[ { "image_filename": "designv11_62_0003210_joe.2018.8514-Figure1-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0003210_joe.2018.8514-Figure1-1.png", "caption": "Fig. 1 ZM cat-head type tower", "texts": [ " Therefore, it is necessary to study the insulator discharge characteristics when the water falls from the box of bird guard. The effects of box of bird guard used in 220 kV transmission line on the electric field and potential distribution on transmission line insulators are studied here. Based on the results of the simulation, a reasonable design scheme for the box of bird guard scupper is proposed. 2.1.1 Transmission line tower: The selected line tower is a ZM cat-head type tower, which is used in 220 kV transmission line and its cross arm is 30 m above the ground. The structure of the tower is shown in Fig. 1. The type of the composite insulator is FXBW4/220-100, the relative permittivity of the insulator umbrella shed is 3.5. The diameter of the grading ring is 400 mm, pipe diameter is 50 mm, and the shielding depth is 60 mm. 2.1.2 Box of bird guard: Based on the actual structure of the tower, the box of bird guard is a conical structure, and the diagram of the box is shown in Fig. 2. Box of bird guard with metal material and fiberglass material were simulated, respectively, to study the influence of material on the electric field and potential distribution of the insulator, the relative permittivity of the fiberglass box of bird guard is 4" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0001522_icra.2011.5980118-Figure2-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0001522_icra.2011.5980118-Figure2-1.png", "caption": "Fig. 2. Data flow in the filter system. For each joint i and time t, the states Sj xi,t and variances Sj Pi,t are transformed from each space Sj to Sk , second the movement is added (see Sec. III), third the estimated states and variances are combined with the current respective measurements.", "texts": [ " In order to model a (robotic) arm with inaccurate movements that is monitored by noisy measurements, we implement Kalman filters. For each frame Sj Fi, we use a different filter Sj Ki. We have implemented transformations between these spaces whose dependencies are shown in Fig.1. In each iteration, first, the information from each separate space is combined. Next, current movement information is integrated. Finally, sensory measurements are considered. This iterative information exchange is illustrated in Fig. 2. At each discrete point in time t, each transformation in Fig. 1 is modeled by two hard-coded mappings, one for the states Sjxi,t 7\u2192 SjSkxi,t (5) and one for the covariance-matrices P SjPi,t 7\u2192 SjSkPi,t = ( \u2202Skxi,t \u2202Sjxi,t \u2217 \u2202Skxi,t \u2202Sjxi,t ) \u00b7 SjPi,t, (6) where \u2202Skxi,t \u2202 Sjxi,t is the Jacobian-matrix for transforming from space Sj to Sk, \u2217 denotes the element-wise matrix product, and \u00b7 a matrix-multiplication. A second term can occur in (6) if uncertain origins need to be modeled (see Sec. II-B", " In order to move from the current arm position PSyi,t in a straight line to the goal, we determine a current desired arm position PSg\u2032 i,t: PSg\u2032 i,t =PS yi,t + min { PSvmax, |PSgi \u2212 PSyi,t| } \u00b7 PSgi \u2212 PSyi,t |PSgi \u2212 PSyi,t| , (30) where the desired movement distance is capped by the maximum velocity vmax. As the arm is moved by manipulating its joint angles, this desired arm position has to be transformed from PS to LP resulting in a desired position in joint angle space: PSLPg\u2032 i,t. The difference of the current arm position LPyi,t from this desired position is used as movement plan in angular space: LPdxi,t = PSLP g\u2032 i,t \u2212 LPyi,t. (31) This is done in each iteration after combining the state information of all spaces (cf. Fig. 2). After the planned arm movement is executed, the states Skxi,t of the Kalman filters have to be updated. Because of the noisy quality of the movement, the uncertainty in all state estimates increases at this point. Also, as the movement itself is not measured in any way and the planned movement dxi,t is only known in posture space, it has to be transformed to the other spaces Sk 6= LP . Jacobians (6) could be used for this transformation, however, to avoid singular Jacobians and their non-existing inverse [10] as well as errors due to linearization, we propound a different approach" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0001608_isam.2011.5942312-Figure2-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0001608_isam.2011.5942312-Figure2-1.png", "caption": "Figure 2. Schematic drawing of experimental apparatus", "texts": [ " Two regimes characterized by distinct hydrodynamic pressure are expected in this new design. Under the condition of IHP, the rotor is supported by the RB, protecting the HB from progressive contact wear, as shown in Fig. 1a. While under the condition of SHP, such as at working speeds, the rotor is levitated by the HB and rotates inside the RB radial clearance. Therefore only the HB supports the rotor (Fig. 1b). Both the clearances of the HB and the RB are drawn at an exaggerated scale in Fig. 1 for better visualization. III. TEST RIG The schematic view of the test rig is shown in Fig. 2. Two pedestals spaced approximately 150mm apart support the stator of the rotor-bearing system. This paper is supported by National Basic Research Program of China (973 Program, Grant No. 2009CB724404) The JRHB is composed of the HB and the cylindrical roller bearing (CRB) located on the right. The HB has a symmetric four lobe fixed geometry configuration with a nominal bore diameter of 48mm and a diameter clearance of 30\u03bcm. The bearing designation of the CRB is NU207E with a common class of bearing tolerance and basic radial clearance section" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0000197_978-3-642-28768-8_65-Figure2-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0000197_978-3-642-28768-8_65-Figure2-1.png", "caption": "Fig. 2 Tooth meshing deflections", "texts": [ " In this way the coupling can be done even with the generator working as a motor when the generator speed is lower than the synchronous one. Meshing forces have been included as user functions (GFORCE) defined following the approach in previous work by the same authors for a low-speed ordinary transmission [5]. Tooth contacting forces are assumed to be contained in the transmission plane, perpendicular to the tooth surface. Thus, tooth deflection can be obtained from the relative displacements of each gear center xi,yi,zi,\u03b8xi,\u03b8yi,\u03b8zi, using the following expression (see Figure 2): ( cos sin cos sin ) cos ( cos sin cos sin ) sin ( ) i ij i ij i zi j ij j ij j zi ij i i xi ij i yi ij j j xj ij j yj ij ij ij ij x y x y z z e t \u03b4 \u03d5 \u03d5 \u03c1 \u03b8 \u03d5 \u03d5 \u03c1 \u03b8 \u03b2 \u03c1 \u03b8 \u03d5 \u03c1 \u03b8 \u03d5 \u03c1 \u03b8 \u03d5 \u03c1 \u03b8 \u03d5 \u03b2 = \u22c5 + \u22c5 + \u22c5 \u2212 \u22c5 \u2212 \u22c5 + \u22c5 \u22c5 + \u2212 \u22c5 \u22c5 \u2212 \u22c5 \u22c5 \u2212 \u2212 \u22c5 \u22c5 \u2212 \u22c5 \u22c5 \u22c5 + (2) Where eij(t) represents the periodic static transmission error due to profile errors, \u03b2ij is the helix angle on the pitch cylinder and \u03d5ij the normal pressure angle. Once tooth deflections are known, meshing force is obtained by multiplying the resulting deflection by the gear pair stiffness", " The placement of the chains around the floating platform defines the yaw stiffness which is one of the most critical features in the design of this kind of supports. A preliminary analysis of the hydrostatic buoyancy behaviour provides rotational stiffness as well as the translational one in the vertical direction. These features are defined in the MBS model by a GFORCE element where the damping associated with each translational and rotational degree of freedom is included. Mooring lines are also analysed in a previous stage and then included in the model as three non linear springs arranged in the xz plane (see Fig 2) every 120 degrees. The MBS model described previously has been applied to a known 750 kW wind turbine drive train, the NREL 750 GRC [6]. No details are presented in this work about the parameters and data related to the example except those not appearing in the cited reference. In Table 1 the data corresponding to the meshing stiffness of each gear pair are presented. These data have been obtained following the approach proposed in [5]. Bearings for the main shaft are modelled as BUSHING joints and only present supported by the main shaft bearing closest to the carrier and the planets themselves" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0002136_2017-01-1506-Figure21-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0002136_2017-01-1506-Figure21-1.png", "caption": "Figure 21. Overview on the applications.", "texts": [], "surrounding_texts": [ "First application is the robustness test of the ESC system. However, such tests could also be used for other control systems during the development phase. Next application are transitions where the surface changes suddenly from a dry surface to slippery spots, e.g. polished ice. Additional difficulties arise when one vehicle side remains on dry surface, the other side enters a slippery surface. Classical example for that is the \u03bc-split-braking (Figure 20). In simulation tools, this is often done by using the tyre characteristic on a dry surface and then scaling the force down with a certain factor e.g. 02. Some tyre models as MF-Tyre already have this idea implemented as scaling factors for the Pacejka parameters can be directly changed from outside. This was topic of the European funded research project VERTEC (\u201cVehicle, road, tyre and electronic control systems interaction: increasing active vehicle safety by means of a fully integrated model for behaviour prediction in potentially dangerous situations\u201d [11]) Other application is a test where the scaling factors change slowly from 120% to 80% during the same test manoeuvres This shows how good the controller can adapt itself to the changing surface conditions. This figure summarise the application proposals by showing the surface change over a special surface e.g. scraped ice. Summarising this chapter: The authors presented some applications for the generic tyre parameters." ] }, { "image_filename": "designv11_62_0003356_978-981-13-5799-2_12-Figure12.109-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0003356_978-981-13-5799-2_12-Figure12.109-1.png", "caption": "Fig. 12.109 Tire model (reproduced from Ref. [153] with the permission of Tire Sci. Technol.)", "texts": [ "107 shows that the equivalent shear strain in the soil is concentrates in a narrow band between the top and bottom boxes. Figure 12.108 shows good agreement between the numerical simulation and experimental results. (2) Prediction of the traction performance in different soil conditions The gross traction (GT) and the motion resistance (Rc) can be evaluated from the stress obtained on the surface contacting with soil in the numerical simulation. The tire model used in this simulation is shown in Fig. 12.109. The tire size is 540/ 65R30, the inflation pressure is 240 kPa, and the load is 32.86 kN. The tire is modeled using Lagrangian elements, while a soil bin having a depth of 0.326 m is modeled using Eulerian elements. The coefficient of friction between the tire and the soil surface is ignored. Figure 12.110 shows the calculated rut shape of a tire with a slip ratio of nearly 100% and the calculated cone penetration resistance of soft soil. The tire sinkage is nearly 110 mm, and the result is in good qualitative agreement with common phenomena" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0002354_1350650117710813-Figure10-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0002354_1350650117710813-Figure10-1.png", "caption": "Figure 10. Schematic diagram of shaft in Cartesian coordinates.", "texts": [ " Calculation of dynamic coefficients in a hydrodynamic bearing considering five degrees of freedom for a general rotor-bearing system. J Tribol Trans ASME 1999; 121: 499\u2013505. 25. Das S, Guha SK and Chattopadhyay AK. On the steady-state performance of misaligned hydrodynamic journal bearings lubricated with micropolar fluids. Tribol Int 2002; 35: 201\u2013210. 26. Lund JW. Review of the concept of dynamic coeffi- cients for fluid film journal bearings. J Tribol Trans ASME 1987; 109: 37\u201341. Appendix 1 As shown in Figure 10, coordinates of the shaft center OC are XC J , Z C J , and in dimensionless form are XC, ZC . Dimensionless film thickness at the shaft section across OC can be represented as h \u00bc 1 XC cos x ZC sin x \u00f028\u00de When the shaft rotates around X and Z with angles X, Z, since the rotating angles are very small (sin X X, sin Z Z), dimensionless coordinates of points on axis of the shat can be represented using their Y coordinates y as XC \u00fe Z y, ZC X y\u00de. Take OA for example, its coordinates can be represented as XC \u00fe Z Lg \u00fe 1 2LB , ZC X Lg \u00fe 1 2LB \u00de" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0002326_icmech.2017.7921149-Figure3-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0002326_icmech.2017.7921149-Figure3-1.png", "caption": "Fig. 3. Concept of vehicle demonstrator with ICC for three mechatronic subsystems", "texts": [ " Individuals from non-academic sectors seconded to academic partners are being exposed to recent knowledge in scientific methods on vehicle system modelling and control. Fig. 2 specifies several channels of knowledge sharing through secondments (staff exchanges) between the project partners. The proposed combination of research, development and training processes results in new ICC and related vehicle systems and testing procedures, which are discussed in subsequent sections. V. INTEGRATED CHASSIS CONTROL OF VEHICLE DEMONSTRATOR The training activities under discussion led to the concept of the target vehicle demonstrator shown on Fig. 3. This concept was firstly introduced in [8] and includes three mechatronic systems. The functions of the brake torque controller cover the base brake and wheel slip control and are realized for a decoupled electro-hydraulic brake-by-wire system. Such system has the brake pedal travel sensor to measure the driver\u2019s brake demand pbr_dem, which is then processed in the brake torque controller to estimate individual brake pressure pbr_est for each calliper in accordance with actual operational conditions" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0003254_s00170-019-03312-1-Figure6-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0003254_s00170-019-03312-1-Figure6-1.png", "caption": "Fig. 6 Curvilinear surfaces for a boundary contour model (a) 2 contour beads, with 25% overlap (0.025 in. width, 0.050 in. width, and (b) one contour bead (0.050 in. height, 0.10 in. width). Note the start-stop points", "texts": [ " Ra \u00bc Rsphere 2 \u03c0 2 \u2212cos\u22121 1\u2212 \u03c02 16 1=2 \u2212 \u03c0 4 1\u2212 \u03c02 16 1=2 ! \u00f04\u00de When emulating an AM process by layering the spherical beads, theoretical degenerate cases can exist for this model when there are non-tangent entities between slice heights if the bead where to be modeled as a sphere. This is an important observation, as this condition is true when employing an \u201cobround\u201d or elliptical bead, and introduces complexity into the model. Additional complexity is introduced when considering the situation in 3D: curvilinear surfaces (Fig. 6), multiple boundary layers, and the raster fill characteristics (Fig. 1) will influence the surface roughness values. The raster fill angle, and a changing raster fill angle between layers introduces additional surface textures that are required to be included for shallow inclination angles. Ahn et al. [18] have done initial investigations, but this needs to be explored deeply as the periodic nature of the layering process suggests that virtual tools could be developed to predictive the surface roughness" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0001699_cbo9780511780509.003-Figure2.33-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0001699_cbo9780511780509.003-Figure2.33-1.png", "caption": "Figure 2.33. The Poincare\u0301 map of the Duffing oscillator, with m = 1 kg, c = 10 Ns/m, k = \u22125 kN/m, h = 50 MN/m3, f0 = 30 N, and \u03c9 = 120 rad/s.", "texts": [ " This occurs at an excitation frequency of \u03c9 = 100 rad/s, for example, and the response is shown in Figure 2.31. Also shown is the DFT of the time response, which clearly highlights the significant response at one third of the excitation frequency. At other excitation frequencies \u2013 for example, at \u03c9 = 120 rad/s \u2013 the response is not periodic and may be quasiperiodic or chaotic. Figure 2.32 shows the response for an excitation frequency of \u03c9 = 120 rad/s, and the broadband response shows that the response is chaotic. One method used to analyze such a response is the Poincare\u0301 map, shown in Figure 2.33 for the same parameter values as used in Figure 2.32. To generate a Poincare\u0301 map, the system is simulated in the time domain until the transient response has decayed. The at https:/www.cambridge.org/core/terms. https://doi.org/10.1017/CBO9780511780509.003 Downloaded from https:/www.cambridge.org/core. The University of Melbourne Libraries, on 19 Jan 2017 at 20:35:23, subject to the Cambridge Core terms of use, available simulation is then continued and the resulting displacement, x, and velocity, x\u0307, are sampled at the same frequency as the excitation. These data are then plotted on the phase plane; that is, the velocity is plotted against the displacement. If the response is periodic, then the Poincare\u0301 map consists of a finite number of discrete points. Figure 2.33 shows a chaotic response, which has no periodicity. Another feature of stable damped linear systems is that the steady-state response of an unforced system is zero, which is not necessarily true for a nonlinear system. Consider a system described by the Van der Pol equation mx\u0308 + c(1 \u2212 \u03b3 x2)x\u0307 + kx = 0 (2.143) where the damping coefficient, c, is negative. If the system is displaced by a small amount from the equilibrium position with zero displacement and velocity, the vibrations grow because the damping is negative" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0000155_sps.2013.6623589-Figure3-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0000155_sps.2013.6623589-Figure3-1.png", "caption": "Fig. 3. Distribution of rotors with selected direction of rotation.", "texts": [], "surrounding_texts": [ "A quadcopter flight control is performed by changing the rotors speed. Each rotor produces a lift force and torque. If all rotors rotate with the same angular velocity - the clockwise rotors balance torque of anticlockwise rotors causing no rotation in the yaw axis. Acceleration of the axis pitch and roll can be achieved without changing the value of the yaw axis. For example, if we want quadcopter to fly forward we need to increase the angular velocity of the motor 3 against motor 1 at unchanged speed of motors 2 and 4." ] }, { "image_filename": "designv11_62_0003666_iemdc.2019.8785290-Figure4-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0003666_iemdc.2019.8785290-Figure4-1.png", "caption": "Fig. 4. Chained Schwarz-Christoffel mapping between infinite strip (canonical domain) and the desired geometry (physical domain). The mapping function h is calculated analytically while the mapping g is performed using numerical quadrature.", "texts": [ " Assembling the mapping function is not difficult, however, solving the integral and finding the prevertices wi for anything but the most basic geometries is. Reppe [14] introduced the idea of integrating the mapping function numerically, allowing for much more complex geometries. The prevertices wi can be obtained by making use of the fact that the integral between two prevertices must equal the corresponding polygon side length\u2223\u2223\u2223\u2223\u2223 \u222b wi+1 wi f \u2032 dw \u2223\u2223\u2223\u2223\u2223 = \u2223\u2223zi+1 \u2212 zi \u2223\u2223 . (2) Driscoll and Trefethen [15], [16] expanded on this idea. Chaining two mappings allows to map an infinite strip to an arbitrarily shaped physical domain as seen in fig. 4. The mapping function h(w) = \u03be = e\u03c0w (3) was obtained analytically for a strip of height 1. This strip map is at the core of the calculation method proposed in this work. When mapping a field solution from the canonical to the physical domain, the field is scaled and rotated by Bz = Bw \u00b7 f \u2032 |f \u2032|2 = Bw \u00b7 dz dw\u2223\u2223\u2223 dzdw \u2223\u2223\u22232 . (4) There are many implementation details, such as how to obtain the multiplicative constant K, how to perform fast and accurate numerical quadrature, how to solve and precondition the parameter problem and how to map inverse numerically. Driscoll\u2019s Schwarz-Christoffel Toolbox for Matlab [17], [18] and the accompanying documentation present a great resource to get into numerical Schwarz-Christoffel mappings. One useful concept when trying to model end region fields in an electrical machine is the fringing air-gap flux density. Fig. 4 shows this effect as calculated using a SchwarzChristoffel mapping. The air-gap flux density remains constant along the homogeneous part of the stator and rotor cores and fringes out into the end region. For the calculation shown in fig. 4, the clamping plate is assumed to be on the same magnetic scalar potential as the stator core. The rotor is on a different magnetic scalar potential, chosen to ensure the air-gap flux density in the homogeneous part to match the amplitude calculated by some other means. The actual field calculation is done by calculating the flux density inside the canonical domain (infinite strip) Bw = B\u03b4 \u00b7 \u03b4, (5) which is constant inside the entire domain and mapping it into the physical domain using equation (4)", " When implementing the method proposed in this paper, it needs to be made sure that the air-gap flux density can be calculated accurately, as it is proportional to the clamping plate field. While lacking a way to account for the effect of a magnetic clamping plate and other geometric features such as a pole shoe overhang or a stator core stepping, Traxler-Samek\u2019s approach [1] manages to account for 3D effects through the use of line conductor segments along the air-gap center carrying different currents matching the air-gap flux density in phase. One drawback of using a Schwarz-Christoffel mapping as presented in fig. 4 is the lack of any kind of 3D effects. Section V is going to show the impact of not accounting for 3D flux paths. Before getting into the method of accounting for 3D flux paths in a Schwarz-Christoffel-based field calculation, the FE models used to validate the calculations need to be introduced. Fig. 5 shows one of the FEM models used. It features a volumetric field winding carrying a fixed current and a virtually infinitely permeable (\u00b5r = 10000) stator core, rotor core and clamping plate. Clamping bolts and the associated holes through the clamping plate as well as non-magnetic pressure fingers (\u00b5r \u2248 1) were not modeled, as exploratory transient FE calculations showed them not to have a meaningful impact on the clamping plate field", " 3D FLUX PATHS Fig. 6 shows 3D flux paths in a FEM simulation. Despite setting the flux tube starting points almost dead center on a pole face, only flux tubes originating on the pole shoe actually enter the stator core or clamping plate. Flux tubes originating further down (in radial direction) move directly from one pole to the neighboring pole. This is in strong contrast to the behavior observed in the fringing air-gap flux density as calculated using a Schwarz-Christoffel mapping presented in fig. 4. Here, all flux lines originating on the rotor surface meet the stator core or clamping plate. From this observation alone it becomes obvious that the SC-based clamping plate field calculation must be producing results that are too large. To quantify the dependency, the clamping plate field was integrated over its surface along one pole pitch and its error was calculated as \u03b53D = \u03a6SC \u03a6FE \u2212 1. (7) Errors above zero indicate the Schwarz-Christoffel-based calculation to produce too large results. Fig" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0000744_isie.2013.6563755-Figure2-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0000744_isie.2013.6563755-Figure2-1.png", "caption": "Fig. 2. Cross section of the thrust wire", "texts": [ " However, it seems that the performance of the flexible actuator depends on shape changes of thrust wire. There are two reasons that affect the performance characteristics. One is the friction inside the thrust wire. Even though the finest materials are used in the manufacturing process to minimize the friction effect, it cannot be avoided completely. Even the Coulomb friction may exist and contributes to the friction between the inner wire and the outer tube. The other fact is the backlash. This backlash occurs due to the spacing between the inner wire and the outer tube as in Fig. 2. Since a sufficient space is necessary for the inner wire to move freely, this space has a minimum value to which the manufacturer can bring down. As a result, there exists a space for the inner wire to bend inside the outer tube when applied with a sufficient force. This ultimately resulted in a backlash between the two endpoints of the thrust wire. Therefore the motion transmission is affected by the spacing inside the wire. Recently there have been several attempts to overcome the friction inside the wire and compensate the backlash effect [11], [12]", " To achieve both force control target and position control target bilateral control is essential. Acceleration is taken as for the control reference so as to make it easier to control the two information, position and force at the same time. Control targets of bilateral control system is achieved by position tracking between master robot xm and slave robot xs with the law of action and reaction between the forces of master and slave. III. SYSTEM MODELING The cross section of the thrust wire used is shown in Fig. 2 and the measured dimensions of the wire are listed in table I. There is an air gap exists between the inner wire and the outer tube of a thrust wire. There are several reasons for the backlash to occur inside a thrust wire. One reason is the clearance between the inner wire and the inner wall of the outer tube. This allows the inner wire to bend inside the outer tube when the wire is pushed. Since the inner wire consists of twisting several strands, the length of the inner wire is slightly shorter than that of it when the wire is pulled by applying a tension" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0002621_aim.2017.8014122-Figure4-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0002621_aim.2017.8014122-Figure4-1.png", "caption": "Figure 4. Foot position put in front of stairs too far", "texts": [ " As the user walks close to stairs and prepare to climb on by using a predefined trajectory, there will be two possible dangerous situations happened. As shown in Fig. 3, if the forward leg is placed in front of stairs too closely, on the one hand, it will collide at next step because the space is too small to move leg along with a fixed stair walking trajectory. On the other hand, if it is placed too far, after the next step, the foot cannot be put on stair safely so that the single leg is difficult to prevent it from falling, as shown in Fig. 4, the reaction force to the CoP create an enormous torque causing the falling. When healthy people are climbing on stairs, they can move fluently with their inertia force because they have more powerful and flexible muscle and relative faster walking velocity to support climbing up, but the elderly and the spinal cord injuries don\u2019t have enough muscle strength and inertia force to preserve their climbing motion continuously. Thence, they need to avoid that their fictitious Zero Moment Point (FZMP), the red point, is located outside the next step area to create an unbalanced moment for falling like in Fig. 4, and the green point of CoP is the center of pressure of foot [10] [11]. By considering both walking trajectory, the trajectory of level walking is from even ground to even ground, and the trajectory of stairs walking is from stairs to stairs, therefore, a trajectory between even ground and stairs is required as a bridge to connect them. Therefore, a safety distance between former foot and stair before the user starting to climb up is also required as a switch to create an ideal trajectory from even ground to stairs" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0000394_09596518jsce982-Figure4-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0000394_09596518jsce982-Figure4-1.png", "caption": "Fig. 4 Case 1: no intersection between dig and excavator", "texts": [ " If RE (m) is the length of the boom\u2013arm pair reduced to carriage rotation point and extended to its outmost position then, under the assumption that the excavator is always positioned at a point lying outside dig borders, several distinct cases should be examined in order to make conclusions about the locus (or loci) of points that correspond to the maximum volume of soil removal. In any case, the factors that should be taken into account in order to determine the point loci are: (a) dig dimensions w and l, (b) the length of the excavator\u2019s boom\u2013arm pair RE, (c) the coordinates (xE, yE) of the carriage rotation point, and (d) angles hE, hR, hL that define the EUZ of the excavator. In case 1, excavator is located at (xE, yE), and since xE{w\u00f0 \u00de\u00a2RE with xE, REw0, there is no intersection between the dig\u2019s surface and the excavator\u2019s work space (see Fig. 4). Similar cases are: (a) xEzRE\u00f0 \u00de\u00a10 with xEv0, REw0, (b) yE{l\u00f0 \u00de\u00a2RE with yE, REw0 and (c) yEzRE\u00f0 \u00de\u00a10 with yEv0, REw0. Case 2 corresponds to the situation when the circle of radius RE formed by the excavator\u2019s end intersects only one side of the dig. This can be expressed by the inequality l\u00a22Re sin hs~2Re sin cos{1 xE{w RE \u00f02\u00de And xe{w\u00f0 \u00de\u00a1Re \u00f03\u00de The soil volume that can be removed is analogous to the surface of the circular segment formed by circular sector and triangle EFG. The sector surface is given as R2 Ehs 2 and the triangle EFG has a surface equal to xE{w\u00f0 \u00deRE sin hs , therefore Vs~ 2hs hmax Vmax~ 2hs hL{hR\u00f0 \u00deVmax \u00f04\u00de with hs~cos{1 xE{w RE " ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0001298_eeeic.2011.5874589-Figure1-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0001298_eeeic.2011.5874589-Figure1-1.png", "caption": "Figure 1. Circuit magnetic of induction motor", "texts": [ " Thus, we use the magneto-evolutionary formulation. The magnetic circuit is represented in two dimensions. We have chosen a section which is perpendicular to the axis of rotation of the machine. The magneto-evolutionary model is represented by the following equation: ( ) ( )HtroJ Atro tro dt Ad e +=\u239f\u239f \u23a0 \u239e \u239c\u239c \u239d \u239b + \u03bc \u03c3 1 (1) A : magnetic vector potential (Wb/m) J : current density uniform (A/m) \u03bc : magnetic permeability (H/m) H : magnetic field (A/m) e\u03c3 : electrical conductivity ( m.1 \u03a9 ) T : Time (seconds) The magnetic circuit is shown below (Figure.1). This motor has been mounted in a test-bench at the laboratory AMPERE Lyon1. The use of the sliding area (rotating zone), defined in Flux2D, allows us to consider the rotation of rotor 978-1-4244-8782-0/11/$26.00 \u00a92011 IEEE in magneto-evolutionary study, without making a new mesh of the machine at each position of the rotor. The distribution of the winding machine is shown in Figure 1 according to a polar periodicity. The rotating air gap is represented (fig.1). We have modelled the whole geometry, in order to be able to take into account the fault, since we lose the symmetry of the machine. III. ELECTRICAL CIRCUIT OF STATOR Electrical circuit of stator (fig.2) contains coils, which will be linked to magnetic scheme. They represent the conductors who will be included in the stator slots. Between two coils connected, there is a resistance representing the resistance of the head coil (2) and an inductance which represents the leakage inductance slot and head coils (3)" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0001433_pi-a.1962.0130-Figure2-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0001433_pi-a.1962.0130-Figure2-1.png", "caption": "Fig. 2.\u2014General arrangement of small high-speed permanent-magnet generator for missile.", "texts": [ " Skilled handling is essential for dismantling and reassembling permanent-magnet machines to avoid loss or weakening of magnetism; if, in the process, some magnetism is lost, special magnetizing equipment will usually be required. Small permanent-magnet alternators are highly suitable for use in missiles, and in aircraft main generators, as a source of auxiliary power for excitation and operation of control and protective gear. In the latter application they can be mounted within the main-generator housing without adding appreciably to its weight or adversely affecting its general arrangement. Fig. 2 shows the general arrangement of a small high-speed alternator designed for short-range missile requirements, incorporating a two-pole permanent-magnet rotor. This generator delivers 600 VA at 0-9p.f. 200 V 3-phase 400c/s at 24000r.p.m. It weighs 3$lb. Its load is sensibly constant and no voltagecontrol equipment is required. (6) INDUCTION GENERATORS (6.1) General In modern aircraft, constant-frequency a.c. power is supplied by synchronous generators driven at constant speed by an air turbine or by one of a variety of hydraulic constant-speed devices" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0000521_j.camwa.2010.05.002-Figure12-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0000521_j.camwa.2010.05.002-Figure12-1.png", "caption": "Fig. 12. Two-link robot.", "texts": [ " Consider the large-scale unknown system that contains two interconnected two-in-two-out subsystems described as follows: \u03a31 : x\u03071(t) = f1(x1(t))+ g1(x1(t)) [u1(t)+ h12(x2(t \u2212 \u03c412))] ; y1(t) = C1x1(t), (68) \u03a32 : x\u03072(t) = f2(x2(t))+ g2(x2(t)) [u2(t)+ h21(x1(t \u2212 \u03c421))] ; y2(t) = C2x2(t), (69) where u1(t) = [ u1,1(t) u1,2(t) ] , u2(t) = [ u2,1(t) u2,2(t) ] , x1(t) = [ x1,1(t) x1,2(t) x1,3(t) x1,4(t) ]T , x2(t) = [x2,1(t) x2,2(t) x2,3(t) x2,4(t)]T . The first Subsystem\u03a31 of the large-scale system is given by two-link robot (Fig. 12), which is described as follows: The dynamic equation of the two-link robot system can be expressed as follows: M(q)q\u0308+ C(q, q\u0307)q\u0307+ G(q) = \u0393 , (70) where M(q) = [ (m1 +m2)l21 m2l1l2(s1s2 + c1c2) m2l1l2(s1s2 + c1c2) m2l22 ] , C(q, q\u0307) = m2l1l2(c1s2 \u2212 s1c2) [ 0 \u2212q\u03072 \u2212q\u03071 0 ] , G(q) = [ \u2212(m1 +m2)l1gr s1 \u2212m2l2gr s2 ] , and q = [ q1 q2 ]T , q1, q2 are the angular positions, M(q) is the moment of inertia, C(q, q\u0307) includes coriolis and centripetal forces, G(q) is the gravitational force, \u0393 is the applied torque vector" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0000345_icce.2010.5418795-Figure1-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0000345_icce.2010.5418795-Figure1-1.png", "caption": "Fig. 1. A Golf Swing Model.", "texts": [ " Finally, we can estimate the golf club\u2019s loci and attitudes using the estimated golf club\u2019s movement when hitting a ball. B. Golf Swing Model Some constraints are necessary because real golf swing is complicated to estimating of golf club\u2019s loci and attitudes. The constraints of the golf swing model on this paper are below: 1) The movement of golf club only exists on the plane that makes \u03b80 degree to the ground, 2) The swing plane is fixed while a swing, 3) The sensor and the golf club head are always on the swing plane, 4) A player\u2019s shoulder is fixed while a swing. The golf swing model is shown as Fig.1. A player grip a 3- axis acceleration sensor. And at the front of the sensor, there is a virtual golf club. Loci and attitudes of the golf club\u2019s head are applied to a golf game. When starting a golf swing, a player points the club to the ball. And after that, a player makes a motion that consists of a backswing and a downswing. When the player does backswing, the \u03b81 and the \u03b82 are decrease to the clockwise. And when the player does downswing, the \u03b81 and the \u03b82 are increase to the counter-clockwise" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0000216_ropec.2013.6702722-Figure4-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0000216_ropec.2013.6702722-Figure4-1.png", "caption": "Fig. 4. Representation of the components of the aerodynamic forces on a transversal section blade", "texts": [ " The regulation of the aerodynamical forces (thrust and Load) in the fixed wing propeller is done through the control of rotating velocity. The effect of the thrust on the fixed wing geometry can be described by the blade element theory, under following assumptions: 1) it acts like a small symmetrical blade (angle of attack, pitch and chord are not depending of the radius) 2) transversal section in the blade is constant. Then, the differentials of load force and thrust force are at the same pressure point, as shown in Fig. 4. In Fig. 4 dT, dH, dL and dD, are differentials of the thrust force, load force, lift and drag force, respectively. The expression to compute the lift and drag is presented in [13] ,)sin)(cos)(( 2 1 2drUCCcdT DL \u03b1\u03b1\u03b1\u03b1\u03c1 \u2212= (11) ,)cos)(sin)(( 2 1 2drUCCcdH DL \u03b1\u03b1\u03b1\u03b1\u03c1 \u2212= (12) where U is defined as (vi+\u03c9mr); and \u03b1 =arctan(vi/\u03c9mr). CL and CD are the coefficients of lift and drag respectively. The coefficients for lift and drag are dependent of the angle of attack. In order to tuning these parameters, it is supposed that the curves describing the dynamic of the coefficients for lift and drag are centered at origin and are approximated as follows: 1) CL=l0+l1\u03b1 (l1 is the slope of the lift curve) and 2) CD=d0+d1\u03b1+d2\u03b12, which is a quadratic function versus the angle of attack \u03b1" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0000954_s1001-6058(10)60054-6-Figure5-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0000954_s1001-6058(10)60054-6-Figure5-1.png", "caption": "Figure 5 illustrates the resolved vortices by different modeling approaches. One is SA-DDES by Spalart and another is our URANS based on SST. As expected, SST-DDES are able to resolve much richer vortex motions, whereas the URANS simulation has seemingly claimed rather \u201cstiff\u201d shear layers after the flow is detached from bluff-body surfaces. The main reason is the excessively viscous when URANS is applied to the massive separation flow. Then the future work is to calculate this flow using advanced RANS/LES hybrid methods with higher-order and low-dissipation scheme.", "texts": [], "surrounding_texts": [ "The tandem-cylinder is a prototype for interaction problems commonly encountered in airframe noise configurations (e.g., the oleo and hoses on a landing gear). The flow has been studied in a series of experiments performed in NASA Langley Research Center[2,3]. Simulation of TCs can help testing the capability of turbulence modeling approaches, spatial and temporal methods to reproduce properly complex flow phenomena, such as the transition on the two cylinders, separation of turbulent boundary layer, free shear layer instability, the interaction of unsteady wake of the front cylinder with the downstream one and unsteady massively separated flow between the cylinders and in the wake of the rear cylinder, etc. In this article, typically unsteady-RANS and advanced DDES with higher-order low-dissipation scheme are applied to calculate the unsteady pressure fields. The mesh provided by NTS is shown in Fig.7. The total size of the grid in the XY plane is 82,000 cells. The grid size between the cylinders is almost isotropic (about 0.01D) to calculate the small scale vortices. The spanwise size of the domain for Roe scheme are 1D and 4D with \u0394z=0.033D and for STVD is 3D with \u0394z=0.02D. Then the overall grids are 2.78 million for 1D, 10.9 million for 4D and 12.4 million for 3D, respectively. At the time of submitting this paper, the DDES computation with low-dissipation and highorder scheme for the last grids is under-going and only some instantaneous results are presented here. the surface. Longer spanwise length can obtained more reasonable results. Of course, to achieve the well matched results with the measurements, longer spanwise length should be applied. As we known, the flow is extremely unsteady. Then the pressure fluctuations on the surface or in the wake should be investigated to verify the calculation methods. The root mean square (RMS) of the pressure on the rear cylinder is presented in figure 9a. From this figure, we can see that the smaller spanwise calculate the Cp,rms much higher results than that of 4D and experiments. At the same time, we can find that the Cp,rms shows two peak values where \u03b8 are 45\u00b0 (due to vortices shedding from the front cylinder) and 120\u00b0 (due to the separation), respectively. Figure 9b presents the comparisons on frequency and amplitude of the pressure fluctuation for the sample where \u03b8 is equal to 45\u00b0 on the rear cylinder surface. From this figure, the results of frequency, especially the amplitude of the power-spectral-density (PSD) for 4D spanwise length can match the measurements much better than that of 1D spanwise length. And the primary frequency is about 180Hz, which is corresponding to the vortices shedding frequency from the front cylinder and also is the frequency of impingement on the rear cylinder surface by the former-mentioned vortices. (a) (b) Fig.9 Cp,rms and PSD at location of 45 degree on rear cylinder As mentioned before, the dissipation of the numerical scheme has a significant influence of the information of small scale of structures. The high-frequency pressure fluctuations are easily cut off when large numerical dissipation schemes are used to calculate the unsteady flow. Sometime, the numerical dissipation can exceed the physical viscosity. Then, the advanced DDES combined with large dissipation scheme maybe can\u2019t obtain acceptable results and their performance look like those of URANS. Figure 10 presents the comparisons on spanwise vorticity (with the same color legends) among numerical simulations using MUSCL, S6WENO5 with lowdissipation and measurements. Fig. 9(a) is the result of Roe scheme with MUSCL interpolation, 9(b) is the results of S6WENO5 with low-dissipation and 9(c) is the experimental results. From the comparisons, although both numerical simulations are DDES, low-order upwind MUSCL scheme smoothes almost all the small scale structures and only extreme large structures are reserved. The excited results can be obtained using S6WENO5 with \u03b5=0.12. The spanwise vorticity can match the measurements much better than that of low-order scheme. Plenty of small scales of turbulence structures are captured although the vortices breakdowns a little downstream in the gap region. Figure 11 demonstrates four snapshots of iso-surface of Q, which is defined as -(Sij 2-Wij 2) and its value is - 60. From these figures, very small scale of 3-D spanwise structures are presented colored by streamwise velocities. It visibly displays the capability of the simulations to resolve fine-grained turbulence (consistent with the grid used), and exhibits the complex although in general, similar vortical structures calculated by the different simulations. CONCLUSIONS Preliminary results about the flows are calculated using URANS around the rudimentary landing gear. At he same time, flows around the tandem cylinders are deeply investigated using DDES with different scheme and acceptable results, especially the instantaneous structures are obtained with high-order scheme with low-dissipation. a) Our in-house code (UNITS) is successfully applied to calculate the complex unsteady flows around the RLG configuration based on arbitrary multi-blocks of structured mesh. URANS can present reasonable pressure distribution on the RLG and it provides the initial flow-fields for the advanced turbulence modeling methods, such as DDES or iDDES in the future. b) DDES could be used to simulate the unsteady flow around the tandem cylinders, and spanwise length has a great effect on the mean pressure coefficients. c) The spanwise length is also have a significant influence on Cp,rms and PSD on the cylinders surface. Larger spanwise length is more reasonable because Lz is equal to 14D in experiment. d) The order of scheme should be increased and the dissipation of scheme should be decreased when the spanwise correlation parameters and information, such as small vortex structures, Cp,rms, and so on, are hoped to calculate accurately. e) Transition trip on the front cylinder surface is applied in experiment but fully turbulence is assumed in our computations. Therefore, more attention on the fix-transition should be paid and more reasonable results are hoped to achieve. Fig.11: Q=-60 iso-surface around tandem cylinders" ] }, { "image_filename": "designv11_62_0002922_edpe.2017.8123247-Figure10-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0002922_edpe.2017.8123247-Figure10-1.png", "caption": "Fig. 10 Rotor and stator of measured switched reluctance motor", "texts": [], "surrounding_texts": [ "This work was supported by the Slovak Research and Development Agency under the contract No. APVV-15-0750. The results of this project were obtained through the financial support of the Ministry of Education, Youth and Sports in the framework of the targeted support of the \"National Programme for Sustainability I LO1201\" and the OPR&DI project \"Centre for Nanomaterials, Advanced Technologies and Innovation CZ.1.05/2.1.00/01.0005\"." ] }, { "image_filename": "designv11_62_0002230_ut.2017.7890291-Figure1-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0002230_ut.2017.7890291-Figure1-1.png", "caption": "Fig. 1. The concept of recursive Newton-Euler algorithm", "texts": [], "surrounding_texts": [ "The redundant UVMS has various combinations of joint velocities that do not affect the given velocity profile of an end-effector and this could induce the self-motion of the vehicle. Hence, by using the degrees of redundancy, the desired trajectories of vehicle and joints can be determined without affecting its motion in the task space. When the velocity profile of an end-effector is given, the joint velocities of a manipulator and the velocity of an underwater vehicle can be determined by the redundancy resolution. ( ) ( ) ( )( )# # 3 3 B t\u03b7 \u03b3\u00d7\u0398 = \u0398 \u22c5 + \u2212 \u0398 \u0398J I J J (6) where J# is the Moore-Penrose pseudo-inverse matrix of the Jacobian matrix J and this can be calculated as according to Eq. (7). ( ) ( ) ( ) ( ) 1# T T \u2212 \u0398 = \u0398 \u0398 \u0398 J J J J (7) The first term in Eq. (6) is the particular solution, which is determined from the given velocity of an end-effector. The second term in Eq. (6) is the homogeneous solution, which is obtained by the projection onto the null-space of Jacobian matrix J. It represents the self-motion which does not affect the task motion. Hence, the vector \u03b3 can be defined in order to obtain the optimal solution for the specified performance index. To secure the stability of the system, it is required to move the object mass to the mass center of the whole system. ZMP algorithm is to stabilize the underwater vehicle by controlling the mass and moment of inertia of the manipulator links and the object to the mass center of the UVMS. In order to put the position calculated by ZMP algorithm, the distance between the mass center position of the UVMS and the ZMP position is set as the performance index as in Eq. (8). In this paper, the performance index is minimized in conjunction with a redundancy analysis, ( ) 2 2 1 1 1 1 1 , ( ) 2 2p zmp zmp p zmp xr x\u03b7\u039e \u0398 = \u2212 \u0394 = \u2212 \u0394W W (8) where Wp1 represents the weight matrix. The manipulator of the UVMS is redundant such that additional redundancy except the used one in deciding ZMP position can be used to other jobs such as the trajectory planning. For this kind of jobs, additional performance index for the planning of manipulator joint angle trajectories can be created using another performance index, as shown in Eq. (9), ( ) 2 3 , 2 2 ,int1 ,max ,min ,max ,min , ,int 1 2 , 2 2 \u03b8 \u03b8 \u03b8 \u03b8 \u03b8 \u03b8 \u03b8 \u03b8 \u03b8 = \u2212 \u039e \u0398 = + \u2212 = = i i mid p ii i i i i i mid i W (9) where \u03b8i,max and \u03b8i,min represent maximum and minimum limit angle of the i-th link respectively, and Wp2 represents the weight matrix. In the paper, the limit angle of the joint angle was assigned as a constraint considering actual application of the manipulator. To satisfy both constraints, a performance index is proposed that manipulator joint angle trajectory can be created to improve the stability of the whole UVMS within the limited range of manipulator joint angle[4], which is expressed as Eq. (10), ( ) 2 3 ,2 1 2, ,int1 1 1 ( ) 2 2 i i mid p zmp x p i ii x \u03b8 \u03b8 \u03b8= \u2212 \u039e \u0398 = \u2212 \u0394 + W W (10) The solution in null space can be obtained by conducting a redundancy analysis using Gradient Projection Method(GPM) for the proposed performance index as shown in Eq. (11). 1 2 ; , , , T N \u03b3 \u03ba \u03b8 \u03b8 \u03b8 = \u2212 \u22c5\u2207\u039e \u2202\u039e \u2202\u039e \u2202\u039e\u2207\u039e = \u2202 \u2202 \u2202 (11) where \u03ba and \u2207\u039e represent a positive gradient constant and the gradient vector of performance index respectively. 4. Experiments 4.1 System composition To implement the proposed ZMP algorithm and redundancy resolution method, a testbed composed of a floating vehicle which represent simple underwater vehicle attached with a redundant manipulator was developed. As shown in Fig. 2(b), this manipulator has three DOF (pitch-pitch-pitch) pitching redundant motion. To drive the each joint of the manipulator, 3 motors and its drivers are installed in the control box which can be waterproof. And the driving forces for each joint are transferred by pulley-timing belt mechanism from the motor to each joint. To measure the pitch-angle and pitch-acceleration, attitude and heading reference system(AHRS) is installed in the control box. The floating vehicle is composed of buoyancy materials and equipped with the developed manipulator. It floats on the water freely and allows to see how manipulator\u2019s motion affect to the stability of the whole system. One directional motion using the reduced mini-sized floating vehicle was implemented for the proposed algorithms. The stability of the system is identified by distance between the mass center position of the UVMS and the ZMP position in X axis(Xzmp). Table 1 shows the specifications of the floating vehicle and the manipulator. In order to control the manipulator with 3 DOF, a control system was developed as shown in Fig. 3. Control system consists of a main controller, and joint motor controller, motor driver, AHRS sensor. The main controller based on ARM processor does a role of editing the control algorithm, sending the order of operation signals to the motor controller for the joint actuators of the manipulator through CAN communication. The main controller sends continuous motion signals to motor controllers of the joint actuators according to a trajectory planning. 4.3 Trajectory following experiment To verify trajectory following performance of the redundancy resolution(RR) algorithm, an optimization for the proposed performance index as Eq. (10) was conducted. For comparison with the RR algorithm, a pseudo inverse (PI) algorithm was also applied to the testbed with the same initial conditions and desired trajectory. For both cases, the end-effector is commanded to follow a line from the initial position(0mm, 550mm) to the destination position(100mm, 550mm) in the X-Z plane. This command excites the stability of the manipulator attached at the floating vehicle. Since the excitation generates the pitching motion of the system, the mass center of the whole system should be controlled to move the mass center position. Under this condition, the RR and PI algorithm with ZMP algorithm were applied to control the system. According to the test results, Fig. 4 indicates that both algorithms generate similar and acceptable tracking errors in the PI and RR cases with 5 mm position error in Z axis. And the second line in Fig. 4 represents the distance between the mass center position of the whole system and the Xzmp. If Xzmp approaches to zero, the whole system goes to the stable state. So, the plot shows that application of the PI case does not decrease Xzmp during following the desired trajectory. On the other hand, the RR algorithm makes Xzmp decrease monotonically. Figure 5 shows configurations of manipulator trajectory using PI and RR in X-Z axis. In the case of using RR, each joint changes the joint angle to reduce the Xzmp with keeping the end-effector position. The result shows that the RR with the ZMP algorithm improves the stability of UVMS while forcing the manipulator to follow the trajectory and performs other tasks using redundancy according to constraints of the performance index." ] }, { "image_filename": "designv11_62_0003526_6.2019-3489-Figure2-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0003526_6.2019-3489-Figure2-1.png", "caption": "Fig. 2 SBPW2 JWB geometry.", "texts": [ " This pressure signature is used as an input for a propagation algorithm, which simulates the propagation through the atmosphere and calculates the pressure signature on the ground. Finally, the loudness is calculated from the ground pressure signature. This section describes the programs and tools used for the simulations and the optimization in detail. Near Field Pressure Signature Ground Pressure Signature Fig. 1 Prediction of the sonic boom. The second Sonic Boom Prediction Workshop provided different geometries and grids. One of the geometries is the JAXA Wing Body (JWB), shown in Fig. 2. It was chosen for the study because it has an increased complexity compared to the axisymmetric body used for the preceding study and it can be considered an interim stage for a full aircraft configuration. Due to the threedimensional influence of the wing on the pressure field the assessment of off-track pressure signatures also becomes relevant. The JWB geometry was designed by Ueno et al. [24] to match the equivalent area of the NASA concept 25D [24]. The nose of the wing body geometry is axisymmetric" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0002082_978-3-319-49137-0_1-Figure3-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0002082_978-3-319-49137-0_1-Figure3-1.png", "caption": "Fig. 3 The channel electrode luminescence cell from Ref. [52]", "texts": [ " Their configuration permitted detection of the emitted light at 90\u00b0 and avoided scattering effect from the faces of the cell. Unfortunately, the equilibration time is longer than the one required for OTTLE cells and shows the pseudo-thin-layer nature of the long optical path electrochemical cells. However, bunch of experiments can be easily performed with this type of cells, particularly for stable electroactive species. Following on from cells mainly limited to stable species, OTTLE cells have been redesigned by Compton et al. to work as flow cells (Fig. 3), in hydrodynamic voltammetry experiments, especially for complex electrochemical processes [52\u2013 54]. Hydrodynamic voltammetry techniques using such cells facilitate the analysis of fluorescent solution-phase electrogenerated species, by modeling, from a proposed model, their spatial and temporal distributions in the flow cell, and finally correlate the fluorescence intensity to the electrode current. In the beginning of 2000s, a new evolution of thin-layer electrochemical cells, able to operate at variable thicknesses, was designed" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0000047_s13272-013-0095-7-Figure1-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0000047_s13272-013-0095-7-Figure1-1.png", "caption": "Fig. 1 HALAS system components", "texts": [ " Furthermore, the modelling of the slung load system used for the later controller design is presented in combination with a first system analysis. The stability analysis includes parameter variations of cable length, load mass and load suspension point position. The paper closes with a brief summary and gives an outlook on the next planned steps in this research project. The pilot assistance system HALAS will be demonstrated during the flight tests with DLR\u2019s research helicopter ACT/ FHS. The HALAS system consists of several components (see Fig. 1) which are described in detail in the following sections. DLR is operating a unique airborne research rotorcraft, the ACT/FHS [36] (see Fig. 2). It is based on an EC135 helicopter. The main feature is a combination of a fullauthority, quadruplex fly-by-wire/light control system with smart actuators and a simplex experimental system. It completely replaces the original flight control system and provides a high degree of flexibility. The experimental system consists of three computer systems, additional sensors, and two stations, one for the evaluation pilot and one for the flight test engineer (see Fig" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0000986_j.proeng.2011.03.137-Figure10-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0000986_j.proeng.2011.03.137-Figure10-1.png", "caption": "Figure 10: A planar mechanism.", "texts": [ " In figures 5-9 we can see the independent movements Tx, Ty, Tz, Rx, Ry of the extreme element 1 of the open chain 0-3-4-2-6-5-1, associated to the loop. So, the extreme element 1 of the open chain, associated to the loop, has the spatiality 5 (Tx, Ty, Tz, Rx, Ry); b1=5. When we rejoin the kinematic chain we observe that each degree of freedom of every joint modifies the movement of the next kinematic element, and so, the number of passive degrees of freedom of the joints becomes 6. An example of mobility calculus of a mechanism with one passive element We consider the mechanism shown in figure 10, where AB=DC, AB//DC, BC=AD, BC//AD. Let\u2019s suppose that all the elements are disjoined and become free in space, except one. Only one segmented frame remains fixed, the second becomes mobile (Figure 11). After segmentation and fictional motion in space (or in plane) the number of temporarily mobile elements becomes equal to the number of kinematic joints, i.e. four. In this phase, the number of degrees of freedom of the system is 6 4, for a spatial movement, or 3 4, for a planar movement. We rejoin the elements by rotational joints of V class, including the temporarily segmented frame (Figure 12) and the constraints of the joints are eliminated 5 4 and 2 4 respectively. The extreme element (0) of the open chain has the spatiality three: Rx, Ty, Tz (whether for a spatial or a planar movement); in order to compose the mechanism (Figure 10), three degrees of freedom will be eliminated. If we calculate the mobility of the mechanism by using Eq. (5), for a spatial temporary movement, we obtain the mobility 1 (Eq. (8)) and for a planar temporary movement we obtain the same result (Eq. (9)). M=6m-5p-b1= 134546 (8) M=3m-2p-b1= 134243 (9) The mobility of the mechanism is one. A point E, situated in the middle of the element BC, describes a circular trajectory, with the dimension of the radius equal to the dimension of the element AB. We join an element EF on the point E (Figure 14)" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0002774_pvp2017-65992-Figure6-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0002774_pvp2017-65992-Figure6-1.png", "caption": "FIGURE 6. AXIAL STRESS ALONG THE LONG AXIS DURING THE FIRST PASS, MIDWAY THROUGH THE SECOND PASS, AND AFTER COOLING FOR VARIABLE LASER POWER THIN WALLED BUILD. CORRESPONDING TEMPERATURE PROFILES ARE SHOWN FIGURE 6.", "texts": [ " The 2000 W laser model showed a larger melt pool, and significantly higher temperatures in both the deposited region and the substrate compared to the 500 W model (Figure 5). For the variable laser power thin walled build model, stresses began to build during deposition, and slightly increased in magnitude following cooling to room temperature. Generally, the magnitudes of axial stress increased once reaching the final build state at room temperature as compared to during the build 4 Copyright \u00a9 2017 ASME Downloaded From: http://proceedings.asmedigitalcollection.asme.org/ on 11/10/2017 Terms of Use: http://www.asme.org/about-asme/terms-of-use (Figure 6). Axial stress magnitudes were lower near the beginning and end of each deposition pass as compared to the middle of the pass. Temperature profiles with the variable laser power thin walled model show a relatively constant and small melt pool size and some substrate heating near the melt pool (Figure 7). For the variable laser power thin walled build model, approximately 61% of the volume of the substrate remained within 50 K of ambient temperature. The axial stress solution shows apparent periodicity that corresponds with the solution time step", " Several obstacles to quick and precise LENS simulation remain, and must be addressed in parallel with experimental validation prior to accurate process simulation. For some simulations, inverted elements due to high thermal gradients occurred. Alternative time stepping methods and/or adaptive meshing may do a better job of resolving these high gradient regions without significant additional computational expense. Variable time stepping in which time step size is a function of laser beam diameter and laser speed may reduce periodicity in the solution (Figure 6, 7). Additionally, there exists an inherent mesh dependence in this activation scheme where in sub element sized features are not captured. Additional work is underway to better understand and quantify the effects of mesh resolution in both the deposition block and the substrate. Significant computational resources are required for these simulations which may present an obstacle when simulating larger LENS builds with the fidelity required to resolve the high stress and/or temperature gradients" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0000431_amr.566.197-Figure1-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0000431_amr.566.197-Figure1-1.png", "caption": "Fig. 1 Experimental setup and samples:", "texts": [ " In the investigation of surface and subsurface cracking mechanisms in metals fracture mechanics is used. A typical mechanism is described later in text of this work. However, these well studied models cannot be applied for plastic bearings as the mechanical and physical characteristics of metals and polymers differ significantly. In the present work, microscopic holes were introduced onto the surface of bottom bearing parts, and cracks growing from these artificial defects were observed. RCF tests were conducted using a thrust type machine shown in Fig. 1 (a). All tests were carried out in water-lubricated conditions as shown in Fig. 1(b). The water in the tank circulated continuously throughout the test. Fig. 1 (c) is a photo of the top bearing race and a retainer with alumina balls. Retainers with nine alumina balls were used. The ball diameter was 9.525 mm (3/8\u201d). The geometry of the bearings was standard #51305 (JIS B 1513). The dimensions were: outer diameter of 52mm, an inner diameter of 25 mm and pitch circle diameter of 38.5mm. Fig. 1 (d) is a photo of the bottom race. The arrows indicate the position of the artificial holes drilled by end mill. Fig. 1(e) shows the hole dimensions: the diameter of 300\u00b5m, the depth of 180\u00b5m. Both the races and the retainers were made of unreinforced PEEK (VICTREX \u00a9450G). The physical and mechanical properties of PEEK are shown in Table 1. All rights reserved. No part of contents of this paper may be reproduced or transmitted in any form or by any means without the written permission of Trans Tech Publications, www.ttp.net. (ID: 141.211.4.224, University of Michigan Library, Media Union Library, Ann Arbor, USA-14/07/15,01:57:31) 200N, 100N, and 50N until reaching 1" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0001363_rast.2011.5966982-Figure8-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0001363_rast.2011.5966982-Figure8-1.png", "caption": "Figure 8. The projections on t-z plane.", "texts": [ " The circle with radius d1 is expressed by the following equation ( ) ( )22 2 1 1 1 (9)x yx p y p d\u2212 + \u2212 = The following equation is written to define the intersection on x-y plane: 1 2 2 2 1 1( ) ( ) (10)x yx p mx k p d\u2212 + + \u2212 = This quadratic equation possesses two reel roots in the case \u0394>0. These reel roots correspond to the point t1 and t2 enabling to calculate the radius of the circle located on the t-z plane which is the projection of the sphere centered at P1. The radius of the circle is given by the following equation: 2 1 (11) 2 t t r \u2212 = To determine P2, an additional intersection on the t-z plane shown in Fig. 8 between the circle with radius r and the circle with radius r2 must be existed. This intersection occurs when the following relation is satisfied: 2 2 (12)r r l r r\u2212 \u2264 \u2264 + The radius r2 of the circle centered at O2 is given by the subsequent equation; 2 2 2 3 3 (13)= \u2212 br L L while the distance Lb3 between B3 and O2 (see Fig. 4, 6 and 7) is defined as the following: 2 2 2 3 4 3 (14) 2 + \u2212=b L L LL L \u03b8 is the angle between the line P1 \u0131\u0131O2 and t axis. \u03b8 is given by the following equation: 1 0 tan (15) \u239b \u239e \u03b8 = \u239c \u239f\u239c \u239f \u239d \u23a0 zp a t where 1 2 0 2 (16)+ = t t t l, a,b and h are the distances, as shown in Fig. 8. These expressions include the following relations: 2 2 0 1 (17)+= zt Pl 2 2 2 2 2 (18)+ \u2212 = r l r a l (19)= \u2212b l a 2 2 (20)= \u2212h r a The projection on t-z plane of vertex P2 (p2x, p2y, p2z) is 2 2 2( , )\u0131\u0131 \u0131\u0131 \u0131\u0131P tp zp and 2 \u0131\u0131tp , 2 \u0131\u0131zp are given by the following equations: 2 cos( ) (21)= \u03b8\u0131\u0131tp b 2 sin( ) (22)= \u03b8\u0131\u0131zp b The coordinates of points A and B on t-z plane are 2 .sin( ) (23)= + \u03b8\u0131\u0131 At tp h 2 .cos( ) (24)= + \u03b8\u0131\u0131 Az tp h 2 .sin( ) (25)= \u2212 \u03b8\u0131\u0131 Bt tp h 2 .cos( ) (26)= \u2212 \u03b8\u0131\u0131 Bz tp h The coordinates (x03, y03) of O2 are given by the following equations: 02 23 3 cos( ) (27)= + \u03c0 \u2212 \u03b1b bx x L 02 3 23 sin ( ) (28)= + \u03c0\u2212 \u03b1b by y L where (xb3, yb3) are the coordinates of B3" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0000359_1.4004623-Figure2-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0000359_1.4004623-Figure2-1.png", "caption": "Fig. 2 Thermal model of shell-to-channel bolted joint Fig. 3 Flange sector with bolt", "texts": [ " The first flange is welded to the cylindrical shell, whereas the other one to the channel, while the tube sheet is installed in between. One gasket is installed on each side of tube plate in order to provide a tightness of the bolted joint. Every flange includes a hub and a ring. These two parts are forged together. The analytical proposed model considers heat flow through each element of the joint assembly to estimate the temperature distribution. Every element is represented by a simplified geometry the thermal behaviour of which is described by a resistance obtained using heat transfer principles. Figure 2 shows the developed heat transfer model of a heat exchanger bolted joint. The heat flow through the 011207-2 / Vol. 134, FEBRUARY 2012 Transactions of the ASME Downloaded From: http://pressurevesseltech.asmedigitalcollection.asme.org/ on 01/28/2016 Terms of Use: http://www.asme.org/about-asme/terms-of-use material of the assembly occurs by conduction, whereas the cooling of the flanges by the ambient air takes place by convection. In the same way, the fluid in contact with internal walls transfers heat to the assembly walls by convection. It is to be noted that all heat flows related to the shell side of the bolted joint are marked with a star. The upper and lower parts of Fig. 2 represent, respectively, the channel and the shell of the heat exchanger. The bulk temperature of the fluid inside the channel is Tc, whereas the bulk temperature inside the shell is Ts. Heat is transferred to the internal wall of the bolted joint by convection. The external walls of the bolted joint are cooled by ambient air. The heat flow Q1 from the hotter fluid to the internal walls of the channel is divided in two flows Q6 and Q3. While the flow Q3 dissipates in the atmosphere by convection, the flow Q6, through the hub, is transferred to the ring by conduction" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0001609_0954406211427833-Figure6-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0001609_0954406211427833-Figure6-1.png", "caption": "Figure 6. Initial position of cutter and gear blank for a concave gear: (a) centre tooth: cutter axis and gear blank axis in the same line; (b) teeth at the inner ring: gear blank axis at 12 with the cutter axis; and (c) teeth at the outer ring: gear blank axis at 24 with the cutter axis.", "texts": [ " Since the cutter reaches its at University of Ulster Library on May 14, 2015pic.sagepub.comDownloaded from extreme positions 16 times in one revolution of the work blank, 16 crests are formed on the periphery of the tooth surface generated by CAD simulation. Generation of tooth on concave gear Each tooth of the concave gear has to be machined individually. Initially, the cutter is placed in such a way that the reference plane of the cutter lies at a position tangential to the reference sphere of the gear blank, as shown in Figure 6(a). While simulating the centre tooth, cutter and gear blank axes are in a horizontal direction. Different incremental motions applied to cutter and gear blank during the simulation as presented in Table 2. If a smoother surface is required, the number of steps for one complete rotation about the Z-axis has to be increased and this will considerably increase the simulation time. For simulation of machining of the teeth which are at the inner ring, the work solid is tilted by 12 about the X-axis with respect to its sphere centre (Figure 6(b)). Similarly, for simulating the next teeth at the inner ring, the work solid is first indexed by 60 about the Z-axis and then tilted by 12 about the X-axis. The same steps are followed for the rest of the teeth at the inner ring. For machining the teeth which are at the outer ring, the work solid is tilted by 24 about the X-axis with respect to the sphere centre (Figure 6(c)). Similarly, for simulating the other teeth at the outer ring, initially, the work solid is first indexed by 30 about the Z-axis and then tilted by 24 about the X-axis. The same steps are followed for the rest of the teeth at the outer ring. CAD model after simulation for the centre tooth of the concave gear is shown in Figure 7. By highlighting the tooth surface, the curvature on it can be seen distinctly. Figure 8(a) shows the CAD-simulated profile obtained by taking an axial section of the centre tooth along with the superimposed profile obtained analytically following the procedure outlined in \u2018Analytical method of tooth surface generation\u2019 section" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0003814_codit.2019.8820321-Figure2-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0003814_codit.2019.8820321-Figure2-1.png", "caption": "Fig. 2. Quadrotor and reference frame.", "texts": [ " The authors based the predictions on learning from prior experience. Later, the prediction system was incorporated into the flight control loop as a feedforward term on the throttle command [21]. They showed a significant improvement for some common obstacles. However, the prediction system is prone to either underestimate or overestimate the disturbance produced by dissimilar or unseen obstacles. We consider a quadrotor that is assumed to be symmetrical and the center of mass lies in its geometric center. The body frame (xb, yb, zb) is defined in Figure 2, and the origin of this frame is expressed in a world frame (xw, yw, zw) fixed in inertial space. Since there are no large deviations from hovering in attitude, the ZYX Euler angles are used to define the roll, pitch and yaw angles (\u03c6, \u03b8, \u03c8). In order to build the dynamic model, we should obtain physical properties like the mass, the moment of inertia matrix and motor specifications. However, for simplicity and safety, off-the-shelf quadcopters are usually sold with inner stabilization system (Figure 3), which means that there is no way to control the angular speed of motor directly" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0002777_aer.2017.106-Figure1-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0002777_aer.2017.106-Figure1-1.png", "caption": "Figure 1. (Colour online) Illustration for helicopter states and coordinates.", "texts": [ " Finally, the main conclusions of this paper are remarked in Section 5. 2.0 SYSTEM MODELLING AND IDENTIFICATION In this paper, Mettler\u2019s linearised helicopter model(20,21) is adopted to describe the dynamics of the TREX-600 small-scale unmanned helicopter. The accurate model parameters are obtained by the time-domain system identification method based on our helicopter\u2019s flight data. This helicopter model consists of 6-Degrees-of-Freedom (DOF) rigid body dynamics and extra servo/rotor dynamics. As illustrated in Fig. 1, the dynamic motion of the helicopter is defined by the state vector x = [u v w \u03b8 \u03c6 q p as bs r]T in the body frame which contains the fuselage velocities (u, v, w), the angular rates (p, q, r), the main rotor flapping angles (as, bs) and the attitude angles (\u03c6, \u03b8, \u03c8). The state-space model governed by Equation (1) contains a system matrix with 18 unknown parameters and a control matrix with 7 unknown parameters. This is obviously a multi-variable linear system with strong coupling. The state vectors are physically shown in the body frame, and the input vectors can be written as u = [\u03b4col \u03b4lat \u03b4lon \u03b4ped ]T , where \u03b4col is the collective pitch and impacts heave motion (along ZB-axis); \u03b4lat is the lateral cyclic and impacts roll motion(around XB-axis) then the lateral translation (along YB-axis); \u03b4lon is the longitudinal cyclic and dominates pitch motion (around YB-axis) then the longitudinal translation (along XB-axis); \u03b4ped is the tail rotor input and affects yaw motion (around ZB-axis) then the heading D ow nl oa de d fr om h tt ps :// w w w " ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0002208_j.proeng.2017.02.180-Figure1-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0002208_j.proeng.2017.02.180-Figure1-1.png", "caption": "Fig. 1. Diagram and beam model of a hydraulic cylinder with damping.", "texts": [ " Nomenclature T kinetic energy V potential energy \u03b4WN virtual work of non-conservative forces originating from damping Wmn (xmn, t) transverse displacement of beams that model cylinder and piston rod Umn (xmn, t) longitudinal displacement of beams that model cylinder and piston rod Emn Young's modulus for individual beams E* mn material viscosity coefficient (internal damping) Jmn moment of inertia in beam cross-sections Amn cross-sectional areas of the beams mn beam material density ce external viscous damping coefficient (ce = 0 for m = 2 and n = 1) cR constructional viscous damping coefficient \u03bc dimensionless constructional damping coefficient \u03b7 dimensionless internal damping coefficient \u03bd dimensionless external damping coefficient lmn length of the beams LC total length of the hydraulic cylinder (LC = l11+ l12+ l22) P cylinder loading force (P = 0 at the length l12) PC critical load of the cylinder \u03c9* the complex eigenvalue of the system (\u03c9* = Re(\u03c9*) + Im(\u03c9*)i) kR stiffness coefficient of rotational spring k dimensionless stiffness coefficient of rotational spring kS stiffness coefficient of translational spring xmn spatial coordinates t time 1i m,n = 1,2 A scheme of the considered system is presented in Fig. 1. In adopted four-beams-model, two of the beams model a cylinder tube (l11, l12) and two - piston rod (l21, l22) in the cylinder. The liquid in the cylinder was adopted as the medium of load transfer between the piston and the cylinder along the length filled with liquid. The liquid rigidity in the cylinder was modelled by the translational spring kS. Formulation of the boundary problem was carried out using the Hamiltion's principle: 0)( 2 1 2 1 t t N t t dtWdtVT (1) The vibration equations for individual beams are known and have the following form: 0 ),( ),(),(),( 2 2 2 2 4 4 * t txW c t txW A x txW P x txW t EEJ mnmn e mnmn mnmn mn mnmn mn mnmn mnmnmn (2) and: 0 ),(),( 2 2 2 2 * t txU A x txU t EEA mnmn mnmn mn mnmn mnmnmn (3) Solutions of equations (2) and (3) are in the form: ti mnmnmnmn exwtxW * )(),( (4) ti mnmnmnmn exutxU * )(),( (5) Substitution of (4) and (5) into (2) and (3) leads to, respectively: 0)()()( 2 mnmnmnmn II mnmnmn IV mn xwxwxw (6) 0)()( 2 mnmnmnmn II mn xuxu (7) where: *2* ** )( i A c iEEJ A mnmn e mnmnmn mnmn mn , )( ** 2* iEE mnmn mn mn , mnmnmn mn JiEE P )( ** (8) Geometrical boundary conditions and continuity conditions are given by: 0)((0))((0) 222221121211 lwwlww IIII , )0()( 121111 II wlw , )0()( 222121 II wlw , (0)(0))( 21121111 wwlw , (0))()( 2221211212 wlwlw , (0))( 222121 ulu , (0))( 121111 ulu , 0)0(11u , )0()0( 1221 uu (9) The natural boundary conditions of the system studied: )0()0()0()( 11 * 111111 ** 1111 I R I R II wicwkwJiEE , )()((0))( 111111 ** 11111212 ** 1212 lwJiEEwJiEE IIII , )()()()( 2222 * 2222222222 ** 2222 lwiclwklwJiEE I R I R II , )()()0()( 212121 ** 21212222 ** 2222 lwJiEEwJiEE IIII , 0)0()0()()0()()()()( 212121 ** 21211212 ** 12121111111111 ** 1111 IIIIIIIIIII PwwJiEEwJiEElPwlwJiEE , 0)0()()()()()( 2222 ** 2222212121 ** 2121121212 ** 1212 IIIIIIIII wJiEElwJiEElwJiEE , )0()()()( 2222 ** 2222212121 ** 2121 II uAiEEluAiEE , )0()0()( 212121 ** 2121 ukuAiEE S I , PluAiEE I )()( 222222 ** 2222 , )()((0))( 111111 ** 11111212 ** 1212 luAiEEuAiEE II (10) The solution of equations (6) and (7) are expressed in the form of functions: xi mn xi mn x mn x mnmn mnmnmnmn eCeCeCeCxw 4321)( (11) xi mn xi mnmn mnmn eCeCxu 65)( (12) where: mnmnmn mnmn mnmn mnmn mn , 42 , 42 4242 (13) The boundary problem is solved numerically for the eigenvalues \u03c9*" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0003356_978-981-13-5799-2_12-Figure12.114-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0003356_978-981-13-5799-2_12-Figure12.114-1.png", "caption": "Fig. 12.114 Physical meaning of the Mohr\u2013Coulomb failure criterion", "texts": [ "64) may not be satisfied in a tire. Equation (12.87) (Fig. 12.113) h\u00f0x\u00de \u00bc j\u00fe t w x h2 \u00bc 1 w Zw 0 h2\u00f0x\u00dedx \u00bc 1 w Zw 0 j\u00fe t w x 2 dx \u00bc j2 \u00fe j t\u00fe t2=3 Note 12.5 Eq. (12.99) The Mohr\u2013Coulomb failure criterion can be determined in shear tests for various normal stresses r. The Mohr\u2013Coulomb failure criterion is that when Mohr circle comes into contact with the line of Eq. (12.99), with plastic deformation occurring at this point. Suppose that the Mohr circle comes into contact with the Mohr\u2013 Coulomb failure criterion at point A of Fig. 12.114a and the angle between the raxis and the line connecting point A with the center of the Mohr circle is 2h (where the counterclockwise direction is positive). A(r, s) in Fig. 12.114a expresses the stress field in the coordinate system at angle h (clockwise direction) measured from the direction of the maximum stress r1. Hence, the shear stress s is applied on plane A\u2013A in Fig. 12.114b. Considering similarly at point A\u2032 expressed by A\u2032(r, \u2212 s) in Fig. 12.114a, the shear stress \u2212s is applied on plane A\u2032\u2013A\u2032 with an angle p/2 \u2212 h measured from the direction of the maximum stress r1 in Fig. 12.114b. According to internal friction theory, planes A\u2013A and A\u2032\u2013A\u2032 will become sliding planes. Note 12.6 Eqs. (12.102), (12.113) and (12.121) for the Heat Flow of a Semi-infinite Solid [160] The governing equation for the heat conduction of a semi-infinite solid is @T @t \u00bc a @2T @z2 0 z\\1; t 0; T \u00bc T0 0 z\\1; t \u00bc 0: \u00f012:222\u00de The solution to Eq. (12.222) is T\u00f0z; t\u00de \u00bc T0erf z 2 ffiffiffiffi at p ; \u00f012:223\u00de where erf(z) is defined by erf z\u00f0 \u00de \u00bc 2ffiffiffi p p Zz 0 e n2dn: \u00f012:224\u00de Suppose that initial condition and boundary condition are given by T \u00bc T0 0 t l=V ; T \u00bc Tm z \u00bc 0: \u00f012:225\u00de The solution to Eqs" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0003356_978-981-13-5799-2_12-Figure12.13-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0003356_978-981-13-5799-2_12-Figure12.13-1.png", "caption": "Fig. 12.13 Pressure distribution of tires having different fore\u2013aft spring rates in braking", "texts": [ " This is because the adhesion region in the driving condition is larger than that in the braking condition, because the pressure distribution in the driving condition is higher at the trailing edge as shown in Fig. 12.10. Figure 12.12 shows the calculation of the fore\u2013aft force of tires for different dynamic fore\u2013aft spring rates Rx. As Rx increases, the maximum tractive force decreases while the maximum braking force increases. The magnitude of the slip ratio for the maximum driving force is smaller than that for the maximum braking force. Note that the static fore\u2013aft spring rates Rx is discussed in 6.2.5. Figure 12.13 presents conceptual figures for the pressure distribution of tires with different fore\u2013aft spring rates in the braking condition to clarify the mechanism. When the same torque is applied to tires with large and small fore\u2013aft spring rates, the fore\u2013aft displacement of the tire with a large fore\u2013aft spring rate is smaller than that of the tire with the small fore\u2013aft spring rate. Hence, the slope of the contact pressure distribution of the tire with the large fore\u2013aft spring rate is less steep than that of the tire with the small fore\u2013aft spring rate" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0003071_iscid.2017.115-Figure1-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0003071_iscid.2017.115-Figure1-1.png", "caption": "Figure 1. The vehicle kinematic model", "texts": [ " In this paper, after considering the actual initial parking state, which always has a deflection angle to the parking space, a path planning algorithm is proposed for parallel parking from arbitrary initial angle. The path is generated based on two tangential arcs of different radii. And the feasibility of the proposed algorithm is verified on the MATLAB. II. THE VEHICLE KINEMATIC MODEL The vehicle is running at low speed when parking, so it is assumed to move with non-sliding. The vehicle structure mentioned in this paper is the Ackerman steering structure. The vehicle kinematic model is illustrated in Fig.1. In the reference coordinate system, the vehicle rotates with a certain radius at the center , which is the intersection of three perpendiculars of the direction of the front wheels and the vehicle. represents the distance between the midpoint of the rear wheel and the center. represents the distance between point B and the center. represents the distance between point C and the center. is defined as the steering angle. is defined as the orientation angle according to global coordinate system. represents the constant movement speed of the vehicle. 2473-3547/17 $31.00 \u00a9 2017 IEEE DOI 10.1109/ISCID.2017.115 55 As shown in Fig.1, the geometric relationships between , and are presented as follows: (1) Because of the vehicle structure constraints, the vehicle will rotate with the minimum radius when the front wheel gets the maximum steering angle. The minimum radius and the maximum steering angle are defined as and respectively. Simultaneously, the minimum radius of and , which are defined as and can be obtained. There have been many researches about the Ackerman steering structure of the vehicle [9], and the vehicle kinematic model is shown as follows (2) III" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0002860_mnl.2017.0515-Figure1-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0002860_mnl.2017.0515-Figure1-1.png", "caption": "Fig. 1 Photographs of the commercial drop cell with Pt thin-film electrodes and microfluidic chip holder a Drop cell and Pt thin-film electrodes b Microfluidic chip holder All the instruments were provided by MicruX Technologies", "texts": [ " Phosphate Micro & Nano Letters, 2018, Vol. 13, Iss. 3, pp. 302\u2013305 doi: 10.1049/mnl.2017.0515 buffer solution (PBS, pH 7.0) was used to disperse the AgNPs. All the solutions injected into the microchip channels were filtrated with the film of 0.22 \u03bcm. 2.2. Instrumentation: Electrochemical experiments were carried out in a drop cell (MicruX Technologies, Asturias, Spain) controlled by the CHI 660E Electrochemical Work Station (Chenhua Instruments, Shanghai, China). The drop cell with Pt thin-film electrodes was shown in Fig. 1a. Three Pt thin-film electrodes were used as the working, reference, and counter electrode. The MCE instrument (HVStat) was provided by MicruX Technologies (Asturias, Spain), which included the main instrument, reusable holder, and microfluidic chip with integrated electrodes. The main instrument unit integrated a high-voltage power supply with a maximum voltage of \u00b13000 V and a potentiostat for amperometric measurements. The system was connected to the computer by Bluetooth and controlled by MicruX Manager Software to allow the automatic control of the experiments, which simplifies the works with microchips. The instrument was complemented with a reusable microfluidic chip holder (MicruX Technologies, Asturias, Spain) inside which the microfluidic chip was placed provides an easy way for the usage of the MCE. Fig. 1b shows the photograph of the holder. All the electric contacts for the high-voltage, detection electrodes, and the reservoirs for buffer and samples were integrated in the holder. 