncoupled Sch\u00f6nflies Motions", + "5.1 Fully-Parallel Topologies 445" + ] + }, + { + "image_filename": "designv11_35_0001419_978-94-007-7401-8-Figure6.11-1.png", + "original_path": "designv11-35/openalex_figure/designv11_35_0001419_978-94-007-7401-8-Figure6.11-1.png", + "caption": "Fig. 6.11 Maximaly regular fully-parallel PMs with Sch\u00f6nflies motions of types 3PPPR1RPPaPa (a) and 3PPPR-1RPPaPat (b) defined by MF = SF = 4, (RF) = (v1; v2; v3;xb), TF = 0, NF = 12, limb topology P\\P ??P\\||R, P\\P ??P ??R and R||P||Pa||Pa (a), R||P||Pa||Pat (b)", + "texts": [ + " 2PRPRR-1PRPR-1RPaPaP (Fig. 6.9a) P||R\\P\\||R\\R (Fig. 4.1b) P||R\\P\\||R (Fig. 4.1e) R||Pa||Pa||P (Fig. 5.4k) Idem no. 1 6. 2PRPRR-1PRPR1RPaPatP (Fig. 6.9b) P||R\\P\\||R\\R (Fig. 4.1b) P||R\\P\\||R (Fig. 4.1e) R||Pa||Pat||P (Fig. 5.4l) Idem no. 1 7. 2PRRRR-1PRRR-1RPaPaP (Fig. 6.10a) P||R||R||R\\R (Fig. 4.2d) P||R||R||R (Fig. 4.1g) R||Pa||Pa||P (Fig. 5.4k) Idem no. 1 8. 2PRRRR-1PRRR1RPaPatP (Fig. 6.10b) P||R||R||R\\R (Fig. 4.2d) P||R||R||R (Fig. 4.1g) R||Pa||Pat||P (Fig. 5.4l) Idem no. 1 9. 3PPPR-1RPPaPa (Fig. 6.11a) P\\P ??P\\||R (Fig. 4.1a) P\\P ??P ??R (Fig. 4.1b) R||P||Pa||Pa (Fig. 5.4m) The last revolute joints of limbs G1, G2 and G3 and the revolute joint between links 7D and 9 have superposed axes. The actuated prismatic joints of limbs G1, G2 and G3 have orthogonal directions 10. 3PPPR-1RPPaPat (Fig. 6.11b) P\\P ??P\\||R (Fig. 4.1a) P\\P ??P ??R (Fig. 4.1b) R||P||Pa||Pat (Fig. 5.4n) Idem no. 9 11. 2PPRRR-1PPRR-1RPPaPa (Fig. 6.12a) P\\P\\||R||R\\R (Fig. 4.2a) P\\P\\||R||R (Fig. 4.1d) R||P||Pa||Pa (Fig. 5.4m) Idem no. 9 (continued) 612 6 Maximally Regular Topologies with Sch\u00f6nflies Motions Table 6.5 (continued) No. PM type Limb topology Connecting conditions 12. 2PPRRR-1PPRR-1RPPaPat (Fig. 6.12b) P\\P\\||R||R\\R (Fig. 4.2a) P\\P\\||R||R (Fig. 4.1d) R||P||Pa||Pat (Fig. 5.4n) Idem no. 9 13. 2PRPRR-1PRPR-1RPPaPa (Fig" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_35_0001579_2015-24-2526-Figure3-1.png", + "original_path": "designv11-35/openalex_figure/designv11_35_0001579_2015-24-2526-Figure3-1.png", + "caption": "Figure 3. MBD engine model with balancer gear drive", + "texts": [ + " This method is widely used in MBD simulations based on flexible multi body approach [4] and can consider dynamic behavior of complex structure in relevant frequency range [5] depending about FE-model quality and considered vibration phenomena. Typically, in today's standard engine development process crankshaft dynamic and strength analysis is previously performed using flexible multi-body model. This model is extended with the MBU for purpose of detail investigations within early development phase. The model includes fully flexible crank train and balancer drive and main structural parts of engine block and bottom end (Figure 3). FEM models used in MBD analysis are shown as schematic wireframe figures. As mentioned for the previous TVA analysis, the structured model could also be used within the 3D MBD analysis. In the specific case a 3D FEM model is preferred as it provides more detailed geometry information for the MBU, especially the gears and their connection to the crankshaft. The flexible multi-body simulation are done with software AVL EXCITE PowerUnit (see [6], [7]), which considers a set of flexible or rigid structures (called as bodies), which are coupled via contact models (named as joints)" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_35_0000636_j.mechmachtheory.2021.104432-Figure1-1.png", + "original_path": "designv11-35/openalex_figure/designv11_35_0000636_j.mechmachtheory.2021.104432-Figure1-1.png", + "caption": "Fig. 1. Package and drivetrain topology of BEV equipped with a twist-beam rear axle. (a) Schematic visualisation of the package space in the underbody of a BEV with a front central drive and a twist-beam rear axle. (b) Design space within the underbody of a B-segment vehicle using a conventional twist-beam and a reversed twist-beam installation.", + "texts": [ + " Although this drivetrain topology is the most conventional compared to vehicles with an ICE, a wide variety of different suspension systems is used throughout the compact and subcompact vehicle segments. A topology that is used in the A- (mini cars), B- (small cars) and C-segments (medium cars) combines a central drive with a MacPherson strut at the front axle and utilises a twist-beam axle (TBA) at the rear. The battery pack is located in the underbody and limited by the front subframe and the cross beam at the rear (Fig. 1(a)). This type of architecture can be found in numerous industrial applications, such as [7] (A-Segment), [8] (B-Segment) and also in research projects like [9]. Another example in which a semi-rigid rear axle is used in combination with a front central drive can be found in [10]. Here, a vehicle concept is shown that results from the * Corresponding authors. E-mail addresses: Tobias.Niessing@uni-siegen.de (T. Niessing), Xiangfan.Fang@uni-siegen.de (X. Fang). Contents lists available at ScienceDirect Mechanism and Machine Theory journal homepage: www", + " In a BEV, however, these components are no longer required, which is why this reversed installation of the TBA enables an e lr Unit vector from the left-to-right body mount e RU,r Unit vector of the rotational axis of right revolute joint RU etotal Total error of the objective function fi DOFs of joint i f(P(t)) Objective function g Number of overall joints hcg Centre-of-gravity height i Index of an arbitrary joint/ hardpoint k Number of knuckles l Number of links lWB Wheelbase P(tj) Matrix including all current hardpoints P(t0) Matrix including all hardpoints at design position Pinitial(t0) Initial set of hardpoints at design position Pmin, local(t0) Hardpoint set at design position leading to a local minimum P i(tj) Current position of hardpoint P i P i(tj+1) Upcoming position of hardpoint P i r Number of redundant DOFs r lr Connection vector from left to right body mount r RURL Connection vector from RUto RL r RUWC Connection vector from RUto WC t0 Initial time step (design position) tj Current time step tj+1 Upcoming time step TIR Transformation matrix from reference frame R to inertial frame I v K,r Translational velocity of the right knuckle evaluated at WC v i(tj,P(tj)) Velocity of P i as function of the current hardpoints v RL,r Velocity at the right spherical joint RL v RU,r Velocity at the right revolute joint RU v Ur Velocity of the isolated right body mount v Ur, \u2016 Velocity of the isolated right body mount projected onto translational DOF v Ur, \u22a5 Velocity of the isolated right body mount perpendicular to the translational DOF wIC Weighting factor for error function eIC zRU Distance of bushings at RU to the ground T. Niessing and X. Fang Mechanism and Machine Theory 165 (2021) 104432 increased, coherent and regularly shaped design space in the vehicle underbody. This design space can be used directly to package a larger drive battery. Fig. 1(b) shows the battery design space in the underbody of a typical B-segment vehicle with a conventional and a reversed installation of a TBA. In addition to the package advantage, the trailing arm and the cross member surround the rear end of the battery design space and can therefore be integrated into the safety concept in the event of a rear or side crash. Since the body mounts are located behind the wheel patch, structural toe-in behaviour under lateral forces can also be achieved [18]. However, reversing the twist-beam results in a negative anti-lift behaviour while braking", + " For this purpose, an additional longitudinal link and flexible joints to the knuckle are introduced. With these, the reversed TBA is integrated into a longitudinal Watt\u2019s linkage (see Fig. 2(a) and (b)). The trailing links and arms are designed to create an IC that is located in front of the wheels. The projections of these trailing links and arms intersect at the new centre of motion. Additionally, this virtual IC should be located above the WC to ensure a desirable positive WC recession during jounce. Fig. 1(b) shows the increased and connected package space in the underbody that can be provided by the use of a reversed TBA or the MLTA, respectively. In this work, a Ford Fiesta as a typical B-segment vehicle was chosen as a reference. The available volume for the battery can be increased by up to 30% compared to the available package space with a common TBA. Additionally, it can be seen that the cross beam and trailing arms of the suspension surround the sides and rear end of the battery. As already stated in [18], the cross beam can be designed to act like an additional load path for the side impact by partly transferring forces from the impact side to the opposite side of the vehicle", + " However, it should be noted that in the case of amplified pitching, an increased rebound motion is expected, which indicates that the MLTA offers a self-stabilising effect. The moving IC also leads to a different behaviour of some suspension characteristics for the different loading states, especially compared to a TBA with a body-fixed IC. A comparison of selective ideal kinematic suspension characteristics is shown in Fig. 10. The loading states are given with respect to the masses of the representative BEV shown in Fig. 1(b). Also, the given values are normalised by the results obtained at the design loading (two passengers in the front and one in the rear) for the conventional TBA. As already T. Niessing and X. Fang Mechanism and Machine Theory 165 (2021) 104432 shown in Fig. 8(b), the MLTA can achieve a positive WC recession in contrast to the TBA. This allows more options regarding the longitudinal elasticity of the body bushing at U and L to improve driving comfort (Fig. 10(a)). While the roll steer gradient of the TBA increases with a growing mass, the corresponding value of the MLTA decreases (Fig" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_35_0001419_978-94-007-7401-8-Figure5.40-1.png", + "original_path": "designv11-35/openalex_figure/designv11_35_0001419_978-94-007-7401-8-Figure5.40-1.png", + "caption": "Fig. 5.40 3PaPPaR-1RPPaPa-type fully-parallel PMs with uncoupled Sch\u00f6nflies motions defined by MF = SF = 4, RF\u00f0 \u00de \u00bc \u00f0v1; v2; v3;xb\u00de, TF = 0, NF = 30 limb topology R||P||Pa||Pa and Pa||P\\Pa\\kR, Pa||P\\Pa||R (a), Pa\\P\\\\Pa\\kR, Pa\\P\\\\Pa||R (b)", + "texts": [ + "44, 5.45, 5.46, 5.47, 5.48, 5.49, 5.50, 5.51, 5.52, 5.53, 5.54, 5.55, 5.56, 5.57, 5.58, 5.59, 5.60, 5.61, 5.62, 5.63, 5.64, 5.65, 5.66, 5.67, 5.68 No. PM type Limb topology Connecting conditions 1. 3PaPPaR-1RPaPaP (Fig. 5.39a) Pa||P\\Pa\\kR (Fig. 5.4a) The last joints of the four limbs have superposed axes/directions Pa||P\\Pa||R (Fig. 5.4d) R||Pa||Pa||P (Fig. 5.4k) 2. 3PaPPaR-1RPaPaP (Fig. 5.39b) Pa\\P\\\\Pa\\kR (Fig. 5.4b) Idem No. 1 Pa\\P\\\\Pa||R (Fig. 5.4c) R||Pa||Pa||P (Fig. 5.4k) 3. 3PaPPaR-1RPPaPa (Fig. 5.40a) Pa||P\\Pa\\kR (Fig. 5.4a) The revolute joint between links 7D and 9 of G4 limb and the last revolute joints of limbs G1, G2 and G3 have superposed axes Pa||P\\Pa||R (Fig. 5.4d) R||P||Pa||Pa (Fig. 5.4m) (continued) 5.1 Fully-Parallel Topologies 531 Table 5.4 (continued) No. PM type Limb topology Connecting conditions 4. 3PaPPaR-1RPPaPa (Fig. 5.40b) Pa\\P\\\\Pa\\kR (Fig. 5.4b) Idem No. 3 Pa\\P\\\\Pa||R (Fig. 5.4c) R||P||Pa||Pa (Fig. 5.4m) 5. 3PaPPaR1RPaPatP (Fig. 5.41a) Pa||P\\Pa\\kR (Fig. 5.4a) Idem No. 1 Pa||P\\Pa||R (Fig. 5.4d) R||Pa||Pat||P (Fig. 5.4l) 6. 3PaPPaR1RPaPatP (Fig. 5.41b) Pa\\P\\\\Pa\\kR (Fig. 5.4b) Idem No. 1 Pa\\P\\\\Pa||R (Fig. 5.4c) R||Pa||Pat||P (Fig. 5.4l) 7. 3PaPPaR-1RPPaPat (Fig. 5.42a) Pa||P\\Pa\\kR (Fig. 5.4a) Idem No. 3 Pa||P\\Pa||R (Fig. 5.4d) R||P||Pa||Pat (Fig. 5.4n) 8. 3PaPPaR-1RPPaPat (Fig. 5.42b) Pa\\P\\\\Pa\\kR (Fig. 5.4b) Idem No", + "6a) The revolute joint between links 7D and 11 of G4 limb and the last revolute joints of limbs G1, G2 and G3 have superposed axes Pa\\Pa||Pa||R (Fig. 5.6b) R||P||Pa||Pat (Fig. 5.4n) 2. 3PaPaPaR1RPPaPat (Fig. 5.70) Pa\\Pa||Pa\\\\R (Fig. 5.6c) Idem No. 1 Pa\\Pa||Pa||R (Fig. 5.6b) R||P||Pa||Pat (Fig. 5.54n) 3. 3PaPPR1CPaPa (Fig. 5.71a) Pa||P\\P\\kR (Fig. 5.2a) The revolute joint between links 6D and 8 of G4 limb and the last revolute joints of limbs G1, G2 and G3 have superposed axes Pa||P\\P\\\\R (Fig. 5.2d) C||Pa||Pa (Fig. 5.4o) 4. 3PaPPR1CPaPa (Fig. 5.71b) Pa\\P\\kP\\kR (Fig. 5.2b) Idem No. 3 Pa\\P\\kP\\\\R (Fig. 5.2c) C||Pa||Pa (Fig. 5.40) 5. 3PaPPR1CPaPat (Fig. 5.72a) Pa||P\\P\\kR (Fig. 5.2a) Idem No. 3 Pa||P\\P\\\\R (Fig. 5.2d) C||Pa||Pat (Fig. 5.4p) 6. 3PaPPR1CPaPat (Fig. 5.72b) Pa\\P\\kP\\kR (Fig. 5.2b) Idem No. 3 Pa\\P\\kP\\\\R (Fig. 5.2c) C||Pa||Pat (Fig. 5.4p) (continued) 534 5 Topologies with Uncoupled Sch\u00f6nflies Motions Table 5.5 (continued) No. PM type Limb topology Connecting conditions 7. 2PaPRRR1PaPRR1CPaPa (Fig. 5.73a) Pa\\P\\\\R||R\\\\R (Fig. 5.3a) Idem No. 3 Pa\\P\\\\R||R (Fig. 5.2e) C||Pa||Pa (Fig. 5.40) 8. 2PaPRRR1PaPRR1CPaPa (Fig. 5.73b) Pa||P\\R||R\\\\R (Fig. 5.3b) Idem No. 3 Pa||P\\R||R (Fig. 5.2f) C||Pa||Pa (Fig. 5.40) 9. 2PaPRRR1PaPRR1CPaPat (Fig. 5.74a) Pa\\P\\\\R||R\\\\R (Fig. 5.3a) Idem No. 3 Pa\\P\\\\R||R (Fig. 5.2e) C||Pa||Pat(Fig. 5.4p) 10. 2PaPRRR1PaPRR1CPaPat (Fig. 5.74b) Pa||P\\R||R\\\\R (Fig. 5.3b) Idem No. 3 Pa||P\\R||R (Fig. 5.2f) C||Pa||Pat (Fig. 5.4p) 11. 2PaRPRR1PaRPR1CPaPa (Fig. 5.75) Pa\\R\\P\\kR\\R (Fig. 5.3c) Idem No. 3 Pa\\R\\P\\kR (Fig. 5.2g) C||Pa||Pa (Fig. 5.40) 12. 2PaRPRR1PaRPR1CPaPat (Fig. 5.76) Pa\\R\\P\\kR\\R (Fig. 5.3c) Idem No. 3 Pa\\R\\P\\kR (Fig. 5.2g) C||Pa||Pat (Fig. 5.4p) 13. 2PaRRRR1PaRRR1CPaPa (Fig. 5.77) Pa\\R||R||R\\kR (Fig. 5.3d) Idem No. 3 Pa\\R||R||R (Fig. 5.2h) C||Pa||Pa (Fig. 5.40) 14. 2PaRRRR1PaRRR1CPaPat (Fig. 5.78) Pa\\R||R||R\\kR (Fig. 5.3d) Idem No. 3 Pa\\R||R||R (Fig. 5.2h) C||Pa||Pat (Fig. 5.4p) 15. 3PaPPaR1CPaPa (Fig. 5.79a) Pa||P\\Pa\\kR (Fig. 5.4a) The revolute joint between links 6D and 9 of G4 limb and the last revolute joints of limbs G1, G2 and G3 have superposed axes Pa||P\\Pa||R (Fig. 5.4d) C||Pa||Pa (Fig. 5.40) 16. 3PaPPaR1CPaPa (Fig. 5.79b) Pa\\P\\\\Pa\\kR (Fig. 5.4b) Idem No. 15 Pa\\P\\\\Pa||R (Fig. 5.4c) C||Pa||Pa (Fig. 5.40) 17. 3PaPPaR1CPaPat (Fig. 5.80a) Pa||P\\Pa\\kR (Fig. 5.4a) Idem No. 15 Pa||P\\Pa||R (Fig. 5.4d) C||Pa||Pat (Fig. 5.4p) 18. 3PaPPaR1CPaPat (Fig. 5.80b) Pa\\P\\\\Pa\\kR (Fig. 5.4b) Idem No. 15 Pa\\P\\\\Pa||R (Fig. 5.4c) C||Pa||Pat (Fig. 5.4p) (continued) 5.1 Fully-Parallel Topologies 535 Table 5.5 (continued) No. PM type Limb topology Connecting conditions 19. 3PaPaPR1CPaPa (Fig. 5.81) Pa\\Pa\\kP\\\\R (Fig. 5.4f) Idem No. 15 Pa\\Pa\\kP\\kR (Fig. 5.4h) C||Pa||Pa (Fig. 5.4o) 20. 3PaPaPR1CPaPa (Fig. 5.82) Pa\\Pa\\\\P\\\\R (Fig. 5.4g) Idem No. 15 Pa\\Pa\\\\P\\kR (Fig. 5.4e) C||Pa||Pa (Fig. 5.40) 21. 3PaPaPR1CPaPat (Fig. 5.83) Pa\\Pa\\kP\\\\R (Fig. 5.4f) Idem No. 15 Pa\\Pa\\kP\\kR (Fig. 5.4h) C||Pa||Pat (Fig. 5.4p) 22. 