2.3. Microfluidic chip design: The cross-shaped single-channel SU-8/Pyrex microfluidic chip with integrated electrochemical Micro & Nano Letters, 2018, Vol. 13, Iss. 3, pp. 302\u2013305 doi: 10.1049/mnl.2017.0515 detector based on three 100 \u03bcm Pt thin-film electrodes with a 100 \u03bcm gap among them (MicruX Technologies, Asturias, Spain) was employed" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0000915_046017-Figure1-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0000915_046017-Figure1-1.png", "caption": "Figure 1. \u2018Compact\u2019 arc deposition system inside a 1.6-cell RF superconducting electron injector.", "texts": [ " It is expected that the photocathodes prepared in such an optimized way will be further examined in terms of resonant quality and QE in a TESLA-like electron injector. Vacuum arc devices used in the early stage of National Centre for Nuclear Research (NCBJ) photocathode group activity 0031-8949/14/014071+05$33.00 1 \u00a9 2014 The Royal Swedish Academy of Sciences Printed in the UK included the following filters of micro-droplets: diaphragms in a \u2018compact\u2019 arc deposition system introduced directly into a 1.6-cell niobium cavity (figure 1) and a 60\u25e6 knee-type magnetic filter (figure 2). The former was used to coat a 3.5 mm diameter, thin-film lead photocathode directly on the rear wall of a RF cavity. The deposition device was contained in a grounded stainless steel capsule mounted inside the cavity. The top part of the capsule was terminated with a niobium mask placed at a short distance (<1 mm) from the back wall. The position and size of a circular opening in the mask determined the position and size of the photocathode to be deposited" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0003623_1.4044296-Figure7-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0003623_1.4044296-Figure7-1.png", "caption": "Fig. 7 CAD model showing the turbocharger with all load sensors installed", "texts": [ " A plastic adapter was fabricated to attach to the outer race with a light press fit. Weights hung from a beam attached to the plastic adapter applied a known torque to the outer race. The geometry of the calibration device, weight applied, and diameter of the outer race were used to calculate the force applied to the load cell. Figures 6(a) and 6(b) illustrate the calibration device and calibration curve. Overall Setup and Test Procedures. The two axial load sensors and the AR pin sensor allowed for simultaneous monitoring of the loads exerted in the TC during operation. Figure 7 presents a diagram of all three sensors installed in the TC. The data acquisition system included quarter Wheatstone bridge circuitry for measuring strain. The sampling rate used for data collection was 10 kHz. The top speed achievable by the TC in steady-state operation was approximately 55 krpm. Turbochargers experience a variety of operating conditions when coupled directly to an internal combustion engine including rapid acceleration, periods of heavy compressor loading, and coast downs when the engine load is rapidly reduced" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0001344_pedstc.2011.5742411-Figure5-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0001344_pedstc.2011.5742411-Figure5-1.png", "caption": "Figure 5. A rotor bar", "texts": [], "surrounding_texts": [ "This zone consists of the stator windings elements. The current density is assumed to have constant distribution in the cross section of the winding conductors. The field equations for the elements of this zone are stated in (24) and (25) [14]. 0 (24) f f f f f co S AU R I R N ds t \u03c3 \u2202= + \u2202\u222b (25) where, (26) The discretized forms of (24) and (25) are as (27) and (28), respectively. (27) \" \" (28) where, [ ]1 1 1 6 TTco co f fs N N DP ds S S \u03c6= =\u222b (29) [ ]\" 1 1 1 6 co co f fs N l N lDQ ds S S \u03c6= =\u222b (30)" ] }, { "image_filename": "designv11_62_0001482_iros.2011.6095088-Figure3-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0001482_iros.2011.6095088-Figure3-1.png", "caption": "Fig. 3. Local evolution of the backbone curve at a point is defined by an instantaneous rotation of magnitude along an axis at an angle to the axis in body coordinates.", "texts": [ " A side effect of this constraint is that the deformation of a whole sheet can be described by the bending occurring on a single backbone drawn across the sheet, as shown in Fig. 2. Such a curve, running end-to-end between two links of a robot, can be parameterized in terms of its arc length, 0, . The local coordinate frame of this backbone curve forms a rotation matrix, , whose local axes are depicted in Fig. 2, with tangent to the backbone and perpendicular to the backbone but along the sheet. Again, there is assumed to be no extension or shearing in the plane, only bending of magnitude about a principal axis of rotation at some angle, , as shown in Fig. 3. The change in as a function of is given by the well-known relationship: (1) The matrix is the skew-symmetric form of the rotation rate vector, 0 cos 0cos 0 sin 0 sin 0 (2) The evolution of the backbone position, , is also defined in terms of the local coordinate axes as described by , 100! \" (3) The body frame velocity, \" , corresponds to moving at a constant rate down the backbone curve. Put together, (1) and (3) describe the shape of the backbone curve as a linear, time-varying (LTV) differential equation in #$ 3 , & 0 1 ' & 0 1 ' & \" 0 0 ' (4) Because the backbone curve runs between the edges of the sheet connecting two robot links, solving (4) on 0, can be used to calculate the kinematic relationship, , between the links" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0003015_2017-36-0413-Figure4-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0003015_2017-36-0413-Figure4-1.png", "caption": "Figure 4. Geometry of the final component of the rear bumper.", "texts": [ "66 = (\ud835\udc4411 + \ud835\udc4422 \u2212 2\ud835\udc4412 \u2212 2\ud835\udc4466)\ud835\udc50\ud835\udc5c\ud835\udc60 2\ud835\udf03\ud835\udc60\ud835\udc56\ud835\udc5b2\ud835\udf03 + \ud835\udc4466(\ud835\udc60\ud835\udc56\ud835\udc5b 4\ud835\udf03 + \ud835\udc50\ud835\udc5c\ud835\udc604\ud835\udf03) (7) The shell elements are modeled as two-dimensional structures because their two dimensions are much larger than the thickness, so the coordinates of the thickness are eliminated from the governing equations. For this consideration, the first theory of shell elements was used as first-order shear deformation theory (FSDT). The theory was validated experimentally by the following considerations: (a) the aspect ratio, which is the ratio of the smallest surface to the thickness, must be greater than 10 and (b) the stiffness of the laminas do not differ by more than two orders of greatness. [4] Based on these considerations, the displacement of a generic point B, according to Figure 4, can be represented by the displacement and rotation in the average surface in C, according to the Equations 8-10. \ud835\udc62(\ud835\udc65, \ud835\udc66, \ud835\udc67) = \ud835\udc620(\ud835\udc65, \ud835\udc66) \u2212 \ud835\udc67\u2205\ud835\udc65(\ud835\udc65, \ud835\udc66) (8) \ud835\udc63(\ud835\udc65, \ud835\udc66, \ud835\udc67) = \ud835\udc630(\ud835\udc65, \ud835\udc66) \u2212 \ud835\udc67\u2205\ud835\udc66(\ud835\udc65, \ud835\udc66) (9) \ud835\udc64(\ud835\udc65, \ud835\udc66, \ud835\udc67) = \ud835\udc640(\ud835\udc65, \ud835\udc66) (10) The functions on the right side of the equations are functions at the two-dimensional level, giving an analysis of the shell theory in two dimensions. On the left side, the displacements are three-dimensional functions and correspond to the material variables. Equations 11 and 12 represent the membrane stresses in the x and y directions, and represent the elongation of the shell in these directions", " Unidirectional materials oriented at 0\u00b0 have greater resistance to rupture, this resistance tends to decrease according to the increase of the orientation angle of the fiber, as shown in Figure 3. [9] The study used as a reference the dimensions of the Mercedes Benz Atego 1719 4x2-truck model with total gross weight of 16000kg. [10] The bumper is made of AISI 1020 carbon steel with dimensions of 2400mm in length, 200mm in height and 2mm in thickness. The vertical support bracket also uses the same material, with ratios 60x600x150mm and 2mm thick. Figure 4 shows the final component of the rear bumper assembly of the truck. For the glass fiber reinforced epoxy composite the dimensions of the main beam and flaps were the same, except for the thickness of the laminate, which in this case will be 3mm (1mm each layer). The support used in the simulations is the same consisting of metallic material. Boundary conditions: force and supports The value of the defined force was of 500N and was applied in five points of the frontal area to a distance of 75mm, from the base of the rear bumper" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0002601_gt2017-64123-Figure12-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0002601_gt2017-64123-Figure12-1.png", "caption": "Figure 12. Scheme of gas-dynamic problem formulation", "texts": [ " A more exact observance of the flow radial nonuniformity at the impeller inlet is obtained by using a sub-model that includes an unchangeable intake sector model. Periodicity is applied to the air-gas channel sectors and inlet boundaries. Prismatic layers are created for the blade and the fillet surfaces, with the exception of the trailing edge, hub, and peripheral outline. A fine mesh is created for the inlet and outlet edges, the gap, and the surfaces forming the fillet. A general scheme for the gas-dynamic problem formulation is shown in Figure 12. The air intake model contains 4 million cells. The impeller air-gas channel model may contain from 8 to 10 million cells, depending on the current configuration. A high-quality flow rate computation is obtained for the efficiency, pressure ratio, and gas-dynamic stability factors using a sub-model to perform the gas-dynamic calculations for several points of the characteristics of the selected mode. In this work, the time of call is reduced by carrying out the sub-model calculations for three points of the characteristic [12, 13]" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0000093_amr.452-453.211-Figure1-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0000093_amr.452-453.211-Figure1-1.png", "caption": "Fig. 1 Sketch of machining a globoidal cam Fig. 2 Coordinate systems", "texts": [ " The profile error of a globoidal cam being machined on a NC machine tool with two coordinated rotational axes, resulting from the rotational deviation of location of the part to be machined, is computed. The experiment of machining a globoidal cam is done. Then the experimental result is compared with the computed result, and they are mostly identical. Machining method In this paper, a globoidal cam is machined on a NC machine tool with two rotational axes, which are coordinated each other, and two translational axes. The machining method of the globoidal cam is shown in Fig.1. The part to be machined shown in Fig.1 rotates about the A-axis and oscillates about the B-axis. The distance between the A-axis and the B-axis is referred to as the central distance, and it is controlled and modified by the translational motion along the W-axis. Before machining a globoidal cam, the modification of the central distance must be accomplished. The feed movement of the cutting tool is controlled by the translational motion along the Z-axis. The rotation about the A-axis and the oscillation about the B-axis are coordinated each other", " No part of contents of this paper may be reproduced or transmitted in any form or by any means without the written permission of Trans Tech Publications, www.ttp.net. (ID: 132.239.1.231, University of California, San Diego, La Jolla, USA-13/04/15,03:46:12) For building model, three Cartesian coordinate systems are built. They are shown in Fig.2. And they are all right-handed coordinate systems. The coordinate system O0x0y0z0 is a coordinate system for the fixed frame. The origin point O0 is the crossing point of the central axis of oscillation of the part to be machined, i.e., the B-axis shown in Fig.1, and the central axis of rotation of the cutting tool. The O0z0 axis is aligned with the central axis of rotation of the cutting tool, i.e., the Z-axis shown in Fig.1. The O0y0 axis is aligned with the central axis of oscillation of the part to be machined, i.e., the B-axis shown in Fig.1. The coordinate system O1x1y1z1 oscillates about the Baxis. The O1z1 axis is aligned with the central axis of rotation of the part to be machined, i.e., the Aaxis shown in Fig.1. The O1x1 axis is aligned with the line O1O0, i.e., the W-axis shown in Fig.1. The coordinate system O2x2y2z2 is determined by rotating the coordinate system O1x1y1z1 about the O1z1 axis with the rotational angular displacement equal to the rotational angular displacement of the part to be machined about the A-axis of the machine tool. The coordinate system Oxyz is connected to the part to be machined and it is determined by rotating the coordinate system O2x2y2z2 about the O2z2 axis with the rotational angular displacement equal to the rotational deviation of location of the part to be machined about the A-axis of the machine tool. In Fig. 2, \u03b1 designates the rotational angular displacement about the Oz axis, i.e., the rotational angular displacement about the A-axis shown in Fig. 1; \u03c5 designates the oscillating angular displacement about the O0y0 axis, i.e., the oscillating angular displacement about the B-axis shown in Fig. 1; \u2206\u03b1 designates the rotational deviation of the part to be machined about the A-axis of the machine tool. The point P in Fig. 2 is a point on the central axis of the cutting tool. By generating cutting, the point P generates the pitch profile of the globoidal cam machined. And b designates the distance between the point P and the point O0; d designates the diameter of the cutting tool; a designates the distance between the origin of the coordinate system Oxyz and the origin of the coordinate system O1x1y1z1, i" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0000176_s11668-010-9343-x-Figure2-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0000176_s11668-010-9343-x-Figure2-1.png", "caption": "Fig. 2 Macrograph of the transverse fracture and longitudinal cracking", "texts": [ " The criterion for determining case depth described in the Chinese standard (GB 9450) [1] is the depth of material with hardness greater than HV550. The case depth on the external circle and the internal hole surfaces are specified as 0.8-1.2 mm and more than 0.5 mm, respectively, and the sum of case depth on the external circle and the internal hole surfaces is specified to be less than 2.5 mm. The piston-pin fractured into two pieces transversely and the smaller piece was lost. The remainder of the fractured piston-pin is shown in Fig. 1. The macro-fracture features are shown in Fig. 2. From the fracture surface, longitudinal cracking was found on the piston-pin (labeled crack A and crack B in Fig. 2). Transverse fracture occurred at the contact location of the piston pin with the connecting-rod. The fracture surface was worn and significantly deformed, but the crack propagation marks along the circumference were observed. From the crack propagation marks, it is inferred that there are two crack origins on the transverse fracture, which are situated close to the internal hole of the pistonpin at the intersections of the transverse fracture with the longitudinal cracks A and B. Two cracks propagated X" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0003066_hnicem.2017.8269562-Figure1-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0003066_hnicem.2017.8269562-Figure1-1.png", "caption": "Fig. 1. Basic Mechanics of a Quadrotor - Hover", "texts": [ " This also has a quadratic relationship with respect to the angular velocity. The constants kf and km are the lift and the drag coefficients.[9] Mi = kM\u03c92 i (2) On the other hand, there is a force that keeps the quadrotor to move down which is equal to the mass times the acceleration due to gravity, mg. During hovering, the motor speeds compensate for the weight of the quadrotor. kF\u03c9 2 i = 1 4 mg (3) The weight can determine the basic operating speeds for each rotor and the individual torques. \u03c4i = kM\u03c92 i (4) As shown in Figure 1, the basic mechanics is shown where the four rotors have the same angular velocities \u03c9 during hovering. The resultant force F will be in equilibrium at 0. F = 4\u2211 i=1 Fi \u2212mg (5) Considering one axis, say z-axis, increasing the rotor speeds mean an acceleration going up while decreasing it means the other way. So the combination of the thrust and the weight determines where the quadrotor will go.[9] The section will discuss the about the use of FLC in the study and how it will be related to the FTC" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0003775_978-3-030-26326-3-Figure3.6-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0003775_978-3-030-26326-3-Figure3.6-1.png", "caption": "Fig. 3.6 Left: Spherical wrist. q2 is the wrist angle. Right: Robot scheme showing the bounding cone of all possible forearm axes. This cone is augmented by adding the wrist joint limit to obtain the possible TCP z-axis bounding cone", "texts": [ " In addition, for each of these M IK attempts, we can extract additional information, such as manipulability [36] at the obtained pose, percentage of orientations found for a given 3D Cartesian point, etc. For the feasible orientations of a robot arm in a Cartesian point, several geometrical shapes to represent the valid orientations have been proposed in literature [37]. Among these shapes, cones are probably the best choice, due to their simplicity and easy characterization. In fact, for a robot with a spherical wrist (see Fig. 3.6a), the Tool Center Point (TCP) stays within a cone whose axis is the rotation axis of the first degree of freedom of the wrist (namely, the forearm axis). 3.2 Bimanual Arm Positioning 45 Moreover, discarding the rotation around the TCP z-axis, we propose to collect the set of valid forearm axes at a certain Cartesian point P \u2208 R 3, which will be enclosed by a cone, and compute the Bounding Cone (BC) that contains them all with the algorithm proposed in [3]. Also, if the wrist angle has symmetric limits, its aperture can be added to the BC angle, yielding a cone that contains all the TCP z-orientation axis that the robot can reach at the given position (see Fig. 3.6b). With this approach, we obtain a mesh for the workspace, encoding all the information gathered when computing the reachable positions such as manipulability, percentage of solutions found, etc. plus the obtained bounding cone containing all the possible z-axis of the TCP. We can see an IK solutions map over the workspace of a WAM robot (see Table1 in [30] for its dimensional parameters) in Fig. 3.7. Multiple-arm cooperative tasks provide the capability of performing tasks that would be impossible or, at least, much more difficult to accomplish with only one arm" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0001159_jjap.51.030206-Figure2-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0001159_jjap.51.030206-Figure2-1.png", "caption": "Fig. 2. (Color online) Schematic illustration of the fabrication of the microslit. The SU-8 photoresist which is spincoated on the silicon substrate is exposed to UV radiation through a photomask. If the width of the pattern on the mask is W, then the width of the cured photoresist after development is W \u00fe W given the UV diffraction. Therefore, if two semicircular patterns are on the mask and the distance between the patterns is smaller than W, the patterns are connected to each other. After molding the photoresist with PDMS, a narrow microslit is generated. With increased UV exposure, W increases and a wide microslit is obtained.", "texts": [ " After staining, cells were cultured for 2 days, detached from the culture flask by the addition of trypsin, washed, and then dispersed in HVJ-E fusion kit buffer (Ishihara Sangyo). A poly(dimethylsiloxane) (PDMS)-glass microchip was fabricated by standard microfabrication techniques with some modifications.12) The PDMS part possessed a recessed microchannel pattern (13 m depth) and a weir containing a microslit. The gap between the weir and the glass substrate was 2 m. The width of the microslit (W) was controlled by the size of the corresponding pattern on the template, and by the exposure time of the photoresist layer to the UV beam using a photomask (Fig. 2). The PDMS part was bonded to a slide glass without any surface treatment. Pairing processes were visualized using a fluorescence microscope (Olympus IX71) equipped with an EB-CCD camera (Hamamatsu C7190). The temperature and humidity of the microscope stage was controlled using a temperature controller (Tokai Hit MATS555RO) and cell culture system (Tokai Hit INU-ONICS-F1). The B6G-2 cells and the 3T3 cells (103 cells L 1 for each) were introduced from different inlets into the microchannel by applying a negative pressure ( 5 kPa) to the outlet of the channel using a vacuum pump (DAL-5D, Ulvac) for 3min at room temperature" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0002185_i2017-11352-9-Figure6-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0002185_i2017-11352-9-Figure6-1.png", "caption": "Fig. 6. 3D configuration diagrams of the corresponding thin elastic rods under the special points. (a) x = HH1; (b) x = OO1; (c) x = HH2; (d) x = OO2; (e) x = HH3. \u03b1 indicates the value of the helix angle.", "texts": [ " n > 0 represents the n-th phase diagram to the right of the reference phase diagram, and n < 0 represents the |n|-th phase diagram to the left of the reference phase diagram. n = 0 represents the reference phase diagram. As shown in fig. 5, the saddle points, the center point, and the critical orbit can be simplified as Hi, Oi, Li. To better understand the relationship between the special points (in table 1 and fig. 5) and the configurations of rods, the spatial configuration of the corresponding rods is drawn using these special points as shown in fig. 6. As shown in fig. 4, in the phase diagram of each cycle, the heteroclinic loops are asymmetric heteroclinic loops, consist of the left (represented by Ln,1 and Ln,2) and right (represented by Ln,3 and Ln,4). All saddle points of the two heteroclinic loops are asymmetrically referenced to the center point. The phase diagram formed by these four heteroclinic orbits has two extreme points Hn,1 and Hn,3 on the \u03b8(t)-axis. The \u03b8(t)-coordinates of the two extreme points meet the following condition: Hn,1 = \u22122\u03c0 + Hn,3, (6) i", " 9(a), and (d) exhibits a similar pattern as fig. 9(b). We found that after passing the known point A, the elastic rod had a spiral growing configuration similar to the one described above. Point A satisfies the boundary condition (12), and the spiral angle corresponding to point A on the rod is zero degrees. In addition, the corresponding end of the rod presents a parallel shape to the central line of constrained cylinder, which is similar to a straight rod and is expressed as \u201cS\u201d. It matches the shape drawn in fig. 6(c). The direction of the elongating elastic rod along the constraint cylinder is correlated with the value \u03b8, and the third equation in eq. (13) was analyzed from a mathematical perspective. If cos \u03b8 > 0, the coordinate value of Z increases; if cos \u03b8 < 0, the coordinate value of Z decreases. As shown in fig. 9(a) and (c), the corresponding values of cos \u03b8 on the heteroclinic orbit are initially greater than zero when the rod elongates; after passing the known point A, the values of cos \u03b8 are smaller than zero", " Whether the elastic rod is elongating as a left helix or a right helix along the constrained cylindrical is also related to the value of \u03b8. Z/Y in eq. (13) was analyzed from the mathematical perspective; if cos \u03b8 > 0, the elastic rod extends as a right helix, if cos \u03b8 < 0, the rod extends as a left helix. As shown in fig. 9(a) and (b), when the heteroclinic orbit falls into the saddle point H1, corresponding to cos \u03b8 < 0, the rod is shown as a left helix, which is in accordance with the configuration in fig. 6(a). As shown in fig. 9(c) and (d), when the heteroclinic orbit falls into the saddle point H3, corresponding to cos \u03b8 < 0, the rod is shown as a left helix, which is in accordance with the configuration in fig. 6(e). Therefore, it is reasonable to assume that at the minimum value of the dip angle \u03b8(t) \u2212\u03c0 < H1 < \u2212\u03c0/2, the corresponding values of cos \u03b8 on the heteroclinic orbit should always be smaller than zero as the rod elongates, while at saddle point H1, cos \u03b8 < 0, the elastic rod extends downward along the constraint cylinder as a left helix, similar to the configuration in fig. 6(b). If the dip angle \u03b8(t) is at the maximum value 0 < H3 < \u03c0/2, the corresponding values of cos \u03b8 on the heteroclinic orbit will always be greater than zero as the rod elongates, while at saddle point H3, cos \u03b8 > 0, such that the elastic rod extends upward along the constraint cylinder as a right helix, similar to the configuration in fig. 6(d). This result may also explain the sudden left or right helix shift during the tendril climbing process, and it is similar to the explanation of the tendril perversion problem using heteroclinic orbit theory [10]. This conjecture will be validated in subsequent quantitative studies. Comparing the graphical representation of the theoretical analysis in fig. 9 and the actual shape of cucumber tendrils in fig. 2 and fig. 10, fig. 10 corresponds to the form shown in fig. 9, where the cucumber tendril was climbing up but owing to a change in the external force, heteroclinic bifurcation occurred" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0003191_1.4042509-Figure18-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0003191_1.4042509-Figure18-1.png", "caption": "Fig. 18 Variables of air pumping loss Pumping area", "texts": [ " Received October 22, 2018; Accepted manuscript posted January 16, 2019. doi:10.1115/1.4042509 Copyright (c) 2019 by ASME Downloaded From: https://gasturbinespower.asmedigitalcollection.asme.org on 01/26/2019 Terms of Use: http://www.asme.org/about-asme/terms-of-use Journal of Engineering for Gas Turbines and Power GTP-18-1661, Arisawa, 22 , , ,min( ) 2 0.5 / 2m in out g in out h in h out in inQ B B M j S j S C C z , (28) where j is the gear backlash, S is the top land length of the gear, hC is the top clearance of the gears (refer to Fig. 18), and z is the number of gear teeth. In Eq. (22), the loss coefficient of the orifice ( ) is calculated using an empirical equation that depends on the area ratio 0 /A A [16]: 2 0.375 2 0 0 01 / 0.707 1 / /A A A A A A . (29) The orifice area 0A , which is the area of the teeth valley, is approximated as 0 1 2 2 2 in out g h h A M , (30) where gM is the gear module, and the final coefficient \u201c2\u201d accounts for both sides. Thus, sV in Eq. (23), p in Eq. (22), F in Eqs. (20) and (21), and sP in Eqs", " 15 Schematics showing the side flow loss, air pumping loss, and vortex loss, Acc ep te d Man us cr ip t N ot C op ye di te d Journal of Engineering for Gas Turbines and Power. Received October 22, 2018; Accepted manuscript posted January 16, 2019. doi:10.1115/1.4042509 Copyright (c) 2019 by ASME Downloaded From: https://gasturbinespower.asmedigitalcollection.asme.org on 01/26/2019 Terms of Use: http://www.asme.org/about-asme/terms-of-use Journal of Engineering for Gas Turbines and Power GTP-18-1661, Arisawa, 37 which comprise of the windage loss Fig. 16 Process of side flow loss Fig. 17 Variables of side flow loss Fig. 18 Variables of air pumping loss Fig. 19 Comparison of calculated and experimental losses Fig. 20 Comparison of calculated and the experimental oil dynamic losses at different oil jet supply rates Fig. 21 An example of accuracy enhancement by using the improved oil dynamic loss model (Shroud 1, oil jet supply = 7.4 liters/min, = 1.9) Fig. 22 Comparison of calculated and experimental windage losses at different input speeds Acc ep te d Man us cr ip t N ot C op ye di te d Journal of Engineering for Gas Turbines and Power" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0003190_978-3-030-05321-5_8-Figure8.8-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0003190_978-3-030-05321-5_8-Figure8.8-1.png", "caption": "Fig. 8.8 Typical structure of hydraulic oscillating motor", "texts": [], "surrounding_texts": [ "Compared to conventional hydraulic machines, robots have multiple degrees of freedom in motion and fine control for highly advanced tasks. From this perspective, low friction and high power are the main challenges for hydraulic components in robotic usage, as these contribute to the multiple degrees of freedom with a rapid response and fine control in precision work.\nThis section introduces our work on hydraulic components for tough robots. We focus on low friction and high power, especially on 35-MPa low-friction and highpower actuators, a power pack for autonomous driving of hydraulic robots, and peripheral equipment for hydraulically driven systems, including hoses and couplings.\nIn general, conventional hydraulic actuators can be classified into hydraulic cylinders and oscillating motors, as shown in Figs. 8.7 and 8.8, respectively. Hydraulic actuators have several advantages: high output density (high force or torque ratio to their body mass), simple driving in bidirectional motion using control valves, and high back drivability because of the lack of reduction gears. These features bring higher shock resistance and environmental robustness than drive systems composed of electric motors and reduction gears. However, in spite of the advantages of hydraulic actuators, most conventional hydraulic actuators are not suitable for robotic usage because of their large weight and bulky size. In addition, those conventional", "actuators need operating pressures of 0.3\u20130.8MPa even under no load because of the high sliding resistance owing to the packing that prevents leakage of the working fluid.\nFor robotic usage, we have developed lightweight hydraulic actuators using lightweight alloys whose regular pressure is 35MPa. Sliding resistance was reduced by appropriate selection of structures and materials for the sealings, minimizing surface roughness of the sliding parts, and strict control of dimensional tolerance and machining temperature. The developed actuators achieved a no-load drive with an operating pressure of 0.15MPa with almost no leakage of the working fluid.\nAs shown in Fig. 8.9, the developed cylinder successfully achieved twice the output density of that of cylinders regulated by Japanese Industrial Standards (JIS) and the International Organization for Standardization (ISO). Note that the output density is the ratio of the maximum thrust force to the mass of the cylinder here.\nBased on the discussion above, we developed cylinders for hydraulic robots operated by 35MPa whose diameter is 20\u201360mm. Figure8.10a, b are examples of the developed cylinders. These cylinders are applied to a tough robotic hand (the details of which will be mentioned in Sect. 8.6).", "Figure8.11a shows the result of measuring the sliding pressure of the developed cylinder (Fig. 8.10a). In the measurement, the cylinder drove with no load for its full stroke of 100mm.\nFigure8.11a, b show the time response of the operating pressures and the position of the pistons on the pushing side and the pulling side, respectively.\nAlthough the pressure increases near the stroke end point, the operating pressure achieved is <0.01MPa.\nOscillating torque actuators that can be operated with 35MPa were also developed, and the output density reached higher values, as shown in Fig. 8.12. The developed motors were designed to drive with no load operated by<0.2MPa. The vane packing was molded out of a special resin (Fig. 8.13). The developed oscillating actuator, depicted in Fig. 8.14, has one vane, with a rotating angle of 270\u25e6, and an output torque of 160\u2013670Nm. A 360\u25e6 oscillating motor was also developed for the high requirements in robotic usage (Fig. 8.15). This oscillating motor can generate 600 N m of torque with an applied pressure of 35MPa. Using this type of hydraulic motor can enable a design with a smaller outer diameter, which is advantageous for a compact hydraulic robot design." ] }, { "image_filename": "designv11_62_0000248_s40194-013-0079-6-Figure2-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0000248_s40194-013-0079-6-Figure2-1.png", "caption": "Fig. 2 Experimental coaxial setup of the plasma torch with a hollow cathode", "texts": [ " In order to differentiate the influences inside the arc column and on the workpiece surface, we modified our coaxial plasma-laser torch to switch the laser spot position slightly related to the electrode axis. Any interaction inside the arc column such as increased electric conductivity should not be influenced, but the temperature distribution is influenced by the relative movement of the laser spot and the arc. The experiments were carried out using a modified plasma torch ABIPLAS WELD 150 with coaxial laser transmission of laser radiation, Fig. 2. This setup reaches best stability for the plasma arc and best absorption at the surface for laser radiation. In order to enable laser deflection in direction of feed, the previous coaxial setup was developed further with a wider drilled hole in the tungsten electrode. Therefore, leading or trailing positions of the laser spot on the surface of the workpiece could be adjusted. Plasma nozzles of \u00d8 1.2 and \u00d8 2.6 mm were used. The working distance between plasma nozzle and workpiece was constant at 6 mm" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0002274_s12206-017-0308-9-Figure12-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0002274_s12206-017-0308-9-Figure12-1.png", "caption": "Fig. 12. (a) Sketch illustrates the cross-section of the final part after finishing processes; (b) sketch shows the isometric view of the part (all dimensions in mm).", "texts": [ " Because of the fact that it is depending on the open source system, which gives the operator capability to tune the process parameters and control the Deposition tool (DT) movements during the deposition process and give the system more flexibility to produce different geometries. So it is possible to claim that the current SMD system is appropriate for workshop to fabricate small and medium size parts. Finally a typical special aerospace part used in aerospace applications was selected as a case study to evaluate the cost and time savings using TIG + Wire SMD and to provide an idea about the benefits of this technique. Fig. 12 represents the sketches showing the dimensions of the part to be manufactured. Total time consumed during deposition of each layer of the part was 53 seconds. So the total deposition time was measured as 42.4 minutes whereas the machining time consuming during finishing the part is 32 minutes. And so the total time for fabricating the finished part is about 1 hour and 14.4 minutes. Whereas the time consumed during machining process of the same finished part from solid block is 56 minutes without calculating time of cutting process for the piece from a desired shaft or billet" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0003485_1350650119851658-Figure2-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0003485_1350650119851658-Figure2-1.png", "caption": "Figure 2. Schematic diagram of lubrication conditions of the whole ring pack.", "texts": [ " hsuppl represents the oil film thickness on the cylinder liner at the inlet position under the starved lubrication condition. For the way to solve the starved lubrication, the interested readers can refer to the works of Gu et al.25,26 In addition, it should be mentioned that equation (3) could be simplified under some certain circumstances. As reported by Guo et al.,24 when the piston is near the dead center, the inlet boundary condition can be simplified as: h xin, y\u00f0 \u00de \u00bc hsuppl. 3. For piston ring pack/cylinder liner system Figure 2 shows the schematic diagram of lubrication conditions of the whole ring pack. When the whole ring pack is considered, the oil flow pattern is indispensable to be illustrated. As shown in Figure 2, when the ring pack moves along the cylinder liner, the oil on the liner fills into the clearance between the ring and the liner and generates the hydrodynamic pressure. If no external oil is provided, the portion of the oil left behind by the preceding ring would maintain its volume and be delivered to the next ring of the ring pack.27 If the amount of oil is more than sufficient to fill the clearance between the ring and the liner, the excess oil will accumulate in front of the ring. In contrast, if the amount of the oil is insufficient to fill the clearance between the ring and the liner, the ring is said to be operating under the starved lubrication.27 In the ring pack, the oil control ring (or say the third ring) is used to prevent the inflow of excess oil. Therefore, under engine condition, most ring packs are operating under the starvation condition.27 As shown in Figure 2(a), when the piston moves down, the bottom rail of the oil control ring (or say the third ring) would see the amount of oil. However, for all other rings, the oil supply over the engine cycle is based on the oil film left behind by the preceding ring, according to the work of Gulwadi.27 The corresponding relationship about the flow rate of each ring can be expressed as follows qsuppl 5 q3inlet \u00bc q3outlet 5 q2inlet \u00bc q2outlet 5 q1inlet \u00bc q1outlet \u00f04\u00de where the superscript of the flow rate (q) represents the kth ring (k \u00bc 1, 2, 3). The subscript inlet denotes the flow rate in leading edge of the particular ring, while the subscript outlet represents the flow rate in the trailing edge. The subscript suppl represents the flow rate for oil supply. When the engine is during the down-strokes (the piston moves down), the expression of qsuppl \u00bc Uhsuppl can be used. The flooded or starved condition for each ring is dependent on oil film history based on liner oil transport. As shown in Figure 2(b), during up-strokes, the scraping effect of the top ring (or say the first ring) is responsible for oil accumulation at the leading edge. As the piston moves up, the film available for lubricating the top ring (or say the first ring) is based on the oil left behind by it during the previous downstroke.27 The corresponding relationship about the flow rate of each ring is expressed as qsuppl 5 q1inlet \u00bc q1outlet 5 q2inlet \u00bc q2outlet 5 q3inlet \u00bc q3outlet \u00f05\u00de For the up-stroke, the flow rate of oil supply qsuppl should be based on the oil left behind by it during the down-stroke, owing to the engine performance" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0000022_j.jappmathmech.2012.05.003-Figure2-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0000022_j.jappmathmech.2012.05.003-Figure2-1.png", "caption": "Fig. 2.", "texts": [ "7) ith a finite region of attraction, which contracts to a point as \u2192 \u20130. To prove this, it is sufficient to find the minimum point of function (4.5) and then use Tikhonov\u2019s theorem. Note that the absence of tationary values of the reaction in cases when function (4.5) is monotonic is caused by the unboundedness of all the phase trajectories f system (4.4), i.e., u \u2192 \u221e as \u2192 \u221e. An unbounded increase in the reaction indicates a shock, which leads to abrupt changes in the arameters , and and function (4.5). Figure 2 shows graphs of function (4.5) for the cases listed in Lemma 4. The system displays regular behaviour, i.e., the constraint eaction is specified uniquely, only in the first of these cases. In cases b and d the positive energy is unbounded below, indicating a shock or any initial values of the reaction. Such a shock is also possible in cases c and e along with the stable value of the reaction in the potential ell determined from formula (4.6) or (4.7), respectively. xample. 2,10Two heavy point masses M1 and M2 with unit masses are joined by a rigid weightless rod of length l", " When the equations f motion are written, the system is released from the constraint y = 0, which is replaced by the sum of the normal reaction N (which is easured in the downward direction) and the friction force F. We will assume that the point M1 slides to the right. Then F = \u2212 |N|. The inetic energy of the system and the gravitational potential energy have the form e write the equations of motion in Lagrange form (4.8) Expressing the acceleration y\u0308 from system (4.8), we obtain formula (4.3), where (4.9) It is seen that > 0; therefore, depending on the values of and , cases a, c and d (Fig. 2) are possible in this system. The regular case corresponds to the inequality (4.10) A similar analysis can be performed for the case of unilateral constraint (2.12) with kinetic friction (4.1). In this case, N \u2265 0 for q1 = 0, nd N = 0 for q1 > 0; therefore, when u > 0, the terms that are first-order in u on the right-hand side of system (4.4) vanish (accordingly, in xpression (4.5) the second-order terms vanish). These terms correspond to the constraint reactions (4.11) Figure 4 presents graphs of function (4" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0001227_tim.2010.2046600-Figure2-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0001227_tim.2010.2046600-Figure2-1.png", "caption": "Fig. 2. Structure of the transducer and the planar coil.", "texts": [ " (7) Use of (4) and (6) in (7) gives e = N\u2211 j=1 ej = \u2212 N\u2211 j=1 d\u03d5j dt = \u2212 N\u2211 j=1 \u03bc0u cos \u03b1 16\u03c0ab di dt \u00d7 \u222b\u222b D ln ((xfj \u2212 x\u2032)2 + (yfj \u2212 y\u2032)2) ((xnj \u2212 x\u2032)2 + (ynj \u2212 y\u2032)2) dx\u2032dy\u2032 (8) where N is the total turn numbers of the coil, i is the current flowing through the conductor, and \u03bc0 is the permeability in vacuum. Based on the coil parameters and the relative position between the coil and the busbar, xnj , xfj , ynj , and yfj are known. The value of e can be calculated using (8). When the three-phase currents are measured with the planar coils, the result can be affected by the nearby phase currents. For example, in Fig. 2, ib and ic may contribute to the voltage output of the coil a. In Fig. 2, the electromotive force of coil a consists of eaa (contributed by ia), eba (contributed by ib), and eca (contributed by ic); they are respectively generated by the three-phase current. Using (8) gives eaa = N\u2211 j=1 eaaj = N\u2211 j=1 ( \u2212\u03bc0u cos \u03b1 16\u03c0ab dia dt \u00d7 \u222b\u222b Daa ln (xfj\u2212x\u2032)2+(yfj\u2212y\u2032)2 (xnj\u2212x\u2032)2+(ynj\u2212y\u2032)2 dx\u2032dy\u2032 ) (9) where Daa ={(x\u2032, y\u2032)|\u2212(3a+d)\u2264x\u2032 \u2264\u2212(a+d),\u2212b\u2264y\u2032 \u2264b}, and where d is the spacing between two conductors, as shown in Fig. 2. We also have eba = N\u2211 j=1 ebaj = N\u2211 j=1 ( \u2212\u03bc0u cos \u03b1 16\u03c0ab dib dt \u00d7 \u222b\u222b Dba ln (xfj\u2212x\u2032)2+(yfj\u2212y\u2032)2 (xnj\u2212x\u2032)2+(ynj \u2212y\u2032)2 dx\u2032dy\u2032 ) (10) where Dba = {(x\u2032, y\u2032)| \u2212 a \u2264 x\u2032 \u2264 a,\u2212b \u2264 y\u2032 \u2264 b}, and eca = N\u2211 j=1 ecaj = N\u2211 j=1 ( \u2212\u03bc0u cos \u03b1 16\u03c0ab dic dt \u00d7 \u222b\u222b Dca ln (xfj\u2212x\u2032)2+(yfj\u2212y\u2032)2 (xnj\u2212x\u2032)2+(ynj\u2212y\u2032)2 dx\u2032dy\u2032 ) (11) where Dca = {(x\u2032, y\u2032)|a + d \u2264 x\u2032 \u2264 3a + d,\u2212b \u2264 y\u2032 \u2264 b}. Therefore, the total electromotive force of the planar coil a is ea = eaa + eba + eca. (12) In (12), eba and eca can be taken as the influence with regard to eaa. Taking this into account, the uncertainty of the planar coil a in the measurement values of current ia in Fig. 2 is \u03b5% = Ea\u2212aa Ea \u00d7 100 (13) where Ea\u2212aa and Ea are the RMS values of (ea \u2212 eaa) and ea. The output of coil b with eab, ebb, and ecb, and coil c with eac, ebc, and ecc, can be obtained by using the same method in deducing the expressions given by (9)\u2013(11). B. Insulation Structure of the Transducer The software package Ansoft SB is used to simulate the potential and electric field intensity distribution inside the transducer. For the convenience of simulation, it is suggested that the transducer is placed above a grounded metal plane, and an HV of 42 kV is applied to the output terminals [the terminals are shown in Fig", " The transducer\u2019s output depends on the busbar current and its relative position to the busbar. The relative errors due to misalignment were reported in [6]. When the horizontal misalignment between the transducer and busbar is less than 5 mm, the relative error is less than about 1.5% [6]. The influence of nearby busbars to the transducers in field operation should be considered, and it will produce an additional error. In field applications, the busbar has a cross section with sides of length 80 mm and 8 mm, and the spacing between two adjacent busbars d shown in Fig. 2 is equal to 130 mm. Supposed that the three-phase current is\u23a7\u23a8 \u23a9 ia = I0 cos(\u03c9t \u2212 120\u25e6) ib = I0 cos \u03c9t ic = I0 cos(\u03c9t + 120\u25e6) (14) where \u03c9 = 2\u03c0f , and f = 50 Hz. In view of (9)\u2013(11), and (14), we can get the output of each of the three-phase coils due to the contribution of the other two adjacent busbars. When the amplitude I0 of the current is 500 A, the additional error of coil a due to the influence of ib and ic, and that for coil b due to ia and ic, and that for coil c due to ia and ib, are listed in Table I" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0002302_b978-0-12-803581-8.10351-0-Figure31-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0002302_b978-0-12-803581-8.10351-0-Figure31-1.png", "caption": "Fig. 31 Displacement due to suspension loading.", "texts": [ " The intermediate bulkhead stress and displacement for the suspension load was determined by conducting a numerical analysis. One side of the bulkhead was constrained as it would be in the chassis. The other side was loaded with a 650 lb force applied through the bulkhead. The bulkhead in actuality would be constrained along the outer surface of the bonding flange, by the means of the chassis. Thus, the results of this form of loading is conservative and even so, the bulkhead had failure index of 0.25 (safety factor of 4 based on ultimate strength), shown in Fig. 31. The rack indent and the bolt holes were loaded with 1000 lb force through the bulkhead. The results of the rack load concluded that the bulkhead had failure index of 0.33 (safety factor of 3.03 based on ultimate strength), as shown in Fig. 32. The main design of the roll hoop is to prevent the driver from becoming crushed by the chassis in the event of a rollover crash. To increase the efficiency of this structure required by rules, it was also used to react loads from the suspension. Under the rules of the competition that this monocoque was originally designed for, the main forward roll hoop was constrained to be a metal tube structure" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0001672_978-1-4614-0222-0_34-Figure4-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0001672_978-1-4614-0222-0_34-Figure4-1.png", "caption": "Fig. 4 The first four in-plane mode shapes", "texts": [], "surrounding_texts": [ "The modal testing of the SRT model was recently studied [24]. Overall, a satisfactory agreement was found between the experimentally measured frequencies and those predicted in the finite element approach, Tables 9 and 10. The results indicate that in the out of plane direction, the FE models are slightly stiffer than the experimental model. While in the in-plane direction, except for the first mode, the experimental model is stiffer than the FE model with the lumped mass and is softer than the model with flap. A few sources of discrepancies could be indentified in this approach explaining the aforementioned differences between experimental and predicted frequencies. In the modal testing, since maintaining the internal pressure was done manually, it is less likely to achieve the same pressure at all times. Variable pressure contributes to variable stiffness which alters the modal parameter of the structure. Owing to presence of very small punctures in the SRT, the air leakage made the distribution of pressure non-uniform in the torus. In the finite element model, the pressure was assumed to be constant in the entire torus. Owing to the extremely small values of the film thickness, any increase in the thickness significantly affects the stiffness and mass of the structure and eventually the eigenvalues. For example, the average seams\u2019 thickness was above six times the shell film thickness. Furthermore, the method of fabrication, as well as the extensive repair that was done on the flaps, resulted in variations of the thickness of the joined regions. This contributed to non-uniform stiffness and mass distribution in both the torus and the flaps, affecting the modal parameters of the SRT. The flaps were not completely flat and had wrinkles at various locations. In the finite element study, the flaps were assumed to be flat with a uniform thickness equal to the average thickness of the inner and outer flaps." ] }, { "image_filename": "designv11_62_0003190_978-3-030-05321-5_8-Figure8.44-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0003190_978-3-030-05321-5_8-Figure8.44-1.png", "caption": "Fig. 8.44 Hand mode and bucket mode of the hydraulic tough hand", "texts": [ " Between the upper and lower overhanging plates, a 4-port swivel joint is located, of which, only two ports are in this application. The two fingers on the inner side are fixed to the palm box at the finger roots. In this paper, the outer-side fingers and inner-side fingers are termed as \u201crotating fingers\u201d and \u201cfixed fingers,\u201d respectively. As mentioned above, the hand has two operation modes, and the change in the mode, or in other words, the \u201ctransformation\u201d, is carried out by rotating the two rotating fingers around their vertical spinning axes in the opposite direction by approximately 160\u25e6, as shown in Fig. 8.44. In the hand mode, the rotating fingers and the fixed fingers face each other, but in the bucket mode they line up laterally. In the bucket mode, in order to form a concave shape, all the root joints are wide open, and the two tip joints of the fixed fingers are closed. In this case, the opening angle of the fixed fingers is designed to be larger than that of the rotating fingers. Furthermore, the rotating fingers are designed to be shorter than the fixed fingers so that all the four fingertips can be arranged in line to ensure satisfactory bucket performance" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0002621_aim.2017.8014122-Figure5-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0002621_aim.2017.8014122-Figure5-1.png", "caption": "Figure 5. Foot position put in front of stairs too far", "texts": [ " Therefore, a safety distance between former foot and stair before the user starting to climb up is also required as a switch to create an ideal trajectory from even ground to stairs. For simplification of calculations, the range of ZMP we set is from the toe part of the apparatus to the projection point of center of knee joint of the user. The objective of this work is to add a new control programing to ensure the apparatus can change its current walking trajectory to another walking trajectory by using ultrasonic sensor for climbing up when it starts to approach the stair and finish the climbing motion, as shown in Fig. 5, which the red dotted is the original level walking trajectory of the swing leg and the modified level walking trajectory is presented by the green one, and make sure the swing leg can be placed on stair area after changing its walking trajectory, and, finally, to achieve the purpose that the apparatus can climb up the stairs safely along with stairs walking trajectory and change back to the level walking trajectory after finishing climbing. For the sack of achieving the function to keep the safety distance by the walking assistance apparatus, distance sensors are needed to be mounted on the toe part of the apparatus to measure the distance between it and the stair like Fig" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0000786_aero.2013.6496960-Figure9-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0000786_aero.2013.6496960-Figure9-1.png", "caption": "Figure 9 \u2013 Attachment of sensors", "texts": [ " the gaps between the propulsion subunits and those between expansion plates, are covered with dustproof material (aluminum evaporation sheets) to prevent soil from entering a subunit. Additional rubber friction sheets are placed on the outer surface of the expansion plates, increasing the friction forces while the robot is moving inside a launcher (see Fig. 13). Table 1 shows the specification for each subunit. The sensor (FlexiForce, Nitta) is attached between the link and the expansion plate (Figure 9). Each link has a sensor, and an expansion plate has two sensors. Each subunit has a total of eight sensors, each of which measure the vertical push force against the wall. The output voltage from a sensor is amplified with an amplified circuit, and it is measured by a computer through an A/D converter. Pushing Force Model of Expansion Plate inside Cylindrical Shape The pushing force of the expansion plates is modeled here. Figure 10 shows the inner structure of the dual pantograph. A propulsion subunit expands and contracts in a radial direction along with the contraction and extension of plates in the axial direction" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0001223_03091902.2013.785608-Figure19-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0001223_03091902.2013.785608-Figure19-1.png", "caption": "Figure 19. COMSOL model depicting the pressure differences across the spirometer design at Time\u00bc 2 s illustrating the pressure change during measurement.", "texts": [ "69 2\u20135/160 1\u201315/160 Proto Board 100 69 0.69 10 10 3/400 Plywood 100 63 0.63 200 200 Pine 30.48 m 100 1 Large-Diameter Flat Washers 200 16 0.08 Wood Screws #12 200 400 52.8 0.132 Steel Machine Screw Hex Nut 400 24 0.06 Load Cells 2 compressive load cell (16375097) strain gauges from technoweighindia.com 200 780 3.9 Plastic Circuit Cover 100 88 0.88 Polyurethane 1150 oz 30 0.3 Total 9.432 Figure 16. Spirometer without mouthpiece. and an outlet of 0 kPa. Figure 18 shows the pressure at Time\u00bc 0 s, while Figure 19 shows the pressure at Time\u00bc 2 s. From the readings, it can be seen that a pressure drop occurs along the narrow cylinder, after the eccentric cone that can be easily detected by the sensor. The user interface for the spirometer was designed to properly engage both the operator and the patient using the device. This was done through a graph and fill bar as shown in Figure 19. The graph of flow rate helped the operator in evaluating spirometer read-out. If the operator saw that the flow rate was non-continuous or that there was severe signal noise, he or she could better determine the cause for error so that they could re-explain or demonstrate how to properly hold the device. Additionally, the software displayed the forced vital capacity (FVC), FEV1/FVC ratio and maximum flow rate. These three numbers could be seen on the user interface, just to the right of the graph of flow rate vs time" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0000703_j.talanta.2010.06.027-Figure1-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0000703_j.talanta.2010.06.027-Figure1-1.png", "caption": "Fig. 1. Schematic of the experimental FI system.", "texts": [ " The present paper reports on the development of a general mathematical model of the Nafion optode incorporating PAN in the case of Cu2+ determination and the experimental verification of this model under FI conditions. The experimental determination of the model parameters (e.g. diffusion coefficients, equilibrium constants, molar absorptivities) whose numerical values were not available in the literature is also outlined. S.D. Kolev et al. / Talanta 82 (2010) 1156\u20131163 1157 2 2 a w m w t c t e i t t [ c w b E w w r n a D w c 2 i m w T m d I . Theory .1. Determination of the diffusion coefficient of CuSO4 in queous solutions The dispersion of CuSO4 in the experimental FI system (Fig. 1) as described mathematically by the axially dispersed plug flow odel (Eq. (1)) [17]. \u2202c \u2202t = DL \u22022c \u2202z2 \u2212 u \u2202c \u2202z (1) here c is the concentration of CuSO4, z is the axial distance, u is he average linear flow velocity and DL is the axial-dispersion coeffiient which in the general case is an empirical parameter analogous o the molecular diffusion coefficient in Fick\u2019s laws. All flow-through sections of the experimental FI system had qual diameters and the volumes of the measuring cell and the njection loop were much smaller than the reactor volume", " The total bsorbance of the membrane is equal to the sum of the absorbances f the chemical species mentioned above: = \u0131 { \u03b5LHcm LH + \u03b5LH+ 2 cm LH+ 2 + \u03b5CuL+ cm CuL+ } (10) here \u03b5i and cm i are the molar absorptivity and the average oncentration of the i-th absorbing chemical species (Eq. (10a)), espectively. m i = 1 \u0131 \u222b \u0131 cm i dx (10a) 0 The initial distribution of LH within the optode membrane prior o Cu2+ measurements was calculated by the mathematical model 23] developed earlier by us for the extraction of LH into Nafion embranes. Fig. 3. Schematic of the optode flow-through measuring cell. The mathematical model of the FI system (Fig. 1) incorporating the optode flow-through measuring cell (Fig. 3) consists of the mathematical descriptions of the following physicochemical subsystems: (1) mass-transfer of the Cu2+ ions in the aqueous solution towards the optode membrane; (2) mass-transfer of the Cu2+ ions and other mobile chemical species within the optode membrane; (3) ion-exchange at the membrane/solution interface; (4) protonation and complexation reactions within the optode membrane. The mathematical descriptions of these four interrelated subsystems are outlined in the following paragraphs", " Apparatus The spectrophotometric measurements were performed with Shimadzu UV-240 UV\u2013visible spectrophotometer (Japan) with n OPI-2 interface and 10 mm cells. The thickness of the Nafion\u00ae embranes was measured by a calibrated microscope (Nikon abophot 2, Type 104, Japan) with \u00b1 0.005 mm accuracy. Flame tomic absorption spectrometric (AAS) measurements were caried out by a GBC Model 933 spectrometer (Australia). .3. FI system The determination of the CuSO4 diffusion coefficient was carried ut in a single-channel FI system (Fig. 1) incorporating a peritaltic pump (Alitea C4, Sweden), a rotary injection valve (Model 020, Rheodyne), Teflon tubing (i.d. 0.5 mm, Sigma\u2013Aldrich), and home-made flow-through measuring cell [8]. In the FI optode easurements, an electrically actuated and computer controlled ultiposition selection valve (Model E10, Valco Instruments, USA) llowing switching between a carrier and several reagent soluions was inserted upstream of the peristaltic pump (Fig. 1). The ength of the straight Teflon tube (referred to as the reactor) conecting the injection valve with the measuring cell (Fig. 1) was .01 m. The average internal diameter of the reactor (0.543 mm) nd the sample volume (14.75 L) were determined from the volme of deionized water which filled the corresponding tubing. The eservoir of the carrier solution was immersed in a water bath therostated at 25 \u25e6C (Thermoregulator TH5, RATEK, Australia). Water rom the water bath was also circulated constantly through the hermostating jacket of the reactor of the FI system (Fig. 1). The volumetric flow rate was varied in the range from 0.200 to .600 mL min\u22121 and was measured before and after each experient by collecting and weighing the effluent over 5 min. The data acquisition and system control were performed by PC with a PCL-818H data acquisition card (Advantech, Taiwan) unning a program in C\u00ae (Microsoft) developed as part of this study. The FI determination of the diffusion coefficient of CuSO4 nvolved the use of a home-made conductimetric flow-through easuring cell described in detail elsewhere [31] and connected o a conductivity meter (Model 2100, TPS Pty Ltd, Australia)", "37 \u00d7 10\u22123 mol L\u22121 were injected into a deionized ater carrier stream flowing at 0.306 mL min\u22121. When the transient bsorbance curve reached a pseudo steady-state value the optode embrane was regenerated by sequentially flowing 2.00 mol L\u22121 Cl solution for 20 s, deionized water for 10 s and 4.00 mol L\u22121 odium acetate for 300 s. . Results and discussion .1. Determination of the diffusion coefficient of CuSO4 in queous solutions The repeatability of the FI experiments involving the injection of tandard CuSO4 solutions into a carrier stream of deionized water Fig. 1) was excellent since the concentration curves monitored at he same flow rate were almost indistinguishable from each other. he excellent agreement between the experimental concentration urves and the model predictions that best fit them is illustrated n Fig. 4. The diffusion coefficient of CuSO4 (Dm) obtained at 25 \u25e6C nd flow rates of 0.245, 0.369 and 0.522 mL min\u22121 was 8.71 \u00d7 10\u221210, .78 \u00d7 10\u221210 and 8.76 \u00d7 10\u221210 m2 s\u22121, respectively, with an average alue of 8.75 \u00d7 10\u221210 m2 s\u22121. .2. Determination of the diffusion coefficients and the on-exchange equilibrium constants of Cu2+ and CuL+ in Nafion embranes in their Na+ form The self-diffusion coefficients and the ion-exchange equilibrium onstants of cations Cu2+ and CuL+ were determined by fitting the odel to the experimental extraction data" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0001896_detc2013-13238-Figure8-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0001896_detc2013-13238-Figure8-1.png", "caption": "Figure 8. POSITION OF THE MEASURING PLANES.", "texts": [ " The Campbell diagram has been calculated in the machine speed operating range and a close-up is shown in Figure 2 in order to highlight the theoretical critical speeds, while Table 1 shows also the qualitative description of the corresponding mode shapes. The mode shapes that are prevalently vertical are shown in Figure 3 to Figure 7. The considered turbo-generator unit is equipped by measuring probes for vertical vibration only, in correspondence of seven measuring planes, which are close to the bearings, as shown in figure 8. These probes are composed of a proximity probe, which measures the gap between the shaft and the casing, and of a seismic probe, which measures the absolute displacement of the case, arranged in series. The condition monitoring system performs the vector sum between the shaft relative and the case absolute vibration in order to obtain the shaft absolute vibration. Finally, the signal is order tracked and the first three harmonic components are stored. 0 10 20 30 40 50 60 70 -10 -8 -6 -4 -2 0 2 4 6 8 10 RAFT\u00a9 4" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0002302_b978-0-12-803581-8.10351-0-Figure18-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0002302_b978-0-12-803581-8.10351-0-Figure18-1.png", "caption": "Fig. 18 Bottom shell (belly pan).", "texts": [ " The rim stiffener, shaded in red, consists of 11 layers around the cockpit opening. The 0\u201390 and 745 are plain weave and the 0 is unidirectional. 0290;745;0;0;0;0; 0; 0;0;0;0\u00bd In areas where the top does not bond to the side shells of the monocoque, honeycomb core is added to increase the flexural stiffness. The top shell consists of four layers of plain weave surrounding a 6 mm (\u00bc00) thick Nomex core in the region shaded in blue in Fig. 17. This region defines the top portion of the structural sidepod. 0290;745; core; 745;0290\u00bd Fig. 18 shows the bottom shell. For the most part the bottom shell is used to carry the outside skin loads from the left side shell to the right side shell. The region that is not in direct contact with the side shells, creates the lower portion of the structural sidepods and acts as the entry to the underbody diffuser. The bottom shell gives added strength where it is bonded to the monocoque and consists of two layers of plain weave that is permanently bonded to the left and right shells of the monocoque" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0001752_detc2011-48226-Figure3-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0001752_detc2011-48226-Figure3-1.png", "caption": "FIGURE 3. A CONSTRAINT SINGULAR CONFIGURATION OF THE 4-RUU PM.", "texts": [ " Points i and k correspond to two directions orthogonal to z and, therefore, these points belong to a line that cannot pass through point j unless i and k are coincident. Consequently, all terms A j vanish simultaneously iff points i, k, l, and m become all coincident. As a result, the 4-RUU PM reaches a constraint singularity iff: m1 \u2016 m2 \u2016 m3 \u2016 m4 (13) In such a configuration, the constraint wrench system of the manipulator degenerates into a 1-system and the moving platform gains one extra DOF, namely, the rotation about an axis directed along the common direction of mi (i = 1, . . . ,4), as shown in Fig. 3. In such a critical configuration, if the moving platform rotates about an axis of direction z, then the robot will come back to a non-singular configuration. 5 Copyright c\u00a9 2011 by ASME Downloaded From: http://proceedings.asmedigitalcollection.asme.org/pdfaccess.ashx?url=/data/conferences/idetc/cie2011/70657/ on 05/14/2017 Terms of Use: http://www.asme.org/about-asme/terms-of-use On the other side, in a constraint singular configuration, if the moving platform rotates about an axis of direction mi, the revolute joints attached to the moving platform will no longer be directed along z" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0002231_2013.10065-Figure5-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0002231_2013.10065-Figure5-1.png", "caption": "Figure 5. Vehicle movement model.", "texts": [ " If the slave is conducting tillage or planting, generally it can\u2019t change the pathway. Therefore, the speed-down must be chosen as a solution for obstacle avoidance. On the other hand, in case of hey transportation on a meadow, the slave can change the pathway to reach hey position. This solution is much better than speeddown in respect of work efficiency. SIMULATION Vehicle model To verify the performance of obstacle avoidance algorithm, the computer simulation was carried out. The vehicle movement model was built for this simulation. From Figure 5, the basic vehicle movement can be explained using Eqns (5) and (6). rf bYaYdtdrI 22/ \u2212=\u22c5 rf YYrdtdmV 22)/( +=+\u03b2 where, m and I are the mass and the moment of inertial of the tractor, respectively. \u03b2 ateral slip, r is yaw rate, a is the distance from the front axis to the center of the gravity, an is the distance from the rear axis to the center of the gravity. Yf and Yr is lateral force on the rear tires. is l d, b (5) (6) front and Where, Kf and Kr are cornering powers and \u03b1f and \u03b1r are slip angle on the front and rear tires which can be calculated using Eqns (9) and (10)" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0003838_acc.2019.8815221-Figure10-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0003838_acc.2019.8815221-Figure10-1.png", "caption": "Fig. 10. Proof-of-the-concept demonstration. Using EEG control, a subject should utilize a mentally emulated (1-to-2)(2) EEG redundant demultiplexer to control two motors of a robot arm to move the arm from a start region A towards a goal region B avoiding an obstacle C along the way.", "texts": [ " 9 shows an EEG demultiplexer with an EEG sentence encoded in two frames. In this realization, both the address and the data d are decoded from the same EEG frame. Address will be binary valued, and data d will be integer valued. In each EEG frame the subject tends to generate an EEG feature (e.g., increased amplitude of the alpha rhythm) that will be interpreted by both the EEG-to-address and the EEG-to-data decoders. To give a proof of the EEG demultiplexer concept, a demonstration scenario is designed as shown in Fig. 10. The subject observes a scene in which a robotic arm should move from a start region A to a goal region B, while avoiding an obstacle C along the way. The scenario uses a 5-motor robotic arm in which only two motors are allowed to move (i.e., two degrees of freedom are allowed), while the other three motors are fixed in some predefined positions. The controlled motors are motor (torso) for azimuth (horizontal) movement, and motor (wrist) for elevation (vertical) movement. The robot arm moves horizontally, and at the point in which the robotic arm reaches the obstacle, its horizontal projection should be less than , where is the horizontal distance between the robot arm base and the obstacle" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0001620_978-1-4614-3475-7_2-Figure2.28-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0001620_978-1-4614-3475-7_2-Figure2.28-1.png", "caption": "Fig. 2.28 Example 2.5", "texts": [ " The coordinate yC is yC = Mx A = a\u2212 sin a cos a 4(1\u2212 cosa) . The MATLAB program is given by syms a x y % dA = dx dy % A = int dx dy ; 0 0", " x = x0 k = Kmax l = 0 1: while dist(x1,x0)> \u03b5 do 2: while interp(V \u2217,k,x)> interp(V \u2217,k\u22121,x) do 3: k = k\u22121 4: i = mod(k,2)+1 5: j\u2217 = argmin j {interp(V \u2217,k,x+h f (x, i,\u03c9 j))} 6: x = x+h f (x, i,\u03c9 j\u2217) 7: \u03a9[l] = \u03c9 j\u2217 8: I[l] = i 9: l = l +1 return [W, I] Starting from initial state x0, Algorithm 3 performs the synthesis of feedback controls, and returns the vector \u03a9 of controls and the vector of system configurations I that allow reaching the target state x1. In loop 1-9, the synthesis of feedback controls is obtained by a reiterated Euler integration step until the distance between the current and target state is less than the allowed error \u03b5 . In the internal loop 2-3, the current value function is updated. Note that function interp(V,x) evaluates the value cost function at x as a multilinear interpolation of the vector V of the values of the cost function on grid vertices. Referring to the car-like model of Figure 1, set \u03b4max = 0.44 rad and l = 2.57 m. Setting v\u2212 = \u22121 and v+ = 1 in subsystems (3), (4), the maximum angular velocity is given by 1 l v+ tan\u03b4max = 0.183 s\u22121, so that W = {\u22120.183,0.183}. Let x1 a target state and let \u0393 be an ellipsoid centered at x1 with semi axes rx = ry = 0.12 m and r\u03b8 = 0.08. Figure 4 shows the path computed by Algorithm 3 for a parallel parking maneuver scenario using a different number of maximum maneuvers Kmax. In Algorithm 2, we used 525231 vertices to approximate the torus \u03a9 = [\u22128,23]\u00d7 [\u221210,10]\u00d7 [0,2\u03c0), with target state x1 = [2,8" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0003716_6.2019-4392-Figure20-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0003716_6.2019-4392-Figure20-1.png", "caption": "Fig. 20 Printed CAD Model of Fuel Inlet", "texts": [ "525 mm (0.375\u201d) hex head torque wrench fitting. More specifics regarding engine assembly can be found in Section VII, Subsection B Engine Assembly. Fig. 19 Section cut of printed CAD model of pintle, showcasing pilot hole design, hex head, and throat constriction feature. D ow nl oa de d by U N IV E R SI T Y O F G L A SG O W o n Se pt em be r 2, 2 01 9 | h ttp :// ar c. ai aa .o rg | D O I: 1 0. 25 14 /6 .2 01 9- 43 92 Fuel Inlet The fuel inlet was the simplest out of the five components. As seen in Fig. 20, the fuel inlet did not have any strict considerations made for printing. Primarily, the reason for the fuel inlet being printed was to maintain the material compatibility with the rest of the components, specifically the nozzle. The geometry of the fuel inlet could be easily machined (Fig. 21), but since there was extra space on the build plate, it made it more efficient to print it. Additionally, it served as a good test bed for our machinist to gain experience on machining heat treated Inconel 718 before continuing onto the more critical components (nozzle and chamber)" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0003190_978-3-030-05321-5_8-Figure8.24-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0003190_978-3-030-05321-5_8-Figure8.24-1.png", "caption": "Fig. 8.24 3D CAD model of 3 DoF wrist mechanism", "texts": [ "23 shows an example of an experimental device to induce strong external shock and vibrations to an artificial muscle. We applied an antagonistic drive system to verify the shock resistance by concrete chipping [22]. Hydraulic actuators areflexible and light artificialmuscles for robots that cangenerate very high forces due to the hydraulic pressure of oil. In addition, these muscles can be placed and twisted freely. In this section, we propose a 3 DoF wrist mechanism that is compact, flexible and light, which consists of hydraulic artificial muscles. Figure8.24 shows a 3D CAD model of the 3 DoF wrist mechanism. The Pitch axis and the Yaw axis in Fig. 8.24 are universal joints, and it is possible to operate the system with a 2 DoF using four hydraulic artificial muscles. In addition, the Roll axis that rotates the entire wrist is operated by two hydraulic artificial muscles. This Roll axis is the so-called twist rotation of the wrist, and the operation is realized by arranging the artificial muscle to wind around the structure by utilizing the flexibility of the hydraulic artificial muscle. In this example, an air pressure of 0.5MPa was applied for operation using all six hydraulic artificial muscles" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0002196_j.proeng.2017.01.226-Figure1-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0002196_j.proeng.2017.01.226-Figure1-1.png", "caption": "Fig. 1. Four-DOF rolling bearing dynamic mode.l", "texts": [ " Peer-review under responsibility of the organizing committee of the 13th Global Congress on Manufacturing and Management simulation value based on the model are calculated, and they have a good coherence. The effects of rotating speed and radial load on the vibration acceleration of the bearing is analyzed, and the vibrating acceleration of the bearing is also measured by a vibrating measurement instrument. 2. The 4-DOF bearing dynamic mathematical model The 4-DOF rolling bearing dynamic model [2], is shown as Fig. 1. In the model, it is assumed that the axial clearance and lubrication at various interfaces are ignored [3], and the bearing inner ring is fixed on the rigid shaft. The chamfer has little effect on the bearing internal stress, so it can be ignored. The general equations of motion for the 4-DOF bearing mathematical dynamic model can be expressed as formula (1). 1 1 1 1 1 1 11 2 2 2 2 2 2 22 0 0 x x y y x y m x c x F w m y c y F w m x c x F m y c y F (1) In which, m1, m2, c1 and c2 respectively represent the mass of shaft and inner race, the mass of outer race, damping factor of inner race and outer race; x1 and y1 respectively represent the displacements of the mass center of inner race; x2 and y2 respectively represent displacements of the mass center of outer race; wx and wy respectively represent the horizontal force and vertical force applied to the inner race of the bearing; Fx and Fy respectively represent the contact forces of the bearing in the x-direction and y-direction" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0003952_s12206-019-0936-3-Figure1-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0003952_s12206-019-0936-3-Figure1-1.png", "caption": "Fig. 1. Physiological structure of a locust.", "texts": [ " To solve the aforementioned problems, dynamic models for takeoff, flight, and landing buffering of the robot is established. The effect of structural parameters on the jumping performance of the robot is also analyzed in detail. The accurate takeoff parameters and stable jumping of the prototype indirectly confirm the correctness of the dynamic modeling method. This study provides a theoretical basis for the design of robots with good jumping performance. An experiment is conducted to understand the entire jumping process of a locust. The experimental object is an oriental migratory locust, which is shown in Fig. 1. The locust has three pairs of legs: front, middle, and hind legs. The front and middle legs are mainly used for walking, supporting the body, and landing buffering. The hind leg is the mutated leg, which is mainly used for jumping [17]. Fig. 2 shows the takeoff process of the locust. In the entire takeoff process, the angle between the tibia and femur (\u03c62) of the hind leg increases rapidly to realize jumping. The coordinate origin of coordinate system o2-x2y2z2 (Fig. 2(a)) is assumed to coincide with the contact point between the jumping leg and ground" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0000916_s1068371210100020-Figure4-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0000916_s1068371210100020-Figure4-1.png", "caption": "Fig. 4. Installation diagram for movements measuring: (1) probe microscope; (2) drive; (3) platform; (4) cantilever.", "texts": [ " Nevertheless, the presence of dry friction enables one to achieve a quies cent state during the formation of the gently sloping leading edge of the control signal. The NTMDT probe microscope (Zelenograd) as well as the PPIP 214 micromovement controller with M 214 inductive transducer were used to define the values of the linear movement of the screw in the pre sented drive. In measurements using a probe micro scope, a detecting head 1 (SMENA50 SEMI) and drive 2 were mounted on a heavy base 3 (Fig. 4). Microscope cantilever 4 was placed above the abutting end of the screw. Movements were registered in height measuring mode (height signal) via semicontactive techniques. The results are presented in Fig. 5. Mea suring device nodes, including the screw in drive 1, were grounded in order to reduce the capacitive pickup influence. Nevertheless, the existing interfer ence level did not permit one to conduct measure ments of extra small (less than 1 nm) step movements. A PPIP 214 micromovement controller was used to define the average value of these movements and the combined shift of the drive screw after the series of control signals was measured, after which the average value of the step movement was determined" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0001553_appeec.2011.5749147-Figure2-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0001553_appeec.2011.5749147-Figure2-1.png", "caption": "Figure 2. Outline and Dimension of Flow Channel (unit\uff1amm)", "texts": [ " Nafion 117 is chosen as the PEM in this study because it has excellent proton transmission ability. Here, the permanganate solution, 0.1mol/L, is used as the cathodic electron accepter. The concept of flow channel has been widely used in fuel cells for enhancing the power output because it can provide an effective flow motion and a larger flow mixing will provide a better power performance of fuel cell, but rarely applied to MFCs. According to these demonstrations, the concept of convergent flow channel with/ without plates, shown in Figure 2, will be utilized in MFC corresponding to the different inlet flow condition for realizing its effect in this study. Here, the outline of the chamber for SMFC will be made in a convergent type because it can increase the flow velocity along the downstream and the flow convection effect will be enhanced. The optimal dimension of the flow channel with/ without plates has been confirmed by Taguchi method, shown in Figure 2, and the largest flow mixing will be confirmed by numerical simulation. Under the natural flow condition, the motion is random for electron and proton. Therefore, the flow convection in convergent flow channel seems to be able to enhance the motion of electron and proton, a better power performance of SMFC could be expected reasonably. In addition, the effect of vortex flow induced behind the plates in the flow channel on the power generation will be investigated in this study. III. EXPERIMENTAL PROCEDURES Flow velocity and the resultant hydraulic retention time (HRT) are important two factors and need to be investigated before the operation of MFC in order to be successfully scaled up for wastewater treatment [26]" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0003574_icnsc.2019.8743340-Figure5-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0003574_icnsc.2019.8743340-Figure5-1.png", "caption": "Fig. 5. Schematic diagram of oscillation points in the same side.", "texts": [ " Case1: end beginq q\u2212 is odd number, which indicate that these two points are in the different sides of the chattering, as shown in Fig.4. It is obvious that ( , )end beginq q l\u03c1 > , thus set ( , )end beginq q n l \u03c1 = , ( , )d end beginm q q n l\u03c1= \u2212 \u00d7 ,and divide the line ( , )begin endL q q into n equal parts, namely n+1 new route points 1 2 1 1+ ( - ) 1+ ( - ) begin begin end begin n begin end begin n end q q q q q q n nq q q q n q q+ = = \u2212= = (5) Case2: end beginq q\u2212 is even number, which indicate that these two points are in the same sides of the chattering, as shown in Fig.5. If ( , )end beginq q l\u03c1 \u2264 , set ( , )s end beginm q q l\u03c1= \u2212 , 1 beginq q= and 2 endq q= , otherwise the processing means is the same as Case1. After eliminating the chattering route points, the length of the new obtained path is ( 1)all d sL N l m m= \u2212 \u00d7 + + (6) where allN is the number of all obtained route points after eliminating chattering, including the starting point and the target point. V. SIMULATION RESULTS In the following simulations, the paths are found by assuming that the robot moves at constant speed, and the resultant virtual force applied to it only determines the direction of its motion" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0000065_s1068371210080092-Figure5-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0000065_s1068371210080092-Figure5-1.png", "caption": "Fig. 5. Graphs generated according to Eq. (7).", "texts": [ "\u2013= = x3I\u00b71 xeI1'\u2013 x4I2\u2013 jE\u00b7 ;\u2013= jx4I1\u2013 I2z11'' I2' R11+ + 0;= I\u00b71x2 I1'x5 I\u00b72x2'\u2013 x1'I2'+ + jE\u00b7 ;= I1'\u2013 jx4' I2' z22'' I\u00b72R11+ + 0,=\u23a9 \u23aa \u23aa \u23aa \u23a8 \u23aa \u23aa \u23aa \u23a7 where x3 = xm \u2013 2xC1 + xe; x4 = xm \u2013 2xC1; = R11 + jx4; = R11 + j x5 = xm + xe \u2013 2xC1; = xm \u2013 2 Equation set (6) can easily be solved using the directed graphs method. For this purpose, we rewrite the set of equations as follows: (7) Figures 5a and 5b depict the current graphs gener ated according to Eq. (7). If we delete node 4, we obtain the graph presented in Fig. 5b. If we carry out some transformations according to the mentioned graph (Figs. 5a, 5b), we determine the loop currents and If we know the loop currents, z11'' z22'' x4' ; x4' xC1' . I\u00b71 jE\u00b7 Y3\u2013 I1'a1 I\u00b72a2;+ += I\u00b72 I1ja3 I2' a4;\u2013= I\u00b71' jE\u00b7 Y5\u2013 I1a5\u2013 I\u00b72a6 I\u00b72a7;\u2013+= I\u00b72' I\u00b71ja3 I\u00b72a9.\u2013= I\u00b71, I\u00b71' , I\u00b72, I\u00b72' , I\u00b73. 450 RUSSIAN ELECTRICAL ENGINEERING Vol. 81 No. 8 2010 MAMEDOV et al. we are able to determine the complex representation of real currents that pass through each branch as fol lows: (10) where The instantaneous currents that pass through the electric circuit are (9) where If we take into account Eqs" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0002156_j.triboint.2017.02.022-Figure1-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0002156_j.triboint.2017.02.022-Figure1-1.png", "caption": "Fig. 1. Example of contacting elements in relative motion, loaded by force F and separated by a thin lubricating film. Also shown is the tangent plane of the contact, P, and the contact area, S.", "texts": [ " However, a critical error in the calculation of the minimum film thickness is revealed, which may invalidate existing formulae in the literature under some kinematical conditions. For example, there are cases where an EHD film of nearly zero thickness is found whilst published formulae predict a sufficiently thick film, a result with obvious repercussions on wear. Moreover, significant errors on the traction coefficient and contact efficiency are found in some cases, which bring additional concerns over the use of some published formulae and studies that ignore tangential tractions intrinsically. An example of two solids in contact and in relative motion is shown in Fig. 1. The solids are loaded by a normal force, F. The contact may be Nomenclature a1 \u03bd s \u03c0E(2\u22122 ) /( )2 a2 \u03bd s \u03c0G(2 \u22121) /(2 ) b auxiliary variable (Eq. (A2)) c correction (Eqs. (10) or (13)) cx, cy coefficients; either 0 or 1 cz,x, cz,y auxiliary variables (Eq. (A3)) czz,x, czz,y auxiliary variables (Eq. (A4)) c1, c2 constants in the fluid mass-density function d sum of roughness local heights of the contact countersurfaces dz,x, dz,y auxiliary variables (Eq. (A5)) dzz,x, dzz,y auxiliary variables (Eq. (A6)) D R R R x R y+ \u2212 \u2212 \u2212 \u2212x y x y 2 2 2 2 Dx, Dy contact-ellipse semi-axial lengths E elastic modulus f, fx, fy functions (Eqs. (7)\u2013(9)) F contact normal load (Fig. 1) G E/(2+2\u03bd); shear modulus h film thickness (Fig. 1; Eq. (1)) hc, hmin central and minimum film thickness H height of the variator in Fig. 7 i, j; is, js discrete coordinates of square m \u03b1 \u03b710 /[5. 1\u2219ln( )+49. 