3PaPaPR1CPaPat (Fig. 5.84) Pa\\Pa\\\\P\\\\R (Fig. 5.4g) Idem No. 15 Pa\\Pa\\\\P\\kR (Fig. 5.4e) C||Pa||Pat (Fig. 5.4p) 23. 2PaPaRRR1PaPaRR1CPaPa (Fig. 5.85) Pa\\Pa||R||R\\\\R (Fig. 5.5a) The revolute joint between links 6D and 10 of G4 limb and the last revolute joints of limbs G1, G2 and G3 have superposed axes Pa\\Pa||R||R (Fig. 5.4i) C||Pa||Pa (Fig. 5.40) 24. 2PaPaRRR1PaPaRR1CPaPa (Fig. 5.86) Pa\\Pa||R||R\\kR (Fig. 5.5b) Idem No. 23 Pa\\Pa||R||R (Fig. 5.4i) C||Pa||Pa (Fig. 5.40) 25. 2PaPaRRR1PaPaRR1CPaPat (Fig. 5.87) Pa\\Pa||R||R\\\\R (Fig. 5.5a) Idem No. 23 Pa\\Pa||R||R (Fig. 5.4i) C||Pa||Pat (Fig. 5.4p) 26. 2PaPaRRR1PaPaRR1CPaPat (Fig. 5.88) Pa\\Pa||R||R\\kR (Fig. 5.5b) Idem No. 23 Pa\\Pa||R||R (Fig. 5.4i) C||Pa||Pat (Fig. 5.4p) 27. 3PaPaPaR1CPaPa (Fig. 5.89) Pa\\Pa||Pa\\kR (Fig. 5.6a) The revolute joint between links 6D and 11 of G4 limb and the last revolute joints of limbs G1, G2 and G3 have superposed axes Pa\\Pa||Pa||R (Fig. 5.6b) C||Pa||Pa (Fig. 5.4o) 28. 3PaPaPaR1CPaPa (Fig" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_35_0001419_978-94-007-7401-8-Figure2.62-1.png", + "original_path": "designv11-35/openalex_figure/designv11_35_0001419_978-94-007-7401-8-Figure2.62-1.png", + "caption": "Fig. 2.62 4PaRPPa-type fully-parallel PM with coupled Sch\u00f6nflies motions defined by MF \u00bc SF \u00bc 4; RF\u00f0 \u00de \u00bc \u00f0v1; v2; v3;xd\u00de, TF = 0, NF = 30, limb topology Pa\\R||P\\Pa", + "texts": [ + "22n) The axes of the revolute joints connecting links 7 and 8 of the four limbs are parallel 42. 4PaPaPR (Fig. 2.54) Pa||Pa||P\\R (Fig. 2.22o) Idem No. 4 43. 4PRPaPa (Fig. 2.55) P\\R\\Pa||Pa (Fig. 2.22p) Idem No. 5 44. 4PPaRPa (Fig. 2.56) P||Pa\\R\\Pa (Fig. 2.22q) Idem No. 33 45. 4PPaPaR (Fig. 2.57) P||Pa||Pa\\R (Fig. 2.22r) Idem No. 4 46. 4PaPPaR (Fig. 2.58) Pa||P||Pa\\R (Fig. 2.22s) Idem No. 4 47. 4RPPaPa (Fig. 2.59) R\\P||Pa||Pa (Fig. 2.22t) Idem No. 15 48. 4PaPPaR (Fig. 2.60) Pa\\P\\\\Pa\\kR (Fig. 2.22u) Idem No. 4 49. 4PaRPaP (Fig. 2.61) Pa\\R\\Pa\\kR (Fig. 2.22v) Idem No. 34 50. 4PaRPPa (Fig. 2.62) Pa\\R||P\\Pa (Fig. 2.22w) Idem No. 34 51. 4PaPRPa (Fig. 2.63) Pa\\P||R\\Pa (Fig. 2.22x) Idem No. 33 52. 4PaPaPR (Fig. 2.64) Pa\\Pa\\\\P||R (Fig. 2.22y) Idem No. 4 53. 4PaPaRP (Fig. 2.65) Pa\\Pa\\\\R||P (Fig. 2.22z) Idem No. 41 54. 4RPaPaP (Fig. 2.66) R\\Pa\\\\Pa\\\\P (Fig. 2.22a0) Idem No. 15 55. 4RPaPPa (Fig. 2.67) R\\Pa\\kP\\\\Pa (Fig. 2.22b0) Idem No. 15 56. 4RPPaPa (Fig. 2.68) R||P\\Pa\\\\Pa (Fig. 2.22c0) Idem No. 15 57. 4PRPaPa (Fig. 2.69) P||R\\Pa\\\\Pa (Fig. 2.22d0) Idem No. 5 58. 4PaCPa (Fig. 2.70) Pa\\C\\Pa (Fig" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_35_0001603_301-Figure10-1.png", + "original_path": "designv11-35/openalex_figure/designv11_35_0001603_301-Figure10-1.png", + "caption": "Fig. 10. Schematic diagram of pin and ring wear machine", + "texts": [ + " It is clear, therefore, that the achievement of adequate lubrication depends on the properties of the solid materials as well as on the properties of the oil. Wear experiments in lubricated conditions are very sensitive to the conditions of lubrication, and it is easier to obtain a picture of the way in which material properties influence wear from experiments made in unlubricated conditions. Also, of course, the study of the wear of unlubricated materials has its own intrinsic importance. The apparatus needed is simpler and many useful conclusions can be drawn from experiments using a pin and ring apparatus such as is shown in Fig. 10. Two general types of wear can be observed, a A , rotating shaft; B, ring, s i n . diametei; C, flat-ended f in . diameter pin, pressed under load against ring. mild type and a severe type. In severe wear the surfaces remain metallic in appearance; there is extensive tearing, the wear fragments are of the order of tenths of a millimetre in size, and the structure of the rubbing metals is disrupted to a comparable depth. In mild wear, the surfaces often become discoloured, the wear fragments are usually smaller than one micron and, in a normal section, there is no distortion of the underlying structure visible in the optical microscope" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_35_0000125_s00170-021-07485-6-Figure2-1.png", + "original_path": "designv11-35/openalex_figure/designv11_35_0000125_s00170-021-07485-6-Figure2-1.png", + "caption": "Fig. 2 Generation principle and equivalent process for power skiving", + "texts": [ + " The rotation of cutter must ensure that the cutting edge is sliding inside the slot of workpiece along the feeding direction. The aforementioned working model of power skiving indicates the engagement of each tooth cut off a layer of material in slot with the help of both meshing motion and feeding motion, while the succession of engagements produces the whole slot via the feeding motion. This revels that the skiving process can be taken as a forming machining process for helical or spur gear like drawn in Fig. 2, in which the swept volume of one cutting edge (SVC) relative to the gear slot during engagement works as the wheel and its instant contact curve generates the slot surface following the feed motion. Consequently, one can deduce the generating process of power skiving as follows: each cutting edge generates one or several points on the desired surface at every engaging moment, and these points further develop a forming curve on the desired slot surface during each engagement of one cutting tooth" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_35_0000068_j.matpr.2021.01.065-Figure2-1.png", + "original_path": "designv11-35/openalex_figure/designv11_35_0000068_j.matpr.2021.01.065-Figure2-1.png", + "caption": "Fig. 2. Retreaded truck tire cut section.", + "texts": [ + " variations and their effect on CGR are studied. (R- Ratio of min stress to maxi stress) This illustrates a fully relaxing or non-relaxing condition. The effect of variables on fatigue CGR is studied by changing one at a time or all factors at once. Hence progressive model can be developed as per the required CGR considering the effect of variations in influencing factors A rubber specimen of the first re-treaded tire is taken from the belt edge area for testing for fatigue crack propagation as shown in Fig. 2 &Fig. 3In this paper sample is taken from the sidewall of the tirefor testing or fatigue life of tire rubber due to aging of rubber [29]. De Mattia flexing machine is used to test rubber specimen of size As shown in Fig. 5. Resistance to cracking is produced either by extension depending on the relative adjustment of the stationary and movable grips. The tests for dynamic fatigue are designed to stimulate the continually repeated distortions received in service by many rubber articles, such as a tire. These distortions may be produced by extension. The effect of the distortions is to weaken rubber until surface cracking or actual rupture the occurs. The choice of the type of strain is optional but notations shall be made of the type used, giving full details of the relative position of the grip and the travel [29]. The test specimen is cut at the edge with a very sharp razor blade. A Crack of 3 mm is cut in the middle of the edge of the specimen Fig. 2c. The test specimen is clamped in the De-Mattia machine, shown in Fig. 5. The temperature of the machine is set to the atmospheric temp of 25o C. The machine is started and the number of cycles for the crack to reach a specified growth is recorded. The part of tire rubber which undergoes continuous flexing during its life cycle is the belt edge. Hence attention was concentrated on the flexing resistance at the belt edge as shown in Fig. 2.Tensile property is tested by this method. Tension tests are made on a power \u2013driven Universal Tensile Testing Machine(UTM). In (UTM), grips are provided with pneumatic pressure ie air is circulated. The inlet pressure adjusts automatically with pneumatic grips, allowing the gripping force to remain constant even though during a test the specimen thickness changes significantly. This machine has capacity to produce a uniform rate of grip separation of 500 mm/min.. The temperature for testing shall be 25 \u00b1 2 C", + "06x2 + 30915x + 7264.9 , =1205.54 J 0.3719 6.0255 17 62(25\u201317) 4.75 y = -405.49x2 + 24980x + 1402.5, =393.10 J 0.1668 7.833 25 62(25\u201325)4.75 0 0 12.65 is adjusted in such a way that it gives the required strain in the specimen in a uniaxial tension test. The Actual size of the specimen is excluding the rubber required for holding in the jaws of the De-Mattia Machine. The specimen is run till it is failed and readings are recorded for strain and stress for selected. Lengths of crack as shown in Fig. 2& Fig. 4. A graph is plotted between fatigue cycles and crack length to calculate the CGR.[ SED-Strain energy density, CGR- crack growth rate,] On the De-Mattia machine, different specimens are tested to study the growth of the crack .as shown in Table 1. Specimen of rubber strip with 3 mm pre-cut and frequency of 5 Hz is taken for explaining the procedure of testing. UTM machine is used for measuring strain and stress at different crack. First strain energy density and then tearing energy are calculated" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_35_0000738_slct.202101957-Figure20-1.png", + "original_path": "designv11-35/openalex_figure/designv11_35_0000738_slct.202101957-Figure20-1.png", + "caption": "Figure 20. Diagram of fluorescent-off sensing membrane for detecting free chlorine in water.", + "texts": [ + " Since the concentration of chlorine in water is related to water quality, it is necessary to control its concentration within a certain range. Because highly oxidized chlorine-containing solutions can quenched the fluorescence of CDs. Ruiz et al.[25] hydrothermally treated a mixture of 1,3,6-trinitropyrene, ammonia and hydrazine hydrate to obtain CDs and wrapped CDs in polyacrylonitrile (PAN) nanofiber membranes through electrospinning technology. The CDs-PAN membrane can respond quickly in the chlorine concentration range of 10\u223c600 \u03bcM and the detection limit is 2.00 \u03bcM (Figure 20). Therefore, the membrane can evaluate water quality. Agricultural products need oxygen to breathe, so the change of oxygen concentration in the packaging is an indicator to measure the freshness of agricultural products. Xu et al.[27] used silicon as the substrate, tris (4,7-diphenyl-1,10phenanthroline) ruthenium (II) dichloride (RuDPP) as the oxygen-sensitive probe, and silane-functionalized carbon dots (SiCDs) are oxygen-sensitive luminous bodies, and oxygensensitive photoluminescent fiber membranes are prepared by a combination of sol-gel method and electrospinning technology" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_35_0000686_j.mechmachtheory.2021.104433-Figure7-1.png", + "original_path": "designv11-35/openalex_figure/designv11_35_0000686_j.mechmachtheory.2021.104433-Figure7-1.png", + "caption": "Fig. 7. Epicyclic gear-rack train.", + "texts": [ + " This gives the following geometrical relationship between the length of the links lbDE, lbAC, lbCD in the equivalent mechanism and the radiuses Rb, rb of the pitch circles in the epicyclic bevel gear train. \u23a7 \u23aa \u23aa \u23aa \u23aa \u23a8 \u23aa \u23aa \u23aa \u23aa \u23a9 tan\u03b4b = rb/Rb \u03b1b1 = arcsin(sin\u03b1bcos\u03b4b) \u03b1b2 = arctan(tan\u03b1bsin\u03b4b \u03b1b3 = arcsin(sin\u03b1bsin\u03b4b) lbDE = rbcos\u03b1b1 lbBC = Rbcos\u03b1b3 lbCD = rbsin\u03b1b1 + Resin\u03b1b3 (13) For the epicyclic gear-rack train that is a new design by assuming the planet gear having an infinite diameter, a base coordinate system, sun rack-gear coordinate system, rack arm coordinate system, and planet rack coordinate system are built, as shown in Fig. 7. Note that the z-axis of all coordinate systems is omitted because they are all perpendicular to the moving plane. Similar to the epicyclic external gear train, the transmission relationship of the epicyclic gear-rack train can be written as Eq. (14). \u03c9r1Rr \u2212 \u03c9r2Rr = vr3 (14) where \u03c9r1, \u03c9r2 represent the scalar angular velocities of the sun rack-gear, rack arm rotating around the zr0-axis respectively. vr3 represents the velocity of the planet rack moved along with the ye3-axis. Rr refers to the pitch radius of the sun rack-gear" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_35_0000686_j.mechmachtheory.2021.104433-Figure2-1.png", + "original_path": "designv11-35/openalex_figure/designv11_35_0000686_j.mechmachtheory.2021.104433-Figure2-1.png", + "caption": "Fig. 2. Structures of the spatial epicyclic gear train - epicyclic bevel gear train.", + "texts": [], + "surrounding_texts": [ + "Three typical epicyclic gear trains [47], i.e. an epicyclic external gear train, an epicyclic gear-rack train, and an epicyclic bevel gear train are shown in Figs. 1 and 2. As illustrated in Figs. 1 and 2, the epicyclic gear train consists of a sun gear, a planet gear, arm, and a ring gear. The planet gear rotates about the sun gear and mounts on the arm, which itself rotates relatively to the sun gear. Meanwhile, the sun gear and the planet gear rotate about their axes of rotation. As in most cases, input and output are carried out by a sun gear and a planet gear, the ring gear does not need to be considered. For the mobility of an epicyclic gear train, it can be given as Mj = \u03bb(L \u2212 1) \u2212 \u2211\u03bb\u2212 1 i=1 (\u03bb \u2212 i)Ji (j= e, r, b) (1) where Mj represents the mobility of an epicyclic gear train. When j is e, it refers to an epicyclic external gear train. When j is r, it refers to an epicyclic rack gear train. When j is b, it refers to an epicyclic bevel gear train. L is the number of links including the ground link. Ji gives the number of joints that provide i mobility. Coefficient \u03bb gives the mechanism type. For the planar mechanism, coefficient \u03bb is 3 and for the spatial mechanism, coefficient \u03bb is 6. In an epicyclic external gear train, there are four links, three revolute joints, and one higher gear pair. Since it is a planar gear train, \u03bb should take the value of 3. Further, L=4, J1=3, and J2=1. This gives mobility two from the above equation. In an epicyclic gear-rack train, there are four links, two revolute joints, one prismitc joint and one higher gear-rack set. Since it is a planar gear train, \u03bb should take the value of 3. Further, L=4, J1=3, and J2=1. This gives mobility two from the above equation. In an epicyclic bevel gear train, there are four links, three revolute joints,and one higher bevel gear set. As it is a spatial gear train, \u03bb H. Yuan et al. Mechanism and Machine Theory 166 (2021) 104433 should take the value of 6. Further, L=4, J1=3, and J5=1. This gives mobility two from the above equation. The higher pair of epicyclic