317]9 0 n contact efficiency (Eq. (A15)) Nx, Ny numbers of partitions in Ox and Oy p pressure P tangent plane of the contact (Fig. 1) r radius of the variator in Fig. 7 rx, ry radii of curvature of the roller Rx, Ry radii of curvature s side semi-length of square S contact area (Fig. 1) sgn(\u2022) sign function ur, ud tangential velocities of the roller and the output disc U (ur+ud)/2; rolling velocity V ur\u2212ud; nominal sliding velocity w normal elastic displacement (Eq. (2)) wp pressure-induced normal elastic displacement (Eq. (3)) ws normal elastic displacement (Eqs. (6) and (11)) wx, wy traction-induced normal elastic displacements (Eqs. (4) and (5)) x, X, y, Y, z coordinates Greek symbols \u03b1 pressure-viscosity coefficient \u03b2 x X y Y( \u2212 ) + ( \u2212 )2 2 \u03b3, \u03b4 coefficients in the \u03c4L-function \u03b5 normalized correction (Eq", " \u23a7\u23a8\u23a9 \u23ab\u23ac\u23ad \u222cw x y \u03bd \u03c0E\u03b2 p X Y \u03bd \u03c0G\u03b2 X x \u03c4 X Y Y y \u03c4 X Y dXdY ( , ) = 1\u2212 ( , ) + 2 \u22121 4 [( \u2212 ) ( , ) + ( \u2212 ) ( , )] zx zy 2 2 (2) where \u03b2 x X y Y\u2254 ( \u2212 ) + ( \u2212 )2 2 , \u03bd is the Poisson's ratio, E and G = E/ (2 + 2\u03bd) are the elastic and shear moduli, respectively, and \u03c4zx and \u03c4zy are the tangential tractions in directions Ox and Oy, respectively. The integration in Eq. (2) is carried out over the entire contact area, which, for lubricated contacts, is normally larger than the Hertzian (dry) contact area. Tractions \u03c4zx and \u03c4zy result from the shearing of layers of the medium such as a liquid lubricant in the case of Fig. 1, as well as from any shearing of interlocked roughness asperities. Whilst it is obvious that tangential tractions affect normal displacements via Eq. (2) and through those the film thickness h in Eq. (1), the vast majority of publications, including textbooks on related topics, simply ignore them. They retain only the first term in the integrand of Eq. (2), namely the term involving pressure p, which is equivalent to setting \u03c4zx=\u03c4zy=0 without justification, effectively dealing with a frictionless contact" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0003741_012077-Figure2-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0003741_012077-Figure2-1.png", "caption": "Figure 2. The first folding mechanism with overlapped compact, partially deployed and totally deployed configurations.", "texts": [ " Modern Technologies in Industrial Engineering VII, (ModTech2019) IOP Conf. Series: Materials Science and Engineering 591 (2019) 012077 IOP Publishing doi:10.1088/1757-899X/591/1/012077 If we consider 1 the rotational angle of the driver link, the area A1 of a wall formed by this mechanism in the deployed configuration will be equation (3): 1 2 1 sin2 AB lA (3) considering that ADAB ll . Another view of the mechanism with overlapped compact, partially deployed and totally deployed configurations is presented in figure 2. Two examples of deployable structures that may use this mechanism will be given here. The first example is a mobile house (see figure 3). Another application of the mechanism described here road barrier, represented in folded, partially deployed and totally deployed configurations (figure 4). Modern Technologies in Industrial Engineering VII, (ModTech2019) IOP Conf. Series: Materials Science and Engineering 591 (2019) 012077 IOP Publishing doi:10.1088/1757-899X/591/1/012077 The first simulations have been done, considering the coordinates of the mechanism nodes in the fully extended configuration as illustrated in figure 5" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0000244_icma.2012.6285106-Figure2-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0000244_icma.2012.6285106-Figure2-1.png", "caption": "Fig. 2 One joint mechanical structure.", "texts": [ " In this section, mechanical design, sensory systems and hardware architecture of the joint are outlined. To reduce the weight of joint and increase the joint output torque, the structure of the motor adding harmonic drive gear 1868978-1-4673-1278-3/12/$31.00 \u00a92012 IEEE Proceedings of 2012 IEEE International Conference on Mechatronics and Automation August 5 - 8, Chengdu, China is adopted to build the joint. The frameless motor and special short and light weight harmonic drive gear are used. Most mechanical parts are made of aluminum. The 3D model of one joint mechanical structure is shown in Fig.2. Each electronic module is of the same type and consists of a motor drive board, the FPGA control board, the power board and sensor conditioning board. All these circuit boards, motor and harmonic drive gear are placed inside the housing to save the space. All housings are made of aluminum and are designed to transfer thermal energy from motor and power electronics to themselves. To overcome joint flexibility, increase joint sensory capability and reliability, earch joint is equipped with multisensor as shown in Table I" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0002541_rnc.3892-Figure1-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0002541_rnc.3892-Figure1-1.png", "caption": "FIGURE 1 Input nonlinearity function", "texts": [ " The nonlinear function \u0393(u) is defined as follows: \u0393(ui) = \u23a7\u23aa\u23a8\u23aa\u23a9 \ud835\udf13+ i (ui)(ui \u2212 u+ i ), if ui > u+ i 0, if u\u2212 i \u2a7d ui \u2a7d u+ i \ud835\udf13\u2212 i (ui)(ui + u\u2212 i ), if ui < u\u2212 i , (2) where \ud835\udf13+ i (ui) > 0 and \ud835\udf13\u2212 i (ui)> 0 are unknown right and left nonlinear functions of ui, u+ i and u\u2212 i are break points of the input nonlinearity, ui represents the i \u2212 th input of the plant, i = 1, 2, \u00b7 \u00b7 \u00b7 ,m. Assumption 1. The specific expressions of right and left nonlinear functions \ud835\udf13+ i (ui) and \ud835\udf13\u2212 i (ui) are unknown, but \ud835\udf13+ i (ui) and \ud835\udf13\u2212 i (ui) satisfy { \ud835\udefc+i \u2a7d \ud835\udf13+ i (ui) \u2a7d \ud835\udefd+i , ui > u+ i \ud835\udefc\u2212i \u2a7d \ud835\udf13\u2212 i (ui) \u2a7d \ud835\udefd\u2212i , ui < u\u2212 i , (3) where ( \ud835\udefc+i , \ud835\udefd+i , \ud835\udefc\u2212i , \ud835\udefd\u2212i ) , called gain reduction tolerances, are known positive constants. The input nonlinearity of \u0393(ui) is shown in Figure 1. For the convenience of control design, the input nonlinearity \u0393(ui) can be redefined as the following form: \u0393(ui) = mi(t)ui + di(t), where mi(t) = { \ud835\udf13+ i (ui), ui > 0 \ud835\udf13\u2212 i (ui), ui \u2a7d 0 , (4) and di(t) = \u23a7\u23aa\u23a8\u23aa\u23a9 \u2212\ud835\udf13+ i (ui)u+ i , ui \u2a7e u+ i \u2212mi(t)ui u\u2212 i < ui < u+ i \ud835\udf13\u2212 i (ui)u\u2212 i , ui \u2a7d u\u2212 i . (5) Since (1) is nonlinear nonminimum, the relative degree of (1) is r, and r < n, n is the order of the system. Following the input/output linearization approach proposed in Isidori,16 the nonlinear system (1) can be partially linearization" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0002302_b978-0-12-803581-8.10351-0-Figure14-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0002302_b978-0-12-803581-8.10351-0-Figure14-1.png", "caption": "Fig. 14 Structural sidepod components. Reproduced from Weidner, L.R., Radford, D.W., Fitzhorn, P.A., 2003. A multi-shell assembly approach applied to monocoque chassis design. In: SAE 2002 Transactions, Journal of Passenger Cars \u2013 Mechanical Systems, p. 2486.", "texts": [ " However, with the multi-shell design, the strakes can be readily molded, directly into the left/right shells. A cross-section view of the lower strake is seen in Fig. 12 and of the upper strake in Fig. 13. In addition to these strakes, which generate a crush structure at the leading edge region of the sidepods, a portion of the sidepod itself is incorporated into the structure of the monocoque. This structural sidepod is generated as a portion of the top and bottom shells and resembles short wings as seen in Fig. 14. For additional reinforcement and integration, the upper shell is bonded to the strake (Fig. 13). The structural sidepod portions of the top and bottom shells are joined by an additional vertical panel on each side of the car that completes the outer structure. These vertical panels are used for hardware mounting (radiator, etc.) in the current design (Fig. 14). The horizontal portions of the structural sidepods that are integral with the top and bottom shells are used to enhance the lateral stiffness of the monocoque and to generate an integral side crush zone. This construction would not be possible in a fabrication approach that only makes use of either a top/bottom, or left/right pair of tools. This current section expands upon the simplified conceptual design of a multi-shell composite monocoque described in the previous section, by describing, in more detail, the design development of the various composite components required for the structure of the chassis" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0001286_iros.2011.6094587-Figure7-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0001286_iros.2011.6094587-Figure7-1.png", "caption": "Fig. 7. Tracing of normal axes and manifolds in experimental trials with the PA10 manipulator: ni axes converge at the constant point Oh, i.e. the shoulder joint.", "texts": [ " More specifically, methods involving nonlinear optimization are quite dependent on the accuracy of starting estimate of solution. Experiments are intended for validation of correctness of the procedure, consistency with simulation data and evaluation of accuracy in real systems. The experimental setup includes a 7 DoF Mitsubishi PA-10 manipulator and the 7the markers are placed before joint 6, so qj are assigned only to relevant joints SMART passive markers tracker by BTS Bioengineering. A set of N = 4 pivoting axes with M = 10 markers positions each is obtained from robot motion (Fig. 7), tracking the corresponding markers positions p\u0303ij for the new method as in procedure hand-eye from manifolds (Fig. 5). A set of additional P = 31 robot poses is generated uniformly varying the joint rotation values, and capturing the corresponding p\u0303ij as well. Such P poses are used for estimating X\u0302 and Z\u0302 for comparison algorithms. The magnitude noise is measured during the execution of robot poses averaging the difference between sampled and nominal length of segments among the cluster of markers" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0001342_0954406213479272-Figure2-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0001342_0954406213479272-Figure2-1.png", "caption": "Figure 2. Lateral view of harmonic drive-based gear pump\u20142: flexible gear; 3: rigid gear; 4: boom; 6: left side plate; 7: mandrel; 8: drive shaft; 9: pipe; 10: right side plate; c and e: inlet holes; j: suction port; d and f: outlet holes; k: discharge port.", "texts": [ "comDownloaded from Two mating gears, having the same tooth width and circular pitch but different tooth number (the tooth number of the flexible gear is slightly less than that of the rigid gear, usually less than two teeth), turn in an inner chamber surrounded by the casing and end cover. The rigid gear outside is the driving gear. The speed ratio uGR of rigid gear to flexible gear is uGR \u00bc ZR=ZG \u00f02\u00de Crescent-shaped sealing blocks (one is on the left and the other on the right, as shown in Figure l) are setup in radial direction between the addendum circle of the rigid gear and the addendum arc surface of the flexible gear. Two gears match closely with two side plates (6 and 10 in Figure 2) in the axial direction. Two sealing blocks (5 in Figure 1) are fixed on the side plates by position pins, two end covers are binded in the casing (1 in Figure 1) using bolts. The flexible gear, rigid gear, sealing blocks and side plates separate the space between the rigid gear and the flexible gear into four independent airtight cavities: two cavities marked A for suction and two cavities marked B for expelling fluid (see Figure l). From Figures 1 and 2, it can be seen that the two gears, sealing blocks and side plates constitute closed cavities A and B to realize the suction and expelling of fluid" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0002311_ls.1385-Figure1-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0002311_ls.1385-Figure1-1.png", "caption": "FIGURE 1 Structure of hydrostatic bearing and coordinates establish", "texts": [ " The dynamic characteristics of the rotor considering velocity slip are analyzed and compared with those in nonslip condition. The effect law of dynamic characteristic of the rotor influenced by velocity slip in microscale is analyzed. The model of dynamic system of the rotor with the mass eccentricity is established, and the rotor axis orbit in different eccentricity mass distance is analyzed. This provides a theoretical basis to improve the dynamic stability of a rotor\u2010bearing system. The structure of hydrostatic bearing is shown in Figure 1A. With a certain load acting on the rotor, the rotor axis and the bearing center are no longer coincided, and a certain deviation will exist. The final equilibrium position of the rotor is O1, e is the position deviation of 2 centers and angle deviation is \u03c6 setting the 2 axis lines as the starting point. The pressure of lubricating oil is Ps flowing into oil chamber of 1, 2, 3, and 4. Figure 1B shows the coordinate establishment of oil film flows, x, y, and z represent the oil film circumferential, radial, axial direction, respectively. Table 1 for the hydrostatic bearing parameter table. Dynamic characteristics of hydrostatic bearing reflects the change of oil film force when the rotor is deviated from the original balance position and does micro vibration near the position. Therefore, dynamic analysis and calculation must be based on the Reynolds equation in unsteadiness condition. The Formula 1 in essay15 is the Reynolds equation when the rotor is under steady operation, the right side of the equation is 0" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0001609_0954406211427833-Figure7-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0001609_0954406211427833-Figure7-1.png", "caption": "Figure 7. Simulated centre tooth of a concave gear (tooth surface highlighted for clarity).", "texts": [ " The same steps are followed for the rest of the teeth at the inner ring. For machining the teeth which are at the outer ring, the work solid is tilted by 24 about the X-axis with respect to the sphere centre (Figure 6(c)). Similarly, for simulating the other teeth at the outer ring, initially, the work solid is first indexed by 30 about the Z-axis and then tilted by 24 about the X-axis. The same steps are followed for the rest of the teeth at the outer ring. CAD model after simulation for the centre tooth of the concave gear is shown in Figure 7. By highlighting the tooth surface, the curvature on it can be seen distinctly. Figure 8(a) shows the CAD-simulated profile obtained by taking an axial section of the centre tooth along with the superimposed profile obtained analytically following the procedure outlined in \u2018Analytical method of tooth surface generation\u2019 section. Since the difference cannot be seen clearly, the deviations in the X-coordinate values are plotted along different Y-coordinate values, taking the analytical profile as a reference" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0002636_icma.2017.8015910-Figure1-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0002636_icma.2017.8015910-Figure1-1.png", "caption": "Fig. 1 Mechanical structure of planetary roller screw The parameters of planetary roller screw are shown in Table.1", "texts": [ " This model provides a new way to calculate the PRSM axial stiffness. Use the FEM software to verify the contact model. The entire axial stiffness is tested by experiment and an empirical formula. From the experiment results, it can be seen that this model is more accurate than empirical formula. Compared with empirical formula, several parameters are considered in this stiffness model. The simulation results shows this model is more convenient than FEM. This paper provides a better method for planetary roller screw parametric design. Fig. 1 is the mechanical structure of planetary roller screw, it mainly consists of 11 rollers, a screw, a nut and two planet carriers. Table.1 Parameters of planetary roller screw Geometric parameter Reference value Pitch diameter of the screw (mm) 21 Bottom diameter of the screw (mm) 20.82 Full diameter of the screw (mm) 21.14 Thread number of the screw 5 Helix angle of screw (\u00b0) 1.74 Pitch diameter of the nut (mm) 35 Thread number of the nut 5 Helix angle of the nut (\u00b0) 1.04 Pitch diameter of the roller (mm) 7 Thread number of the roller 1 Helix angle of the roller (\u00b0) 1" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0003912_lpt.2019.2942646-Figure2-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0003912_lpt.2019.2942646-Figure2-1.png", "caption": "Fig. 2. Detailed description of the proposed intraocular lens positioning hook: (a) the overall structure; (b) configuration of the sensitive elements; (c) photograph of the intraocular lens positioning hook prototypes; (d) magnified views of the bonding parts; (e) the magnetron sputtering machine.", "texts": [], "surrounding_texts": [ "Intraocular lens positioning hooks made of titanium alloy are selected as bases for prototyping. As shown in Fig.1 (a), the sensing elements of FBGs are located in the connecting rod, which serves as an active sensing portion is sensitive to both the tangential force terms (Fx and Fy). The detailed layout of the FBGs is shown in Fig.1 (b). Four FBGs labeled as FBG1, FBG2, FBG3, FBG4 are arranged around the circumference of the connecting rod at uniform intervals of 90\u25e6. These four FBGs were fabricated along single model optical fibers by using phase mask technology. The effective length and the reflectivity of each FBG are about 4 mm and 80%, respectively. FBG1 and FBG3 are located in the XOZ plane, whereas FBG2 and FBG4 are located in the YOZ plane. These four FBGs are attached flush to the surface of the connecting rod, such that the entire length of each grating is evenly pasted, and the polymer coating of each grating has been stripped using a carbon dioxide laser. After encapsulation, the central wavelengths of FBG1, FBG2, FBG3 and FBG4 are 1546.217 nm, 1549.105 nm, 1552.080 nm and 1555.317 nm, respectively. In this letter, two intraocular lens positioning hook prototype has been fabricated, as shown in Fig.1 (c). The bonding part of one prototype has been plated with a layer of titanium coating using a magnetron sputtering machine. While the other prototype has not been plated. The detailed presentations of the bonding parts of these two hooks are shown in the magnified view in Fig.1 (d), and the magnetron sputtering machine (BESTEC Wuhan-II Sputtering System) for coating is shown in Fig.1 (e). When Fx is exerted on the end of the hook, FBG1 and FBG3 will be stretched and compressed, respectively; and experience opposite strain values of the same magnitude due to the symmetrical layout. Thus we can assume the Fx -induced wavelength shifts of FBG1 and FBG3 are \u03bb(F x\u2212bending) and - \u03bb(F x\u2212bending), respectively. Whereas FBG2 and FBG4 are located in the neutral layer, making these two FBGs insensitive to Fx . When Fy is exerted at the end of the hook, the connecting rod produces both bending and torsional deformations. As response to the torsional deformation, FBG1-FBG4 exhibit identical wavelength variations which can be assumed as \u03bb(Fy\u2212torsion). Whereas the Fy-induced bending deformation will cause FBG2 and FBG4 to be stretched and compressed, respectively, which results in opposite wavelength shift values of FBG2 and FBG4 namely \u03bb(Fy\u2212bending) and - \u03bb(Fy\u2212bending), respectively. Moreover, suppose the temperature variation is T , then the T -induced wavelength shifts of these four FBGs are identical, it can be labeled as \u03bb( T ). Therefore, the wavelength shifts of FBG1 and FBG3 ( \u03bb1 and \u03bb3) can be described as: \u03bb1 = \u03bb(F x\u2212bending) + \u03bb(Fy\u2212torsion) + \u03bb( T ) (1) \u03bb3 = \u2212 \u03bb(F x\u2212bending) + \u03bb(Fy\u2212torsion) + \u03bb( T ) (2) Thus, the following equation can be obtained: \u03bb1 \u2212 \u03bb3 = 2 \u00b7 \u03bb(F x\u2212bending) = 2 \u00b7 Ka \u00b7 Fx (3) where Ka represents the sensitivity of FBG1 under Fx . Thus, the differential operation of the resulting wavelength shift values of FBG1 and FBG3 ( \u03bb1- \u03bb3) can determine Fx after calibration. Similarly, Fy can be represented by the difference between the wavelength shift values of FBG2 and FBG4 ( \u03bb2- \u03bb4). From the above analysis, we can conclude that \u03bb1- \u03bb3 and \u03bb2- \u03bb4 can respectively represent Fx and Fy after calibration. The torsion effect and the temperature cross sensitivity have all be eliminated after the subtraction operations." ] }, { "image_filename": "designv11_62_0002865_we.2149-Figure7-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0002865_we.2149-Figure7-1.png", "caption": "FIGURE 7 Model of the cage hinge with design parameters and loads", "texts": [ " Therefore, the total number of springs, which are evenly distributed around the rotational axis in a radial orientation, and their dimensions are chosen as the design parameters that need to be determined TABLE 1 Design parameters of the initial design Design parameter Initial value Number of springs 12 Length of springs 2 m Inner diameter 1 m Outer diameter 2 m Thickness of springs 40 mm load step (LS) 1 2 3 4 5 torsion (\u03b1) axial force (F axial ) root bending moment (M root , M rootUlti ) FIGURE 8 Load steps to fulfill the design requirements (see Figure 7). In detail, these are the number of springs as well as their length, thickness, and width (the inner and outer diameter of the hinge). A preselected set of parameters, the initial design, is listed in Table 1, which is adapted to our needs in the sections that follow. The goal of the dimensioning process is to adapt the cage hinge to withstand the external loads while providing the elastic angular range of\u00b13\u25e6 by torsion. The torsion (embodied by rotation angle \ud835\udefc) induces stresses in addition to the previously determined external loads, the axial force Faxial, and the root bending moment Mroot (see Figure 7). The pitch moment does not need to be considered because it is passed through the hinge. Depending on its direction of action, it will either support the torsion of the cage hinge or counteract it. In both cases, it is the actuation drives rather than the springs that are loaded\u2014this becomes clearer when looking at the whole concept in the sections that follow. Different combinations of the loads are applied in several load steps (LSs) (see Figure 8), to investigate the individual and combined influences on the stresses and strains in the springs" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0000418_detc2011-47599-Figure5-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0000418_detc2011-47599-Figure5-1.png", "caption": "FIGURE 5. Forming a polyhedron containing the origin for forceclosure of planar cable-driven closed chains", "texts": [ " (29), and force-closure condition of a general n-DOF closed chain is mathematically described as follows: \u2200W \u2208\u211c n, \u2203ti > 0, 3 AT = W (30) Proposition: An n-DOF cable-driven closed chain must have a minimum of n+1 cables to achieve force-closure. From literature, achieving force-closure can be equivalently represented as column vectors ais forming a convex hull containing the origin (see [4] for proof). Looking at eqn. (29), which is the force/moment equilibrium equation of the planar two-3R cable-driven closed chain, W represents the 3-D orthogonal space of moments, which is spanned by the column vectors ai (see Fig. 5). For force-closure, the ais must form a convex hull (i.e., a polyhedron), and it must encompass the origin. The smallest polyhedron is a tetrahedron, which requires a minimum of 4 vertices, where each vertex represents a ai. As seen from Fig. 5, a minimum of four ais are necessary to form the polyhedron and encompass the origin. This shows that a minimum of four cables are necessary for the 3- DOF planar cable-driven closed chain to achieve force-closure. Similarly, extending the above analogy to an n-DOF cabledriven closed chain, it would require a minimum of n+1 cables to form a convex hull in n-dimensional space and achieve forceclosure. 6 Copyright c\u00a9 2011 by ASME Downloaded From: http://proceedings.asmedigitalcollection.asme.org/ on 02/01/2016 Terms of Use: http://www" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0002881_978-3-319-66697-6_32-Figure1-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0002881_978-3-319-66697-6_32-Figure1-1.png", "caption": "Fig. 1 Spiral bevel gear model", "texts": [ " (2012) studied the dynamic behavior of bevel gear transmission in presence of teeth defects in stationary conditions. This chapter will investigate the effect of local damage on the dynamic behavior of bevel gear which is a major component in the gearbox of an excavator in nonstationary conditions by modeling single stage bevel gear transmission using a lumped parameter model. The simulation of the dynamic behavior is achieved in the time domain using an implicit Newmark\u2019s time-step integration scheme. We consider a dynamic model of spiral bevel gear divided into two rigid blocks as shown in Fig. 1. Each block has five degrees of freedom: three translations xi, yi, zi (i = 1, 2) and two rotations \u03b8m and \u03b81 for the pinion and \u03b82 and \u03b8r for the wheel. The gear bodies are assumed to be rigid cone disks and the shafts are considered with torsional stiffness. The mesh stiffness is modeled by a linear stiffness ke t\u00f0 \u00de acting along the line of action. The driving torque is related to the motor speed as following (Wright 2005): Cm = Tb 1+ sb \u2212 s\u00f0 \u00de2 a s \u2212 bs2 \u00f01\u00de where a and b are constant properties of the motor, sb and Tb are the slip and torque at breakdown, and s is the proportional drop in speed given by: s=1\u2212 \u03a9 \u03a9s \u00f02\u00de where \u03a9s is the synchronous velocity of the motor" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0003941_j.vacuum.2019.108992-Figure9-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0003941_j.vacuum.2019.108992-Figure9-1.png", "caption": "Fig. 9. Contact angle (\u03b8) made by water drops on substrate after each fabrication step. (a) Polished Si wafer; (b) Etched Si wafer; (c) Etched Si wafer with 400 nm Fe2O3 coated; and (d) Etched Si wafer with Fe2O3 coating followed by ZnO coating.", "texts": [ " It is envisaged that the multiscale feature size of different coatings together with inward bending will help in efficient trapping of incident light through multiple reflections causing loss of energy until it gets fully absorbed in UV\u2013Vis-IR region. V. Ghai et al. Vacuum 171 (2020) 108992 Hydrophobicity of absorbing surfaces is an important consideration to avoid their environmental assisted corrosion in solar PV applications [46]. To this end, contact angle of multilayered assembly with an average absorption of more than 98.7 in 300\u20132000 nm range is measured after each fabrication step in Fig. 9. The plane polished wafer shows affinity towards water droplets with a contact angle of 48\ufffd [47]. KOH etching further increases the hydrophilicity of polished wafer with a lower value of contact angle of 23\ufffd (Fig. 9b). Even though contact angle increases up to 70\ufffd (Fig. 9c) after Fe2O3 coating on etched wafer, it still remains in the hydrophilic region. However, the behavior of assemble switches from hydrophilic to hydrophobic after deposition of zinc oxide nanostructures on iron oxide layer with a contact angle of around 135\ufffd as shown in Fig. 9d. The hydrophobicity of optically graded assembly reveals that it will not absorb any water from surroundings and hence will resist moisture-induced degradation of absorbing surface. This is a desirable feature for the developed surface for its industrial applications where degradation of absorbing surfaces is predominantly caused due to moisture absorption from the surroundings. In summary, we have studied the role of surface texturing on light absorption. Various etching solutions (KOH, HNA, KOH followed by HNA and HNA followed by KOH) are used for texturing the silicon wafer V" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0002131_10402004.2017.1285970-Figure2-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0002131_10402004.2017.1285970-Figure2-1.png", "caption": "Figure 2: Schematic representation of the Maxwell model", "texts": [ " The ultrasonic reflection coefficient is, then, correlated to the acoustic and viscoelastic properties of the solid-liquid interface. Lubricating oils are, usually, non-Newtonian and an algorithm that reflects the viscoelastic behaviour of these fluids is needed. It is widely accepted that the viscoelastic Maxwell model describes sufficiently well the interaction between an ultrasonic shear wave and a non-Newtonian liquid with a dominant relaxation time at a solid-liquid interface [12, 13, 14]. Figure 2 schematically shows the interaction between a solid particle ACCEPTED MANUSCRIPT 6 and a liquid particle at a solid-liquid interface with a spring-dashpot Maxwell model when an oscillatory shear stress is applied. The damper element models the relaxation effects of a viscoelastic system as ultrasonic shear occurs at the solid-liquid boundary, while the spring element is used to model the instantaneous materials deformation. Lamb [6] obtained the viscoelastic properties of a Maxwell liquid under oscillatory shear as: ( ) [ ( ) ] {[ ( ) ] ( )} (3) ( )[ ( ) ] {[ ( ) ] ( )} (4) Where is the infinite shear modulus and G\u2019 and G\u2019\u2019 are derived from the reflection coefficient by combining equations (1) and (2)" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0002226_ut.2017.7890311-Figure1-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0002226_ut.2017.7890311-Figure1-1.png", "caption": "Figure 1 Equipment layout of SBPAUV", "texts": [ " SBP is an acoustic instrument to investigate precisely geological layer structure under several tens of meters below the seabed mainly for exploration of hydrothermal deposits. The second AUV (called general purpose type AUV) can change the payload according to the purpose of investigation. Multi beam echo sounder (MBES) is equipped for survey of seabed topography in the second AUV. And high altitude image mapping device will be equipped for survey of seabed resource. Table 1 shows the main specification of the two AUVs. Lengths of both AUVs are less than 4 meters. Figure 1 and Figure 2 show the equipment layout of each AUV. Two AUVs are equipped with the same acoustic positioning device and underwater acoustic communication device for the simultaneous operation of multiple AUVs. 2.2 Patterned behavior The AUVs can behave underwater with the combination of predefined behavior pattern [2]. The patterned behavior control helps AUV\u2019s operators to set up mission plan easier. Table 2 shows predefined behavior patterns of our AUVs. 3. Sea Trial After finishing the assembling, the land test and the water tank test, we conducted the sea trial for each AUV in Suruga Bay" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0001166_gt2012-69356-Figure6-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0001166_gt2012-69356-Figure6-1.png", "caption": "Figure 6: Rig set-up schematic for the calibration run", "texts": [ " High pressure room temperature air is introduced in the middle section through a swirl-plate and leaks out through the seals to downstream chambers. These chambers are connected to maintain the same downstream pressure to avoid excessive axial thrust on the bearings. The swirl plate can be adjusted to test the seal under varying inlet swirl conditions, but this feature was not used during the reported tests. The mass flow rate is measured upstream of the rig by a Coriolis flow-meter to get accurate readings for a high turn-down ratio. In the calibration run, two identical labyrinth seals are tested as shown in Figure 6, and an effective clearance map is constructed for different upstream and downstream pressures and rotor speeds. In the Compliant Plate Seal test run, indicated in Figure 7, the high pressure air is introduced in the middle section through a swirl\u2013plate and leaks out to the downstream chambers \u2013 on one side through the Compliant Plate Seal that is being tested and on the other side through the calibrated labyrinth seal. The rig and test conditions are described in brief in Table 2. The total leakage is measured upstream of the rig, and the leakage through the Compliant Plate Seal is calculated by subtracting the leakage through the calibrated labyrinth seal" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0003077_icetcct.2017.8280315-Figure1-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0003077_icetcct.2017.8280315-Figure1-1.png", "caption": "Fig. 1: Rotary inverted pendulum [15]", "texts": [ " Here in this paper two control strategies of adaptive control are made, one of which is explicit self- tuning regulator and the other is an MRAC technique to stabilize the rotary inverted pendulum. After the brief introduction in section I, section II gives the modeling of the rotary inverted pendulum system, section III discusses the control strategies of STR and MRAC following which, in section IV the implementation is discussed. Section V draws the results and conclusion followed by references. II. ROTARY INVERTED PENDULUM The rotary inverted pendulum which has been used in this paper is given in fig. 1. This pendulum was invented by K. Futura at Tokyo Institute of Technology, which is why it is often called Furuta pendulum [14]. It has two degrees of freedom with a single actuator; the pendulum arm is connected to a rigid link at one end and another side to the pendulum. The servo motor provides the necessary actuating signal or the control signal. is the pendulum arm angle and is the pendulum angle. The assumptions considered for modeling of the system are [16]: \u2022 Zero initial conditions are inferred when system initiates from equilibrium state" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0003985_icamechs.2019.8861668-Figure9-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0003985_icamechs.2019.8861668-Figure9-1.png", "caption": "Fig. 9. Mechanism of the leg when flexed", "texts": [ " The joint will move freely when the vertical force is not applied on it , and this will allow to passively avoid obstacles by rotating the joint using reactive forces from the obstacles as shown in Fig. 8. However,, when a vertical force is applied to support the weight of the robot the rubbers of the joint engages, and the joint becomes locked. Therefore, the stiffness of the leg becomes high and the leg can transmit the force from the motor to the ground as a driving force. The advantage of the proposed mechanism in comparison to the conventional mechanism is that the joint can be locked at any angle as shown Fig. 9. Thus, it can produce a driving force when moving through a narrow space. The mechanism of the leg when flexed is given in Fig. 9. The trunk is illustrated in Fig. 10. It is flexible and can adapt to the rough ground during locomotion. The trunk was made of duplex bellows as shown in Fig. 10 i.e. smaller bellows inside larger bellows for the realization of different stiffness i.e. for small flexure and large flexure, respectively. The trunk is made up of series of connected rectangular prisms, and has high stiffness in the vertical direction for upliftment of its weight and low stiffness in the horizontal direction to allow for the robot to turn" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0002669_s11837-017-2597-y-Figure1-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0002669_s11837-017-2597-y-Figure1-1.png", "caption": "Fig. 1. (a) Front view of Arcam A2 machine with door and front panel open. (b) Thermal image from IR camera. Temperatures are generated with the same emissivity value of 0.26,6 so the powder temperatures in the image are not representative of actual powder temperature. Cylinder labels, direction label, and scale bar are added for reference. (c) Side view of IR camera mounted above chamber view port.", "texts": [ " The beam is then used to selectively melt the powder in the areas dictated by the original computer-aided design (CAD) model. The build platform is lowered by a specified layer thickness (50 lm or 70 lm for Ti64), more powder is spread, and the fusion process is repeated until the part is complete.5 To investigate temperatures at the surface of parts during production, the University of Texas at El Paso (UTEP) has instrumented an Arcam A2 machine with an FLIR SC645 infrared (IR) camera, as shown in Fig. 1, and developed software necessary to collect images at specified times during and after material fusion for each layer.2,6 To minimize metallization of the viewport and provide the best possible imaging conditions, a shutter inside the build chamber has been incorporated into the system. Coupled with prior research on view factors specific to the Arcam A2 machine and the emissivity of solidified Ti64, surface temperatures can be extracted from IR images.7 This information DOI: 10.1007/s11837-017-2597-y 2017 The Minerals, Metals & Materials Society enables monitoring of absolute surface temperatures after a layer has been melted", " CONTINUOUS MONITORING OF A SINGLE LAYER All builds in this work were performed at UTEP using Arcam-supplied Ti64 powder and the standard Ti64 theme for 50 lm layers on the Arcam A2 machine with software version 3.2.121. Builds were made using a standard Arcam build plate, and measurements were taken at build heights greater than 3 mm to avoid transient effects in the initial layers of the build. An initial assessment of temperatures at the top surface of the Arcam A2 powder bed during fusion was conducted by continuously monitoring a single, 15-mm-diameter cylinder at 20 Hz using the thermography setup described in Fig. 1. Figure 2 shows temperatures extracted from the IR images by averaging a 12-pixel-diameter (16 mm2) area at the center of each cylinder. The temporal and/or spatial resolution of the IR camera are not adequate to see melt-pool-scale phenomena, resulting in the maximum temperatures detected by the camera of roughly 1300 C even though melting is occurring. The squares in Fig. 2 represent temperature data taken from the solid cylinder surface, while the triangles represent data from the cylinder surface covered with powder", " The highlighted area labeled \u2018\u2018Shutter Open\u2019\u2019 illustrates when the camera views the top surface of the build if monitoring of a substantial portion of a build is attempted. Thermal images were then taken for a build of four, 12-mm-tall, 15-mm-diameter cylinders surrounding a representative industry component. For this build, a shutter system was used and images were only taken when the electron beam was off to minimize metallization of the camera viewport.6 A representative thermal image from this build is shown in Fig. 1b. With the understanding that temperatures before spreading and after heating should be similar, the temperature of the solidified cylinders just before powder spreading is used as the surface temperature for the subsequent layer. As with the continuous monitoring, temperatures were extracted by averaging a 12-pixel-diameter area at the center of each cylinder. The results of tracking the surface temperature from 3.5\u2013 8.25 mm of build height are shown in Fig. 3, showing a clear upward trend. The start plate temperature rapidly decreased from a maximum of 780 C at 0 mm build height to 600 C at 3" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0003913_s00773-019-00675-8-Figure1-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0003913_s00773-019-00675-8-Figure1-1.png", "caption": "Fig. 1 Beam element model", "texts": [ " The ship studied here is a modern bulk carrier with a single screw, and its turning performance is first simulated by CFD, as well as the propeller dynamic loads during the turning. Then, the results of CFD simulation is inputted into the following shaft structure analysis. In the calculation of shaft alignment, ship propulsion shafting is often simplified as a multi-support statically indeterminate continuous beam with a variable cross-section subjected to external force and moment. The spatial beam element is used to establish the finite element model for the alignment calculation of ship propulsion shafting, as indicated in Fig.\u00a01. The beam model has two nodes, namely, i and j, and each node has six degrees of freedom, which corresponds to six node forces. However, the force along and the torque around the direction of the beam element need not be considered in the calculation. Thus, eight degrees of freedom is considered for each element to reduce the calculation scale. On the basis of the basic principle of the finite element method, the displacement mode of the element is complete and compatible. Therefore, the displacement and rotation of the node are continuous, which can be expressed as a certain polynomial" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0001329_ls.121-Figure1-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0001329_ls.121-Figure1-1.png", "caption": "Figure 1. The general view of a lubricated contact.", "texts": [ " In the boundary layers (inlet and exit zones), two systems of equations for the main terms of asymptotic expansions for unknown quantities will be obtained. Furthermore, some asymptotic formulas for lubrication fi lm thickness will be derived. The regimes of starved and fully fl ooded lubrication will be analysed. Copyright \u00a9 2010 John Wiley & Sons, Ltd. Lubrication Science 2010; 22:251\u2013289 DOI: 10.1002/ls Let us consider a steady plane isothermal EHL problem for a heavily loaded contact of two moving smooth elastic cylinders (see Figure 1). The gap between the cylinders is assumed to be much smaller than the contact size, which, in turn, is much smaller than the cylinder\u2019s radii. The lubricant is an incompressible fl uid with Newtonian rheology. In this case, the problem equations can be reduced to the following (derivation details can be found in Hamrock2) d dx h dp dx u u h p x p x dp x dx i e e 3 1 2 12 2 0 0 \u03bc \u2212 +\u23a1 \u23a3\u23a2 \u23a4 \u23a6\u23a5 = ( ) = ( ) = =, ( ) , h x h x x R E p t dt p t dt Pe e x t x t x x x x e i e i e ( ) ( ) ln , ( ) ,| |= + \u2212 \u2032 + \u2032 =\u2212 \u2212\u222b 2 2 2 2 \u03c0 \u222b (1) where u1 and u2 are the lower and upper cylinder surface velocities, R\u2032 and E\u2032 are the effective radius and elasticity modulus, 1/R\u2032 = 1/R1 \u00b1 1/R2, R1 and R2 are cylinder\u2019s radii (signs + and \u2212 are chosen in accordance with cylinders\u2019 curvatures), 1/E\u2032 = 1 / E \u20321 + 1 / E \u20322, E \u2032j = Ej/(1 \u2212 v2 j), j = 1, 2, E1, \u03bd1 and E2, \u03bd2 are cylinder\u2019s elastic material Young\u2019s moduli and Poisson\u2019s ratios, respectively, P is the normal force acting along the z-axis and applied to the cylinders, xi and xe are the inlet and exit coordinates of the contact, p(x) and h(x) are pressure and gap, respectively, \u03bc(p) is lubricant viscosity, he is the lubrication fi lm thickness at the exit point of the contact xe, i" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0000903_s1061934811050091-Figure1-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0000903_s1061934811050091-Figure1-1.png", "caption": "Fig. 1. Cyclic voltammograms of a 1 \u00d7 10\u20133 M solution of K3Fe(CN)6 in a 0.1 M solution of KCl obtained on a renewed (CVA 1, 1, \u0394\u0415p = 59 mV) electrode made of a graphite\u2013epoxy composite and on similar electrodes deac tivated for three days in air (CVA 2, \u0394\u0415p = 90 mV); (v = 0.1 V/s).", "texts": [ " The solutions of \u041a3Fe(CN)6 were pre pared by dissolving a precisely weighted portion just before the experiment; oxygen was removed with argon. The dependence of the current of the reduction peak of K3Fe(CN)6 on the ionic strength of the solu tion was studied by direct voltammetry at 24 \u00b1 2\u00b0\u0421. Reversibility of the electrode process The reversibility of the system was estimated by the difference in the potentials of anode and cathode peak \u0394Ep = \u0415a \u2013 \u0415c in the cyclic voltammograms (CVA) of solutions con ( ) ( ) 3 4 6 6 Fe CN Fe CN .\u2212 \u2212 Fe CN( )6 3\u2013/Fe CN( )6 4\u2013. ( ) ( ) 3 4 6 6 Fe CN Fe CN \u2212 \u2212 taining K3Fe(CN)6 only (Fig. 1). The supporting solu tions were 0.1, 0.5, 1.0 M solutions of KCl and a 0.5 M solution of K2SO4. The values of \u0394Ep in the system are known [2, 4\u201311] for these media and different carbon electrodes. The same method was used for the investigation of the influence of the alkali metal cations on the value of \u0394Ep in 1.0 M solutions of LiCl, NaCl, KCl, and CsCl. The value of the reversibility parameter \u0394Ep in the system obtained in the solutions of alkali metal chlorides on a GEC electrode are provided in Table 1", " The value \u0394Ep = 56 mV, which is close to the theoretical one, obtained [8] only ay a carbon paste electrode produced by pressing GP 38 graphite powder without a binding component. The surface of the GEC electrode only partially loses its activity when kept in air, just as it happens with all carbon electrodes. For example, after storing the electrode in air for 20, 40, 120 min, and 3 days. the val ues of \u0394Ep in the CVA of 10\u20133 M solution of K4Fe(CN)6 in a 1 M KCl solution were 61, 62, 68, and 90 mV, respectively (Fig. 1). The mechanical renewal of the electrode by cutting a surface layer directly in the solu tion reduced \u0394Ep to 58 \u00b1 1 mV regardless of the dura tion of its stay in air. In the case of the deactivated GEC 2 electrode, a strong dependence of \u0394Ep on the nature of the alkali metal cation was observed (Table 1, rows 8\u201310); it was many times stronger than the theoretical one. The increase in \u0394Ep in 1 M solutions of KCl, NaCl, and LiCl from 215 to 570 mV (Table 1, rows 8\u201310) indi cates that the rate constant of the electrode process ks decreases" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0000863_ilt-03-2011-0018-Figure1-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0000863_ilt-03-2011-0018-Figure1-1.png", "caption": "Figure 1 Squeeze film configuration of wide parallel plates lubricated with a non-Newtonian Rabinowitsch fluid", "texts": [ " In the present study, the non-Newtonian effects of Rabinowitsch fluids on the squeeze film performances between wide parallel rectangular plates are mainly concerned. By using a small perturbation method, a closedform solution of the squeeze film characteristics is derived for the parallel plates considering the non-Newtonian effects of cubic stresses. Comparing with the Newtonian-lubricant parallel plates, the effects of non-Newtonian cubic-stress flow rheology upon the squeeze film characteristics are presented through the variation of the non-dimensional nonlinear factor accounting for the non-Newtonian pseudoplastic and dilatant effects. Figure 1 shows the squeeze film configuration of wide rectangular parallel plates lubricated with a non-Newtonian Rabinowitsch fluid. The upper plate is approaching the lower one with a squeezing velocity V \u00bc 2dh/dt. It is assumed that the film thickness is thin, the fluid inertia is small, and the body forces are absent. Then the incompressible continuity equation and the momentum equation reduce to the following: \u203au \u203ax \u00fe \u203aw \u203az \u00bc 0 \u00f01\u00de \u203ap \u203ax \u00bc \u203atxz \u203az \u00f02\u00de \u203ap \u203az \u00bc 0 \u00f03\u00de In the Rabinowitsch fluid model, the constitutive equation is expressed as: txz \u00fe kt3xz \u00bc m \u203au \u203az \u00f04\u00de where k denotes a nonlinear factor accounting for the nonNewtonian effects" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0003229_1350650119826440-Figure1-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0003229_1350650119826440-Figure1-1.png", "caption": "Figure 1. EHL line contact geometry used in the analysis.", "texts": [ " First, to investigate by means of full EHL numerical simulations the time-dependent evolution of film thickness and pressure perturbations introduced by moving subsurface inclusions. Second, to verify the rapid, semi-analytical model24 predictions by direct comparison with full numerical simulations. For the sake of simplicity but without loss of generality, the analysis is limited to a single cylindrical inclusion of varying stiffness in a line contact configuration. The EHL contact configuration considered is illustrated in Figure 1. It consists of a rigid cylinder of radius Rx rolling over a heterogeneous elastic solid that contains a single cylindrical inclusion of radius Ri parallel to the overrolling cylinder. Both cylinders are assumed to be infinitely long so that a line contact geometry can be adopted. The rolling entrainment velocity is along the x-axis, whereas the applied normal load F acts along the z-axis. The proposed configuration has been selected only for convenience to the finite element model (ease of meshing, symmetric geometry, etc", " As such, for the case of smooth surfaces treated here the film thickness is given as h \u00bc hc \u00fe \u00f0w\u00fe wi\u00de exp\u00bdi x\u00f0x u1t\u00de exp\u00bdi yy \u00f09\u00de with hc indicating the central film thickness with smooth surfaces and wi the elastic displacements corresponding to the inclusion. As can be seen in equation (9), these displacements are decomposed into individual sinusoidal waves using Fourier transform. The procedure needs to be performed for each time step because wi depends on the applied pressure, which for this purpose follows the smooth Hertz distribution. However, note that the scheme can easily be modified to include roughness.24 It is also important to mention that the semi-analytical solver is a full 3D model, so to solve a 2D problem as the one described by Figure 1 some adaptations are required. For instance, pressure variations along the y direction (direction normal to the paper) should be zero. Thus, a 3D configuration meeting this requirement was set-up. The geometrical domain used was in x 1.3mm, in y 2.5mm, in z 0.821mm, divided in a mesh of 193 385 129 equidistant points in each direction respectively, thus in total 9,585,345 points in the whole domain. The mesh in time has no effect since the present time solution does not require information from a previous time, so the solutions were carried out for the specific time locations" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0000196_2013-01-1907-Figure2-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0000196_2013-01-1907-Figure2-1.png", "caption": "Figure 2. Illustration of the disc-pad interface (Aab) with leading and trailing edges.", "texts": [ " For each pad set with the same pad material composition, is measured under the same representative on-brake conditions, with no resonant growth. The same disc with \u03b6(t) is employed. The measured \u03b6(t) is found to be approximately 20 \u03bcm peak-to-peak. The static compliance of the system is experimentally quantified (using load-deflection measurements) to understand variation in stiffness parameters due to the change loading interfaces within the caliper. Each disc and caliper system is comprised of a disc-pad contact interface, as illustrated in Figure 2 where Aab represents the interface between the (rotating) disc and pad, and through which a point on the disc enters (leading edge) and then leaves (trailing edge) the caliper. Both inboard and outboard pads are assumed to have equivalent disc-pad contact properties, with a center of contact defined at an effective point R on the disc. The rotational speed of the disc is defined as \u03a9 in revolutions per min (rpm). The brake torque T developed at Aab subjected to the boundary conditions of the caliper is defined by both a mean torque and an alternating peak-to-peak torque component, given by (1) For the sake of simplicity, is assumed not to be varying with time, which is appropriate for the constant pressure braking events considered in this study" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0001482_iros.2011.6095088-Figure5-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0001482_iros.2011.6095088-Figure5-1.png", "caption": "Fig. 5. The law of sines can be used to find the distance from the backbone of the flexure to the point where two bending rules intersect, forming an angular section of a cone.", "texts": [ "2 cos (7) # / log 1 , / log 1 / 2/ (8) The dimensionless group / and the shape factor # / have been separated out of (6) because they capture several salient features of the bending energy of the strip. The first feature of note about the shape factor, plotted in Fig. 4, is that it will go to infinity as / approaches either 1 or -1. This divergence can be explained by ascribing some real meaning to /. A way of visualizing the local deformation of a sheet is to see a thin slice of the sheet as lying tangent to a cone, at an angle from the backbone curve and a distance from the vertex of the cone, as illustrated in Fig. 5. The law of sines can be used to find by considering the triangle shown in Fig. 4 for a small displacement \u0394 along the backbone, sin 342 5 \u0394 sin 6 \u0394 7 (9) It is possible to solve for by taking the limit of (9) as \u0394 approaches 0, lim \u0394 cos sin 6 \u0394 7 cos (10) The local energy of the strip at is equal to the energy of the infinitesimal cone over the section tangent to backbone curve. At the vertex of the cone, the local bending energy becomes infinitely large, because the cone\u2019s radius of curvature goes to 0" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0000358_iccse.2010.5593596-Figure2-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0000358_iccse.2010.5593596-Figure2-1.png", "caption": "Figure 2. Trace spline of point", "texts": [ " A correctional Newton-Raphson iterative algorithm is employed in every step of integration. [1-3] Using databases of mechanical parts, joints and constraints, forces and motions, a model of multi rigid bodies can be built according to actual physical parameters of mechanical systems. The elementary process of analysis with ADAMS software is shown in Figure 1. V. EXAMPLES A linkage mechanism is a kind of widely used mechanism. The Main Box of ADAMS provides modeling tool (Icon and ) for building a linkage mechanism. The operation of \u2018Create Trace Spline\u2019 as shown in Figure 2 is used by the designer to hold movement of a trace of every point on the linkage mechanism, and to judge whether it meets the requirements of the design. Besides the simple 4-bar mechanism, more complicated linkage mechanisms can be rapidly built and the advanced structure design can be completed. ADAMS provides the function of \u2018Table of points\u2019 (shown in Figure 3), in which coordinate values of every key point (e.g., link points) are input. With this function, a parametric model will be achieved" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0001374_eeeic.2011.5874721-Figure2-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0001374_eeeic.2011.5874721-Figure2-1.png", "caption": "Figure 2. Schematic representation of dynamic eccentricity.", "texts": [ " In this study, we consider one of the signal processing method, namely, Time-frequency representation (TFR),will be used to study the effect of dynamic air-gap eccentricity on the torque profile of an SRM. Any new idea or result, is a positive step to advance the SRM's diagnosis field. 978-1-4244-8782-0/11/$26.00 \u00a92011 IEEE II. MAGNETOSTATIC ANALYSIS USING FEMM A Dynamic eccentricity occurs when the center of the rotor is not at the center of rotation and the minimum air-gap revolves with the rotor as introduced in Fig. 2. The risk of high levels of dynamic eccentricity is a mechanical contact between the rotor and the stator resulting in considerable damage for the machine. To investigate the effect of dynamic eccentricity on the 6/4 switched reluctance behaviour, the motor is simulated utilizing 2-D finite element analysis by Finite Element Method Magnetics (FEMM) package1. Using finite elements method is a priori appealing for solving complex problems with a better accuracy. Iron material was used in the structure of the stator and rotor cores with the following static B-H curve shown in Appendix" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0000086_amr.423.143-Figure3-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0000086_amr.423.143-Figure3-1.png", "caption": "Figure 3 : Scheme of the paper support", "texts": [ " The same process was used to get unitary fibres from the bundles. The hemp sample which was tested in this study was harvested on the third period of 2009 and grew in Aube which is located in France. It\u2019s referenced as 09ITC N\u00b046 3/8. This kind of hemp was subjected to a higher water portion during the growing. The tension tests showed in this paper can be apply to unitary fibres as well as bundles. The fibres are assembled with a paper support which is pierced in its middle by a rectangular hole showed in Figure 3. For these tests, the length of hole L1 =10mm was used. Tow glue points at both ends of sample keep the fibre on the support. The lengths L2, L3 and L4 were fixed respectively to 60, 6 and 10mm, which allows to set up the support in the tension machine. A micro-tensile test machine Kammrath & Weiss (Fig.4) was used to carry out to put in tension very small samples. The moving velocity of the machine was 1\u00b5m.s -1 . The load cell capability is 50N with an uncertainty of 1/100N. With this kind of tension test machine, the measure errors caused by the rigidity of the machine can be disregard" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0002336_17452759.2017.1325132-Figure12-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0002336_17452759.2017.1325132-Figure12-1.png", "caption": "Figure 12. Moulds and shell with detailed symmetries.", "texts": [ " Graphic dependences of elements\u2019 amount for calculations according to the symmetry of the object. objects while the universal conversion algorithm computes most quantity elements. Apparently, the time for the transformation process with thin layers can be significantly decreased if one uses the combination of stated methods for objects such as a pen in Figure 11(a) with asymmetric, rotary and linear symmetry, Figure 11(b). Moreover, these features are widely applicable for moulds and casting used in different industries when the raster ribbon points\u2019 value is inverted. Figure 12, showing moulds Figure 13. Screenshot for transformation of pen into a ribbon within eight parts. and shell, illustrates that the share of asymmetry is minimal relative to linear and rotary symmetries. Figures 13 and 14 demonstrate screenshots of the transformation for a pen and shell into the spiral coordinate system, where red colour denotes rotary symmetry, green colour denotes linear symmetry and blue colour denotes asymmetry. Simplified, length compressed scale, conversion of these objects shown in Figure 15" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0001195_1.3552319-Figure3-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0001195_1.3552319-Figure3-1.png", "caption": "FIGURE 3: Emissivity calibration system.", "texts": [], "surrounding_texts": [ "The emissivity of the measured material is really important. A dedicated experimental device has been set up to obtain results as precisely as possible. It is compound of an oven which heats the material sample, a thermocouple which records temperatures on the measured face of the sample (supposed homogenous) and a protective tube which ensures a neutral atmosphere (nitrogen) and absorbs parasite radiations. During all these experiments a FLIR camera records the luminance levels emitted by the sample for a given temperature value. By this way it is possible to link the luminance level and the temperature of the face. The following curves (Figure 4) present the 316L and 100Cr6 emissivity recorded with our experimental device, for a ground surface and luminance acquisition made perpendicularly to the emitting surface. After each test the measured surface is controlled to detect any traces of oxidation. 1056 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: 139.184.14.159 On: Tue, 11 Aug 2015 15:34:19" ] }, { "image_filename": "designv11_62_0001059_1.4001772-Figure17-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0001059_1.4001772-Figure17-1.png", "caption": "Fig. 17 Special case for the second order, three-DOF,planar motion of a rigid body when the angular velocity and the angular acceleration of the body is zero", "texts": [ " 2.4 Case: = =0 . When both the angular velocity and angular acceleration of the mobile platform are zero, the IC for acceleration goes to infinity Eq. 16 . The acceleration of point E cannot be computed using the acceleration IC i.e., using Eq. 20 . However, using Eq. 17 , we can see that the acceleration of point E is instantaneously equal to the acceleration of the point of reference P. It means that all the points in the body move with same linear velocity up to the second order, as shown in Fig. 17. Fig. 15 Special case for the second order, three-DOF planar AUGUST 2010, Vol. 2 / 031015-11 ashx?url=/data/journals/jmroa6/27999/ on 03/15/2017 Terms of Use: http://www.asme.org/about-asme/terms-of-use R F m 0 Downloaded Fr eferences 1 Tisius, M., Pryor, M., Kapoor, C., and Tesar, D., 2009, \u201cAn Empirical Ap- proach to Performance Criteria for Manipulation,\u201d ASME J. Mech. Rob., 1 3 , p. 031002. 2 Kapoor, C., and Tesar, D., 2006, \u201cIntegrated Teleoperation and Automation for Nuclear Facility Cleanup,\u201d Ind" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0003065_gtindia2017-4534-Figure1-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0003065_gtindia2017-4534-Figure1-1.png", "caption": "Figure 1. Shaft element: (a) FE discretization (b) Crack orientation", "texts": [ " Finite element formulations of a uniform FG (SS/ZrO2) rotating shaft system has been modeled considering two-nodded Timoshenko beam element with four DOFs (two translational and two rotational DOFs) at each node. In the present study, breathing crack is considered during rotation of the FG shaft to accurately predict the dynamic responses of the rotor system. The shaft internal damping is neglected in the present work. Modeling of cracked FG shaft element The cracked FG shafts with simply supported ends are discretized using FE and shown in Fig. 1(a), where total length L , element length eL and diameter D of the shaft and the depth of crack located at distance cL from the left end of the shaft. Fig. 1(b) shows the any arbitrary crack orientation of the shaft. Fig. 2(a) shows the shaft element subjected to shear forces 1 2 5 6, , and P P P P and bending moments 3 4 7 8, , and P P P P , rotating at speed , v and w are the translation displacements along Y and Z direction, and and are the rotational displacements about Y and Z at a distance s from the left end of a cross-section of the shaft. Fig. 2(b) shows the circular section of the cracked shaft having crack orientation o180 with crack half-width b " ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0001661_tasc.2011.2109050-Figure1-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0001661_tasc.2011.2109050-Figure1-1.png", "caption": "Fig. 1. Conveyor system with electromagnets.", "texts": [ " This is because magnetic fluxes move to other positions due to flux creep effect. In order to solve the problem, pulse-field magnetization is applied to the system to improve the levitation height and the dynamics of the levitated conveyor [5]. Then, pulse-field magnetization is not sufficient due to installed small electromagnets. Then, in this paper we install bigger electromagnets above the levitated conveyor and apply pulse-field magnetization to the conveyor from the upper superconductor surface. Fig. 1 shows a total illustration of the magnetically levitated conveyor system. The system consists of a levitated conveyor, a pair of magnetic rails, two electromagnets for pulse-field magnetization and driving the levitated conveyor, and power supplies. Manuscript received August 02, 2010; accepted January 16, 2011. Date of publication February 28, 2011; date of current version May 27, 2011. The authors are with the Department of Applied Science for Integrated System Engineering, Graduate School of Engineering, Kyushu Institute of Technology, Fukuoka 804-8550, Japan (e-mail: komori_mk@yahoo" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0000261_iros.2010.5651258-Figure3-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0000261_iros.2010.5651258-Figure3-1.png", "caption": "Fig. 3. Model of planar telescopic-legged biped robot with feet", "texts": [ " Human foot is, however, formed of inverse shape and this implies that it can drive the stance leg backward only. In other words, human foot is not suitable to drive the body forward. Based on the observations, in this paper, we extend our 978-1-4244-6676-4/10/$25.00 \u00a92010 IEEE 4477 gait generation method to a planar telescopic-legged biped robot with feet incorporating a brake spring and numerically investigate the gait descriptors, especially the behavior of ZMP. Through numerical simulations, we discuss the anterior-posterior asymmetry of human foot from the ZMP point of view. 1) Dynamic equation: Fig. 3 shows the planar telescopiclegged biped model. Let the stance leg and swing leg be Leg 1 and Leg 2, and qi = [ xi zi \u03b8i bi ]T be the generalized coordinate vector for Leg i. The corresponding dynamic equation becomes M i(qi)q\u0308i + hi(qi, q\u0307i) = 04\u00d71, (1) and we then augment them by adding the holonomic constraint forces and control inputs. The two legs are connected at the hip joint. We also assume that the foot mass and thin are ignorable or the foot dynamics does not affect the walking motion at all" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0001660_isam.2011.5942364-Figure11-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0001660_isam.2011.5942364-Figure11-1.png", "caption": "Figure 11. Induced deformation of the thin-walled hollow shaft", "texts": [ "63 \u00b5m is measured at the beginning up to a cutting length of lc = 966 m. After the final cut at a cutting length of lc = 1208 m, the wall thickness of the hollow shaft is reduced to t = 1.5 mm. The average surface roughness increased to Rz = 3.19 \u00b5m. 4) Roundness of the hollow shaft The longitudinal turning operations were conducted until a wall thickness of t = 1.5 mm was reached. Measuring the roundness after the final cut shows a strong influence of the chosen clamping situation on the contour of the workpiece (Fig. 11). The forces of the clamping jaws at the spindle deformed the workpiece in radial direction. At the center point the roundness was affected as well, but less in comparison to the opposite end face. Instead of an ideal round form, an oval form is recognized. In the middle of the workpiece (axially), the lowest roundness values of 39.0 \u00b5m were measured, whereas at the end faces roundness values of 113.2 \u00b5m, and 977.5 \u00b5m, respectively, have been recorded. The material of the thin-walled hollow shaft may offer a low Young\u2019s modulus like the undeformed raw material" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0000765_gt2010-22852-Figure6-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0000765_gt2010-22852-Figure6-1.png", "caption": "Figure 6. Solid model of turboalternator simulator system", "texts": [ " 5, the alternator rotor bend mode would occur at speeds above 182,000 rpm for the selected design when operating at 116,000 rpm. Additionally, the rotor rigid modes and coupling modes all occur below 34,000 or at approximately 30% of operating speed. This design approach therefore provides substantial margin both for rigid and bending critical speeds. In order to demonstrate dynamic stability and full speed operation of the system at least risk and cost a lowtemperature, high-speed simulator as shown in Fig. 6 was designed. This simulator was designed to have the same dynamics as the completed turboalternator but would run at ambient temperature conditions as opposed to elevated temperatures. Since test operation was to be conducted at room temperature and initial tests were to have a nonmagnetic rotor installed, an impulse turbine was designed and sized to have inertial properties similar to the projected high temperature turbine. The assembled test rig hardware is shown in Fig. 7. This section of the paper will describe the test rig setups used for rotor-bearing system dynamic testing of the turbine drive and coupled rotor-bearing systems" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0003045_eptc.2017.8277474-Figure1-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0003045_eptc.2017.8277474-Figure1-1.png", "caption": "Fig. 1. The model structure and gantry motion", "texts": [ " High precision motion platform is widely used in these equipment, so its motion precision directly affects the performance of equipment, then affects the quality of industrial products and production efficiency. Dual-linear motor driven motion platform refers to drive two motors to complete one direction of movement. It is a typical H-drive structure of high precision motion platform, which is widely used in electronic packaging equipment. [1] The motion platform is equipped with two linear motors at both ends of the Y direction to drive the beam in the X direction, as shown in Fig. 1. This structure has some advantages, including large driving force, high load and high bandwidth. Although both sides of the Y direction have the same driving and executing structure, during the process of the motor movement, the synchronous motion of two linear motors will be affected inevitably by the existence of unbalanced forces, external disturbances and mechanical parameter errors. Therefore, when this motion platform is applied in high precision equipment, higher precision synchronization control is required" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0002681_s40010-017-0396-z-Figure4-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0002681_s40010-017-0396-z-Figure4-1.png", "caption": "Fig. 4 E. Coli bacteria while it swimming and tumbling", "texts": [ " coli bacteria consist of a plasma membrane, cell wall and a capsule. The cell is about 1 lm diameter and 2 lm in length and the weight is about 1 picogram. The E. coli bacterium has a control system that enables it to search for food and avoid noxious environments [18]. The control mechanism of E. coli bacteria is described by following 4 steps. (1) Swimming and tumbling The E. coli bacterium has set of rigid flagella that enable it to swim. It can move in two different ways: swim and tumble. The flagellum can rotate clockwise and counterclockwise as shown in Fig. 4. If the flagellum rotates clockwise then the bacteria tumble and if it is anticlockwise then the bacteria swim. The Bacteria can swim up to a maximum no of Ns steps. (2) Chemotaxis The chemotaxis step is the combination of swimming and tumbling for an example if an E. coli bacterium is in some substance that is neutral then the flagella simultaneously alternates between counter clockwise and clockwise, thus the bacteria will alternately tumble and swim. The maximum no of the swim within a chemotaxis is Ns and when the swim steps stop the tumble action takes place" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0002480_978-3-319-60399-5_7-Figure3-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0002480_978-3-319-60399-5_7-Figure3-1.png", "caption": "Fig. 3 The combined scheme of the operating and machine-tool meshing for the centre of the bearing contact arranged on the middle plane of the gear", "texts": [ " Modification of the gearwheel tooth flank along the tooth length (called longitudinal localization) is aimed at compensation for these errors. The most common methods for longitudinal localization of the contact for worm gears of the 6th\u201312th degrees of accuracy is the application of hobs having an increased diameter [5, 12\u201315]. And it is important to assign such a diameter of the hob and manufacturing tolerances that allow for compensating the manufacturing errors of gearbox parts and for obtaining the gear with a specific degree of accuracy and the required quality parameters. Figure 3 shows the combined scheme of meshing of the orthogonal worm gear and the machining of gearwheel teeth by a hob with an increased diameter. The diameter d0 of the pitch cylinder of the hob is greater than the diameter d1 of the pitch cylinder of the worm, however, these cylinders make contact with each other at the pitch point F. Let us consider the version of the arrangement of the nominal centre of the bearing contact (the point \u041a) on the middle plane of the gear. The left side of this figure shows the section of the gear by a plane perpendicular to the worm axis and passing through the gearwheel axis", " The normal line to any helical surface is skewed with its axis at a distance equal to the radius rb of the base cylinder of the equivalent involute worm: rb \u00bc 0:5mx1z1=\u00f0tan2ax1 \u00fe tan2c1\u00de0:5 \u00f014\u00de That is why the straight line passing through the point \u041a tangential to the circle with the radius rb represents a projection of the common normal line at the design point. Simultaneously, this tangent line shows the direction of displacement of the centre of the bearing contact along the tooth flank at worm rotation. In the picture on the right side of Fig. 3, the axis of the worm O1\u2013O1 is arranged horizontally, and the axis of the hob O0\u2013O0 is rotated with respect to it at an angle Dc. The shaded area is the section of the worm thread by the plane A\u2013A tangential to the pitch cylinder of the worm. This section is in contact with the section of the thread of the hob-generating surface by the same plane at the design point \u041a. The straight line \u041aC0 represents the common normal line to longitudinal profiles of threads. This line makes up the angle c1 (equal to the pitch lead angle of the worm) with the worm axis O1\u2013O1 and the angle c0 (equal to the pitch lead angle of the hob-generating surface) with the projection of the hob axis", " Values R1, R0 are determined by the expressions R1 \u00bc 0:5d1=\u00f0tan ax1cos3c1\u00de;R0 \u00bc 0:5d0=\u00f0tan ax0cos3c0\u00de \u00f015\u00de where ax1, ax0 are axial pressure angles of the worm thread and the generating surface of the hob, respectively. In order to arrange the design point \u041a on the pitch cylinder of the worm at the middle plane of the gear, the hob axis should be rotated towards increasing the tooth helix angle at an angle Dc equal to Dc \u00bc c1 c0: \u00f016\u00de When the worm is rotating in the direction shown by the arrow x, the active part is that side of the profile which passes through the point \u041a; and the \u201cEnter\u201d and \u201cExit\u201d points of the worm out of the meshing are shown on the right side of Fig. 3. The arrow of the convexity (curvature) d1 of the surface of the worm thread along the gearwheel tooth in the considered section is determined to be d1 \u00bc b22=\u00f08R1cos2c1\u00de : \u00f017\u00de As for the generating surface of the hob having an increased diameter due to the increase of the curvature radius of the thread by the value DR = R0 \u2212 R1, the arrow of the convexity in the same section will be decreased by the value Dd \u00bc b22DR=\u00f08R2 1cos 2c1\u00de; \u00f018\u00de where the minus sign shows that deviations of Dd and DR have opposite signs" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0002302_b978-0-12-803581-8.10351-0-Figure35-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0002302_b978-0-12-803581-8.10351-0-Figure35-1.png", "caption": "Fig. 35 Roll hoop full enclosed by both halves of the main bulkhead.", "texts": [ " The inner flange follows the roll hoop contour with an offset to allow the front half of the bulkhead shell inner flange to fit between the hoop and back half, thus bonding the back and the front together forming a single hollow bulkhead with the steel tube structure inside. The front half is configured in a u-shape to combine the two halves as stated above with the inner flange, and the outer bonding flange is then tucked between the hoop and the chassis. The front vertical wall is the same as the back. This allows for a full 360 degree bonding of the roll hoop, and provides sufficient bonding to the chassis as shown in Fig. 35 and the full assembly of the roll hoop into the bulkhead shown in Fig. 36. The main bulkhead reinforces the chassis cross-section from distorting under load and, in addition to providing rollover protection, supports localized loads imposed by the suspension. The inner bonding flanges of the composite halves closely follow the roll hoop contour to allow the largest opening possible for the driver in the chassis. Following the roll hoop inner perimeter allows for only a 0.05000 loss in the perimeter due to the composite bulkhead thickness and required adhesive" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0002758_icems.2017.8055934-Figure1-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0002758_icems.2017.8055934-Figure1-1.png", "caption": "Fig. 1. Cross section of prototype machine.", "texts": [ " Then, a no-slot model is used for further investigation. It is found that the UMF with rotor eccentricity is mainly determined by the amplitude of open circuit airgap flux density and its sensitivity to the airgap length. Moreover, it shows that the UMF varies slightly with pole numbers in machines having parallel and radial magnetizations, but much more significantly when Halbach array is employed, especially when the magnet is relative thick. The cross section of a prototype machine is shown in Fig. 1, and the detailed parameters are listed in Table I. It should be noticed that only the 12-slot/10-pole machine is shown, while the machine with other pole number will also be investigated which shares the same stator and rotor specifications. In addition, machines having different magnetizations, i.e. parallel, radial and Halbach magnetizations, are considered and the corresponding open circuit flux density distributions are shown in Fig. 2. The magnet thickness is chosen as 9mm to show the characteristics of different magnetizations clearly" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0000458_amr.591-593.1879-Figure1-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0000458_amr.591-593.1879-Figure1-1.png", "caption": "Fig. 1 Modeling unbalance force in models. (a) Equivalent unbalance force and its components. (b) Unbalance modeling method in 3D solid element rotor models.", "texts": [ " Modal and unbalance response analysis is performed on the rotor-bearing system with the dynamic support stiffness considered. Concluding remarks are given in section 4. Modeling Method. ANSYS is one of the commercial finite element softwares which permit modeling rotors using 3D solid elements including gyroscopic effects. For unbalance response analysis in ANSYS, the unbalance effects in 3D solid rotor model is still modeled using equivalent unbalance force similar to what is done in beam element model. The equivalent unbalance force as depicted in Fig.1a can be defined as a static force as shown in table 1 where x is the assumed spin axis [8]. In table 1, F0 is the amplitude of the force. For unbalance, the amplitude is equal to the mass times the distance of the unbalance mass to the spin axis (me in Fig. 1a). \u03b1 is the phase of the force, needed only when several such forces, each with a different relative phase, are defined. For unbalance forces, multiplication of the amplitude of the static forces by the square of the spin velocity (w 2 in Fig. 1a) is unnecessary. ANSYS performs the calculation automatically at each frequency step. In beam element model, as mentioned in prior section, the unbalance force is directly applied to the node at the right axial location. But in 3D solid element model, one single node is not enough to describe the unbalance force rotating around the spinning axis. In order to accurately describe the unbalance force in 3D solid model, an equivalent modeling method is presented in this paper. One circle of nodes of the model disk where unbalance is assumed is selected and four nodes are used to apply the unbalance force, two nodes in each direction as depicted in Fig. 1b. The resultant force in each direction is twice the amplitude of the component force applied on the nodes and the resultant force in each direction is the same as the corresponding force in table 1. In this way the equivalent unbalance force acts as it is applied on the center node and rotates around the spinning axis. Other vector synthesis effects, such as moment caused by component tilt, can also be modeled with the method presented here. Numerical Example. A simple example is considered to illustrate the validity and accuracy of the unbalance equivalent modeling method in 3D solid rotor model" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0001363_rast.2011.5966982-Figure9-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0001363_rast.2011.5966982-Figure9-1.png", "caption": "Figure 9. The vertex P3 moving in the circle with radius r0.", "texts": [ " The projections on x-y plane of points A and B can be written as 2 2sin( ) (29)= + \u03b1A O Ax x t 2 2cos( ) (30)= + \u03b1A O Ay y t 2 2sin( ) (31)= + \u03b1B BOx x t 2 2cos( ) (32)= + \u03b1B BOy y t xA , xB and yA , yB are the solutions of the coordinates (p2x) and (p2y), respectively while zA and ,zB are the solutions of p2z. Each solution of (p2x, p2y, p2z) is accepted for the vertex P2, if the associated constraints are satisfied. 3) The Determination of Position of Vertex P3: Given the coordinates of vertices P1 and P2 calculated above, the geometric relations among P1, P2, and P3 are utilized to figure out the coordinates (p3x, p3y, p3z) of vertex P3. For a fixed P1 and P2, P3 moves in a circle centered at the point O\u0131 with the radius r0, shown as in Fig. 9. The points on the circle with the radius r0, where P3 moves are utilized to determine the leg lengths of L5 and L6 through inverse kinematics. Providing that the determined leg lengths satisfy the constraints of the leg lengths and joints, the points are included to the solution set of the vertex P3. In order to determine vertex P3, the coordinate frame O \u0131(x\u0131 y\u0131 z\u0131) is defined. The origin of the coordinate frame O\u0131(x\u0131 y\u0131 z\u0131 ) is located at the center of the circle with the radius r0. while y\u0131 axis passes through the line P1P2 and x\u0131 axis lying parallel to the x-y plane" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0003970_s11370-019-00294-7-Figure9-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0003970_s11370-019-00294-7-Figure9-1.png", "caption": "Fig. 9 Parallel RRR-to-serial manipulator a a 3-RRR parallel manipulator and b its derived serial manipulator", "texts": [ " (41)T\u0302 = \ufffd s\u0302 23 s\u0302T 23 (S\u03022)1 s\u0302 31 s\u0302T 31 (S\u03022)2 s\u0302 12 s\u0302T 12 (S\u03022)3 \ufffd\u23a1\u23a2\u23a2\u23a3 v1 v2 v3 \u23a4\u23a5\u23a5\u23a6 , 1 3 and inserting (42) into (41) yields where 1, 2 and 3 can be regarded as the joint angular speeds of the three revolute joints of 3-RRR planar serial manipulator generating the same end-effector twist as shown in Fig.\u00a08b. It is noted that the three columns of forward Jacobian of serial manipulator in (43) are lines expressed in axis coordinates. Therefore, it is remarked that the Jacobian of parallel manipulator in (39) composed of three lines expressed by ray coordinates is converted into lines in axis coordinates. Another common 3-RRR parallel manipulator has been shown (with active revolute joints only) in Fig.\u00a09a. The instantaneous first-order kinematics of the platform, which is common in three chains, can be represented as Here, we assume that the first joint of each chain is active joint. Applying reciprocal screws to each chain, we have (42) \ud835\udf141 = v1 S\u0302T 23 (s\u03022)1 \ud835\udf142 = v2 S\u0302T 31 (s\u03022)2 \ud835\udf143 = v3 S\u0302T 12 (s\u03022)3 , (43)T\u0302 = \ufffd S\u030223 S\u030231 S\u030212 \ufffd\u23a1\u23a2\u23a2\u23a3 \ud835\udf141 \ud835\udf142 \ud835\udf143 \u23a4\u23a5\u23a5\u23a6 , (44)T\u0302 = \ud835\udf141i(S\u03021)i + \ud835\udf142i(S\u03022)i + \ud835\udf143i(S\u03023)i, where ( 1)1, ( 1)2 and ( 1)3 are the angular velocities acting about the axis coordinates (S\u03021)1, (S\u03021)2 and (S\u03021)3, , respectively", " (s\u030223) T 2 T\u0302 = (\ud835\udf141)2(S\u03021) T 2 (s\u0302 23 )2 (s\u030223) T 3 T\u0302 = (\ud835\udf141)3(S\u03021) T 3 (s\u0302 23 )3 (46) (\ud835\udf141)1 = (s\u030223) T 1 T\u0302 (s\u030223) T 1 (S\u03021)1 (\ud835\udf141)2 = (s\u030223) T 2 T\u0302 (s\u030223) T 2 (S\u03021)2 (\ud835\udf141)3 = (s\u030223) T 3 T\u0302 (s\u030223) T 3 (S\u03021)3 (47) \u23a1\u23a2\u23a2\u23a2\u23a3 (\ud835\udf141)1 (\ud835\udf141)2 (\ud835\udf141)3 \u23a4\u23a5\u23a5\u23a5\u23a6 = \u23a1 \u23a2\u23a2\u23a2\u23a2\u23a2\u23a2\u23a2\u23a2\u23a2\u23a3 (s\u0302T 23 )1 (s\u0302T 23 )1(S\u03021)1 (s\u0302T 23 )2 (s\u0302T 23 )2(S\u03021)2 (s\u0302T 23 )3 (s\u0302T 23 )3(S\u03021)3 \u23a4 \u23a5\u23a5\u23a5\u23a5\u23a5\u23a5\u23a5\u23a5\u23a5\u23a6 T\u0302 . (48) (v23)1 = (s\u030223) T 1 T\u0302 (v23)2 = (s\u030223) T 2 T\u0302 (v23)3 = (s\u030223) T 3 T\u0302 (49) \u23a1\u23a2\u23a2\u23a2\u23a3 (v23)1 (v23)2 (v23)3 \u23a4\u23a5\u23a5\u23a5\u23a6 = \u23a1\u23a2\u23a2\u23a2\u23a3 (s\u0302T 23 )1 (s\u0302T 23 )2 (s\u0302T 23 )3 \u23a4\u23a5\u23a5\u23a5\u23a6 T\u0302 . (50)T\u0302 = (jT )\u22121v. 1 3 Putting the value for the inverse of jT by using \u201cAppendix\u201d, we have which is analogous to the twist generated by the angular velocities of 3-RRR serial manipulator as where 1, 2, and \u03c93 are the three joint angular velocities of serial manipulator as shown in Fig.\u00a09b and they are related as It is noted that the three columns of Jacobian of serial manipulator in (52) are lines expressed in axis coordinates. Therefore, it is remarked that the Jacobian of parallel manipulator in (47) composed of three lines expressed by ray coordinates is converted into axis coordinates. The Jacobian matrix is an inherent property of a manipulator that effectively highlights the interconnected kinematics of serial and parallel architectures, as described in Sect.