gear trains is the gear meshing, it needs to be equivalent to the lower pair for a kinematic model. A link perpendicular to the instantaneous screw axis and collinear with the mesh line is used to replace the higher pair by calculating the position of the instantaneous screw axis [48] to obtain the mesh points [49]. In this way, equivalent mechanisms are established in this section." + ] + }, + { + "image_filename": "designv11_35_0001419_978-94-007-7401-8-Figure6.3-1.png", + "original_path": "designv11-35/openalex_figure/designv11_35_0001419_978-94-007-7401-8-Figure6.3-1.png", + "caption": "Fig. 6.3 Maximaly regular fully-parallel PMs with Sch\u00f6nflies motions of types 3PPPR-1RUPU (a) and 2PPRRR-1PPRR-1RUPU (b) defined by MF = SF = 4, (RF) = (v1; v2; v3;xb), TF = 0, NF = 4 (a), NF = 2 (b), limb topology R\\R\\R\\P\\||R\\R and P\\P\\\\P\\||R, P\\P\\\\P\\\\R (a), P\\P\\||R||R\\R, P\\P\\||R||R (b)", + "texts": [ + " rG4 0 0 19. MGi (i = 1, 3) 4 5 20. MG2 4 4 21. MG4 4 4 22. (RF) (v1; v2; v3;xb) (v1; v2; v3;xb) 23. SF 4 4 24. rl 0 0 25. rF 12 14 26. MF 4 4 27. NF 6 4 28. TF 0 0 29. Pp1 j\u00bc1 fj 4 5 30. Pp2 j\u00bc1 fj 4 4 31. Pp3 j\u00bc1 fj 4 5 32. Pp4 j\u00bc1 fj 4 4 33. Pp j\u00bc1 fj 16 18 a See footnote of Table 2.2 for the nomenclature of structural parameters 588 6 Maximally Regular Topologies with Sch\u00f6nflies Motions Table 6.3 Structural parametersa of parallel mechanisms in Figs. 6.3 and 6.4 No. Structural parameter Solution Figure 6.3a Figures 6.3b and 6.4 1. m 16 18 2. pi (i = 1,3) 4 5 3. p2 4 4 4. p4 6 6 5. p 18 20 6. q 3 3 7. k1 4 4 8. k2 0 0 9. k 4 4 10. (RG1) (v1; v2; v3;xb) (v1; v2; v3;xa;xb) 11. (RG2) (v1; v2; v3;xb) (v1; v2; v3;xb) 12. (RG3) (v1; v2; v3;xb) (v1; v2; v3;xb;xd) 13. (RG4) (v1; v2; v3;xa;xb;xd) (v1; v2; v3;xa;xb;xd) 14. SGi (i = 1, 3) 4 5 15. SG2 4 4 16. SG4 6 6 17. rGi (i = 1, 2, 3) 0 0 18. rG4 0 0 19. MGi (i = 1, 3) 4 5 20. MG2 4 4 21. MG4 6 6 22. (RF) (v1; v2; v3;xb) (v1; v2; v3;xb) 23. SF 4 4 24. rl 0 0 25" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_35_0001419_978-94-007-7401-8-Figure5.52-1.png", + "original_path": "designv11-35/openalex_figure/designv11_35_0001419_978-94-007-7401-8-Figure5.52-1.png", + "caption": "Fig. 5.52 2PaPaRRR-1PaPaRR-1RPaPaP-type fully-parallel PM with uncoupled Sch\u00f6nflies motions defined by MF = SF = 4, RF\u00f0 \u00de \u00bc \u00f0v1; v2; v3;xb\u00de, TF = 0, NF = 28, limb topology Pa\\Pa||R||R\\kR, Pa\\Pa||R||R and R||Pa||Pa||P", + "texts": [ + " 3PaPaPR-1RPPaPat (Fig. 5.49) Pa\\PakP\\\\R (Fig. 5.4f) Idem No. 3 Pa\\Pa\\kP\\kR (Fig. 5.4h) R||P||Pa||Pat (Fig. 5.4n) 16. 3PaPaPR-1RPPaPat (Fig. 5.50) Pa\\Pa\\\\P\\\\R (Fig. 5.4g) Idem No. 3 Pa\\Pa\\\\P\\kR (Fig. 5.4e) R||P||Pa||Pat (Fig. 5.4n) 17. 2PaPaRRR-1PaPaRR1RPaPaP (Fig. 5.51) Pa\\Pa||R||R\\\\R (Fig. 5.5a) Idem No. 1 Pa\\Pa||R||R (Fig. 5.4i) R||Pa||Pa||P (Fig. 5.4k) (continued) 532 5 Topologies with Uncoupled Sch\u00f6nflies Motions Table 5.4 (continued) No. PM type Limb topology Connecting conditions 18. 2PaPaRRR-1PaPaRR1RPaPaP (Fig. 5.52) Pa\\Pa||R||R\\kR (Fig. 5.5b) Idem No. 1 Pa\\Pa||R||R (Fig. 5.4i) R||Pa||Pa||P (Fig. 5.4k) 19. 2PaPaRRR-1PaPaRR1RPaPatP (Fig. 5.53) Pa\\Pa||R||R\\\\R (Fig. 5.5a) Idem No. 1 Pa\\Pa||R||R (Fig. 5.4i) R||Pa||Pat||P (Fig. 5.4l) 20. 2PaPaRRR-1PaPaRR1RPaPatP (Fig. 5.54) Pa\\Pa||R||R\\kR (Fig. 5.5b) Idem No. 1 Pa\\Pa||R||R (Fig. 5.4i) R||Pa||Pat||P (Fig. 5.4l) 21. 2PaPaRRR-1PaPaRR1RPPaPa (Fig. 5.55) Pa\\Pa||R||R\\\\R (Fig. 5.5a) The revolute joint between links 7D and 10 of G4 limb and the last revolute joints of limbs G1, G2 and G3 have superposed axes Pa\\Pa||R||R (Fig" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_35_0000433_tie.2021.3078396-Figure2-1.png", + "original_path": "designv11-35/openalex_figure/designv11_35_0000433_tie.2021.3078396-Figure2-1.png", + "caption": "Fig. 2: Simple model of RoK-3.", + "texts": [ + " Therefore, we propose a method on how to generate a jumping motion given the flight time and the x-direction velocity at takeoff. For convenience, we independently generated vertical and horizontal motions. Of course, the two motions affect each other; however, we ignored the effects for simplicity of motion generation. This paper only deals with trajectory generation for standing jumping in the x-direction and does not consider the y-direction; however, the method can be regarded as the same as the x-direction. To solve the inverse dynamics, we use a simple robot model in the sagittal plane as shown in Fig. 2. Fig. 3 shows the entire system of the robot as a block diagram. The vertical COM trajectory (zdescom, z\u0307 des com, z\u0308 des com) is obtained offline through a nonlinear optimization process, and the horizontal COM trajectory is generated through MPC every step period according to the operator\u2019s command which is desired horizontal COM velocity (x\u0307descom). The obtained vertical and horizontal COM trajectories are converted into desired linear momentuma (P desx , P desz ) and added with \u03b4P desx obtained from the ZMP-based balance controller [24]" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_35_0000095_s00170-021-06769-1-Figure6-1.png", + "original_path": "designv11-35/openalex_figure/designv11_35_0000095_s00170-021-06769-1-Figure6-1.png", + "caption": "Fig. 6 Knife holder layout. 1 \u2013 fixing block; 2 \u2013 force sensor; 3 \u2013 knife holder; 4 \u2013 knife, 5 \u2013 wood sample", + "texts": [ + " The samples had a width of 20 mm, length (= cutting length lc) amounted to 100 mm. For examination, the sample is mounted to the aluminum holder using woodscrews from the bottom side. To avoid initial hit excitation, which can influence measured force values, the sample is squeezed between two polyurethane blocks (homogeneous density of 530 kg/m3). Samples were examined in cutting direction B according to Kivimaa [3]. For cutting examinations, a standardized single knife (LEITZ, Turnblade knife, HW-05: 30 \u00d7 12 \u00d7 1.5) was used. It was placed to the knife holder (Fig. 6) with a constant cutting angle \u03b4 = (\u03b1 + \u03b2) of 60\u00b0. The knife holder (Fig. 7) has been optimized to achieve the highest natural frequency possible. The topology optimization of shape was used to achieving the best stiffness\u2013mass ratio. The knife holder is made of high stiffness steel. It weighs only 146 g; nevertheless, it can resist the stress of approx. 1000 MPa from the force sensor pre-load. Experiments were performed at speeds of 10\u201380 m/s, in steps of 10 m/s. Investigated chip thicknesses were 0", + " In the scale of moisture content, it changes equal to 0.1\u20130.18 % per 20 min. Since the experiment duration was very short (max. 20 min), the influence was negligible. In case of longer examinations (in range of hours), the moisture content change must be taken into consideration, or experiment cuts have to be distributed to more samples, where each will be picked from a regulated climate just prior to the test. A major advantage of the test set-up is the opposite configuration of the knife (Fig. 1, Fig. 6). A force sensor is usually placed beneath a sample. This design is normally chosen because of its simplicity. Mounting of a spindle does not enable placing a sensor in-between knife and holder or holder and drive. Commonly, this assembly is because of the performance and accuracy considered untouchable. From the unique design of the device, the sensor can be placed on both sides. The decision to place it below the knife holder has been done considering the weight-stiffness ratio of free mass. The knife holder with a knife represents the free mass which is supposed to be minimized" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_35_0000477_j.procir.2021.05.085-Figure4-1.png", + "original_path": "designv11-35/openalex_figure/designv11_35_0000477_j.procir.2021.05.085-Figure4-1.png", + "caption": "Figure 4. Non-assembly with normal support placed at everywhere", + "texts": [ + " It noticed that the tree support takes 281 min to complete the non-assembly which does not seem to be reasonable. This type of support structure and placement has also increased the weight of the support material to 3.91 g. In this support structure, the printer build supports for each overhanging feature like joints and circular features in the nonassembly. However, the printer starts to print the support structure from the most recent or upper layers of the features of the non-assembly can be seen in Fig. 4. In this structure type, support material and build time are comparatively lesser than tree structure. It has been observed that shifting from Tree support to normal support, build time of the non-assembly is significantly decreased to 182 min and support material is reduced to 2.36 g. It has been observed that supports build in the circular features of the non-assembly are stiff and rigid and difficult to remove due to strong adherence with the actual part surface as shown in Fig. 4. Removal of support structures from the joints and fragile features needs more attention which seems to be a time-consuming activity after the printing. It is also observed that the support removal process affects the part surface and sometimes cause damage upon removal. Movement of the joints is also affected by the improper removal of the support structure which is a serious concern of this support structure as compared to tree structure support. This case is similar to case 1 to a certain extent of support placements (support everywhere and touching build plate)" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_35_0001419_978-94-007-7401-8-Figure2.35-1.png", + "original_path": "designv11-35/openalex_figure/designv11_35_0001419_978-94-007-7401-8-Figure2.35-1.png", + "caption": "Fig. 2.35 Fully-parallel PMs with coupled Sch\u00f6nflies motions of types 4PaRPP (a) and 4PaPRP (b) defined by MF \u00bc SF \u00bc 4; RF\u00f0 \u00de \u00bc \u00f0v1; v2; v3;xd\u00de, TF = 0, NF = 18, limb topology Pa\\R||P\\P (a) and Pa||P\\R||P (b)", + "texts": [ + "21k) The before last joints of the four limbs have parallel axes 12. 4PPaPR (Fig. 2.31b) P||Pa\\P|| R (Fig. 2.21l) Idem No. 4 13. 4PPaPR (Fig. 2.32a) P\\Pa||P\\\\R (Fig. 2.21m) Idem No. 4 14. 4PPaRP (Fig. 2.32b) P\\Pa\\kR\\kP (Fig. 2.21n) Idem No. 4 15. 4RPPaP (Fig. 2.33a) R\\P||Pa\\kP (Fig. 2.21o) The first revolute joints of the four limbs have parallel axes 16. 4RPPPa (Fig. 2.33b) R||P\\P||Pa (Fig. 2.21p) Idem No. 15 17. 4PaPPR (Fig. 2.34a) Pa||P\\P||R (Fig. 2.21q) Idem No. 4 18. 4PaPPR (Fig. 2.34b) Pa\\P\\kP\\kR (Fig. 2.21r) Idem No. 4 19. 4PaRPP (Fig. 2.35a) Pa\\R||P\\P (Fig. 2.21s) Idem No. 11 20. 4PaPRP (Fig. 2.35b) Pa||P\\R||P (Fig. 2.21t) Idem No. 5 21. 4PaPRP (Fig. 2.36a) Pa\\P||R\\P (Fig. 2.21u) Idem No. 11 22. 4PaRPP (Fig. 2.36b) Pa\\R\\P\\kP (Fig. 2.21v) Idem No. 5 23. 4PCPa (Fig. 2.37a) P\\C\\Pa (Fig. 2.21w) Idem No. 5 24. 4PPaC (Fig. 2.37b) P||Pa\\C (Fig. 2.21x) The last joints of the four limbs have parallel axes 25. 4CPPa (Fig. 2.38a) C\\P||Pa (Fig. 2.21y) The first joints of the four limbs have parallel axes 26. 4PaPC (Fig. 2.38b) Pa||P\\C (Fig. 2.21z) Idem No. 24 27. 4PaCP (Fig. 2.39) Pa\\C\\P (Fig. 2.21z1) Idem No" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_35_0001419_978-94-007-7401-8-Figure3.8-1.png", + "original_path": "designv11-35/openalex_figure/designv11_35_0001419_978-94-007-7401-8-Figure3.8-1.png", + "caption": "Fig. 3.8 Overactuated PMs with coupled Sch\u00f6nflies motions of types 4PPRR (a) and 4PRPR (b) defined by MF \u00bc SF \u00bc 4; RF\u00f0 \u00de \u00bc v1; v2; v3;xd\u00f0 \u00de, TF = 0, NF = 6, limb topology P\\P||R||R (a) and P\\R||P||R (b)", + "texts": [ + "1a) The prismatic joints of the four limbs have parallel directions 2. 4RPRR (Fig. 3.4b) R||P||R||R (Fig. 3.1b) The first revolute joints of the four limbs have parallel axes 3. 4RPRR (Fig. 3.5a) R||P||R||R (Fig. 3.1c) Idem No. 2 4. 4RRPR (Fig. 3.5b) R||R||P||R (Fig. 3.1d) Idem No. 2 5. 4RRPR (Fig. 3.6a) R||R||P||R (Fig. 3.1e) Idem No. 2 6. 4PPRR (Fig. 3.6b) P\\P\\kR||R (Fig. 3.1i) Idem No. 1 7. 4RRRP (Fig. 3.7a) R||R||R||P (Fig. 3.1f) Idem No. 2 8. 4RPRP (Fig. 3.7b) R\\P\\kR||P (Fig. 3.1g) Idem No. 2 9. 4PPRR (Fig. 3.8a) P\\P||R||R (Fig. 3.1h) The last revolute joints of the four limbs have parallel axes 10. 4PRPR (Fig. 3.8b) P\\R||P||R (Fig. 3.1p) Idem No. 9 11. 4RRPP (Fig. 3.9a) R||R||P\\P (Fig. 3.1j) Idem No. 2 12. 4RPRP (Fig. 3.9b) R||P||R\\P (Fig. 3.1k) Idem No. 2 13. 4RPPR (Fig. 3.10a) R||P\\P\\kR (Fig. 3.1l) Idem No. 2 14. 4RPPR (Fig. 3.10b) R\\P\\kP||R (Fig. 3.1m) Idem No. 2 15. 4PRRP (Fig. 3.11a) P\\R||R||P (Fig. 3.1n) The last prismatic joints of the four limbs have parallel directions 16. 4PRPR (Fig. 3.11b) P\\R||P||R (Fig. 3.1o) Idem No. 9 17. 4PRRP (Fig. 3.12a) P||R||R\\P (Fig. 3.1v) The first prismatic joints of the four limbs have parallel directions 18" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_35_0001419_978-94-007-7401-8-Figure5.43-1.png", + "original_path": "designv11-35/openalex_figure/designv11_35_0001419_978-94-007-7401-8-Figure5.43-1.png", + "caption": "Fig. 5.43 3PaPaPR-1RPaPaP-type fully-parallel PM with uncoupled Sch\u00f6nflies motions defined by MF = SF = 4, RF\u00f0 \u00de \u00bc \u00f0v1; v2; v3;xb\u00de, TF = 0, NF = 30 limb topology Pa\\Pa\\kP\\\\R, Pa\\Pa\\kP\\kR and R||Pa||Pa||P", + "texts": [ + "4c) R||P||Pa||Pa (Fig. 5.4m) 5. 3PaPPaR1RPaPatP (Fig. 5.41a) Pa||P\\Pa\\kR (Fig. 5.4a) Idem No. 1 Pa||P\\Pa||R (Fig. 5.4d) R||Pa||Pat||P (Fig. 5.4l) 6. 3PaPPaR1RPaPatP (Fig. 5.41b) Pa\\P\\\\Pa\\kR (Fig. 5.4b) Idem No. 1 Pa\\P\\\\Pa||R (Fig. 5.4c) R||Pa||Pat||P (Fig. 5.4l) 7. 3PaPPaR-1RPPaPat (Fig. 5.42a) Pa||P\\Pa\\kR (Fig. 5.4a) Idem No. 3 Pa||P\\Pa||R (Fig. 5.4d) R||P||Pa||Pat (Fig. 5.4n) 8. 3PaPPaR-1RPPaPat (Fig. 5.42b) Pa\\P\\\\Pa\\kR (Fig. 5.4b) Idem No. 3 Pa\\P\\\\Pa||R (Fig. 5.4c) R||P||Pa||Pat (Fig. 5.4n) 9. 3PaPaPR-1RPaPaP (Fig. 5.43) Pa\\Pa\\kP\\\\R (Fig. 5.4f) Idem No. 1 Pa\\Pa\\kP\\kR (Fig. 5.4h) R||Pa||Pa||P (Fig. 5.4k) 10. 3PaPaPR-1RPaPaP (Fig. 5.44) Pa\\Pa\\\\P\\\\R (Fig. 5.4g) Idem No. 1 Pa\\Pa\\\\P\\kR (Fig. 5.4e) R||Pa||Pa||P (Fig. 5.4k) 11. 3PaPaPR1RPaPatP (Fig. 5.45) Pa\\Pa\\kP\\\\R (Fig. 5.4f) Idem No. 1 Pa\\Pa\\kP\\kR (Fig. 5.4h) R||Pa||Pat||P (Fig. 5.4l) 12. 3PaPaPR1RPaPatP (Fig. 5.46) Pa\\Pa\\\\P\\\\R (Fig. 5.4g) Idem No. 1 Pa\\Pa\\\\P\\kR (Fig. 5.4e) R||Pa||Pat||P (Fig. 5.4l) 13. 3PaPaPR-1RPPaPa (Fig. 5.47) Pa\\Pa\\kP\\\\R (Fig. 5.4f) Idem No. 3 Pa\\Pa\\kP\\kR (Fig" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_35_0001419_978-94-007-7401-8-Figure3.76-1.png", + "original_path": "designv11-35/openalex_figure/designv11_35_0001419_978-94-007-7401-8-Figure3.76-1.png", + "caption": "Fig. 3.76 4PPPaR-type overactuated PMs with coupled Sch\u00f6nflies motions defined by MF \u00bc SF \u00bc 4; RF\u00f0 \u00de \u00bc v1; v2; v3;xd\u00f0 \u00de, TF = 0, NF = 18, limb topology P\\P||Pa\\\\R (a) and P\\P\\kPa\\\\R (b)", + "texts": [ + " 3.72) Pa||R||P||R (Fig. 3.50b0) Idem No. 1 31. 4PaPRP (Fig. 3.73a) Pa||P||R\\P (Fig. 3.50c0) Idem No. 1 32. 4PaRPP (Fig. 3.73b) Pa||R\\P\\kP (Fig. 3.50d0) Idem No. 1 33. 4PRPPa (Fig. 3.74a) P\\R\\P||Pa (Fig. 3.50e0) The second joints of the four limbs have parallel axes 34. 4PRPPa (Fig. 3.74b) P\\R\\P\\Pa (Fig. 3.50f0) Idem No. 33 35. 4PRPaP (Fig. 3.75a) P\\R\\Pa\\\\P (Fig. 3.50g0) Idem No. 33 36. 4PPRPa (Fig. 3.75b) P\\P\\\\R\\Pa (Fig. 3.50h0) The third joints of the four limbs have parallel axes 37. 4PPPaR (Fig. 3.76a) P\\P||Pa\\\\R (Fig. 3.50j0) The last joints of the four limbs have parallel axes (continued) 3.2 Topologies with Complex Limbs 359 Table 3.5 (continued) No. PM type Limb topology Connecting conditions 38. 4PPPaR (Fig. 3.76b) P\\P\\kPa\\\\R (Fig. 3.50k0) Idem No. 37 39. 4PaPPR (Fig. 3.77a) Pa||P\\P\\\\R (Fig. 3.50l0) Idem No. 37 40. 4PaPPR (Fig. 3.77b) Pa\\P\\kP\\\\R (Fig. 3.50m0) Idem No. 37 41. 4PaRPR (Fig. 3.78) Pa\\R\\P\\kR (Fig. 3.50n0) Idem No. 37 42. 4PaPRR (Fig. 3.79a) Pa||P\\R||R (Fig. 3.50o0) Idem No. 37 43. 4PaPRR (Fig. 3.79b) Pa\\P\\\\R||R (Fig. 3.50p0) Idem No. 37 44. 4PaRRP (Fig. 3.80a) Pa\\R||R\\P (Fig. 3.50q0) Idem No. 23 45. 4RPaRR (Fig. 3.80b) R\\Pa\\kR||R (Fig. 3.50s0) Idem No. 37 46. 4PaRRR (Fig. 3.81) Pa\\R||R||R (Fig" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_35_0000686_j.mechmachtheory.2021.104433-Figure23-1.png", + "original_path": "designv11-35/openalex_figure/designv11_35_0000686_j.mechmachtheory.2021.104433-Figure23-1.png", + "caption": "Fig. 23. Metamorphic epicyclic gear-rack train based on the combination constraint.", + "texts": [ + " 19(a)), the metamorphic epicyclic bevel gear train transforms to rotation configuration which is similar to the example discussed above and has mobility 2 H. Yuan et al. Mechanism and Machine Theory 166 (2021) 104433 By combining geometrical constraint and force constraint, a set of metamorphic epicyclic gear trains can be obtained based on a combination constraint. As shown in Figs. Figures 21-23, the combination constraint can provide both geometrical limitation and constraint forces. The metamorphic epicyclic gear-rack train in Fig. 23 is used to demonstrate the configuration transformation. In the rotation configuration, the driving force drives the planet rack overcoming the constraint force, and the spring in combination constraint is compressed or released. The angular velocity \u03c9r2 is not equal to \u03c9r1 the corresponding motion branch has mobility 2 as discussed in Fig. 14(b). When the driving force is equal to the constraint force or the planet rack is locked by geometrical limitation, the metamorphic epicyclic gear-rack train transforms to revolution configuration" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_35_0000535_tmag.2021.3085750-Figure3-1.png", + "original_path": "designv11-35/openalex_figure/designv11_35_0000535_tmag.2021.3085750-Figure3-1.png", + "caption": "Fig. 3. Illustration of Fourier-series based hysteresis model. (a) Simplified view for illustrating the variable flux principle. (b) Numerical representation.", + "texts": [ + " BJ-13 2 distributions under different magnetization states (MSs) in Fig. 2. It can be observed that the air-gap flux can be flexibly adjusted by applying a transient magnetizing or demagnetizing current pulse. Meanwhile, due to the variable flux capability, the losses at different speeds and loads can be also manipulated to realize efficiency improvement over a wide operating range. The nonlinear hysteresis characteristics of the LCF PMs used in the proposed HMC-VFMM can be characterized by a simple Fourier-series based hysteresis model as shown in Fig. 3(a). The PM working point can track along different recoil lines by applying a specific remagnetizing or demagnetizing current pulse and finally stabilize at the intersect of the recoil line and load line. The corresponding numerical model shown in Fig. 3(b) is utilized to analytically reveal the balanced bidirectional-magnetization effect. B The developed Fourier-series based hysteresis model [9] consists of a set of recoil lines and the major loop, where the recoil lines can be simplified by a set of linear B-H functions as follows: 0 , 1, 2,3,... r ri B H B i = + = (1) where \u03bc0 and \u03bcr are the vacuum permeability and the relative permeability of LCF PM respectively; H is the positive magnetic field intensity, and Bri represents the corresponding ith remanence of the sets of hysteresis loops", + " The left branch of the proposed Fourier-series based hysteresis model can be expressed as: ( ) l H f B= (4) The right branch can be derived from the left one owing to the central symmetry of the hysteresis loop: The values of coefficients of the Fourier-series based hysteresis model can be obtained by the experiments [9]. As a result, the operating point of LCF PM can be specified by the intersection of load line and recoil line. From the equivalent magnetic circuit calculation, it can be deduced that for series VFMMs, the load line is shifted to the right compared to the parallel VFMMs [8]. For the HMC design, the load line is placed between the parallel and series cases shown in Fig. 3(b), which allows a better balance between the magnetizing and demagnetizing current amplitudes. Fig. 4 (a) and (b) show the equivalent circuits of the HMC- VFMM subject to remagnetizing and demagnetizing processes. Fm21, Fm22 and Fm1 represent the magnetomotive force (MMF) of the HCF PMs of series and parallel branch, and the full magnetization MMF of LCF PM respectively. Rm21, Rm22 and Rm1 represent the magnetic reluctance of the PMs accordingly. Rg is the magnetic resistance of the airgap, and Fd is the required MMF for MS manipulation" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_35_0001419_978-94-007-7401-8-Figure3.77-1.png", + "original_path": "designv11-35/openalex_figure/designv11_35_0001419_978-94-007-7401-8-Figure3.77-1.png", + "caption": "Fig. 3.77 4PaPPR-type overactuated PMs with coupled Sch\u00f6nflies motions defined by MF \u00bc SF \u00bc 4; RF\u00f0 \u00de \u00bc v1; v2; v3;xd\u00f0 \u00de, TF = 0, NF = 18, limb topology Pa||P\\P\\\\R (a) and Pa\\P\\kP\\\\R (b)", + "texts": [ + " 3.74b) P\\R\\P\\Pa (Fig. 3.50f0) Idem No. 33 35. 4PRPaP (Fig. 3.75a) P\\R\\Pa\\\\P (Fig. 3.50g0) Idem No. 33 36. 4PPRPa (Fig. 3.75b) P\\P\\\\R\\Pa (Fig. 3.50h0) The third joints of the four limbs have parallel axes 37. 4PPPaR (Fig. 3.76a) P\\P||Pa\\\\R (Fig. 3.50j0) The last joints of the four limbs have parallel axes (continued) 3.2 Topologies with Complex Limbs 359 Table 3.5 (continued) No. PM type Limb topology Connecting conditions 38. 4PPPaR (Fig. 3.76b) P\\P\\kPa\\\\R (Fig. 3.50k0) Idem No. 37 39. 4PaPPR (Fig. 3.77a) Pa||P\\P\\\\R (Fig. 3.50l0) Idem No. 37 40. 4PaPPR (Fig. 3.77b) Pa\\P\\kP\\\\R (Fig. 3.50m0) Idem No. 37 41. 4PaRPR (Fig. 3.78) Pa\\R\\P\\kR (Fig. 3.50n0) Idem No. 37 42. 4PaPRR (Fig. 3.79a) Pa||P\\R||R (Fig. 3.50o0) Idem No. 37 43. 4PaPRR (Fig. 3.79b) Pa\\P\\\\R||R (Fig. 3.50p0) Idem No. 37 44. 4PaRRP (Fig. 3.80a) Pa\\R||R\\P (Fig. 3.50q0) Idem No. 23 45. 4RPaRR (Fig. 3.80b) R\\Pa\\kR||R (Fig. 3.50s0) Idem No. 37 46. 4PaRRR (Fig. 3.81) Pa\\R||R||R (Fig. 3.50r0) Idem No. 37 47. 4PaPRP (Fig. 3.82a) Pa||P\\R\\P (Fig. 3.50t0) Idem No. 23 48. 4PaPRP (Fig. 3.82b) Pa\\P\\\\R\\P (Fig. 3" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_35_0000504_tmag.2021.3087267-Figure2-1.png", + "original_path": "designv11-35/openalex_figure/designv11_35_0000504_tmag.2021.3087267-Figure2-1.png", + "caption": "Fig. 2. Flux tubes indicating 3D flux paths of the fringing air-gap flux density in an FEA simulation.", + "texts": [ + " This effect is referred to as the fringing air-gap flux density and it is the main contributor to end-region losses in salient-pole synchronous machines. This field effect is proportional to the air-gap flux density\u2019s amplitude, discounting the impact of saturation on the end-region field distribution. The key influencing factors on the fringing air-gap flux density are the end-region geometry itself, most notably the axial rotor overhang, as well as the pole pitch. Larger pole pitches increase the fringing air-gap flux density. Fig. 2 shows this effect in a magnetostatic 3D FEA simulation. B. Calculation Approach The field calculation is based on numerically solved Schwarz\u2013Christoffel mappings (a type of conformal mapping) as pioneered by Reppe [12] and later improved and opensourced by Driscoll [13]\u2013[15]. Relying on conformal mappings has the advantage of being computationally inexpensive but comes with its own drawbacks. First and most importantly, only a 2D geometry can be considered. Therefore, we need to find a way to incorporate the 3D flux paths shown in Fig. 2 into the 2D representation of the end-region geometry (more on that later). Second, all flux lines are parallel to the surfaces. Therefore, we have to make one of two choices. A material can either have the same permeability as air or be perfectly permeable \u00b5r \u2192 \u221e. The stator and rotor core as well as the clamping plate are made from magnetic steel and are assumed to be perfectly permeable. This leaves the pressure fingers at the permeability of air. Lastly, there is no way to incorporate the impact of eddy currents on the fringing air-gap flux density in the field calculation", + " The Schwarz\u2013Christoffel mapping now takes an infinite strip (canonical domain) and maps it to the desired geometry (physical domain) using the mapping function f . Details on how this works can be found in a previous paper [1]. Between the two sides of the infinite strip, the amplitude of the fundamental spatial harmonic of the air-gap flux density B\u03b4 is assumed. The air-gap flux density is then mapped into the physical domain. C. 3D Effects: Flux Sink Concept The key issue with using a 2D representation of the endregion to calculate the fringing air-gap flux density is that it is intrinsically a 3D effect as can be seen in Fig. 2. Therefore, the flux sink concept is introduced. The idea behind this concept is to add an additional edge to the 2D geometry that absorbs those flux lines that would have taken 3D paths. Authorized licensed use limited to: China Jiliang University. Downloaded on June 30,2021 at 09:40:22 UTC from IEEE Xplore. Restrictions apply. 0018-9464 (c) 2021 IEEE. Personal use is permitted, but republication/redistribution requires IEEE permission. See http://www.ieee.org/publications_standards/publications/rights/index" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_35_0000506_tia.2021.3088768-Figure10-1.png", + "original_path": "designv11-35/openalex_figure/designv11_35_0000506_tia.2021.3088768-Figure10-1.png", + "caption": "Fig. 10. Experimental setup.", + "texts": [ + " 9, where the PI coefficients are as same as those in Fig. 8. The PI regulator with additional gains exhibits wider bandwidth in Fig. 9(a), which indicates fast current response. Nevertheless, the high close-loop gain, especially near 500Hz and -500Hz, will cause pulsating current in dynamics. In Fig. 9(b), the PI regulator with additional gains shows less capability to reject the disturbance, and thus will suffer from more current harmonics compared with the proposed method. Experiments are conducted on a DTP PMSM system as shown in Fig. 10. The parameters of the prototype machine and drive system are given in TABLE I, Appendix A. A PM dc machine is mechanically coupled to the test DTP PMSM and the output of dc machine is connected to an adjustable resistor to be served as the load. Fig. 11 shows the block diagram of the overall control system. The inner loops of the control strategy include two current loops shown in Fig. 3. The current references \u2217 and \u2217 are configured to zero to suppress the current harmonics in z1z2 subspace, and the current references \u2217 and \u2217 are determined by the outer speed loop" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_35_0001419_978-94-007-7401-8-Figure5.78-1.png", + "original_path": "designv11-35/openalex_figure/designv11_35_0001419_978-94-007-7401-8-Figure5.78-1.png", + "caption": "Fig. 5.78 2PaRRRR-1PaRRR-1CPaPat-type fully-parallel PMs with uncoupled Sch\u00f6nflies motions defined by MF = SF = 4, RF\u00f0 \u00de \u00bc \u00f0v1; v2; v3;xb\u00de, TF = 0, NF = 19, limb topology Pa\\R||R||R\\kR, Pa\\R||R||R and C||Pa||Pat", + "texts": [ + "4p) 10. 2PaPRRR1PaPRR1CPaPat (Fig. 5.74b) Pa||P\\R||R\\\\R (Fig. 5.3b) Idem No. 3 Pa||P\\R||R (Fig. 5.2f) C||Pa||Pat (Fig. 5.4p) 11. 2PaRPRR1PaRPR1CPaPa (Fig. 5.75) Pa\\R\\P\\kR\\R (Fig. 5.3c) Idem No. 3 Pa\\R\\P\\kR (Fig. 5.2g) C||Pa||Pa (Fig. 5.40) 12. 2PaRPRR1PaRPR1CPaPat (Fig. 5.76) Pa\\R\\P\\kR\\R (Fig. 5.3c) Idem No. 3 Pa\\R\\P\\kR (Fig. 5.2g) C||Pa||Pat (Fig. 5.4p) 13. 2PaRRRR1PaRRR1CPaPa (Fig. 5.77) Pa\\R||R||R\\kR (Fig. 5.3d) Idem No. 3 Pa\\R||R||R (Fig. 5.2h) C||Pa||Pa (Fig. 5.40) 14. 2PaRRRR1PaRRR1CPaPat (Fig. 5.78) Pa\\R||R||R\\kR (Fig. 5.3d) Idem No. 3 Pa\\R||R||R (Fig. 5.2h) C||Pa||Pat (Fig. 5.4p) 15. 3PaPPaR1CPaPa (Fig. 5.79a) Pa||P\\Pa\\kR (Fig. 5.4a) The revolute joint between links 6D and 9 of G4 limb and the last revolute joints of limbs G1, G2 and G3 have superposed axes Pa||P\\Pa||R (Fig. 5.4d) C||Pa||Pa (Fig. 5.40) 16. 3PaPPaR1CPaPa (Fig. 5.79b) Pa\\P\\\\Pa\\kR (Fig. 5.4b) Idem No. 15 Pa\\P\\\\Pa||R (Fig. 5.4c) C||Pa||Pa (Fig. 5.40) 17. 3PaPPaR1CPaPat (Fig. 5.80a) Pa||P\\Pa\\kR (Fig. 5.4a) Idem No. 15 Pa||P\\Pa||R (Fig" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_35_0000132_lra.2021.3062583-Figure2-1.png", + "original_path": "designv11-35/openalex_figure/designv11_35_0000132_lra.2021.3062583-Figure2-1.png", + "caption": "Fig. 2. The flexible wrist of the percutaneous robot consisting of the SMA spring, DF, and PF.", + "texts": [ + " The long tubes act as channels (refer to Fig. 1(c)) for the passage of actuation cables (Osaka Coat Rope Co., Ltd., Japan) with 0.21 mm diameter. This is a low-cost yet effective solution to compensate for the inability of commonly accessible manufacturing techniques to produce sub-millimeter diameter channels with large aspect ratios (>100). These channels serve as dedicated pathways for the cables, preventing cable tangling and more importantly, keeping the center of the lumen clear. As shown in Fig. 2, the SMA spring in the flexible wrist has a length ls of 6 mm, mean coil radius r of 0.875 mm, spring wire diameter d of 0.75 mm, number of coil n = 4, and a lumen diameter of 1 mm. It can convert its material phase between the low stiffness martensite and the high stiffness austenite through temperature variation. Heating it beyond the austenite finish temperature of around 45 \u25e6C raises its stiffness to facilitate the percutaneous needle insertion process while cooling it down below the martensite finish temperature allows it to have reduced stiffness and behave like a flexible spring" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_35_0001419_978-94-007-7401-8-Figure3.35-1.png", + "original_path": "designv11-35/openalex_figure/designv11_35_0001419_978-94-007-7401-8-Figure3.35-1.png", + "caption": "Fig. 3.35 Overactuated PMs with coupled Sch\u00f6nflies motions of types 4RPRRR (a) and 4RRRPR (b) defined by MF \u00bc SF \u00bc 4; RF\u00f0 \u00de \u00bc v1; v2; v3;xd\u00f0 \u00de, TF = 0, NF = 2, limb topology R\\P\\kR\\R||R (a) and R||R\\R||P||R (b)", + "texts": [ + " 4RRPRR (Fig. 3.33a) R\\R\\P\\kR\\R (Fig. 3.3m) Idem No. 14 27. 