\u00a03. Based on the characteristics of common actuation wrench between the two dual structures, these can be evaluated for their performance" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0003985_icamechs.2019.8861668-Figure8-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0003985_icamechs.2019.8861668-Figure8-1.png", "caption": "Fig. 8. Mechanism of the leg", "texts": [ " On the other hand, it should be flexible to passively avoid obstacles when it is not supporting the weight of the body. In order for the leg to meet both requirements, we propose a switching mechanism for the stiffness of the leg as shown in Fig. 7. The proposed leg has a passive joint and is composed of layers of aluminum plates and rubber sheets. The joint will move freely when the vertical force is not applied on it , and this will allow to passively avoid obstacles by rotating the joint using reactive forces from the obstacles as shown in Fig. 8. However,, when a vertical force is applied to support the weight of the robot the rubbers of the joint engages, and the joint becomes locked. Therefore, the stiffness of the leg becomes high and the leg can transmit the force from the motor to the ground as a driving force. The advantage of the proposed mechanism in comparison to the conventional mechanism is that the joint can be locked at any angle as shown Fig. 9. Thus, it can produce a driving force when moving through a narrow space. The mechanism of the leg when flexed is given in Fig" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0002349_1.4983239-Figure10-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0002349_1.4983239-Figure10-1.png", "caption": "FIG. 10. Surface roughness measurement positions.", "texts": [ "12 The results show a decreasing fatigue life with the increasing surface roughness. A low surface roughness is desired to reduce the need of final machining of as-build parts which shall be used as structural parts under cyclic loadings. The evaluation of the surface roughness has been recorded with a Keyence VK-8710 laser scanning microscope according to ISO 4287:1997. The surface roughness had been measured in four areas per manufactured wall. The areas were on the same position on every wall according to the sketch in Fig. 10. The different positions on the wall were chosen to get uniformly distributed measured values. In each of the four areas, three line roughness measurements had been made. In Fig. 11, an J. Laser Appl., Vol. 29, No. 2, May 2017 M\u20acoller et al. 022308-7 example of a measured area with a measured roughness line is shown. The mean value of the measurements for the Rz and the Ra values is plotted in Fig. 12. The surface roughness of the manufactured walls has been evaluated using the introduced laser scanning microscope" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0000666_humanoids.2012.6651523-Figure5-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0000666_humanoids.2012.6651523-Figure5-1.png", "caption": "Fig. 5. Robot model.", "texts": [ " It should be noted that while we describe our solution for the specific example in this paper, our approach is more general and geared towards larger problems of determining the optimal muscle arrangement for desired robot motions. For example, we can easily imagine that our method can determine the optimal muscle arrangement of a robot for multiple motions by using multi-objective optimization methods [13]. III. ROBOT MOTION MODEL AND PAM ACTUATOR MODEL A. Robot Motion Model The robot model is shown in Fig. 5. It was modeled as a four-link robot in the sagittal plane (the x-y plane). The generalized coordinates are defined as the position of the toe ( ) , the angle and the joint angle vector ( ), where is the attitude of the link 1 (foot), and , and are the angles of joint 1 (ankle joint), joint 2 (knee joint) and joint 3 (hip joint) respectively. The robot motion is divided into the following three phases by contact with the ground (phase transitions: A B C). Phase A: The toe and heel are on the ground" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0002441_978-3-319-60867-9_38-Figure2-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0002441_978-3-319-60867-9_38-Figure2-1.png", "caption": "Fig. 2. The loop structure of the robot: (a) joint-and-loop graph; (b) one limb with rotational input.", "texts": [ " The angle between ui and x-axis is represented by \u03bci, \u03bc1 = \u03bc2 = 0, \u03bc3 = \u03bc4 = 2\u03c0/3, \u03bc5 = \u03bc6 = \u22122\u03c0/3. Moreover, unit vectors vi and wi are parallel to the segments AiBi and BiCi, respectively, namely, the unit vectors along the active link and parallelogram in the ith limb. Here and after, vectors i, j and k represent the unit vectors of x-, y- and z-axis, respectively. In the following study, the vertical offsets in the platforms are supposed to be zero for convenience, as they do not affect the planar motion of the end-effector. Figure 2(a) depicts the joint-and-loop graph of the robot, wherein the gray and white boxes represent the actuated and passive joints, respectively. With the Group Theory, the kinematic bond Li of the RUU chain in ith limb, as displayed in Fig. 2(b), is the product as below: Li = R(Ai) \u00b7 R(Bi) \u00b7 T (ni) \u00b7 R(Ci) = X (ui) (1) and the kinematic bonds of the joint 1st and 6th limbs L16 is L16 = T \u00b7 R(N, k) = X (k) (2) Similarly, the kinematic bonds of the 2nd and 3rd limbs L23 is X (k). Subsequently, the intersection of the subgroups for platform 1 results in a Scho\u0308nflies subgroup X (k), namely, the Scho\u0308nflies motion. On the other hand, the kinematic bonds L45 of the closed loop A4\u2013B4\u2013C4\u2013 C5\u2013B5\u2013A5 is L45 = T \u00b7 S (N) (3) and the subgroups of the coupler in the closed loop six-bar linkage is X (k)\u00b7G (k), where G (k) is the planar motion subgroup" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0003666_iemdc.2019.8785290-Figure3-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0003666_iemdc.2019.8785290-Figure3-1.png", "caption": "Fig. 3. Nomenclature around Schwarz-Christoffel mappings", "texts": [ "4404 description non-alloy steel stainless steel structure ferritic austenitic Elastic (Young\u2019s, 190 GPa 200 GPa Tensile) Modulus relative price 9.5 % 100 % 978-1-5386-9350-6/19/$31.00 \u00a92019 IEEE 1824 The Schwarz-Christoffel (SC) mapping maps an infinite half plane (canonical domain) to another geometry (physical domain), conserving angles locally [13]. Conformal mappings such as the Schwarz-Christoffel mapping allow twodimensional field problems to be solved in an approachable geometry and their solution to be transformed into the desired, more complex geometry. Fig. 3 shows the Schwarz-Christoffel mapping being described by a polygon. The polygon\u2019s vertices are given as complex numbers z = x + j \u00b7 y. The mapping function z = f(w) = K \u00b7 \u222b (w \u2212 w1)\u03b11/\u03c0\u22121 \u00b7 (w \u2212 w2)\u03b12/\u03c0\u22121 . . . (w \u2212 wn)\u03b1n/\u03c0\u22121 dw (1) depends on the angles at the polygon\u2019s vertices \u03b1i shown in fig. 3 and the points in the canonical domain corresponding to the the polygon\u2019s vertices wi. Assembling the mapping function is not difficult, however, solving the integral and finding the prevertices wi for anything but the most basic geometries is. Reppe [14] introduced the idea of integrating the mapping function numerically, allowing for much more complex geometries. The prevertices wi can be obtained by making use of the fact that the integral between two prevertices must equal the corresponding polygon side length\u2223\u2223\u2223\u2223\u2223 \u222b wi+1 wi f \u2032 dw \u2223\u2223\u2223\u2223\u2223 = \u2223\u2223zi+1 \u2212 zi \u2223\u2223 " ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0001212_tia.2010.2070052-Figure3-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0001212_tia.2010.2070052-Figure3-1.png", "caption": "Fig. 3. Connection methods. (a) Differential connection. (b) Cumulative connection.", "texts": [ " Therefore, to reduce shaft length, the integration of the bearingless motor and the displacement sensor is an important subject. Fig. 2 shows an arrangement of search coils for rotor displacement estimation. Search coils A and B are set up in the stator teeth located in the x-axis negative direction. Search coils C and D are set up in the stator teeth in the x-axis positive direction. The number of turns of search coils is all the same. The search coil A and B are connected and referred as search coil AB. The search coils C and D are also connected and referred as search coil CD. Fig. 3 shows connection methods. Fig. 3(a) shows a method of connecting search coils in opposite magnetomotive force (MMF) direction. This connection method is referred as a differential connection. Fig. 3(b) shows a method of connecting search coils in parallel MMF direction. This connection method is referred as a cumulative connection. The principle of rotor displacement estimation is based on inductance variation caused by rotor displacement. When the rotor is displaced in the x-direction, then the gap between the stator and the rotor is increased at search coil AB. Magnetic resistance is increased, and the self-inductance of search coil AB is decreased. Therefore, it is possible to estimate the rotor displacement by detecting the self-inductance variations" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0001128_20100906-5-jp-2022.00066-Figure1-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0001128_20100906-5-jp-2022.00066-Figure1-1.png", "caption": "Fig. 1 Interaction field around the agent", "texts": [ " Namely, approach is selected when the neighbour is distant, repulsion is selected when the neighbour is near at hand, and parallel orientation is selected when the neighbour is in an intermediate distance. These distance-reaction rules function to avoid straying away from neighbours in \"approach\", to avoid collision with neighbours in \"repulsion\", and to move together with neighbours to a common destination in \"parallel orientation\". To incorporate these three rules into the air vehicle model, a spherical interaction field is set around the agent as shown in Fig. 1, which is divided into three sub-fields, i.e. approach, parallel orientation, and repulsion fields, and the agent selects one reaction from above-mentioned three reactions in response to the position of a neighbour in the interaction field as shown in Fig. 2. The vectors app, para, and repul, in Fig. 2 denote the body direction vectors of the agent for the approach, parallel orientation, and repulsion, respectively, where app and para point to the direction of the neighbour rj and to the direction in parallel with the body direction of the neighbour j, respectively; repul points to the direction which is the summation of the counter direction of the neighbour -rj and the body direction of the agent i with the weight coefficients c1 and c2, respectively, to avoid a collision with the neighbour (here c1=c2=1 is used)", " When the number of neighbours in the interaction field Nb exceeds this limit (Nb>Nb,max), Nb,max neighbours are selected as targets of interaction. This selection is based on the priority of direction where the agent selects neighbours near the specified direction with higher priority than neighbours distant from this direction. This priority (hereafter denoted as \u201cinteraction directivity\u201d) reflects the property of detection sensor, e.g. microwave or infrared sensor, which usually has the strong sensitivity to the direction where the sensor faces. The vector in Fig. 1 denotes this direction of priority, which usually orients front to detect front neighbours, but can be changed by altering the facing direction of sensor. In addition, the off-sensing area (= blind region) is set at the rear of the agent because sensors are generally set around the head and thus the agent is subject to weak sensitivity in rear. The size of blind region is given by the angle in Fig. 1. If the agent has rear sensors, becomes zero, and if not, takes positive value depending on the difficulty of sensing rear objects. The motion of air vehicle is calculated by using following three dimensional Newton\u2019s equation of motion: for translation, ,coscos ,sincos ,sin a a a ZmgQUPVWm YmgPWRUVm XmgRVQWUm (1) for rotation, , , , 22 NQRIPQIIRIPI MRPIRPIIQI LPQIQRIIRIPI xzxxyyzzxz xzzzxxyy xzyyzzxzxx (2) for the Euler angle of the body, ,seccossecsin ,sincos ,tancostansin RQ RQ RQP (3) where m is the mass of the air vehicle and g is the gravity acceleration; (U, V, W) and (P, Q, R) are the velocity and the angular velocity of the vehicle, respectively, determined on the body-fixed coordinates (Xb, Yb, Zb); () are the Euler angle of the body determined on the ground-fixed coordinates (Xg, Yg, Zg)", "4 Snapshot of calculated motion with other fish around a common circular path to form a ring or cylinder (Parrish & Edelstein-Keshet, 1999; Parrish et al., 2002). When a fish school conducts a straight motion, each fish predominantly follows its front neighbours to avoid straying away from them, whereas in mill formation, each fish follows neighbours that are slightly inward because they are moving on the circular path, indicating the difference in direction of targets to follow between straight motion and mill formation. The directivity of interaction, which is given as the vector in Fig. 1, can therefore be used to mimic these two types of sensing as follows. The direction of vector is usually oriented forward for conducting a straight motion, but can be oriented inwardly with respect to the moving path for loitering. The size of attraction field is also modified to promote the approach motion of agent because the following motion of individuals to foregoing neighbours is the most prevailing behaviour in the mill formation. The size of attraction field is then enlarged by reducing the size of parallel orientation field Rp (Rp=10BL)", " This means the mutual interaction parameter which controls the inter-agent distance or relative position of agents in the group, e.g. the shape of the interaction field, might be a possible internal factor for the change of group shape. The position of neighbours in the field depends on the changes in the geometry of interaction field. In fact, the distance between an agent and its neighbour is subject to the balance between attraction and repulsion similar to the mass-spring system. Therefore, a spherical interaction field (Fig. 1) leads to neighbours being symmetrically distributed around the agent, and in turn when the field shape is modified, the neighbours\u2019 positions will be rearranged asymmetrically, thus resulting in deformation of the group shape. Simulation results confirm this assumption (Fig. 8). The original spherical interaction field yields spherically distributed agents (Fig. 8A). When the interaction field is expanded along either the Yb-axis or the Zb-axis (Figs. 8B or 8C, respectively), however, the distribution of agents expands laterally (Fig" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0000663_robot.2010.5509639-Figure3-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0000663_robot.2010.5509639-Figure3-1.png", "caption": "Fig. 3. Results of Experiment 1, in a world with no obstacles, showing the mean number of resources transported by robots with different populations and field of view configuration. Variance is < 10% for each point.", "texts": [ " Each trial runs for 30 minutes, and the total number of resources delivered at the end of the trial is our performance metric. 10 trials are performed for each of a range of settings of radius df , reception angle \u03b1f , and population size. LOST is deterministic but the local obstacle avoidance and searching is stochastic (for robustness), hence the need for repeated trials. Experiment 1 examined all permutations of df = [1.5, 2.0, 2.5] meters, \u03b1f = [10, 20, . . . 360] degrees, population P = [10, 20, . . . 100] robots. The results of the first experiment are summarized in Figure 3, with mean performance over 10 repeated trials plotted for each [df , \u03b1f , P ] configuration. Error bars are ommitted for clarity: the variance is < 12% in 80% of experiments. The FOV range parameter df appears to have relatively little effect on the performance, but the FOV angle \u03b1f appears to have an important effect. The results show that, with a constant df , a small team of 10 robots has about the same performance for any \u03b1f above 90 degrees. As the population size increases, the performance is better for smaller \u03b1f , until a lower bound is reached" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0003514_978-3-030-20131-9_324-Figure32-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0003514_978-3-030-20131-9_324-Figure32-1.png", "caption": "Fig. 32. Schematic diagram of the process of single-line straight-through weaving beaded cooling pad.", "texts": [], "surrounding_texts": [ "As the alternate threading actions by the hand-weaving method mentioned above is very cumbersome, results in more complicated and low-efficiency to achieve the weaving works on the machine. In order to make the weaving of the beaded cool pad easy to implement on the machine, it is necessary to decompose and transform the weaving route extracted from the hand-weaving process to find a new weaving method. 3.1 Warp and Weft Automatic Weaving Method The warp and weft weaving method is widely used in the weaving of woven fabrics. In this method, the yarns of two systems, which are perpendicularly to each other, are interwoven to form a product according to a certain rule. Among them, the yarn of one system along the length direction is called a warp, and the yarn of one system along the width direction perpendicular to the warp is called a weft. When applied to the weaving of the beaded cooling pad, the yarn is regarded as a string. More specifically, the warp is called a warp string and the weft is named a weft string, as shown in Fig 6. Due to the beaded pad is woven by the \"warp and weft weaving\" method, the bead in the adjacent four pad units will fluctuate greatly, and the maximum fluctuation direction is perpendicular to the plane formed by the warp and weft strings. In order to reduce this fluctuation, a joint is added at the Method Research and Mechanism Design of Automatic Weaving\u2026 2541 intersection of the warp and weft strings. The joint is composed of two equal-diameter holes whose axes are perpendicular to each other and in the same plane. The outer shape is designed to be spherical for aesthetic appearance, so the joint is called a cross ball joint, as shown in Fig. 7. Since the cross-ball joint is added at the intersection of the warp and weft strings, the warp and weft strings won't be interlaced and the weaving process is simple. The string can be straightly passed through the bead and the cross ball joint. The weaving result is illustrated in Fig. 8. Fig. 6. Connection diagram of wrap and weft strings. Fig. 7. Schematic diagram of the cross-ball joint. Stereogram Section graph Fig. 8. Schematic diagram of warp and weft weaving beaded cool pad. 3.2 Lock Stitch Sewing Weaving Method In the sewing machine, there is a locked stitch consisted of two stitches intertwined together, and the interlacing point in the middle of the sewing material can be demonstrated in Fig. 9. Fig. 9. Diagram of the locked stitch. Fig. 10. Effect diagram of lock stitch sewing weaving method. In the lock stitch sewing weaving method, the vertical alignment beads are regarded as the sewing material. And the transverse arrangement beads are placed on the horizontal suture. The weaving result can be demonstrated in Fig. 10, in which the dotted line indicates the suture. In this method, it is firstly needed to insert the vertical alignment beads into a plurality of hooks. The number of hooks is one more than the number of columns of the beaded pad, the next weaving process can be illustrated in Fig.11 [12-13]. Welt string Warp string Stereogram Section graph Welt string Warp string S. Ouyang et al.2542 3.3 Single-line Straight-through Method The single-line straight-through method is a kind of weaving method that is easy to implement on the machine. Firstly, insert all beads that downlink line needed from one row of the beaded pad into the string, and the downlink line can be in any shape. Then the weaving process of every row of the beaded pad can be achieved by the threading movement of the uplink line. Finally, every row of the beaded pad can be joined together to form the desired b eaded pad. And the effect of this method is the same as that of hand-weaving method. For the beaded cooling pad LDm\u00d7p, each row has p units, so the number of beads required for the down line is 2p+1. The specific weaving steps are shown in Fig.12. And repeat steps in Fig.12 until the desired single row of beaded pads with units are woven, as shown in Fig. 13. Next each row of weaving beaded pad is repeated from the first step until the required number of beaded pad are completed, and then the single row of beaded pad are joined together to form the desired beaded cooling pad. To apply this weaving method on the machine, the machine can be designed in two parts, the first part is used to complete step 1 and the second part is used to complete steps 2 -7. It can be seen from the weaving process that the first part only has the downlink line threading, while the second part only has the uplink line threading. The whole process avoids the cumbersome action of repeatedly crossing the uplink and downlink lines. The beaded cooling pad woven by this method has a better stable and reliable string connection. As it ca n only be woven one row or one column at a time, it is easier to implement on a machine and has higher efficiency, providing an effective weaving method for designing and manufacturing a beaded cool pad automatic weaving machine. Method Research and Mechanism Design of Automatic Weaving\u2026 2543" ] }, { "image_filename": "designv11_62_0000661_15325008.2011.596752-Figure3-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0000661_15325008.2011.596752-Figure3-1.png", "caption": "Figure 3. Angles \u02c7di and \u02c7qi in a four-pole salient pole rotor.", "texts": [ " (32) and (33). In the same way, fr2 can be derived: D ow nl oa de d by [ Fl or id a St at e U ni ve rs ity ] at 0 3: 46 0 3 Ja nu ar y 20 15 Computation of Salient Pole Machine Inductances 1515 fr2.\u02c7m; \u0131/ D p X iD1 \" Z i d C.i 1/ q .i 1/ d C.i 1/ q g 1 d .\u02c7; \u02c7m; \u0131/na.\u02c7/d\u02c7 C Z i d Ci q i d C.i 1/ q g 1 q .\u02c7; \u02c7m; \u0131/na.\u02c7/d\u02c7 # D p X iD1 \" na.\u02c7di / Z i d C.i 1/ q .i 1/ dC.i 1/ q g 1 d .\u02c7; \u02c7m; \u0131/d\u02c7 C na.\u02c7qi / Z i d Ci q i d C.i 1/ q g 1 q .\u02c7; \u02c7m; \u0131/d\u02c7 # : (38) The parameters \u02c7di and \u02c7qi are shown in Figure 3. Considering Eqs. (30) and (31), fr2 is obtained as follows: fr2.\u02c7m; \u0131/ D p X iD1 na.\u02c7di /.frd .i d C .i 1/ q ; \u02c7m; \u0131/ frd ..i 1/ d C .i 1/ q; \u02c7m; \u0131/ C na.\u02c7qi /.frq.i d C i q ; \u02c7m; \u0131/ frq.i d C .i 1/ q ; \u02c7m; \u0131// ; (39) where frd and frq are calculated from Eqs. (32) and (33). Function fr is defined as follows: fr .\u02c7m; \u0131/ D fr2.\u02c7m; \u0131/ fr1.\u02c7m; \u0131/ : (40) Considering Eqs. (28), (34), (35), and (40), the MWF of winding b is derived as follows: Mb.\u02c7; \u02c7m; \u0131/ D nb.\u02c7/ fr .\u02c7m; \u0131/: (41) D ow nl oa de d by [ Fl or id a St at e U ni ve rs ity ] at 0 3: 46 0 3 Ja nu ar y 20 15 Replacing Eq" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0002791_b978-0-12-812138-2.00003-9-Figure3.35-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0002791_b978-0-12-812138-2.00003-9-Figure3.35-1.png", "caption": "FIGURE 3.35", "texts": [ " At normal operating speeds, since the rotor frequency is low, the rotor current is almost uniformly distributed over the cross-sectional area of the rotor bar, and thus the effective resistance Reff will be as low as the DC value Rdc. For a deepbar rotor, its effective resistance at starting may be several times greater than at the rated speed as shown in Fig. 3.34. As explained above, for induction motors with a deep-bar rotor, the rotor resistance varies according to the speed so that a satisfactory performance can be achieved both at the start and normal operations. A double-cage rotor is another type of rotor whose effective resistance varies as shown in Fig. 3.35 [4]. It consists of two cages to exaggerate its variation more than the deep-bar. The upper cage has smaller cross-sectional areas than the lower cage. In addition, the upper cage is made up of a higher resistance material than the lower cage. Similar to a deep-bar rotor, the rotor bars in the lower cage have a greater leakage flux and thus, higher leakage inductance. At starting, since the rotor frequency is high, the division of the rotor current is mainly decided by the leakage reactance difference between the two cages" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0002621_aim.2017.8014122-Figure6-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0002621_aim.2017.8014122-Figure6-1.png", "caption": "Figure 6. The simplified diagram of calculation of safety distance with ultrasonic sensor", "texts": [ " 5, which the red dotted is the original level walking trajectory of the swing leg and the modified level walking trajectory is presented by the green one, and make sure the swing leg can be placed on stair area after changing its walking trajectory, and, finally, to achieve the purpose that the apparatus can climb up the stairs safely along with stairs walking trajectory and change back to the level walking trajectory after finishing climbing. For the sack of achieving the function to keep the safety distance by the walking assistance apparatus, distance sensors are needed to be mounted on the toe part of the apparatus to measure the distance between it and the stair like Fig. 6. Distance sensor can be used to do its work by its changing of output voltage, and we use the value of output voltage in programing to control the apparatus\u2019 walking trajectory. Therefore, as the result of the reason above, the precision of measured voltage from sensor is one of the most important factor to choose. We choose TAKEX US-S25AN to be as the distance sensor, which has a relative good performance in a short-range case. An adequate level walking trajectory is get from the previous record about the changes in the angle of each joint of subjects, and, to prevent falling while the user wearing the walking assistance device, the stride is decreased and the step height is increased", " The stairs walking trajectory can be also get with the same way, but the limit of stride need to be controlled more carefully, because a safety stride of stairs walking need to be considered with the safety distance, and stairs length and height. Consequently, the safety distance should be defined first in our work. To find the value of required safety distance, an experiment was held to measure the average data from 10 subjects. These subjects were asked to walk close to stairs from a designated position that the distance from the subject to stair is 318 (mm), and measured their safety distance after one of the subject\u2019s leg put on the first stair. The average value we decide is 191 (mm). The simplified diagram is indicated in Fig. 6. The step length while stairs walking can be decided per the area which the foot contact on the stair and the safety distance. Therefore, the limit value of step length, dstairs can be calculated by Eq. (4). thsafestairs ddd (4) Where, dsafe is the value of the safety distance, and dth is the designated length of the contact area for making sure the walking safety. When the step distance decrease, the stride from even ground to stair also decrease, and the limit stride can be calculated as )(22 thsafestairsMIN dddStride (5) To avoid falling, the minimum stride of walking trajectory from ground to stairs need to be equal or bigger than MINStride " ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0000874_s12541-013-0228-2-Figure5-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0000874_s12541-013-0228-2-Figure5-1.png", "caption": "Fig. 5 Motor driven crank slider simulator", "texts": [ " (7), where M, \u03c9, R are the equivalent mass of all the mechanical components, the angular velocity and the radius of the crank shaft, respectively. Fin = M\u03c92 R (7) With the equivalent mass and the radius of crankshaft constant, the inertia force is related to the squared engine speed. It is very difficult to test the bearing wear with such a large engine, because it\u2019s a very time consuming process. So, instead of a real piston driven crank mechanism, a motor driven crank-slider mechanism shown in Fig. 5 is adopted to investigate the effect of bearing wear and inertia force on the slider displacement. The sensor used here is an eddy current sensor which has a range of 1-6 mm @2 kHz. To mimic a certain amount of accumulated wear the gap between the sensor and the slider can be adjusted through a micrometer. As the motor runs, the slider repeats moving from TDC to BDC. Fig. 6 shows a series of gap data around BDC collected by the gap sensor during a cycle. It is seen that a value of BDC level, which corresponds to the lowest position, can be acquired for each cycle" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0000566_1.4000814-Figure2-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0000566_1.4000814-Figure2-1.png", "caption": "Fig. 2 Horizontal and vertical fins", "texts": [ " gi i=1,2 , . . . ,12 are parameters related to the physical characteristics of the VMUV and surrounding environment, and I is the inertia moment vector. 3.2 Control Surfaces. Although in the VMUV, fin angles are control inputs; they do not appear in the dynamic equations 32 directly. Thus, the knowledge of relation between control inputs and state variables is of crucial importance. These control surfaces affect hydrodynamic coefficients, which appeared in hydrodynamic forces and moment equations 8 Fig. 2 . There is an interface block\u2014DATCOM\u2014which calculates the hydrodynamic coefficients based on the control inputs, the VMUV attitude in water, and other factors 19 . DATCOM is a nonlinear mapping that uses look up tables, mathematical equations, and first block based on \u201ea\u2026 only offline the n Transactions of the ASME Terms of Use: http://asme.org/terms e D r o c T I V J Downloaded Fr xperimental curves to find the hydrodynamic coefficients. Since ATCOM does not have a closed-form mathematical formula, it is equired to use black box identification methods" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0001706_978-94-007-4204-8_14-Figure14.4-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0001706_978-94-007-4204-8_14-Figure14.4-1.png", "caption": "Fig. 14.4 Schematic diagram of a 24-well microbial fuel cell (MFC) array as shown in Hou et al. [126]. Composed of an anode layer (1 anode electrode layer, 2 anode well layer), a proton exchange membrane (3 PEM), and a cathode layer (4 cathode well layer, 5 cathode electrode layer)", "texts": [ " The major limitations thus far in the generation of \u201csuperbugs\u201d is the current \u201clove affair\u201d of many MFC scientists with the known electrogenic organisms such as Geobacter and Shewanella species. For example, if one plugs in the words \u201cShewanella,\u201d \u201cGeobacter,\u201d and \u201cmicrobial fuel cell\u201d into the academic search engine PUBMED, there are 97 hits as of March 2011. Thus, despite the fairly large number of known electrogenic bacteria, the majority of studies involve studies on MFCs powered by these organisms. A recent paper by Hou et al. [126] showed elegantly in a high throughput, 24-well format (Fig. 14.4) that the overall power densities of the top electrogenic bacteria varied only modestly (\u02d910%). Although not at all to be considered a major negative, the metabolic properties of Shewanella and Geobacter species certainly pale with known, highly competitive soil-dwelling organisms such as species of the genera Actinomyces, Bacillus, Burkholderia, and Pseudomonas. Many of the organisms with high electrogenic properties include members of the genera Bacillus, Enterobacter, Arthrobacter, Stenotrophomonas, Aeromonas, Shewanella, Paenibacillus, and Pseudomonas" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0000155_sps.2013.6623589-Figure2-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0000155_sps.2013.6623589-Figure2-1.png", "caption": "Fig. 2. Distribution of the principal axes of flying vehicle.", "texts": [ " The result of the project can be also used to other vehicles, where we need to correct the position. In the project is shown that fuzzy logic does not deteriorate the stability of the vehicle and in some cases it can improve the results. II. FLYING VEHICLE CONTROL Each flying vehicle can be assigned with 3 axes: angle of rotation along the direction of the flight (called roll), angle of transverse rotation (called pitch) and angle of vertical rotation (called yaw). Graphical illustration is presented in Fig. 2. A quadcopter flight control is performed by changing the rotors speed. Each rotor produces a lift force and torque. If all rotors rotate with the same angular velocity - the clockwise rotors balance torque of anticlockwise rotors causing no rotation in the yaw axis. Acceleration of the axis pitch and roll can be achieved without changing the value of the yaw axis. For example, if we want quadcopter to fly forward we need to increase the angular velocity of the motor 3 against motor 1 at unchanged speed of motors 2 and 4" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0003085_intellisys.2017.8324247-Figure1-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0003085_intellisys.2017.8324247-Figure1-1.png", "caption": "Fig. 1. Bipedal locomotion model with compliant legs (only the right leg is shown in this figure).", "texts": [ " In Section 4: a trajectory generation by using ZMP method. Section 5 is devoted to the discussion of obtained simulation results to show the influence of the muscles on the total work of each leg, and the conclusion is drawn in the Section 6. II. MODELING OF A BIPEDAL ROBOT WITH COMPLIANT In order to have an effective analysis on the effect of mono and biarticular muscles on the total work of each leg of robot during walking, we consider the bipedal robot model with springs like mono and biarticular muscles as shown in Fig. 1. The bipedal model shown in Fig. 1 consists of four leg limbs( with upper body, three joints (hip, knee and ankle joints) and four linear springs which are represented by the red dashed lines. The springs S1, S2, S4 correspond to the biarticular muscles: rectus femoris (RF), biceps femoris (BF) and gastrocnemius (GAS) in human legs respectively. Additionally, the spring S3 that correspond to a monoarticular muscle: tibialis anterior (TA) in human legs is also used in this model. The force generated in these springs is calculated as follows: (1) where and denote the stiffness coefficient and the displacement (lengthening or compression) of spring respectively, and in our case the intrinsic damping factor of springs is neglected" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0002302_b978-0-12-803581-8.10351-0-Figure32-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0002302_b978-0-12-803581-8.10351-0-Figure32-1.png", "caption": "Fig. 32 Displacement due to rack loading.", "texts": [ " The bulkhead in actuality would be constrained along the outer surface of the bonding flange, by the means of the chassis. Thus, the results of this form of loading is conservative and even so, the bulkhead had failure index of 0.25 (safety factor of 4 based on ultimate strength), shown in Fig. 31. The rack indent and the bolt holes were loaded with 1000 lb force through the bulkhead. The results of the rack load concluded that the bulkhead had failure index of 0.33 (safety factor of 3.03 based on ultimate strength), as shown in Fig. 32. The main design of the roll hoop is to prevent the driver from becoming crushed by the chassis in the event of a rollover crash. To increase the efficiency of this structure required by rules, it was also used to react loads from the suspension. Under the rules of the competition that this monocoque was originally designed for, the main forward roll hoop was constrained to be a metal tube structure. The design of this tube structure will not be covered in this discussion. However, how it is incorporated into the chassis structure is an important component of the composite chassis design for manufacture" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0000261_iros.2010.5651258-Figure1-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0000261_iros.2010.5651258-Figure1-1.png", "caption": "Fig. 1. Relations between impact posture and potential barrier", "texts": [ " The robot must start walking with a suitable and sufficient initial momentum to overcome the potential barrier and to reach the next impact. It is not easy to guarantee overcoming the potential barrier only with intuitive control laws in limit cycle walking. It is also difficult to start walking from a standing posture smoothly and we must search the suitable initial conditions through a trial and error process. The potential barrier in dynamic gait originally comes from the fact that limit cycle walkers have anterior-posterior symmetric impact posture as shown in Fig. 1 (a). To solve this problem, the author proposed a method for generating a gait that guarantees to overcome the potential barrier by asymmetrizing the impact posture as shown in Fig. 1 (b). The primary purpose of this method is to tilt or shift the robot\u2019s center of mass (CoM) forward for overcoming the potential barrier at mid-stance easily. The easiest way to asymmetrize the impact posture is to extend the stance leg during stance phases using the prismatic joints or knee joints. The author numerically investigated the validity of the proposed method using a telescopic-legged rimless wheel model shown in Fig. 2 (a), and performed parameter study in [6]. We showed the F. Asano is with the School of Information Science, Japan Advanced Institute of Science and Technology, 923-1292 Ishikawa, Japan fasano@jaist" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0000371_10402004.2013.861047-Figure1-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0000371_10402004.2013.861047-Figure1-1.png", "caption": "Fig. 1\u2014Configuration of a hydrodynamic rolling hybrid bearing: (a) the rotor rides on the rolling bearing at low speeds and (b) the rotor is supported by the hydrodynamic bearing at high speeds.", "texts": [ " The aim of this study is to investigate the effects of rolling bearing type and size on the maximum eccentricity ratio in order to provide a theoretical basis for the design of rolling bearing clearance. An analytic model for the maximum eccentricity ratio is presented. An evaluation method of the maximum eccentricity ratio for ball bearings and cylindrical rolling bearings are developed. The effects of the inner and outer diameters of the two bearings on the maximum eccentricity ratio are also investigated. An HRHB consists of a hydrodynamic bearing and a rolling bearing arranged side by side, as shown in Fig. 1. The rolling bearing has a fixed clearance smaller than that of the hydrodynamic bearing. At low speeds, the shaft only rides on the rolling bearing without rubbing against the hydrodynamic bearing, as shown in Fig. 1a. While at high speeds, the hydrodynamic bearing raises and supports the shaft alone, as shown in Fig. 1b. Several magnets are set on the side surface of the outer ring to attract or pull the roller toward the outer raceway; thus, the rolling elements of the rolling bearing do not make contact with the inner ring at high speeds. An analytic model for the maximum eccentricity ratio is developed to study the effects of the rolling bearing on the maximum eccentricity ratio, as shown in Fig. 2. O and OJ in this figure are the centers of the HRHB and the shaft, respectively. It is assumed that the hydrodynamic bearing and the rolling bearing have no misalignment; that is, the clearance circle of the rolling bearing has the same center with that of the hydrodynamic bearing. CH and CR are the radius of the clearance circle of the hydrodynamic bearing and rolling bearing, respectively. It is assumed that the oil film force WH and the rolling bearing reaction WR exerted on shaft are in balance with the external load WE under the supporting state as shown in Fig. 1a. Therefore, the balance equation can be written as follows: \u23a7\u23a8 \u23a9 WR sin \u03b8R \u2212 WH sin \u03b8H = 0 WR cos \u03b8R + WH cos \u03b8H = WE , [1] where WE is the external load; WH and \u03b8H are the magnitude and direction of the oil film force, respectively; and WR and \u03b8R are the magnitude and direction of the rolling bearing reaction, respectively. D ow nl oa de d by [ M os ko w S ta te U ni v B ib lio te ] at 0 6: 36 1 7 Fe br ua ry 2 01 4 TABLE 1\u2014INNER PARAMETERS OF BALL BEARING Inner Ring (mm) Outer Ring (mm) Bearing Designation Diameter of Ball (mm) Groove Diameter Bottom Diameter Groove Diameter Bottom Diameter 6205 7" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0000969_012001-Figure3-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0000969_012001-Figure3-1.png", "caption": "Figure 3. Typical concrete box girder viaduct of spanish high-speed railway lines.", "texts": [ " The bodies are linked together by linear springs and dampers, which simulate the behavior of the two suspension levels of passengers railway vehicles. Vertical and lateral springs and dampers are used for modelling primary and secondary suspensions. In addition, yaw dampers are fitted longitudinally between the body car and bogie to damp hunting motion ([24]). Structures considered into this work are high-speed railway viaducts. A lot of that viaducts can be represented as multi-span continuous beams. They are usually single concrete box girders with two railways tracks (figure 3). Due to that, trains crossing the viaduct introduce an eccentric vertical load over the structure. For this work goal, these structures can be modeled using three-dimensional beam elements into a finite element framework. Beams with shear deformation effect have been used in the calculations [25]. In order to introduce the interaction between vehicle and structure, rigid surfaces are joint to the beam elements. These surfaces will move according to these elements. Thus, the vehicle will transit over these surfaces carrying the loads over the structure (figure 4)" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0001509_1350650111404114-Figure10-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0001509_1350650111404114-Figure10-1.png", "caption": "Fig. 10 The power spectrum and phase diagrams for different rotating speeds with unbalance mass: (a) power spectrum diagram n\u00bc 13 000 r/min; (b) power spectrum diagram n\u00bc 79 000 r/ min; (c) power spectrum diagram n\u00bc 130 000 r/min; (d) the phase diagram n\u00bc 13 000 r/ min; (e) the phase diagram n\u00bc 79 000 r/min; and (f) the phase diagram n\u00bc 130 000 r/min", "texts": [], "surrounding_texts": [ "The cases that consider the unbalance mass are carried out. The unbalance mass acts on the node 4. The mass weight is 0.2 kg, and the eccentricity 0.0001 m. When the rotating speeds is 13 000 r/min, the rotor runs with the stable periodic status, as shown in Fig. 9(a). The amplitude of the motion increases as the rotating speeds increase. When the rotating speeds reaches 79 000 r/min, the stable periodic motion destabilizes. The periodic motion is kept as the rotating speed reaches 118 000 r/min. The phase diagram Figs 10(d) to (f), indicates that the disciplinarian of the motion becomes more complicated. Then, the frequency analysis for the gas-bearing rotor system with unbalance mass is carried out. When the rotating speed is 13 000 r/min, the basic frequency generated from the unbalance mass is the main component, as shown in Fig. 9(d). The amplitude of the basic frequency component is small. As the rotating speed increases to 79 000 r/min, the 0.25 frequency multiplication that the amplitude is bigger than the value of the basic frequency component appears in Fig. 9(e). The lower frequency whirling phenomenon is obvious. As the rotating speed reaches 118 000 r/min, the amplitude of the 0.25 frequency multiplication increases as the rotating speed increases. The amplitude of 0.25 frequency multiplication is the biggest in the system. The power spectral analysis is performed for different rotating speeds, which is shown in Figs 10(a) to (c). It is observed that the frequency component of the power spectral is more complicated than that of when the unbalance mass is ignored. As there is increment in the rotating speed, the power of the basic frequency becomes smaller. On the contrary, the lower frequency component is the main part of the whole frequency component. Also, some certain Proc. IMechE Vol. 225 Part J: J. Engineering Tribology at NORTH CAROLINA STATE UNIV on May 2, 2015pij.sagepub.comDownloaded from Proc. IMechE Vol. 225 Part J: J. Engineering Tribology at NORTH CAROLINA STATE UNIV on May 2, 2015pij.sagepub.comDownloaded from frequency components that exist between the basic and second frequency components appear. As shown in Fig. 11, the rotating speed is taken as the bifurcation parameter for the hybrid gas-bearing system. It is observed that as the rotating speed reaches 70 000 r/min, the stable periodic motion destabilizes. It indicates that the motion changes in the multiply periodic and the quasi-periodic motion occurs. Compared with the destabilization occurring rotating speed shown in Figs 7 and 11, it is indicated that the unbalance mass can be employed to increase the whirling speeds by delaying the occurrence of whirling for self-acting gas-bearing system. However, for the externally pressurized gas bearing shown in Figs 8 and 12, it is observed that the unbalance mass has little effect on the increase of the Hopf bifurcation occurring speed for hybrid gas-bearing system, which is different for the case of self-acting gas-bearing system." ] }, { "image_filename": "designv11_62_0001014_iros.2013.6696949-Figure6-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0001014_iros.2013.6696949-Figure6-1.png", "caption": "Fig. 6: Planar LIP plus flywheel model", "texts": [ " \u2022 In our method, we heuristically choose the contact to add or change in case of support change. We will refine these parameters through push recovery experiments that we led on human subjects. \u2022 Finally, we compared the results of the simple model based approach to human push recovery through the experiments we led on human subjects [15]. We obtained a first promising feedback on the realism of the approach. A. Illustration of the COM and internal kinetic energies using a Linear Inverse Pendulum (LIP) plus flywheel model A planar LIP plus flywheel model is shown in Fig.6. The flywheel represents the inertia about the COM. EKCOM and EKINT are then respectively the translational and rotational kinetic energies. We have: \u2022 EKCOM = 1 2 m x\u03072, EKINT = 1 2 J f \u03b8\u0307 2, J f \u03b8\u0308 = \u03c4 where J f is the inertia about the COM and m is the mass of the model. x and \u03b8 are defined in Fig.6. \u03c4 is the torque generated by the flywheel; it is the RAM about the COM We now consider the case when the COM is moving in the x positive direction and EKCOM and EKINT are minimized respectively with the weights wv and w\u03c9 : 1) If wv >> w\u03c9 : a large positive \u03c4 allows greater maximum horizontal forces used to decelerate the COM. The minimization of EKCOM is though enhanced with the RAM. EKINT is increased and should be minimized before exceeding joint limits. 2) If w\u03c9 >> wv : the priority of EKINT minimization tends to keep the flywheel stationary while the EKCOM is less minimized through a COP control" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0000614_amm.110-116.2940-Figure2-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0000614_amm.110-116.2940-Figure2-1.png", "caption": "Figure 2: Planar force angle stability measure [11]", "texts": [ " The lateral tilting of the two rear wheels in is shown schematically in Figure 1. It may be noted that for lateral tilt ability the wheel-ground contact must be a point contact and hence the wheel must be in the shape of a torus. It is shown by Nilanjan and Ghosal [3-4] that with the capability of lateral tilt, a three wheeled mobile robot can traverse an uneven terrain without slip. In this work, we use the same concept and analyze such a WMR for tip over stability. The force-angle tip is used in this paper and this is discussed next. Tip over axis and its normal. Figure 2 shows force-angle stability measure for planar system whose centre of mass is subjected to net force fr. This force makes an angle \u03b81 and \u03b82, with the two tip over axis normals I1 and I2. The force angle stability measure, \u2200, is given by minimum of the two angles weighted by the magnitude of the force vector for heaviness sensitivity as given below \u2200 = \u03b81 || fr || (3.1) Of all the vehicle contact points with the ground, it is only necessary to consider those outermost points which form a convex support polygon when projected onto the horizontal plane, and these points are referred to as ground contact points" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0000010_afrcon.2011.6072176-Figure3-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0000010_afrcon.2011.6072176-Figure3-1.png", "caption": "Figure 3: Graphical Illustration of Euler Angles [10]", "texts": [], "surrounding_texts": [ "It is shown by Peddle [7] that by the principle of timescale separation the aircraft dynamics can be split into two subsections: The rigid body kinetics and the point mass kinematics. It is also proven by Peddle, that if high-bandwidth inner loop Specific Acceleration Controllers1 can be designed, then from a guidance perspective the UAV reduces to a point mass under the influence of virtual acceleration \u201cactuators\u201d. These virtual actuators are the axial, lateral and normal Specific Accelerations coordinated along the Xw, Yw and Zw axes respectively. In addition, the point mass is also subject to the effect of the inertially fixed gravity vector2 (GI). Furthermore, the direction of the total Specific Acceleration vector is determined by the orientation of the wind axis with respect to the inertial frame. In order to achieve this, a roll rate (Pw) about the Xw can be commanded to rotate the wind axis. Peddle further states that if one desires the aircraft to operate using coordinated flight 3 , the Lateral Specific Acceleration (Bw) should not be used for guidance. Thus, Bw is assumed to be commanded to zero throughout the flight envelope. However, Gaum [8] has illustrated that the Axial Specific Acceleration Controller (Aw) or ASA, is not suitable for guidance control implementation due to the bandwidth limited nature of its underlying actuator (the engine). Thus, from a guidance perspective the only commandable inputs which hold the timescale separation principle are: \u2022 Normal Specific Acceleration (Cw) \u2022 Roll Rate (Pw) In a similar manner, for this paper, the design has also disregarded the airspeed command as a feasible means of guidance for collision avoidance due to the aforementioned sluggish and bandwidth limited nature of the engine. Thus, the mathematic model employed here is outlined by Fig. 4. 1 Specific Accelerations are all the accelerations affecting the aircraft except for gravity. 2 GI = [0 0 g] = [0 0 9.81]m/s2 in the NED axis. 3 Coordinated flight is flying with zero sideslip. 978-1-61284-993-5/11/$26.00 \u00a92011 IEEE The following assumptions have been made for the sake of simplicity: 1) Virtual actuators are available for the outerloop guidance algorithms \u2013 This assumption means that the outerloops can exploit the timescale seperation principle and can ignore the Specific Acceleration dynamics. This implies that the interface to the aircraft will be at a virtual actuator level. 2) Airspeed is held constant \u2013 From a guidance perspective the airspeed is not commandable but is being held constant by some arbitrary control algorithm. Finally, the direction of the total Specific Acceleration vector is determined by the attitude of the wind axis with respect to the inertial frame. In order to achieve this, a roll rate (Pw) about the Xw can be commanded to rotate the wind axis. The dynamics equations are shown below: (1) (2) ! (3) III. COLLISION AVOIDANCE SYSTEM DEVELOPMENT In the forthcoming sections the development of the CASSAM is shown." ] }, { "image_filename": "designv11_62_0002359_j.talanta.2017.05.074-Figure1-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0002359_j.talanta.2017.05.074-Figure1-1.png", "caption": "Fig. 1. Schematic of three-phase boundary gas-phase electrochemical sensor. The thickness of the sol-gel layer is intentionally exaggerated as a visual aid. The true thickness of the layer over the epoxy plug is 100\u2013200 \u00b5m.", "texts": [ " The exposed length of the working electrode was partially covered by capillary action. The sol was dried under ambient conditions for at least 15 h before use, and the solgel was aged for one week prior to use in studies where long term stability was the objective. Longer exposure to the ambient laboratory atmosphere did not change the performance in that the water content was constant. The water content with this formulation was 35% as measured by thermal gravimetric analysis. The initial test of the probe, the design of which is shown in Fig. 1, was on the cyclic voltammetry of static samples of gas-phase phenol in He in the head-space above the silica sol-gel. The phenol vapor was introduced to the He with a syringe pump via the 6-port valve described in the Experimental section. Consistent with previous reports [21,22], an irreversible oxidation at 0.9 V was observed. Peak currents of 0.7 and 3.5 \u00b5A for 1.6 and 16 parts per thousand (volume) phenol, respectively, were developed. After continuous cycling, a passivating film that was visually observable was formed on the Pt working electrode" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0002375_iea.2017.7939175-Figure7-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0002375_iea.2017.7939175-Figure7-1.png", "caption": "Figure 7. A measuring of dimension of nozzle diameter", "texts": [ ") are set as per experiment plan, as shown in Table II. In experiment plan, study of two variable factor at four level requires number of experiments is 16 (24). After the parts were printed the used nozzle was then cut and investigated its cross section by milling machine and microscope, respectively. The used nozzles are cut by milling machine, then their cross-section are measured in order to investigate to variation, as shown in Fig. 6. To determine the wear of nozzle, the diameter of nozzle that prints specimen is measured, as shown in Fig. 7. The percentage of the nozzle wear rate is calculated by the following equation: Wear rate (%) = (|(0.4 \u2013 X)/0.4|) \u00d7100 (3) where X is the dimension of nozzle diameter after print specimen (mm) and 0.4 is the size of nozzle before print specimen (mm). III. RESULT AND DISCUSSION The results were obtained by printing all the four level of two variable factor and controlled fixed factor. The result shown that nozzle shapes related with the variable factor. The diameters of nozzles were expanded when more printed material and feed rate was printed" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0000649_1.4001727-Figure1-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0000649_1.4001727-Figure1-1.png", "caption": "Fig. 1 Reuleaux\u2019s method", "texts": [ "org/about-asme/terms-of-use 2 p r t p i p r t c l t r t 3 M P p r c s e l E o p S 0 Downloaded Fr Reuleaux\u2019s Method Reuleaux\u2019s method 15 uses geometric construction to find the ole of planar displacement using two homologous points of a igid body. Analogously in differential kinematics, the instant cener of zero velocity is found as the intersection point of the lines erpendicular to the velocities of two points of the moving body. In finite kinematics, the pole is the intersection point of bisectng lines of the line segments joining each of the homologous oints P1 , P1 and P2 , P2 , as shown in Fig. 1. The total angle of otation of the rigid body around the pole can be determined from he construction, and thus, the circular motion of the body is reonstructed. The angle of rotation will be the angle between the ine connecting the pole to any point of the body at the first posiion and the line connecting the pole to the same point after the otation. Reuleaux\u2019s method can be used for the kinematic regisration problem 14 . Three-Dimensional Generalization of Reuleaux\u2019s ethod Let La= la , l\u0304a and Lb= lb , l\u0304b be a pair of lines given in their l\u00fccker coordinates that belong to the rigid body before a dis- lacement", " The line coordinates of the screw axis that represents the motion from the first to the second position of the given lines will be the normalized common perpendicular between Ca and Cb, namely S = Ca Cb = s, s\u0304 1 The screw parameters of a helical motion consist of a line and a pitch. We have found the Pl\u00fccker line coordinates of the screw from Eq. 1 . We will use the same construction method to find the pitch of the screw. In the two-dimensional Reuleaux method, the angle of rotation of the rigid body around the pole is the angle between the line connecting any point of the body at the first position to the pole, and the line that connects the corresponding point to the pole after the rotation see Fig. 1 . Following the same line of thinking, we can see that in the three-dimensional case, the angle of rotation of the rigid body around the screw axis will be the angle between the common perpendicular of any line of the rigid body before displacement and the screw axis, and the common perpendicular of the corresponding line and the screw axis after the displacement see Fig. 3 . Let G be the common perpendicular between any line of the body at the first position and the screw axis, and G be the common perpendicular between the corresponding line and the screw axis after the helical motion" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0000017_2011-01-2148-Figure4-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0000017_2011-01-2148-Figure4-1.png", "caption": "Figure 4. Rear Steering About a Point.", "texts": [ " From the data collected the product \u03b41R was calculated for a vehicle with the rear axle centered as shown in Figure 2. This vehicle was equipped with an electronically controlled steerable rear axle using a simple control law steering the rear wheels \u03b43 as a constant proportion of the front steering input. Equation 4 A simple relationship can be formed based on their respective distances from the second (unsteered) axle [11], Equation 5 where l is the length between the first and third axles, and t is then tandem spread distance between the second and third axles. In this manner, as shown in Figure 4 the vehicle effectively turns about a point colinear with the extended second axle defined by its intersection with perpendiculars from the front and rear wheels. SAE Int. J. Commer. Veh. | Volume 4 | Issue 1 (October 2011)4 A multiple regression analysis was performed on the obtained data shown in Figure 2 and Figure 3 to yield modeled estimates of the equivalent wheelbase and equivalent understeer with the rear axle steering turned on and off as reported in Table 1. It is apparent in Figure 2 and Figure 3 that there was an asymmetry present in the vehicle causing different behavior in clockwise and counterclockwise turning; most likely an alignment issue" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0001116_1.3552336-Figure7-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0001116_1.3552336-Figure7-1.png", "caption": "FIGURE 7. (a) Standard model for the thermal system; (b) share of the single thermal resistances on the overall thermal resistance", "texts": [ "9 On: Tue, 11 Aug 2015 06:36:02 A parameter to estimate the expected cooling performance of a tool is the overall thermal resistance Rt. .1 , \u2211++= \u03bb \u03b4 h RR ctt (1) Here, Rt,c is the thermal contact resistance between the tool and the blank, h is the convective heat transfer coefficient between the tool and the fluid in the cooling channel and \u03a3(\u03b4/\u03bb) is the total thermal conductive resistance through the blank and the tool. According to a standard model for the thermal system of a hot stamping tool, shown in Fig. 7, the overall thermal resistance can be calculated according to Eq. 1 with the values given in Tab. 1. that the thermal conduction resistance of the tool \u03b4t/\u03bbt and the thermal contact resistance Rt,c make up the biggest part on the overall thermal resistance share. In order to improve the cooling performance, those parameters have to be optimized. The thermal contact resistance is dependent on the contact pressure, the surface roughness and the thermal conductivities of the touching bodies (Ref. [7])" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0000462_icma.2012.6282838-Figure1-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0000462_icma.2012.6282838-Figure1-1.png", "caption": "Figure 1: Three DOF planar robot manipulator with revolute joints.", "texts": [ "5), crossover rate (\ud835\udc36\ud835\udc5f = 0, 8), and population size (\ud835\udc41\ud835\udc43 = 50) were set based on the values suggested in [25]. As for the discarding parameters, empirical results revealed that a good performance is obtained with \ud835\udeff set equal to the 5% of the population size, \ud835\udefd is set to the 50% of \ud835\udc41\ud835\udc43 , and \ud835\udefc is defined as a Gaussian random variable with mean zero and standard deviation \ud835\udf0e = 2\u2218. B. 3DOFs planar robot manipulator A three degrees of freedom (\ud835\udc5b = 3) robot manipulator with all rotational joints has been considered as schematically illustrated in Fig. 1. The joint configuration vector is defined as ?\u20d7? = (\ud835\udf031, \ud835\udf032, \ud835\udf033) \u22a4, such that, each \ud835\udf03\ud835\udc57 (\ud835\udc57 = 1, 2, 3) represents a joint angle. Since the kinematic configuration constraints the end-effector motion to a 2D Cartesian plane, the pose vector is defined as ?\u20d7? = (\ud835\udc5d\ud835\udc65, \ud835\udc5d\ud835\udc66, \ud835\udf19) \u22a4 (\ud835\udc5a = 3), such that, (\ud835\udc5d\ud835\udc65, \ud835\udc5d\ud835\udc66) represents the end-effector position and \ud835\udf19 is the end-effector orientation respect to the horizontal axis. In Table III, the Denavit-Hartengberg kinematic parameters and the joint limits are summarized. The inverse kinematics of the robot in Fig. 1 admits two solutions for any end-effector pose inside its workspace, namely: elbow-up and elbow-down configurations. The exception occurs when the robot is at a singular joint configuration and thus the inverse kinematics has only one possible solution. This feature is examined by solving the inverse kinematics problem for a set of four different end-effector pose vectors, three of them correspond to non-singular joint configurations (?\u20d7?1,?\u20d7?3,?\u20d7?4) and a fourth to a singularity (?\u20d7?2). As described above, for each end-effector pose in the test bed 100 independent trials were conducted" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0000969_012001-Figure4-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0000969_012001-Figure4-1.png", "caption": "Figure 4. Schema of the structure model used in this work.", "texts": [ " Due to that, trains crossing the viaduct introduce an eccentric vertical load over the structure. For this work goal, these structures can be modeled using three-dimensional beam elements into a finite element framework. Beams with shear deformation effect have been used in the calculations [25]. In order to introduce the interaction between vehicle and structure, rigid surfaces are joint to the beam elements. These surfaces will move according to these elements. Thus, the vehicle will transit over these surfaces carrying the loads over the structure (figure 4). Interaction between vehicles and structure surface will be explain in detail in Section 4 The piers of the viaducts are modeled too using beam elements. Depending on bearing support types used, different relative displacements and rotations will be allowed or constraint. Thus, algebraic equations are used to impose constraints between nodes on the deck and piers. One of the most important points on railway dynamics is the wheel-rail contact. Due to the geometric and kinematic characteristics of wheelsets, wheels and rails, very non-linear effects appear when a vehicle goes out of its stationary circulation regime (wheelset not centered between rails)", " These forces can be written as: FD,z = cz ( dz(y) dy y\u0307 \u2212 z\u0307w,r ) , (6a) FD,\u03b8 = c\u03b8 ( d\u03b8(y) dy \u03b8\u0307 \u2212 \u03b8\u0307w,r ) , (6b) where cz and c\u03b8 are penalty damping coeficients. Numerical implementation of this model has been realised using the commercial software Abaqus. Thus, for solving this problem user implemented features have been coded in combination with standard capabilities. In order to introduce contact forces in wheel-rail interface in the model, additional nodes have been defined. As can be seen in Figure 4, contact nodes exist between vehicle model and rigid surfaces of the structure. One node per wheel has been defined. These nodes represent the rails locations over deck bridge. They will run over deck surface at the same velocity as train. Train model rests over these nodes and they will introduce the deck displacements and rotations on vehicle model. Furthermore, trains loads, including shear loads, are introduced in these nodes. Thus, a new contact interaction is needed for stablishing contact between contact nodes and rigid surfaces" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0000568_s11431-010-4273-0-Figure1-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0000568_s11431-010-4273-0-Figure1-1.png", "caption": "Figure 1 Sketch of air cannon exhaust process.", "texts": [ " air cannon, impulse force, optimization Citation: Xu J, Chen H S, Tan C Q, et al. Numerical and experimental investigations for an air cannon optimization. Sci China Tech Sci, 2011, 54: 345351, doi: 10.1007/s11431-010-4273-0 Air cannons, a kind of de-clogging device which produces impulse force by instantly releasing the compressed air deposited in a pressure vessel, are widely used in the transport pipes of warehouses, docks, furnaces, coal mines, building materials industry and chemical industry. Figure 1 shows the air cannon part section view and exhausting process sketch, including piston sleeve 2, exhausting nozzle 5, and piston 6. The piston 6 is impacted on the inlet of the nozzle 5 by the spring 4, sealing the nozzle inlet with the rubber which is embedded in the piston bottom. On the cylindrical face of the piston sleeve 2, two small holes are drilled upon the piton upper surface for air inflation and several large apertures which are covered by the piston are handled as air flow inlet for exhausting the compressed air by piston moving up", " The compressed air, through oil-water separator and fast electromagnetism valve which is separated with the air cannon, flows into the vessel along the dashed line in the sketch. When the air cannon works, the air in the piston sleeve flows out instantaneously by the fast speed valve and the air in the vessel can not fast flow into the piston sleeve by the two small holes, causing the piston moves up to the cover flange 1 by pressure difference, then the large apertures inlet of piston sleeve exposes and the compressed air in the vessel flows out through the exhausting nozzle along the solid line, generating high speed flow to impact the jams, as shown in Figure 1(b). After the operation, the piston moves back by the spring and the compressed air inflates into the vessel again. The exhausting process of air cannon is about several hundred milliseconds. Its structure and application were discussed in some refs. [1\u20134]. However, the research on the exhausting process and its influence on the air cannon performance are not yet reported, especially on the flow field of air cannon. The air cannon performance is dependent on the impulse force by the compressed air release, which is affected by the vessel volume, air pressure and exhaust process" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0000720_amm.307.304-Figure3-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0000720_amm.307.304-Figure3-1.png", "caption": "Figure 3 Replacing contact of the gears with a cylinder and surface", "texts": [ " The load changes linearly so that the maximum load is applied on the heel and the minimum at toe. Moreover, whether there is one pair of gear or two pair of gear in contact will affect the load distribution for bevel gear. After replacing the bevel gear with multiple spur gears, the load acting on each layer of bevel gear will apply on its equivalent spur gear. In order to determine the Hertzian stress, it is important to determine the equal radii of curvatures of bevel gear at each point. Considering the equivalent geometry as shown in Fig. 3 and using Eqs. 3-5 [5], the variation of radii of curvature illustrated in Fig. 4. 2 2 2 1 2 1 112 EEEeq \u03c5\u03c5 \u2212 + \u2212 = (3) 21 111 rrReq += (4) ( )( ) ( ) \u03c6 \u03c6\u03c6 tan tantan )( 21 21 bb bb eq RR XRXR xR + \u2212+ = (5) where: Ei (i=1,2) module of elasticity Eeq Equal module of elasticity ri (i=1,2) Local radii of curvature at contact point Rbi (i=1,2) Base circle radius Req Equal radius of curvature X Contact point coordinate along line of action \u03bdi (i=1,2) Poisson ratio pressure angle Fig. 5 shows the load distribution in bevel gear teeth" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0001662_vppc.2011.6043205-Figure1-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0001662_vppc.2011.6043205-Figure1-1.png", "caption": "Fig. 1. Flux Relations for Cylindrical Annulus", "texts": [ "27 r , Ta > 100 (10) where Pr = \u03bd \u03b1 is the Prandtl number and \u03b1 is the thermal diffusivity of air. The convection heat transfer, which is determined by the Nusselt number, can be combined with the conductivity heat transfer term in the heat transfer equation to form a new term to calculate both conduction and convection heat transfer [4]: keq (\u03c9r) = Nukair 2 , (11) where kair is the thermal conductivity of air. Once we have determined this conductivity, we exploit the fact that the air gap is a cylindrical annulus, as shown in Fig. 1. Solving Laplace\u2019s equation, and assuming a uniform normal heat flux density Hig leaving the boundary and temperature Tig at the air gap boundaries of the stator and rotor, we achieve the following heat flux/temperature relationship between the stator and rotor: Hsg = keq (\u03c9r) rsg ln ( rsg rrg ) (Tsg \u2212 Trg) , (12) Hrg = keq (\u03c9r) rrg ln ( rsg rrg ) (Tsg \u2212 Trg) (13) C. 1D FEA for Winding End-turns The temperature of the winding end-turns cannot be captured by a 2D FEA model. To capture the dynamics of the end-turns, a simple 1D FEA model is built" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0002691_chicc.2017.8027900-Figure1-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0002691_chicc.2017.8027900-Figure1-1.png", "caption": "Fig. 1: Flight profile", "texts": [ " The tail-sitter UAV is an aircraft that takes off and lands vertically on its tail. The whole aircraft tilts forward use differential thrust or control surfaces to achieve horizontal flight [5]. It can make the transition and don not need extra actuators, which is mechanically simple and saves a huge amount of weight [6]. Moreover, since the tail-sitter UAV land on its tail, it requires relatively stronger tail to be able to withstand landing impacts [5]. The profile of the tail-sitter flight is shown in Fig. 1. There are three different flight modes for the tail-sitter UAV, that is, the vertical fight, the horizontal fight and the *This work is supported by National Natural Science Foundation (NNSF) of China under Grants 61503009, 61333011, 61421063, the Aeronautical Science Foundation of China under Grant 2016ZA51005, the Special Research Project of Chinese Civil Aircraft, and the Fundamental Research Funds for the Central Universities under Grant YWF-14-RSC-101 transition fight. In vertical flight thrust is the dominant force and horizontal control is achieved via rotors rotation" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0001299_pedstc.2011.5742425-Figure3-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0001299_pedstc.2011.5742425-Figure3-1.png", "caption": "Fig. 3. Flux and current vector deviations in x-y reference frame.", "texts": [ " Therefore, the torque deviation is obtained as: (17) ' 2 1 1 1e e e s y y s s yT T T k i i i\u03c8 \u03c8 \u03c8\u23a1\u0394 = \u2212 = \u0394 + \u0394 +\u0394 \u0394\u23a3 By calculating flux deviation from (2) and substituting it in (17) and after long but simple calculations, the torque deviation is given by: (18) ( ) ( ) ( ) ( ) ( ) 2sin sin2 2 sin cos2 2 2 sin cos2 q d e x y m d q d q q d y x m d q d q L LAT k i i B L L L L L L k i i L L L L \u03b4 \u03b4 \u03c8 \u03b4 \u03b4 \u03c8 \u03b4 \u03b4 \u239b \u239e \u239c \u239f\u2032\u0394 = \u2212 \u0394 \u239c \u239f\u23a1 \u23a4\u2212 + + \u2212\u23a3 \u23a6\u239d \u23a0 \u239b \u239e \u239c \u239f\u2032\u2212 \u0394 \u239c \u239f\u23a1 \u23a4\u2212 + + \u2212\u23a3 \u23a6\u239d \u23a0 So, we have: (19) xye ikikT \u0394\u2212\u0394=\u0394 43 where, ( ) ( )[ ( ) ( )[ ] \u239f \u239f \u23a0 \u239e \u239c \u239c \u239d \u239b \u2212++\u2212 = \u239c \u239c \u239d \u239b \u2212++\u2212 \u2212= y qdqdm dq qdqdm dq i LLLL LL kk LLLL LL B Akk \u03b4\u03b4\u03c8 \u03b4\u03b4\u03c8\u03b4 \u03b4 2cossin2 2 2cossin2 2 2sin sin ' 4 ' 3 (20) The deviations of flux vector and current vector have been shown in more details in Fig. 3. As the figure shows the current deviation component (\u0394 ) and flux deviation component (\u0394 ) are aligned with the y-axis and the current deviation \u0394 and flux deviation component \u0394 are aligned with the x-axis. Therefore, it is proved that the radial component of the stator flux linkage deviation is proportional to the x-axis component of the stator current vector deviation and both are responsible for providing the motor flux. Also, the perpendicular component of the stator flux linkage deviation is proportional to the y-axis component of the stator current vector deviation and these are responsible for motor torque development" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0001687_978-90-481-9689-0_68-Figure8-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0001687_978-90-481-9689-0_68-Figure8-1.png", "caption": "Fig. 8 First (a), second (b), and third (c) manufacturing solution for the Geometry 1.B.", "texts": [ "B is taken (for clarity) as an example of this type of 3-UPU TPM. The first manufacturing solution S1, is to rebuilt the platform of the manipulator. This is obtained by disconnecting the platform of this geometry from the legs and rotating it by a suitable angle \u03b1 about the z axis of Sb, then connecting again the legs to the platform still keeping the same base axis directions. This means to manufacture a platform with the revolute axis directions rotated of \u03b1 (clockwise in the example shown in Figure 8a) with respect to the Geometry 1.B. This makes it possible to avoid the leg collision. In Figure 8a, the universal joints on the base and on the platform are represented by points for clarity, and the prismatic ones are omitted. After manufacturing the new platform, the coordinates of the center of the universal joint on the platform Ai, i = 1,2,3, are given by: 601 A.H. Chebbi and V. Parenti-Castelli OpAi = cos\u03b1OpA\u2032 i + sin\u03b1OpA\u2032\u2032 i , with OpA\u2032\u2032\u22a5OpA\u2032; and \u2225\u2225OpA\u2032\u2032 \u2225\u2225 = \u2225\u2225OpA\u2032 \u2225\u2225 (2) where A\u2032 i, i = 1,2,3, are the centers of the universal joints on the platform of the Geometry 1.B. The second manufacturing solution S2, schematically shown in Figure 8b, is to rebuild both the base and the platform of the Geometry 1.B in order to have the coordinates of the centers of universal joints at the base and at the platform, respectively Bi and Ai, i = 1,2,3, see Figure 8b, given as follows: ObBi = ObB\u2032 i + eq1i, and OpAi = OpA\u2032 i + eq4i (3) where B\u2032 i and A\u2032 i, i = 1,2,3, are respectively the center of the universal joints in the base and in the platform of the original Geometry 1.B; q1i and q4i, i = 1,2,3, are respectively the unit vectors of the revolute joints on the base and on the platform, which maintain the same directions of the original Geometry 1.B; e is a given distance between the corresponding center of universal joints in the platform of the Geometry 1.B and the platform rebuilt. The third manufacturing solution S3, schematically shown in Figure 8(c), is to rebuilt the second and the third link of each leg of the Geometry 1.B in order to change the physical position of the prismatic pairs on each leg along EiFi, where the coordinate of points Ei and Fi, i = 1,2,3 are given by: ObEi = ObBi + dq2i, and OpFi = OpAi + dq3i (4) where Bi and Ai, i = 1,2,3, are respectively the centers of the universal joints in the base and in the platform of the Geometry 1.B; q2i and q3i, i = 1,2,3, are respectively the unit vectors of the intermediate revolute joints of the i-th leg; d is a given distance between the directions of the prismatic pairs for the Geometry 1" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0002302_b978-0-12-803581-8.10351-0-Figure38-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0002302_b978-0-12-803581-8.10351-0-Figure38-1.png", "caption": "Fig. 38 Monocoque to rear clip interface general configuration.", "texts": [ " To attach the main roll bar and engine/transmission unit to the back of the monocoque a four bolt interface was developed. The details of this approach are specific to the rules imposed by the form of competition this design was originally intended for and are not directly pertinent to this discussion of the composite monocoque. However, to incorporate these four mounting studs into the rear of the monocoque, significant design effort focused on incorporating metallic inserts into the composite was necessary. The general configuration at the rear bulkhead of the monocoque is shown in Fig. 38. These interface brackets were designed to be bonded into the side shells of the monocoque. They were to be bonded in between the facesheets, in a fashion much like the hardpoints previously described. The four aluminum brackets were designed to carry the loads from both the rear clip interface stud and collar into the walls of the monocoque chassis. The bracket designs are shown in Figs. 39\u201342. Figs. 39 and 40 show the top bracket design. This design was used for both right and left sides; however, since the stud and collar attachment to the rear clip had to be slightly above the bottom of the chassis, the bottom brackets had to be designed separately" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0000918_iros.2013.6697166-Figure1-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0000918_iros.2013.6697166-Figure1-1.png", "caption": "Fig. 1. Hexapod model with left tripod in support", "texts": [], "surrounding_texts": [ "Our model for sagittal plane hexapedal running consists of a body with mass m and mass moment of inertia J , six massless, serial elastic actuated, telescoping legs and serial elastic actuated, rotating hips. A viscous damper is included in parallel to each spring. The legs are arranged in a sprawled posture. Front, middle and hind legs each have a different length. All hip joints are collinear with the body center of mass which is placed slightly behind the hip joint of the middle leg. A second order slip model is implemented for each foot that can be activated or deactivated. Upon activation the slip model moves the foot contact point once the ground reaction force leaves the friction cone. If it is deactivated the foot position is fixed during stance until the lift off condition is fulfilled. Two tripods consisting of left front, left hind and right middle leg (left tripod) as well as right front, right hind and left middle leg (right tripod) are controlled 180\u25e6out of phase with a single frequency periodic feedforward pattern. Hereby, the actuation changes the force free length of the leg and hip springs following a sine based pattern for the telescoping legs and a cosine based pattern for the rotating hip joints. The equations of motion of our hexapedal model are the following, mr\u0308b = fg + 6 \u2211 i=0 fl,i, (1) J\u03b8\u0308b = 6 \u2211 i=0 \u2212\u03c4h,i + 6 \u2211 i=0 (\u2212rh,isin(\u03b8b)flx,i \u2212 rh,icos(\u03b8b)flz,i) (2) \u03c4h,i = { \u2212(kh,i\u2206\u03b8i + dh,i\u2206\u03b8\u0307i) : ci = 1 0 : ci = 0 , (3) fl,i = { \u2212(kl,i\u2206li + dl,i\u2206l\u0307i)er,i + \u03c4h,i li et,i : ci = 1 02\u00d71 : ci = 0 (4) Herein, rb = (xb, zb) is the planar position of the body center of mass (COM) with respect to the world coordinate system. The vector fg represents the gravity force acting on the COM and fl,i, i = 1 . . . 6 are the ground reaction forces of the legs. \u03b8b is the pitch angle of the body and \u03c4h,i are the torques at the hip joints of the respective legs. The distance between the hip joint of the ith leg and the COM is given by rh,i which is positive for the hip being located in front of the COM and negative for the hip being located behind the COM. The parameters kh,i and dh,i are the spring and damping constant of the ith hip joint while kl,i and dl,i are the spring and damping constant of the ith leg. er,i is the radial unit vector of leg i that is directed along the leg towards the hip. et,i is the tangential unit vector of the respective leg that results from rotating er,i 90\u25e6 clockwise. The discrete state ci = 1 indicates ground contact of leg i. The touch down (TD) and lift off (LO) conditions of a leg are TD: zb \u2212 rh,i sin(\u03b8b)\u2212 l0,i(t) sin(\u03b80,i(t) + \u03b8b) <= 0 and LO: flz,i = 0, f\u0307lz,i < 0. The deflection of hip and leg spring is given by \u2206\u03b8i and \u2206li, respectively, which are calculated according to \u2206\u03b8i = \u03b8i \u2212 \u03b80,i \u2212 \u2206\u03b80,i 2 (1\u2212 cos(\u03c6(t) + \u2206\u03c6i)), (5) \u2206li = li \u2212 l0,i \u2212\u2206l0,i sin(\u03c6(t) + \u2206\u03c6i). (6) The terms \u03b80,i(t) = \u03b80,i + \u2206\u03b80,i 2 (1 \u2212 cos(\u03c6(t) + \u2206\u03c6i)) and l0,i(t) = l0,i + \u2206l0,i sin(\u03c6(t) + \u2206\u03c6i) are the force free length of hip and leg spring consisting of a fixed and a time varying component. The fixed components of the rotational hip springs, \u03b80,i are determined manually. In contrast, the fixed components of the translational leg springs, l0,i are calculated such that the feet touch ground for a configuration with a COM height h0, zero body pitch and leg angles according to \u03b8i = \u03b80,i. The time varying part of the force free length of the springs is the feedforward control signal with a time and frequency dependent phase, \u03c6\u0307(t) = \u03c9, and a leg specific fixed phase shift, \u2206\u03c6i. This leg specific phase shift is composed of interleg phase shifts within the same tripod group (\u03c6lf = \u03c6lh+2 \u00b7\u03c60 and \u03c6rm = \u03c6lh+\u03c60, where \u03c60 is a constant) and a 180\u25e6 phase shift with respect to the opposite tripod group. The subscripts lf , rm and lh indicate the specific legs of the left tripod, i.e. left front, right middle and left hind respectively. The right tripod follows the same naming conventions. To be more realistic with respect to slippage we include a second order slip model for the feet which can be activated or deactivated during simulation. If the slip model is activated a small mass is assigned to each foot in stance. Once the horizontal ground reaction force is larger than the static friction force the difference of both forces accelerates the foot mass and thus shifts the ground contact point of the foot. If the horizontal ground reaction force returns into the friction cone the foot motion is decelerated." ] }, { "image_filename": "designv11_62_0000018_2012-01-0620-Figure8-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0000018_2012-01-0620-Figure8-1.png", "caption": "Figure 8. Analyzed Positions for Test 1", "texts": [], "surrounding_texts": [ "For each of these full-scale yaw tests, the authors conducted a speed analysis using two of the models discussed earlier, the BNP/NCB model and CRASH model. Also, a simple model that assumed no braking and another that assumed full wheel lockup were included for comparison. Using diagrams created from our site survey, vehicle models were aligned with the tire mark evidence as depicted in Figures 8 and 9. The segment lengths between positions were measured and the average vehicle sideslip angles were calculated based on the angles at the beginning and end of the segments. To account for braking, the BNP/NCB and CRASH models require the longitudinal slip percentages to be known. Tire mark striations were used to calculate longitudinal slip percentages [Reference 9]. During the phase of the yaws where the parking brake was applied, the rear tires of the Malibu locked while the front tires were unaffected. This lockup was apparent in the tire mark evidence. Figure 10, for instance, shows the tire marks from Test 2. In this figure, each tire mark has been labeled with the tire responsible for depositing it. Longitudinal striations indicating wheel lockup are evident in the early portions of the left and right rear tire marks. Thus, the longitudinal slip percentage of the rear tires were assigned a value of 100% during this phase. Lighter longitudinal striations were also observed in the front tire marks, and these were considered in the analysis. Once the parking brake was released, striations from the rear leading tire mark were used to calculate longitudinal slip percentages and these values were applied to all tires. When the striation analysis indicated that a driven wheel was free rolling, the wheel was assigned a longitudinal slip percentage of 0.5%, to correspond with the measured rolling resistance of a similar vehicle documented in Reference 10. When tire marks from a non-driven wheel indicated no braking, a longitudinal slip value of 0% was used. In Test 1, the service brakes were not applied after the parking brake was released, so the longitudinal slip percentages were assigned values commensurate with free- rolling wheels. In Test 2, we used the longitudinal slip percentages shown in Figure 7. Four different methods were used to calculate the deceleration of the vehicle during each of the segments shown in Figures 8 and 9. These methods are summarized in Table 2. The first model, which we have termed the Simple No Braking Model, is similar to a model presented by Daily, Shigamura, and Daily [Reference 11], neglecting rolling resistance and roadway slope. The Simple No Braking Model, as well as the model presented in Reference 11, do not account for braking, and thus the Simple No Braking Model is used here to demonstrate a lower boundary on the speed in a vehicle sideslip angle analysis. Likewise, the Full Lockup Model is presented to demonstrate an upper boundary. For the analysis in this paper, use of the CRASH model required calculation of the (flong/\u00b5o) term. To calculate this value for a given segment, we first utilized the striation evidence in the tiremarks to calculate a longitudinal slip percentage, then used the NCB equations and the coefficients listed in Table 1 to determine the proportion of longitudinal force. In other words, the longitudinal slip was calculated from striation evidence, and used to calculate QX(s), which is equivalent to (flong/\u00b5o) in this application. Using the deceleration rates calculated with the models in Table 2, speeds at the beginning of each segment were calculated using Equations (10) and (11). These equations yield the speed at the beginning of the segment, given the speed at the end of the segment, the deceleration rate and the segment distance. Equation (10) is in general terms whereas Equation (11) is the common crash reconstruction equation, expressed in units of speed in miles per hour, distance in feet, and acceleration in g's. (10) (11) For Test #1, our analysis of the vehicle speed began at the end of the tire marks and worked back to the beginning. We chose not to begin the analysis of this test at rest because the VBOX data became erratic during the rollout phase at the end of the test. For Test #2, our analysis began at the vehicle rest position and worked back to the beginning of the tire marks. Figures 11 and 12 depict the results of this analysis. In these graphs, vehicle speed is plotted against distance. The solid black line represents the vehicle speed obtained from the VBOX. The other four curves represent the speeds obtained with the four methods listed in Table 2. As was expected, the no-braking model underestimated the vehicle speed over the entire duration of the test. Similarly, using roadway friction for the deceleration value (full lockup) overestimated the speed. For the two yaw tests analyzed in this study, both the BNP/NCB model and CRASH model yielded acceptable estimates for the vehicle speed." ] }, { "image_filename": "designv11_62_0002480_978-3-319-60399-5_7-Figure4-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0002480_978-3-319-60399-5_7-Figure4-1.png", "caption": "Fig. 4 The axial profile of the thread of the modified worm", "texts": [ " In order to provide the conjugacy of these surfaces along the entire height of the thread, the worm should be made quasi-Archimedes and its axial profile should be outlined along the convex curve described by the expression x\u00f0r\u00de \u00bc Z rp1 tan ax\u00f0r\u00de@r; \u00f038\u00de where r is the current radius at the active thread profile, which varies within the range from rp1 to ra1; rp1 = aw \u2212 0.5da2 cos(aa2 \u2212 ax1) is the radius of the lower boundary point of the active profile; and ra1 = da1/2 is the outer radius of the worm. The current value of the axial pressure angle ax is determined from (37) for the current value r and the corresponding lead angles of the worm threads c1(r) = atan (mx1z1/2r) and the hob c0(r) = atan[mx0z0/(2r \u2212 d1 + d0)]. The axial profile of the modified worm is shown in Fig. 4. Its curvature at the current point M is determined by the expression K\u00f0r\u00de \u00bc @ @r ax\u00f0r\u00de\u00bd cos ax\u00f0r\u00de: \u00f039\u00de The required axial profile can be approximated by the arc of a circle, parabola or ellipse, depending on the type of contour follower applied at the worm grinding machine-tool. Its arrow of the convexity fn is determined by the following expression with sufficient accuracy for practical purposes: fn m2 xK r1\u00f0 \u00de= 2 cos2ax1 ; \u00f040\u00de where K(r1) is the curvature calculated according to (39) for the radius r = r1" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0000347_2012-01-1810-Figure2-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0000347_2012-01-1810-Figure2-1.png", "caption": "Figure 2. Schematic of the leading and trailing linings.", "texts": [], "surrounding_texts": [ "The used finite element model herein consists of 5 main parts; the drum and the two shoes and two linings that have been created directly using the ANSYS package (FEM). Figure 1 indicates the meshed coupled drum, shoe and lining after appropriate simplification to the original parts. A solid 45 elements has been chosen from the package library to model the 3-D solid structure that has 8-nodes with three degree of freedom per each node. The drum consists of 11774 brick elements with 18123 nodes, however, the shoe consists of 1260 brick elements with 2666 nodes, and lining contains 2700 brick elements with 2886 nodes. Each lining covers an angle equal to 120\u00b0 of the drum ring that has an internal diameter of 340 mm as shown in Figure 1. However, the shoe and shoe rib cover an angle equal to 140\u00b0 of drum ring that has specifications as indicated clearly in Tables 1 and 2. A drum brake used for commercial vehicles is chosen as an example for the experimental work and then verification of the theoretical work with the tested results. Eight accelerometers are mounted along the drum that is hit by an impulse hummer. The tested signal is then fed from the accelerometers to a FFT for further analysis. The natural frequencies and modes of the drum and lining were collected in two cases. The first case was a free-free drum and free-free lining however the second case was the coupled drum-lining with a hydraulic pressure. The collected data has been analyzed in both cases and then the finite element model is being adjusted to control the difference between the experimental results and theoretical results. Each component's FE model is refined and adjusted to make the analytical results close to experimental modal analysis results [24,25,26,27,28]. An accurate representation of the component models as well as the statically coupled model is important for good correspondence between experimental squeal characteristics and those in simulations because the brake system's propensity to squeal is very sensitive to the geometry of the system and the material properties. Then the natural frequencies and mass-normalized mode shapes for each component are extracted from the modal analysis of the FE models. These modal characteristics of the components are used to replace the FE models to form the coupled system, and the total degrees of freedom are greatly reduced. To ensure the accuracy of the modal representations of the components and the convergence of the stability analysis results, the upper cut-off frequency for individual component modes was selected at least as high as twice of the squeal frequency of interest. It was found that for a system model constructed in this manner, the frequencies of the statically coupled drum-shoes-lining system exhibited convergence in the range of interest. The boundary conditions have been applied to the drum and shoe, and apply the appropriate solve to the model. Many trials were made to adjust the meshing elements for the drum and shoe with the appropriate number of elements. The lowest difference between the experimental work, and predicted FE model has been achieved in models shown in Figures 1 and 2. It was found a maximum difference of \u00b1 3% between the experimental and FEM results for the drum and \u00b1 2.5 % for the shoe with lining. So, this difference in both cases seems to be acceptable to carry on with these models as in Table 3. As it is well-known that there is a small gap between the drum and the two shoes during the rotation of the wheel. However, this gap becomes zero at the full contact between the drum and the shoe. The coupling between the brake shoe and lining assembly and the brake drum is made by the contact between them when the brake is actuated and friction force between the lining and drum is generated. The coupling can be regarded as a \u201ccontact stiffness\u201d modeled by springs connecting the brake shoe assembly and the brake drum [22]. However, the model proposed here expects that the degree of coupling, and thus the contact spring stiffness, will be determined by the contact force between the brake lining and brake drum. This means that the contact stiffness over the whole contact area is dependent not only on the brake force applied, but also on the friction interface pressure distribution. The contact stiffness will therefore vary around the contact surface, being higher as the local contact pressure increases. The coupling between the lining and drum can be represented by springs whose stiffnesses represent the local interface contact pressure and the brake shoe can thus be modeled as being coupled to the drum via two springs, one representing the contact stiffness Kcontact and one representing the brake lining dynamic stiffness Klining [23]. These two springs can then be combined to give a single \u201ccoupling\u201d spring whose stiffness is (1) Figure 3 shows the APDL (Ansys Parametric Design Language) flowchart of a contact analysis of drum brake assembly system. This APDL includes 3 main stages such as preprocessor, solutions and post-processors. The preprocessors step include construction of FE models of drum brake assembly, parametric meshing generation, parametric material definition, parametric boundary conditions, parametric analysis options, parametric solving and parametric post-processing (35). The post-processing step includes the summary, reading and plotting the of required results. Figures 1 and 2 show the used drum brake that contain the main 5 components which are the drum, the two brake shoes (leading and trailing) and the two lining (leading and trailing), participate in the vibrational response of a drum brake system [26-27]. The attached linings to the shoes will be in contact with the drum during braking to produce the friction forces. These leading and trailing shoes can be moved in different direction opposed to each other through hydraulic cylinders which contain two pistons to assist the shoes in the braking action. It is well-known that the friction-induced vibration is generated by the stiffness and friction coupling between the drum and the shoes through the shoe lining. A simplified coupled model that includes the drum, the shoes, and the shoe lining has been modeled to incorporate the effect of different boundary conditions on the occurrence of squeal. Finite Element models are built-up by ANSYS package for the five components of drum brake using 3-dimensional brick elements called SOLID45 [29]. The drum is clamped in the bolt hole positions while the shoe model uses free boundary conditions. To include the inertial and stiffness influences of the shoe lining on modal characteristics (eigenvalues and eigenvectors) of the shoe, the shoe lining is modeled as an integral part of the FE model of the shoe. The equations of motion of the uncoupled system including one drum and two identical shoes can be written as [23, 26, 27 and 30]: (2) Where [\u03c92] is a diagonal matrix of the extracted N natural frequencies of the components, and {q} is an N-vector of generalized coordinates. However, the number of degrees of freedom of the system, N, is equal to the total number of extracted component modes [26]. In considering the coupling between the drum and the shoe through the contact lining, the contact interface between the drum and shoe is discretized into a mesh of 2-dimensional contact elements. The lining material is then modeled as a spring located at the contact elements. So, the equation of motion of the coupled system is as follows [26-27]; (3) Where, [A] and [C] are stiffness contribution due to the lining and shoe supports respectively and [B] arises from friction coupling between the shoes and drum which is asymmetric. In the absence of lining coupling i.e., [A] and \u03bc equal to zero, the eigenvalues are purely imaginary that is being the natural frequencies of the drum components and the shoes coupled through the hydraulic cylinder stiffness and backing plate stiffness. The solution of the Equation 3 gives the eigenvalues of: (4) However; in the presence of the lining stiffness coupling but without friction coupling, the eigenvalues are again purely imaginary and correspond to the natural frequencies of an engaged brake system which is not rotating. This is referred to what is called statically coupled system. In the presence of the lining stiffness coupling but with friction coupling and the [B] is non-symmetric. When all of the eigenvalue are purely imaginary, these correspond to the natural frequencies of an engaged and rotating system. If any of the eigenvalues is complex, it will appear in the form of complex conjugate pairs, one with positive real part and the other with negative real part. The existence of complex roots with positive real parts indicates the presence of mode merging or what is called coupled mode, instability, which causes the brake to squeal. The value of friction coefficient that demarcates stable and unstable oscillations will be referred as a critical value of friction coefficient \u03bccr. The imaginary part of the eigenvalues with a doublet root at this \u03bccr is the squeal frequency and the corresponding mode of the complex structure is the mode shape at this squeal frequency, [26-27]." ] }, { "image_filename": "designv11_62_0001154_1.3552524-Figure4-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0001154_1.3552524-Figure4-1.png", "caption": "FIGURE 4. Schematic illustration showing the hot tear region before and after solidification.", "texts": [ "% and the crack length and consequently hot tear susceptibility. In the model of choice, the main reason of hot tearing is the restriction of solidification contraction by the ceramic shell mold. If the ceramic mold has a good collapsibility at high temperatures, the middle part of the casting is assumed to be in free contraction. Therefore, the casting obtained under no constraints is sound. However, hot tearing occurred at the hot spot under the constraint of a rigid ceramic mold to be placed between the two sides of the casting. Figure 4 schematically shows the interconnection between shell mold and casting during solidification. As the thinner section cools below the temperature where it develops strength, it begins to contract. It has been shown that hot tears occur in weak areas where the strains resulting from contraction are concentrated. When the molten metal can no longer mass-feed toward the hot spot, the contraction strains pull the solid dendrites apart at this point, and hot tears occur. Figure 5 schematically shows the effect of restraint on the final dimension" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0001039_detc2011-48755-Figure1-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0001039_detc2011-48755-Figure1-1.png", "caption": "Figure 1. Quadrotor helicopter configuration with Roll-PitchYaw Euler angles [\u03c6, \u03b8, \u03c8] [4]", "texts": [ " There are four fixed-pitch-angle blades whereas single-rotor helicopters have variable-pitchangle blades. The control of a quadrotor helicopter is performed by varying the speed of each rotor. The configuration and structure of a quadrotor with related hardware/software, especially the Quanser quadrotor (Qball-X4) which was developed in collaboration between Concordia University and Quanser Inc. through an NSERC (Natural Sciences and Engineering Research Council of Canada) Strategic Project Grant (SPG), is presented in the next parts of this paper. In Fig. 1 the concept of a quadrotor helicopter is shown. Each rotor produces a lift force and moment. The two pairs of rotors, i.e., rotors (1, 3) and rotors (2, 4) rotate in opposite directions so as to cancel the moment produced by the other pair. To make a roll angle (\u03c6) along the x-axis of the body frame, one can increase the angular velocity of rotor (2) and decrease the angular velocity of rotor (4) while keeping the whole thrust constant. Likewise, the angular velocity of rotor (3) is increased and the angular velocity of rotor (1) is decreased to produce a pitch angle (\u03b8) along the y-axis of the body frame", "org/about-asme/terms-of-use 4 Copyright \u00a9 2011 by ASME We can present equivalently, x = [x, y, z, \u03c6, \u03b8, \u03c8] and u = [ 1u , 2u , 3u , 4u ] in the vector form as: (12) 00 00 ( ) ( ) , ( ) , 0 0 0 0 0 0 r r g zg f x f x (13) and to define g(x) as follows: cos sin cos sin sin 0 0 0 cos sin sin sin cos 0 0 0 cos cos 0 0 0 ( ) 0 0 0 0 0 0 0 0 0 1 g x l l (14) This section describes the dynamic model of the Qball-X4. The nonlinear models as well as linearized models for the use of controllers development are described. For the following discussion, the axes of the Qball-X4 vehicle are denoted as (x, y, z) and are defined with respect to the vehicle as shown in Figure 1. Roll, pitch, and yaw are defined as the angles of rotation about the x, y, and z axis, respectively. The global workspace axes are denoted as (X, Y, Z) and are defined with the same orientation as the Qball-X4 sitting upright on the ground. 