4RRRPR (Fig. 3.33b) R\\R||R\\P\\kR (Fig. 3.3n) Idem No. 14 28. 4RRRRP (Fig. 3.34a) R\\R||R\\R\\P (Fig. 3.3o) The first and the fourth revolute joints of the four limbs have parallel axes 29. 4PRRRR (Fig. 3.34b) P\\R||R\\R||R (Fig. 3.3p) The second joints of the four limbs have parallel axes (continued) 244 3 Overactuated Topologies with Coupled Sch\u00f6nflies Motions Table 3.2 (continued) No. PM type Limb topology Connecting conditions 30. 4RPRRR (Fig. 3.35a) R\\P\\kR\\R||R (Fig. 3.3q) Idem No. 16 31. 4RRRPR (Fig. 3.35b) R||R\\R||P||R (Fig. 3.3r) Idem No. 16 32. 4RRRPR (Fig. 3.36a) R||R\\R||P||R (Fig. 3.3s) Idem No. 16 33. 4RRRRP (Fig. 3.36b) R||R\\R||R||P (Fig. 3.3t) Idem No. 16 34. 4RRRRP (Fig. 3.37a) R||R\\R||R\\P (Fig. 3.3u) Idem No. 16 35. 4PRRRP (Fig. 3.37b) P\\R\\R||R\\P (Fig. 3.3v) Idem No. 29 36. 4RRPRP (Fig. 3.38a) R\\R||P||R\\P (Fig. 3.3w) Idem No. 16 37. 4RPRRP (Fig. 3.38b) R\\P||R||R\\P (Fig. 3.3x) Idem No. 16 38. 4RRPRP (Fig. 3.39a) R\\R||P||R\\P (Fig. 3.3y) Idem No. 16 39. 4RRRPP (Fig. 3.39b) R\\R||R||P\\P (Fig" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_35_0000395_978-3-030-67750-3_16-Figure9.14-1.png", + "original_path": "designv11-35/openalex_figure/designv11_35_0000395_978-3-030-67750-3_16-Figure9.14-1.png", + "caption": "Fig. 9.14 a Mesh of original component, b mesh of optimized component", + "texts": [ + " NASA is exploring the opportunity of 3D printing in space for tool manufacturing as well to manufacture objects which previously could not be launched to space (Harbaugh 2015). TWI employed the LMD process to manufacture a helicopter engine combustion chamber as shown in Fig. 9.13. The component consists of overhanging geometries but it was built without support structures by utilizing the 5-axes of the LMD printer. The thin walled part showed a density of more than 99.5%. The part was built in 7.5 h with 70% powder efficiency (Hauser 2014). In academic literature Seabra et al. (2016) optimized the topology of an aircraft bracket (Fig. 9.14) to bemanufactured using SLM. Compared to the original part, the new part had 54% reduced material volume and weigh 28% less though the material was changed from aluminum to titanium which resulted in increased factor of safety by 2. In medical industry customization is really favored as the products must be tailored fitted for each patient, available on demand and at a reasonable price. Additive manufacturing fulfills all these demand, thus it is highly preferred by the medical industry. Additivemanufacturing offers patient-specific parts which are strong and lightweight consisting of lattice structures as shown in Fig" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_35_0001572_978-3-319-05371-4_7-Figure7.3-1.png", + "original_path": "designv11-35/openalex_figure/designv11_35_0001572_978-3-319-05371-4_7-Figure7.3-1.png", + "caption": "Fig. 7.3 Longitudinal load transfer force distribution", + "texts": [ + "6) where m, Iz are, respectively, the mass and yaw moment of inertia of the vehicle, Iw is the rotational inertia of each wheel about its axis, R is the effective tire radius, and f , r are, respectively, the distances of the front and rear axles from the vehicle center of mass. In (7.1)\u2013(7.6) Fi j (i = x, y; j = f, r ) denote the longitudinal and lateral force components developed by the tires, defined in a tire-fixed reference frame. These forces depend on the normal loads on the front and rear axles, Fz f and Fzr , given by Fz f = mg r \u2212 hmg\u03bcxr f + r + h(\u03bcx f cos \u03b4 \u2212 \u03bcy f sin \u03b4 \u2212 \u03bcxr ) , (7.7) Fzr = mg f + hmg(\u03bcx f cos \u03b4 \u2212 \u03bcy f sin \u03b4) f + r + h(\u03bcx f cos \u03b4 \u2212 \u03bcy f sin \u03b4 \u2212 \u03bcxr ) (7.8) where h is the distance of the vehicle center of mass from the ground (see Fig. 7.3), and where \u03bc j = D sin(C arctan(Bs j )), \u03bci j = \u2212(si j/s j )\u03bc j , i = x, y; j = f, r, (7.9) for some constants C, B and D. Expression (7.9) is a simplified version of the well-known Pacejka \u201cMagic Formula\u201d (MF) [24] for the tire friction modeling, and combines the longitudinal and lateral motion, thus intrinsically incorporating the non-linear effect of the lateral/longitudinal coupling also known as the \u201cfriction circle\u201d (see Fig. 7.4), according to which, the constraint F2 x, j + F2 y, j \u2264 F2 max, j = (\u03bc j Fz, j ) 2 ( j = f, r) couples the allowable values of longitudinal and lateral tire friction forces" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_35_0001419_978-94-007-7401-8-Figure6.42-1.png", + "original_path": "designv11-35/openalex_figure/designv11_35_0001419_978-94-007-7401-8-Figure6.42-1.png", + "caption": "Fig. 6.42 3PPPaR-1CPaPa-type maximaly regular fully-parallel PMs with Sch\u00f6nflies motions defined by MF = SF = 4, (RF) = (v1; v2; v3;xb), TF = 0, NF = 21, limb topology P\\P\\||Pa||R, C||Pa||Pa and P\\P\\||Pa\\||R (a), P\\P\\||Pa ??R (b)", + "texts": [ + "10b) R||Pa||Pat||P (Fig. 5.4l) Idem no. 1 19. 3PPaPaR-1RPPaPa (Fig. 6.40) P||Pa||Pa\\R (Fig. 4.10a) P||Pa||Pa||R (Fig. 4.10b) R||P||Pa||Pa (Fig. 5.4m) Idem no. 9 20. 3PPaPaR-1RPaPatP (Fig. 6.41) P||Pa||Pa\\R (Fig. 4.10a) P||Pa||Pa||R (Fig. 4.10b) R||P||Pa||Pat (Fig. 5.4n) Idem no. 9 Table 6.14 Limb topology and connecting conditions of the fully-parallel solutions with no idle mobilities presented in Figs. 6.42, 6.43, 6.44, 6.45, 6.46 No. PM type Limb topology Connecting conditions 1. 3PPPaR-1CPaPa (Fig. 6.42a) P\\P\\||Pa\\||R (Fig. 4.8b) P\\P\\||Pa||R, (Fig. 4.8c) C||Pa||Pa (Fig. 5.4o) The last revolute joints of limbs G1, G2 and G3 and the revolute joint between links 6D and 8 have superposed axes. The actuated prismatic joints of limbs G1, G2 and G3 have orthogonal directions 2. 3PPPaR-1CPaPa (Fig. 6.42b) P\\P\\||Pa\\\\R (Fig. 4.8a) P\\P\\||Pa||R, (Fig. 4.8c) C||Pa||Pa (Fig. 5.4o) Idem no. 1 (continued) 6.3 Fully-Parallel Topologies with Complex Limbs 643 Table 6.14 (continued) No. PM type Limb topology Connecting conditions 3. 3PPPaR-1CPaPat (Fig. 6.43a) P\\P\\||Pa\\||R (Fig. 4.8b) P\\P\\||Pa||R, (Fig. 4.8c) C||Pa||Pat (Fig. 5.4p) Idem no. 1 4. 3PPPaR-1CPaPat (Fig. 6.43b) P\\P\\||Pa\\R (Fig. 4.8a) P\\P\\||Pa||R, (Fig. 4.8c) C||Pa||Pat (Fig. 5.4p) Idem no. 1 5. 3PPaPR-1CPaPa (Fig. 6.44a) P||Pa\\P\\\\R (Fig. 4.8d) P||Pa\\P\\||R (Fig" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_35_0001419_978-94-007-7401-8-Figure3.111-1.png", + "original_path": "designv11-35/openalex_figure/designv11_35_0001419_978-94-007-7401-8-Figure3.111-1.png", + "caption": "Fig. 3.111 4PaRRRR-type overactuated PM with coupled Sch\u00f6nflies motions defined by MF \u00bc SF \u00bc 4; RF\u00f0 \u00de \u00bc v1; v2; v3;xd\u00f0 \u00de, TF = 0, NF = 14, limb topology Pa||R\\R||R\\kR", + "texts": [ + "51q) The revolute joints of the parallelogram loops connecting the four limbs to the fixed base have parallel axes 18. 4CPaPa (Fig. 3.106) C||Pa||Pa (Fig. 3.51r) The cylindrical joints of the four limbs have parallel axes 19. 4PaPaC (Fig. 3.107) Pa||Pa||C (Fig. 3.51s) Idem No. 18 20. 4PaCPa (Fig. 3.108) Pa||C||Pa (Fig. 3.51t) Idem No. 18 21. 4RPaRRR (Fig. 3.109) R\\Pa\\kR\\R||R (Fig. 3.52a) Idem No. 13 22. 4RRRRPa (Fig. 3.110a) R||R\\R||R||Pa (Fig. 3.52b) Idem No. 15 23. 4RRRRPa (Fig. 3.110b) R\\R||R\\R||Pa (Fig. 3.52d) Idem No. 15 24. 4PaRRRR (Fig. 3.111) Pa||R\\R||R\\kR (Fig. 3.52c) Idem No. 13 25. 4PaRRRR (Fig. 3.112) Pa||R||R\\R||R (Fig. 3.52e) Idem No. 17 26. 4RRRPaR (Fig. 3.113) R||R\\R\\Pa\\kR (Fig. 3.52f) Idem No. 12 27. 4RRRRPa (Fig. 3.114) R||R\\R||R\\kPa (Fig. 3.52g) Idem No. 12 28. 4PRRRPa (Fig. 3.115) P\\R\\R||R\\kPa (Fig. 3.52h) The second joints of the four limbs have parallel axes 29. 4RRPRPa (Fig. 3.116) R\\R||P||R\\kPa (Fig. 3.52i) Idem No. 12 30. 4RPRRPa (Fig. 3.117) R\\P||R||R\\kPa (Fig. 3.52j) Idem No. 12 31. 4RRPRPa (Fig. 3.118) R\\R||P||R\\kPa (Fig" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_35_0000089_tmag.2021.3060767-Figure1-1.png", + "original_path": "designv11-35/openalex_figure/designv11_35_0000089_tmag.2021.3060767-Figure1-1.png", + "caption": "Fig. 1. Schematic of the water pipeline with UWEH containing direct-drive PMSG with propeller and analyzed slot shapes for final slot selection.", + "texts": [ + " Finally, Manuscript received November 30, 2020; revised January 28, 2021; accepted February 17, 2021. Date of publication February 22, 2021; date of current version May 17, 2021. Corresponding author: J. Chang (e-mail: cjhwan@dau.ac.kr). Color versions of one or more figures in this article are available at https://doi.org/10.1109/TMAG.2021.3060767. Digital Object Identifier 10.1109/TMAG.2021.3060767 a prototype is developed to verify the performance of the proposed FCPMSG under both no-load and load conditions. The schematic of the proposed FCPMSG-based UWEH system is shown in Fig. 1. The system-level design constraints for the proposed system are listed in Table I. As shown in Fig. 1, the FCPMSG in the UWEH system uses a hydraulic bearing system where the rotor of the PMSG contains air chambers on the back yoke to make it a buoyant rotor. The PMSG casing chamber is filled with fluid to create the buoyant force. For the hydraulic bearing, a balanced radial force topology should be selected with high electromagnetic performance. In addition, the proposed direct-drive FCPMSG design should ensure high efficiency and low loss, and low noise and vibration as it will be installed in the urban water pipeline network", + " Based on the topology comparison performed in [6], the 48Slot-40Pole (base 12Slot-10Pole) configuration is selected which uses AA BB CC fractional slot concentrated winding (FSCW) and ensures a high power density, low torque ripple, and losses. As the UWEH system will be installed around the water pipeline, the variation in water flow can result in mechanical ripples and vibration into the PMSG. Thus, to minimize the mechanical disturbance created by the variation in water flow as shown in the enlarged view of the propeller hub in Fig. 1, leaf springs are installed at the pre-stator section to damp out the impact created by the water flow. Considering the geometric restriction mentioned in Table I, the proposed FCPMSG is a rim-driven PMSG model. Thus, the short endwinding length of the FSCW is suitable. However, it results in high spatial MMF harmonics. Furthermore, as shown in Fig. 1, the FCPMSG uses a buoyant rotor which is levitated in the fluid-filled environment with the help of d-axis current 0018-9464 \u00a9 2021 IEEE. Personal use is permitted, but republication/redistribution requires IEEE permission. See https://www.ieee.org/publications/rights/index.html for more information. Authorized licensed use limited to: UNIVERSITY OF WESTERN ONTARIO. Downloaded on May 26,2021 at 09:28:21 UTC from IEEE Xplore. Restrictions apply. control [7]. Thus, it is important to have a balanced radial force over the machine bore and an analysis of the machine loading is important", + " , Ns is the permeance harmonic orders, and Ns is the number of stator slots. For the ease of manufacturing, usually in large renewable energy generators, the stators are designed with open slots. However, for the proposed FCPMSG with FSCW, it will introduce severe slot harmonic components in addition to the MMF harmonics produced by the FSCW. Therefore, for the optimum slot shape selection for the proposed FCPMSG, this section investigates three different slot types, namely, semi-open, open slot, and open slot with wedge (\u03bcr = 5), respectively, as shown in Fig. 1. The aim is to understand the slot shape effect on the airgap magnetic flux density harmonics and the radial force. The spatial distribution of the radial and tangential components of the airgap flux density harmonic components at \u03c9t = 0, under the three different simulation conditions, namely, no-load, only current, and load condition (Iq = Irated, Id = 0) is presented in Figs. 2 and 3, respectively. MST method was used to analyze the effect of these components on the radial force and vibration" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_35_0000165_s40430-021-02894-w-Figure15-1.png", + "original_path": "designv11-35/openalex_figure/designv11_35_0000165_s40430-021-02894-w-Figure15-1.png", + "caption": "Fig. 15 Contact stresses of the pinion for the design of CASE 1", + "texts": [ + " However, the cutter pitch radius has almost no effect on the evolution of the maximum contact stresses for steady contact region \u2018II\u2019 in Fig.\u00a010. The curvilinear gear set with tip relief can efficiently eliminate the severe contact stress caused by the edge contact, and the evolution of maximum contact stresses has almost no connection with the cutter pitch radius. As shown in Fig.\u00a014b, the tip relief and the cutter pitch radius have almost no effect on the evolution of the maximum bending stresses throughout the meshing cycle. Figure\u00a015 shows the contact stresses of the pinion for the design of CASE 1. Figure\u00a015a, c shows that edge contact appears at the addendum of the pinion because tip relief was not applied. Figure\u00a015b, d shows that the edge contact is eliminated by the tip relief. Figure\u00a015a is compared with Fig.\u00a015b. When the tip relief was applied, the maximum contact stress was reduced by 71% at contact position 8. Figure\u00a015c is compared with Fig.