4.1. Actuator Dynamics The thrust generated by each propeller is modeled using the following first-order system model F k u s (15) where u, is the PWM input to the actuator, \u03c9 is the actuator bandwidth and k is a positive gain. These parameters were calculated and verified through experimental studies" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0001123_gt2012-68510-Figure10-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0001123_gt2012-68510-Figure10-1.png", "caption": "Figure 10: Short stack of actual recuperator geometry successfully laminated and sintered using isopropyl alcohol to adhere the layers.", "texts": [ " With the small dimensions of the recuperator, it would be unreasonable to individually place something in each channel before lamination and remove it afterwards. The second lamination method that used a solvent to bond the layers at room temperature had more success. 100 300 500 35 40 45 50 Temperature [C] M a s s [ m g ] -5 0 5 10 T e m p e ra tu re D if fe re n c e [ C ] 5 Copyright \u00a9 2012 by ASME Downloaded From: http://proceedings.asmedigitalcollection.asme.org/ on 01/27/2016 Terms of Use: http://www.asme.org/about-asme/terms-of-use Figure 10 and Figure 11 show samples laminated using isopropyl alcohol to adhere the layers. The layers bonded together well and the micro channels remained straight in general. It was important to minimize the amount of isopropyl alcohol to prevent the piece from deforming. Initially the solvent was applied with an eyedropper, which resulted in excessive amounts of the solvent. The extra solvent caused samples to lose shape and deform while still in the unfired state. To remedy this situation, solvent was applied in a fine mist that dried quickly after the layers were placed on one another" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0000720_amm.307.304-Figure7-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0000720_amm.307.304-Figure7-1.png", "caption": "Figure 7 Hertzian stresses along face width of gear in ABAQUS simulation", "texts": [ " The load distribution as well as the radii of curvature varies along the face width, which result in a constant Hertzian stress along the face width of gear while the Hertzian stress values are variable along the line of action (LOA). As long as double pairs of teeth are in contact the total load is shared between teeth. When the total load is transmitted by engagement of a single pair of teeth, the Hertzian stress has its maximum value. A comparison between present model and a finite element model for a pair of bevel gears has been implemented. One pair of teeth has been modeled using ABAQUS software and the resulted Hertzian stress distribution is shown in Fig.7, which is fairly close to the results achieved using the equivalent gears model shown in Fig.6. [1] E.Buckingham: Analytical Mechanics of Gears (Dover, New York 1949). [2] S.J. Nalluveettil and G.Muthuveerappan: Pergamon Press, Vol. 48, No. 4, pp. 13, 1992. [3] A.H. Elkholy, A.A.Elsharkawy and A.S.Yigit: Mechanics of Structures and Machines vol. 26, pp. 41-61, 2007. [4] V.G. Pavlov: Journal of Machinery Manufacture and Reliability, vol. 40, pp. 443-449, 2011. [5] R. Gohar: Elastohydrodynamics (McGraw\u2013Hill, london, 2001)" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0000189_amr.753-755.1680-Figure4-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0000189_amr.753-755.1680-Figure4-1.png", "caption": "Fig. 4. Ball screw spiral curve Fig.5. Rectangular coordinate system Fig. 6 . (U, V, W) coordinate system", "texts": [ " ( ) 2 2 2 31 1 23 1 2 1 13 2 J P E E \u00b5 \u00b5 \u03b4 \u03c1 \u03b1\u03c0 \u2212 \u2212 = + \u2211 (6) Here, J determined by the elliptic integral of \u03c4 ;\u03b1 is the load distribution coefficient of screw or nut side: can take s\u03b1 or n\u03b1 [2]. When under the effect of axial load, the surface of ball contact with nut and screw occur normal deformations sp\u03b4 , np\u03b4 . So the normal deformation between the Nut raceway surface and the screw raceway surface can be calculated by Eq.7. sp np\u03b4 \u03b4 \u03b4= + (7) Considering the helix angle, in Fig.3, the axial deformation a\u03b4 that screw relative to nut can be calculated by Eq.8. And in Fig.4, the number z of working ball can be calculated by Eq.9. a cos sin \u03b4 \u03bb \u03b4 \u03b2 = (8) 0 cosb d z i d \u03c0 \u03bb = (9) Here, i is the product of circle\u2019s number and columns' number of working ball on the nut. According to Eq.9, in the ideal state, under the effect of the axial load F , the normal load P of each ball is equal, can be calculated by Eq.10. sin cos FP z \u03b2 \u03bb= (10) Because, 1 2 210E E= = (G pa), 1 2 0.3\u00b5 \u00b5= = ,The initial contact angle 45\u03b2 = , combine Eq.7, Eq.8, Eq.9, and Eq.10, we can get the axial deformation of ball screw as Eq" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0003866_112015-Figure1-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0003866_112015-Figure1-1.png", "caption": "Figure 1. Calculated motor grader scheme", "texts": [ " This will allow the prospects for further use of the proposed structural solutions to be evaluated, such as simultaneous energy accumulation and dynamic vibration damping. 2. The description of the simulation model for determining energy losses in the working equipment suspension In order to determine the amount of energy that dissipates when damping oscillations of the working equipment, a simulation mathematical model of a motor grader was developed using MATLAB Simscape Multibody system [21, 22, 23]. The calculated scheme of the mechanical part of the mathematical model is presented in figure 1. The mechanical subsystem of the motor grader includes 5 elements. These are chassis base frame (ridge beam with cab and sub-frame) with m1 mass, front balancing bridge with m2 mass, the left balance beam of rear balancing cart with m3 mass, the right balance beam of rear balancing cart with m4 mass, and the working equipment (traction frame with moldboard) with m5 mass. During the development of a mathematical model of the motor grader, the following assumptions were taken into account [3, 25]: - the motor grader is a hinged articulated multi-unit; - the links imposed on the dynamic system of the motor grader are holonomic and stationary; - elements of metal structures are described as absolutely rigid rods; - the mass-inertial properties of the elements of metal structures of the motor grader are characterized by masses, coordinates of the mass centers, moments of inertia and centrifugal moments of inertia; Mechanical Science and Technology Update IOP Conf" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0000940_msf.735.338-Figure4-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0000940_msf.735.338-Figure4-1.png", "caption": "Fig. 4 Titanium sheet metal heat shield assemblies containing SPF and SPF/DB components", "texts": [ " This geometry is used to provide more thermal protection for the structure above the engines but no longer allows a simple bent piece of sheet metal to act as the part. When heat shields were being developed for a twin aisle airplane, castings were the initial choice. However, the benefits from using sheet metal on the single aisle program convinced the designers to pursue a sheet metal design. Assemblies of sheet metals details are used to produce this structure. These assemblies contain single sheet SPF details as well as two-sheet SPF/DB components as shown in Fig. 4. A portion of these details are produced using a fine grain version of the 6Al-4V alloy which is described in more detail in the next section. Fabrication of SPF Details. The SPF components for the assembly were fabricated in multi-cavity dies in order to get the best utilization of the starting material. The parts to be nested together were chosen so the amount of strain would be similar and the pressure cycle would form all the parts at close to the same strain rate. Representative sheets of details produced using the fine grain 6Al-4V material is shown in Fig" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0002541_rnc.3892-Figure2-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0002541_rnc.3892-Figure2-1.png", "caption": "FIGURE 2 Reference frame of vertical takeoff and landing", "texts": [], "surrounding_texts": [ "In this section, a nonlinear nonminimum system-VTOL aircraft is provided to illustrate the effectiveness of the proposed method in the previous sections. The nonlinear model of VTOL has been introduced in Su and Lin.11 The body-fixed reference frame is shown in Figure (2) and the dynamics of VTOL can be written as\u239b\u239c\u239c\u239c\u239c\u239c\u239d x\u03071 x\u03072 x\u03073 x\u03074 x\u03075 x\u03076 \u239e\u239f\u239f\u239f\u239f\u239f\u23a0 = \u239b\u239c\u239c\u239c\u239c\u239c\u239d x2 0 x4 \u22121 x6 0 \u239e\u239f\u239f\u239f\u239f\u239f\u23a0 + \u239b\u239c\u239c\u239c\u239c\u239c\u239d 0 0 \u2212 sin x5 \ud835\udf00 cos x5 0 0 cos x5 \ud835\udf00 sin x5 0 0 0 1 \u239e\u239f\u239f\u239f\u239f\u239f\u23a0 \u0393(u) = f (x) + g(x)\u0393(u). (13) In the nonlinear dynamics, x1 is the horizontal position y, and x3 is the vertical position z. x2, x4, and x6 are the corresponding velocities of the aircraft, and x5 is the roll angle. \u201c\u22121\u201d is the gravitational acceleration. \ud835\udf00 is a coefficient, which represents the coupling between the rolling moment and the lateral acceleration. The output of the plant is y1 = x1, y2 = x3. The control inputs are the thrust u1 and the rolling moment u2. The input nonlinearity model of VTOL is chose as following: \u0393(u1) = { (u1 \u2212 0.1) (1 + 0.2 \u00d7 sin u1), u1 > 0.1 0, \u22120.1 \u2a7d u1 \u2a7d 0.1 (u1 + 0.1) (1 + 0.2 \u00d7 cos u1), u1 < \u22120.1 (14) \u0393(u2) = { (u2 \u2212 0.3) (1 + 0.1 \u00d7 sin u2), u2 > 0.3 0, \u22120.3 \u2a7d u2 \u2a7d 0.3 (u2 + 0.3) (1 + 0.1 \u00d7 cos u2), u2 < \u22120.3. (15) Then ( \ud835\udefc+1 , \ud835\udefd+1 , \ud835\udefc\u22121 , \ud835\udefd\u22121 ) = (0.8, 1.2, 0.8, 1.2), ( \ud835\udefc+2 , \ud835\udefd+2 , \ud835\udefc\u22122 , \ud835\udefd\u22122 ) = (0.7, 1.3, 0.7, 1.3), \ud835\udf07 = 0.2. The output of the plant is (x1 x3)T , and the nonlinear dynamics (13) is a general form, the internal states are \ud835\udf02 = [ x5 x6 ] , and F(x) = [ 0 \u22121 ] ,G(x) = [ \u2212 sin x5 \ud835\udf00 cos x5 cos x5 \ud835\udf00 sin x5 ] , s(\u03c2, \ud835\udf02, (F(x) + G(x)\u0393(u))) = [ x6 sin x5 \ud835\udf00 ] + [ 0 0 cos x5 \ud835\udf00 sin x5 \ud835\udf00 ] (F(x) + G(x)\u0393(u)) . (16) The details of the transformation can be found in Su and Lin.11 From the analysis of Su and Lin,11 the internal dynamic (16) is unstable, so the plant of VTOL is a MIMO nonlinear nonminimum phase system. The IID of (16) can be constructed by stable inversion or stable system center method. In this section, IID is computed by stable system center method proposed in Shtessel et al.20 For computing IID, \ud835\udf00 is set to be \ud835\udf00 = 0.5, and according to the stable system center method proposed in Shtessel et al,20 the IID \ud835\udf02d = (\ud835\udf02d 1 , \ud835\udf02 d 2 ) T \u225c (xd 5, x d 6) T can be got by solving the following equation: ?\u0308?d + c1?\u0307? d + c0\ud835\udf02 d = \u2212 ( P1?\u0307? d + P0\ud835\udf18 d) , where c0 = 1, c1 = 2, \ud835\udf18d = (0, 1)T ( y\u0308d 1 cos \ud835\udf02d 1 + ( 1 + y\u0308d 2 ) sin \ud835\udf02d 1 \u2212 \ud835\udf02d 1 ) , P0 and P1 are power matrices. The details can be found in Su and Lin.11 In this simulation, the desired reference trajectories are chosen as yd 1 = R cos(\ud835\udf14t), yd 2 = R sin(\ud835\udf14t), where R = 1, \ud835\udf14 = 0.1. Then the desired trajectories and the IID of the plant are shown in Figure 3. After getting IID, define the tracking error as e1 = y1 \u2212 yd 1 = x1 \u2212 xd 1, e2 = y\u03071 \u2212 yd 2 = x2 \u2212 x\u0307d 1, e3 = y2 \u2212 yd 3 = x3 \u2212 xd 2, e4 = y\u03072 \u2212 yd 4 = x4 \u2212 x\u0307d 2, e5 = \ud835\udf021 \u2212 \ud835\udf02d 1 = x5 \u2212 xd 5, e6 = \ud835\udf022 \u2212 \ud835\udf02d 2 = x6 \u2212 xd 6, then the states tracking model can be built\u239b\u239c\u239c\u239c\u239c\u239c\u239d e\u03071 e\u03072 e\u03073 e\u03074 e\u03075 e\u03076 \u239e\u239f\u239f\u239f\u239f\u239f\u23a0 = \u23a1\u23a2\u23a2\u23a2\u23a2\u23a2\u23a3 0 1 0 0 0 0 0 0 0 0 0 0 0 0 0 1 0 0 0 0 0 0 0 0 0 0 0 0 0 1 0 0 0 0 cos x5 \ud835\udf00 0 \u23a4\u23a5\u23a5\u23a5\u23a5\u23a5\u23a6 \u239b\u239c\u239c\u239c\u239c\u239c\u239d e1 e2 e3 e4 e5 e6 \u239e\u239f\u239f\u239f\u239f\u239f\u23a0 + \u23a1\u23a2\u23a2\u23a2\u23a2\u23a2\u23a3 0 0 1 0 0 0 0 1 0 0 cos x5 \ud835\udf00 sin x5 \ud835\udf00 \u23a4\u23a5\u23a5\u23a5\u23a5\u23a5\u23a6 ( F(x) + G(x)\u0393(u) \u2212 [ yd(2) 1 yd(2) 2 ]T ) = Ae + B ( F(x) + G(x)\u0393(u) \u2212 [ yd(2) 1 yd(2) 2 ]T ) . The initial condition x(0) is x(0) = [1.5, 0,\u22120.5, 0.2, 0.28, 0]T . According to the control method proposed in Section 3, suppose \ud835\udf00 = 0.01, by solving LMI (9), the controller gain matrix can be obtained, K = [ 32.39 92.4 23.4 36.5 \u2212102.67 \u221268.09 23.34 36.17 \u221242.59 \u221224.31 \u221229.10 \u221219.3 ] . The adaptive law parameters are set to be q1 = 1, q2 = 10. To compare the tracking performance, a traditional IID-based controller, which does not consider the input nonlinearity of the system, is utilized here for the nonlinear simulation. The traditional controller is ud = 1 G(x) [ \u2212F(x) + yd(r) + Kde ] , where Kd is obtained by LQR methods proposed in Hu et al.10 Firstly, ud is carried on the nonlinear model of VTOL without input nonlinearity. The tracking responses are depicted in Figures 4 to 5, where Figure 4 is the tracking performance and the solid line is the given reference trajectory, and dashed line is the output of the plant under controller ud. Figure 5 is the input of the plant. The tracking errors are small and the input of the plant are smooth, so the traditional controller ud can achieve good performance for nonlinear model of VTOL without input nonlinearity. Secondly, ud is carried on the nonlinear model of VTOL with input nonlinearity (14) to (15). The tracking responses are depicted in Figures 6 to 7. From Figure 6, we can easily see that the tracking error is bigger than the input nonlinearity free case. Thirdly, the proposed adaptive controller uad is carried on the nonlinear model of VTOL with input nonlinearity (14) to (15). To testify the advantage of the proposed adaptive controller, 2 fixed parameter controllers with the form of (10) are adopted here. Controller I, which is marked as u0.01, the value of ?\u0302?0 and ?\u0302?1 in (10) are chosen as ?\u0302?0 = ?\u0302?1 = 0.01. Controller II, which is marked as u0.4, the value of ?\u0302?0 and ?\u0302?1 in (10) are chosen as ?\u0302?0 = ?\u0302?1 = 0.4. u0.01 and u0.4 are all carried on the nonlinear model of VTOL with input nonlinearity, and the simulation results are shown in Figures 8 to 9, where dashdot line is the given reference trajectory, solid line is the output under controller uad, dashed line is the output under controller u0.4, and dotted line is the output under controller u0.01. The inputs of the plant and the estimating results of \ud835\udf0c0 and \ud835\udf0c1 are also given in Figures 10 and 11. Noticing from Figures 8 to 9, tracking error of uad, u0.01, and u0.4 are all accepted small, and tracking performances are better than ud. What is more, the tracking performance of uad is better than u0.01 and u0.4. Through the comparison simulation results, effectiveness of the proposed control methods can be validated." ] }, { "image_filename": "designv11_62_0002755_icems.2017.8056505-Figure3-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0002755_icems.2017.8056505-Figure3-1.png", "caption": "Fig. 3 Magnetic flux of basic study model", "texts": [ " However, since reverse magnetic field component Hm, which acts on the magnetization direction of the parallelogram magnet, is smaller than vector H of the reverse magnetic field that is acting on the rectangular parallelepiped magnet (Eq. (4)), improvement of the demagnetizing field is expected. \uff46 \u03c3 AmLg LmAg Pc= (1) Lm=tm/(cos \u03b1) (2) Am = Wm cos \u03b1 (3) Hm = H cos \u03b1 (4) As verification of the oblique orientation, we examined the relation between the demagnetization resistance and the magnetic flux amount with a basic study model. Fig. 3 shows a 1/2 anti-periodic basic study model, and TABLE I shows the analysis conditions. The rotor has two poles, and a flat plate permanent magnet is embedded in the range of the mechanical angle of 90-degree of the rotor core. The stator is arranged outside of the rotor by an air gap, and the winding is wrapped around the stator tooth and a mechanical angle of 90 degrees at the stator core\u2019s tip. The parallelogram magnet with the changed magnetic field orientation angle is incorporated into the basic study model (Fig. 3). We applied a current that causes the reverse magnetic field to act on the winding and the demagnetization state of the permanent magnet and obtained air-gap magnetic flux density before and after the energization by electromagnetic field analysis software (JMAG-Designer). The demagnetization ratio is shown in Eq. (5). 100 \u00d7 (1-B2/B1) [%] (5) Figure 4 shows the relationship between the air-gap magnetic flux density and the demagnetization limit current ratio for magnetic field orientation angle \u03b1" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0000321_016918610x512631-Figure1-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0000321_016918610x512631-Figure1-1.png", "caption": "Figure 1. CEE made of a cylindrical flexible object from the beginning or a rolled blade spring.", "texts": [], "surrounding_texts": [ "Vertebrate animals such as humans are characterized by a body consisting of muscles, tendons and skeletons. They pull the tendons by actuation of the muscles to move the joints. Changing the internal force among muscles enables variation of joint stiffness so that vertebrate animals can control joint stiffness according to the * To whom correspondence should be addressed. E-mail: kino@fit.ac.jp \u00a9 Koninklijke Brill NV, Leiden and The Robotics Society of Japan, 2010 DOI:10.1163/016918610X512631 D ow nl oa de d by [ U ni ve rs ity o f C on ne ct ic ut ] at 1 1: 45 1 1 Ja nu ar y 20 15 1640 H. Kino, D. Nakiri / Advanced Robotics 24 (2010) 1639\u20131660 situation. A tendon-driven manipulator is modeled on the musculoskeletal structure of vertebrate animals. This system uses wire cables (hereinafter, wires) and actuators to drive the joints instead of using living tendons and muscles [1\u20136]. The characteristic of this manipulator is that the mass/inertia can be reduced; high power and safety can be readily guaranteed because actuators are fixed on the base. Wires can only generate tension. Therefore, redundant actuation is necessary for cases in which tensioners are not used. The redundancy causes the internal force among wires. Therefore, by installing some devices among wires, the tendon manipulator enables variation of joint stiffness through control of the internal force. To date, many devices to control joint stiffness have been developed for tendondriven manipulators. For instance, an elastic pulley is one solution. Usually, such an elastic pulley is built in an actuator unit. Kobayashi and co-workers investigated a nonlinear spring tensioning device (NST) [7, 8] that consists of some springs and some pulleys. They then developed tendon manipulators using them. Koganezawa et al. developed nonlinear elastic systems (NLES) [9] and an actuator unit with a NLES (ANLES) [10] consisting of torsion springs and a ball\u2013screw structure. These devices can change the joint stiffness drastically based on elastic deformation of the springs. However, such devices generally require many mechanical parts, such as guide pulleys, shafts, springs and bearings. The increase of such mechanical parts increases the mass/inertia and mechanical friction, which collectively degrade the positioning performance of the manipulator. To make matters worse, it is difficult for such devices to change the inherent stiffness, because they are intricately built in bodies of manipulators or actuator units so that replacement is difficult. To overcome the these difficulties, we propose a novel mechanism \u2014 a cylindrical elastic element (CEE). The CEE is made of a cylindrical flexible object or rolling of a blade spring, as portrayed in Figs 1 and 2. For control of joint stiffness, CEEs are inserted into the routes of wires in a tendon-driven manipulator (Fig. 3). The top and the bottom of a CEE are fixed, respectively, on a wire and a base. Figures 3 and 4 show that the CEE can transmit wire tension by transforming itself into an elliptic shape. The variation of wire directions based on the deformation can change the joint stiffness depending on internal forces among wires. This system offers many advantages, such as lighter weight, simpler structure, lower cost, fewer mechanical parts and lower friction. Furthermore, it is easy to replace in order to vary the characteristic of the CEE\u2019s stiffness. Our ultimate goal is the development of a tendon-driven manipulator with CEEs, that has multi-d.o.f., such as 6 or 7 d.o.f. like a human arm. In the manipula- D ow nl oa de d by [ U ni ve rs ity o f C on ne ct ic ut ] at 1 1: 45 1 1 Ja nu ar y 20 15 H. Kino, D. Nakiri / Advanced Robotics 24 (2010) 1639\u20131660 1641 tor system with CEEs, however, the relation between wire lengths and joint angles/stiffness has not been clarified because the deformation of CEEs is intricately changed depending on the wire tension. Consequently, the uncertainty of deformation complicates the control of both joint angles and stiffness for the manipulators Figure 2. Photograph of a CEE (65 mm external diameter, 30 mm width, 3 g, polyethylene terephthalate). Figure 3. Overview of a one-link manipulator with two CEEs. The stiffness of CEEs is much lower than that of the wires. In (a), the joint stiffness is low because it is dominated by the stiffness of CEEs. On the other hand, in (b), the joint stiffness becomes high because it is denominated by the wire\u2019s stiffness. D ow nl oa de d by [ U ni ve rs ity o f C on ne ct ic ut ] at 1 1: 45 1 1 Ja nu ar y 20 15 1642 H. Kino, D. Nakiri / Advanced Robotics 24 (2010) 1639\u20131660 Figure 4. Shape deformation of CEEs depending on wire tension. using CEEs. As the first step, this paper aims to establish the solving method of kinematics of a one-link manipulator with CEEs. It is very important to model a CEE mathematically and analyze the deformation for solving the kinematics. To analyze the deformation, the finite element method (FEM) might be one of the useful methods. However, the FEM usually requires much time to calculate in the case using a lot of CEEs, even though it has high precision. Therefore, it is very important to alternatively establish the CEE model characterized in both applicable precision and also less time. The novelty and originality in this paper are (i) a proposal of the tendon-driven manipulator equipped with CEEs to control the joint stiffness, (ii) a proposal of approximate models of a CEE to analyze the deformation and their evaluation, and (iii) establishment of a technique to solve the forward/inverse kinematics of the manipulator. This paper first presents characteristics of a CEE in Section 2. Next, in Section 3, three approximate models of a CEE are proposed using link\u2013spring\u2013 damper elements to analyze the CEE\u2019s deformation to solve the kinematics. The three models\u2019 reliability is examined through simulations and experiments; then the most useful model is selected. Using the approximate model, we specifically examine the numerically kinematic solving framework of a one-link manipulator equipped with two CEEs. In Section 4, we propose a numerical solving method for the forward kinematics, which is the calculation of the joint angle and its stiffness from wire length. Based on numerical forward kinematics, we expand it to derive the inverse kinematics to obtain the wire length from the joint angle and stiffness in Section 5. The precision of the proposed methods is discussed through comparison between experimental results and numerical results. Finally, we conclude our investigation in Section 6." ] }, { "image_filename": "designv11_62_0003064_ssci.2017.8280822-Figure4-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0003064_ssci.2017.8280822-Figure4-1.png", "caption": "Fig. 4. Map and compass operator model of PIO", "texts": [ " The rules are defined with the position Xi and the velocity Vi of pigeon i, and the positions and velocities in a D-dimension search space are updated in each iteration. The new position Xi and velocity Vi of pigeon i at the t-th iteration can be calculated by (7) and (8). V (t)= V (t-1).e +rand.( ( ) ( 1)) (7) ( ) ( 1) V (t) (8) Rt i i g i i i i X t X t X t X t where R is the map and compass factor, rand is a random number, and Xg is the current global best position, and which can be obtained by comparing all the positions among all the pigeons. As shown in Fig. 4, the best positions of all pigeons are guaranteed by using map and compass. Each pigeon can adjust its flying direction by following the pigeon which has best position according to (7). The former flying direction Xi(t-1), which has relation to Vi(t-1).e-Rt in (8). The vector sum of these two arrows is its next flying direction. landmark operator model is presented based on landmarks. In this operator, the number of pigeons is decreased by half in every generation. The pigeons are still far from the destination, and they are unfamiliar the landmarks" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0003913_s00773-019-00675-8-Figure2-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0003913_s00773-019-00675-8-Figure2-1.png", "caption": "Fig. 2 Multi-section model of the after stern tube bearing", "texts": [ ") and the displacement constraint at (1) y1_dsp z1_dsp y1_ang z1_ang y2_dsp z2_dsp y2_ang z2_ang ke = \u23a1\u23a2\u23a2\u23a2\u23a2\u23a2\u23a2\u23a2\u23a2\u23a2\u23a2\u23a2\u23a2\u23a3 12EIZ len3 0 0 6EIZ len2 \u2212 12EIZ len3 0 0 6EIZ len2 0 12EIy len3 6EIy len2 0 0 \u2212 12EIy len3 6EIy len3 0 0 6EIy len2 4EIy len 0 0 \u2212 4EIy len2 2EIy len 0 6EIZ len2 0 0 4EIz len \u2212 6EIz len2 0 0 2EIz len \u2212 12EIZ len3 0 0 \u2212 6EIz len2 12EIZ len3 0 0 \u2212 6EIZ len2 0 \u2212 12EIy len3 \u2212 4EIy len2 0 0 12EIy len3 12EIy len2 0 0 6EIy len3 2EIy len 0 0 12EIy len2 4EIy len 0 6EIZ len2 0 0 2EIz len \u2212 6EIZ len2 0 0 4EIy len \u23a4\u23a5\u23a5\u23a5\u23a5\u23a5\u23a5\u23a5\u23a5\u23a5\u23a5\u23a5\u23a5\u23a6 , (2)K = f , the bearing locations. The unique solution can be achieved by eliminating the appropriate rows and columns corresponding to the specified or \u201csupport\u201d degree of freedom. Given that the after stern tube bearing is long and the centerline of the journal has a deflection, the after tube bearing is divided into several bearing segments, as illustrated in Fig.\u00a02. The midpoint of the journal centerline of the bearing segment is used as the node of the beam element in the calculation model of shafting alignment. To calculate the oil film force of each bearing segment, the nodes are grouped in axial order, and each bearing segment corresponds to a group. Thus, the Y-component Fyi and z-component Fzi of the oil film force for each bearing segment can be obtained. Figure\u00a03 displays that under the action of oil film forces Fx and Fy, journal center position O2 works in an eccentric position relative to bearing center O1" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0003470_1.5096278-Figure1-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0003470_1.5096278-Figure1-1.png", "caption": "Fig. 1. Free body diagrams with and without friction: Standard model (left), without and with friction (middle and right).", "texts": [ " Solutions for arbitrary initial conditions are obtained by strict application of the fundamental principles of mechanics. The validity of these solutions is discussed in context with the loss of contact of the ladder with the wall. A phase space representation is used to visualize for which initial conditions the ladder stays in contact with the wall and at which angle it would lose contact. VC 2019 American Association of Physics Teachers. https://doi.org/10.1119/1.5096278 I. INTRODUCTION The sliding ladder problem relates to the system sketched in Fig. 1. It is the subject of many textbooks1\u20133 and journal papers. In 1985, it was noted by Freeman and Palffly-Muhoray4 that a sole treatment of the kinematics results in an infinite speed when the top of the ladder approaches the floor. After that further works on the statics and dynamics of a ladder in contact with a wall and the floor were published, namely: The statics of a ladder leaning against a rough wall with a vertical external load was studied by Mendelson in 1995.5 Mendelson showed that the system is statically indeterminate and took the axial compression of the ladder and Coulomb\u2019s law of static friction into account to obtain the stability conditions", " This allows for an understanding of which states the ladder loses contact with the wall, and which initial conditions may lead to an immediate loss of contact or to large critical angles. The falling ladder is modeled within the framework of planar rigid-body motion. The ladder is in contact with the wall and with the floor. The contact forces will be analyzed for the case that the ladder has an initial velocity, which may be imposed experimentally by a hammer stroke. The constraint that the ladder be in contact with both the wall and the floor allows us to describe the motion with a single degree of freedom. We choose to use the angle h, shown in Fig. 1. The position vectors xA and xB of the contact points read xA \u00bc xAex \u00bc \u2018 cos\u00f0h\u00deex; xB \u00bc yBey \u00bc \u2018 sin\u00f0h\u00deey; (1) where \u2018 denotes the length of the ladder. The length of the ladder is constant and hence jxB xAj2 \u20182 \u00bc 0, which is identically fulfilled by Eq. (1). For the dynamical analysis, the acceleration aC of the center of mass can be obtained from the position vector xC xC \u00bc \u2018 2 cos h\u00f0 \u00deex \u00fe sin h\u00f0 \u00deey ; (2a) vC \u00bc \u2018 2 _h sin h\u00f0 \u00deex \u00fe cos h\u00f0 \u00deey ; (2b) aC \u00bc \u2018 2 sin h\u00f0 \u00de\u20ach \u00fe cos h\u00f0 \u00de _h2 ex \u00fe cos h\u00f0 \u00de\u20ach sin h\u00f0 \u00de _h2 ey ; (2c) where _h and \u20ach represent the first and second derivatives of h with respect to time. The initial conditions can be expressed in terms of the h as well: h(t\u00bc t0)\u00bc h0 with 0< h0

>>>>>< >>>>>>: 9>>>>>>= >>>>>>; (21) and C\u0302 \u00bc d dt 1V\u03021 1N Xn\u00bcN n\u00bc2 d dt 1 nM\u0302 nV\u0302n n;n 1 ( ) (22) Consider an RCCC spatial four-bar mechanism which has a revolute input joint, two cylindrical intermediate joints, and a cylindrical output joint as shown in generalized form in Fig. 2 where input link 1 has twist a1\u00bc 72 deg and length a1\u00bc 40 mm, coupler link 2 has twist a2\u00bc 85 deg and length a2\u00bc 80 mm, output link 3 has twist a3\u00bc 110 deg and length a3\u00bc 120 mm, and the ground link 4 has twist a4\u00bc 150 deg and length a3\u00bc 170 mm. Suppose that the input speed is a constant 1000 rpm\u00bc 104.7198 rad/s so that 1V\u03021 14 \u00bc 0 0 104:7198 8< : 9= ; (23) 011005-2 / Vol. 4, JANUARY 2012 Transactions of the ASME Downloaded From: http://mechanismsrobotics.asmedigitalcollection.asme.org/ on 07/13/2014 Terms of Use: http://asme" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0003085_intellisys.2017.8324247-Figure9-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0003085_intellisys.2017.8324247-Figure9-1.png", "caption": "Fig. 9. The equilibrium position of bipedal robot with compliant leg modeled by using LMS AMESim.", "texts": [ " 8 that the stiffness coefficients of muscles don\u2019t have any effect on the because the muscles are internal forces and the is an external force. 972 | P a g e 978-1-5090-6435-9/17/$31.00 \u00a92017 IEEE 973 | P a g e 978-1-5090-6435-9/17/$31.00 \u00a92017 IEEE For simplicity and ease of comparison, the four springs use the same stiffness ( ), but four different values of have been used, as explained below. The equilibrium position of bipedal robot has been defined with the hip at the vertical of the ankle (same coordinate for the hip and ankle) [18] as shown in Fig. 9. In order to analyse the effect of springs on the work the motors have to supply, two cases have been simulated: Case 1: No preload of the spring muscles. This means that in the equilibrium position, the springs are unstretched/uncompressed. Case 2: Preload of TA and GAS in such a way to support the weight of the robot. The preload is realized by precompression of the springs in the equilibrium position. This means that in the equilibrium position, with the nominal stiffness chosen, the system is in static equilibrium" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0002074_nigercon.2013.6715667-Figure5-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0002074_nigercon.2013.6715667-Figure5-1.png", "caption": "Fig. 5. Variable speed wind t", "texts": [], "surrounding_texts": [ "2013 IEEE International Conference on Emer\nFrom above, superscript T is a transpos instantaneous input-phase to output-phase t the three-phase matrix converter. and and output phase voltage vectors, and a the input and output phase current vec equations 4 and 5, the output-line voltag currents can be expressed as\nWhere,\nging & Sustainable Technologies for Power & ICT in a Dev\n(2)\n(3)\ne and M is the ransfer matrix of\nare the input nd represent\ntors. Considering es and input-line\nIV. DESCRIPTION O\neloping Society (NIGERCON)\nF THE MODEL SYSTEM\nal Control System for DFIG\ntional Control System for DFIG", "2013 IEEE International Conference on Emer\nFig. 2 shows the DFIG system, where t matrix converter system replaces the co controller shown in Fig 3.\nV. BEHAVIOR OF THE WIND FARM FO VARIABLE SPEED WIND TURBINE S\nFigs 4 and 5 show respectively, a fixed sp\ninduction generator (IG) wind turbine and DFIG system. It is assumed that both machi together [36].\nThe proposed matrix converter system\nachieve the following behavior of the wind f composed of both fixed and variable sp system, during transient and dynamic conditi\nging & Sustainable Technologies for Power & ICT in a Dev\nhe new proposed nventional DFIG\nR FIXED AND YSTEMS\need squirrel cage a variable speed nes are connected\nmust be able to arm (Figs. 6 to 9)\need wind turbine ons.\ns\nurbine\neloping Society (NIGERCON)\nilable wind data [40]\nse of DFIG powers [40]\nng wind speed change [40]", "From Figs 6 to 9, the use of the DFIG variable speed wind turbine in conjunction with the squirrel cage wind turbine seems to show better performance it terms of wind farm stability.\nVI. CONCLUSION\nThe use of a matrix converter control topology for a double fed induction generator (DFIG) wind turbine has been proposed in this paper. This is because of its reduced cost, simple topology, absence of large DC link capacitor and other benefits. The proposed matrix converter for DFIG system was further suggested to stabilize a wind farm composed of DFIG and fixed speed squirrel cage induction generator system. The next stage of this paper is the detailed simulations and experimental validations of the proposed matrix converter control algorithm in comparison with results obtained using the conventional DFIG control system.\n[1] P. W. Wheeler and D. A. Grant, \u201cReducing the Semiconductor Losses in a Matrix Converter,\u201d Industrial Electronic Group, Department of Electrical and Electronic Engineering, University of Bristol, UK, 2007\n[2] M. Venturini, \u201cA New Sine Wave in, Sine Wave out Conversion Technique which Eliminates Reactive Elements\u201d, in; Proceedings of the POWERCON 7, 1980, pp. E3-1-E3-15.\n[3] L. Huber, and D. Borojevic, \u201cSpace Vector Modulated Three-phase Matrix Converter with Input Power Factor Correction,\u201d IEEE Trans. Industrial Application, vol. 31, no. 6, pp. 1234-1246, 1995.\n[4] S. Kwak, \u201cIndirect Matrix Converter drives for Unity Displacement Factor and Minimum Switching Losses,\u201d Electric Power Systems Research, vol. 77, pp. 447-454, 2006.\n[5] J. Lettl and S. Fligl, \u201cPWM Strategy Applied to Realize Matrix Converter System,\u201d PIERS Proceedings, Prague, Czech Republic, August 27-30, 2007.\n[6] J. K. Kang, H. Hara, A. M. Hava, E. Watanabe, and T. Kume, \u201cThe Matrix Converter drives Performance under Abnormal Input Voltage Conditions,\u201d IEEE Trans. on Power and Energy, vol. 17, no. 5, pp. 721- 730, 2002.\n[7] L. Gyugyi L and B. R. Pelly, Static Power Frequency Changers, Wiley, New York, 1976.\n[8] A. Alesina and M. Venturini, \u201cSolid State Power Conversion: A Fourier Analysis Approach to Generalized Transformer Synthesis,\u201d IEEE Transactions on Circuit Systems, vol. CAS-28, no. 4, pp. 319-330, 1981.\n[9] T. Matsuo, S. Bernet, R. Stephen Colby, and T. A. Lipo, \u201cModeling and Simulation of Matrix Converter/Induction Motor Drive,\u201d Mathematics and Computers in Simulation Journal, vol. 46, pp. 175-195, 1998.\n[10] A. Alesina and M. Venturini \u201cAnalysis and Design of OptimumAmplitude Nine-switch Direct ac-ac Converters,\u201d IEEE Trans. on Power Electronics, vol. 4, no. 1, pp. 101-112, 1989.\n[11] S. Sunter and J. C. Clare, \u201cA True Four Quadrant Matrix Converter Induction Motor Drive with Servo Performance,\u201d IEEE PESC \u201996, Baveno, Italy, pp. 146-151, 1996.\n[12] S. Sunter and S. Altas, \u201cPspice Modeling and Simulation of the SPIM Fed by 3-to-2 Phase matrix Converter,\u201d 2nd FAE International Symposium European University of Lefke, TRNC, pp. 451-456, 2002.\n[13] M. Bednar M, \u201cThe Precision Simulation of the First Generation Matrix Converter,\u201d Acta Polytechnica vol. 48, no. 3, pp. 66-70, 2008.\n[14] E. Babaei, S. H. Hosseini, and G. B. Gharehpetian, \u201cReduction of THD and Low Order Harmonics with Symmetrical Output Current for Singlephase ac/ac Matrix Converters. Electrical Power and Energy Systems, vol. 32, pp. 225-235, 2010.\n[15] H. M. Kojabadi, \u201cSimulation and Experimental Studies of Model Reference Adaptive System for Sensorless Induction Motor Drive,\u201d Simulation Modeling Practice and Theory, vol. 13, pp. 451-464, 2005.\n[16] S. Sunter, \u201cSlip Energy Recovery of a Rotor-side Field Oriented Controlled Wound Rotor Induction Motor Fed by Matrix Converter,\u201d Journal of the Franklin Institute, vol. 345, pp. 419-435, 2008.\n[17] S. Sunter and J. C. Clare, \u201cA True Four Quadrant Matrix Converter Induction Motor Drive with Servo Performance,\u201d IEEE-PESC, pp. 146- 151, 1996.\n[18] R. Datta and V. T. Ranganathan, \u201cA Simple Position-sensorless Algorithm for Rotor-side Field-oriented Control of Wound-rotor Induction Machine,\u201d IEEE Trans. Ind. Electronic, vol. 4, pp. 786-793, 2001.\n[19] H. Altun and S. Sunter, \u201cMatrix Converter Induction Motor Drive: Modeling, Simulation and Control,\u201d Electrical Engineering, vol. 86, no. 1, pp. 25-33, 2003.\n[20] H. Altun, \u201cApplication of a Matrix Converter in a Slip Energy Recovery Drive System,\u201d International Journal of Engineering Modeling, vol. 18, no 3-4, pp. 69-80, 2005.\n[21] A. Alesina and M. Venturini, \u201cIntrinsic Amplitude Limits and Optimum Design of 9-switches Direct PWM AC to AC Converter,\u201d Proceedings of PESC Conference Record, Pp. 242-252, 1980.\n[22] R. Beasant, W. Beatie, and A. Refsum, \u201cAn Approach to the Realization of a High Power Venturini Converter\u201d. IEEE Trans., pp. 291-297, 1990.\n[23] A. Zuckherberger, D. Weinstock and A. Alexandrovitz, \u201cSimulation of Three-phase Loaded Matrix Converter,\u201d IEE Proc. Trans. Electrical Power Application, vol. 143, pp. 294-300, 1996.\n[24] P. W. Wheeler, J. C. Clare, L. Empringham, \u201cA Vector Controlled MCT Matrix Converter Induction Motor Drive with Minimized Commutation Times and Enhanced Waveform Quality,\u201d In: Proceedings of the IEEEIAS Annual Meeting, vol. 1, pp. 466-471, 2002.\n[25] L. Wei, and T. A. Lipo, \u201cA Novel Matrix Converter Topology with Simple Commutation,\u201d In: Proceedings of the IEEE-IAS Annual Meeting, vol. 3, pp. 1749-1754, 2001.\n[26] S. Kwak and H. A. Toliyat, \u201cDevelopment of Modulation Strategy for Two-phase AC-AC Matrix Converter,\u201d IEEE Trans. Energy Conversion, vol. 20, no. 2, pp. 493-494, 2004.\n[27] S. Kwak, \u201cDesign and Analysis of Modern Three-phase AC/AC Power Converters for AC Drives and Utility Interfaces,\u201d PhD. Dissertation, Texas A & M University, 2005.\n[28] D. Vincenti, P. D. Ziogas, and R. V. Patel, \u201cAn Analysis and Design of a Forced Commutated Three-phase PWM AC Controllers with Input Unbalance Correction Capability,\u201d In Proceedings IEEE APEC Conference, pp. 487-493, 1992." ] }, { "image_filename": "designv11_62_0001140_s1672-6529(13)60226-7-Figure1-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0001140_s1672-6529(13)60226-7-Figure1-1.png", "caption": "Fig. 1 Rough terrain used in the experiment.", "texts": [ "2, the Chinese mitten crabs with body masses Mb ranging from 35 g to 40 g were selected to analyze velocity, gait patterns, and duty factors. In sections 4.3 and 4.4, all 18 crabs were used to discuss the mechanical energy of the MCs. Five types of artificial terrains with varying roughness, including one smooth terrain and four types of rough terrains, were constructed to study the gait patterns and kinematics of the Chinese mitten crab motion. Each rough terrain consisted of three layers: basement, concrete, and rough terrain surface (Fig. 1). The rough terrain surface was constructed using quartz sand with sizes obeying certain uniform distribution on the concrete with a random surface height distribution to ensure an irregular moving surface. Concrete pavement between the basement and quartz sand was used to fix and prevent the rolling of the quartz sand particles when the dactylopodite of the crab walked on the particles. Different sizes of quartz sand particles (average particle sizes dav = 5.0 mm, 7.0 mm, 12.0 mm and 24.0 mm) were chosen to prepare the rough terrain surfaces to impel the Chinese mitten crabs to clearly modify their locomotion patterns on different rough terrains" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0001607_0021998312463455-Figure12-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0001607_0021998312463455-Figure12-1.png", "caption": "Figure 12. Experiment setup for radial expansion measurement at blocking state.", "texts": [ " The geometry parameters of the actuators can be found in Table 2 in Appendix 2. Hydraulic swage fittings were attached at both ends of the actuator and then was clamped in an ADMET MTEST QuattroTM testing frame, which provides programmable force and displacement control. A high pressure ball-valve was mounted at each end, and a pressure transducer was mounted between the valves to measure the internal pressure. A fiber-optic displacement sensor measuring the actuator\u2019s radial expansion was mounted perpendicular to the actuator axis pointing to the axis as shown in Figure 12. A metalized polyester film was attached to the surface of the actuator to increase the reflection. In the experiments, the internal pressure was gradually increased by using a valve connected to a pressure source. An air pressure tank with a maximum pressure of 1.38MPa (200 psi) was used as the pressure source and provided ample pressure since most of the compaction occurs at low pressures (<100 kPa). The experiment was performed quasistatically to avoid dynamic effects in the results. Data from the pressure transducer, the displacement encoder and the fiber-optic displacement sensor were recorded simultaneously", " While these small variations during pressurization are undesirable, the trends in the experimental results do confirm the effects of compaction on the radial expansion in the blocked state. Tensile test experiments for actuators in the closed-valve scenario As indicated previously, the compaction can decrease the effective stiffness of the closed-valve actuator filled with a high bulk modulus working fluid. Tensile test experiments were carried out for water-filled actuators with the valve closed. In this scenario, the effective stiffness of the actuators is maximized. The experiment setup is the same as shown in Figure 12 except that there was no fiber-optic displacement sensor measuring radial displacement. In the experiment, the actuator was initially pressurized using a pump with both ends fixed. After pressurization, both valves were closed and tests began. The tests were performed quasi-statically. Axial forces, axial displacements, and internal pressures were recorded for each tensile test. Experimental results for actuators made of three types of fibers are shown in Figures 14 to 16 as discrete circles. Computational results with various conditions are also shown in the figures" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0002163_s10483-017-2182-6-Figure2-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0002163_s10483-017-2182-6-Figure2-1.png", "caption": "Fig. 2 Elastic ribbon with rectangular cross section", "texts": [ " The pitch angle \u03d5 is the key role to determine the configuration of the ribbon. It relies on the elastic properties of the material and the geometry of the cross section. Therefore, we examine the effects of the elastic properties and the geometry of the cross section on the pitch angle. For the analysis convenience, we choose two symmetrical cross sections, i.e., rectangle and ellipse, in the following text. The rigidity of structure can be expressed in terms of Young\u2019s modulus E or shear modulus G and the cross-sectional geometry. As shown in Fig. 2, let c1 and c2 be the side lengths of the rectangular cross section. Then, we have[45] B = 1 12 Ec1c 3 2, C = 1 3 Gc1c 3 2. Since G = E/2(1 + \u03bd), we can rewrite Eq. (23) as follows: \u03c6 = arctan ( 2 + 2 (1 + \u03b7)2 \u2212 4 1 + \u03bd )\u2212 1 2 , (24) where \u03bd is Poisson\u2019s ratio. Equation (24) shows that the pitch angle of the helical ribbon depends on the ratio of the intrinsic twisting \u03c90 3 to the twist rate \u03c930 and Poisson\u2019s ratio. Clearly, Eq. (24) requires \u03b7 > \u221a 5/2 \u2212 1. Figure 3 exhibits the change of the pitch angle \u03d5 with the ratio \u03b7" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0001396_icpec.2013.6527760-Figure4-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0001396_icpec.2013.6527760-Figure4-1.png", "caption": "Fig. 4 Distribution of flux lines at the unaligned position", "texts": [ " At this stage it is possible to visualize and manipulate the fields and to extract the machine parameters such as the flux linkages, inductances, induced voltages, torque developed etc., The entire SRM model is subdivided into triangular finite elements. The elements are defined such that the sides of the triangles coincide with the boundary of each material. Fig.2 shows subdivision of problem region. Fig.2 Subdivision of the 8/6 SRM By solving the above created model, the distribution of flux lines obtained by FEA in the aligned and unaligned positions are shown in Fig.3 and Fig.4. III. MATHEMATICAL MODEL OF SRM In order to model/analyze the machine, the SRM is represented by the voltage equation across one phase winding. Since the mutual coupling between phases for doubly salient machine is small, it is neglected[9] and the voltage equation is represented as dt idiRv ),( \u03b8\u03bb+= (1) The flux linkage \u03bb, is a function of phase current i and rotor position \u03b8. The phase winding resistance R cannot be ignored during low speed operation. The phase current may be low during initial excitation, but increases as inductance profile decreases" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0001036_s10015-013-0132-y-Figure2-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0001036_s10015-013-0132-y-Figure2-1.png", "caption": "Fig. 2 Analysis model of walking robot. Link 1: stance leg, Link 2: swing leg, Link 3: torso. a Side view, ml mass of the leg, mf mass of the foot, mh mass of the hip, mt mass of the torso, al length from the foot to the gravity center of the leg, at length from the hip to the gravity center of the torso, ll length of the leg, lt length of the torso, Jl inertia moment of the leg, Jt inertia moment of the torso. b Isometric view", "texts": [ " 1, which consists of three links; the torso, the swing and the stance legs. In an initial state, the robot rotates its swing leg in forward direction as in \u2018 ,\u2019 while the stance leg also rotates on the foot end at the same time. After moving forward of the swing leg as in \u2018`\u2019, both legs alternate their modes between \u2018swing\u2019 and \u2018stance\u2019 with each other as in \u2018\u00b4.\u2019 Then, the new swing leg that was the stance leg in the previous cycle starts its rotating motion. Alternating the swing and the stance modes between legs, the robot move forward resultantly. Figure 2 shows the analysis model of walking robot which consists of three links which are connected by the hip joints. Its physical parameters are given in Table 1. There are two motors at the hip joint for generating control torque. Where, s1 denotes the control torque between the stance leg and the torso, and s2 also denotes that between the swing leg and the torso, respectively. It was assumed that the motion of the robot is constrained in the twodimensional planar space. 2.2 Dynamic model The equation of motion is given by M h\u00f0 \u00de\u20ach\u00fe Cc h; _h \u00fe Cd _h\u00fe g h\u00f0 \u00de \u00bc Bs; \u00f01\u00de where h \u00bc h1 h2 h3f gT and s \u00bc s1 s2f gT denote the general coordinate vector and the torque vector at the motors, respectively" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0003514_978-3-030-20131-9_324-Figure98-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0003514_978-3-030-20131-9_324-Figure98-1.png", "caption": "Fig. 98. Bead string state diagram of touching the movable travel switch.", "texts": [], "surrounding_texts": [ "In chapter 2, three methods of automatic knitting of bead mat are put forward. Among them, Warp and Weft Automatic Weaving Method is the most simple, but the cost is higher and its stability is poor. The lock stitch sewing weaving method is rea lly good, but there exists three difficulties applied in the machine: The cross-sectional size of the hook must be smaller than the size of the bead hole as the hook needle needs to completely pass through the bead, which makes it difficult to hook the string smoothly. It is difficult to guarantee the parallelism between the cross-section of the loop ring and the end of the bead of the first row in steps 4-6. When weaving a larger size beaded cooling pad, threading the string into the longer aligned transverse bead holes becomes hard. So compared with the other two methods, single-line straight-through method, which has good stability and high knitting efficiency and is easy to be realized on the machine, is the focus of the discussion below. Based on this method, an automatic weaving device capable of weaving a beaded cool pad is proposed and designed. 4.1 Feeding Device Design Before weaving the beaded cooling pad, since all the beading arrangements are disordered, it is necessary to design a device that puts the beads into the weaving state in an orderly manner, which is called feeding device. Referring to the feeding mechanism of the firecracker weaving [14-16], Fig.14 shows the schematic diagram of the designed bead feeding mechanism device. Before the device runs, all the beads are placed in the hopper and the two guiding wheels. A small number of longitudinally are placed in a horizontal arrangement. When the two guide wheels rotate in opposite directions, the beads in the hopper will be putted into the guide groove and conveyed to the front of the beading device in an orderly manner. In order to avoid a rigid collision between the feeding device and the ball transported device, the end of the guiding groove is made by a material with better elasticity. 4.2 Design and Working Principle of the Beaded Pad Weaving Device Fig. 15 demonstrated the beaded pad weaving device. Since the figure is only for explaining the movement process of the beaded pad weaving device, the feeding device is not shown in the figure. And there are 7 motors in this device. Control motor A controls the movement of the threading device. Control motor B controls the movement of the movable line of the downlink line. Control motor C controls the movement of the movable stroke switch. Control motor D controls the rotation of the output port and the braided port. Linear motor E S. Ouyang et al.2544 drives the up-line feed ball push block movement. Linear motor F pushes the braided beaded cool pad unit into the braided port. Linear motor G drives the linear motion of the downlink line feed bead block. Before the device is operated, the downlink threading is first performed. After the downlink threading is finished, the string is installed into the beaded pad weaving device. The downlink line with heavy beads at the end is wrapped around the fixed pulley mounted on the frame. Then it passes through the downlink line movable seat, the through hole, the downlink end sleeve, the bead and the braided port successively, to reach the uplink line. And the uplink line is directly connected to the needle of the threading device, the first string of beads are moved to the corresponding position on the weaving port, as shown in Fig.16. After the string installation is completed, the motor A is manually controlled to make the driving roller be located between the two trapezoidal blocks on the movable seat rail of the threading device. The manually controlled linear motor E is to drive the uplink line to send th e beads push block moves, which is external bead conveyed from the feeding device. It causes the holes axis of the bead to coincide with the needle axis of the bead threading device and the up-line bead push block is in the beading state. The specific work ing process of the device is as follows: Method Research and Mechanism Design of Automatic Weaving\u2026 2545 Step 1: Under the drive of the control motor A, the bead threading device moves to the right. When the driving roller moves in the second trapezoidal block on the movable seat rail, the location clamping position of the needle will be changed. As a result, the uplink line smoothly penetrates an external bead provided by the uplink line bead transported device, as shown in Fig.17. Step 2: The bead threading device continues to move to the right. When the movable seat contacts the movable travel switch, the uplink line is just tightened. At this time, some of the motor operation will change as follows: Control motor A reversed means the bead threading device starts to move to the left. Controlling motor B rotated forward means the downlink line movable seat moves to the right for a suitable distance, providing two beads re quired for the next unit downlink line weaving. After that, controlling motor B stops. Linear motor F runs, the weaving beads are pushed into the weaving port and the output port, then moves back to the initial position. Linear motor E reversely drives, the pushing block of uplink line feeding bead returns to the initial position, and is on out feeding condition, as shown in Fig.18. S. Ouyang et al.2546 Step 3: Under the driving of the control motor A, the bead threading device moves to the left. When the threading device contacts the fixed stroke switch, some of the motor operation will change as follows: Linear motor G forward drives, the pushing block of the downlink line feeding bead pushes a shared bead to make the hole axis of the shared bead coincide with the needle axis. Control motor A rotated forward means the bead threading device starts to move to the right. Linear motor E drives forward, the push block of the uplink line feeding bead is in the feeding state. Control motor C rotated forward, the movable travel switch moves to the left for a suitable distance exactly equal to the length of the string required to weave every bead pad unit, as illustrated in Fig. 19. Step 4: When the pushing block of the downlink line in the feeding state, the control motor D is drive to rotate the output port and the braided port counterclockwise by 180\u00b0. Step 5: Under the driving of the control motor A, the bead threading device moves to the right. When the driving roller moves in the first trapezoidal block on the movable seat rail, the location and clamping position of needle will be changed to make the uplink line successfully penetrate into a shared bead provided by the downlink line feeding device. Step 6: The linear motor E is reversely driven to make the pushing block of the downlink feeding bead out of the feeding state, as shown in Fig. 20. Method Research and Mechanism Design of Automatic Weaving\u2026 2547 Step 7: Repeat the actions from steps 1 to 6 until the end of the weaving task. Step 8: When the weaving process is finished, each motor is controlled by software programming to bring the device into an initial state." ] }, { "image_filename": "designv11_62_0003065_gtindia2017-4534-Figure2-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0003065_gtindia2017-4534-Figure2-1.png", "caption": "Figure 2. Shaft element: (a) Coordinate system and general loading (b) Geometry of cracked cross-section", "texts": [ " In the present study, breathing crack is considered during rotation of the FG shaft to accurately predict the dynamic responses of the rotor system. The shaft internal damping is neglected in the present work. Modeling of cracked FG shaft element The cracked FG shafts with simply supported ends are discretized using FE and shown in Fig. 1(a), where total length L , element length eL and diameter D of the shaft and the depth of crack located at distance cL from the left end of the shaft. Fig. 1(b) shows the any arbitrary crack orientation of the shaft. Fig. 2(a) shows the shaft element subjected to shear forces 1 2 5 6, , and P P P P and bending moments 3 4 7 8, , and P P P P , rotating at speed , v and w are the translation displacements along Y and Z direction, and and are the rotational displacements about Y and Z at a distance s from the left end of a cross-section of the shaft. Fig. 2(b) shows the circular section of the cracked shaft having crack orientation o180 with crack half-width b . The LFCs of a FG shaft with a crack using Castigliano\u2019s theorem and Paris\u2019s equations [20] in conjunction with the expression for SIFs are used and the additional displacement c iu due to crack is expressed as 0 2 2 2 4 4 4 1 1 1 ( , ) 1 and ( , ) ( , ) x c i i i i i i i i u J y T dy P J y T AK K K y TE (5) where ( , )J y T is the strain energy density function and SIFs I II III, and K K K corresponding to modes I, II, and III 3 Copyright \u00a9 2017 ASME Downloaded From: http://proceedings.asmedigitalcollection.asme.org/ on 02/06/2018 Terms of Use: http://www.asme.org/about-asme/terms-of-use expressions in Eq. (5) are taken from Tada et al. [20] and 1, 2,3,4 i are load indices. Young\u2019s modulus of elasticity, for the plane stress ( , ) ( , )y T E y TE and for the plane strain 2( , ) ( , ) 1 ( , )y T E y T y TE and {1 ( , )}A y T . Finally, LFCs of an FG shaft with crack orientation o180 are computed as follow (refer Fig. 2(b)) 2 0 , xc b ic ij bj i j u C J y T dy dz P P P (6) where 2 2 x R R z and 22b RR . Here, integrations are performed to compute LFCs of a cracked FG shaft considering ( , )E y T and ( , )y T as a function of y as well asT , unlike homogeneous shaft where these are considered as constants and could be taken out of integration in Eq. (6). Using Eq. (6), LFCs are calculated corresponding to fully open crack ( o180 ) and corresponding local flexibility matrix ,y Tc C is determined using energy methods as c 11 c 22 c c 33 34 c 44 , 0 0 0 , 0 0 , , , , C y T C y T y T C y T C y T sym C y T c C (7) where expressions of the elements of ,y Tc C are provided in Appendix A" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0003210_joe.2018.8514-Figure8-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0003210_joe.2018.8514-Figure8-1.png", "caption": "Fig. 8 Electric field distributions on and around the insulator (without water column)", "texts": [ "2 Air gap breakdown simulation of the insulator: When the water column is falling vertically from the box of bird guard, keep the distance of d as 200 mm, which is the centre of the water column to the centre of the insulator, keep the distance of the upper air gap is 349 mm, and the distance of the lower air gap is 220 mm, under this condition, the length of the water column is 1314.04 mm, the simulation results are shown in Fig. 7. The electric field and potential distributions around the insulator without water column are shown in Fig. 8. The unit of the potential is kV, and the unit of the electric field is kV/m. When there is no water column, the maximum electric field strength around the insulator is 12.76 kV/cm. In the case of a water column, the maximum electric field strength on the insulator is increased to 13.44 kV/cm, and the electric field near the water column is obviously distorted, the maximum electric field strength has reached 23.99 kV/cm, the floating potential on the water column is 74.02 kV, thus the voltage applied on the lower air gap is 123" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0003399_pierc19010804-Figure1-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0003399_pierc19010804-Figure1-1.png", "caption": "Figure 1. A scheme of two sets of cylindrical magnets.", "texts": [ "n some applications, such as eddy current dampers, magnetic refrigerators, micropumps, the calculation of interaction force, torque and field is very important [1\u20134]. The benefit for these applications is that magnetic forces act without physical contact over larger distances than electrostatic, piezoelectric or other schemes [5, 8, 9]. In the literature related to permanent magnets, calculations of forces between magnets of various shapes and geometries have their relevance in the context of several applications. For investigating the magnetic force between two sets of cylindrical permanent magnets (Fig. 1), the force engendered between two parallel magnets should be calculated firstly. The analytical expression of the attractive force between two arrays of cylindrical permanent magnets was derived from the derivative of the total magnetostatic interaction energy with respect Received 8 January 2019, Accepted 19 March 2019, Scheduled 2 April 2019 * Corresponding author: Naamane Mohdeb (mohdeb.naamane@gmail.com). The authors are with the L2EI, Universite\u0301 de Jijel, BP 98, OuledAissa 18000, Algeria" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0002375_iea.2017.7939175-Figure2-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0002375_iea.2017.7939175-Figure2-1.png", "caption": "Figure 2. Schematic diagram of extrusion set", "texts": [ " In recent year, modern FDM technologies offers great potential for a range of other materials including PLA filled bronze, iron or stainless steel [6] and new composite material filaments with metal filled particles have been successfully developed for direct application in FDM process [7]. The most concerned of FDM process parts with composite material filaments is the wear of the nozzle [8]. Almost all filament manufacturers give a vague warning about composite material filaments causing nozzle wear on the nozzle of FDM machine [9]. The extrusion set in FDM machine, as shown in Fig. 2. The systems involves a cold end and a hot end. The cold end is part of an extruder system that pulls and feed the material from the spool, and pushes it towards the hot end. The cold end is mostly gear- or roller-based supplying torque to the material and controlling the feed rate by means of a stepper motor. The hot end is the active part which also hosts the liquefier of the FDM that melts the filament. The hot end consists of a heating chamber and nozzle [10]. The melt flow channel\u2019s shape and length in nozzle are designed for thermoplastic material [11]" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0002414_b978-0-08-101123-2.00008-x-Figure8.4-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0002414_b978-0-08-101123-2.00008-x-Figure8.4-1.png", "caption": "Fig. 8.4 Classification of the radiant heating and cooling system.", "texts": [ " Depending on the position of the piping in the building, an RHC system is classified as follows: (A) Embedded surface systems (pipes placed within a building layer (floor, wall, and ceiling)) that are isolated from the main building structure (B) Thermally activated building system where pipes integrated into the main building structure (ceiling, wall, and floor) (TABS) (C) Radiant panel system (pipes integrated into lightweight panels) A classification of the radiant heating and cooling system is shown in Fig. 8.4 [9]. THE INDOOR CLIMATE OF RADIANT HEATING AND COOLING SYSTEMS By utilizing the radiant heat transfer between the human body and radiant surfaces, a radiant heating system can achieve the same level of thermal comfort at a lower air temperature than a conventional system. A radiant cooling system can provide the equivalent thermal comfort at a higher air temperature and thus has energy-saving potential. Due to the small temperature differences between a heated or cooled surface and the occupied space, the radiant heating and cooling system (RHC) can benefit from the selfregulating effect, which can provide a stable thermal environment for the occupants within the space" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0003613_012066-Figure6-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0003613_012066-Figure6-1.png", "caption": "Figure 6. Contour of von-Mises stress on a coil spring with a defect of 0.2 mm given the minimum load.", "texts": [], "surrounding_texts": [ "Transport vehicles require a good suspension system to dampen vibration, swings and shocks received as they travel along bumpy, hollow, and uneven roads [1]. These conditions are very uncomfortable and may cause accidents. The suspension is also expected to hold the load during some common vehicle maneuvers such as acceleration, braking or deflection while on the road [2]. The coil spring is one of the main components for dampening vibrations and shocks to the load so as to provide comfort and security while the vehicle is in motion [3]. Depending on the condition of their application, coil springs often sustain fatigue failure. This indicates that the tension received below by the coil spring from the maximum stress of the material while sustaining a dynamic load causes fatigue failure [4-8]. The yield strength of the material is also a criterion of failure. Components of automotive suspension must be changed with a traveling distance of 73,500 km, or every five years [9]. The fault of 13.18 % of 24.2 million vehicle tests was recorded [10]. With the development of computing technology, the numerical analysis method has become particularly suitable for use because it will increase the calculation efficiency, the cost-effectiveness as well as save time. Various numerical analysis methods are widely available, but the finite element analysis (FEA) has proven to be reliable in solving problems in the field of continuum mechanics [11]." ] }, { "image_filename": "designv11_62_0001458_j.ast.2012.02.004-Figure1-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0001458_j.ast.2012.02.004-Figure1-1.png", "caption": "Fig. 1. Airframe and dynamic variables.", "texts": [ " The control objective is to force the air vehicle to track a desired motion path generated by the guidance-navigation sys- tem as a reference normal acceleration command for the center of the mass. The problem is first formulated and performance objectives are specified. The zero dynamics of the airframe is utilized to demonstrate the non-minimum phase characteristics of the system. The model assumes constant mass, no roll rate, zero roll angle, no side slip, and no yaw rate. Under these assumptions, the longitudinal nonlinear equation of motion is reduced to two forces and one moment. Using body axis coordinates, Fig. 1, these equations are: Fx = Q SC D F z = Q SCN [ \u03b1(t), \u03b4(t), M(t) ] M y = Q SCM [ \u03b1(t), \u03b4(t), M(t) ] (1) where dynamic pressure is defined as Q = 0.5\u03c1(t)V 2 m(t) = 0.7P0(t)M2(t). Note that the velocity and air density are not assumed to be constant or slow-varying, but a standard atmospheric model is assumed in simulation based on previously reported data [11]. Aerodynamic polynomials resulting from wind-tunnel measurements are given as [6] CN [ \u03b1(t), \u03b4(t), M(t) ] = \u03b21N\u03b13(t) + \u03b22N\u03b1(t) \u2223\u2223\u03b1(t) \u2223\u2223 + \u03b23N [ 2 \u2212 M(t) 3 ] \u03b1(t) + dN\u03b4(t) = cn [ \u03b1(t), M(t) ] + dn\u03b4(t) (3) CM [ \u03b1(t), \u03b4(t), M(t) ] = \u03b21M\u03b13(t) + \u03b22M\u03b1(t) \u2223\u2223\u03b1(t) \u2223\u2223 + \u03b23M [ \u22127 + 8M(t) 3 ] \u03b1(t) + dM\u03b4(t) = cm [ \u03b1(t), M(t) ] + dm\u03b4(t) (4) The numerical values of the various constant parameters in the dynamic equations (2)\u2013(4) are given in Table 1" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0001128_20100906-5-jp-2022.00066-Figure2-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0001128_20100906-5-jp-2022.00066-Figure2-1.png", "caption": "Fig. 2 Interaction rules and corresponding body direction vectors", "texts": [ " These distance-reaction rules function to avoid straying away from neighbours in \"approach\", to avoid collision with neighbours in \"repulsion\", and to move together with neighbours to a common destination in \"parallel orientation\". To incorporate these three rules into the air vehicle model, a spherical interaction field is set around the agent as shown in Fig. 1, which is divided into three sub-fields, i.e. approach, parallel orientation, and repulsion fields, and the agent selects one reaction from above-mentioned three reactions in response to the position of a neighbour in the interaction field as shown in Fig. 2. The vectors app, para, and repul, in Fig. 2 denote the body direction vectors of the agent for the approach, parallel orientation, and repulsion, respectively, where app and para point to the direction of the neighbour rj and to the direction in parallel with the body direction of the neighbour j, respectively; repul points to the direction which is the summation of the counter direction of the neighbour -rj and the body direction of the agent i with the weight coefficients c1 and c2, respectively, to avoid a collision with the neighbour (here c1=c2=1 is used)" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0001477_ciss.2012.6310783-Figure3-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0001477_ciss.2012.6310783-Figure3-1.png", "caption": "Fig. 3. The clustered sensor network example from Fig. 1 with an additional coordination layer that coordinates the medium-access of clusters such as in a standard TDMA approach. This requires bidirectional wireless communication links between the coordination layer and the computation layer as well as between the nodes and the fusion centers in each cluster (represented by arrows with two peaks).", "texts": [ " , yk,2|Ck|) between E |Ck| and \u0393k \u2282 R 2|Ck|+1 that enable each cluster to compute every fk \u2208 C[E|Ck|] because of the existence of representations fk ( xk1 , . . . , xk|Ck| ) = 2|Ck| \u2211 m=0 \u03c8km ( \u2211 n\u2208Ck \u03d5(k) nm(xn) ) , (13) k = 1, . . . , K . Since the compact sets \u0393k will unfortunately differ in general, the pre-processing functions in (13) depend on k such that the common nodes between clusters have to adapt their pre-processing functions in dependency of the fusion center to which they transmit next. This in turn requires the coordinated activation of clusters as indicated in Fig. 3. But to guarantee an appropriate coordination, two-way communication between nodes and fusion centers has to be implemented. In contrast, the proposed computation method from Section IV-A1, depicted in Fig. 2, do not requires any global coordination since all clusters can transmit simultaneously such that one-way communication suffices. Furthermore, due to the universality property of pre-processing functions, cluster topologies are allowed to change without to reconfigure transmitting nodes. If a standard TDMA protocol is used to compute vectorvalued functions in a clustered wireless sensor network, besides the orthogonalized medium-access of the nodes in each cluster, clusters themselves have to be appropriately separated in time (see Fig. 3). Unlike the 2N + 1 transmissions, a standard TDMA would induce KL \u2265 Kmaxk |Ck| separated transmissions to convey the entire raw sensor readings interference-free to the K fusion centers, which subsequently compute the component-functions {fk} K k=1. In contrast to KL, it is obvious that 2N + 1 does not scales with the number of clusters such that huge performance gains are possible. Remark 5. We conclude from Theorem 3 that unfortunately the number 2N +1 of transmissions can not be reduced since otherwise there exists an f \u2208 C[EN ] that is not representable in the form (2)" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0002029_978-3-319-02609-1-Figure7.7-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0002029_978-3-319-02609-1-Figure7.7-1.png", "caption": "Fig. 7.7 Two colliding robots can assemble only if their relative heading \u03b8h is smaller than \u03b1a/2. The blue circular sector represents both the detection and the communication area of the robots.", "texts": [ " 106 7 Model Calibration 7.2.1 Example from Case Study II Case Study II A collision does not necessarily lead to the formation of a bond due to the directionality characterizing the building blocks, be they robots or molecules. In Case Study II for instance, robots must be aligned to some extent in order to aggregate successfully. We approximate this constraint by stating that the absolute value of the heading \u03b8h must be smaller than \u03b1a/2, where \u03b1a is the central angle of the detection and communication sector (see Figure 7.7). Because of the non-holonomic nature of the Alice robots, we assume that there is always at least one robot, which we denote B, that is aligned upon collision (B is the robot that runs into the other). Furthermore, we assume that the absolute value of the heading of B with respect to another robot A is uniformly distributed in [0, \u03c0]. As a result, the probability pa that, upon collision, two robots are properly aligned is pa \u223c \u03b1a/(2 \u00b7\u03c0). Since each robot and each chain has two valid binding sites, the overall probability pb that a bond is formed can be written pb = pc \u00b7 2 pa \u223c= v\u0302 T wd Atotal \u00b7 \u03b1a \u03c0 " ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0000786_aero.2013.6496960-Figure4-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0000786_aero.2013.6496960-Figure4-1.png", "caption": "Figure 4 \u2013 Details of the developed propulsion unit", "texts": [ " Here, we developed an excavation unit that is to be combined with the propulsion unit. Several excavation experiments were performed to obtain realistic requirements for the propulsion unit. Development of an Excavation Unit The excavation unit of the robot bores a hole having the same diameter as that of the propulsion unit, so a tapered auger using a fishtail single-spiral type screw was developed [8]\u2013[9]. It was necessary to equip the auger with a skirt that can assist in smoothly carrying the excavated soil to the transport part. Figure 4 shows the characteristics of the excavation unit. Because peristaltic crawling requires at least three units, we considered it useful to equip this robot with four units. Then, the robot can move with several propulsion patterns and maintain its position with a maximum of two contracted units. The EA was 425 mm long, which is longer than the combined four units used for peristaltic crawling. The total mass of this unit is 5.9 kg. Excavation Experiments In this experiment, we set the pushing forces at 55 N at the beginning of the experiments and the rotation of the EA at 10 rpm" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0003355_rpj-06-2018-0156-Figure1-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0003355_rpj-06-2018-0156-Figure1-1.png", "caption": "Figure 1 Two examples of 3D models with multiple chambers", "texts": [ " For this purpose, a 3D model is usually processed into a model with chambers, which may be hollow or infilled with a variety of geometric patterns such as honeycomb (Lu et al., 2014), struts (Stava et al., 2012), tree (Wang et al., 2013), rhombic cells (Wu et al., 2016). However, given a model with disjoint chambers inside, the powders are trapped in the chambers of the finished model when using the powder-bed AM techniques such as SLM and SLS. This is because these techniques selectively fabricate a region of powders in a layer-by-layer manner, leaving the unselected powders trapped in the interior of the chambers. See Figure 1 for an illustration; after a model with disjoint chambers is fabricated, a set of channels need to be drilled into the model to leak the powders. Albeit straightforward, blind drilling may weaken the strength of the model, and it may not The current issue and full text archive of this journal is available on Emerald Insight at: www.emeraldinsight.com/1355-2546.htm Rapid Prototyping Journal \u00a9 Emerald Publishing Limited [ISSN 1355-2546] [DOI 10.1108/RPJ-06-2018-0156] This work was supported in part by National Natural Science Foundation of China No", " However, all the literature has assumed that the 3D objects are of simple geometry such as points and spheres; no literature has ever considered the shortest spanning tree for arbitrary objects. In this paper, to preserve the interior structure and the appearance of the model, the problem of connecting all the disjoint chambers in any 3D model with a shortest channel network and the least number of outlets is studied. Particularly, the chambers are allowed to be hollow or infilled with proper support structures inside (e.g. the crossbars in Figure 1(b)). The rest of the paper is organized as follows. Section 2 presents the methodology for solving the problem with arbitrary chambers; Section 3 presents the simulation and experimental results of a set of models; Section 4 presents the limitations of our work; Section 5 concludes the paper. In this section, the problem of designing channels with a fixed dumping direction is discussed first, and then an approach for addressing the general problem of channel design with multiple dumping directions is presented" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0001600_tasc.2010.2100801-Figure6-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0001600_tasc.2010.2100801-Figure6-1.png", "caption": "Fig. 6. Measuring bench:", "texts": [ " To define orientation of field sensors inside the measuring unit with reference to the PF1 dummy, measurements of individual field components were carried out [3]. The coil geometry reconstructed from the magnetic measurements was compared with the parameters measured with a level gage and a ruler. A good match of the results proved a feasibility of the proposed technique for coil quality assessment [3]. Subsequently a measuring bench prototype was built. The measuring bench used in the PF1 geometry inspection is shown in Fig. 6. The bench comprises a baseboard 1 where a small-scale coil dummy 2 is installed. The measuring unit 3 is moved around the dummy by means of a stepper 5. The measuring unit is equipped with 12 single-axis field sensors grouped in sets of 3 so as to provide three-axial field measurements. The electronic interface provides control inputs and signal amplification/conditioning. Two 16-bit performance, 8-channel analog universal input modules are used to convert analog signals to digital outputs. The data are recorded and stored to the PC hard drive to be used for further processing, analysis, and display" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0000418_detc2011-47599-Figure2-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0000418_detc2011-47599-Figure2-1.png", "caption": "FIGURE 2. Screw notations of an n-DOF cable-driven closed chain", "texts": [ ", twists, reciprocal wrenches, external wrenches and cable wrenches) 2. Select the active joints and arrange the cable routings accordingly 3. Determine the joint torques required to sustain the external wrenches 4. Determine the joint torques provided by the cable wrenches based on the routing configuration 5. Equilibrate the joint torques between the external wrenches and cable wrenches 6. Determine force-closure at a particular pose by checking if there exists a set of positive cable tension As shown in Fig. 2, the following notations will be used for the twists and wrenches in the examples of this paper: \u2022 $\u0302k: j represents the unit screw of twist of jth joint in kth chain \u2022 $\u0302k: jr represents the unit screw of reciprocal wrench of jth joint in kth chain \u2022 $\u0302i: j\u2212h represents the unit screw of ith cable wrench, acting from jth link to hth link and j 6= h \u2022 $ewk: j represents the screw of external wrench acting on jth link in kth chain \u2022 $ep represents the screw of external wrench acting on end- effector \u2022 ti represents the tension scalar of ith cable Figure 3 presents the planar two-3R cable-driven closed chain that will be used as an example to carry out the proposed force-closure analysis based on reciprocal screw theory" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0000564_s12204-011-1147-y-Figure1-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0000564_s12204-011-1147-y-Figure1-1.png", "caption": "Fig. 1 The schematic of meshing for the wedge module", "texts": [ " According to the classical welding theory, in the temperature field model, the main consideration is given to the heat source data, the thermal boundary conditions and the materials\u2019 thermodynamic properties. In the stress-strain field model, the main consideration is given to the thermal stress, the mechanical boundary conditions and loading data and the stress-strain curve[1-2]. As the repair of big workpiece\u2019s circular orbit has been always a difficult problem, in this paper some ring is the subject investigated. Figure 1 shows the schematic of meshing for the wedge module. The largest external diameter is 1.96m, and its height is 0.75m. The bulk material is the 1045 steel, and the cladding material is Ni60 self-flux powders. The thermal and physical parameters are collected through the reference book. When the shape is very complicated and the volume is large, there is a large difficulty in calculation. So, considering the symmetry of the ring, one wedge module is selected to simplify the calculated process. The first cladding ring is on the bottom side of the wedge module, and the laser moves from the plate 1 to plate 2. The cladding parameters are laser power of 4 kW, scanning speed of 4mm/s, laser facular diameter of 4mm and powder feeding speed of 350mg/s. Assuming the cross-section of overlay is a trapezium with the bottom margin length of 4 mm and the height of 1mm. The 10\u25e6 wedge module is divided into 26 124 elements, as shown in Fig. 1(a)[3]. As the materials at these points tend to crack most easily, the four key points are selected to be the subject investigated. Points A and B are the center points on the surfaces of base and coating respectively. Points C and D are both adjacent points on the base/coating interface. In the numerical modeling process, the coordinate system is set up in such way that z represents circumferential direction in which the laser moves, while x and y are the radial and axial directions as shown in Fig. 1(b). Temperature field is the foundation of the stress and strain field in the calculating process. According to the welding theory, the temperature field is mainly influenced by the heat source data, boundary condition and thermodynamic property[4-6]. Therefore, in the present calculation, the thermal loads are applied on the body elements of the coating and the surface element of the base in the forms of body heat generation rate and surface heat flux density respectively. The loads are applied by the means of step, which means that when some elements are exposed by the laser, the loads exist" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0000569_1.4004098-Figure1-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0000569_1.4004098-Figure1-1.png", "caption": "Fig. 1 Schematic of test setup", "texts": [ " Impact and friction-induced acoustic emission is considered as the stress wave or solid sound emanating from the regions of localized deformation of a solid structure. The piezo- electric element, or the acoustic emission sensor, transforms the particle motion produced by an elastic wave into an electrical signal. Mechanical deformation and fracture are the primary sources of AE, which generate elastic waves that travel through the structure to the sensor element. The experiment is carried out with a conventional glide test spindstand. The test setup is illustrated in Fig. 1. The sliders used in the study were 30% pico sliders made from Al2O3\u2013TiC with a thickness of 0.3 mm. A PZT element with the same length and width as that of the slider was sandwiched between the slider and the suspension tab [19]. The air-bearing surface (ABS) surface of the head is coated with carbon overcoat. The air-bearing surface (ABS) has central trailing pad with width of 300 lm. The nominal Journal of Tribology JULY 2011, Vol. 133 / 031901-3 Downloaded From: http://tribology.asmedigitalcollection" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0000710_s10958-013-1525-0-Figure5-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0000710_s10958-013-1525-0-Figure5-1.png", "caption": "Fig. 5", "texts": [ " The subsequent increase in the rotational velocity not necessarily leads to the appearance of plastic state along the circle \u03c1 = \u03b2 in D 2 but changes the values of the coefficients of the characteristic equation [10]: A1 = 2 \u2032\u2032C3 \u2212 2(\u03bd2 + 3) \u2032x , A2 = A1 + 8 \u2032x , \u2032\u2032C3 = s \u2212 8\u03b22\u0393 + 3(\u03bd2 + 3)(1\u2212 \u03b22 )\u23a1\u23a3 \u23a4\u23a6 \u2032x /3{ }{1\u2212 \u03b2\u22122}\u22121 If the characteristic equation turns into the identity \u0394(\u03b20 ) \u2261 0 prior to satisfying the condition \u2032\u2032T2 = \u2032\u2032C3(1+ \u03b2\u22122) + (\u03bd2 + 3 \u2212 (3\u03bd2 + 1)\u03b22) \u2032x = 1 , (14) then this situation is called the special case 01\u00b0 : D1(p)D 2(e) ; the loss of stability for the corresponding critical values of the angular velocity \u03c9\u2217 := \u03c93 and the radius of the plastic zone \u03b20\u2217 := \u03b2 (Fig. 5). On the contrary, if the inequality \u2032\u2032T2 < 1 turns into equality (14) on attainment of the rotational velocity \u03c93 , them the loss of stability of the disk D is realized according to scenario 02\u00b0 : D1(p)D 2(pe) (Fig. 6). In this special case, in the characteristic equation, we have \u0394(\u03b20 ) = 0 , \u03b20 > \u03b2 , \u2032\u2032C3 = 1\u2212 \u03bd2 + 3 \u2212 (3\u03bd2 + 1)\u03b202\u23a1\u23a3 \u23a4\u23a6 \u2032x{ }{1+ \u03b20\u22122} \u22121 , \u03c9\u2217 2 = \u03c942 = 24q22 2 + \u03b2\u03b20\u22121(s \u2212 1)(1+ \u03b202 ){ } 3[\u03bd2 + 3 \u2212 (3\u03bd2 + 1)\u03b202 ](1\u2212 \u03b202 ){ + 3(\u03bd2 + 3) \u2212 (3\u03bd2 + 1)\u03b202 + 8\u03b23\u03b20\u22121(\u0393 \u2212 1)\u23a1\u23a3 \u23a4\u23a6(1+ \u03b202)}\u22121 for \u03b201\u2217 := \u03b2 and \u03b20\u2217 := \u03b20 " ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0000002_10402000903491275-Figure2-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0000002_10402000903491275-Figure2-1.png", "caption": "Fig. 2\u2014A typical representation of a single asperity contact.", "texts": [ " [1], [2], and [12] and replace r by (R\u03b4)1/2 or (2R\u03b4)1/2, as applicable, and the average flash temperatures for an elastically and plastically deformed asperity in the low-speed regime can be given respectively by: avle = \u03b6\u00b5v 4\u03ba1 [ K\u03b4 \u2212 ( 6\u03b3K \u221a R\u03b4 )1/2 ] [16] avlp = \u03c0\u03b6\u00b5v 2 \u221a 2\u03ba1 [ H \u221a R\u03b4 \u2212 \u03b3 \u221a R \u03b4 ] [17] where the subscripts avle and avlp refer to average low-speed elastic and average low-speed plastic, respectively. For fastmoving heat sources the flash temperature is given by Eq. [13] and its maximum value would occur when the trailing edge of the moving body B traverses a distance of 2r (Fig. 2). Substituting t = 2r/v in Eq. [13] we have maxh and averaging over the traverse distance we have avh.. These may be given by maxh = 2 \u221a 2 qr1/2 (\u03c0\u03ba\u03c1cv)1/2 [18] and avh = 4 \u221a 2 3 qr1/2 (\u03c0\u03ba\u03c1cv)1/2 [19] respectively. Now, replacing q by \u03b6\u00b5Pev \u03c0r2 and \u03b6\u00b5Pp v \u03c0r2 for elastically and plas- tically deformed contact spots respectively in Eq. [19] and then substituting Pe and Pp from Eqs. [1] and [2] with r equal to (R\u03b4)1/2 or (2R\u03b4)1/2, as applicable, the average temperature rise for an elastically and plastically deformed asperity contact in the high- speed regime can be given respectively by: avhe = 4 \u221a 2 3\u03c03/2 \u03b6\u00b5 ( v \u03ba1\u03c11c1 )1/2 [ K ( \u03b43 R )1/4 \u2212 (6\u03b3K)1/2 ] [20] avhp = 4\u03c0.23/4 3 \u03b6\u00b5 ( v \u03ba1\u03c11c1 )1/2 [ H (R\u03b4)1/2 \u2212 \u03b3 ( R \u03b43 )1/4 ] [21] where the subscripts avhe and avhp refer to average high-speed elastic and average high-speed plastic, respectively. The maximum local average temperature within the whole contact zone (Fig. 2) must be the largest value of the average temperatures given in Eqs. [16], [17], [20], and [21]. However, in view of the involvement of a large number of thermal and mechanical parameters it is difficult to maximize these temperatures with respect to a single parameter. For this purpose it is convenient to nondimensionalize the equations and plot the nondimensional temperature against nondimensional load with certain parametric variations. Using the nondimensional scheme followed in the section on Contact Loading and Deformation in the Presence of Surface Forces, Eqs" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0003948_0731684419879235-Figure1-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0003948_0731684419879235-Figure1-1.png", "caption": "Figure 1. Self-developed component four-point bending fixture for coupon specimens.", "texts": [ " Strain gauges were then installed on some specimens at the centre of the upper and lower faces (details in Section Experimental Results). The experiments on this work were performed on a very high speed servohydraulic testing machine (Instron, Norwood, USA), which is capable of testing at the multiple speeds required for this study. By using the same machine for all the testing speeds, the effect of the fixture/machine on the measured results was kept constant throughout the entire study. A four-point bending fixture30 was developed which can be observed in Figure 1. It consists of an impactor and two supports. The radius of the loading noses and supports is 5mm, the loading span is 27mm, and the support span was set at 80mm to keep a loading span to support span ratio of 1/3. These dimensions were determined using the ASTM D627231 standard as reference. The specimen displacement was measured using the Digital Image Correlation software ARAMIS (GOM, Braunschweig, Germany) together with two highspeed cameras type FASTCAM SA-X (Photron, Tokyo, Japan). The force was measured using a strain gauge type BTMC-1-D16-003LE (Tokyo Sokky Kenkyujo, Tokyo, Japan) integrated in the supports which were calibrated in accordance to the norm ISO7500" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0001342_0954406213479272-Figure1-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0001342_0954406213479272-Figure1-1.png", "caption": "Figure 1. Front view of harmonic drive-based gear pump\u20141: casing; 2: flexible gear; 3: rigid gear; 4: boom; 5: sealing block; A: suction cavity; B: discharge cavity.", "texts": [ "comDownloaded from Two mating gears, having the same tooth width and circular pitch but different tooth number (the tooth number of the flexible gear is slightly less than that of the rigid gear, usually less than two teeth), turn in an inner chamber surrounded by the casing and end cover. The rigid gear outside is the driving gear. The speed ratio uGR of rigid gear to flexible gear is uGR \u00bc ZR=ZG \u00f02\u00de Crescent-shaped sealing blocks (one is on the left and the other on the right, as shown in Figure l) are setup in radial direction between the addendum circle of the rigid gear and the addendum arc surface of the flexible gear. Two gears match closely with two side plates (6 and 10 in Figure 2) in the axial direction. Two sealing blocks (5 in Figure 1) are fixed on the side plates by position pins, two end covers are binded in the casing (1 in Figure 1) using bolts. The flexible gear, rigid gear, sealing blocks and side plates separate the space between the rigid gear and the flexible gear into four independent airtight cavities: two cavities marked A for suction and two cavities marked B for expelling fluid (see Figure l). From Figures 1 and 2, it can be seen that the two gears, sealing blocks and side plates constitute closed cavities A and B to realize the suction and expelling of fluid. Accordingly, two axial holes, c, for suction (corresponding to lower pressure cavity A), and two axial holes, d, for discharge (corresponding to higher pressure cavity B) are cut symmetrically on the side plates" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0003105_interjenercleanenv.2017020710-Figure1-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0003105_interjenercleanenv.2017020710-Figure1-1.png", "caption": "FIG. 1: Press-granulator scheme: 1) hopper-agitator, 2) paddle conveyor, 3) system of feeding pipes, 4) cylindrical die, 5) pressing roll, 6) cylindrical channel, 7) cutting metal knife", "texts": [ " These technologies allow one to obtain fuels with different characteristics due to the combination of different biomaterials and additives (Berghel et al., 2013). The fractional composition of complex pellets has a substantial effect on the fuel energy related characteristics. Therefore, the determination of the optimal fractional composition of such pellets is an urgent goal. The fuel pellets for experimental research were manufactured by a two-roll granulator of capacity G = 2 t/h, the scheme of which is shown in Fig. 1. The diameter of the granules is d = 8 mm. Sawdust of hardwood, milling peat of cotton\u2013sphagnum type with a level of decomposition R = 30% were used as the raw material for the production of granules. The mass content of sawdust in the composition (Mwood) relative to the initial mass varied in the range 0\u2013100% in 25% steps. The drying of the raw material before pressing was carried out in a continuous drying drum at the heat carrier temperature of 180oC at the inlet and of 120oC at the outlet. The procedure for obtaining fuel pellets is as follows. A prepared mixed raw material is loaded into a hopper-agitator 1 of the granulator (Fig. 1). It is uniformly fed from the agitator into a cylindrical granulating die 4 through the system of feeding pipes 3 using a conveyor 2. In the granulator operation mode, the die 4 is rotated at a frequency of n = 200 rpm. There are freely rotating cylindrical rolls 5 with a clamping mechanism and notches on the working surface inside the die. The raw material is pushed to the die inner walls by centrifugal forces during the rotation. Then it gets into the space between the roller and the die and becomes pressed into the cylindrical channels 6" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0003968_oceanse.2019.8867403-Figure1-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0003968_oceanse.2019.8867403-Figure1-1.png", "caption": "Fig. 1. The earth-fixed frame and the body-fixed frame", "texts": [ " The adaptive law is used to design and adjust the parameters of the fuzzy controller online. Finally, the actual control law with unknown term estimation based on adaptive fuzzy logic is proposed. The stability of the DP system is proved by Lyapunov theory. Simulation studies on a FPSO vessel are provided in different cases and compared with the ordinary sliding mode controller, it illustrates the adaptability, robustness and effectiveness of the control method. II. MATHEMATICAL MODEL Consider the earth-fixed frame NOE and the body-fixed frame b bx O y\u2032 in horizontal plane shown in Figure 1, the variables [ ]= , , T x y\u03b7 \u03c8 denote the position ( ),x y and the heading \u03c8 in the earth-fixed frame; and [ ]= , , T u v r\u03bd denote the surge velocity u , sway velocity v and yaw velocity r in the body-fixed frame, respectively. The 3-degrees-of-freedom (surge, sway, and yaw) nonlinear model of the FPSO vessel is established as follows: ( ) ( ) ( ) ( ) J M C D d t \u03b7 \u03b7 \u03c5 \u03c5 \u03c5 \u03c5 \u03c5 \u03c5 \u03c4 = + + = + (1) where 3 3= A RBM M M R \u00d7+ \u2208 , RBM is the generalized rigid body inertia matrix and AM is the generalized add mass inertia matrix; ( ) ( ) ( ) 3 3 RB AC C C R\u03c5 \u03c5 \u03c5 \u00d7= + \u2208 , ( )RBC \u03c5 and ( )AC \u03c5 are the skew-symmetric Coriolis-centripetal matrix of the rigid body and the added mass respectively; ( ) ( ) 3 3 L NLD D D R\u03c5 \u03c5 \u00d7= + \u2208 , LD and ( )NLD \u03c5 are the linear and the nonlinear damping coefficient matrix; \u03c4 represents the control input of the vessel; ( )d t represents the forces and moment of the environment disturbances; and ( )J \u03b7 is a rotation matrix" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0000443_s10778-010-0311-7-Figure1-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0000443_s10778-010-0311-7-Figure1-1.png", "caption": "Fig. 1", "texts": [ " The solutions to the bending problem for plastic plates with a free hole known from the literature are related to the dynamic behavior of the following plates of constant thickness: annular plates [12, 14, 15], curvilinear plates [5, 6], circular plates with an arbitrary hole [9], and to the limiting analysis of annular [1] and square [10, 16] plates. We will analyze the dynamic behavior of hinged and clamped arbitrary curvilinear plates with varying thickness and an arbitrary free hole. An elliptic plate with a free rhombic hole and with piecewise-linear thickness function will be considered as an example. 1. Model, Assumptions, and Equations of Motion. Let us consider a hinged or clamped perfect rigid-plastic plate with varying thickness and arbitrary smooth convex boundary L 1 (Fig. 1). There is a free hole with an arbitrary boundary L 2 in the middle of the plate. The plate is subject to a high explosive load P t( ) uniformly distributed over the surface of the plate, instantaneously peaking (P P t max ( ) 0 ) at the initial time t 0 and then rapidly decreasing. The boundary L 1 of the plate is described by the parametric equation x x 1 ( ) , y y 1 ( ) , 0 2 . The radius of curvature of the boundary L 1 is given by R( ) L x y x y 3 1 1 1 1 ( ) / ( ) , where L x y( ) ( ) ( ) 1 2 1 2 , ( ) ( ) / . Let, for the sake of determinacy, the plate be symmetric about the x-axis and have y-dimension no greater than the x-dimension. Let also the inward normals drawn from L 1 do not intersect inside the plate (Fig. 1). This condition is satisfied if 0 2 D R( ) ( ) , (1.1) where D 2 is the distance along the normal to L 1 from the boundary L 1 to the boundary L 2 . We choose curvilinear orthogonal coordinates ( 1 , 2 ) related to the Cartesian coordinates (x y, ) by x x y L 1 2 1 1 2 2 ( ) ( ) / ( ) , y y x L 1 2 1 1 2 2 ( ) ( ) / ( ) (2). (1.2) International Applied Mechanics, Vol. 46, No. 3, 2010 304 1063-7095/10/4603-0304 \u00a92010 Springer Science+Business Media, Inc. S. A. Khristianovich Institute of Theoretical and Applied Mechanics of the Siberian Branch of the Russian Academy of Sciences, 4/1 Institutskaya St", " The curves 1 const 0 are at a distance 1 from the boundary L 1 inside the plate and have radius of curvature 1 2 1 R( ) . Such curves are called parallel or equidistant [11]. The straight lines 2 const are perpendicular to the boundary L 1 (radius of curvature 2 ). In this case, the boundary L 1 is described by the equation 1 0 , 0 2 2 , while the boundary L 2 by 1 2 2 D ( ), 0 2 2 . The following equalities are valid: arctan[ ( ) / ( )]x y 1 2 1 2 , d L R d [ ( ) / ( )] 2 2 2 , L R d d ( ) ( ) 2 20 2 2 0 2 2 , (1.3) where is the angle between the 1 -direction and the x-axis (Fig. 1). The thickness hof the plate is a function of the parameter 1 and varies symmetrically about the mid-surface of the plate. The rigid\u2013plastic plate can follow two deformation patterns depending on P max . If the load does not exceed the ultimate level (low loads), the plate remains at rest. As in the case of constant thickness [6], if the load is slightly above the ultimate level (moderate loads), the plate rotates around the boundary L 1 through an angle 1 and deforms into some ruled surface. This is pattern 1" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0000267_s11517-013-1052-7-Figure2-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0000267_s11517-013-1052-7-Figure2-1.png", "caption": "Fig. 2 a\u2013e The rototranslation sequence that transforms the position and orientation of the instrument in RefS, to its position and orientation in AbsS, according to Eq. (2a). f The spherical local reference frame, TipS, adopted to describe the position of the points of the stripe on M0. g A point of the stripe, PS,i,abs, and its projection, ps,i, on the camera image plane", "texts": [ " In the following, we will assume that distortions have been corrected through an adequate procedure and that the interior (cx, cy, fx, fy) and exterior (RC, TC) parameters have been determined through calibration. 2.2 Instrument model We suppose that the section of the tracked instrument can be approximated by a straight segment and two spherical markers, M0 and M1, are attached to its axis. The vectors M0,abs = [X0, Y0, Z0]T and M1,abs = [X1, Y1, Z1]T indicate the 3D position of the markers center in AbsS. A local reference frame, RefS, is attached to the instrument, centered half-way between M0 and M1, with the RefX axis aligned with the instrument axis (Fig. 2). In RefS we have M0,ref = [-d/2 0 0]T and M1;ref \u00bc d=2 0 0\u00bd T, where d = ||M0,abs\u2013M1,abs||. We also suppose that a narrow black stripe has been painted onto the surface of M0 and that the stripe axis is contained in the RefZX plane. The position and orientation of the instrument with respect to AbsS is described by the 6-tuple [XT,YT,ZT,RX, RY,RZ] that represents the roto-translation that brings AbsS coincident with RefS. The roto-translation is defined by three sequential anti-clockwise rotations around the absolute X, Y and Z axes (Fig. 2a\u2013d). This can be described by the matrices RX, RY and RZ, followed by a translation T = [XT, YT, ZT]T, [10, 24] (Fig 2e). The relationship between a 3D point in AbsS and RefS is then: Mabs \u00bc RZRYRXMref \u00fe T \u00bc RMref \u00fe T \u00f02a\u00de Mref \u00bc RT XRT YRT Z Mabs T\u00f0 \u00de \u00bc RT Mabs T\u00f0 \u00de \u00f02b\u00de where R = RZRYRZ is the rotation matrix. Axial rotation RX is the first one applied to the points: it does not influence the final position of the center of M0 and M1 as they both lie along the X axis in RefS. As a consequence, we can rewrite (2a) for M0 and M1, omitting RX as: Mi;abs \u00bc RZRYMi;ref \u00fe T i \u00bc 0; 1 \u00f03\u00de 2.3 Computation of the instrument motion The instrument motion is described by the temporal sequence of sextuples that transforms the point coordinates from RefS to AbsS", " Therefore, from the 3D absolute position of M0 and M1, all the dofs of the instrument but the axial rotation are computed through (4) and (5a, 5b). 2.4 Computation of the axial rotation We first identify on the camera images the pixels, fps;i \u00bc \u00bdxS;iyS;i gi\u00bc1::N ; that belong to the stripe painted on M0. To this aim we consider the largest dark blob contained inside the circle associated to the marker. We then exploit the correspondence between each 2D point ps,i of the stripe and its corresponding 3D point over the marker\u2019s spherical surface, Ps,i,abs (Fig. 2g). To this aim we set an additional local reference frame, TipS, centered in M0 with the axes parallel to those of RefS (Fig. 2f). In TipS the two markers have coordinates M0;tip \u00bc 0 0 0\u00bd T and M1;tip \u00bc d 0 0\u00bd T . Equation (2a) can be rewritten for each stripe point as: PS;i;abs \u00bc RZRYRX PS;i;tip \u00fe d=2 0 0\u00bd T \u00fe T \u00f06\u00de The position of the stripe point in TipS can also be written in spherical coordinates as: PS;i;tip \u00bc R3D sin bi sin ai cos bi cos ai cos bi\u00bd T \u00f07\u00de where ai and bi represent the longitude and latitude of the ith stripe point (Fig. 2f). Differently from what happened to the markers center, a rotation around the RefX axis does affect the final position of the stripe points. Suppose now that the instrument undergoes only a rotation around the RefX axis, (RX = 0, RY = RZ = 0, T = 0). In this case, the position of the stripe points in TipS is described, after the rotation, as: RXPS;i;tip \u00bc 1 0 0 0 cos RX sin RX 0 sin RX cos RX 2 4 3 5PS;i;tip \u00bc R3D sin bi\u00f0 \u00de sin ai Rx\u00f0 \u00de cos bi cos ai Rx\u00f0 \u00de cos bi 2 64 3 75 \u00f08\u00de Comparing (7) and (8), it can be seen that an axial rotation of the instrument corresponds to a translation of a stripe point by RX along the a direction in the (a, b) plane, independently from the longitude bi of the point. The position of PS,i,abs, can also be obtained as the intersection of the line r (passing through the image point ps,i and the projection center of a camera, T), and the marker\u2019s spherical surface (Fig. 2g). In particular, for any point P of r we have P = TC ? kvi, where: vi\u00bcRC xs;i cx fx ys;i cy fy 1 h iT , ffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi xs;i cx fx 2 \u00fe ys;i cy fy 2 \u00fe1 s \u00f09\u00de vi is a versor containing the director cosines of r in AbsS and k represents the distance of P from TC. Since PS,i,abs lies on r, we can write: PS;i;abs \u00bc TC \u00fe kivi \u00f010\u00de where ki is the distance of PS,i,abs from TC", " cos b: a high value of da/dxS is associated to points that are very sensitive to errors in their position on the image. In particular, the analysis of Eq. (14) of Appendix B highlights that the reliability of a decreases in the following cases: (i) b ? \u00b1 90 . In this case, cos(b) ? 0 and da/ dxS ? ?. These are the points close to the poles of the marker sphere, where the distance from the axis of rotation tends to zero and the longitude a of the point tends to become undefined. (ii) D ? 0. These are the points close to the border of the marker on the 2D image. In this case the line r in Fig. 2g is almost tangent to the marker surface and a small change in xS corresponds to a large change of the latitude, a. (iii) k ? ?. This is the case of a very far marker, which produces an image with a small diameter: a small error on image position corresponds to a large displacement on the marker surface. These theoretical observations have been validated through simulations with the same set-up adopted in the \u2018\u2018Results\u2019\u2019 section. Markers with a radius of 7.5 mm were randomly positioned and oriented (for a total of 1,000,000 different samples) at a distance from 50 to 800 mm from the projection center, TC, of a camera, with an image plane of infinite dimension and a focal length of 1,000 pixels" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0002587_icca.2017.8003129-Figure3-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0002587_icca.2017.8003129-Figure3-1.png", "caption": "Fig. 3. The relationship between U, \u03b3i, \u03b2i and \u03d5i.", "texts": [ " Moreover, the estimated error of target center r\u0303i(t) also exponentially converges to 0. Proof: We first show that \u03be\u0303i(t), \u2200i, exponentially converges to 0. By (14) and Lemma 1, for any i the estimated errors of target position \u03be\u0303i(t) exponentially converges to 0 if and only if the signal \u03d5\u0304i(t) is PE, that is, there exist three positive constants \u03c31, \u03c32 and T such that for any t0 \u2208 R\u22650 \u03c31 \u2264 \u222b t0+T t0 (UT \u03d5\u0304i(t)) 2dt \u2264 \u03c32 (25) is satisfied. \u03b3i(t) is defined as the angle between the unit vector U and \u03d5\u0304i(t) as shown in Fig. 3. Similarly, \u03b8i(t) and \u03b2i(t) are shown in Fig. 3. We assume that if the directions of the angle \u03d5\u0304i(t), \u03b8i(t) and \u03b2i(t) are counter-clockwise then they are positive; otherwise, they are negative. From (25) we further have that \u03c31 \u2264 \u222b t0+T t0 cos2\u03b3i(t)dt \u2264 \u03c32. (26) Due to cos(\u00b7) \u2264 1, (26) always has an upper bound. Thus we only need to show that the lower bound of (26) is always satisfied for any t0 \u2208 R\u22650. Since U is the unit constant vector, thus the angle \u03b2i(t)\u2212 \u03b3i(t) is constant. From Fig. 3 we have that d\u03b3i(t) dt = d\u03b2i(t) dt . (27) Since \u03c6\u0304i(t) is obtained by \u03c0/2 clockwise rotation of \u03c6i(t), so \u03b2i(t) \u2212 \u03b8i(t) = 3\u03c0/2 is also a constant. Thus we have that d\u03b2i(t) dt = d\u03b8i(t) dt . (28) Since the speed of agent i along \u03d5\u0304i(t) is \u03b1i (refer to (10)), the distance between xi(t) and \u03bei(t) is \u03c1i(t), we have that d\u03b8i(t) dt = \u03b1i \u03c1i(t) . From (27) and (28) one gets d\u03b3i(t) dt = \u03b1i \u03c1i(t) . (29) According to Proposition 1, we know that for any i there exists Di > 0 such that \u03c1i(t) \u2264 Di, \u2200t \u2265 0. From (29) one further gets d\u03b3i(t) dt \u2265 \u03b1i Di , namely, \u03b3i(t) \u2265 \u03b3i(0) + \u03b1it Di , \u2200t \u2265 0" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0001166_gt2012-69356-Figure14-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0001166_gt2012-69356-Figure14-1.png", "caption": "Figure 14: Schematic of the offset rotor used to create a", "texts": [ " The high speed videos of the Compliant Plate Seals indicate that self\u2013correcting behavior is relatively insensitive to variations in intermediate plate front gap and back gaps. Figure 13 shows non\u2013contact operation was achieved for two extreme cases in which (a) Intermediate Plate \u2013 Front Gap was four times the Intermediate Plate \u2013 Back Gap and (b) Intermediate Plate \u2013 Back Gap was four times the Intermediate Plate \u2013 Front Gap. D) Response of seal to high-frequency transients In one test, an offset rotor as shown in Figure 14 was used to test seal response to high frequency transients (1/rev). The rotor is segmented into sections \u201cA\u201d and \u201cB\u201d. The calibration labyrinth seal mates with section \u201cA\u201d of the rotor and the corresponding geometric axis \u201cC\u201d is aligned with the axis of rotation. The Compliant Plate Seal mates with the section \u201cB\u201d of the rotor, and the corresponding geometric axis \u201cD\u201d is offset from the axis of rotation by a certain amount \u201coffset\u201d as shown in Figure 14. In this set up, the rotor diameter has the same diameter as the seal inner diameter. With the rotor assembled as shown, the seal has interference at the bottom dead center of amount \u201coffset\u201d and the seal has clearance at the top dead center of amount \u201coffset\u201d. As the rotor rotates, the seal is subjected to a rotor transient of distance \u201c2 X offset\u201d mils/rev, that are significantly larger than those observed in turbomachinery. In one of the tests, an offset rotor with a radial offset of 0.508 mm was used and the seal was subject to a transient of 1" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0000086_amr.423.143-Figure9-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0000086_amr.423.143-Figure9-1.png", "caption": "Figure 9: Real boundary profile of the fibre cross section", "texts": [ " Secondly, to take into account this measured profile to determinate the mechanical properties of a unitary fibre [8]. Thereby a profile detection and cross section measure method of fibre was developed and investigated. The method consists in taking several pictures at different tested sample orientations. A total of 5 pictures were taken at 0\u00b0, 36\u00b0, 72\u00b0, 108\u00b0 and 144\u00b0 using a specific mounting showed in Figures 7 and 8. The data which are gotten by numerical imaging treatment are sufficient to have an exactly description of cross section (Fig.9). As hemp fibre doesn\u2019t have a geometrical standard profile, it\u2019s necessary to determine a specific geometry of each sample tested. From the geometric data points, five different ways of fibre modelling was investigated, considering an average circular cross section (method 1) or an average polygonal cross section (method 2), a cross section at the failure location being circular (method 3) or polygonal (method 4) and finally considering 3D fibre model. The four first methods define the cross section like circular or polygonal and constant along the fibre (Fig" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0001600_tasc.2010.2100801-Figure1-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0001600_tasc.2010.2100801-Figure1-1.png", "caption": "Fig. 1. One of the eight double pancakes of the PF1 coil.", "texts": [ " The field perturbations above this limit can cause plasma instability resulting in plasma disruption [1]. Main sources of the field asymmetry in tokamaks are possible misalignments and distortions of the ITER coils from their ideal geometry associated with the accepted manufacture/assembly tolerances. One of Russian Federation\u2019s responsibilities for the ITER project is manufacturing and procurement of the coil PF1, which belongs to the ITER poloidal field magnet system intended for the control of the plasma shape and position. PF1 consists of 8 double pancakes (Fig. 1) and will be fabricated and tested at Efremov Institute in St.Petersburg, Russia. The desired performance is ensured if proper techniques and instrumentation would be applied to control the coil quality. So the choice of the techniques is one of the critical issues. As an example, the dimensional inspection technique and environment of the PF1 coil could be mentioned. The current Manuscript received August 03, 2010; accepted November 29, 2010. Date of publication January 28, 2011; date of current version May 27, 2011" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0002774_pvp2017-65992-Figure7-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0002774_pvp2017-65992-Figure7-1.png", "caption": "FIGURE 7. TEMPERATURE PROFILES DURING THE FIRST PASS, MIDWAY THROUGH THE SECOND PASS, AND AFTER COOLING FOR THE VARIABLE LASER POWER THIN WALLED BUILD. CORRESPONDING AXIAL STRESSES ARE SHOWN IN FIGURE 6.", "texts": [ " Generally, the magnitudes of axial stress increased once reaching the final build state at room temperature as compared to during the build 4 Copyright \u00a9 2017 ASME Downloaded From: http://proceedings.asmedigitalcollection.asme.org/ on 11/10/2017 Terms of Use: http://www.asme.org/about-asme/terms-of-use (Figure 6). Axial stress magnitudes were lower near the beginning and end of each deposition pass as compared to the middle of the pass. Temperature profiles with the variable laser power thin walled model show a relatively constant and small melt pool size and some substrate heating near the melt pool (Figure 7). For the variable laser power thin walled build model, approximately 61% of the volume of the substrate remained within 50 K of ambient temperature. The axial stress solution shows apparent periodicity that corresponds with the solution time step. Additionally, a jagged pattern near the boundary of activated elements is also evident and correlates with the solution time step size. The cylindrical button model was run through 22.5 seconds of LENS build time. With progressively increasing build time, maximum principal stress and minimum principal stress magnitudes and distributions propagated with deposition, but showed similar magnitudes and distributions throughout the build time (Figure 8)" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0002350_0954406217712280-Figure3-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0002350_0954406217712280-Figure3-1.png", "caption": "Figure 3. FEM model mesh close to tooth flank. Point A has fixed translation and is rotated with small angular increments. Point B has fixed translation and a torsional moment.", "texts": [ " It is then exported as a STEP file and imported into the FEM software Abaqus. Here the geometry is used to create a plane strain model of the pinion and the wheel. The two wheels are positioned on a fixed distance and fixed translation boundary conditions. Since this is a two-dimensional model, only rotation out of the plane is active. The load applied as a moment is set on the larger gear and a fixed rotational is set on the smaller pinion gear. Individual interactions between each pair of gear, see Figure 3 teeth are defined the back side of the tooth are excluded this to avoid possible contact on both sides. As friction model, a simple Columb friction has been used with coefficient of friction set to zero. Each gear wheel has been partitioned close to the tooth root and along the contact interface. This allows an easier way to set a denser mesh especially at the tooth contact. Elements used are a variation of linear and quadratic with a mesh density at the interface of 50\u2013200 mm. The model has 30 k\u2013120 k degrees of freedom" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0001123_gt2012-68510-Figure3-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0001123_gt2012-68510-Figure3-1.png", "caption": "Figure 3: Complete recuperator assembly comprised of 36 stacks measuring 257 layers tall.", "texts": [ " The sintering process heats the tapecast material to 1,500C, forming a coherent ceramic mass through atomic diffusion. Once the sintering was completed, the individual segments may be 2 Copyright \u00a9 2012 by ASME Downloaded From: http://proceedings.asmedigitalcollection.asme.org/ on 01/27/2016 Terms of Use: http://www.asme.org/about-asme/terms-of-use bonded to each other using a high temperature cement, or held in place in a recuperator assembly by some other method such as mechanical clamping. As seen in Figure 3, each segment was equivalent to a 10 degree section of the recuperator. Therefore, the entire recuperator consists of 36 segments with each segment measuring 257 layers tall. The 257 layer tall segments are comprised of 64 four-layer wafers plus a single \u201cend cap\u201d layer. In the green state (meaning before sintering), the material was very flexible and was prone to break during handling. Adding sacrificial material around the exterior and interior supports made the layers more durable and helped maintain straight channels throughout fabrication" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0001260_ecce.2011.6064134-Figure1-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0001260_ecce.2011.6064134-Figure1-1.png", "caption": "Figure 1. Dc current decay standstill test wiring diagram for an Axially Laminated Anisotropic rotor Reluctance Synchronous Machne with 4 poles", "texts": [ " The space\u2013 vector model of RSM in rotor reference frame is: \u03bb = + + \u03c9\u03bbs s s s s d V i R j dt (1) \u03c3 \u03c3\u03bb = \u03bb + \u03bb = + + +s d q d md s q mq sj i ( L L ) ji ( L L ) (2) where: s d qV =V +jV \u2013 stator voltage, s d qi =i +ji \u2013 stator current vector, s\u03bb \u2013 stator flux vector, Rs \u2013 stator phase resistance, Ls\u03c3, Ldm, Lqm \u2013 the leakage and d\u2013q axis magnetizing inductances, and \u03c9r is the rotor speed. The RSM electrical parameters were determined after performing dc current decay tests with the rotor aligned along d axis, and, respectively, along q axis, at standstill. The electrical diagram of the test is illustrated in Fig. 1 As can be seen from Fig.1 the current decay occurs in phases B and C, which are actually connected in series, while in phase A different values of constant current, icc, are injected. The d-q fluxes were calculated using Equation (3): , (2 ) / 2d q s DR idt V dt\u03bb = \u22c5 \u22c5 +\u222b \u222b (3) where: Rs \u2013 phase resistance; i - decay current, VD \u2013 diode voltage drop. The \u03bbd\u2013Id and \u03bbq\u2013Iq dependencies are illustrated in Fig. 2 a and b. The cross coupling saturation is visible mainly in axis d, while Lq is rather constant for an ALA rotor, as shown in Fig" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0003399_pierc19010804-Figure3-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0003399_pierc19010804-Figure3-1.png", "caption": "Figure 3. Model of the permanent magnet by the Coulombian approach.", "texts": [ " The dimensions of the system are shown in Fig. 2. 2a and 2A are the radii of magnet 1 and magnet 2. 2C and 2c are the thicknesses of the magnets. The cylinder shown in Fig. 2 is uniformly magnetized in the z direction. The modeling of magnets can be based on Coulombian method, and the Coulombian approach replaces the magnet by two surfaces distribution of fictitious magnetic charge with surface density \u03c3\u2217 = M \u00b7 n [20], where M is the magnetization, and n is the unit vector normal to the surface. Fig. 3 shows a rectangular permanent magnet with uniform magnetization in the z-direction. For circular surface, we can approximate circle with several rectangles. We will approximate the graph by dividing the interval into \u201cn\u201d subintervals, each of width, \u0394x = (2a)/n. The rectangle height is established by evaluating the values of f(c), as shown for the typical case x = c, where the rectangle height is f(c). In Fig. 4, we approximate the surface using inner rectangles (each rectangle is inside the curve)" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0001620_978-1-4614-3475-7_2-Figure2.27-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0001620_978-1-4614-3475-7_2-Figure2.27-1.png", "caption": "Fig. 2.27 Example 2.4", "texts": [ "60), after integration, it results xC ( 1 2 R2 ) (2\u03b1) = ( 1 3 R3 ) (2sin(\u03b1)) , or xC = 2Rsin(\u03b1) 3\u03b1 . (2.61) For a semicircular area, \u03b1 = \u03c0 2 , the x-coordinate to the centroid is xC = 4R 3\u03c0 . For the quarter-circular area, \u03b1 = \u03c0 2 , the x-coordinate to the centroid is xC = 4R \u221a 2 3\u03c0 . Example 2.4. Find the coordinates of the mass center for a homogeneous planar plate located under the curve of equation y = sinx from x = 0 to x = a. Solution A vertical differential element of area dA = ydx = (sinx)(dx) is chosen as shown in Fig. 2.27a. The x-coordinate of the mass center is calculated from (2.15): xc \u222b a 0 (sin x)dx = \u222b a 0 x(sin x)dx or xc {\u2212cos x}a 0 = \u222b a 0 x(sin x)dx, or xc(1\u2212 cos a) = \u222b a 0 x(sin x)dx. (2.62) The integral \u222b a 0 x(sin x)dx is calculated with \u222b a 0 x(sin x)dx = {x(\u2212cos x)}a 0 \u2212 \u222b a 0 (\u2212cos x)dx = {x(\u2212cos x)}a 0 + {sin x}a 0 = sin a\u2212 a cos a. (2.63) Using (2.62) and (2.63) after integration, it results xc = sin a\u2212 acosa 1\u2212 cosa . The x-coordinate of the mass center, xC, can be calculated using the differential element of area dA = dxdy, as shown in Fig. 2.27b. The area of the figure is A = \u222b A dxdy = \u222b a 0 \u222b sin x 0 dxdy = \u222b a 0 dx \u222b sin x 0 dy = \u222b a 0 dx {y}sin x 0 = \u222b a 0 (sin x)dx = {\u2212cos x}a 0 = 1\u2212 cos a. The first moment of the area A about the y-axis is My = \u222b A xdA = \u222b a 0 \u222b sin x 0 xdxdy = \u222b a 0 xdx \u222b sin x 0 dy = \u222b a 0 xdx {y}sin x 0 = \u222b a 0 x(sin x)dx = sin a\u2212 a cos a. The x-coordinate of the mass center is xC = My/A. The y-coordinate of the mass center is yC = Mx/A, where the first moment of the area A about the x-axis is Mx = \u222b A ydA = \u222b a 0 \u222b sin x 0 ydxdy = \u222b a 0 dx \u222b sin x 0 ydy = \u222b a 0 dx { y2 2 }sin x 0 = \u222b a 0 sin2 x 2 dx = 1 2 \u222b a 0 sin2 dx" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0000394_09596518jsce982-Figure7-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0000394_09596518jsce982-Figure7-1.png", "caption": "Fig. 7 Case 3: excavator\u2019s work space intersects two dig sides", "texts": [ " The middle of this segment has a coordinate that equals l/2. Similar results may be obtained for the remaining three sides of the dig. In Fig. 6 all corresponding line segments (PQ) and lines (e\u2013e9 for vertical sides and h\u2013h9 for horizontal) are depicted. The four segments are the loci of positioning points that provide maximum surface and thus maximum volume of removed soil. In case 3 it is assumed that the circle of radius RE formed by the excavator\u2019s end intersects two sides of the dig. Two distinct subcases can be distinguished, as shown in Fig. 7. In the first subcase (Fig. 7(a)), xE and yE have their values in the ranges f , wzRe\u00f0 \u00de and l{Re cos sin{1 f {w Re , l respectively. The re- movable soil volume is analogous to the surface Sc (shown by hatching in figure) that can be calculated as follows Sc~SezSa{St \u00f06\u00de where Sa is the surface of circular sector and Se, St are the surfaces of triangles BFH and EHG respectively. Since Sa~ R2 Ehs 2 ~ R2 E 2 sin{1 l{ye Re z cos{1 xe{w Re \u00f07\u00de Se~ 1 2 BH\u00f0 \u00de BF\u00f0 \u00de ~ 1 2 l{ye\u00f0 \u00de{ tan sin{1 l{ye=Re\u00f0 \u00de xe{w\u00f0 \u00de 2 tan sin{1 l{ye=Re\u00f0 \u00de\u00bd \u00f08\u00de St~ 1 2 EK\u00f0 \u00de HG\u00f0 \u00de ~ 1 2 Re sin cos{1 xe{w Re z xe{w\u00f0 \u00de tan sin{1 l{ye Re xe{w\u00f0 \u00de \u00f09\u00de As seen, surface Sc is a function of Re, xE, yE, w and l and is maximized when both xE and yE present JSCE982 Proc. IMechE Vol. 224 Part I: J. Systems and Control Engineering at University of Otago Library on September 4, 2014pii.sagepub.comDownloaded from minimum values in their ranges (see above). This implies that point E3 on line e\u2013e9 with coordinates f , l{Re cos sin{1 f {w Re is the point that pro- vides maximum soil removal volume for the subcase under consideration for frontal excavation orientation. In the second subcase (see Fig. 7(b)), xE and yE have their values in the ranges w, wzRe cos sin{1 g{l Re and g , lzRe\u00f0 \u00de respectively, where g is the coordinate of line h\u2013h9 that is similar to line e\u2013e9 (see previous analysis). The removable soil volume is analogous to the surface Sc (shown by hatching in figure) that can be calculated as follows Sc~Sa{Se{St \u00f010\u00de where Sa is the surface of circular sector and Se, St are the surfaces of triangles EFB and EBG respectively. Since Sa~ R2 Ehs 2 ~ R2 E 2 cos{1 ye{l Re z cos{1 xe{w Re { tan{1 xe{w ye{l { tan{1 ye{l xe{w \u00f011\u00de Se~ 1 2 ye{l\u00f0 \u00de Re sin cos{1 ye{l Re { xe{w\u00f0 \u00de \u00f012\u00de St~ 1 2 xe{w\u00f0 \u00de Re sin cos{1 xe{w Re { ye{l\u00f0 \u00de \u00f013\u00de As seen, surface Sc is a function of Re, xE, yE, w, and l and is maximized when both xE and yE are minimum values in their value domains" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0003355_rpj-06-2018-0156-Figure5-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0003355_rpj-06-2018-0156-Figure5-1.png", "caption": "Figure 5 A 2D illustration of particle Pi = [di,1, di,2, di,3, di,4]", "texts": [ " In this approach, there is a built-in learning mechanism: the birds learn the position of the current best bird and adjust their flying directions accordingly. The difficulty of exploiting the PSO to solve the problem lies in two aspects: the first part is the definition of \u201cparticle\u201d and the second part is the evolution of the particles. Formally, the PSO algorithm for solving the system is presented as follows. First of all, a swarm of particles are potential solutions. In this case, refer to Figure 5 for an illustration, a particle Pi is defined as an array of directions [di,1, di,2, . . ., di,k], the directions are applied on M in sequence and each direction corresponds to a maximal connected component of chambers and channels. Each particle Pi is associated with two vectors, i.e. the velocity vector Vi = [vi,1, vi,2, . . ., vi,k] and the position vector Xi = [di,1, di,2, . . ., di,k], where the velocity vector is used to move a particle from one position to another. The initial velocity is given as [0, 0, " ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0000685_coase.2012.6386315-Figure5-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0000685_coase.2012.6386315-Figure5-1.png", "caption": "Fig. 5. Schematic diagram of a pneumatic artificial muscle manipulator.", "texts": [ " Additionally, the longer distance from the client to the server causes a longer communication time, mainly due to the increase in the number of routers that data pass through. From these results, we choose UDP/IP and consider that the Internet communication needs 45 ms at the maximum. With regard to the communication error, as it occurred at 0.05% in 6600 measurements with UDP/IP, we neglect its occurrence in this study. For a robot manipulator driven by three PAMs, we built a test apparatus of a networked JIT control system based on the concept described in Section II-B. Fig. 5 is a schematic diagram of the PAM manipulator. The coordinate systems Ob \u2212 xbybzb and Oe \u2212 xeyeze shown in Fig. 5 are referred to here as the base coordinate system and the end coordinate system, respectively. As in the figure, three PAMs connect the movable end part with the fixed base part in parallel through spherical bearings, and a parallel mechanism is formed. Thus, the manipulator has both the advantages of PAMs (such as safety for humans) and those of parallel mechanisms (such as compact as well as high degree-offreedom motions). Link 1 is attached to the center of the base part, and only its length is freely adjustable" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0003679_j.mechmachtheory.2019.05.026-Figure2-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0003679_j.mechmachtheory.2019.05.026-Figure2-1.png", "caption": "Fig. 2. Hob section profile of single cutting tooth.", "texts": [ " In the homogenous coordinate notation system used in the present study, a position vector a x i + a y j + a z k is written in the form of a column matrix a = [ a x a y a z 1 ] T . Similarly, the unit directional vector is written as n = [ n x n y n z 0 ] T . The drawing approach provides a feasible means of obtaining the critical points of a hob. However, it is not easily implemented in computer code. Consequently, the present study employs an algebraic method to derive the hob profile instead. Fig. 1 shows the complete hob profile cross section perpendicular to the rotation axis, C 0 . Fig. 2 shows the section profile of a single cutting tooth on the hob with a negative rake angle \u03c8 and outer radius R . The origin of the hob coordinate system ( xyz ) 0 is located at the rotation axis C 0 . For a brand new hob, the following basic design parameters are known in advance: (1) the relief parameter s cs ; (2) the outer radius R = R 0 ; and (3) the number of hob gashes n gash . Given the value of n gash , the interval angle \u03c6 between neighboring gashes is obtained simply as \u03c6 = 2 \u03c0 n gash . (1) The clearance angle \u03b7 of the hob can then be determined as tan \u03b7 = s cs R 0 \u03c6 . (2) However, while some hobs provide the relief parameter s cs along the rotation radius line (see Fig. 2 ), other hobs provide an alternative parameter s \u2032 cs along the rake surface. The relationship between the two parameters is discussed in the following. Consider an arbitrary point P \u2032 t on the clearance surface projection curve S cs . The distance r \u03b8 between P \u2032 t and the rotation axis C 0 varies as a function of the rotation angle \u03b8 in accordance with r \u03b8 = C 0 P \u2032 t = R ( 1 \u2212 \u03b8 tan \u03b7) . (3) Substituting Eq. (2) and R = R 0 into Eq. (3) , the distance r \u03b8 for a brand new hob is given by r \u03b8 | brand new hob = R 0 \u2212 s cs \u03b8 \u03c6 ", " (4) Hence, the equation of the clearance surface projection curve S cs is obtained as S cs = \u23a1 \u23a2 \u23a2 \u23a3 \u2212 r \u03b8 | brand new hob S\u03b8 r \u03b8 | brand new hob C\u03b8 0 1 \u23a4 \u23a5 \u23a5 \u23a6 = \u23a1 \u23a2 \u23a2 \u23a2 \u23a2 \u23a2 \u23a2 \u23a3 \u2212 ( R 0 \u2212 s cs \u03b8 \u03c6 ) S\u03b8 ( R 0 \u2212 s cs \u03b8 \u03c6 ) C\u03b8 0 1 \u23a4 \u23a5 \u23a5 \u23a5 \u23a5 \u23a5 \u23a5 \u23a6 , (5) where S and C denote sine and cosine, respectively. The coordinates of the tip point P tip of the hob are given by P tip = [ 0 R 0 1 ] T . Consequently, the rake surface projection line L rs can be described using the following parametric equation: L rs = P tip \u2212 t Lrs \u23a1 \u23a2 \u23a3 S\u03c8 C\u03c8 0 0 \u23a4 \u23a5 \u23a6 , (6) where t Lrs is the parameter of this parametric line equation. It should be noted that if the hob is brand new, the value of the outer radius R = R 0 . The equation of the neighboring rake surface projection line L rs 1 (see Fig. 2 ) can be obtained by rotating L rs around the center point C 0 through the gash interval angle \u03c6, i.e., L rs 1 = Rt (z, \u03c6) L rs , (7) where Rt ( z, \u03c6) is a rotation matrix about the z-axis, as shown in Eq. (A4) . Solving the equality equation of Eqs. (5) and (7) (i.e., S sc = L rs 1 ) with R = R 0 , the intersection point P cs can be determined and the relationship between s cs and s \u2032 cs obtained. In other words, if the alternative relief parameter s \u2032 cs along the rake surface is given rather than the relief parameter s cs along the rotation radius line, the intersection point P cs can be determined by solving S sc = L rs 1 using a numerical method", " In general, the initial value of R = R 0 for a brand new hob reduces after routine re-sharpening. As a result, the relief parameter s cs is no longer valid. However, the clearance angle \u03b7 remains the same. In preparing the mold for the injection molding process, the hob is required to cut a gear with a whole tooth height H g , where H g is the distance between points P tip and P Hg (see Fig. 3 ). The end point P rs of the cutting edge on the rake surface can be obtained by rotating point P Hg about the rotation axis C 0 (refer to Fig. 2 ). Hence, the parametric equation of arc r c can be expressed as r c = (R \u2212 H g ) \u23a1 \u23a2 \u23a2 \u23a3 S t rc C t rc 0 1 \u23a4 \u23a5 \u23a5 \u23a6 , (8) where t rc is the parameter of this parametric line equation. The rake surface projection line L rs is given in Eq. (6) . Solving the equality equation of Eqs. (6) and (8) (i.e., L rs = r c ), the intersection point P rs can be analytically determined as P rs = \u23a1 \u23a2 \u23a2 \u23a2 \u23a2 \u23a3 S\u03c8 \u221a 2 H 2 g \u22124 H g R + R 2 + R 2 C(2 \u03c8) \u221a 2 \u2212 RC\u03c8S\u03c8 C\u03c8 \u221a 2 H 2 g \u22124 H g R + R 2 + R 2 C(2 \u03c8) \u221a 2 + R S 2 \u03c8 0 1 \u23a4 \u23a5 \u23a5 \u23a5 \u23a5 \u23a6 , (9) where H g is the whole tooth height of the machined gear" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0002302_b978-0-12-803581-8.10351-0-Figure26-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0002302_b978-0-12-803581-8.10351-0-Figure26-1.png", "caption": "Fig. 26 Bulkheads vertical walls and bonding flanges.", "texts": [ " The enlarged bonding flanges allow for fasteners from the rocker post to have a greater thickness to pass through to reduce the amount of shear stress the composite sees. The bracket indent allows for a smaller bracket to be used to reduce the amount of stress from the suspensions loading. All of this must be packaged into a small enough volume to not interfere with the driver\u2019s shins, which are directly below this bulkhead. The bulkhead reinforces the chassis in torsion by the means of two vertical walls and the bonding flanges shown in Fig. 26. The two vertical walls extend from each side of the chassis enclosing the upper portion of the chassis at the suspension box. The bonding flanges increase the overall thickness of the chassis material, thus helping to boost stiffness in this area. Fig. 27 demonstrates how the bulkhead reduces the chassis ability to twist. The localized loads that the bulkhead sees include a 650 lb lateral load from the suspension and a 1000 lb lateral load from the steering rack as shown in Fig. 28. The suspension load also utilizes the U-shape of the bulkhead in reducing the load transfer" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0000708_imece2013-65269-Figure1-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0000708_imece2013-65269-Figure1-1.png", "caption": "Fig. 1. Schematic diagram of electrospinning apparatus", "texts": [ " The solution was stirred overnight to generate a homogeneous solution. Prepared solution was then used for nanofiber fabrication using electrospinning. In order to fabricate fibers with different diameters, three different flow rates of 1, 2, and 3mL/hr were used and the rest of parameters were kept fixed. A voltage of 25 kV DC was used for all three fibers. The distance of 3 cm was used between the syringe tip and the collector. All other electrospinning factors were kept constant. The basic set up for the electrospinning process is illustrated in Fig. 1. The interdigitated electrodes were soaked in ethanol for 10 minutes and washed with DI water and air dried to create a contamination free area. Three fabricated electrospun polyaniline nanofibers and manifolds were integrated on top of the gold interdigitated electrodes using biocompatible adhesives to create three electrospun polyaniline nanofiber based biosensors for detection of COX-2. A different biosensor was 2 Copyright \u00a9 2013 by ASME Downloaded From: http://proceedings.asmedigitalcollection" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0003254_s00170-019-03312-1-Figure16-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0003254_s00170-019-03312-1-Figure16-1.png", "caption": "Fig. 16 (a) Standard boundary-fill pattern with three measurement regions which will provide differing surface roughness results, (b) process variations to the boundary contours and raster-fill angle b, (c) a block with 3 mm height, 6 mm width, 15\u00b0 sides, 0% overlap, 45\u00b0 start angle, 90\u00b0 rotation angle and 0.25 mm bead height, 0.5 mm bead width, 0% overlap, 45\u00b0 start angle, 90\u00b0 rotation angle, (obround representation). (e) torus front view, 3D view, and sliced model showing the internal packing", "texts": [ " At this point, only the surface boundary conditions have been discussed, and when a critical angle has been reached, offset contour geometry assumed. As boundary contours are deposited, and then a raster-scan fill pattern is utilized to fill the solid regions, the influence of rasterfill pattern needs to be explored. This is done in the next section. The number of boundary contours and the fill raster angle will significantly alter the surface roughness. The default configuration for the FDM process with one boundary curve, and a 45\u00b0 raster fill pattern is illustrated in Fig. 16 (a). The surface roughness results for the red line will be different from the yellow and green lines: the surface contour defined by the red line will include a void; whereas, the yellow line has two extended flatter sections, and the dashed green line includes a corner condition, but this influences only one bead. Multiple boundary curves and fill patterns are shown in Fig. 16 (b) (top surface for a simple block). The interior regions will have consistent surface roughness results, the other regions will have differing results due to the multiple boundary contours, voids, long continuous sections, and the raster fill angle b. Two case studies realistically illustrate the general challenges for assessing the surface roughness: (i) a block with 15\u00b0 slope for each edge, a 6 mm bead width, and 3 mm bead height, and a 45\u00b0 initial fill angle, and a 90\u00b0 rotation angle, or an FDM variant with a 0.25 mm bead height and 0.50 mm bead width, and (ii) a torus with 1.25 mm bead height, 2.5 mm bead with, and a 10\u00b0 initial start angle, with a 30\u00b0 rotation angle (Fig. 16 (c) and (d)). It can be seen that there is no dominant lay or pattern emerges for the block example, and differing results would occur for the torus example (including the stop-start zone, a known problem with the FDM and other material extrusion processes). The surface roughness is determined for the contour defined by the red line cross section in Fig. 17(a) for the 90\u00b0 shoulder case and the complete virtual case (Fig. 17(b)) for various bead counts (2, 3, 5, 10). The beads are .25 mm in height, and 0" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0001446_icelmach.2012.6349981-Figure9-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0001446_icelmach.2012.6349981-Figure9-1.png", "caption": "Fig. 9. PMSG steady state operation (a) steady state PMSG electric model; (b) phasor diagram for generation action and angle definitions of the PMSG.", "texts": [ " (14) In order to relate the PMSG model with the rectifier model, the dq0 transformation is applied at (9), which gives the model shown in Fig. 8. The similarity between the electric models of the PMSG and the rectifier connected to the grid is evident. The main difference between the electric models are for machines with different inductances values in the direct and quadrature axis, which will be considered in the following. The control actions for CSC with lagging compensation of the stator currents are determined based on the steady state model of the PMSG shown in Fig. 9(a) [25]. Thus, the compensation must be applied based on the value of the inductance of the quadrature axis, since its reactance is lagging the stator currents. Thus, the rectifier voltages in the \u03b1\u03b2-plane are ~v\u03b1\u03b2,s = kIVdc~i\u03b1\u03b2,s \u2212 \u03c9\u0302rL\u0302q~i\u03b1\u03b2,s,q, (15) where~i\u03b1\u03b2,s,q=[\u2212i\u03b2,s i\u03b1,s] T , L\u0302q is the estimated quadrature inductor value and \u03c9\u0302r is the estimated rotor angular speed. Note that the direct axis inductance changes the amplitude of the machine internal voltage (E\u0307a) in steady state as shown in Fig. 9(b). But, its knowledge is not required for the application CSC with lagging compensation as shown in (15). The closed loop control system behavior is analyzed in the dq-plane, where the PMSG quantities are constant variables in steady state and matrix algebra can be applied. In the following analysis the rotor speed \u03c9r and its estimation is considered a constant known value. Another consideration is that the current control loops (inner loops) present much faster dynamics than the power control loop (outer loop), which modified the power reference given by kI " ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0002302_b978-0-12-803581-8.10351-0-Figure46-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0002302_b978-0-12-803581-8.10351-0-Figure46-1.png", "caption": "Fig. 46 Backboard in chassis assembly.", "texts": [ " The preliminary design for the ergonomically contoured backboard can be seen in Fig. 45. From this figure it can be seen that the design for the backboard is curved both horizontally and vertically for the drivers back, but in fact it is actually molded somewhat oversize to enable custom seat inserts for various drivers. In addition to supporting the driver, the molded backboard provides an upper \u201cshelf\u201d as a location to mount the shoulder safety harness. The molded backboard is shown in the monocoque in Fig. 46. This design was extensively modeled to ensure structural integrity, especially in the safety harness mounting locations. An additional outcome of this more complex geometry was an increase in overall stiffness of the monocoque with the backboard bonded in place as an integral component. The primary structural shells of the multi-shell monocoque design and internal bulkheads have been described. However, there remained a need to control the flow of air into the sidepod areas, which would be critical for not only cooling, but for downforce generated by underbody aerodynamics" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0002980_978-94-007-7194-9_71-1-Figure3-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0002980_978-94-007-7194-9_71-1-Figure3-1.png", "caption": "Fig. 3 Kinematic representation of floating-base systems. The root body of the tree structure of the mechanism is free floating in a reference inertial frame R0", "texts": [ " Legged robots are generally modeled from the control point of view as systems composed of rigid bodies, arranged in a tree structure with a base body as their root, called floating base. The displacement of the robot in space is captured with respect to the position and orientation of a reference frame Rb attached to this body, with respect to a given reference inertial frame R0, called world frame. Being free-floating systems, the base is henceforth treated as linked with a 6-DoF virtual unactuated joint to the world, defining the pose qb 2 SE.3/ of Rb with respect to R0, with SE.3/ the special Euclidean group, as illustrated in Fig. 3. The associated twist b is in R 6. The equations of motion for such systems can be derived [43] from the Lagrange formalism and take the form\" Mb Mbj MT bj Mj # \u201e \u0192\u201a \u2026 M.q/ P b Rqj \u201e\u0192\u201a\u2026 P C nb nj \u201e\u0192\u201a\u2026 n.q; / C gb gj \u201e\u0192\u201a\u2026 g.q/ D 06 S C c; (1) where q, called generalized coordinates, parameterizes the configuration of the freefloating system. For the sake of simplicity, legged robots being generally articulated around revolute joints, joints of the tree structure are assumed to evolve in linear configuration spaces in this chapter: that is, qj 2 R n parameterizes the joint configurations in the joint space R n, with n the degree of freedom of the tree structure, and q 2 SE" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0001704_978-3-642-19373-6_6-Figure6.1-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0001704_978-3-642-19373-6_6-Figure6.1-1.png", "caption": "Fig. 6.1 Two ways of doing work on a neutral polymeric gel. Mechanical loads are applied by hanging a field of weights to the network. Chemical loads are applied using a field of pumps to inject solvent molecules into the gel.", "texts": [ " Imagine attaching a field of markers to the polymer chains. While the choice of a reference state is more or less arbitrary, here we simply take the dry polymer network under no mechanical load as the reference state, and name each marker by its coordinate X in the reference state. In the current state at time t, the marker X moves to a place with coordinate x(X, t). We measure both the volume element dV (X) and the area element N(X)dA(X) of integrals in the reference configuration, where N(X) is the unit outward-normal vector. Figure 6.1 illustrates two ways of doing work upon a piece of neutral gel: application of a mechanical force (e.g. by hanging to the polymer network a weight), and attaching to the gel a source of solvent molecules (e.g. by connecting the gel to a reservoir through a pump). Let us first consider the consequence of a field of mechanical forces. In the current state, let b(X, t)dV (X) be the force on a volume element and t(X, t)dA(X) be that on an area element. When the polymer network deforms by \u03b4x, the field of forces does work \u222b b \u00b7 \u03b4xdV + \u222b t \u00b7 \u03b4xdA" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0001189_jtam-2013-0001-Figure4-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0001189_jtam-2013-0001-Figure4-1.png", "caption": "Fig. 4. Working detail and the band-saw blade", "texts": [ " It is calculated by the following expression: (4) Re 3 = 1 2 Pu = Rn b + R\u03a3 2 , where Pu is the feeding force, Rn b is the normal force with which the band-saw blade loads up the detail. This force has the following relation with respect to the force P \u03c4 b : (5) Rn b = mP \u03c4 b , Unauthenticated Download Date | 5/15/17 2:19 PM where m is a coefficient, dependent on the state of the band-saw blade (m = 0.5 for a sharp band-saw blade and m = 1 for a very blunt band-saw blade). R\u03a3 is the total resistance force on the woodworking detail. This force is calculated for each individual case. Figure 4 shows the woodworking detail and the forces and the velocities described above. The weight of the feeding wheel 3 is: (6) Ge 3 = m3g, where m3 is its mass and g is the acceleration of gravity. The belt pulley position 2 is shown in Fig. 5, below: This belt pulley transmits the driving moment from the electric motor to the basic shaft. The force P e 2 creates the driving moment Me 2z with respect to the axis of rotation. It is calculated from the known expression, cited below: (7) Me 2z = P e 2 r2" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0000568_s11431-010-4273-0-Figure3-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0000568_s11431-010-4273-0-Figure3-1.png", "caption": "Figure 3 3-D sketch of the air cannon model.", "texts": [ " However, the exhausting process is the unsteady flow of complicated feature, especially in the nozzle inlet, so the actual impulse force is less than the theoretical analytic result. Above analysis indicates that vessel pressure, piston sleeve inlet, nozzle diameter and Laval nozzle with different parameters are the key factors for air cannon performance. Numerical simulation is carried out to investigate the influence of these parameters. The computation is performed using NUMECA software with three-dimensional turbulent Navier-Stokes equation. The air cannon model is shown in Figure 3 and a butterfly grid is used to improve grid quality. About two million grid points are used with y+ in a range of 1\u201310. The turbulent model is Standard k- model. Multi-grid and implicit residual iteration method is used. Computation results show that the nozzle mass flow, outlet static pressure and the impulse force increase almost linearly with the vessel pressure. The impulse force increases with the length of piston sleeve inlet as shown in Figure 4, especially when the length increases from 37 to 51 mm" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0001607_0021998312463455-Figure4-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0001607_0021998312463455-Figure4-1.png", "caption": "Figure 4. Compressive forces distributed on the intersection area.", "texts": [ " FA and FN are the two components of FC. is the bending angle of yarn D. Thus, equilibrium equations at point M can be written as FT cos \u00fe FA \u00bc FT FN \u00bc FT sin \u00fe P Aint \u00f01\u00de where P is the internal pressure of the braided composite actuator and Aint is the overlapping area by two intersected yarns. The out-of-plane component FN is the compressive force for fiber compaction. The compressive forces are simplified to point forces and we assume that the compressive forces are distributed on a small parallelogram area, as shown in Figure 4. As the fibers rotate, the area of the parallelogram changes, which is included in this model. In the figure, d is the width of the fiber yarn and is the braiding angle by the fiber and the generator of the cylinder. As shown in Figure 5, the solid lines and the dashed lines are the yarns before and after compaction, respectively. The thickness of the sleeve is reduced by . If the fiber yarns are inextensible, the length of the element edge is increased. Because the element dimension is much smaller than the actuator\u2019s radius, the element can be treated as flat" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0002491_00325899.2017.1344451-Figure13-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0002491_00325899.2017.1344451-Figure13-1.png", "caption": "Figure 13. Images showing (a) the extrusion tool made of AISI H13 grade manufactured by SLM, (b) rubber product extruded using the tool.", "texts": [ " CAD models of the tool components used as die for plastic extrusion: (a) base component (b) and (c) other component of the tool shown in different angles. Courtesy of Trelleborg Sealing Profiles AB. D ow nl oa de d by [ U ni ve rs ity o f Fl or id a] a t 0 9: 37 0 5 A ug us t 2 01 7 machining. Functional testing of the tool for production of rubber profiles was carried-out in Trelleborg Sealing Profiles AB (Trelleborg, Sweden) with satisfactory results. The final extrusion tool and the extruded rubber profile are shown in Figure 13. By means of shear cell testing it is possible to correlate between certain flow parameters of the powder and its flowability during the recoating operation in laser based powder bed fusion processes. This investigation demonstrates that the combination of bulk density with the degree of cohesion and the flow function can be used as indicators for powder flowability during SLM processing. Both tumbling and thermal treatment improve the flow properties of the H13 tool steel powder investigated in this study" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0002642_978-3-319-66866-6_2-Figure8-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0002642_978-3-319-66866-6_2-Figure8-1.png", "caption": "Fig. 8. Relevant design guidelines for the optimized model of the pedal crank", "texts": [ " The optimization of the preliminary design results in a homogeneous stress distribution with maximum stresses (r 240 N/mm2) below the allowable values. The preliminary design is finally evaluated in comparison to the design guidelines. Therefore, the orientation and position of the model in the process chamber has to be defined. After weighing the criteria for production time, accuracy, loading capacity due to anisotropy, post-process effort and avoiding damage by the coater, the orientation and position is set as depicted in Fig. 8. Due to the horizontal positioning various areas result, which are limited by design guidelines. For example, down skin surfaces - normal vector is negative in respect to z direction - have to be investigated with regards to necessity of support structures. Supports are unavoidable between the component and building platform, but can be easily removed in post-process. In the components\u2019 interior, supports are forbidden, because removing them is impossible. Consequently, cavities as well as internal structures are limited by maximum overhangs and angles. Considering this restriction, all down-skin surfaces can be manufactured without lowering in negative z direction during manufacturing. Besides specific values, general guidelines for removing excess material have to be considered. As depicted in Fig. 8, cleaning openings are provided on a slightly loaded area of the surface. The sizing of the openings is limited by the minimum diameter in z direction. Furthermore, the consideration of minimizing stress peaks and avoiding supporting structures is necessary during designing. The optimized model is manufactured with AlSi10Mg using an Eosint M280 machine. As depicted in Fig. 9-a, the process parameters for core and skin exposure are set. Figure 9-b shows the result from manufacturing process after thermal (300 \u00b0C for 2 h) and mechanical post-processing (removing support structures and shot peening)" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0002200_978-3-319-51222-8-Figure6.20-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0002200_978-3-319-51222-8-Figure6.20-1.png", "caption": "Fig. 6.20 Gravity load in a wall plate a principal stresses b stresses \u03c3yy in some horizontal sections (BE-solution)", "texts": [ " Better to round out the corners of the plate, then the streamlines can rotate, and this way, they have it easier to balance the vertical load, see Figs. 6.16 and 6.17. No need to develop infinite stresses. Note also that couples pose no problem, see Fig. 6.18. But a cantilever plate is not so special. Even at such seemingly harmless points as the reentrant corners of openings in a wall plate, see Fig. 6.19, singularities develop. Normally, the meshes are too coarse for the singularities to shine through, but when you really go all the way as in Fig. 6.20, then you begin to notice the infinite stresses. can be relevant to the design (stress intensity factors). At corner points, solutions often become singular. This happens also to influence functions, and so, they too are affected by these singularities because in some hidden sense also finite elements are boundary elements and the negative effects of the singularities on the boundary are transmitted via the influence functions into the domain. Singularities pollute the FE-solution via this \u2018boundary element\u2019 mechanism" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0000437_s11223-013-9461-2-Figure1-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0000437_s11223-013-9461-2-Figure1-1.png", "caption": "Fig. 1. Scheme of the rotor and the ball bearing: A and B are supports.", "texts": [ " This paper also proposes the equations obtained to determine the nonlinear radial and axial reaction force in ball bearings as well as the way of their linearization. The objective of the present paper is to investigate the nonlinear free vibrations of the elastic rotor with multi-disks that are located asymmetrically relative to the shaft supports. The rotor rotates on angular ball bearings. The rotor model considers the nonlinear characteristics of ball bearings. The NNM method is used to investigate the vibrations. 1. Equations of System Motion. Figure 1a shows the system with N D disks under study. The origin of coordinates is at the support A. Denote the shaft length by l and displacement of the shaft center line in the direction of coordinate axes by ux , u y , and uz . The shaft vibrations are expanded by the natural modes of the linear system [5]: u t x t n l x t l x t l l u x n N N n N y ( , ) ( )sin ( ) ( ) , 1 2 1 ( , ) ( )sin ( ) ( ) , t y t n l y t l y t l l n n N N N 1 1 2 (1) 316 0039\u20132316/13/4503\u20130316 \u00a9 2013 Springer Science+Business Media New York Podgornyi Institute of Problems of Mechanical Engineering, National Academy of Sciences of Ukraine, Kharkov, Ukraine", " The components of the elastic reactions of the bearing along the x y z, , axes [8] are of the following form: P K x y z zx N b [ cos cos cos sin ( )sin ] / 1 0 3 2 cos cos , cos cos cos sin ( P K x y z zy N b [ 1 0 3 2 1 )sin ] cos sin , cos cos cos / P K x yz N b [ sin ( )sin ] sin , / z z0 3 2 (5) K P N z b 3 2 0 0 3 2 5 2/ / sin . Here, x , y , and z are the displacements of the center of the internal ring relative to the center of the external ring, is the number of the generalized coordinate of the bearing journal, Fig. 1b shows the angles and , and N b are the number and amount of balls, respectively, P0 is the axial preload force, z0 is the axial displacement of the internal ring relative to the external ring depending on the force action P0 determined from geometric relations within contact zone: z R w w dr b0 1 22 ( )sin , where Rr is the radius of the races in the bearing rings, w1 and w2 are the convergences between internal and external rings and the ball towards the contact line caused by preliminary axial contraction, and db is the ball diameter" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0002450_978-3-319-60867-9_17-Figure4-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0002450_978-3-319-60867-9_17-Figure4-1.png", "caption": "Fig. 4. Independent variables of functions (1) and (4)", "texts": [ " In previous experimental studies on the walking of the Bioloid robot [8] it was observed that the balancing of the robot was more sensitive to changes in values of the lateral displacements dyp of the pelvis. Consequently, in this work we use the displacements dyps and dypd of point Op of the pelvis in direction yW, associated to the single support phase and double support phase, respectively. Thus, we define: w1 dyps \u00f07\u00de w2 dypd \u00f08\u00de The meaning of these independent variables in the motion of the biped can be appreciated in Fig. 4. The six sets of dyps and dypd considered are given in Table 2. Thus, minimizing function (6) by using the function fmincon of Matlab\u00a9 for these sets of parameters we obtain b0 \u00bc 0:020, b1 \u00bc 0:015 and b2 \u00bc 0:085. The function fmincon is based on the Interior Point algorithm [9]. Such a function minimize constrained non-linear functions. The plot of function (4) with the obtained values of b0, b1, b2 is shown in Fig. 5. The optimal values of w1 and w2 are gotten by using the partial derivatives of (4) with respect to w1 and w2" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0002900_slct.201701018-Figure5-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0002900_slct.201701018-Figure5-1.png", "caption": "Figure 5. Typical setup of the all solid contact Cs+ ISE.", "texts": [ " Plot of ISE EMF versus time was constructed and it is shown in Figure 3. Mean response time of ISE was found to be about 20 seconds at limiting slope value of 1 mV/min (DE/Dt). PANI was deposited on the platinum (Pt) electrode by electrochemical method employing cyclic voltammetry (CV) technique using Pt electrode as a working electrode, Ag/AgCl/ 3 M KCl as a reference electrode and Pt wire as a counter electrode. The cyclic voltammogram of PANI deposition (15 cycles) is given in Figure 4. The typical setup of the all solid contact ISE is given in Figure 5. Details of response for Cs+ with two types of ISEs i. e. ISE 12 with PANI and ISE 13 with Au nanoparticles doped PANI are given in Table 1 as well as in Figure 2. It can be seen from the Table 1 that best response was observed with ISE 13 i. e. Au nanoparticles doped PANI. Hence this ISE was selected for further studies. The developed ISE gave linear response in the range 10 7- 10 2 M of Cs+ with Nernstian slope of 56.3 0.8 mV/decade. The detection limit of ISE is 7.7 x 10 8 M Cs+. The response time of ISE 13 was estimated as mentioned above and it is about 20 seconds" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0003660_iemdc.2019.8785146-Figure4-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0003660_iemdc.2019.8785146-Figure4-1.png", "caption": "Fig. 4 Winding pattern of a multi-three-phase PMSM", "texts": [ " We believe that we can construct a motor system that is robust against major manufacturing variations. First, we discuss the theory regarding the harmonic component of no-load stator voltages that has the static eccentricity information and next explain some results of the magnetic analysis using Finite Element Analysis (FEA). 978-1-5386-9350-6/19/$31.00 \u00a92019 IEEE 671 II. THEORETICAL CALCULATION In this chapter, we formulate theoretically the relationship between no-load stator voltage and static eccentricity. An example of a multi-three-phase PMSM is shown in Fig. 4. The angle between two coils neighboring each other is described as [degree]. The coils of each phase (U, V, W) are wound as shown in Fig. 4 and the phases are identified by three colored arrows. The flowchart of detecting the static eccentricity is shown in Fig. 5. First, we detect line-line voltages per group and derive the maximum and minimum amplitude of the line-line voltages in all groups; that is, the derived line-line voltages are maximum and minimum in the line-line voltages of all groups. The difference between the maximum and minimum means the amount of the static eccentricity, so we can finally identify the direction of the static eccentricity using these and the position of the coils", " However, with static eccentricity, Auk, Avk, and Awk are different from each other and \u03b8u(N), \u03b8v(N), and \u03b8w(N) are not equal, so Auv N , Avw N , and Awu N are not equal to 0 and the harmonic with N=3n is generated. In other words, since the harmonics with N=3n of the line-line voltages ( Auv N , Avw N , Awu N ) increase remarkably when static eccentricity exists, their detection is easier than the other harmonics. Considering line-line voltages ( Auv N , Avw N , Awu N ) in a group with static eccentricity, the harmonic with N=3n of the line-line voltage with neighboring coils (for example, U and V in Group1 of Fig. 4) or where coils are put across the minimum or maximum magnetic gap is nearly 0 because of the small difference between the harmonics with N=3n of the phase voltages (Au N , Av N , Aw N , and \u03b8u N , \u03b8v N , \u03b8w N ). Therefore, it reaches the minimum in all harmonics with N=3n of the line-line voltages. On the other hand, the harmonic with N=3n of the line-line voltage in which coils are not neighbors (for example, W and U in Group1 of Fig. 4) becomes larger than that of the other coils, so it becomes the maximum. In other words, when static eccentricity exists, the ratio of the maximum and minimum of the harmonic with N=3n of the line-line voltages ( Auv N , Avw N , Awu N ) increases remarkably because its denominator converses in 0. It is therefore useful as an indicator of static eccentricity. In order to inspect the facility of detection, we establish the theoretical model of the motor with static eccentricity shown in Fig. 6. The magnetic flux density at the magnetic gap with static eccentricity Bg and that without static eccentricity Bgm of the coil placed at a position where the mechanical angle is \u03b2 [degree] can be described as equations (13) and (14), where Br [T] is the residual magnetic flux density, hm [mm] is the magnet thickness, gm [mm] is the magnetic gap length without static eccentricity, gr [mm] is the amount of static eccentricity, \u03bcrec is the recoil permeability (=1" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0000806_6.2010-1206-Figure6-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0000806_6.2010-1206-Figure6-1.png", "caption": "Figure 6. CAD representation of the fuel cells and solar panels powered concept", "texts": [], "surrounding_texts": [ "C. Fuel Cells Propulsion Fuel cell propulsion has been demonstrated by Boeing2 on a Dimona aircraft that has been\nmodified in order to keep the weight down. Other studies on the usage of fuel cells have been presented1-3. The usage of fuel cells have been proposed in many case for mide size UAV4, and Georgia Tech have been flying a UAV demonstrator powered by fuel cells.