\u00a015d. When the tip relief was applied, the maximum contact stress was reduced by 73% at contact position 10. Figure\u00a016 shows the von Mises stress distribution on the fillet surface of the pinion for CASE 1 for different contact positions. The maximum bending stress at each contact position is symmetrically distributed at both ends of the gear. The maximum bending stress appears at contact position 11. This means that the bending stress is the peak value when tooth contact occurs at the pitch cylinder of the gear tooth surface" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_35_0000775_j.engfailanal.2021.105672-Figure5-1.png", + "original_path": "designv11-35/openalex_figure/designv11_35_0000775_j.engfailanal.2021.105672-Figure5-1.png", + "caption": "Fig. 5. Initial fatigue cracks.", + "texts": [ + " The post-accidental visual inspection of the failed landing gear revealed that the aft cross tube was fractured into two pieces. The broken, separated part was removed and taken for examination. Fig. 3 shows fractured aft cross tube and fracture surface. Visual analysis of fractured surface performed on macroscopic level, Fig. 3., revealed that the fracture was initiated at the bottom, position 6o\u2019clock, on the inner surface of the tube. The characteristic beach marks are observed indicating fatigue fracture (areas A in the Fig. 4 and Fig. 5), where crack propagated up to a half of the wall thickness. This region is clearly separate from rougher region with V-shape chevron marks (area B in the Fig. 4). Tracing back the observed chevron marks, the fracture origin area was also indicated to be in the sector A of fracture surface. The region opposite to cracks initiation area (marked area C in the Fig. 4) is the final, fast fracture area. Macroscopic examination of the aft cross tube surface revealed numerous deep scratches as the inner surface defects. These preexisting cracks are identified to be in the circumferential and in the longitudinal directions, where some of them are shown in Fig. 6. In Fig. 7 one of the initiation places, is shown which is presented in Fig. 5. Crack initiation has semi elliptical shape, approx. 2.3 mm and 0.7 mm deep in the tube wall (red arrow). Further propagation is revealed by presence of typical ratchet marks (yellow arrows). The crack was initiated in the inner side of the cross tube and propagated through tube wall. Observation of crack propagation area at low magnification by SEM, revealed a numerous secondary cracks (yellow arrows in Fig. 8). The rough mate region identified as overload (Fig. 9) revealed transgranular fracture surface, with secondary cracking on coarse second phase particles (zone C, Fig", + " The static deflections of the helicopter skid landing gear, when a helicopter is parked on the ground, is such that the aft cross tube is exposed to bending load, inducing tension on the underside and compression on the upper side. Resulting stress field of the critical cross-section is shown at Fig. 15. The fractographic analysis revealed features associated with progressive propagation under initial fatigue on approximately 50% cross-section of the wall tube. The fatigue was initiated at two pre-existing cracks on the inner surface of the aft cross tube (Fig. 5). Identified beach, ratchet and chevron marks indicate initiation sites and cracks growth directions. The presence of two initial sites (multiple origins) indicates that cross tube operated at high nominal stress. When the fatigue crack reached the critical size, fast fracture occurred as the last phase. Macroscopic inspection on the inner surface of the cross tube revealed the significant presence of large cracks in the circumferential direction, as well as other surface unacceptable defects (Figs, 7\u20139)" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_35_0001419_978-94-007-7401-8-Figure5.54-1.png", + "original_path": "designv11-35/openalex_figure/designv11_35_0001419_978-94-007-7401-8-Figure5.54-1.png", + "caption": "Fig. 5.54 2PaPaRRR-1PaPaRR-1RPaPatP-type fully-parallel PM with uncoupled Sch\u00f6nflies motions defined by MF = SF = 4, RF\u00f0 \u00de \u00bc \u00f0v1; v2; v3;xb\u00de, TF = 0, NF = 28, limb topology Pa\\Pa||R||R\\kR, Pa\\Pa||R||R and R||Pa||Pat||P", + "texts": [ + "51) Pa\\Pa||R||R\\\\R (Fig. 5.5a) Idem No. 1 Pa\\Pa||R||R (Fig. 5.4i) R||Pa||Pa||P (Fig. 5.4k) (continued) 532 5 Topologies with Uncoupled Sch\u00f6nflies Motions Table 5.4 (continued) No. PM type Limb topology Connecting conditions 18. 2PaPaRRR-1PaPaRR1RPaPaP (Fig. 5.52) Pa\\Pa||R||R\\kR (Fig. 5.5b) Idem No. 1 Pa\\Pa||R||R (Fig. 5.4i) R||Pa||Pa||P (Fig. 5.4k) 19. 2PaPaRRR-1PaPaRR1RPaPatP (Fig. 5.53) Pa\\Pa||R||R\\\\R (Fig. 5.5a) Idem No. 1 Pa\\Pa||R||R (Fig. 5.4i) R||Pa||Pat||P (Fig. 5.4l) 20. 2PaPaRRR-1PaPaRR1RPaPatP (Fig. 5.54) Pa\\Pa||R||R\\kR (Fig. 5.5b) Idem No. 1 Pa\\Pa||R||R (Fig. 5.4i) R||Pa||Pat||P (Fig. 5.4l) 21. 2PaPaRRR-1PaPaRR1RPPaPa (Fig. 5.55) Pa\\Pa||R||R\\\\R (Fig. 5.5a) The revolute joint between links 7D and 10 of G4 limb and the last revolute joints of limbs G1, G2 and G3 have superposed axes Pa\\Pa||R||R (Fig. 5.4i) R||P||Pa||Pa (Fig. 5.4m) 22. 2PaPaRRR-1PaPaRR1RPPaPa (Fig. 5.56) Pa\\Pa||R||R\\kR (Fig. 5.5b) Idem No. 21 Pa\\Pa||R||R (Fig. 5.4i) R||P||Pa||Pa (Fig. 5.4m) 23. 2PaPaRRR-1PaPaRR1RPPaPat (Fig. 5.57) Pa\\Pa||R||R\\\\R (Fig", + " The last joints of of limbs G1, G2 and G3 have superposed axes Pa\\Pa||Pa||R (Fig. 5.6b) R\\R\\R\\P\\kR\\R (Fig. 5.1b) 28. 3PaPaPaR-1RUPU (Fig. 5.62) Pa\\Pa||Pa\\\\R (Fig. 5.6c) Idem No. 27 Pa\\Pa||Pa||R (Fig. 5.6b) R\\R\\R\\P\\kR\\R (Fig. 5.1b) 29. 3PaPaPaR-1RPaPaP (Fig. 5.63) Pa\\Pa||Pa\\kR (Fig. 5.6a) Idem No. 1 Pa\\Pa||Pa||R (Fig. 5.6b) R||Pa||Pa||P (Fig. 5.4k) (continued) 5.1 Fully-Parallel Topologies 533 Table 5.4 (continued) No. PM type Limb topology Connecting conditions 30. 3PaPaPaR-1RPaPaP (Fig. 5.64) Pa\\Pa||Pa\\\\R (Fig. 5.6c) Idem No. 1 Pa\\Pa||Pa||R (Fig. 5.6b) R||Pa||Pa||P (Fig. 5.54k) 31. 3PaPaPaR1RPaPatP (Fig. 5.65) Pa\\Pa||Pa\\kR (Fig. 5.6a) Idem No. 1 Pa\\Pa||Pa||R (Fig. 5.6b) R||Pa||Pat||P (Fig. 5.4l) 32. 3PaPaPaR1RPaPatP (Fig. 5.66) Pa\\Pa||Pa\\\\R (Fig. 5.6c) Idem No. 1 Pa\\Pa||Pa||R (Fig. 5.6b) R||Pa||Pat||P (Fig. 5.54l) 33. 3PaPaPaR-1RPPaPa (Fig. 5.67) Pa\\Pa||Pa\\kR (Fig. 5.6a) The revolute joint between links 7D and 11 of G4 limb and the last revolute joints of limbs G1, G2 and G3 have superposed axes Pa\\Pa||Pa||R (Fig. 5.6b) R||P||Pa||Pa (Fig. 5.4m) 34. 3PaPaPaR-1RPPaPa (Fig. 5.68) Pa\\Pa||Pa\\\\R (Fig. 5.6c) Idem No. 33 Pa\\Pa||Pa||R (Fig. 5.6b) R||P||Pa||Pa (Fig. 5.54m) Table 5.5 Limb topology and connecting conditions of the fully-parallel solutions with no idle mobilities presented in Figs. 5.69, 5.70, 5.71, 5.72, 5.73, 5.74, 5.75, 5.76, 5.77, 5.78, 5.79, 5.80, 5.81, 5.82, 5.83, 5.84, 5.85, 5.86, 5.87, 5.88, 5.89, 5.90, 5.91, 5.92 No. PM type Limb topology Connecting conditions 1. 3PaPaPaR1RPPaPat (Fig. 5.69) Pa\\Pa||Pa\\kR (Fig. 5.6a) The revolute joint between links 7D and 11 of G4 limb and the last revolute joints of limbs G1, G2 and G3 have superposed axes Pa\\Pa||Pa||R (Fig. 5.6b) R||P||Pa||Pat (Fig. 5.4n) 2. 3PaPaPaR1RPPaPat (Fig. 5.70) Pa\\Pa||Pa\\\\R (Fig. 5.6c) Idem No. 1 Pa\\Pa||Pa||R (Fig. 5.6b) R||P||Pa||Pat (Fig. 5.54n) 3. 3PaPPR1CPaPa (Fig. 5.71a) Pa||P\\P\\kR (Fig. 5.2a) The revolute joint between links 6D and 8 of G4 limb and the last revolute joints of limbs G1, G2 and G3 have superposed axes Pa||P\\P\\\\R (Fig. 5.2d) C||Pa||Pa (Fig. 5.4o) 4. 3PaPPR1CPaPa (Fig. 5.71b) Pa\\P\\kP\\kR (Fig. 5.2b) Idem No. 3 Pa\\P\\kP\\\\R (Fig. 5.2c) C||Pa||Pa (Fig. 5.40) 5. 3PaPPR1CPaPat (Fig. 5.72a) Pa||P\\P\\kR (Fig. 5.2a) Idem No. 3 Pa||P\\P\\\\R (Fig. 5.2d) C||Pa||Pat (Fig. 5.4p) 6. 3PaPPR1CPaPat (Fig. 5.72b) Pa\\P\\kP\\kR (Fig. 5.2b) Idem No", + " 23 Pa\\Pa||R||R (Fig. 5.4i) C||Pa||Pat (Fig. 5.4p) 26. 2PaPaRRR1PaPaRR1CPaPat (Fig. 5.88) Pa\\Pa||R||R\\kR (Fig. 5.5b) Idem No. 23 Pa\\Pa||R||R (Fig. 5.4i) C||Pa||Pat (Fig. 5.4p) 27. 3PaPaPaR1CPaPa (Fig. 5.89) Pa\\Pa||Pa\\kR (Fig. 5.6a) The revolute joint between links 6D and 11 of G4 limb and the last revolute joints of limbs G1, G2 and G3 have superposed axes Pa\\Pa||Pa||R (Fig. 5.6b) C||Pa||Pa (Fig. 5.4o) 28. 3PaPaPaR1CPaPa (Fig. 5.90) Pa\\Pa||Pa\\\\R (Fig. 5.6c) Idem No. 27 Pa\\Pa||Pa||R (Fig. 5.6b) C||Pa||Pa (Fig. 5.54o) 29. 3PaPaPaR1CPaPat (Fig. 5.91) Pa\\Pa||Pa\\kR (Fig. 5.6a) Idem No. 27 Pa\\Pa||Pa||R (Fig. 5.6b) C||Pa||Pat (Fig. 5.4p) 30. 3PaPaPaR1CPaPat (Fig. 5.92) Pa\\Pa||Pa\\\\R (Fig. 5.6c) Idem No. 27 Pa\\Pa||Pa||R (Fig. 5.6b) C||Pa||Pat (Fig. 5.54p) 536 5 Topologies with Uncoupled Sch\u00f6nflies Motions No. Structural parameter Solution Figure 5.7 Figures 5.8, 5.9, 5.10 Figures 5.11 1. m 20 22 22 2. pi (i = 1, 3) 7 8 7 3. p2 7 7 7 4. p4 4 4 6 5. p 25 27 27 6. q 6 6 6 7. k1 1 1 1 8. k2 3 3 3 9. k 4 4 4 10. (RG1) (v1; v2; v3;xb) (v1; v2; v3;xa;xb) (v1; v2; v3;xb) 11. (RG2) (v1; v2; v3;xb) (v1; v2; v3;xb) (v1; v2; v3;xb) 12. (RG3) (v1; v2; v3;xb) (v1; v2; v3;xb;xd) (v1; v2; v3;xb) 13. (RG4) (v1; v2; v3;xb) (v1; v2; v3;xb) (v1; v2; v3;xa;xb;xd) 14" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_35_0000529_012088-Figure14-1.png", + "original_path": "designv11-35/openalex_figure/designv11_35_0000529_012088-Figure14-1.png", + "caption": "Figure 14. Final Position of the Robot in XYZ Plane after implementing 6th order polynomial method (Left side)", + "texts": [], + "surrounding_texts": [ + "It should be noticed that compared to Cartesian scheme, joint space scheme gives smoother trajectory. Cubic polynomial works faster than 5th and 6th order polynomial methods, but accuracy is lesser. 5 th order trajectory gets settled quickly but leaves some amount of steady state error. 6th order trajectory creates a little amount of overshoot, but accuracy is extremely good. 6 th order polynomial satisfy zero angular acceleration and angular velocity conditions at the boundaries. 6 th order method protects joint actuators from achieving instantaneous velocity and infinite acceleration. Compared to 5 th order and cubic polynomials, 6 th order polynomial gives lesser jerks and vibration to the joint links, as a result food can be safely deliver without being wasted. In future, trajectory planning can be done using S-curve method and linear segment parabolic blend for multiple via points. Also the number of degrees of freedom can be taken six to avoid robot from singularities. The optimization algorithm can also be used to develop inverse kinematics algorithm, which can have a single solution instead of multiple solution. ICASSCT 2021 Journal of Physics: Conference Series 1921 (2021) 012088 IOP Publishing doi:10.1088/1742-6596/1921/1/012088" + ] + }, + { + "image_filename": "designv11_35_0000797_9781119526483.ch14-Figure14.43-1.png", + "original_path": "designv11-35/openalex_figure/designv11_35_0000797_9781119526483.ch14-Figure14.43-1.png", + "caption": "Figure 14.43 Propulsion of diatoms via many small explosions (2020). Nuclear propulsion via hundreds of \u201csmall\u201d nuclear bombs: Project Orion (conceived by Stanislaw Ulam, 1946), an analog of diatom propulsion via explosive hydration of raphe fibrils. Top left: artist\u2019s conception [14.420]. Top right: Principle [14.315]: The \u201cPulse unit injection\u201d would correspond to a raphe carrying raphe fibrils that explosively hydrate on being exposed to water. Middle right: A space faring diatom [14.119] [14.124] Lyrella esul propelled by explosions [14.134]; Lyrella esul [14.338] with kind permission of David A. Siqueiros Beltrones. Bottom: Project Orion configuration [14.420], all others from NASA, in public domain).", + "texts": [ + " Our inability to time-resolve the jerky, high accelerations of single diatoms [14.313], first observed by Lesley Edgar (1979) [14.87] (Figure 14.9), has led to the notion that something very erratic and sudden is occurring as a diatom moves. We proposed that this could be due to explosive hydration of individual raphe fibrils as they leave the raphe and join the diatom trail, which led to our project to try to hear these predicted miniexplosions [14.420] [14.443]. An analogous mechanism was conceived for space travel by Stanislaw Ulam (1909- 1984) [14.422] in 1946 (Figure 14.43), with whom I later did postdoctoral research [14.118]: \u201cThe idea of using a series of explosive pulses to propel a rocket vehicle can be traced back to Hermann Ganswindt (1899)\u2026. A proposal for use of fission-based explosives was first made by Stanislaus Ulam in 1946\u2026. detonating a 20-kiloton nuclear device 10 meters away from two -1-meter-diameter, graphite coated steel spheres\u2026. [sent them] several kilometers from ground zero; the spheres were recovered, with only a few thousandths of an inch of graphite ablated from their surfaces\u2026" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_35_0001554_j.apenergy.2015.05.058-Figure1-1.png", + "original_path": "designv11-35/openalex_figure/designv11_35_0001554_j.apenergy.2015.05.058-Figure1-1.png", + "caption": "Fig. 1. (a) The schematic diagram of air\u2013cathode l-Liter MFC (A: medium inlet/ outlet). (b) The schematic diagram of micro-channel for fabricating microparticles. B: DADMAC monomer and microorganism mixture. C: soybean oil. D: KCl solution. E: waste oil. F: outlet for micro-particles.", + "texts": [ + " However, the diffusion length for solute to travel between electrode and bulk becomes longer in macroscale MFCs and the rate of diffusion can become much slower than surface reaction rate. It will be challenging to provide similar surface to volume Please cite this article in press as: Chen Y-Y, Wang H-Y. Polyelectrolyte microp bial fuel cell. Appl Energy (2015), http://dx.doi.org/10.1016/j.apenergy.2015.05 ratio or diffusion length in macroscale MFCs as those in microscale counterparts; therefore, reduced current density of macroscale MFC is anticipated. 2.1. Air\u2013cathode l-Liter MFC construction The schematic of the air\u2013cathode l-Liter MFC (lMFC) is shown in Fig. 1a. The air\u2013cathode containing a carbon/polytetrafluoroethy lene (PTFE) layer on the airside and a platinum catalyst layer on the waterside was fabricated as described by Cheng and Logan [20]. The anodal electrode was composed of a glass slide coated with a thin gold layer. The electrodes were connected to the external resistance by the silver enamel covered wire. The polydimethylsiloxane (PDMS) well having volume of 200 lL was prepared as the micro-liter chamber for anolyte. All components of the cell were connected together by glue", + " The mixed culture microorganisms were obtained from the seacoast of Taiwan and have been proven to have electro-activity in a lab-scale H-type MFC [17]. The medium for culturing microorganisms contained 2 g/L CH3COONa, 5.24 g/L NH4HCO3, 6.72 g/L NaHCO3, 0.125 g/L K2HPO4, 0.01388 g/L MnSO4 5H2O, 0.1 g/L MgCl2 6H2O, 0.025 g/L FeSO4 7H2O, 0.005 g/L CuSO4 5H2O and 0.000125 g/L CoCl2 6H2O. All reagents were purchased from Sigma\u2013Aldrich. The microparticles were generated using droplet microfluidic system shown in Fig. 1b. The microfluidic device was prepared articles for enhancing anode performance in an air\u2013cathode l-Liter micro.058 by standard soft-lithographic techniques [21] and the details can be found in our previous report [22]. Microdroplets contained 65 wt% monomer solution (DADMAC, abbrev. diallyldimethylammonium chloride), 0.014 g/mL photoinitiator, 0.014 g/mL cross-linker and 100 lL/mL microbial solution which was concentrated by 2 mL microbial suspension with optical density (OD) of 0.206 at 550 nm", + " 2 shows that microorganisms were successfully incorporated and evenly distributed in the particle. The average intensity of fluorescence in each particle is about 0.18 \u00b1 0.74 a.u. in Day 0. anisms on the anode taken in (a) Day 0 and (b) Day 1. ode taken in Day 7 with (a) 350 , (b) 2.21k , (c) 10k and (d) 35k . articles for enhancing anode performance in an air\u2013cathode l-Liter micro.058 0 0.05 0.1 0.15 0.2 0 1 2 3 4 5 6 7 Day Microparticle Biofilm The microparticles containing microorganisms were incubated in the air\u2013cathode lL-microbial fuel cell (lMFC, Fig. 1a) for examining their performance as anode. Another set of examination using identical lMFC setup but microorganism biofilm, instead of microorganisms in poly(DADMAC) microparticles, was also carried out. The output voltage of both lMFCs is shown in Fig. 3. The output voltage of the lMFC with microorganisms encapsulated in the microparticle was steady with an average value of 14.46 \u00b1 1.63 mV in the first 28 h. In the lMFC with microorganism biofilm, the output voltage increased rapidly in the first 12 h and decreased rapidly in the subsequent 16 h", + " After the 28th hour, the voltage output maintained relatively stable with an average value of 13.34 \u00b1 2.92 mV. The output voltage from microorganism biofilm in the first 28 h had higher variation than in the subsequently 140 h and this was probably caused by the developing biofilm. Conversely, the output voltage from microparticles remained relatively consistent along the whole 168 h. The unseeded ter MFC with microorganism biofilm taken in (a) Day 1 and (b) Day 7. microparticle and bare electrode barely produced voltage output (Supplementary Fig. 1). After the 28th hour, the output voltage of the lMFC with microparticles gradually increased and became much higher than the lMFC with microorganism suspension. The highest output voltage reached 22.81 \u00b1 3.71 mV in the lMFC with microparticles, while microorganism biofilm only produced 11.45 \u00b1 1.81 mV. This validates that the poly(DADMAC) microparticles have great potential in improving the performance of MFC. To investigate the role of poly(DADMAC) in enhancing the electricity output, its biocompatibility and ability in decreasing the resistance of anode were examined by microscopy, cyclic voltammetry, anode polarization curve, and EIS" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_35_0001419_978-94-007-7401-8-Figure2.38-1.png", + "original_path": "designv11-35/openalex_figure/designv11_35_0001419_978-94-007-7401-8-Figure2.38-1.png", + "caption": "Fig. 2.38 Fully-parallel PMs with coupled Sch\u00f6nflies motions of types 4CPPa (a) and 4PaPC (b) defined by MF \u00bc SF \u00bc 4; RF\u00f0 \u00de \u00bc \u00f0v1; v2; v3;xd\u00de, TF = 0, NF = 18, limb topology C\\P||Pa (a) and Pa||P\\C (b)", + "texts": [ + " 2.34a) Pa||P\\P||R (Fig. 2.21q) Idem No. 4 18. 4PaPPR (Fig. 2.34b) Pa\\P\\kP\\kR (Fig. 2.21r) Idem No. 4 19. 4PaRPP (Fig. 2.35a) Pa\\R||P\\P (Fig. 2.21s) Idem No. 11 20. 4PaPRP (Fig. 2.35b) Pa||P\\R||P (Fig. 2.21t) Idem No. 5 21. 4PaPRP (Fig. 2.36a) Pa\\P||R\\P (Fig. 2.21u) Idem No. 11 22. 4PaRPP (Fig. 2.36b) Pa\\R\\P\\kP (Fig. 2.21v) Idem No. 5 23. 4PCPa (Fig. 2.37a) P\\C\\Pa (Fig. 2.21w) Idem No. 5 24. 4PPaC (Fig. 2.37b) P||Pa\\C (Fig. 2.21x) The last joints of the four limbs have parallel axes 25. 4CPPa (Fig. 2.38a) C\\P||Pa (Fig. 2.21y) The first joints of the four limbs have parallel axes 26. 4PaPC (Fig. 2.38b) Pa||P\\C (Fig. 2.21z) Idem No. 24 27. 4PaCP (Fig. 2.39) Pa\\C\\P (Fig. 2.21z1) Idem No. 11 28. 4PPaPaR (Fig. 2.40) P\\Pa\\\\Pa\\kR (Fig. 2.22a) Idem No. 4 29. 4PPaRPa (Fig. 2.41) P\\Pa||R\\Pa (Fig. 2.22b) The axes of the revolute joints connecting the two parallelogram loops of the four limbs are parallel (continued) 2.2 Topologies with Complex Limbs 175 Table 2.3 (continued) No. PM type Limb topology Connecting conditions 30. 4PRPaPa (Fig. 2.42) P\\R||Pa\\Pa (Fig. 2.22c) Idem No. 5 31. 4RPPaPa (Fig. 2.43) R\\P\\kPa\\\\Pa (Fig" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_35_0001391_s00170-021-07282-1-Figure3-1.png", + "original_path": "designv11-35/openalex_figure/designv11_35_0001391_s00170-021-07282-1-Figure3-1.png", + "caption": "Fig. 3 Schematic diagram of HSMS radial loading", + "texts": [ + " Figure 2a shows the structure and composition of the contact loading device in detail, which is composed of a mechanical device and a measurement and control device. The mechanical device is mainly composed of a cylinder, floating joint, sliding pillar, sliding base and four rolling bearings. The sliding pillar and the sliding base fixed on the experimental rig constitute a sliding pair. The sliding pillar is pushed up by the cylinder (SMC CQ2L40-50D) so that the outer rings of the rolling bearings installed on the sliding base contact the rotor of the HSMS to achieve loading. A more detailed loading principle is shown in Fig. 3. The radial loading force is adjusted by the air pressure of the cylinder. The air pressure of the cylinder is provided by an air compressor in Fig. 2b. The maximum loading force can be calculated as 845 N by multiplying the maximum air pressure of the air compressor by 0.8 Mpa by the cylinder area 1056 mm2. To avoid mechanical interference caused by coaxiality errors during installation, a floating joint (SMC JA20-8-125) is used to connect the cylinder and force sensor. Oil\u2013air lubrication system not only has a lubricating effect but also takes away much heat via the air, which is used not only to lubricate and cool the HSMS bearings, but also to lubricate and cool the contact position between the rolling bearings and the loading rod" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_35_0000214_j.measurement.2021.109266-Figure1-1.png", + "original_path": "designv11-35/openalex_figure/designv11_35_0000214_j.measurement.2021.109266-Figure1-1.png", + "caption": "Fig. 1. Images of aluminium extrusion dies: a) front view of the insert solid die; b) assembled die sets.", + "texts": [ + " In this study, the solid die design consists of a front chamber, die, support, and assembled sets parts. A profile cross-section with a jagged and flat surface was used in order to monitor the behaviour of the solid die. The components of the die set, which do not affect the die surface and mechanical properties of the product (front chamber and supports), were manufactured using 1.2344 hot-work tool steel to make cheaper the total cheaper cost by using traditional manufacturing methods. Later, the insert die, which was manufactured with the DMLS process was assembled. Fig. 1 shows, the images of the insert solid die and the assembled dies are shown. Having been drafted, the solid die was converted to .stl format in order for the EOS DMLS device to process it by the tool path. The solid die was manufactured with a 0.04 mm standard layer thickness and the EOS parameter. After the part manufacturing was completed, the powders on the part were first cleaned in the manufacturing cabin. Then the support structures of the part were removed after the part had been cut by using a wire erosion device (Fig" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_35_0000797_9781119526483.ch14-Figure14.16-1.png", + "original_path": "designv11-35/openalex_figure/designv11_35_0000797_9781119526483.ch14-Figure14.16-1.png", + "caption": "Figure 14.16 Left: In 2015: \u201cWe propose a model [for gliding of Flavobacterium johnsoniae] in which a pinion, connected to a rotary motor, drives a rack (a tread) that moves along a spiral track fixed to the rigid framework of the cell wall. SprB [a cell-surface adhesin], carried by the tread, adsorbs to the substratum and causes the cell to glide\u2026. Tethered cells pinwheel around a fixed axis, suggesting that a rotary motor that generates high torque is a part of the gliding machinery [14.334]\u2026. If 90 nm is the radius of a pinion rotating 3Hz (the maximum speed of rotation when a cell is tethered) then that pinion can drive a rack (a tread carrying adhesins) at 1.5 \u00b5m/s, which is the speed that cells glide. This suggests that nature has not only invented the wheel, it also has invented the likes of a microscopic snowmobile\u2026. A model of the gliding machinery. (a) A cross-sectional view of a cell with a rotary gliding motor (blue), a mobile tread (green), a stationary track (red), and an adhesin (magenta). The rotary motor and the track are anchored to the peptidoglycan (PG), and the track is wound spirally around the cell. The rotary motor drives a pinion that engages a mobile tread (rack) that slides along the track. The adhesin, SprB, is attached to the tread and moves with it. The dimension d is the distance between the axis of rotation of the motor and the center of the track, and r is the radius of the pinion. (b) A side view of a cell with a rotary motor powering the motion of a tread carrying SprB\u201d [14.332] [14.333]. O.M. = outer membrane, C.M. = inner cell membrane. Right: A rack and pinion.", + "texts": [ + " He experimented with plasmolysis, which reversibly halted motility (but not necessarily per [14.79] [14.93] [14.141] [14.142] [14.158] [14.164] [14.177]), and mentions the \u201celegant\u201d centrifugation study of Fritz Legler & Hellmuth Schindler (1939) [14.211] which showed that the cytoplasm could be moved to one end of the cell and that 90% of the subsequent movements were in that direction. These experiments suggest that cell membrane contact with raphes is necessary for movement. A snowmobile-like tread has been proposed for the gliding motility of Flavobacterium [14.330] (Figure 14.16). The motors driving the gliding motility of Flavobacterium are membrane embedded structures that rotate and secrete protein, lectin or polysaccharide adhesins.\u00a0This prokaryote has abilities to reverse direction, pivot and flip [14.232], analogous to pennate diatoms [14.29]. The adhesins both rotate and move along the cell surface in a closed helical loop, leaving a polysaccharide \u201croad\u201d on the substrate [14.232], analogous to the diatom trail. What drives them along this track is still an open question [14.178], but now has a working hypothesis: \u201c\u2026we propose a model where a molecular rack and pinion-like assembly actuates the motion of SprB on the cell surface\u201d [14.331] (Figure 14.16). The rotation is driven by a proton-motive force [14.268], i.e., chemiosmosis [14.202]. One wonders whether the following 2019 observation means, contrary to the authors\u2019 interpretation, that we are not done considering a tank tread mechanism for diatom motility, or some variant in which they lay down and then immediately take up and reprocess their own trail, outdoing the broom-headed dog in Alice in Wonderland [14.75] (Figure 14.17): \u201cSurprisingly, we also observed that adhesive trails are not deposited behind every moving diatom cell, implying that diatom gliding does not require the deposition of a permanent adhesive but rather that temporary substrate adhesion is sufficient for gliding" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_35_0001419_978-94-007-7401-8-Figure5.83-1.png", + "original_path": "designv11-35/openalex_figure/designv11_35_0001419_978-94-007-7401-8-Figure5.83-1.png", + "caption": "Fig. 5.83 3PaPaPR-1CPaPat-type fully-parallel PM with uncoupled Sch\u00f6nflies motions defined by MF = SF = 4, RF\u00f0 \u00de \u00bc \u00f0v1; v2; v3;xb\u00de, TF = 0, NF = 30 limb topology Pa\\Pa\\kP\\\\R, Pa\\Pa\\kP\\kR and C||Pa||Pat", + "texts": [ + "4d) C||Pa||Pat (Fig. 5.4p) 18. 3PaPPaR1CPaPat (Fig. 5.80b) Pa\\P\\\\Pa\\kR (Fig. 5.4b) Idem No. 15 Pa\\P\\\\Pa||R (Fig. 5.4c) C||Pa||Pat (Fig. 5.4p) (continued) 5.1 Fully-Parallel Topologies 535 Table 5.5 (continued) No. PM type Limb topology Connecting conditions 19. 3PaPaPR1CPaPa (Fig. 5.81) Pa\\Pa\\kP\\\\R (Fig. 5.4f) Idem No. 15 Pa\\Pa\\kP\\kR (Fig. 5.4h) C||Pa||Pa (Fig. 5.4o) 20. 3PaPaPR1CPaPa (Fig. 5.82) Pa\\Pa\\\\P\\\\R (Fig. 5.4g) Idem No. 15 Pa\\Pa\\\\P\\kR (Fig. 5.4e) C||Pa||Pa (Fig. 5.40) 21. 3PaPaPR1CPaPat (Fig. 5.83) Pa\\Pa\\kP\\\\R (Fig. 5.4f) Idem No. 15 Pa\\Pa\\kP\\kR (Fig. 5.4h) C||Pa||Pat (Fig. 5.4p) 22. 3PaPaPR1CPaPat (Fig. 5.84) Pa\\Pa\\\\P\\\\R (Fig. 5.4g) Idem No. 15 Pa\\Pa\\\\P\\kR (Fig. 5.4e) C||Pa||Pat (Fig. 5.4p) 23. 2PaPaRRR1PaPaRR1CPaPa (Fig. 5.85) Pa\\Pa||R||R\\\\R (Fig. 5.5a) The revolute joint between links 6D and 10 of G4 limb and the last revolute joints of limbs G1, G2 and G3 have superposed axes Pa\\Pa||R||R (Fig. 5.4i) C||Pa||Pa (Fig. 5.40) 24. 2PaPaRRR1PaPaRR1CPaPa (Fig. 5.86) Pa\\Pa||R||R\\kR (Fig. 5.5b) Idem No" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_35_0001419_978-94-007-7401-8-Figure6.38-1.png", + "original_path": "designv11-35/openalex_figure/designv11_35_0001419_978-94-007-7401-8-Figure6.38-1.png", + "caption": "Fig. 6.38 3PPaPaR-1RPaPaP-type maximaly regular fully-parallel PM with Sch\u00f6nflies motions defined by MF = SF = 4, (RF) = (v1; v2; v3;xb), TF = 0, NF = 30, limb topology P||Pa||Pa\\R, P||Pa||Pa||R and R||Pa||Pa||P", + "texts": [ + " 3PPaPR-1RPPaPa (Fig. 6.34) P||Pa\\P\\R (Fig. 4.8d) P||Pa\\P\\||R (Fig. 4.8e) R||P||Pa||Pa (Fig. 5.4m) Idem no. 9 14. 3PPaPR-1RPPatPa (Fig. 6.35) P||Pa\\P\\\\R (Fig. 4.8d) P||Pa\\P\\||R (Fig. 4.8e) R||P||Pa||Pat (Fig. 5.4n) Idem no. 9 15. 2PPaRRR-1PPaRR-1RPPaPa (Fig. 6.36) P||Pa||R||R\\R (Fig. 4.9a) P||Pa||R||R (Fig. 4.8g) R||P||Pa||Pa (Fig. 5.4m) Idem no. 9 16. 2PPaRRR-1PPaRR-1RPPaPat (Fig. 6.37) P||Pa||R||R\\R (Fig. 4.9a) P||Pa||R||R (Fig. 4.8g) R||P||Pa||Pat (Fig. 5.4n) Idem no. 9 17. 3PPaPaR-1RPaPaP (Fig. 6.38) P||Pa||Pa\\R (Fig. 4.10a) P||Pa||Pa||R (Fig. 4.10b) R||Pa||Pa||P (Fig. 5.4k) Idem no. 1 18. 3PPaPaR-1RPaPatP (Fig. 6.39) P||Pa||Pa\\R (Fig. 4.10a) P||Pa||Pa||R (Fig. 4.10b) R||Pa||Pat||P (Fig. 5.4l) Idem no. 1 19. 3PPaPaR-1RPPaPa (Fig. 6.40) P||Pa||Pa\\R (Fig. 4.10a) P||Pa||Pa||R (Fig. 4.10b) R||P||Pa||Pa (Fig. 5.4m) Idem no. 9 20. 3PPaPaR-1RPaPatP (Fig. 6.41) P||Pa||Pa\\R (Fig. 4.10a) P||Pa||Pa||R (Fig. 4.10b) R||P||Pa||Pat (Fig. 5.4n) Idem no. 9 Table 6.14 Limb topology and connecting conditions of the fully-parallel solutions with no idle mobilities presented in Figs" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_35_0001419_978-94-007-7401-8-Figure5.17-1.png", + "original_path": "designv11-35/openalex_figure/designv11_35_0001419_978-94-007-7401-8-Figure5.17-1.png", + "caption": "Fig. 5.17 3PaPPR-1RPaPatP-type fully-parallel PMs with uncoupled Sch\u00f6nflies motions defined by MF = SF = 4, RF\u00f0 \u00de \u00bc \u00f0v1; v2; v3;xb\u00de, TF = 0, NF = 21, limb topology R||Pa||Pat||P and Pa||P\\P\\kR, Pa||P\\P\\\\R (a), Pa\\P\\kP\\kR, Pa\\P\\kP\\\\R (b)", + "texts": [ + "2d) R||Pa||Pa||P (Fig. 5.4k) 2. 3PaPPR-1RPaPaP (Fig. 5.15b) Pa\\P\\kP\\kR (Fig. 5.2b) Idem No. 1 Pa\\P\\kP\\\\R (Fig. 5.2c) R||Pa||Pa||P (Fig. 5.4k) 3. 3PaPPR-1RPPaPa (Fig. 5.16a) Pa||P\\P\\kR (Fig. 5.2a) The revolute joint between links 7D and 9 of G4 limb and the last revolute joints of limbs G1, G2 and G3 have superposed axes Pa||P\\P\\\\R (Fig. 5.2d) R||P||Pa||Pa (Fig. 5.4m) 4. 3PaPPR-1RPPaPa (Fig. 5.16b) Pa\\P\\kP\\kR (Fig. 5.2b) Idem No. 3 Pa\\P\\kP\\\\R (Fig. 5.2c) R||P||Pa||Pa (Fig. 5.4m) 5. 