\nHere 3 different configurations are presented and compare to the Lancair Legacy. The main differences between them being a compromise between range or higer cruise speed.\nAmerican Institute of Aeronautics and Astronautics\n7", "D. Fuel Cell and Solar Panel Propulsion One major challenge with solar power aircraft is the take off where the highest power is\nneeded, and that case will often be one man drivers for the final design. By combining a fuel cell, able of delivery higher power than solar panel alone, during the take off, the aircraft can be design around the available solar power at cruise condition.\nSimilar concept have been presented by Lisa-Aviation, the concept has a tendency to be close\nto a motor glider. But the extra power available from the fuel cells reduced the issues with having only solar power to rely on. One alternative to that solution could be combining solar power and batteries, but that solution does not offer any extra benefit compare to combining fuel cells and solar cells. The characteristics of the presented aircraft are displayed in table 5.\nAmerican Institute of Aeronautics and Astronautics\n8", "V. Demonstrator Two demonstrators have been built to further investigate the concepts. The first one is a solar panels powered aircraft (visible on the right hand side in Fig. 7) while the second carries a small hydrogen fuel cell capable of producing 3W (left hand side in Fig. 7).\nAmerican Institute of Aeronautics and Astronautics\n9" ] }, { "image_filename": "designv11_62_0003671_02670836.2019.1647383-Figure1-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0003671_02670836.2019.1647383-Figure1-1.png", "caption": "Figure 1. Positions of test samples in prepared CrCuFeMnTi alloy ingot. The thicknesses of the test samples are all 3mm. The dimensions of A1, B1, C1, A2, B2, andC2plates are all larger than 10\u00d7 10mm2 to ensure rectangular sheets of dimensions 10\u00d7 10\u00d7 3mm3 can be processed. \u20181\u2019 and \u20182\u2019 in the arch-shaped sample D represent different paths of the microstructure observations.", "texts": [ " Therefore, owing to the lack of understanding of the inhomogeneity of HEAs ingots fabricated by vacuum arc melting, the present work aims to investigate the microstructures, chemical distribution, and hardness of a dual-phase as-cast CrCuFeMnTi HEA [26] ingot from the surface to the interior, and provides an effective way to improve this microstructure inhomogeneity. The equiatomic CrCuFeMnTi alloy ingot was prepared by vacuum arc melting, and the elements of Cr, Cu, Fe, Mn, and Ti with purity higher than 99% were used as raw materials. The melting processes were conducted under anAr atmosphere. The button-shaped alloy ingot was remelted four times in a water-cooled Cu crucible to ensure fullmixing of the compositional elements. For the subsequent study on microstructure inhomogeneity, four different areas, indicated by \u2018A\u2019, \u2018B\u2019, \u2018C\u2019, and \u2018D\u2019 in Figure 1 were chosen. In particular, \u2018A\u2019 and \u2018B\u2019 can be considered as the same position owing to symmetry and \u2018C\u2019 is located at the edge of the alloy ingot. Six rectangle sheets of dimensions of 10\u00d7 10\u00d7 3mm3 (cut from the regions \u2018A1\u2019, \u2018B1\u2019, and \u2018C1\u2019, which are close to the surface of the alloy ingot, and regions \u2018A2\u2019, \u2018B2\u2019, and \u2018C2\u2019, which are located at the centre of the ingot in Figure 1) and an arch shape of thickness 3mm (\u2018D\u2019 in Figure 1) were prepared from the as-cast alloy. The microstructures of the alloy were characterised using an OLYMPUS GX71 optical microscope (OM) and ZEISS SIGMA scanning electron microscope (SEM), and chemical analyses were implemented by energydispersive spectrometer (EDS) installed in the SEM. Before themicrostructure observations, the sample surfaces were processed by an etchant consisting of 2% HF, 3% HNO3, and 95% H2O (by volume). Hardness tests were conducted using an HV-50A Vickers hardness tester with a load of 10N and holding time of 15 s. Finally, the electrochemical corrosion properties of variousCrCuFeMnTi alloy specimenswhichwere cut from different areas (Figure 1) were evaluated by acquiring their potentiodynamic polarisation curves in a 3.5% NaCl solution, using an electrochemical workstation (AUTOLAB) at room temperature. A three-electrode system was used for this experimental setup: the working, reference, and auxiliary electrodes were the alloy specimens, a saturated calomel electrode, and a platinum electrode, respectively. A polarisation scan rate of 100mVmin\u22121 was employed in this work. Microstructure inhomogeneity of as-cast HEAs The OM images of the as-cast CrCuFeMnTi alloy along path \u20181\u2019 and \u20182\u2019 in the arch-shaped specimen D are presented in Figure 2. Referring to the sampling position of specimen D (Figure 1), there is a trend of declining cooling rate from positions \u2018I\u2019 to \u2018VI\u2019 and \u2018I\u2019 to \u2018III\u2019 along paths \u20181\u2019 and \u20182\u2019, respectively. Typical dual-phase microstructures with dendritic (DR) and interdendritic (ID) regions are observed for the as-cast CrCuFeMnTi alloy, and there is an obvious microstructure coarsening with the decline in the cooling rate (from \u2018I\u2019 to \u2018VI\u2019 Figure 2. Microstructures of as-cast CrCuFeMnTi alloy along paths \u20181\u2019 and \u20182\u2019 in the arch-shaped specimen D. Figure 3. SEM images for as-cast CrCuFeMnTi alloy along path \u20181\u2019 with different depth from the surface: (a) 2mm; (b) 5mm; (c) 9mm; (d) 13mm; (e) 16mm; (f ) 19mm", " In order to further investigate the microstructure inhomogeneity in the as-cast CrCuFeMnTi alloy and the effect of the heat treatment improvement, electrochemical corrosion testswere performed to estimate the effect of the microstructures in different areas of the alloy on the corrosion properties, and the corresponding sampling positions including \u2018A1\u2019, \u2018B1\u2019, and \u2018C1\u2019 Figure 7. Changes of chemical compositions for DR (a) and ID (b) regions with the depth from the surface for as-cast and asannealed CrCuFeMnTi alloys. are shown in Figure 1. The potentiodynamic polarisation curves for various alloy specimens in a 3.5% NaCl solution are shown in Figure 8. The measured values of the electrochemical parameters of these alloy specimens are listed in Table 1. It is clearly evident that there is no obvious passivation zone in the as-cast specimens immersed in a 3.5% NaCl solution, whereas the polarisation curves for the annealed specimens show small steady-state passivation regions. This indicates that the annealed CrCuFeMnTi alloy would present better corrosion resistance owing to the formation of a dense passive film on the alloy surface to prevent polarisation current increasing with the increasing potential. All the specimens in the different sampling positions exhibit corrosion potentials in the range from \u22120.33 to \u22120.24V, and corrosion current densities between 1.27 \u00d7 10\u22126 and 2.36\u00d7 10\u22126 A cm\u22122. According to the sampling positions of the specimens shown in Figure 1, \u2018A1\u2019 and \u2018C1\u2019, which are cut near the surface of the alloy, ingot undergo a higher solidification rate than \u2018A2\u2019 and \u2018C2\u2019 during the alloy preparation process. Meanwhile, compared with \u2018A2\u2019, the cooling rate in the \u2018C2\u2019 region would be higher owing to its near-edge sampling. Hence, the cooling rates in descending order for the six as-cast specimens can be considered as: A1 (=B1) \u2248 C1 > C2 > A2 (=B2). Combining Figure 8 and Table 1, it can be seen that the corrosion tendencies of the alloy specimens with higher cooling rates (\u2018A1\u2019 and \u2018C1\u2019) are significantly lower than those of specimens with lower cooling rates (\u2018A2\u2019 and \u2018C2\u2019), because there are obviously higher corrosion potentials in the \u2018A1\u2019 and \u2018C1\u2019 specimens" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0000836_detc2013-13594-Figure1-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0000836_detc2013-13594-Figure1-1.png", "caption": "Figure 1. ILLUSTRATIONS OF: (a) MESH COUPLING OF HYPOID GEAR PAIR, AND (b) CONTACT CELLS ON ENGAGING TOOTH SURFACE.", "texts": [ " The load sharing characteristic is examined by load dependent contact ratio and mesh stiffness, and the torque load effect on dynamic responses is also examined. The hypoid gear mesh model, including mesh point, lineof-action, mesh stiffness and kinematic TE can be derived from the load distribution results calculated by a 3-dimensional quasi-static loaded tooth contact analysis program for hypoid and spiral bevel gears [8]. This program combines the semianalytical theory with finite element method [9], which can solve the gear tooth contact problem very efficiently. A hypoid gear pair with multiple contact interfaces is shown in Figure 1(a), and Figure 1(b) shows contact cells on each tooth pair. The position vector of the contact cell in the mesh coordinate system is ),,( iziyixi rrrr , the contact force is if , and the normal vector is ),,( iziyixi nnnn . The total contact force is calculated by summing the contact forces on each contact cell (assuming there are N contact cells) , 1 N i iijj fnF ,2 z xj jFF ),,,( zyxj (1) The line-of-action vector zyx nnnL ,, can be obtained from ,/ FFn jj ).,,( zyxj (2) The total contact moment is given by N i iii nrfM 1 )( (3) The mesh point mmm zyxR ,, can be obtained from N i i N i iiy m f fr y 1 1 (4a) 4 ,/)( ymxzm FyFMx (4b) ,/)( xmzym FxFMz (4c) The translational loaded and unloaded transmission errors Le and 0e are the projections of corresponding angular transmission error along the line of action" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0002329_s40997-017-0082-4-Figure3-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0002329_s40997-017-0082-4-Figure3-1.png", "caption": "Fig. 3 Dimensional accuracy of the shaft geometry (top left pictures show point cloud image of complete shaft)", "texts": [ " The results of the roughness measurements, their averages and standard deviations for the shafts are provided in Table 6. \u2013 Because of dimensional deviation of the shafts, 3D scan of cylindrical shaft geometry was performed by an optic camera. Shafts specimens with three values of the roughness have been scanned, and point cloud images were created. Then, some surfaces were fitted on the measured points of the points cloud images in the Geo-Magic Studio software. Geo-Magic Studio software was used to fit the best surfaces to the contact area of the shafts. Figure 3 shows contour plot of deviations of the fitted surfaces. This figure represents less than 1-lm variance in contact surfaces. It should be noted that the surface fitting in software is a kind of smoothening of surfaces and deviation contour represents maximum distance between fitted surface and original place of the measured cloud points model. Digitalization of the profile and surface fitting was used to construct a model with the precision of 1 lm. In Fig. 3, the gray area in the center of the shaft represents that this surface is considered as a smooth face and there are not any measured points for surface fitting. It should be remembered that just contact surfaces of the joint parts were effective in the numerical analysis. So it was not necessary to model non-contact parts of the shafts, precisely. Noncontact parts of the real shaft at the start and end portions were deleted and assumed as perfect surfaces for further analysis. Derived models from Geo-Magic Studio were shells" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0002774_pvp2017-65992-Figure5-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0002774_pvp2017-65992-Figure5-1.png", "caption": "FIGURE 5. DIFFERENTIAL THERMAL RESPONSES OF THIN WALL MODEL AFTER 1.5 DEPOSITION PASSES WITH 2000 WATT AND 500 WATT LASER POWERS USING 0.001 M AND 0.0025 M LASER BEAM DIAMETERS, RESPECTIVELY.", "texts": [ " For one 0.05 second time step, higher laser power and a larger beam diameter resulted in a larger volume of activated elements (i.e. more deposited material) and higher substrate and deposited part temperatures as shown in Figure 4. Later in the build process, thermal differences between 2000 W and 500 W laser powers were also evident. The 2000 W laser model showed a larger melt pool, and significantly higher temperatures in both the deposited region and the substrate compared to the 500 W model (Figure 5). For the variable laser power thin walled build model, stresses began to build during deposition, and slightly increased in magnitude following cooling to room temperature. Generally, the magnitudes of axial stress increased once reaching the final build state at room temperature as compared to during the build 4 Copyright \u00a9 2017 ASME Downloaded From: http://proceedings.asmedigitalcollection.asme.org/ on 11/10/2017 Terms of Use: http://www.asme.org/about-asme/terms-of-use (Figure 6). Axial stress magnitudes were lower near the beginning and end of each deposition pass as compared to the middle of the pass" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0000558_icrera.2012.6477470-Figure2-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0000558_icrera.2012.6477470-Figure2-1.png", "caption": "Figure 2. Conception picture of tidal power generation system.", "texts": [ " Attempts have been made to investigate and artificially reproduce the system of lubrication in natural synovial joints. Typical application areas are in the medical field [3-5], in which next generation joint prosthesis is expected to demonstrate ultralow friction and wear. Natural synovial fluid is a water-based liquid composed of biodegradable constituents. The proposed environmentally friendly bearing produces low friction in various lubrication modes and has potential for application in systems such as a tidal power generation system (Figure 2), where mechanical loss should be reduced to increase power generation efficiency and marine pollution should be avoided. The hydration lubrication observed in articular cartilage has been frequently reproduced using gels or porous materials as bearing materials. One such material is a polyvinyl alcohol (PVA) hydrogel, which is one of the few polymers with hydrophilic properties. However, PVA hydrogel has a shortcoming in that it is a low wear-resistant material. Therefore, in this work, improvement of the wear resistance was attempted by performing a chemical cross-linking reaction of PVA to produce polyvinyl formal (PVF) [6]" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0000197_978-3-642-28768-8_65-Figure1-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0000197_978-3-642-28768-8_65-Figure1-1.png", "caption": "Fig. 1 MBS model and details of the drive train", "texts": [ " Moreover, the MBS approach provides more flexibility to include additional phenomena such as the excitations coming from aerodynamic blade loads or the flexibility of different components (shafts, tower, etc.) although in this work the attention is focused on the dynamics of the gear train. The model developed includes a planetary stage and two ordinary stages with variable meshing compliance as well as bearing flexibility. Special attention is given to the run up and emergency stop as dynamic forces are critical in these maneuvers. Information about the behavior of loads induced by the dynamic interactions is analyzed. The whole MBS model developed in MSC-ADAMS, is presented in Fig. 1 with details of the gearbox and kinematics scheme. The blades with the hub are considered as a rigid body with lumped mass and inertia. Wind loads are included as external forces on the rotor hub neglecting aerodynamic loads due to the structural behaviour of blades and the aero-elastic coupling. Buoy restoring forces are included at the base of the tower using specific Single Forces defined in the MBS code environment. Regarding gearbox, shafts and gears are lumped and assumed to be rigid bodies while gear meshing forces are modelled by a variable stiffness spring following the approach described later. The rotor hub is connected thorough the main shaft to the carrier of the planetary stage (see Fig 1). The carrier moves planets and transmits the power by the fixed ring and the sun to the low-speed shaft. Then, two ordinary stages are used to increase the rotational speed, up to the desired value at the generator side which is included by means a representative mass. In order to simplify, all shafts have been considered as rigid and the main shaft is the only one supported by flexible bearings, while rotational joints are used for the others. Flexible bearings are included in the MBS model as bushing joints defining the corresponding stiffness and damping values neglecting cross terms" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0001348_tasc.2010.2042160-Figure1-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0001348_tasc.2010.2042160-Figure1-1.png", "caption": "Fig. 1. The cross-section of an induction motor and its flux plot at", "texts": [ " One column of matrix is obtained: (36) Similarly, if the charge excitations are set as: (37) we have: (38) If the charge excitations are set as: (39) we have: (40) After the matrix is obtained, the matrix can be obtained using (32). With this algorithm, the coefficient matrix of the FEM equations can be kept unchanged. Only a multi right hand side (RHS) problem needs to be solved. By using the multi-RHS algebraic solvers, the computing time required to extract the capacitance matrix can be greatly reduced. The developed FEM is applied to evaluate the effects of a LC filter of an inverter driven 250 kW, 380 V, 2-pole, 50 Hz induction motor (Fig. 1). The carrier frequency of pulse-width-modulation (PWM) is 4 kHz. Step voltages representing the PWM wavefronts with a rise-time of 50 ns and a magnitude of 400 V are applied to the windings. The responses of the nodal voltages on the conductors are shown in Fig. 2. It reveals that the voltage drops on the conductors are non-uniform as the rise-time is very short because of capacitance effect. The responses of the . nodal voltages with a -shaped LC filter (40 mH, 60 ) connected between the output of the inverter and the stator windings are shown in Fig" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0002302_b978-0-12-803581-8.10351-0-Figure122-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0002302_b978-0-12-803581-8.10351-0-Figure122-1.png", "caption": "Fig. 122 Schematic of fixturing developed to assemble the monocoque shells and bulkheads.", "texts": [ " Since the monocoque shells were built from the outside surface inward, and the bulkheads were built from the inner surface of the bonding flange outward, the potential exists for interference. Some of this was able to be evaluated during dry assembly. The following section discusses the assembly of the monocoque. A jig to assemble the monocoque was created that allowed accurate assembly. It consists of pieces for the right and left halves, as well as a bulkhead placement jig as shown schematically in Fig. 122. Tolerance screws were located that support the monocoque in the jig, as well as take up any misalignment of the pieces from fabrication. The monocoque shells were used to adjust the jig and set the tolerance screws. The top and bottom supports are removable to allow the adhesion of the top and bottom pieces, while the monocoque side shells are in the jig. The jig can also be inserted into the oven for accelerated curing of the adhesive. All of the monocoque pieces were dry fit to insure accurate placement" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0000224_kem.504-506.1305-Figure1-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0000224_kem.504-506.1305-Figure1-1.png", "caption": "Figure 1: (A)Workpiece dimensions [mm]; (B) dynamometer equipment fixture; (C) tool-holder; (D) machining operation.", "texts": [ " [4], conducted dry turning experiments to clearly identify the tool wear mechanisms when a commonly used coated cemented carbide tool cuts Inconel 718. Jindal et al. [5] studied the relative merits of PVD TiN, TiCN and TiAlN coatings on cemented carbide substrate (WC\u20146wt.% Co alloy) in the turning of Inconel 718. Prengel et al. [6] performed Inconel 718 turning tests with a coolant and different PVD coated carbide cutting tools. Experimental set-up and operating parameters In the present work an Inconel 718 ring was used for the experimental tests. The ring (Fig. 1 A) was obtained by laser welding of two forged rings produced by a triple melt process consisting of vacuum induction melting (VIM) followed by electroslag consumable electrode remelting (ESR) and vacuum arc consumable electrode remelting (VAR) (AMS 4530 M). The chemical composition and the physical properties are those indicated in the standard SAE-AMS 5663M - 2004. Cutting tests with high performance mineral oil-based coolant (Blaser Swisslube Cut 2000 Universal) have been conducted on a high speed CNC MAZAK Q-turn lathe using S-quality rhomboidal 55\u00b0 coated carbide tool with ISO code DNMG 15 06 16 S quality and the following geometry characteristics: rake angle \u03b3 = 1.7\u00b0, clearance angle \u03b1= 0\u00b0 and approach angle \u039ar = 90\u00b0. The tool composition for the bulk and coating are: \u2022 Bulk composition: Co (10.2 wt.%); WC (89.3 wt.%); TaC (0.2 wt.%); NbC (0.3 wt.%) \u2022 Coating composition and layers thickness: TiCN (2.2 \u00b5m); Al2O3 (1.5 \u00b5m); TiN (0.5 \u00b5m). The tool holder was mounted on a Kistler 9257 piezoelectric dynamometer (Fig. 1 B), for forces detection: Fcutting, (Fx) and Fthrust (Fz). Moreover, a customized steel support (Fig. 1 C) was designed and realized to fix the ring in the lathe as illustrated in Fig. 1 (D). The workpiece has been initially worked to remove the surface unevenness and to bring the outer diameters to the reference value of 347 mm. During the test the ring was machined by different working steps, reducing the outside diameter from 347 mm to 332 mm. For the industrial privacy that characterizes this work, the range of the values than those experimentally used is reported in Table 1 as dimensionless values for the DOC and indicated with the terms: very low (VL), low (L), medium (M), high (H) and very high (VH)" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0002791_b978-0-12-812138-2.00003-9-Figure3.59-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0002791_b978-0-12-812138-2.00003-9-Figure3.59-1.png", "caption": "FIGURE 3.59", "texts": [ " The maximum torque occurs at a torque angle of 90 electrical degrees. This is called the pull-out torque, which indicates the maximum value of torque that a synchronous motor can develop without pulling out of synchronism. In general, its value varies from 1.25 to 3.5 times the full-load torque. In actual practice, however, the motor will never operate at a torque angle close to 90 electrical degrees because the stator current will be many times its rated value at this operation. Next, we will discuss the salient pole rotor synchronous motor. As we can see in Fig. 3.59, a salient pole synchronous motor has a nonuniform air gap due to the protruding rotor poles. In this type of motor, the amount of the flux produced by the stator current varies according to the position of the rotor. A higher flux is generated along the poles in comparison to between the poles because the reluctance of the flux path is low along the poles and high between the poles. To consider the saliency of the rotor in a synchronous motor model, we need to define the d and q axes as the following" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0003590_1.5112554-Figure2-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0003590_1.5112554-Figure2-1.png", "caption": "FIGURE 2: Hybrid bearing bushing in half section and meshed FE model.", "texts": [ " In this paper, the Dang Van stress criterion is used [12]. A detailed methodological description of the calculation is presented in earlier work by the authors [13, 14]. Using the aforementioned materials, the hybrid bearing bushing is modelled by means of the finite element (FE) software Ansys APDL. With a parametric FE model of the hybrid bearing bushing, stress states in the joining zone and every other part of the bearing bushing can be calculated as a function of the local geometry (layer height of the steel). Figure 2 shows a half section of the hybrid bearing bushing with the FE model overlayed. A 36\u00b0 wide ring segment with a complete rolling element was modelled to represent the contact. The contact area between rolling element and bearing bushing is discretized with a mapped mesh of element type CONTA173 for the rolling element and TARGE170 for the bearing bushing. The element size in the contact is 38x20x17 \u00b5m (lengt x width x height) and the coefficient of friction is \u00b5 = 0.015 [15]. The non-lifedeterming part zones are discretised by element type SOLID185", " The residual stress, the layer thickness of the case hardening steel and the load are varied, whereby the influence of the parameter variation on the fatigue life is investigated. An incremental variation of layer thickness h from 0.5 to 4.5 mm was carried out. The parameter h is defined as the distance between the contact point of the rolling elements/steel raceway of the bearing bushing and the aluminium. It thus corresponds to the height of the case hardening steel in radial direction (purple in Fig. 2). The layer thickness of the manufactured bearing bushing is approx. 4.5 mm. Different residual stress depth profiles, resulting from the manufacturing processes as shown in Fig. 3, can be regarded in the simulation. The external load was applied as axial force Fa on the inner rings shoulder and varied between 15 kN and 8.9 kN. This corresponds to a high to very high load on the (commercial) angular contact ball bearing without any tailored forming component. 040020-3 Figure 4 shows the resulting von-Mises equivalent stress distribution in the contact zone between rolling element and bearing bushing" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0003741_012077-Figure3-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0003741_012077-Figure3-1.png", "caption": "Figure 3. Mobile house in: (a) compact configuration (on a trailer); (b) deployed configuration [9].", "texts": [ "1088/1757-899X/591/1/012077 If we consider 1 the rotational angle of the driver link, the area A1 of a wall formed by this mechanism in the deployed configuration will be equation (3): 1 2 1 sin2 AB lA (3) considering that ADAB ll . Another view of the mechanism with overlapped compact, partially deployed and totally deployed configurations is presented in figure 2. Two examples of deployable structures that may use this mechanism will be given here. The first example is a mobile house (see figure 3). Another application of the mechanism described here road barrier, represented in folded, partially deployed and totally deployed configurations (figure 4). Modern Technologies in Industrial Engineering VII, (ModTech2019) IOP Conf. Series: Materials Science and Engineering 591 (2019) 012077 IOP Publishing doi:10.1088/1757-899X/591/1/012077 The first simulations have been done, considering the coordinates of the mechanism nodes in the fully extended configuration as illustrated in figure 5. We have to note that no dimensional synthesis of this mechanism has been realized till now" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0000010_afrcon.2011.6072176-Figure7-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0000010_afrcon.2011.6072176-Figure7-1.png", "caption": "Figure 7: Specific Acceleration Matching Algorithm augmented by Gaum in [8]", "texts": [ " 2) Specific Acceleration Transformation Algorithm (SATA) \u2013 This algorithm matches the specific acceleration required in the inertial frame (\u03a3Ic) to an equivalent Specific Acceleration in the wind axis frame given the constraint of zero sideslip. Thus, to obtain a lateral acceleration with respect to inertial space, the UAV will simply roll and command the required NSA. Moreover in [8], Gaum has augmented the algorithm to remove the Axial Component due to the bandwidth limit constraint previously stated. The augmented SATA then outputs an applicable CWc and an orientation as shown in Fig. 7. The SAM Architecture is normally employed with outerloop commands being generated by an inertial Position or Velocity control loop. Unfortunately, these loops command a change in total velocity, which will result in a requirement for a change in airspeed and in turn, an axial acceleration. However, due to the bandwidth limits on the ASA, the CASSAM was expected not to employ this virtual actuator in its design. As a result of this constraint, a different approach needs to be considered; the 3D Velocity Guidance Controller addresses this issue" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0001363_rast.2011.5966982-Figure3-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0001363_rast.2011.5966982-Figure3-1.png", "caption": "Figure 3. The vertex P1.", "texts": [ " Each pair of the six legs is attached to one vertex of moving platform. Bi and Pj are the centers of the joints located on the fixed and moving platforms, respectively. Geometric relations among vertices P1, P2, P3 and other parameters were presented by Nanua et al. [16]. The methodology consists of three steps. In each step, the position of one of the vertices is determined. The details of steps are presented in the following sections: 1) The Determination of Position of Vertex P1: The coordinates (p1x, p1y, p1z) of vertex P1 in Fig. 3 are determined by varying lengths of L1, L2 and \u04241 with respect to the constraints of L1, L2, and joints. Lbi and rj are the distances between Bi and Oj, and between Pj and Oj, respectively, as shown in Fig. 3. It is necessary to express Lb1, Lb2 and r1 for vertex P1 in terms of leg lengths. These expressions include the following: 2 2 2 1 2 1 (1) 2b L L LL L + \u2212= 2 1 (2)b bL L L= \u2212 2 2 1 1 1 (3)br L L= \u2212 The coordinates (x01, y01) of O1 are given by the following equations: 01 1 11 cos( ) (4)b bx x L= + \u03c0 \u2212 \u03b1 01 1 11sin( ) (5)b by y L= + \u03c0\u2212\u03b1 where (xb1, yb1) are the coordinates of B1 and \u03b11 is the angle between x axis and O1, as shown in Fig. 4. The coordinates (p1x, p1y, p1z) of vertex P1 are given by the following equations: 1 1 1 1 1 )cos sin( (6)x Op x r= \u2212 \u03c6 \u03c0 \u2212 \u03b1 1 1 1 1 1)cos cos( (7)y Op y r= + \u03c6 \u03c0 \u2212\u03b1 1 1 1sin (8)zp r= \u03c6 where \u04241 determined by considering the limitations of joints is the angle between the planes of x-y and the triangle B1P1B2" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0003874_jae-190056-Figure1-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0003874_jae-190056-Figure1-1.png", "caption": "Fig. 1. Geometry of the studied eddy-current couplings: (a) Exploded view (b) 2-Dgeometry.", "texts": [ " As the accurate analytical model can be used to analyze performance and optimize design parameters, the analytical models of electromagnetic field and torque for flux-concentration cage-type eddy-current couplings with slotted conductor rotor topology is developed in polar coordinates. Based on the accurate sub-domain method, the magnetic vector potential in each domain is obtained using the separation of variables method. Then the explicit expressions for the flux density distribution and output toque are given. The nonlinear FEA and prototype test are applied to validate the torque model. In the end, the improved particle swarm optimization is introduced to optimize the torque performance. The real geometry of the considered couplings is shown in Fig.\u00a01. As shown in Fig.\u00a01, the PM rotor is inside the conductor rotor; each magnet is circumferentially magnetized and inserted into iron cores; the conductor rotor is slotted and filled with protrusions of back iron. The geometrical parameters of prototype to be investigated are as follows: (1) The inner radius of the PM rotor is R1; (2) The outer radius of the PM rotor is R2; (3) The inner radius of the conductor rotor is R3; (4) The outer radius of the conductor rotor is R4; (5) The numbers of conductor spokes and pole-pairs are respectively Q and p", "\u00a07 shows the comparison results of error rate with 2D-FEM, analytical model with R\u2013N correction, and analytical model without R\u2013 N correction. Figure\u00a08 clearly shows the analysis results above. For the slip speed case, without R\u2013 N correction, the analytical results are close to those obtained with 2D-FEM, where the nonlinear characteristic of the materials is considered. For the high speed case, the deviations between (63) and 3D-FEM results will increase, but within the scope of the acceptance. As shown in Fig.\u00a01, the device is essentially a\u00a03D cylindrical topology. Thus the curvature effects have to be taken into consideration in analytical model. However, it is neglected in most of the literature\u00a0for simplifying model. To consider a dimensionless number \ud835\udf06, which is defined as the ratio of the radial excursion of the magnets around the mean radius to the pole pitch\u00a0[9]: \ud835\udf06 = \ud835\udc452 \u2212 \ud835\udc451 \ud835\udf0f with \ud835\udf0f = \ud835\udf0b(\ud835\udc451 + \ud835\udc452) 2\ud835\udc5d . (64) The value of \ud835\udf06 reflects the curvature situation of such devices, the larger \ud835\udf06 is, the more significantly the curvature is" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0001264_fbw.2011.5965559-Figure2-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0001264_fbw.2011.5965559-Figure2-1.png", "caption": "Fig. 2. Render from the CAD model of the quadrotor.", "texts": [ " This article presents in Section II the most relevant considerations taken into account during the design and construction of the quadrotor. In Section III, the system modelling and the design of a con troller are presented. Later, in Section IV, a Lyapunov Krasovskii based theorem is used to determine if the delay added by the wireless communication affects the system's stability. The Fly-By-Wireless implementation details are discussed in Section V and the experimental results are presented in Section VI. The quadrotor used for the purpose of the research presented in this article was designed and constructed by the authors. Figure 2 shows a render made in the Computer Aided Design (CAD) software package used in the design of the vehicle. Besides the main obvious design requirements such as the weightlthrottle ratio and stiffness, an extra effort was made to add certain features such as a high modularity and flexibility. The two latter characteristics are key in the development of a vehicle that is going to be used in several different kinds of tests that might involve the need of adding and subtracting significant hardware parts such as sensors and on-board processing units, among others", " A Microstrain 3DM-GX I Inertial Measurement Unit (lMU) is used to measure the at titude angles and a Maxbotics LV-MaxSonar EZ4 is used to determine the altitude. Quadrotor helicopters have six degrees of freedom with respect to an inertial reference frame: they can move along the three directional coordinate axis and rotate with respect to each of them. They are controlled by changing the thrust produced by each of its pro pellers and using the counter-rotative torque produced by its rotation. Figure 2 illustrates a quadrotor's motion and the forces acting upon it. The convention is that \u00a2 (roll) is defined as the rotation around the x axis, () (pitch) as the rotation around y, and 'IjJ (yaw) as the rotation around z. Propellers I and 3 rotate clockwise, while 2 and 4 rotate counter-clockwise. Note that only four linearly independent control commands can be given, and therefore it is an underactuated system. Assuming that the motors are producing a similar throttle and the overall thrust is close to the weight of the quadrotor, we can define the following control inputs to the speed command for each motor mi as follows: { ml = Ul - U2 + U4 m2 = Ul - U3 - U4 m3 = Ul + U2 + U4 m4 = Ul + U3 - U4 (1) where Ul, U2, U3 and U4 are the outputs of the altitude, roll, pitch and yaw contollers, respectively" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0000477_s00202-011-0204-8-Figure2-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0000477_s00202-011-0204-8-Figure2-1.png", "caption": "Fig. 2 Schematic representation of a cage rotor, showing also the position of the accelerometers used for the dynamical measurements described in Sect. 3", "texts": [ " A simple mathematical model is developed here. It can be analytically solved, and is able to accurately evaluate the main dynamical features of the considered structure. A preliminary version of the model was presented in [10], and used to estimate the main eigenfrequencies of undamaged rotors. The model is here extended to the case of damaged rotors, and is used to compute both the main structural eigenmodes and eigenfrequencies and the mechanical stresses at the junction between the bars and endrings, shown in detail in Fig. 2, which occur during normal service. Moreover, it is employed to design and validate a solution with the goal of increasing the torsional stiffness of the motor, thus enhancing its dynamical characteristics. 2.1 Damaged rotor A squirrel cage rotor like the one shown in Fig. 1 is here considered. It is composed of a rotor stack made of iron sheets pressed together and clamped on the shaft. Equally spaced slots are punched into the rotor stack, parallel to the rotor axis, hosting N longitudinal copper bars, with \u03b1i indicating the angular position of the i-th bar. The copper bars overhang both ends of the rotor stack by some centimetres, and their tips are connected to the copper end-rings, electrically shorting all of the bars together as shown in detail in Fig. 2. In what follows, the right and left end-rings are indicated by the letters r and l, respectively, whereas the rotor stack is indicated by the letter s. In order to develop a simple dynamical model, the rotor is schematised as shown in Fig. 3. A Cartesian frame is introduced, with the origin at the centre of the left-end cross section of the rotor, the x1 and x2 axes in the cross section plane and x3 coincident with the rotor axis. The end-rings and rotor stack are assumed to behave like rigid bodies, with mr and ml representing the masses of the two end-rings and Jr , Jl and Js representing the axial inertia moment of the two end-rings and of the rotor stack, respec- tively", " The effective bar overhang length has then been experimentally estimated by measuring the lowest eigenfrequency of single bars of a cage rotor, inserted into the rotor stack and disconnected from the end-rings. The distance from the external edge of the hammer blow imprint up to the bar tip was different on the two ends, for this particular rotor; the measured lengths were 58.3 and 55.2 mm, respectively for bars connected to the right and left end-ring. In order to measure their lowest eigenfrequency the bars have been equipped with an unidirectional accelerometer glued near the bar tip, with axis oriented along the circumferential direction, as indicated in Fig. 2 with label A1. Free vibrations have been induced by means of a hammer blow delivered on the tip of the bar overhang, parallel to the accelerometer axis as shown in Fig. 2. The accelerometer signal has been recorded by an integrated data acquisition system. The lowest eigenfrequency has been determined by performing a numerical FFT of the accelerometer signal and, consequently, the equivalent bar length has been estimated by employing the Timoshenko classical theory for shear deformable vibrating beams [11]. Results relevant to a set of bars showed that the clamping condition was attained almost 1 cm far apart from the external edge of the hammer blow imprint. Successively, two end-rings have been connected to the bar tips by the usual brazing process, thus obtaining a complete assembly, denoted as rotor UR. The free vibrations of the rotor UR have been then measured. To this end an accelerometer has been bonded with glue to the lateral surface of one end-ring, with axis oriented along the radial direction as indicated in Fig. 2 with label A2. Free vibrations have been induced by a hammer blow on the end-ring lateral surface, as indicated in Fig. 2, holding the rotor suspended with cables on its shaft. The accelerometer signal is reported in Fig. 5, whereas its FFT in Fig. 6, showing the first four resonance frequencies of the rotor. The measured values of these resonance frequencies appear in the first row of Table 2; the second row of that table contains the corresponding theoretical eigenfrequencies computed with the two decoupled set of Eqs. (8) and (9) valid for undamaged rotors, revealing an excellent agreement. The theoretical eigenfrequencies have been computed by using the rotor parameters reported in Table 1 and the bar overhang lengths lr = 40" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0001033_cca.2013.6662871-Figure3-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0001033_cca.2013.6662871-Figure3-1.png", "caption": "Fig. 3. The reaction wheel pendulum", "texts": [ " A cascaded structure of NMPC and input-output feedback linearization, has been proposed in [15]. The slower internal dynamics of the system is controlled by the predictive controller, while the fast input-output dynamics is controlled by feedback linearization. In any cascaded structure with NMPC in the outer loop, the same problems as in the direct applications of NMPC might occur - the time delay introduced by the duration to solve the optimization problem can cause stability and performance drawbacks. The reaction wheel pendulum (Fig. 3) consists of a bar, that can be rotated (by angle \u03b8) around the bearing point, located at one end. On the free end, the stator of a dcmotor is mounted. The rotor of the motor is connected to the reaction wheel via a gearbox. The control signal is the exciting voltage of the dc-motor u \u2208 [\u221224V,+24V]. The motor produces torque, which acts on both, reaction wheel and bar. The experimental setup (Fig. 3(b)) consists of a Bernecker & Rainer (B&R) reaction wheel pendulum, a B&R controller of type X20-1484-1 equipped with appropriate I/O to read the encoder values and provide the control signal as PWM modulated voltage. The equations of motion are given by a torque balance (cf. [13] for the detailed model derivation for our specific pendulum), and are defined in state space form as x\u0307 = f (x,u) = \u23a1\u23a2\u23a2\u23a2\u23a2\u23a2\u23a2\u23a2\u23a2\u23a2\u23a2\u23a2\u23a2\u23a2\u23a3 x2 \u2212 1 J [ mgl sin(x1)+Ka (u\u2212 krx3\u2212 kx2) ] 1 Ja [Ka (u\u2212 krx3\u2212 kx2)] x3 \u23a4\u23a5\u23a5\u23a5\u23a5\u23a5\u23a5\u23a5\u23a5\u23a5\u23a5\u23a5\u23a5\u23a5\u23a6 (1) for the state vector x= [\u03b8 \u03b8\u0307 \u03d5\u0307 \u03d5]T and with the torque constant Ka = k r \u03b7d\u03b7g Ra " ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0003653_j.promfg.2019.06.199-Figure3-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0003653_j.promfg.2019.06.199-Figure3-1.png", "caption": "Fig. 3. GE Aviation Challenge Bracket: a) finished part design; b) LPBF design; c) L-PBF build; d) L-PBF part after WEDM separation; e) solid support floor; f) bracket with solid support floor removed", "texts": [ " When a WEDM process is used to separate an L-PBF part from the build plate, electrical shorts often occur when the wire has to traverse through the lattice with trapped powder. This results in frequent stoppages of the process. To eliminate this problem, a preferred practice is to leave a solid metal floor along the complete bottom of the support network. This provides a solid metal pathway for the wire to traverse through. A typical thickness for this floor is 2mm. When the part is separated, approximately 1 mm of the original floor remains attached to the part. As an illustration of these attributes, consider the bracket illustrated in Fig. 3. The design was pulled from the bank of entries to the GE jet engine bracket challenge (https://grabcad.com/challenges/ge-jet-engine-bracketchallenge/results) and slightly modified to become a producible, functioning part. The part design modification and overall process design was done as collaboration between AM process engineers at CIMP-3D and the authors. The authors identified surfaces of the net shape part (Fig. 3a) that would require machining to be functional or to remove the interface to supports. The AM process engineer chose the build orientation of the part, and subsequently added solid metal supports and 0.5 mm thick machining stock to the net shape part as shown in Fig. 3b. The solid metal supports consist of a 2 mm thick wall that wraps around the bottom periphery of the part as well as a 2 mm thick solid floor. Using MAGICS software (www.materialise.com), the AM process engineer subsequently defined the lattice supports to be used to support the overhanging part surfaces, including those within the underlying support system. Once this was done, six brackets were printed from Inconel 718 powder using an EOS M280 L-PBF machine along with a system of metal testing specimens. The build took 25.5 hours to complete. The build is illustrated in Fig. 3c. As can be seen, each bracket and each specimen sits upon a support network that does not extend beyond the projected boundary of the part. Using parameters recommended by Morris Technologies [10] and the thermal processer, the build was thermally stress relieved at 1950\u00b0F +/- 25\u00b0F for 90 + 15/-0 minutes, and then allowed to thermally coast to room temperature. The parts were subsequently separated from the build plate using a WEDM process. This left a layer of solid floor (\u22481 mm thick) attached to the build plate, and yielded the bracket preform illustrated in Fig. 3d. The planar bottom of the underlying support network is illustrated in Fig. 3e. The support network with the floor stripped away is illustrated in Fig. 3f. The solid wall, lattice support, and trapped powder can be observed. The vast majority of the trapped powder is packed and partially sintered into a compacted mass. Some other relevant information pertaining to this build is that the theoretical mass of the net shape part illustrated in Fig. 3a is 132.2 g. The theoretical mass of this part with added solid supports and machining stock (see Fig. 3b) is 213.9 g. The actual masses of the six parts separated from the build plate ranged between 246 grams to 249 grams with an average weight of 247.2 grams. This net increase in mass was due to the addition of the metal lattice supports and trapped powder and the loss of approximately 1 mm thickness of the solid floor. Based on hardness data derived from the tensile samples, the hardness of the brackets after stress relief was HRC 36. With regard to added value, the L-PBF process, thermal stress relief process, and WEDM process added $339, $80, and $53 respectively to the unit value of each bracket" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0002552_978-3-319-65298-6_21-Figure6-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0002552_978-3-319-65298-6_21-Figure6-1.png", "caption": "Fig. 6. Table tennis-picking machine", "texts": [ " On the table tennis picking research, Yang Yi [16], from Beijing Institute of Technology proposes an autonomous cleaning robot with indoor global camera. The robot can find the ball, pick up the ball, pour away the ball and charge automatically with the control system. With the path planning method of \u201cMinimum Exercise Cost Short Distance Ball Priority\u201d, the robot has two modes of operation, namely single-machine operation and multi-machine collaboration. It applies to different sizes of balls, especially for tennis and table tennis. In 2008, Northeastern University invented an autonomous table tennis-picking robot (Fig. 6a) [17]. Use vision technology to locate the table tennis, to select the final path and to pick up the ball through the front holding hand. In 2014, Southeast University invented a cleaning table tennis-picking robot (Fig. 6b) [18]. It is composed of walking mechanism, picking mechanism, control system and power supply system. With machine vision and infrared sensor, it can complete the identification, tracking and picking up table tennis. Clean the table tennis ball into the channel by the picking gripper in the front, with automatic obstacle avoidance, automatic tracking and other functions. In 2014, Guangzhou University of Technology developed a suction-type table tennis-picking robot with the function of obstacle avoidance (Fig. 6c) [19]. It can locate of the table tennis by the CMOS image sensor on the real time, so that robot can quickly reach the table tennis to do the picking work. Infrared sensors can make the robot effectively to avoid obstacles and enhance the ability of adapting to the environment. A badminton-picking robot is also searched through the network (Fig. 7) [20]. It mainly includes monocular camera, the brush roller, synchronous belt lifting mechanism, the ball barrel and the ARM control system. It has the function of image recognition and visual navigation" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0000018_2012-01-0620-Figure2-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0000018_2012-01-0620-Figure2-1.png", "caption": "Figure 2. Direction of feff,i, QX(\u03b1,s), Qy(\u03b1,s) in Relation to the Velocity Direction and \u03b1", "texts": [ " (2) (3) Cs and C\u03b1 are determined from the coefficients in the BNP equations and are defined as follows: (4) (5) Once the normalized components of the tire forces are calculated, an effective deceleration rate for each tire can be calculated with the following equation [Reference 3]. (6) This equation projects the tire force components onto the vehicle velocity direction to determine the portion of these tire forces that is opposing the velocity at any particular point, a concept that is illustrated from a top down view in Figure 2. In this equation, \u00b50 is the nominal roadway coefficient of friction and feff,i is the effective deceleration rate at the ith wheel position. An effective deceleration rate should be calculated for each wheel position. Once an effective deceleration rate is obtained for each wheel position, they can be combined into an effective deceleration rate for the vehicle as a whole with the following equation. In this equation, Wtotal is the total weight of the vehicle, Wi is the static weight on each wheel position and N is the total number of wheels" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0002985_1350650117748261-Figure1-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0002985_1350650117748261-Figure1-1.png", "caption": "Figure 1. A model of freely falling steel ball.", "texts": [ " All previous studies on starved EHL problems were only concerned with the motion in the tangential direction. In this work, the authors built a mathematical model considering the effect of oil starvation to reveal the oil film formation during the impactrebound process. By incorporating the Elrod algorithm,10 the Reynolds equation used by the authors6 is rewritten to simulate the impact-rebound process with a thin layer of oil. The results are compared with those based on a fully flooded assumption. The schematic diagram of a ball freely falling onto a steel plane is shown in Figure 1. In this figure, hoil is the initial oil layer on the infinite plane, h 0 is the initial height of the falling ball. In this study, a rectangular Cartesian coordinate is employed even though the contact is axis-symmetric. This is because the available code based on a Cartesian coordinate is versatile, considering the possibility of surface roughness or elliptical contact. By incorporating the Elrod algorithm,10 the dimensional Reynolds equation of the pure squeeze motion6 under starved condition can be rewritten as @ @x h3 @p @x \u00fe @ @y h3 @p @y \u00bc 12 @ @t \u00f0 h\u00de \u00f01\u00de with a complementary condition13 as p\u00f01 \u00de \u00bc 0 \u00f02\u00de where 0< (x,y,t)4 1, is the fractional film content representing the extent as to how the gap between the two solids are filled by lubricant, \u00bc hoil/h" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0000937_ijmic.2010.035275-Figure2-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0000937_ijmic.2010.035275-Figure2-1.png", "caption": "Figure 2 Single flexible link manipulator", "texts": [ " Many methods have been proposed in literatures to reduce chattering including continuous approximation of the discontinuous control signal. A continuous approximation of sgn( )\u03c1 in (27) is ( )tanh \u03c1 \u03b5 or alternatively \u03c1 \u03c1 \u03b4+ where 0\u03b5 > and 0 1.\u03b4< < However, using this continuous function can increase the ultimate error bound, which is proportional to the value of \u03b5 and .\u03b4 The block diagram of the proposed method is shown in Figure 1. 5 Simulation examples 5.1 Flexible link manipulator (FLM) The proposed controller is applied to stabilise the FLM system depicted in Figure 2 (Wang and Vidiasagar, 1991). Points on the beam have their position fixed by the variable x, which is the distance of that point from the hub of the motor driving the beam. The elastic deformation at x is given as ( , )a x t and ( , )y x t is the net movement of that point. In other words ( , )( , ) ( ) a x ty x t t x \u03b8= + where \u03b8 is the angular rotation of the beam and 1 ( , ) ( ) ( ) n i i i a x t q t x\u03c6 = =\u2211 in which ( )i x\u03c6 are the clamped-free eigenfunctions, and n is the number of considered resonance frequencies" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0001374_eeeic.2011.5874721-Figure4-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0001374_eeeic.2011.5874721-Figure4-1.png", "caption": "Figure 4. Flux distribution and density shadow due to excitation of phase a in the case of 40% eccentric rotor to the right.", "texts": [ " To investigate the effect of dynamic eccentricity on the 6/4 switched reluctance behaviour, the motor is simulated utilizing 2-D finite element analysis by Finite Element Method Magnetics (FEMM) package1. Using finite elements method is a priori appealing for solving complex problems with a better accuracy. Iron material was used in the structure of the stator and rotor cores with the following static B-H curve shown in Appendix. Number of turns in each phase equal to 120 and the winding of phase was excited with a current magnitude of 2A. Fig.3 and fig.4 , reveals the magnetic flux distribution and density shadow in the case of the healthy machine and in the case of 40% dynamic eccentricity, respectively. From the comparing of the healthy and faulty case we observe that the flux density has increased with increasing the relative dynamic eccentricity. The main data of the studied switched reluctance machine are given in Table I. The static torque and the magnetizing flux linkages are obtained at different rotor positions from 0 to 90\u00b0 taking rotational steps of 5\u00b0 where the rotor moves from unaligned to fully aligned position" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0001460_j.sna.2010.05.007-Figure5-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0001460_j.sna.2010.05.007-Figure5-1.png", "caption": "Fig. 5. Actual views of Micro USM I/II.", "texts": [ " This means, the wiring problem of ultrasonic motors could e reduced. In addition, this method does not require any conroller at each motor like smart actuators or intelligent actuators 11]. However, we must be aware of the limitation of the proposed ethod, i.e., the driving frequency of each motor should be differnt. The required difference of driving frequencies is related to the echanical Q factor of each motor. . Motor characteristics In this study, we use two kinds of ultrasonic motors to conrm the effectiveness of the proposed driving method. Fig. 5 and able 1 show the traveling wave ultrasonic motors, Micro USM I nd Micro USM II (Canon Inc., Japan), used in this study. As menioned in Section 3, the operational parameters for ith ultrasonic able 1 pecification of Micro USM I/II. Micro USM I Micro USM II Maximum torque (Nm) 6.9 \u00d7 10\u22123 7.8 \u00d7 10\u22123 Maximum rotational speed (rpm) 950 700 Size (mm2) \u00d8 11 \u00d7 25 \u00d8 10 \u00d7 10 Mass (g) 11.3 5 Driving voltage (Vp-p) 12\u201320 12\u201320 Natural frequency (kHz) 36\u201337 56\u201357 t) An sin(2 fnt + ) (2) motor are the amplitude Ai and frequency fi of ith component of the superimposed signal" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0000104_icra.2013.6630660-Figure4-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0000104_icra.2013.6630660-Figure4-1.png", "caption": "Fig. 4. The resource constrained delta-wing UAV used in the experimental demonstration of the DDDAS technique. The 0.8 m wingspan UAV is equipped with a GPS sensor for position, a roll rate sensor, XBee wireless communication radio, and an autopilot to control the craft during autonomous mode.", "texts": [ " As the puff diffuses, sensors in scheme 1 do not yield information about the puff\u2019s outer regions thus resulting in a poorer overall estimation of the puff compared to sensors under the guidance of schemes 2 and 3. In this section we describe the experimental platform used to demonstrate the feasibility of the dynamic data-driven feedback control system. A complete description of the system can be found in [7]. Following the description of the system, the results from the field deployment are presented. The mobile sensor used in the experimental validation is shown in Figure 4. The aircraft has a wingspan of 0.8 m and weighs less than 0.5 kg. The UAV is equipped with the CUPIC custom autopilot [26]. The autopilot system contains a GPS sensor, a roll rate gyro, and a wireless radio to transmit and receive information. The autopilot controls the horizontal position of the aircraft by varying the roll angles while the pilot controls the altitude through pitch and throttle. The autopilot system also includes a complementary groundstation which is comprised of a laptop and an XBee wireless radio" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0000728_robio.2011.6181627-Figure2-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0000728_robio.2011.6181627-Figure2-1.png", "caption": "Fig. 2. Model: (a) Frontal-plane ankle strategy, (b) Lift-leg strategy.", "texts": [ " We call this strategy the Cross-leg step strategy (Fig. 1 (c)). Sometimes, with a relatively large disturbance, humans reacted by directly taking a step either aside or across to attain a balance posture resulting in a Side-step strategy (Fig. 1 (d)) or Cross-leg step strategy (Fig. 1 (c)), without involving the Lift-leg strategy. III. IMPLEMENTATION WITH A SMALL HUMANOID ROBOT The video clip shows how we implemented the ankle and the lift-leg strategies with a small humanoid robot HOAP-2. We used the two planar models shown in Fig. 2. Note that the robot does not have a joint corresponding to the waist joint of the human model. Proceedings of the 2011 IEEE International Conference on Robotics and Biomimetics December 7-11, 2011, Phuket, Thailand An inverted pendulum model was used for the ankle strategy. This was possible because the two legs move in parallel. The two phases of motion, disturbance response and balance recovery, can be realized with the help of a virtual spring and damper attached between the vertical and the pendulum (see Fig. 2(a)). We implemented reaction patterns in response to two types of disturbances: impact force and continuous force. In the case of an impact force acting on one of the robot\u2019s shoulders, the magnitude of the impact is measured by the internal acceleration sensor of the robot located near the total CoM. In the case of a continuous disturbance force acting on one of the shoulders, the ankle joint torque is derived from the foot moment balance equation using the ZMP position. Experiments: The first half of the video shows that the robot reacts appropriately to the two types of disturbances and restores the balance posture smoothly. Note also that the robot can react to a disturbance applied during a balance recovery phase. The lift-leg strategy is realized with the help of the Reaction Null Space method [7] which was developed to tackle reaction control in the field of space robotics. Next, referring to the model in Fig. 2 (b), We note that only the support leg\u2019s ankle and the support leg\u2019s hip joints will be considered. The swing leg\u2019s hip and the swing leg\u2019s ankle joints are fixed. The motion disturbance response and balance recovery, can be realized with the help of a virtual spring and damper attached between the vertical and the pendulum. The support leg\u2019s hip joint is used for disturbance response and balance recovery. The support leg\u2019s ankle controls center of mass based on RNS method. C. Integration of Frontal-plane ankle strategy and Lift-leg Strategy We need to ensure smooth transitions between the Frontalplane ankle strategy and Lift-leg strategy during balance control in response to a continuous disturbance" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0001455_1350650112439260-Figure2-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0001455_1350650112439260-Figure2-1.png", "caption": "Figure 2. Pressure pads mounted on the sides of the pinion (motion reveals the oil flow).", "texts": [ " The average roughness (Ra) of the tooth flanks, measured in the radial direction before tests, was in the range 0.21\u20130.28mm (Lt\u00bc 1.5mm, cut off\u00bc 0.25mm). The FZG test gearbox was prepared to test spur gears, thus no axial loads are generated. The low loss gears have a helix angle that will promote an axial load for which the cylindrical roller bearings are not prepared. To overcome this issue, a set of pressure pads was used to allow testing helical gears and absorb the corresponding axial loads. The pressure pads are two side discs mounted on the pinion\u2019s shaft, as shown in Figure 2. Their purpose is to transfer axial loads through the lateral faces of the gear pinion and wheel without the need of tapered bearings, in this manner the cylindrical roller bearings of the gearbox only support the radial loads from the helical gears. The pressure pads have influence in the power losses because they promote lubricant drag, as shown in Figure 2, and modify the way how the contact is fed. The experimental results were divided into two: the first showing the churning power loss tests and the second the load power loss tests. The first part will present the influence of the speed, geometry, oil type and oil level in churning losses, while the second the influence of the low loss geometries and oil type on power losses for different loads and speeds. The results of the power loss tests are expressed by the stabilization temperature that is determined as the oil bath temperature subtracted by the room temperature (Ts\u00bcToil Troom)", "8,15,16 The pressure pads are responsible for churning losses so its lateral area was considered and added to the lateral area of the pinion, thus they were also considered for the churning losses Pspl \u00bc 30 n 1 2 n 30 2 Ai d 2 3 \" # 2 h d 0:45 Voil d 0:1 Fr 0:6 Re 0:21 \" # \u00f04\u00de The gear churning losses are also dependent on the sense of rotation.16 When the sense of rotation drags the oil to the beginning of the tooth contact, the churning losses increase. Changenet and Pasquier16 propose an increase of churning losses, P, represented in equation (5) to take into account this phenomena P\u00bc 1 2 n3 Ai d 2 3 :17:7 Fr0:68 u 1 u8 1 h d=2 \u00f05\u00de When observing Figure 2, it is clear that the pressure pads promote the drag of a high amount of oil to the entrance of the contact and thus promote a similar effect as the sense of rotation that drags the oil immediately to the contact. Hence, this effect was taken into account for the energetic balance. The pressure pads used to allow the test of helical gears promote another type of churning losses, the losses generated by the lubricant flow between the pad and the wheel. It was assumed that a Couette flow takes place and the losses for that flow were calculated according to equation (6), which already took into account the existence of two pads Pcou \u00bc 2 u y i \u00f06\u00de In equation (6), vi represents the relative speed between the pressure pad and the wheel, for each zone of the domain considered" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0003970_s11370-019-00294-7-Figure5-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0003970_s11370-019-00294-7-Figure5-1.png", "caption": "Fig. 5 Serial-to-RRR parallel conversion a serial manipulator, b derived parallel manipulator RRR", "texts": [ "\u00a04b. In summary, when the instantaneous twist T\u0302 of parallel platform is specified, the actuator speeds v1, v2, v3 can be computed as (30). The geometry of parallel manipulator converted from serial manipulator can be diverse. Here, RRR-type parallel manipulator is considered. The linear velocities v1, v2, v3 along the direction of each second link can be obtained from the angular velocities ( 1)1, ( 1)2, and ( 1)3 of the first joint of each chain of the 3-DOF RRR parallel manipulator shown in Fig.\u00a05. where (\u22c5)i denotes a vector in the ith chain of RRR parallel manipulator and the mutual moment term (e.g., s\u0302T 23 (S\u03021)1 ) (32)jT = \u23a1\u23a2\u23a2\u23a2\u23a3 s\u0302T 23 s\u0302T 31 s\u0302T 12 \u23a4\u23a5\u23a5\u23a5\u23a6 = \u23a1\u23a2\u23a2\u23a3 c1 c2 c3 s1 s2 s3 p1 p2 p3 \u23a4 \u23a5\u23a5\u23a6 T . (33) v1 = (\ud835\udf141)1(s\u0302 T 23 (S\u03021)1) v2 = (\ud835\udf141)2(s\u0302 T 31 (S\u03021)2) v3 = (\ud835\udf141)3(s\u0302 T 12 (S\u03021)3), denotes the shortest distance between two vectors ( \u0302s 23 and (S\u03021)1 ). Geometrically, they are the shortest distance between two vectors. Putting the values of v1, v2, v3 into the equation will result in the following derived parallel-type manipulator. It is noted that the three rows of (34) are lines expressed in ray coordinates as shown in Fig.\u00a05. Therefore, it is remarked that the Jacobian of serial manipulator composed of three lines expressed by axis coordinate is converted into ray coordinate. It is also remarked that the mutual moment term (e.g., s\u0302T 23 (S\u03021)1 ) in the denominator of (34) provides the optimum choice to design the first link of each chain. That is because the freedom to locate (S\u03021)i is diverse as shown in Fig.\u00a06, while the line denoted as s\u0302T 23 is fixed along the link. It is also remarked that the static relation, the dual relation of (34), is denoted as (34)" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0003660_iemdc.2019.8785146-Figure11-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0003660_iemdc.2019.8785146-Figure11-1.png", "caption": "Fig. 11 Analysis models of multi-three-phase IPMSMs", "texts": [ " In addition, quantitatively, there is a difference between the analysis value and the theoretical calculation value and, in particular, a large difference is observed at 6p36s. This seems to be because the theoretical model assumes all the harmonics have a gap magnetic flux density described as equation (15) and cannot consider variation of the magnetic saturation of the stator core with the change in the magnetic gap length and permeance due to the slot opening. We next performed the same examination as in Section III. A with the IPMSM structure and verified the validity of our theory. The analytical models of the IPMSM are shown in Fig. 11. They are structured with flat magnets embedded in a perfectcircle-shaped rotor. Their other characteristics are the same as the SPMSM models and Table 1. Waveforms of the no-load stator voltage as a result of FEA are shown in Fig. 12, and the results of Fourier transformation on those waveforms are shown in Table 3 and Fig. 13. In Table 3, the maximum and minimum voltage are chosen from the no-load line-line voltage in all groups. Since the magnetic pole of the rotor and teeth of the stator are roundly and axially symmetrical in these analytical models, even-order harmonics are not generated, so only the odd-order harmonics are shown in Table 3 and Fig" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0002302_b978-0-12-803581-8.10351-0-Figure49-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0002302_b978-0-12-803581-8.10351-0-Figure49-1.png", "caption": "Fig. 49 The four primary monocoque shells, left, right, top and bottom, plus nose, engine cowl, and sidepod inlets. Reproduced from Radford, D.W., Fuqua, P.C., Weidner, L.R., 2004. Tooling development for a multi-shell monocoque chassis design. In: 36th International SAMPE Technical Conference, San Diego, CA, November 15\u201318, 2004, pp.1063\u20131077.", "texts": [ " The carbon fiber prepreg that was chosen was the Toray T700S-12K-50C plain weave and the corresponding Toray T700G-12K-31E unidirectional material. The corresponding composite prepreg cure parameters were critical in the selection of mold materials and ultimately, in the tooling materials. The decision was to follow common industrial practice, using carbon fiber/epoxy for the mold material to help match the thermal expansion of the planned prepreg monocoque shells. The primary monocoque shells are shown in the exploded view of Fig. 49. Outer surfaces of each of the shells were to be the as-molded finish. It was determined that a soft tooling material, in this case a urethane foam, could be used in conjunction with a liquid epoxy for the molds which would cure at room temperature, but was capable of a freestanding postcure to a degree that would enable a 1501C cure temperature. In addition to the geometry dictated by the structural demands of the chassis and of the individual monocoque shells, the CAD design had to incorporate draft angles, in this case of approximately 3 degree", " From a tooling design perspective, this had limited effects on the machining; however, these flat, parallel \u2013 no draft \u2013 surfaces created additional complexity during mold manufacture, and will be further discussed during the discussion of mold manufacture. The previous sections have described the basic concept of developing a multi-shell composite monocoque chassis to investigate potential improvements in manufacturability. Some level of design detail has also been included in previous sections to indicate the design complexity and the additional composite components that are required to actually meet the structural requirements, beyond the primary composite shells of the multi-shell configuration shown in Fig. 49. This section will discuss the soft tooling developed for the primary monocoque shells, as well as some of the hard tooling designed for the manufacture of bulkheads. Soft tooling, as discussed in this section, refers to positive master patterns machined from urethane tooling foam and coated with a thin, sprayed-on epoxy surfacing coat. These tools were designed from the CAD models of the original monocoque shells and ultimately machined from large blocks of urethane tooling foam on a five-axis router" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0002749_physreve.96.042603-Figure1-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0002749_physreve.96.042603-Figure1-1.png", "caption": "FIG. 1. Flagellated bacterium model and channel geometry. The flagellated microswimmer is modeled as a system of two rigid bodies: the head (a prolate ellipsoid) and the tail (an helical tubular structure). The head and tail axes lie along the same direction, and the tail can rotate with respect to the head. The system has 7 degrees of freedom, namely the position xj of the junction between the head and the tail, the orientation of the head with respect to the fixed reference frame, and the rotation angle \u03c6 of the tail with respect to the head. In the plane channel geometry, the only relevant degrees of freedom are the distance between the junction and the wall zj and the angle between e\u03021 and x\u0302.", "texts": [ " In an intermediate range of tumbling frequencies, D must have at least one maximum that is found to occur approximatively when \u03b1 is comparable to the frequency of the stable circular orbit. In contrast, in strong confinement, the swimmer does not stabilize on circular orbits during the run phase. Hence, for each value of \u03b1 the trajectory is a sequence of straight lines and the same arguments holding for the free-space case applies, namely D \u223c 1/\u03b1. The microswimmer is composed of two rigid bodies, the head and the tail; see Fig. 1. The head is a prolate ellipsoid with a1 and a2 the longitudinal and transverse semiaxis, respectively. The swimmer head position is identified by the junction between head and tail xj, which is also taken as the origin of the body reference frame {e1,e2,e3}, where the unit vector e1 lays along the longer ellipsoid axis. The body translational velocity is denoted by U, while H is the angular velocity of the head. The tail is modeled as a helix of wavelength \u03bb, amplitude A, and axis et = \u2212e1, L being the axial length of the tail which has a cross section of radius aT ", " At each time step, given the torque \u03c4M , the system is solved and the swimmer velocities can be employed to update the microswimmer configuration. Following Shum et al. [21,40], the system symmetries can be 042603-2 exploited to set up a more efficient approach. In brief, since the system is invariant under rotation around the z axis and translation in x and y directions, it is convenient to precalculate the velocities when the bacteria sits in specific positions in the reduced space ( ,zj ) with the (pitch) angle between the e1 body axis and the bottom wall and zj the z coordinate of the junction position xj; see Fig. 1. The tail rotational velocity \u03c6\u0307(t) = T is much larger than the other velocities involved in the problem. This calls for averaging on the tail rotation \u03c6; thus, the swimmer translational velocity U and its angular velocity H depend only on and zj . The bacterium trajectory during the run phase is then calculated integrating the rigid-body kinematics equations of the swimmer head, where U and H are interpolated from the precalculated values on a discrete set of ,zj pairs. The tumbling phase is achieved by randomly rotating e1 by an angle \u03c8 to a new orientation e\u2032 1 which lies on a cone of axis e1 and semiamplitude \u03c8 [41]" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0000529_eptc.2011.6184475-Figure1-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0000529_eptc.2011.6184475-Figure1-1.png", "caption": "Figure 1 (a) Schematic diagram of a flexographic printing roller assembly; (b) a typical printing stage on a roll-to-roll flexographic printer", "texts": [ " [4] Flexographic printing is one of the potential roll-to-roll patterning technologies that can be adopted to perform additive deposition of functional materials. Flexography (often abbreviated to flexo) is a type of graphic printing process which utilizes a flexible elastomeric plate. It is an improved version of letterpress or rubber stamp that can be used for printing on almost any type of substrate including plastic, metallic films and paper. It is also widely used for printing on the non-porous substrates required for various types of food packaging. A typical flexo printing roller assembly is illustrated as in Figure 1. The flexo setup consists of an inking tray or chamber (where ink is stored), anilox roller (which captures accurate ink volume in the arrays of microcavities/cells), plate roller (which receives ink from anilox roller onto raised surfaces) and impression roller (where the ink transfer occurred when the plate contacts the substrate). Plate Roller Impression Roller Anilox Inking tray Substrate (a) The flexible elastomeric plate making process is similar to photolithography process in semiconductor and electronics industry" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0001059_1.4001772-Figure7-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0001059_1.4001772-Figure7-1.png", "caption": "Fig. 7 kth order kinematics of a general rigid body", "texts": [ " In case of velocity, there is no component along the adius vector. Thus the case needs to be handled separately. In eneral, for kth order motion Mk is the component that is directed long the radius vector. When Mk is zero, k=90 deg. This can be erified for second order kinematics when the angular velocity f the body is zero while the angular acceleration is nonzero, . In hat case, 2=90 deg and the Eq. 47 results in a division by zero 0/0 . This is a special scenario that yields important information f the instantaneous motion of the body. Figure 7 illustrates the properties graphically, which shows a eneral kth order motion of a rigid body. E1, E2, and E3 are three istinct points on the body. At a particular point in time, the IC for th order is located at Ik. For discussion purpose, let the three oints be such that the orientations of k 1 and k 2 are instantaeously equal k 1 = k 2 . This is true when E1, E2, and Ik are ollinear. Also let the magnitudes of k 1 and k 3 be equal. This s true when the two points are on a circle with the IC as the rigin. The parameters that have constant value for all points in he body are colored red, where as the parameters that are depenent on the location of the point are colored green in Fig. 7. We an see that m1 and m2 are parallel and m1 and m3 have the same agnitude. In other words, all the points on a radial line will have he same orientation for respective motion properties and all the oints at same radial distance will have equal magnitudes. This hows the geometric nature of the formulation. We have presented here an algebraic formulation to represent he kth order motion of a rigid body in terms of the motion proprty of a general point E in the body by using the corresponding C as the reference" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0003036_imece2017-70301-Figure6-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0003036_imece2017-70301-Figure6-1.png", "caption": "Figure 6. Structure of the measured spindle", "texts": [ " The specifications of measurement system are shown in Tab.1. Table 1. Specifications of the measurement system Parameter value Loading force Fmax 200 N Coil current Imax 10 A Swept range 0~1000 Hz Diameter of dummy tool 40 mm Air gap 1 mm As shown in Fig.5, an experimental spindle test rig is established, and the developed measurement system is used to measure the stiffness and natural frequency of the spindle. The experimental spindle is driven by a motorized spindle, and the structure is shown in Fig.6. The springs in the measured spindle are used to apply constant pressure preload to the bearing. In Tab.2, the specifications of the measured spindle are listed. As the heat generation in the bearings is the major heat source of the measured spindle, several temperature sensors are installed in the spindle to monitor the temperature of the outer ring of bearings when spindle rotates. spindle, the bias current 0i is set as 5 A, and the control current shift from 0 A, 1 A, 2 A, 3 A to 4 A. Next, the control current is reduced to 0A one by one again" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0002082_978-3-319-49137-0_1-Figure2-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0002082_978-3-319-49137-0_1-Figure2-1.png", "caption": "Fig. 2 Left a Gold-coated Teflon working electrode insert, areas 1 and 2 as in C. b Completed cell diagram showing (a) Pt auxiliary electrode, (b) SCE reference electrode. (c) Radiation path, (d) Teflon bottom spacer, (e) quartz cuvette and (f) gold-coated Teflon working electrode. c Top view of the cell showing radiation paths in absorption and fluorescence experiments, areas 1 and 2 as in A (from Ref. [48]). Right Design of a two-body spectrofluoroelectrochemical cell with right angle detection (from Ref. [51])", "texts": [ " Despite results in agreement with the potential of F-SEC in terms of sensitivity, the fluorescence detection is mainly limited by the light scattering from the front face of the cell leading to a low signal-to-noise ratio and to poor reproducibility. To overcome limitations encountered with OTTLE cells, e.g., short pathway inducing low signal intensities, \u201clong optical path electrochemical\u201d (LOPE) cells based on gold resinate film electrode [48] or reticulated vitreous carbon electrode (RVC) [49\u201351] were designed (Fig. 2). Their configuration permitted detection of the emitted light at 90\u00b0 and avoided scattering effect from the faces of the cell. Unfortunately, the equilibration time is longer than the one required for OTTLE cells and shows the pseudo-thin-layer nature of the long optical path electrochemical cells. However, bunch of experiments can be easily performed with this type of cells, particularly for stable electroactive species. Following on from cells mainly limited to stable species, OTTLE cells have been redesigned by Compton et al" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0000208_icosp.2010.5656122-Figure2-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0000208_icosp.2010.5656122-Figure2-1.png", "caption": "Figure 2. Mount position of tri-axles accelerometers.", "texts": [ " Transfer function from input jF to output iX should be estimated as well. ___________________________________ 978-1-4244-5900-1/10/$26.00 \u00a92010 IEEE The purpose of this measurement was to acquire vehicle interior noise ( struP ) and acceleration responses ( iX ). The work was performed in a semi anechoic room on 4 wheel chassis dyno with smooth surface. Sound pressure at driver position was chosen as target interior noise and four tri-axles accelerometers were mounted at four spindle noses respectively representing for 12 transfer paths (figure 2). A four cylinder front wheel driven compact car with automatic gearbox has been utilized in this work. In order to exclude other noise sources the car was driven by chassis dyno with engine idle (due to automatic gearbox) in neutral gear during measurement. Target interior noise at constant speed from 50km/h to 120km/h with 10km/h interval was investigated (figure 3) and the result showed a positive relationship between interior noise and speed. Subjective evaluation of interior noise on road was carried out during measurement" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0000022_j.jappmathmech.2012.05.003-Figure4-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0000022_j.jappmathmech.2012.05.003-Figure4-1.png", "caption": "Fig. 4.", "texts": [ "9) It is seen that > 0; therefore, depending on the values of and , cases a, c and d (Fig. 2) are possible in this system. The regular case corresponds to the inequality (4.10) A similar analysis can be performed for the case of unilateral constraint (2.12) with kinetic friction (4.1). In this case, N \u2265 0 for q1 = 0, nd N = 0 for q1 > 0; therefore, when u > 0, the terms that are first-order in u on the right-hand side of system (4.4) vanish (accordingly, in xpression (4.5) the second-order terms vanish). These terms correspond to the constraint reactions (4.11) Figure 4 presents graphs of function (4.11) (for convenience an additive constant was added) for the following cases: a) > \u2013 , < 0; ) > \u2013 , > 0; c) < \u2013 , < 0; ) < \u2013 , > 0. L b d E T r a o A R 1 1 1 1 A.P. Ivanov / Journal of Applied Mathematics and Mechanics 76 (2012) 142\u2013 153 153 emma 5. a) If > \u2013 and < 0, in a system with kinetic friction and a unilateral constraint, the normal constraint reaction is uniquely defined: (4.12) ) If > \u2013 and > 0, the contact is broken. c) If < \u2013 and < 0, the reaction increases without limit (shock occurs)" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0002668_0892705717722189-Figure3-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0002668_0892705717722189-Figure3-1.png", "caption": "Figure 3. Testing setup for the flatwise tension test after DIN53292.", "texts": [ " It is possible to pre-heat the core in a second air-circulating oven before the installation into the female mould. Once transfer plate and skin are integrated into the picture frame (installed at the lower press platen), the mould is closed. To avoid core collapsing, the pressure is kept low at p\u00bc 0.2 MPa and the process is controlled by the penetration depth hpenetration. Closing of the mould is stopped when female and male mould touch. Mechanical testing is performed to evaluate the mechanical properties of the skin to core interface. The tension test in the flatwise plane after DIN532928 (see Figure 3) is selected due to the small specimen dimensions and the nominal specimen preparation. The sandwich specimens are adhesively bonded to aluminium profiles by means of a pasty thermoset adhesive DP490 (3M, Manchester, UK), which can be integrated in a universal testing machine Instron 5566 (Instron, Norwood, Massachusetts, USA), equipped with a 10-kN load cell. Testing is performed using a constant speed of 0.5 mm/ min at room temperature and a relative humidity of 50%. Failure modes of the sandwich specimens are assessed based on the evaluation according to DIN EN ISO 10365,9 which differentiates between cohesive failure at the centre of the core (Coh), cohesive failure in the boundary layer of the core (B" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0000408_rm2012v067n02abeh004786-Figure6-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0000408_rm2012v067n02abeh004786-Figure6-1.png", "caption": "Figure 6. \u2018Cut\u2019 singularity. Figure 7. \u2018Cup\u2019 singularity.", "texts": [], "surrounding_texts": [ "Here we mainly follow the papers [25]\u2013[28] in describing generic singularities of the boundary of the set of local transitivity of control systems on a plane and in three-dimensional space. The basic types of control systems are examined, namely, the cases where the indicatrices I(x) \u2282 TxM are: 260 A.A. Davydov and V.M. Zakalyukin \u2022 smooth curves or surfaces which are embedded (or immersed) in the linear spaces TxM of tangent vectors at points x \u2208 M ; \u2022 or smooth curves or surfaces with boundaries and corners embedded in the spaces TxM ; \u2022 or images of smooth maps of general form of a smooth closed manifold U with dimension at least two; \u2022 or, finally, images of smooth maps of a manifold U with boundary and corners. We recall that a corner on a manifold is a subset of it which is diffeomorphic to a quadrant of the arithmetic space of the corresponding dimension. The fact that the manifold U of controls has boundaries or corners can be interpreted as a simulation of constraints imposed on the control in the form of inequalities. This kind of control system is widely used in applications. Let us consider a point x of the manifold Mn as a parameter and fix some trivialization of the tangent bundle of M . We obtain a family of submanifolds I(x) embedded in Rn. Accordingly, by the boundary \u03a3 of the set of local controllability we shall mean the set of parameters x for which 0 \u2208 Rn belongs to the boundary of the convex hull of the corresponding submanifold I(x). We recall that the convex hull of a (closed) subset of an affine space is the intersection of all closed half-spaces containing this set. Note also that the boundaries of the convex hulls of smooth submanifolds in general position can have singularities. Their classification, which is closely related to the singularities of the boundaries of the zone of local transitivity, is also discussed below. 