3PaPPR1RPaPatP (Fig. 5.17a) Pa||P\\P\\kR (Fig. 5.2a) Idem No. 1 Pa||P\\P\\\\R (Fig. 5.2d) R||Pa||Pat||P (Fig. 5.4l) 6. 3PaPPR1RPaPatP (Fig. 5.17b) Pa\\P\\kP\\kR (Fig. 5.2b) Idem No. 1 Pa\\P\\kP\\\\R (Fig. 5.2c) R||Pa||Pat||P (Fig. 5.4l) 7. 3PaPPR-1RPPaPat (Fig. 5.18a) Pa||P\\P\\kR (Fig. 5.2a) Idem No. 3 Pa||P\\P\\\\R (Fig. 5.2d) R||P||Pa||Pat (Fig. 5.4n) 8. 3PaPPR-1RPPaPat (Fig. 5.18b) Pa\\P\\kP\\kR (Fig. 5.2b) Idem No. 3 Pa\\P\\kP\\\\R (Fig. 5.2c) R||P||Pa||Pat (Fig. 5.4n) 9. 2PaPRRR-1PaPRR1RPaPaP (Fig. 5.19a) Pa\\P\\R||R\\\\R (Fig. 5.3a) The last joints of the four limbs have superposed axes/directions. The revolute joints between links 5 and 6 of limbs G1, G2 and G3 have orthogonal axes Pa\\P\\\\R||R (Fig" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_35_0001419_978-94-007-7401-8-Figure3.61-1.png", + "original_path": "designv11-35/openalex_figure/designv11_35_0001419_978-94-007-7401-8-Figure3.61-1.png", + "caption": "Fig. 3.61 Overactuated PMs with coupled Sch\u00f6nflies motions of types 4PaRPP (a) and 4RPaPP (b) defined by MF \u00bc SF \u00bc 4; RF\u00f0 \u00de \u00bc v1; v2; v3;xd\u00f0 \u00de, TF = 0, NF = 18, limb topology Pa||R||P\\P (a) and R||Pa||P\\P (b)", + "texts": [ + " 3.50e) Idem No. 1 6. 4PPRPa (Fig. 3.56b) P\\P\\kR||Pa (Fig. 3.50f) Idem No. 1 7. 4PRPPa (Fig. 3.57a) P\\R||P||Pa (Fig. 3.50g) Idem No. 1 8. 4PRPaP (Fig. 3.57b) P\\R||Pa||P (Fig. 3.50h) Idem No. 1 9. 4RPRPa (Fig. 3.58a) R||P||R||Pa (Fig. 3.50i) Idem No. 1 10. 4RRPPa (Fig. 3.58b) R||R||P||Pa (Fig. 3.50j) Idem No. 1 11. RRPaP (Fig. 3.59) R||R||Pa||P (Fig. 3.50k) Idem No. 1 12. 4RPaPP (Fig. 3.60a) R||Pa\\P\\kP (Fig. 3.50l) Idem No. 1 13. 4RPPPa (Fig. 3.60b) R||P\\P\\kPa (Fig. 3.50m) Idem No. 1 14. 4PaRPP (Fig. 3.61a) Pa||R||P\\P (Fig. 3.50n) Idem No. 1 15. 4RPaPP (Fig. 3.61b) R||Pa||P\\P (Fig. 3.50o) Idem No. 1 16. 4RPPaP (Fig. 3.62a) R\\P\\kPa||P (Fig. 3.50p) Idem No. 1 17. 4PPPaR (Fig. 3.62b) P\\P\\kPa||R (Fig. 3.50q) Idem No. 1 18. 4PPPaR (Fig. 3.63a) P\\P||Pa||R (Fig. 3.50r) Idem No. 1 19. 4PPaRR (Fig. 3.63b) P||Pa||R||R (Fig. 3.50s) Idem No. 1 20. 4PPaRP (Fig. 3.64) P\\Pa||R||P (Fig. 3.50t) Idem No. 1 21. 4PPaPR (Fig. 3.65) P\\Pa||P||R (Fig. 3.50u) Idem No. 1 22. 4PPaPR (Fig. 3.66a) P||Pa\\P\\kR (Fig. 3.50v) Idem No. 1 23. 4PPaRP (Fig. 3.66b) P||Pa\\R\\P (Fig. 3.50i0) The before last joints of the four limbs have parallel axes 24" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_35_0001552_j.nahs.2015.06.002-Figure1-1.png", + "original_path": "designv11-35/openalex_figure/designv11_35_0001552_j.nahs.2015.06.002-Figure1-1.png", + "caption": "Fig. 1. Examples of the time evolution of switched system (7) under the state-dependent Switching Law 1. (Left) n = 2, (Right) n = 3.", + "texts": [ + " Assume that the nominal system (7) is GPS, we would like to find a state-dependent switching law such that (6) is G\u03b5-PAS. Problem 2. Show that a stabilizing switching law for the Problem 1, with ci = fi(0), \u03b5-practically stabilizes (1). Firstly, consider the state-dependent switching law for system (7) as follows Switching Law 1 ([14]). \u2022 If x(t) \u2208 Int(Ck), 1 \u2264 k \u2264 n, the subsystem k + 1 is active. \u2022 If x(t) \u2208 Int(Cn+1), the subsystem 1 is active. \u2022 If x(t) \u2208 Ci Cj, i =\u0338 j, the active subsystem is the one of the cone where the trajectory enters (see Fig. 1). Notice that Switching Law 1 uses the boundary of the cone Ck as switching surfaces. To see this clearly, consider the nominal system (7) for n = 1. For (7) to be stabilizable it is required 2 i=1 Ci = R, which means that if c1 > 0 and c2 < 0, C1 = {x \u2208 R : x = \u03bb2(\u2212c2), \u03bb2 \u2265 0} and C2 = {x \u2208 R : x = \u03bb1(\u2212c1), \u03bb1 \u2265 0}. According to Switching Law 1 every time that x(t) \u2208 Int(C1), the system will evolve along Subsystem 2; and when x(t) \u2208 Int(C2), the system will evolve along Subsystem 1, making x = 0 a common cone boundary where the system switches" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_35_0001419_978-94-007-7401-8-Figure5.36-1.png", + "original_path": "designv11-35/openalex_figure/designv11_35_0001419_978-94-007-7401-8-Figure5.36-1.png", + "caption": "Fig. 5.36 3PaPaPR-1RUPU-type fully-parallel PMs with uncoupled Sch\u00f6nflies motions defined by MF = SF = 4, RF\u00f0 \u00de \u00bc \u00f0v1; v2; v3;xb\u00de, TF = 0, NF = 22 limb topology R\\R\\R\\P\\kR\\R and Pa\\Pa\\kP\\\\R, Pa\\Pa\\kP\\kR (a), Pa\\Pa\\\\P\\\\R, Pa\\Pa\\\\P\\kR (b)", + "texts": [ + "35a) Pa||P\\Pa\\kR (Fig. 5.4a) The first revolute joint of G4 limb and the last revolute joints of limbs G1, G2 and G3 have parallel axes. The last joints of of limbs G1, G2 and G3 have superposed axes Pa||P\\Pa||R (Fig. 5.4d) R\\R\\R\\P\\kR\\R (Fig. 5.1b) (continued) 530 5 Topologies with Uncoupled Sch\u00f6nflies Motions Table 5.3 (continued) No. PM type Limb topology Connecting conditions 8. 3PaPPaR1RUPU (Fig. 5.35b) Pa\\P\\\\Pa\\kR (Fig. 5.4b) Idem No. 7 Pa\\P\\\\Pa||R (Fig. 5.4c) R\\R\\R\\P\\kR\\R (Fig. 5.1b) 9. 3PaPaPR1RUPU (Fig. 5.36a) Pa\\Pa\\kP\\\\R (Fig. 5.4f) Idem No. 7 Pa\\Pa\\kP\\kR (Fig. 5.4h) R\\R\\R\\P\\kR\\R (Fig. 5.1b) 10. 3PaPaPR1RUPU (Fig. 5.36b) Pa\\Pa\\\\P\\\\R (Fig. 5.4g) Idem No. 7 Pa\\Pa\\\\P\\kR (Fig. 5.4e) R\\R\\R\\P\\kR\\R (Fig. 5.1b) 11. 2PaPaRRR1PaPaRR1RUPU (Fig. 5.37) Pa\\Pa||R||R\\\\R (Fig. 5.5a) The last joints of of limbs G1, G2 and G3 have superposed axes. The revolute joints between links 7 and 8 of limbs G1, G2 and G3 have orthogonal axes Pa\\Pa||R||R (Fig. 5.4i) R\\R\\R\\P\\kR\\R (Fig. 5.1b) 12. 2PaPaRRR1PaPaRR1RUPU (Fig. 5.38) Pa\\Pa||R||R\\kR (Fig. 5.5b) Idem No. 11 Pa\\Pa||R||R (Fig. 5.4j) R\\R\\R\\P\\kR\\R (Fig. 5.1b) Table 5.4 Limb topology and connecting conditions of the fully-parallel solutions with no idle mobilities presented in Figs" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_35_0000807_s12206-021-0724-8-Figure2-1.png", + "original_path": "designv11-35/openalex_figure/designv11_35_0000807_s12206-021-0724-8-Figure2-1.png", + "caption": "Fig. 2. FE models of structure: (a) entire engine block; (b) crankcase housing; (c) single-bearing housings.", + "texts": [ + " The leakage of the oil is obtained as follows: 2 23 3 0 0/2 /2 d\u03b8 d\u03b8 12 12= =\u2212 \u2202 \u2202 = + \u2202 \u2202\u222b \u222bo z B z B h p h pQ R R z z \u03c0 \u03c0 \u03b7 \u03b7 . (13) The structure finite element (FE) model is established in this section. The FE model scale influences the calculation accuracy, and three different scale FE models can be selected for the calculation of the main bearing lubrication characteristics of a V-type six-cylinder engine, namely, the entire engine block, crankcase housing, and single bearing housings, as shown in Fig. 2. The crankcase housing model has a smaller scale and higher stiffness accuracy than the engine block and single bearing housing models. Simulations are conducted using different FE models, and the results show that the PACP of the main bearing in the first two cases is consistent, as shown in Fig. 3. The maximum relative difference of the average PACP is less than 10 %. Therefore, the crankcase housing model is selected in the simulation. Fig. 4 shows the FE model of the crankcase housing and crankshaft" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_35_0000498_2050-7038.12961-Figure16-1.png", + "original_path": "designv11-35/openalex_figure/designv11_35_0000498_2050-7038.12961-Figure16-1.png", + "caption": "FIGURE 16 Magnetic field density distribution plots under the no-load condition for (A) 4p24s, (B) 4p36s, and (C) 4p42s machine models", + "texts": [ + "3 0.4 0.5 C o g g in g T o rq u e [ N m ] Time [s] 4p24s 4p36s 4p42s FIGURE 20 Cogging torque for the investigated machine models To investigate the performance of the employed machine models for the validation of the proposed simplified brushless field-excitation scheme under no-load condition, a DC of 1A is used for the rotor field excitation and the machine is operated at 1800 rpm. The magnetic field density distribution plots of the studied machine models under no-load condition are shown in Figure 16; however, Figure 17 illustrates the induced back-EMF of the machine models under such a condition. The magnitude of induced back-EMF for the 4p24s machine model is 47.00 Vrms. This magnitude for 4p36s and 4p42s machine models is 69.55 Vrms and 76.63 Vrms, respectively. The harmonic contents of the induced back-EMF for the studied machine models under no-load condition are shown in Figure 18. Figure 19 illustrates the total harmonic distortion (THD) of the induced back-EMF which shows that the THD for the 4p42s model is around 10" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_35_0000596_s43452-021-00255-x-Figure14-1.png", + "original_path": "designv11-35/openalex_figure/designv11_35_0000596_s43452-021-00255-x-Figure14-1.png", + "caption": "Fig. 14 a Focus tension around the rivet hole, b elastic stress distribution in notches (Sr 9)", + "texts": [ + " Here, 174,214 elements have been used because of the convergence of the solution to the problem as well as the appropriate number of meshes that speed up the analysis (see Fig.\u00a011). Numerical analysis for the elastic region shows that the greatest stress concentration occurs around the rivet holes. As the number of rivets increases, the joint stiffness increases, and as a result, the maximum stress in the structure decreases. The von Mises stress contour and stress distribution in the internal layer of the joint, 9-rivet layout, are shown in Figs.\u00a012 and 13. According to Fig.\u00a0 14a observed that the maximum von Mises stress on the rivet equals 74.25\u00a0MPa. Also, in Fig.\u00a014b, the highest amount of the elastic stress distribution occurs in a lateral surface of holes (53.66\u00a0MPa). Results of simulation for the double-lap adhesive joint using ABAQUS software was showed that the highest amount of stress in the adhesive layer equal 17.61\u00a0MPa and occurs around holes and external edges (Fig.\u00a015a, b). 1 3 1 3 In the simulation of an adhesive joint, what is seen is the uniformity of tension that occurs at the surface of the joint. Also, creating tension concentration around the rivet holes in the structure is a feature of these rivet joints" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_35_0000750_j.matpr.2021.07.468-Figure4-1.png", + "original_path": "designv11-35/openalex_figure/designv11_35_0000750_j.matpr.2021.07.468-Figure4-1.png", + "caption": "Fig 4. Schematic of Fused Deposition Modelling [30].", + "texts": [ + " This process is compatible with 2D printing methods avoiding tedious standard 3D printing methods. It can also deliver speedy prototyping of structures that are not feasible with a layer-by-layer printing process. Demonstrated soft grippers have the potential for soft robotic applications [28]. Boley et al. presented the application of EGaln nanoparticles in flexible electronics. They consist of inkjetprinted EGaln nanoparticles on an elastomeric glove to create arrays of strain gauges [29]. FDM, also known as fused filament fabrication (FFF), is the most widely used (Fig 4) 3D printing process due to its cheap fabrication cost and ease of use. Here thermoplastic filament is heated, melted and deposited onto a build platform through a nozzle, one layer at a time. It gradually cools down and solidifies while being fused with the top and bottom layers of filament. A wide range of structures with variations in size and ink Design can be designed using the FDM method. Some of the most commonly used FDM filament materials are polycarbonate and polylactic acid. Another advantage of FDM, aside from being cost-effective and easy to use, is the extremely low post-processing requirements" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_35_0000222_s12369-021-00769-7-Figure2-1.png", + "original_path": "designv11-35/openalex_figure/designv11_35_0000222_s12369-021-00769-7-Figure2-1.png", + "caption": "Fig. 2 Machine model using SimMechanics", + "texts": [ + " With the generation of XML file, SimMechancis automatically generates an STL body geometry file, for each body represented as a CAD assembly part (Fig.\u00a01a). These STL files were generated for SimMechancis visualization (Fig.\u00a01b). For complete CAD translation, XML file was converted to SimMechancis model. With the help of STL body geometry files, bodies with their original CAD assembly shapes could be visualized. To validate the mechanism design of active orthosis in terms of kinematic and kinetic constraints, a SimMechancis model of the active orthosis is generated from the CAD assembly (Fig.\u00a02). The SolidWorks to SimMechancis translator is used for developing this model. SimMechancis has got additional advantages of coordinate alignment, mass and inertia properties which it directly imports from CAD assembly (Fig.\u00a03). The robot prototype is powered by a relatively new class of actuators know as Pneumatic Muscle Actuators (PMA). PMA are lightweight and have high power to weight and power to volume ratios as compared to conventional electromagnetic actuators. PMA behave similar to skeletal muscles and can only provide unidirectional pulling forces" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_35_0001419_978-94-007-7401-8-Figure3.23-1.png", + "original_path": "designv11-35/openalex_figure/designv11_35_0001419_978-94-007-7401-8-Figure3.23-1.png", + "caption": "Fig. 3.23 Overactuated PMs with coupled Sch\u00f6nflies motions of types 4PCR (a) and 4RCP (b) defined by MF \u00bc SF \u00bc 4; RF\u00f0 \u00de \u00bc v1; v2; v3;xd\u00f0 \u00de, TF = 0, NF = 6, limb topology P\\C||R (a) and R||C\\P (b)", + "texts": [ + "26, 3.27, 3.28, 3.29, 3.30, 3.31, 3.32, 3.33, 3.34, 3.35, 3.36, 3.37, 3.38, 3.39, 3.40, 3.41, 3.42, 3.43, 3.44, 3.45, 3.46, 3.47, 3.48, 3.49 No. PM type Limb topology Connecting conditions 1. 4RRC (Fig. 3.21a) R||R||C (Fig. 3.2a) The cylindrical joints of the four limbs have parallel axes 2. 4RCR (Fig. 3.21b) R||C||R (Fig. 3.2b) Idem No. 1 3. 4PCP (Fig. 3.22a) P\\C\\P (Fig. 3.2c) Idem No. 1 4. 4CPP (Fig. 3.22b) C\\P\\\\P (Fig. 3.2d) Idem No. 1 5. 4PPC (Fig. 3.22c) P\\P\\\\C (Fig. 3.2e) Idem No. 1 6. 4PCR (Fig. 3.23a) P\\C||R (Fig. 3.2f) Idem No. 1 7. 4RCP (Fig. 3.23b) R||C\\P (Fig. 3.2g) Idem No. 1 8. 4RPC (Fig. 3.24a) R\\P\\kC (Fig. 3.2h) Idem No. 1 9. 4PRC (Fig. 3.24b) P\\R||C (Fig. 3.2i) Idem No. 1 10. 4CRP (Fig. 3.25a) C||R\\P (Fig. 3.2j) Idem No. 1 11. 4CPR (Fig. 3.25b) C\\P\\kR (Fig. 3.2k) Idem No. 1 12. 4RPC (Fig. 3.26a) R\\P\\kC (Fig. 3.2i) Idem No. 1 13. 4CRR (Fig. 3.26b) C||R||R (Fig. 3.2m) Idem No. 1 14. 4RRRRR (Fig. 3.27a) R\\R||R\\R||R (Fig. 3.3a) The first and the last revolute joints of the four limbs have parallel axes 15. 4RRRRR (Fig. 3.27b) R||R\\R||R\\R (Fig" + ], + "surrounding_texts": [] + } +] \ No newline at end of file