3.1. Formulation of results. We start with the simplest case of a control system on a two-dimensional manifold M . In the tangent plane at each point of the manifold let there be specified a set I(x) of admissible velocities which is a smooth closed curve. We recall [5] that the boundary of the convex hull of a generic smooth curve on the plane at any point is either the germ of a smooth curve (a convex curve or a straight line) or the germ of a curve which can be transformed via some diffeomorphism of the plane into the germ of the graph of the function y = f(x) at the origin, where f(x) = { 0 for x 6 0, x2 for x > 0. If the curve has endpoints, then the list mentioned above should also include the singularity that corresponds to the graph of the function g(x) = |x|. For a generic two-parameter family of smooth curves, the boundary \u03a3 of the set of local transitivity can have only one more singularity distinct from the singularities of the convex hull of an individual curve in general position [25], namely, the indicated singularity of the graph of the function y = |x| at the origin. It turns out that this list coincides exactly with the list of generic local singularities (up to diffeomorphisms of the plane) of the boundary of the zone of local transitivity for control systems on the plane whose indicatrices are smooth curves with endpoints. Moreover, it is the same as the list of singularities of convex hulls and boundaries of the zones of local transitivity in the case where the indicatrices are the images Controllability of non-linear systems 261 of a manifold U of arbitrary dimension under smooth maps which form families in general position. This being the case, it can be assumed that the manifold U has boundaries and corners. The techniques of the proofs of these (and some other) facts are discussed below. Let us turn to the main case of a three-dimensional manifold M . First we shall assume that the indicatrix is a C\u221e-smooth closed space curve. The list of generic local singularities of the convex hull of a curve in R3 up to diffeomorphisms of the ambient space was obtained in [29] and [6]. It includes the first six of the normal forms presented below in Theorem 1. In order to describe the singularities of the boundaries of the convex hulls of space curves with endpoints we introduce the following notation. A closed convex simple piecewise smooth curve \u03b3 that lies in the plane z = 1 of the space R3 = {(x, y, z)} is called a simple pivot curve if it consists of alternating segments of straight lines and strongly convex arcs and is C1-smooth at their common points. A curve \u03b3 is called a pivot curve with k corners if it contains the sides of k angles less than \u03c0 connected by convex pieces consisting of straight segments and arcs and is C1-smooth at all common endpoints except the vertices of the angles. Some sides may be common for adjacent angles. If we replace a straight-line side of an angle by a convex curve arc in this definition, then we obtain the definition of a pivot curve with k corners and a curved side. The conic surface formed by the segments that join the origin with the points of a pivot curve which is either simple (k = 0) or has k = 1, 2, . . . corners (possibly, with curved sides) will be called a k-cone. We note that the germs of these cones at the origin have functional invariants with respect to the action of diffeomorphisms. Indeed, the tangent vectors at the apex form the tangent cone and sweep out a convex curve on the sphere of directions. The tangent cone is subject only to linear transformations under the action of diffeomorphisms. Theorem 1 [27]. The germ at an arbitrary point of the convex hull boundary of a generic connected space curve with endpoints can be reduced by an appropriate diffeomorphism of R3 to the germ at the origin of the graph of one of the following functions z = fi(x, y), where i = 1, . . . , 7 (Figs. 2\u20135): 262 A.A. Davydov and V.M. Zakalyukin 1) f1 = 0 (the germ of a smooth surface); 2) f2(x, y) = |x| (edge); 3) f3(x, y) = { 0 for x 6 0, x2 for x > 0 (adjacency); 4) f4(x, y) = x2 for y 6 x, x > 0, y2 for y > 0, y > x, 0 for y 6 0, x 6 0 (bow); 5) f5(x, y) = 0 for y 6 0, x 6 0, x2 for y 6 \u2212x, x > 0, y2 for y > 0, y 6 \u2212x, 1 2 (x2 + y2)\u2212 y \u2212 x for x + y > 0 (stern); 6) f6(x, y) = minz\u2208R{z4 + xz2 + yz} (truncated swallowtail); 7) f7(x, y) = y2 + x for x > 0, y2 for y 6 0, x 6 0, (1\u2212 x)y2 for y > 0, x 6 0 (bend); or to the germ at the origin of a k-cone with k = 0, 1, 2. Remark 2. Generic singularities of the convex hull of a closed space curve (without endpoints) belong to the first six classes mentioned in this theorem. Remark 3. The smooth surface 1) may be strictly convex, ruled (Fig. 2, point B), or flat (Fig. 2, point A). The \u2018edge\u2019 2) arises at a generic point of the initial curve itself (Fig. 2, point C). The singularity 3) appears, in particular, at the points of adjacency of a ruled surface and a flat one (Fig. 2, point D). The germs 4) and 5) correspond to vertices E and F of the flat triangles which inevitably appear on the boundary of the convex hull of the curve. The cone apex coincides with the endpoint (Fig. 5, point K). The truncated swallowtail appears at a point of the curve such that the tangent line at this point also intersects the curve at some other point (Fig. 3, point J). Controllability of non-linear systems 263 The list of generic local singularities of the boundary of the transitivity zone for a control system on a three-dimensional manifold whose indicatrix is a closed space curve is given by the following theorem. Theorem 2. For a generic family of smooth curves rm : S1 \u2192 R3 depending on a three-dimensional parameter m = (x, y, z) \u2208 R3 the list of all local singularities on the boundary \u03a3 of the transitivity zone (considered up to diffeomorphisms of R3) consists of the germs 1)\u20137) mentioned in Theorem 1 and the germs 9), 12), 14) described in Theorem 3 below. For a space curve with endpoints the list of generic singularities of the transitivity zone is as follows. Theorem 3. For a generic three-parameter family of connected space curves with endpoints, the germ at any point of the boundary \u03a3 of the transitivity zone can be reduced by an appropriate diffeomorphism either to one of the germs 1)\u20137) of the boundaries of the convex hulls of generic curves in R3 listed in Theorem 1, or to the germ at the origin of the graph of one of the following functions : 8) f8(x, y) = 0 for x 6 0, y 6 0, y2 for y2 > x, y > 0, x for y 6 \u221a x , x > 0 (cut); 9) f9(x, y) = 0 for y 6 0, x + \u03b1y 6 0, (x + \u03b1y)2 1 + \u03b12 for y 6 \u03b1x, x + \u03b1y > 0, \u03b1 \u0338= 0, y2 for y > 0, x 6 0, x2 + y2 for x > 0, y > \u03b1x (adjacency of four surfaces, a cup); 10) f10(x, y) = \u2212y for y > 0, 0 for y 6 0, x > y, y(x\u2212 y)2 for y 6 0, x 6 y (helmet); 11) f11(x, y) = { |x| for y > 0, |x|+ y2 for y < 0 (butterfly); 264 A.A. Davydov and V.M. Zakalyukin or to the germ at the origin of the union of three surfaces with boundaries determined by the conditions 12) z = 0, y 6 0, y = x2, z 6 \u22124x2, z = \u2212t2, y = 1 4 z + tx, 0 6 t 2x 6 1 (adjacency with Whitney umbrella); or to the germ at the origin of the union of two surfaces with common boundary given by the conditions (book) 13) z = 0 for y, x > 0 or for x 6 0, y > 1 4 x2, z = 2t3 + xt2 for 3t2 + 2tx + y = 0, t > max { 0,\u2212x 2 } ; or to the germ at the apex of the lateral surface of a pyramid with n ridges for n = 3, 4, 5, that is, to the germ of the boundary of the domain given by the inequalities 14) x, y, z > 0 for n = 3, x, y, z > 0, z \u2212 x + y > 0 for n = 4, x, y, z > 0, z \u2212 x + y > 0, z \u2212 \u03b1x\u2212 \u03b2y > 0 for n = 5 (for n = 5 the normal form has two scalar invariants \u03b1 and \u03b2); or, finally, to the germ of a k-cone with k = 0, 1, 2. Controllability of non-linear systems 265 Remark 4. The boundaries a and b of the surfaces I and II of the germ 12) are given by the equations z = 0, y = 0 and y = x2, z = \u22124x2, respectively. These curves are tangent to each other at the origin, which is their unique common point. The surface III is a piece of a Whitney umbrella bounded by these curves and tangent to the surfaces I and II along a and b (Fig. 9). The boundary H(Ix) of the convex hull of a generic C\u221e-smooth surface Ix may have singularities which are only C1-smooth. Generic local singularities of H(Ix) [5] are either germs of a smooth surface or germs of a C1-smooth surface of \u2018adjacency\u2019, \u2018cup\u2019, or \u2018truncated swallowtail\u2019 type. For surfaces with boundary and corners, the following assertions hold. Theorem 4. The list of generic local singularities of the convex hull boundary of a smooth compact surface in R3 with a smooth boundary consists, up to diffeomorphisms, of the singularities 1)\u20137), 9), and 11). Theorem 5. Generic local singularities of convex hull boundaries for smooth compact surfaces in R3 with boundaries and corners are either the singularities 1)\u20137), 9), 10), or k-cones with arbitrary k. Theorem 6. The list of local singularities (up to diffeomorphisms of the parameter space) of the boundary of the set of local transitivity for a generic three-parameter family of smooth compact surfaces in R3 consists of the germs 1)\u20136), 9), 12), 14) of normal forms. Theorem 7. The list of local singularities (up to diffeomorphisms of the parameter space) of the boundary of the set of local transitivity for a three-parameter generic family of smooth compact surfaces with a smooth boundary consists of the singularities 1)\u201314). Theorem 8. The list of local singularities (up to diffeomorphisms of the parameter space) of the boundary of the zone of local transitivity for a generic three-parameter family of smooth compact surfaces with boundary and corners consists of all the singularities 1)\u201314) and k-cones with arbitrary k = 1, 2, . . . . Finally, let us consider the case where the indicatrices in three-dimensional space are the images of smooth maps of a given compact n-dimensional manifold Un of controls. 266 A.A. Davydov and V.M. Zakalyukin Theorem 9. The list of local singularities (up to diffeomorphisms of the target space) of the convex hull boundary of the image of a generic smooth map V : Un \u2192 R3 of a smooth n-dimensional compact manifold Un coincides with the list of local singularities of the convex hull boundary of a compact space curve for n = 1, or with the list of local singularities of the convex hull boundary of a smooth compact surface for n > 2, respectively. For n = 2 the list of local singularities of the convex hull boundary of a smooth closed surface also contains the \u2018adjacency with Whitney umbrella\u2019 singularity 12) described in Theorem 3. Theorem 10. The list of local singularities (up to diffeomorphisms of the parameter space) of the boundary of the zone of local transitivity for a generic threeparameter family of smooth maps V : Un \u2192 R3 of a smooth compact manifold Un of dimension n > 3 coincides with the list of germs 1)\u20136), 9), 12), and 14) of normal forms in Theorem 6. Theorem 11. The list of local singularities (up to diffeomorphisms of the target space) of the convex hull boundary of the image of a generic map V : Un \u2192 R3 of a smooth n-dimensional compact manifold Un with boundary and corners coincides with the list of local singularities of the convex hull boundary of a compact space curve with endpoints for n = 1, or with the list of local singularities of the convex hull boundary of a smooth compact surface with boundary and corners for n > 2, respectively. For n = 2 the list also contains the \u2018adjacency with Whitney umbrella\u2019 singularity 12) described in Theorem 3. Theorem 12. The list of local singularities (up to diffeomorphisms of the parameter space) of the boundary of the zone of local transitivity for a generic threeparameter family of smooth maps V : Un \u2192 R3 of a smooth n-dimensional compact manifold Un with boundary and corners coincides with the list of generic local singularities of the convex hull boundary of a smooth compact space curve with endpoints for n = 1, or with the list of generic local singularities of the convex hull boundary of a smooth compact surface with boundary and corners for n > 2, respectively. Remark 5. The lists described above show that generically the boundary of the convex hull and the boundary of the zone of local transitivity are Lipschitz. Apparently, this is true in any dimension. Remark 6. Moreover, in all cases we have the following statement, which provides an example of the principle [19] that \u2018good cases\u2019 dominate in many control-theory constructions, in contrast to Arnold\u2019s principle of \u2018fragility of the good\u2019, which is typical in singularity theory [30]. Consider a germ K of the generic transitivity zone whose base point corresponds to the origin being in a C1-smooth germ of the convex hull. Then either the boundary \u03a3 is smooth, or K is on the \u2018larger\u2019 side of the germ of the complement of \u03a3, which cannot be embedded in a half-space in any smooth local coordinates. Remark 7. The boundaries of the convex hulls of curves and surfaces contain ruled and flat pieces. These features are lost under a diffeomorphism: different arrangements of flat domains and straight lines on ruled pieces may correspond to diffeomorphic singularities. The statements of Theorems 1\u201312 contain only this rough Controllability of non-linear systems 267 classification up to diffeomorphisms. However, all possible cases of the affine structure of singularities can be found by the methods described below. 3.2. Legendre transformations and support hyperplanes. In this section we present the basic constructions and propositions ([25], [27]) which are involved in the proofs of the theorems formulated in the previous section and which may prove to be useful in the study of singularities in many other problems. First, let us note that if the indicatrix is given as the image of a map Rn \u2192 Rm with n > m, m = 2, 3, then, according to Thom\u2019s and Mather\u2019s classical results on the classification of singularities of generic maps (see, for instance, [17]), in the generic case only points of smoothness of the visible contour can appear on the boundary of the convex hull of the indicatrix. The only exception is the case of the \u2018Whitney umbrella\u2019 singularity R2 \u2192 R3. In all other cases it suffices to consider the indicatrices which are embedded curves and surfaces (possibly, with boundaries and corners), and their families. Surfaces with boundaries and corners, as well as curves with endpoints, can be treated uniformly as particular cases of stratified submanifolds embedded in R3. A collection J = {I1, . . . , Is} of closed embedded submanifolds I kj j (strata) of dimensions kj will be called a stratified submanifold if it contains a unique stratum of highest dimension, and any other stratum of lower dimension belongs entirely to some other stratum of higher dimension. For indicatrices in three-dimensional space we only deal with one-dimensional manifolds I1 with endpoints or with a smooth surface I2 in R3 on which there is a smooth curve I1 (a boundary) or two mutually transversal curves I1 1 and I1 2 (the sides of a corner) whose intersection point is treated as a separate stratum. However, the main constructions described below remain valid in the general multidimensional case. Thus, in the general case to a point q of a k-dimensional submanifold I we assign the set (diffeomorphic to the sphere S3\u2212k\u22121) of germs at q of all co-oriented planes tangent to the submanifold at this point. All such germs form a smooth Legendrian submanifold LI \u2248 I\u00d7S3\u2212k\u22121 in the space ST \u2217R3 of co-oriented contact elements. Forgetting the base point of the germ, we obtain the projection \u03c0\u2217 on the space of all co-oriented planes of this submanifold, which is a Legendre map of the submanifold LI to the dual space R\u03023. The image (wavefront) I\u0302 = \u03c0\u2217(LI) of this projection is called the Legendre transform of the initial submanifold I, or the dual surface to I. Denote by I\u0302A the germ of the dual surface to I at the set of all planes tangent to I at the point A. The Legendre transform for a stratified submanifold J is defined as the collection J\u0302 of the Legendre transforms \u03c0\u2217(LIj ) of all the strata Ij . A manifold Jc with boundary or corners is regarded as a subset of the stratified submanifold determined by the corresponding inequalities. Therefore, the corresponding Legendrian submanifold LJc is taken to be the subset of tangent elements at the points of Jc, and the Legendre transform J\u0302c is formed by the corresponding subsets of the dual surfaces of the strata of J . The space of all co-oriented planes in R\u03023 is fibred over the sphere of unit normals (with one-dimensional fibre of parallel planes). The boundary H(\u0393) of the convex hull of the compact subset \u0393 is determined by the set P (\u0393) of support planes of \u0393. Namely, for any unit normal we choose, among all the parallel planes intersecting 268 A.A. Davydov and V.M. Zakalyukin this hull, a plane P that corresponds to the maximum value of the coordinate whose gradient is directed along this normal. We note that a plane P \u2208 R\u03023 is a support plane of J if and only if it is tangent to some of the strata Ij at one or more points, or in other words, belongs to the image \u03c0\u2217LJ . For any support plane P we denote by SP = \u03b3\u22121(P ) the set of points at which this plane is tangent to \u0393. We call it the base of P and its points the base points. The number \u00b5P of distinct points of the base SP will be called its multiplicity. Thom\u2019s transversality theorem [17] for the space of multi-jets of maps or families of maps imposes some restrictions on the possible values of the multiplicity of support planes, types of Legendre maps \u03c0\u2217, and so on. Moreover, the condition that the origin O belongs to the support plane under consideration also imposes some constraints on the possible types of generic singularities of the Legendre map. For instance, for a generic surface \u0393 the multi-germs \u03c0\u2217 |SP are Legendre stable. Each germ is Legendre equivalent either to a germ with a singularity of A1 type (in a neighbourhood of such a point the surface is the graph of a Morse function and its quadratic form is negative definite) or to a germ with a singularity of A3 type (which corresponds to the Legendre map of the graph of the function h = x4 + (y\u2212 x2)2). The number of points in each base SP is at most three for singularities of A1 type and at most one for a singularity of A3 type. Thus, it is feasible to list all possible singularities of dual surfaces and to classify singularities of the convex hulls and the transitivity zones using the following advantageous properties of the Legendre transform. 1. The Legendre transformation of a generic hypersurface repeated twice is the identity map. Indeed, a Legendrian submanifold has two Legendre projections: \u03c0\u2217 and the standard projection \u03c0 : ST \u2217R3 \u2192 R of the fibre bundle. A smooth Legendrian submanifold LI is uniquely determined by its wavefront I\u0302 provided that the regular points of \u03c0\u2217 are dense in LI [31]. Hence the repeated Legendre transformation of the hypersurface I\u0302 yields the same Legendrian submanifold LI and the projection \u03c0, whose image \u03c0(LI) coincides with I. 2. The Legendrian submanifold LI2 of a smooth surface I2 has a regular projection on the surface itself: the tangent plane at each point is unique. In a neighbourhood of the convex hull of the base SP of a support plane P tangent to I2, the convex hull boundary H(I) is determined by the support planes which are close to P . The germ of the dual surface (I\u03022, P ), where P is the plane tangent to the generic surface I2 at a point q, is smooth provided that q is not a parabolic point. At a generic point of a parabolic line \u03b4, this germ is diffeomorphic to a semicubic cylinder (that is, to the bifurcation diagram of an A2 singularity). At isolated points of the line \u03b4, the Legendre transform has a singularity of A3 type, and the germ of (I\u03022, P ) is diffeomorphic to a swallowtail. We note that the interior points of A2 type cannot belong to the convex hull, since the tangent plane divides the surface in a neighbourhood of the point q. 3. The Legendrian submanifold LI1 for a smooth curve I1 is swept out by the circles Sq containing the tangent elements to I1 at its points q. The dual surface I\u03021 is ruled: in the affine chart the circles Sm correspond to straight lines. The set of support planes P tangent to the curve at a fixed point q forms a connected closed arc Eq on Sq. A support plane P passing through q \u2208 I1 can be rotated about the Controllability of non-linear systems 269 tangent to the curve until it touches the curve at some other point or becomes an osculating plane. The germ of the convex hull near the point q of a one-dimensional stratum I1 is determined by the germ Su(I1) on the arc Eq of support planes passing through q. On a generic curve some isolated points may have simple flattening (in a neighbourhood of such a point in the canonical Frenet frame the curve has the parametric form q1 = t, q2 = t2 + \u00b7 \u00b7 \u00b7 , q3 = t4 + \u00b7 \u00b7 \u00b7 ), while at all other regular points the germ of the curve has the canonical form q1 = t, q2 = t2 + \u00b7 \u00b7 \u00b7 , q3 = t3 + \u00b7 \u00b7 \u00b7 . In a neighbourhood of the osculating plane, which is given in both cases by the equation q3 = 0, the dual surface has a singularity of A3 type (swallowtail) in the flattening case and a singularity of A2 type (semicubic cylinder) in the regular case. The osculating plane at a regular interior point (one which is not an endpoint) cannot be a support plane: the curve is located on both sides of the plane (however, at an endpoint the osculating plane can be a support plane). 4. Any point q \u2208 R3 corresponds in the dual space to the plane q\u0302 consisting of all planes in R3 passing through q. The germ of the convex hull in a neighbourhood of a point q which is a zero-dimensional stratum of J (that is, a corner vertex or a curve endpoint) is determined by the germ of Su(J\u0302) on the convex domain Uq \u2282 q\u0302 consisting of all support planes passing through q. 5. The Legendre transforms I\u03021 and I\u03022 of two strata I1 \u2282 I2 are tangent along their intersection I\u03022 \u2229 I\u03021, which consists of the tangent planes to the stratum I\u03021 at the points of I\u03022. Indeed, the surfaces I\u0302i = \u03c0\u2217(LIi ), i = 1, 2, intersect along the set of common tangent planes at the points of the smaller stratum I1. Due to the involution property of the Legendre transformation, the planes tangent to \u03c0\u2217(LI) are points of the submanifold I itself. Therefore, the planes tangent to I\u03021 and I\u03022 at their intersection points correspond to the common points in I1 \u2229 I2 and hence coincide, as required. If the quadratic form of the surface I\u03022 is non-degenerate in the directions tangent to I2 and transversal to I1, then the tangency of I\u03021 and I\u03022 is of the first order. 6. A point Q \u2208 R3 corresponds in the dual space to the plane Q\u0302 consisting of all planes in R3 passing through Q. A point Q belongs to the convex hull boundary of a submanifold B if and only if the plane Q\u0302 is negatively supporting for the subset Su(B) of support planes of B: the open negative half-space of Q\u0302 does not contain points of Su(B), but the plane itself contains such points. In particular, the point O belongs to a convex surface X \u2282 R3 if and only if the dual plane O\u0302 is tangent to the dual surface X\u0302. 7. If O belongs to the convex hull of the base of the support plane for some generic three-parameter family of submanifolds B of dimension k, then the singularities of the Legendre transform at each base point are generic singularities of the Legendre transforms of k-dimensional submanifolds. Indeed, the dimension of the convex hull of the base consisting of l points is less than or equal to l\u2212 1. In order for the origin O to belong to the convex hull of the base, at least 3\u2212 l + 1 conditions must be satisfied. For l points to fall on the osculating plane, l conditions must hold. A degenerate point appears on the base if at least one condition holds. Thus, in this case more than three independent conditions must be satisfied, which is generically impossible for three parameters. 270 A.A. Davydov and V.M. Zakalyukin Now a classification of generic singularities of the boundary of the transitivity zone is obtained by listing all possible positions of the origin in the base of a support plane, describing the corresponding families of dual surfaces and their supporting subsets, and, finally, applying the following results in singularity theory. A diffeomorphism \u03a6: R\u03023 \u2192 R\u03023 mapping the plane O\u0302 onto itself and preserving its positive half-space is called admissible. If \u03a6 maps the support part Su(J1) of the Legendre transform of a stratified manifold J1 to the Legendre transform Su(J2) of another stratified manifold, then, clearly, the origin O belongs to the convex hull boundary H(J1) if and only if it belongs to H(J2). Thus, a family of admissible diffeomorphisms \u03a6m, fibred over a diffeomorphism of the parameter space, acts on the parameter-dependent families of surfaces I\u0302j and maps the respective transitivity zones to each other. Reducing the equations of the dual strata I\u0302j to normal forms via admissible diffeomorphisms, we obtain normal forms of the transitivity zone. For this purpose, consider the action of the diffeomorphism \u03a6m as a contact transformation on the product of the equations of all the components [32]. Denote by Pi(x), i = 0, . . . , n, the polynomials of degree ki in x \u2208 R of the form P0 = xk0 + \u2211k0\u22122 j=0 xja0j , Pi = xki + \u2211ki\u22121 j=0 xjaij , i = 1, . . . , n. Denote by a \u2208 RN the vector of coefficients of all these polynomials. We note that P0 is the standard miniversal deformation of the singularity Ak0\u22121, and Pi for i = 1, . . . , n are versal deformations of the singularities Aki\u22121. Lemma. Let gi(x, b), i = 0, . . . , n, be functions of x with parameters b \u2208 RN such that the values of gi(x, 0) and the values of their derivatives up to order ki \u2212 1 vanish at the origin (the function gi(x, 0) has singularity Aki\u22121 at the origin), and let g(x, b) = \u220fn i=0 gi(x, b). Then there is a contact equivalence consisting of a diffeomorphism (x, b) 7\u2192 ( X(x, b), B(b) ) and a non-zero function \u03d5(x, b) which reduces the function g to the form g(x, b) = \u03d5(x, b) \u220fn i=0 Pi(X, B). The assertion of the lemma is equivalent to the versality of the map described below with respect to a special group of equivalences. Consider the map G : Rk \u2192 Rn+1, G : x 7\u2192 ( g0(x), . . . , gn(x) ) , and in the target space consider the collection Y = \u22c3 {yi = 0} of coordinate hyperplanes. Let D be the group of diffeomor- phisms of Rn+1 which preserve Y ; that is, they have the form \u03b8 : (y0, . . . , yn) 7\u2192( h0(y)y0, . . . , hn(y)yn ) . Two maps G1 and G2 are called Y -contact equivalent if for some family of diffeomorphisms \u03b8x in the group D that depend on a parameter x we have G2(x) = \u03b8x \u25e6G1(X(x)) for some change of variables x 7\u2192 X(x). For a family Gm : Rk \u00d7 RN \u2192 Rn+1 of maps G depending on parameters a in RN , a parameter-dependent contact equivalence is defined in a natural way: the parameters a are replaced via a diffeomorphism by parameters b; a diffeomorphism of the form x 7\u2192 X(x, a) acts on the variables x; a family of diffeomorphisms \u03b8x,a : Rn+1 \u2192 Rn+1 acts on the target space Gm. The notions of versal and infinitesimally versal deformations of a map G are straightforward. Obviously, for this group D of transformations an analogue of the versality theorem holds, since the group is geometrical in the sense of J. Damon. We note that the map P\u0303 : (x, a) 7\u2192 (P0, . . . , Pn) is an infinitesimally versal deformation of the map x 7\u2192 (xk0 0 , . . . , xkn n ). The versality of the map P\u0303 implies the assertion of the lemma: under an equivalence in the group D each component Controllability of non-linear systems 271 is multiplied by a non-zero factor; hence the product is multiplied by a non-zero factor." ] }, { "image_filename": "designv11_62_0001014_iros.2013.6696949-Figure1-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0001014_iros.2013.6696949-Figure1-1.png", "caption": "Fig. 1: Simple Model", "texts": [ " The first algorithm is a fall indicator that determines the necessity of support change to stabilize the perturbed model. The second is a support change algorithm that computes the new stabilizing contact configuration, in case of support change. We first review the simple model. We then describe the push recovery approach and finally we review the two algorithms used by the approach. A. Review of the simple model and related assumptions The simple model is detailed in [16] and is illustrated in Fig.1. It has completely actuated massless limbs. It consists of a point mass m at the COM and of n non-coplanar contact surfaces. We assume a null inertia of the body about the COM. The control variables of the system are the ground reaction wrenches with Wi = ( ft i \u03c4 t i )t being the wrench at the ith contact and fi \u2208 IR3 and \u03c4 i \u2208 IR3 being respectively the contact force and torque. The dynamics of the model are described by NewtonEuler equations and the model is subject to several constraints: \u2022 The COP of each contact belongs to the contact surface", " Review of the Fall indicator [17] Given an initial contact configuration of the model and a COM perturbed state ( XBR, X\u0307BR ) , the fall indicator informs whether the simple model can be stabilized while maintaining the same contact configuration. Bretl generates in [2], for a given contact configuration, a convex static stability region over which the COM of a static system must lie. A direct fall indicator consists of a linear step by step optimization formulated in terms of the variable Wk (details in [17]) and that minimizes the simple model (Fig.1) kinetic energy, which is a point mass kinetic energy, along the direction of perturbation d collinear to X\u0307BR (See Fig.3) while satisfying the model dynamics and constraints; if the model reaches a static state inside the initial static region (before reaching Xbd), then the system can remain in a static state and is therefore stabilized without support change; otherwise, the COM goes beyond Xbd and the static region should change to include the COM and allow it to stop inside the new static region; a support change is then necessary to obtain a new different static region" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0002758_icems.2017.8055934-Figure6-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0002758_icems.2017.8055934-Figure6-1.png", "caption": "Fig. 6. Cross section of simplified slotless model.", "texts": [ " 50 100 150 200 250 1 2 3 4 5 6 7 8 9 10 U M F ( N ) Magnet thickness (mm) parallel on-load parallel noload Halbach onload Halbach noload Radial onload Radial noload (a) 12-slot/10-pole magnetizations. In order to investigated in more details, the model needs to be simplified. Due to the fact that the reason of UMF with rotor eccentricity is the unequal airgap length instead of slotted harmonics and asymmetric windings, a no-slot model with linear soft magnetic material under no-load condition is employed to get rid of the influence of slotting effect, armature field and saturation. The cross section of simplified model, i.e. slotless model, is shown in Fig. 6 which has the same rotor comparing with the conventional model, and the UMFs of slotless model is shown in Fig. 7. As can be seen, the UMFs of slotless model and conventional model have very similar characteristics, and the difference is mainly due to the saturation of the soft magnetic material. Consequently, the slotless model will be employed for further investigation. The UMF due to rotor eccentricity origins from the unbalanced airgap flux density distribution caused by the unequal airgap length which is shown in Fig" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0003202_978-981-13-3119-0_29-Figure1-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0003202_978-981-13-3119-0_29-Figure1-1.png", "caption": "Fig. 1 Reference frames and hydrodynamic forces on the robotic gliding fish", "texts": [ " A robotic gliding fish is a miniature underwater glider with an idea of exploring the platform by fitting fins/tails/control planes in future. The tail fin (hereby designated as rudder) is a control surface and one of the sources for forces and moments. The model of the gliding fish is considered as a rigid body which also includes two internal movable masses (one for roll and another for pitch control) and a bladdertype buoyancy engine (which will be studied in future work). The deflected rudder provides forward thrust, side force and yaw moment. Figure 1 shows the reference frames and hydrodynamic forces on the glider, and Fig. 2 shows themass distribution. A total of three reference frames are designated, and the body position, velocity and other parameters are defined as per one of these three axes. These are the inertial frame I0, the body-fixed frame B0 and the flow frame \u03c00. The body-fixed frame (B0) of reference has its origin at the geometry centre of the vehicle. Bxb axis is along the Table 1 (continued) Notation Description Notation Description Kq3 Yaw moment coefficient with respect to V2 M\u03b1 MP Pitch moment with respect to \u03b1V2 S Surface area M\u03b2 My Yaw moment with respect to \u03b2V2 M\u03b4 My Yaw moment with respect to \u03b4V2 Fig. 2 Mass distribution of the robotic gliding fish body\u2019s longitudinal axis pointing forward (towards the head of the fish); the Byb and Bzb axes are formed as per the right-hand orthonormal principle and are as shown in Fig. 1. In the inertial frame, I0(xyz), I0z axis is along the gravity direction and I0x/I0y are defined in the horizontal plane and origin is a fixed point in space. The stationary mass of the robotic gliding fish ms comprises of three parts: hull mass mh, point mass mr used for roll control with displacement rp and ballast mass mb. Here we assume that the buoyancy engine is positioned and effects of change in buoyancy occur at the origin of body axis (being the CG of the fish). The pitch control massmp is located at rp distance from body-fixed axis origin" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0001258_icra.2013.6631391-Figure6-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0001258_icra.2013.6631391-Figure6-1.png", "caption": "Fig. 6. Underactuated biped model with knees and upper body", "texts": [ " In this case, the effect of speeding-up attained with the use of semicircular feet would be the primary factor. From Fig. 5 (b), we can see that \u03bbI4 tends to increase and become positive where m2 is sufficiently large. The value of m2 where \u03bbI4 reaches zero from negative becomes smaller as mH decreases. As in the previous result, it is thought to be causally related to speeding-up of the generated gait but the details are still unclear. More investigations are necessary. Here, we briefly report the analysis results of the effects of an upper body. Fig. 6 shows the model; we add an upper body to the underactuated bipedal model of Fig. 1 incorporating the joint torque, u4, between the upper body and the swing leg. We apply an output PD control to the upper body angle, \u03b85, for maintaining it at the desired one, \u03b85d [rad]. The walker can generate stable level gaits by choosing the parameters of the upper body and PD gains. The physical parameters except the upper body ones were chosen as listed in Table I. Fig. 7 shows the simulation results of level dynamic walking where mt = 10 [kg], Lt = 1" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0000404_amm.29-32.1602-Figure2-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0000404_amm.29-32.1602-Figure2-1.png", "caption": "Figure 2. The rollingbearing test setup. Figure 3. Schematic diagram of the bearing test setup. The vibration data were acquired under heavy-load with 660 rpm of the motor speed. The", "texts": [ " The FKNN is a simple algorithm that is used to assign class membership as a function of the pattern distance from its K nearest neighbors and those neighbors\u2019 memberships in the possible classes [15]. As a result, the FKNN has ability to deal with problems where pattern data overlap. Thus, the FKNN is suitable for the rolling bearing multi-fault state identification. The detail of the FKNN algorithm was given in literature [14]. A flow chart of the diagnosis method proposed is illustrated in Fig. 1. Experimental analysis of the 6308 type rolling bearing with man seeded faults has been carried out using an experimental setup. The fault simulator setup with sensor is shown in Fig. 2 and the Schematic diagram of the test setup shown in Fig. 3. A variable speed DC motor with speed up to 3000 rpm is the basic drive. The sensor used is two piezoelectric accelerometers (CA-YD-106) which are mounted on the flat surface horizontally and vertically along the radial direction of the rolling bearing housing, respectively. The software DASP is used for recording the signals. The vibration was measured under eight different rolling bearing fault conditions: pattern 1-roller cage fracture (RCF), pattern 2- roller spalling (RS), pattern 3-roller erosion (RE), pattern 4-inner race spalling (IRS), pattern 5-inner race erosion (IRE), pattern 6-outer race erosion (ORE), pattern 7-inner and outer race spalling (IORS), pattern 8-roller and outer race spalling (RORS), respectively" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0003985_icamechs.2019.8861668-Figure5-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0003985_icamechs.2019.8861668-Figure5-1.png", "caption": "Fig. 5. Definition of the joint angles", "texts": [ " '&% %' & % (& % ) )* + ) , --- III. PROPOSED ROBOT We propose a multi-legged robot with a flexible leg mechanism that can switch the stiffness of the legs during locomotion. The mechanism of the proposed robot is presented in Fig. 4. The robot has nine segments and each segment has two legs as shown in Fig. 4. Each segment has two motors; one for vertical movement and the other for horizontal movement. The same pattern of movement with a phase delay was repeated, and the locomotion pattern was realized as shown in Fig. 5 , Fig. 6. Repetition (1) and (2) of the simple pattern reduced amount of information which had to be processed to allow for the control of the multiple legs. However, typically in complex environments, the robot when moving must adapt to the terrain. In the proposed robot, the adaption during locomotion can be realized through the flexibility of the trunk and the legs. The robot can passively adapt to the rough terrain by utilizing its flexibility and through the repetition of a simple locomotion pattern" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0003985_icamechs.2019.8861668-Figure11-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0003985_icamechs.2019.8861668-Figure11-1.png", "caption": "Fig. 11. Mechanism of change in direction", "texts": [ " The trunk was made of duplex bellows as shown in Fig. 10 i.e. smaller bellows inside larger bellows for the realization of different stiffness i.e. for small flexure and large flexure, respectively. The trunk is made up of series of connected rectangular prisms, and has high stiffness in the vertical direction for upliftment of its weight and low stiffness in the horizontal direction to allow for the robot to turn. Two wires were installed on both sides of the trunk and were connected to an active pulley with a servomotor as shown in Fig. 11. This allowed for the robot to turn by pulling the wires using the servomotor. The trunk was moved by wires, and this made it to flexible even when the wire was being pulled. Therefore, the robot could passively adapt to rough ground using its flexibility. & A prototype was developed to evaluate the applicability of the proposed leg mechanism. The prototype is shown in Fig. 12, and its specifications are given in Tables I and II. To demonstrate the effectiveness of the proposed robot, we developed a prototype and conducted experiments" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0002642_978-3-319-66866-6_2-Figure2-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0002642_978-3-319-66866-6_2-Figure2-1.png", "caption": "Fig. 2. Relevant design guidelines concerning the example of honeycomb structures", "texts": [ " In form of knowledge storages, such as checklists or design catalogues, these guideline provides standard values to ensure manufacturability [19]. For example, boundary conditions for minimal wall thickness or diameter, depending on the building direction of a component, are defined [20, 21]. Design guidelines are partially available for different parameter sets, which describe a specific machine and material. Furthermore, general statements are available, which define optimal orientation, arrangement or positioning of components in the process chamber or the necessity for cleaning openings [18]. As shown in Fig. 2, using the example of a honeycomb, minimum wall thicknesses, realizable overhangs and 45\u00b0 angles to avoid support structures have to be considered. Furthermore, the element size b is limited by a minimal diameter. A material database to specify the anisotropic behavior of a selective laser melting component is defined. Using this database during computer-aided modeling and simulation, the consideration of building directions is possible. Analogous to conventionally processed materials, powder alloys for selective laser melting differ in mechanical properties depending on the post-process" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0003835_ccdc.2019.8832418-Figure7-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0003835_ccdc.2019.8832418-Figure7-1.png", "caption": "Fig 7. The minimum distance in link coordinate", "texts": [ " 1 1 1 - - + - A A B A x x R d y R y y R z z R d \u2264 + \u2264 \u2264 \u2264 + (12) Assuming that link AB and obstacle sphere are intersected, the minimum distance is a scalar without any change in any coordinate. The solution of the collision point K is equal to obtaining the Jacobian matrix 0J of the collision point [12]. To calculate the Jacobian matrix 0J accurately, we select vector product method to get scalar KA . Therefore, it is easier to calculate it directly in the local coordinate than convert the minimum distance to the base coordinate as shown in Fig 7. From the intersection test of links and obstacles, we know that the coordinate of A is ' ' '( , , )A A Ax y z and the coordinates of the B is ' ' '( , , )B B Bx y z , and the central coordinate is 1 1 1 1D ( , , )x y z . Therefore the collision point is ' ' 1( , , )A AK x y z , the distance between point K and the point A is ' 1 AKA x x= \u2212 and the minimum distance is ' min 1 Ad y y= \u2212 . The coordinate value of the target point can be directly used to obtain the result. 4186 The 31th Chinese Control and Decision Conference (2019 CCDC) Through the above analysis, we can see that the minimum distance mind applied to the obstacle avoidance of redundant robots can save the time of environmental detection, which satisfies the requirements on real-time performance" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0002450_978-3-319-60867-9_17-Figure1-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0002450_978-3-319-60867-9_17-Figure1-1.png", "caption": "Fig. 1. Kinematic scheme of legs of the Bioloid humanoid", "texts": [ " The formulation of the optimization problem and the process to solve it are presented in third section. Then, this method is applied for walking optimization of the Bioloid robot for a rectilinear path. Finally the conclusion of the work is presented. 2 Specification of a Walking In the walking pattern of a biped robot, the desired poses for both the pelvis and the oscillating (or free) foot are specified with respect to a world\u2019s frame (xW-yW-zW) as time functions. The points for position specification of these bodies are Op (pelvis) and Of (free foot), showed in Fig. 1. The positions are given in Cartesian coordinates. For orientation with respect to the world\u2019s frame, the Bryant angles k, l and m are applied to frames xp-yp-zp and xf-yf-zf attached to the pelvis and the free foot, respectively. Both frames and the world\u2019s frame are shown in Fig. 1. The Bryant angles correspond to successive rotations applied in the order x-y-z to a frame whose orientation initially matches the world\u2019s frame in order to obtain the desired orientations. The equations that define all the coordinates as time functions are those proposed in [7]. Some of the main walking parameters are appreciated in Fig. 2. The cycle of a step is composed by two phases: single support phase (SSP) and double support phase (DSP). The first one is achieved during a period TS, when only one foot is in contact with the floor while the other foot is moving forwards" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0001087_amr.591-593.1519-Figure1-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0001087_amr.591-593.1519-Figure1-1.png", "caption": "Fig. 1 Structure of a pendulum-driven spherical robot Fig. 2 Simplified model for the longitudinal motion", "texts": [ " Spherical robots are more versatile, less exposed to physical conditions, and they have greater resistance against object collisions. Moreover, the exoskeleton provides an efficient cover for the driving mechanisms and sensory equipments inside it. Spherical robots can be categorized into different types according to their driving mechanisms [1]. Compared with other types of spherical robots [2-4], a pendulum-driven spherical robot [5-7] has a simpler structure, which makes it easier to be maneuvered. The structure of a pendulum-driven spherical robot is shown in Figure 1. Longitudinal motion is a basic form of locomotion of pendulum-driven spherical robots. In this paper, a hierarchical sliding mode control approach is presented for stable control of the longitudinal motion. In the proposed controller, a double layer structure is used to guarantee the stability of the whole system, and sliding surfaces are utilized to drive the output tracking errors to zeros. We start with a simplified planar model, only considering no slip longitudinal motion on flat surfaces" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0000822_1754338x10392310-Figure2-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0000822_1754338x10392310-Figure2-1.png", "caption": "Fig. 2 Ball roll test set-up for: (a) Australian football, and (b) the change in the ramp for the cricket ball roll tests", "texts": [ " The frictional force and the coefficient of friction were calculated again; the change between the frictional forces for the pre- and post-drag measures was recorded as abrasion, and the coefficient of friction was recorded as friction. The cricket ball used for all these tests was a twoseam ball of weight 160 \u00b1 10 g, the Australian football was 720\u2013730 mm in circumference and inflated to The test used to measure the ball roll involved rolling a ball, both an Australian football and a cricket ball, down a 1 m ramp with a 45\u25e6 inclination. The dimensions of the ramp were in accordance with BS EN 12234:2002 [9]. The changes in the ramp for the cricket test can be seen in Fig. 2. The ball was released by the operator removing their fingertips from the ball, Proc. IMechE Vol. 225 Part P: J. Sports Engineering and Technology at PURDUE UNIV LIBRARY TSS on May 31, 2015pip.sagepub.comDownloaded from Proc. IMechE Vol. 225 Part P: J. Sports Engineering and Technology at PURDUE UNIV LIBRARY TSS on May 31, 2015pip.sagepub.comDownloaded from and the cricket ball was positioned with the seam of the ball across the ramp. Electronic timing gates were used to measure the change in the velocity over a distance of 1" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0002885_978-3-319-66697-6_96-Figure2-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0002885_978-3-319-66697-6_96-Figure2-1.png", "caption": "Fig. 2 Numerical model", "texts": [ " They are connected through two rigid shafts, inner shaft called sun shaft and external shaft called carrier shaft and motor drive. This test bench is well described in (Hammami et al. 2015, 2014). The external load is obtained by the adding mass on the arm which is fixed on the free ring. Accelerometers are mounted on the reaction and the test ring as well as on the reaction and test carrier. Data obtained are transferred to the acquisition system \u2018LMS test lab.\u2019 A numerical model referred to the experimental test bench is presented (Fig. 2). It is a tridimensional model based on that developed of Karray et al. (2016). The model composed with two-stage planetary gear the components of each stage are the carrier (c) the ring (r) the sun (s) and the planets (p1, p2, p3). The planet set are supposed as identical and equally spaced. These components are modeled as rigid bodies with mass mij, inertia Iij.each components is supported by bearing with stiffness kbijk, where i = c, r, s, p1, p2, p3; j = r, t. in direction k = x, y, z, \u03c6, \u03a8,\u03b8 For each stage, the ring gear and the sun gear are respectively linked to the three planets via mesh stiffness Krpt1, Krpt2, Krpt3 and Kspt1, Kspt2, Kspt3 for the test gear and Krpr1, Krpr2, Krpr3 and Kspr1, Kspr2, Kspr3 for the reaction gear", " Each shaft is characterized, respectively, by axial stiffness ksa kca; flexural stiffness ksf kcf and torsional stiffness kst kct. The equation of motion of this model is: Mq\u0308+Cq \u0307+ \u00f0Kb +Ke\u00f0t\u00de\u00deq=F\u00f0t\u00de+T , \u00f01\u00de where q is the vector degree of freedom, M is the mass matrix, C damping matrix, Kb is the bearing, and shaft stiffness matrix, Ke(t) is the time-varying stiffness matrix, F(t) is the excitation force due to the rotation of the carrier and T is the external force vector applied to the system. The computation of dynamic response in presence of gravity of carrier is done according to Fig. 2. The gravity of carriers introduces a variation on the distances between sun-planets and ring planets during the running of the system. Then, the values of gear mesh stiffness decrease as the distances between gears. Figure 3 shows the evolution of gear mesh stiffness ring-planets during one period of rotation of carrier with effects of gravity of carrier. Results presented are obtained for the input rotational speed of motor 1498.5 rpm, and an external load 100 Nm. An external force due to the rotation of carrier with period Tc/3 = 0" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0002236_s12204-017-1827-3-Figure2-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0002236_s12204-017-1827-3-Figure2-1.png", "caption": "Fig. 2 The typical structure of quad-rotor", "texts": [ " One approach to reduce the chattering is to replace the discontinuous signal function with a continuous approximation: \u03c3/(\u2016\u03c3\u2016 + \u03b4), where \u03b4 is a positive constant. 3 Example The experiments proposed in this paper are based on Qball-X4 quad-rotor semi-physical simulation platform produced by Quanser Inc. Figure 1 shows the typical structure of micro quadrotor which has four propellers in cross configuration. A linearized model used for controller design can be developed from the nonlinear model of the quad-rotor. For the following discussion, the axes of the quad-rotor are shown in Fig. 2. Roll, pitch, and yaw are defined as the angles of rotation about the x, y, and z axes, respectively. The global axis distances are also denoted as x, y and z and they are defined with the same orientation as the quad-rotor sitting upright on the ground. In the quad-rotor semi-physical simulation platform, the host computer is connected to lower computer (the quad-rotor body) with wireless network according to the network delay and the packet loss in the actual situation, and the time delay in state is considered", " Before that, we model the actuator dynamics first. The thrust generated by each propeller is modeled by the following first-order system: F = K \u03c9 s + \u03c9 u, (26) where u is the pulse width modulation (PWM) input to the actuator, \u03c9 is the actuator bandwidth, and K is a positive gain. A state variable v will be used to represent the actuator dynamics, and it is defined as follows: v = \u03c9 s + \u03c9 u. (27) Then we obtain the state space equations: v\u0307 = \u2212\u03c9v + \u03c9u. (28) The motion of the quad-rotor in the vertical direction (see Fig. 2) is affected by all the four propellers. The dynamic model of the quad-rotor height can be written as Mz\u0308 = 4F cos \u03b8 cos\u03c6 \u2212 Mg, (29) where F is the thrust generated by each propeller; M is the total mass of the device; \u03b8 and \u03c6 represent the roll and pitch angles, respectively; g is the gravity acceleration. As expressed in Eq. (29), if the roll and pitch angles are nonzero, the overall thrust will not be perpendicular to the ground. Assuming these angles are close to zero, the dynamics equations can be linearized to the following state space form: \u23a1 \u23a2 \u23a2 \u23a3 z\u0307 z\u0308 v\u0307 \u23a4 \u23a5 \u23a5 \u23a6 = \u23a1 \u23a2 \u23a2 \u23a3 0 1 0 0 0 4K M 0 0 \u2212\u03c9 \u23a4 \u23a5 \u23a5 \u23a6 \u23a1 \u23a2 \u23a2 \u23a3 z z\u0307 v \u23a4 \u23a5 \u23a5 \u23a6 + \u23a1 \u23a2 \u23a2 \u23a3 0 0 \u03c9 \u23a4 \u23a5 \u23a5 \u23a6u + \u23a1 \u23a2 \u23a2 \u23a3 0 \u2212g 0 \u23a4 \u23a5 \u23a5 \u23a6 " ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0001454_msf.768-769.628-Figure1-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0001454_msf.768-769.628-Figure1-1.png", "caption": "Fig. 1: Sketch of the preparation showing the weld layers.", "texts": [ " E410NiMo was deposited on base metal 1 and the other two filler alloys on base metal 2. The chemical compositions of the base metals and the weldments (including dilution) are presented in table 1. Carbon, Sulfur, Oxygen and Nitrogen were determined by combustion and IGF (ASTM E1019), other elements by ICAP-EAS (ASTM E1479). Five layers of each filler metal were deposited using a Scompi robot in an 18 mm height 45 o Vpreparation with a root radius of 6 mm made in two plates of 238 mm x 125 mm x 31 mm (BM 1) and one of 225 mm x 150 mm x 38 mm (BM 2) (see Fig. 1). No external restraint was applied to the plates during welding. An oscillation of the welding torch was produced by the robot to get only one pass per layer. Welding parameters are presented in table 2. The contour method has been used to characterize the residual stresses. This residual stress measurement method, developed by Prime [9, 10], has been used for numerous welding residual stress characterizations [3, 11-15]. Based on a superposition principle, the contour method consists in the imposition of measured distortion profile of a newly cut surface to a numerical model" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0001400_interact.2010.5706149-Figure2-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0001400_interact.2010.5706149-Figure2-1.png", "caption": "Figure. 2. Tracked vehicle model", "texts": [ " By controlling the lengths of the manipulator, the upper platform follows the trajectory commands generated by the system and the driver. III. VEHICLE DYNAMICS MODEL In this section, the tracked vehicle model is developed to predict the vehicle's longitudinal and lateral response due to the driver\u2019s inputs. A simplified vehicle dynamic model [5] is considered with torsion bar suspension system and for simplicity all other sources of dynamic forces like engine vibration, structural flexibility, transmission vibration, etc., are ignored. Fig. 2 shows the tracked vehicle model at rest on the XI-YI plane. When the vehicle is not moving, the load under the road wheels are given as n gM gMP s uji 2 , (4) where i, j denotes the road wheel nos. on the left & right side respectively, n denotes the total no. of wheels on each side, Ms is the mass of the vehicle body, Mu is the unsprung mass of the suspension system and g is the gravitational acceleration. The wheels are assumed to be evenly spaced and L gives the length between and first and last wheel", " When the vehicle has forward acceleration ax at time t, the normal forces on the road wheels is given by Lnn shamni n gM gMP zxs uji )1( 123 2 , (5) where h is the height of the vehicle's center of mass when it is at rest and sz is the shift of the vehicle's center of mass along the positive direction of zc at time t. From (5) the deflection of the first wheel and last wheel is given by s zx f LKnn shamn )1( 13 (6) s zx r LKnn shamn )1( 13 (7) where Ks indicates the stiffness of the road wheel suspension. At the next instant tt , where 10 t , the shift of the vehicle's center of mass due to pitching of the vehicle body is 2 1 rf zs (8) and the rotation angle of coordinate frame Fv along Yv axis is L rf y (9) Considering the lateral motion of the vehicle in the Fig. 2, when the vehicle is not moving, the static forces exerted by the ground on the left and right tracks are D dgM gMnW Rs uL (10) D dgM gMnW Ls uR (11) Considering the vehicle having a velocity of vx in the positive direction of Xv and angular velocity z in the positive direction of Zv , the centripetal acceleration is zx x y v v a 2 (12) Similar to the longitudinal case, we can obtain relative quantities at time t as D shaM D dgM gMnW zysRs uL (13) D shaM D dgM gMnW zysLs uR (14) s zys RL KDn shaM (15) where L and R denotes the deflection of the left and right suspensions" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0000130_j.physc.2013.04.021-Figure11-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0000130_j.physc.2013.04.021-Figure11-1.png", "caption": "Fig. 11. The three regions of the HTS winding. Winding parts I and III are winding end. Winding parts II and IV are the parts in slots.", "texts": [ " Therefore, adding some air space between HTS coil and primary core is an effective method to improve critical current of HTS windings in the AFIM. The magnetic field in slot will affect the AC losses of HTS windings. Using previous results of magnetic fields in HTS windings, the AC losses of HTS windings in HTS AFIM can be calculated. First, we assume the relative permeability of primary core is infinity so that there is no leakage field around winding end except self-field from each tape. Then, classify the HTS winding into four parts: winding part I, winding part II, winding part III and winding part IV, as shown in Fig. 11. Winding parts I and III are winding end. Winding parts II and IV are the parts in slots. Then, improve Kim\u2019s model and get the nonlinear equation system [10]: Jc \u00bc J0 B0 B0 \u00fe jBx\u00f0x; y\u00de \u00fe Bexj where c\u00f0x\u00de < jyj < H I=D 2 R H c\u00f0x\u00de J0 B0 B0 \u00fe jBx\u00f0x; y\u00de \u00fe Bexj dy 2c\u00f0x\u00de where jyj < c\u00f0x\u00de 8>>>< >>>: \u00f018\u00de External magnetic field Bex applied on parts I, II, III and IV are different. For parts I and III, the value of Bex can be calculated by FE method; for parts II and IV, the value of Bex is the maximum perpendicular magnetic field calculated previously" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0003737_s12553-019-00355-y-Figure1-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0003737_s12553-019-00355-y-Figure1-1.png", "caption": "Fig. 1 Schematic representation of ball-on-disc tribometer. (1) AISI 316L ball, (2) Lubricant, (3) UHMWPE disc, and (4) AISI 316L support", "texts": [ " For clarity and simplicity purposes, we designate the kinematics for which SRR>3 as super-sliding. To improve our understanding of protein assisted lubrication (PAL) during the rubbing of AISI 316L on UHMWPE under super-sliding conditions when lubricated with FBS solutions at 37 \u25e6C, we performed a systematic study of the behavior of the COF within the chosen experimental intervals. Presumably, under these experimental conditions, protein interactions should have less of an effect on lubrication, leading to EHL as the expected lubrication regime for the tribosystem. 2Materials andmethods Figure 1 depicts a schematic representation of the experimental apparatus, described by Barceinas et al. [24]. In this experimental setup, an AISI 316L stainless steel ball (diameter: 19.05 mm, maximum surface roughness of 0.127 \u03bcm), is loaded against the face of an UHMWPE disc, backed with an AISI 316L stainless steel disc. The instrument only takes integer values for load L, in newton. The lubricant cup possesses a temperature control graded in Celsius. The ball-on-disc tribometer measures the COF as a function of SRR = |Vr/Vm|, where the relative and entrainment speeds at the contact point are defined as |Vr | = |V1 \u2212 V2| and |Vm| = |(V1 + V2)/2|, respectively" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0002082_978-3-319-49137-0_1-Figure4-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0002082_978-3-319-49137-0_1-Figure4-1.png", "caption": "Fig. 4 Schematic diagram of the thin-layer spectroelectrochemical cell from Ref. [55]", "texts": [ " Hydrodynamic voltammetry techniques using such cells facilitate the analysis of fluorescent solution-phase electrogenerated species, by modeling, from a proposed model, their spatial and temporal distributions in the flow cell, and finally correlate the fluorescence intensity to the electrode current. In the beginning of 2000s, a new evolution of thin-layer electrochemical cells, able to operate at variable thicknesses, was designed. A first version [55], dedicated to UV\u2013Vis and fluorescence spectroscopies, was proposed by Yu et al. and is composed of a working electrode disk inserted normally to the emission detector, in a Teflon body up to an experimental chamber. The excitation light and the emission detector were placed at 90\u00b0 to each other. This type of flow cell is illustrated in Fig. 4. A second version, proposed by Levillain et al. [14, 25] and inspired by the works of Salbeck [56] and Wertz [44], can be used in different configurations due to its wide versatility and dedicated to UV\u2013Vis, IR or fluorescence spectroscopies. Regarding F-SEC, the excitation light makes a 30, 45 or 60\u00b0 angle with the electrode surface and the emission light is recorded normal to the surface. The path length can range from few micrometers (thin-layer conditions) to few millimeters (diffusion layer conditions)" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0001286_iros.2011.6094587-Figure2-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0001286_iros.2011.6094587-Figure2-1.png", "caption": "Fig. 2. Conventional setups for the pose capturing: (a) eye-in-hand and (b) hand-to-eye configurations.", "texts": [ " Molinari Tosatti are with the Institute of Industrial Technologies and Automation (ITIA) of Italian National Research Council (CNR), via Bassini, 15 - 20133 Milan, Italy federico.vicentini@itia.cnr.it the sensor/tool position in feedback controls, like in visual servoing applications ([3], [4] and op. cit.) and medical robotics ([5], [6], [7], [8]). Remarkably in this latter use, the accuracy of the calibration is a key outcome for the whole control process, namely when a fine positioning of tools w.r.t. a surgical targets is required. The problem display two main configurations, namely eye-in-hand and hand-toeye formulations. The former (Fig. 2-a) considers a camera mounted on the robot end-effector. The latter case (Fig. 2-b) considers the capturing system detached from the robot end effector, e.g. fixed cameras. However, the capturing system does not necessarily involve cameras, like most medical applications. In these cases the eye frame can be retrieved from a cluster of reflective active/passive markers attached to any link (hand) of the robot or attached directly on tools possibly carried by hand, e.g. surgical endoscopes, like in [9]. In the foregoing discussion, the setup involves a passive marker tracking system and remarks about its usage in the hand-eye context are given in V", " In addition, the present work consider a framework of explicit calibration, i.e. the vision system is assumed to be independently calibrated, unlike a remarkable number of recent works that cope with the intrinsic calibration of the camera parameters (in eye-inhand setup) together with the estimate of X and Z (see [10], [11], [12]). Once the tracking system is set in {c} frame, T{ce} is the homogeneous transform displaying the pose of {e} on a robot link, as the robot moves, according to one of the 978-1-61284-456-5/11/$26.00 \u00a92011 IEEE 3327 two conventions in Fig. 2. The same poses are measured in the robot {r} frame, providing paired measures of same poses attained by pairs of frames under calibration. At least 2 different poses are required for solving the problem in the unknowns [2]. Considering N different robot poses, the coordinate frames relationship (see Fig. 3) can therefore be expressed as: AiX = ZBi (1a) AijX = XBij (1b) where X = T{he}, Z = T\u22121{cr}, Ai = T{rh}|i and Bi = T{ce}|i, for any i = 1, . . . , N poses, as well as Aij = T\u22121{rh}|iT{rh}|j , Bij = T\u22121{ce}|iT{ce}|j express {ij} pairs of poses used for obtaining separate solutions for X and Z from the elimination of one of the two unknowns in (1a)" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0000964_j.triboint.2013.07.003-Figure1-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0000964_j.triboint.2013.07.003-Figure1-1.png", "caption": "Fig. 1. Squeeze film configuration between a sphere and a plate lubricated with a non-Newtonian couple stress fluid.", "texts": [ " Analytical solutions are obtained for the film pressure, the load capacity and the squeeze film time. Comparing with the case of the non-inertia Newtonian lubricant by Matthewson [3] and the case of the noninertia non-Newtonian lubricant by Lin [11], the squeeze film characteristics between a sphere and a plate taking into account the fluid inertia forces and the non-Newtonian couple stresses are investigated through the variation of the density parameter D and the non-Newtonian couple stress parameter N. 2. Analysis Fig. 1 shows the non-Newtonian squeeze film configuration between a sphere and a flat plate with a separated film height h. The sphere of radius R is approaching the plate with a squeezing velocity: \u2202h=\u2202t. The lubricant in the film region is modeled as a Stokes couple stress fluid [4]. Assume that the fluid is incompressible and that the body couples and the body forces are negligible. According to the thin film hydrodynamic lubrication theory of Cameron [14] and the micro-continuum theory of Stokes [4], the basic equations for the motion of an incompressible couple stress fluid including the convective inertia terms can be expressed in polar coordinates (r,\u03b8,z) as \u03c1 u \u2202u \u2202r \u00few \u2202u \u2202z \u00bc \u2202p \u2202r \u00fe \u03bc \u22022u \u2202z2 \u03b7 \u22024u \u2202z4 \u00f01\u00de \u2202p \u2202z \u00bc 0 \u00f02\u00de 1 r \u2202\u00f0ru\u00de \u2202r \u00fe \u2202w \u2202z \u00bc 0 \u00f03\u00de where \u03c1 is the fluid density, \u03bc is the fluid viscosity, \u03b7 is the material constant responsible for Stokes couple stress fluids, and u and w are the velocity components in the r and z directions, respectively" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0003762_1.5111940-Figure2-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0003762_1.5111940-Figure2-1.png", "caption": "Fig. 2. A free-body diagram showing the internal forces (excluding ground interactions) on one of the particles in the rim of the wheel. The variables are defined in the text.", "texts": [ " D~ri is the twodimensional displacement vector from particle i to particle i\u00fe 1 (or to particle 0 in the case of i\u00bc n 1); Dr\u0302 i is the unit vector in the same direction. Each rim particle is also connected to the central particle mc by a spring with spring constant kc (c for \u201ccentral\u201d) and equilibrium length lc0 and a linear damper with constant bc. D~rc;i is the two-dimensional displacement vector from the central particle to particle i. We define n\u0302i as the unit vector normal to the wheel. This direction bisects the angle exterior to the wheel between D~ri 1 and D~ri. /i is the angle between D~ri 1 and D~ri, with a positive angle shown in Fig. 1. Figure 2 is a free-body diagram showing all of the internal forces (i.e., gravity and excluding ground interactions) acting on one of the rim particles. ~Fi is the force of mass i\u00fe 1 pulling or pushing on mass i and is 721 Am. J. Phys., Vol. 87, No. 9, September 2019 Robert Knop 721 ~Fi \u00bc k \u00f0Dri l0\u00de \u00fe b \u00f0D _~r i Dr\u0302 i\u00de h i Dr\u0302 i: (1) Notice that the damping only considers the component of the relative velocity of the two particles along the direction of their separation. That is, there will only be damping if the distance between the two point masses is changing", " A small but nontrivial fraction is dissipated in the ground damping springs (5%) and in work that the pressure force does on the wheel (8%). The work from pressure would go into the temperature of the gas in the wheel, but this simulation does not include any thermodynamics. It is interesting that the force doing most of the work is different from the external force providing the most impulse. The overall picture of rolling friction in this work is consistent with modern descriptions such as those that appear in Fig. 2 of Ref. 7, Fig. 1 of Ref. 8, and Fig. 1(b) of Ref. 9. That is, there are three external forces on the wheel. Gravity acts downwards from the center of mass. The normal force acts upwards, but at an effective contact point that is forward (in the direction of motion) of the point on the ground directly below the center of mass. Static friction acts backwards; it is the force that slows down the linear speed the wheel, but its torque would tend to increase the rotation rate. The torque from the offset normal force is large enough to counteract the torque from static friction and to slow down the rotation at the rate necessary to maintain rolling without slipping" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0002176_j.procs.2017.01.179-Figure2-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0002176_j.procs.2017.01.179-Figure2-1.png", "caption": "Fig. 2. Prototype of a tensegrity robot with three rods and nine wires.", "texts": [ " The contributions of this study are as follows: (i) the designing of a basic motion experiment of the prototype with three rods and nine wires, as shown in Fig. 1(a); (ii) proposal and verification of a numerical calculation method of the balancing internal force among wires. The remainder of this paper is organized as follows. Section 2 describes the aspect of the prototype robot and the experimental result of basic motion control. Section 3 demonstrates the calculation method of the balancing internal force. Finally, Section 4 concludes the discussion. Fig. 2 shows the proposed prototype, which consists of three rods and nine wires. Although the lengths of the rods are fixed, the lengths and/or tension of the nine wires can be changed using the nine actuators in the three rods. The power supply is given through electric cables from the external environment. Figure 3 shows the structure of a rod, which is made of an aluminum frame (50 \u00d7 50 \u00d7 600 [mm], 950 [g]). Each rod has three actuators inside to reel three wires, and has six wires entirely: three actuated and three unactuated wires" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0000909_12.880386-Figure2-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0000909_12.880386-Figure2-1.png", "caption": "Figure 2. Experimental setup.", "texts": [ " In order to determine the relationship between electrical parameters and curvature of actuators, three different experiments were carried out on each sample: 1) measurement of resistance of both electrodes in a bent, but electrically inactive state of actuator; 2) recording of voltages along electrodes in its electrically active state (during actuator\u2019s work-cycle) a) while bending of actuator was not restricted; b) while actuator was fixed in its initial state. In electrically inactive state the resistance of bent electrodes was measured using potentiostat PARSTAT 2273 from Princeton Applied Research. In order to guarantee homogeneous curvature for the whole length of the CPC sample, a special rig was used. This mechanism allowed uniform bending curvatures in range of 20 to 100 m-1. The signals during single work-cycle of the actuator were recorded using a National Instruments PCI-6120 DAQ card, driven by PC and LabView, as presented in Fig. 2. The level of the driving voltage \u2013 3V \u2013 was close to the electrochemical window of ionic liquid, and the pulse width \u2013 30s \u2013 was determined from the rate of bending of the particular actuator. Data acquisition frequency was 4 samples per second, while the output signals were generated at the sampling rate of 1000 samples per second. In these measurements CPC actuator was in cantilever configuration, as depicted in Figs. 2 and 3. In order to measure voltage drops along the surface electrodes, additional probes with a lightweight clamp were attached to both electrodes at the free tip of the actuator" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0001428_ajpa.21280-Figure1-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0001428_ajpa.21280-Figure1-1.png", "caption": "Fig. 1. The gibbon brachiating beneath the superstrate (drawn from video footage). Brachiation is split into support and swing phases. The single-limb support phase of the right forelimb corresponds to the swing phase of the left forelimb. The swing phase of the right forelimb corresponds to the single-limb support phase of the left forelimb.", "texts": [ "0 kg) was trained to perform spontaneous brachiation beneath a superstrate of steel pipe (5-m long, 5-cm diameter), which was elevated 1.7 m. We defined a 1.2 3 2 3 1.4 m (wide 3 long 3 high) measurement space in the middle of the experimental apparatus. The fixed coordinate frame was defined as follows: the positive x-axis paralleled the superstrate (i.e., the direction of brachiation), the positive y-axis was upward, and the z-axis was positive to the subject\u2019s right. We defined a brachiation gait cycle as beginning with right hand contact and ending with the next right hand contact (Fig. 1). Each brachiation gait cycle was split into support and swing phases; in this study, the support phase is defined as the duration from right hand contact to release, and the swing phase is from right hand release to the next contact. The swing phase of the right hand corresponded to the single-limb support phase of the left hand. Our subject animal never lowered its swinging-arm below the shoulder in our experimental environment (Fig. 1). This form of brachiation seems different from that of other gibbons illustrated in previous studies (e.g., Jungers and Stern, 1984; Usherwood and Bertram, 2003). However, the animal was normal without orthopedic or pathological problems. We measured the right lower trunk and thigh of the gibbon during continuous-contact brachiation with an infrared detecting 3D motion analyzer (ELITE System; BTS, Milan, Italy) at sampling rate of 100 Hz. Three hemispherical markers, which were made of styrene foam and covered with infrared reflective sheets, were attached on the shaved skin of the gibbon at three sites: the anterior superior iliac spine (ASIS), lateral to the hip (close to the greater trochanter; GT), and the lateral epicondyle of the femur (LE)", " RESULTS AND DISCUSSION Gait cycle and phase definitions We acquired the 3D displacements of the ASIS, GT, and LE of the gibbon during brachiation gait cycles (n 5 12). The mean stride time and stride length were 1.38 s and 1.48 m, respectively. The forward velocity in a brachiation gait cycle was 1.07 6 0.05 m/s (mean 6 SD). In a single gait cycle, the duration of the support and swing phases was 58 and 42% of the total time, respectively. Double-limb support phases existed at the initial and last stage of support phase ( 8% of the brachiation gait cycle, when the trunk and limb motions during brachiation had bilateral symmetry) (see Fig. 1). American Journal of Physical Anthropology The mean displacements of the ASIS, GT, and LE are shown in Figure 3, including the fore-aft displacements (x-axis) and velocities (Fig. 3A,B), vertical displacements (y-axis; Fig. 3C), and lateral displacements (z-axis; Fig. 3D). The coordinates of the right wrist at the beginning of a brachiation gait cycle were defined as the origin, that is, (x, y, z) 5 (0, 0, 0). The displacements of the ASIS and GT represent right lower trunk movement. The displacement of the LE represents distal thigh movement", " Consequently, the lower trunk was translated to the right, and the GT was to the right of the ASIS. Conversely, in the swing phase, the trunk was inclined to the opposite side; in this case, the lower trunk was shifted to the left, and the GT was to the left of the ASIS on the right side of body (Fig. 3D). The right lower trunk moved to the handhold in the latter half of both the support and swing phases through the trunk rotation for the next grip, as the body of the brachiating gibbon rotated 1808 around the vertical axis twice during a brachiating cycle (Fig. 1). The lateral deviation of the right lower trunk from the handhold was smaller during the swing phase than during the support phase (Fig. 3D). One possible reason for this difference is that we measured movements of the right side of the lower trunk; this method of measurement shifted the graphs in Figure 3D upward (i.e., to the right side). The lateral movement of the trunk presumably resulted from the passive change of the CoM position by gravity rather than as a result of the active change in the CoM path" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0003947_intercon.2019.8853567-Figure2-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0003947_intercon.2019.8853567-Figure2-1.png", "caption": "Figure 2. Base of the movement in the \u201cZ\u201d axis. Source: Self made", "texts": [ " Travel axis X = 400mm, axis Y = 400mm, axis Z = 400mm. The maximum weight of the head to be displaced is Wz = 26.00 N; important parameter to perform the design. b) Design of the horizontal carriage for the \"Z\" axis: When the load is applied in the same direction of the trip, the resulting load can be separated to calculate the loads in each block of the guides. This type of configuration is usually found in vertical applications. It is important to consider the orientation of the load with respect to the guidance system. The Figure 2 shows the mechanical design of the movement of the Z axis. For the design of this part, Aluminum 6061 was chosen, which a factor of safety N = 2. la Fig. 2 . = (1) Where: : Maximum Stress. N: Factor of safety. : Maximum stress of Aluminum=55MPa. Then, = 55 2 = 27.5 The maximum stress that the carriage Z is going to be exposed was determined. \u2032 = 0.2576 . It is verified that the calculated stress is less than that supported by the material. < 0.2576 < 27.5 c) Design of the horizontal carriage for the \"Y\" axis: The normal force is applied to a horizontal system of arrows or guides. The need of the system is to move linearly and precisely in the two Cartesian axes, for which reason the use of the mechanism shown was chosen" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0003039_s38314-017-0062-x-Figure1-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0003039_s38314-017-0062-x-Figure1-1.png", "caption": "FIGURE 1 The friction coefficient-dependent size of the Kamm friction circle defines the forces which can be transferred between the tires and the road surface (\u00a9 Continental)", "texts": [ " By recognising road conditions and classifying friction coefficients on this basis, it is now possible to adjust the driving strategy to the conditions which are actually prevalent and therefore to expand the safety functions of Advanced Driver Assistance Systems and Automated Driving. System developers have been searching for some decades for a solution that can detect the actual friction coefficient of a road surface as far in advance as possible in order to protect the driver against losing control of the vehicle. For example, the friction circle is so small on roads in winter that the physical boundaries of the vehicle are reached significantly faster than on a dry, grippy road surface, FIGURE 1. As the friction coefficient determines the maximum force between the tires and the road surface, this information is highly important for safe vehicle \u00a9 Continental AUTHORS Bernd Hartmann is Head of the Enhanced ADAS & Tire Interactions Project Group in the Advanced Technology Department within Continental\u2019s Chassis & Safety Division in Frankfurt (Germany). Alfred Eckert is Director of the Advanced Technology Department within Continental\u2019s Chassis & Safety Division in Frankfurt (Germany). Sensor Technology 34 DEVELOPMENT SENSOR TECHNOLOGY handling and also for Advanced Driver Assistance Systems (ADAS)" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0000974_amr.712-715.709-Figure3-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0000974_amr.712-715.709-Figure3-1.png", "caption": "Fig. 3 Laser cladding device positioned on a crankshaft", "texts": [ " The longitudinal feed of the laser nozzle 10 to-wards the crankshaft bearing surface 1 is achieved by means of transmission between the spur gear 11 and the feeding rod - toothed rack 66. The spur gear 11 is fixed on the shaft of a second control motor 8B. The positioning angle or pivoting angle X of the laser nozzle 10 towards the crankpin 1 surface is ensured by the first control motor 8A. The laser nozzle 10 and the first control motor 8A are connected through a bush 9. The laser nozzle 10 is fitted rotatably in carriage 7A by means of a pin 12. The control motors 8A and 8B are secured to the carriages 7A and 7B by means of four fixation screws 13 (see Fig. 3). During the build-up operation, the crankshaft it-self is rotated in the engine by built-in means, convention-ally the service-electrical motor. This motion is aligned with the movements controlled by the control motors 8A and 8B. The laser build-up process includes a step where-by cladding powder or any other cladding material is applied to the damaged surface of the journal. At the same time, the cladding powder is irradiated by the high-energy laser beam. Thus a metallurgical bond between the crank-pin surface and substrate material is achieved by melting both the cladding material and substrate" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0002289_j.robot.2017.04.002-Figure1-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0002289_j.robot.2017.04.002-Figure1-1.png", "caption": "Fig. 1. Front- (a) and rear- (b) wheel driven cases of a kiwi drive mobile robot platform, where Vwheel is the constrained driving velocity of the wheels and V robot is the linear velocity vector. Due to robot dynamics, maximum acceleration of the robot is higher in case (b).", "texts": [ " The model was applied to control a kiwi drive mobile robot and validated by experimental measurements. An open-source Robot Operating System (ROS) catkin C++ package was published to enable the feasible implementation of the results. \u00a9 2017 Elsevier B.V. All rights reserved. The first three-wheeled holonomic platform appeared in 1994, developed by Stephen Killough and Francois G. Pin [1]. Killough\u2019s design used three pairs of wheels mounted in cages, orthogonal to each other, and thereby achieved holonomic movement without using true omniwheels. Fig. 1 shows the platform employing three omnidirectional wheels (instead of actuated caster wheels [2,3]) in a triangular formation, which is generally called kiwi drive. After Killough\u2019s patented invention, several researches focused on the modeling and control of these holonomic platforms. The kinematics were first described by Killough, and the dynamics are presented in many different forms. The path tracking for these robots is developed using differentmethods: Vazquez andVelascoVilla derived the computed-torque control for omnidirectional * Corresponding author", " Mecanum wheel based forklifts or security robots. Regarding results in general for the generation and tracking of time optimal trajectory, mostmethods apply constant velocity and acceleration limits to the robot body. Only a few of the methods deal with the possible maximum velocity and acceleration of the robot, which is rapidly changing duringmotion. Surprisingly, there are caseswhen the robot can go faster in a direction if it rotates during linearmovement [11]. As an illustrative example, Fig. 1 shows a front- and a rear-wheel driven case, where the possible maximum acceleration (and deceleration) difference is often twofold. http://dx.doi.org/10.1016/j.robot.2017.04.002 0921-8890/\u00a9 2017 Elsevier B.V. All rights reserved. In 2002, Robert L. Williams II, Brian E. Carter, Paolo Gallina and Giulio Rosati [12] proposed the extensive modeling of holonomic movement considering wheel-slip. However, their work mostly focused on the slip caused by a single row omniwheel, which is designed for material handling industrial applications instead of mobile robots. They included the latter friction case for handling the non-continuous rolling surface of the single wheel. They assumed that the robot weight is equally distributed on each wheel, which neglects the effect (showed by Fig. 1) when the robot operation is similar to a front- or rear-wheel driven vehicle. Andr\u00e9 Scolari Conceicao, A. Paulo Moreira, Paulo J. Costa presented a method in 2006 for time optimal velocity control that considers the maximum wheel speeds [13], called ideal reference velocities (IRV). This IRVmethod can be adapted andmay work for the dynamics too, but it causes too many different equations for different wheel arrangements. The solution can be very difficult, especially when the robot is over-actuated by four ormorewheels", " Finally, Section 5 summarizes the conclusions. The following conventions are used in the equations for marking scalars, vectors and their absolute value: Vectors are underlined (if not, that represents their absolute value). If the label of a vector is not underlined but has subscript, it represents one component of the vector. General time-optimal theories in the field of ground vehicles and mobile robots often use the friction circle model to model wheel-slip [23]. But as the simple illustration of Fig. 1 shows, the direction-independent friction limits are not appropriate to avoid the wheel-slip of omniwheels. Fig. 2 shows the calculated and verified acceleration phase space of a real kiwi drive platform with transparent body. The solid body inside shows the possible maximum sized direction-independent assumption of the allowed acceleration vector set. As the volumetric difference shows, many valuable capabilities will be lost if the direction-independent assumption is used. Therefore, vector sets of allowed velocities and allowed accelerations have to be defined in a direction-dependent form in the field of all omniwheel based holonomic drives", " Leng, Q. Cao, Y. Huang, A motion planning method for omni- directional mobile robot based on the anisotropic characteristics, Int. J. Adv. Robot. Syst. 5 (4) (2008) 327\u2013340. [18] Y. Liu, J.J. Zhu, R.L. Williams, J. Wu, Omni-directional mobile robot controller based on trajectory linearization, Robot. Auton. Syst. 56 (5) (2008) 461\u2013479. [19] M. Quigley, K. Conley, B. Gerkey, J. Faust, T. Foote, J. Leibs, E. Berger, R.Wheeler, A. Mg, ROS: an open-source Robot Operating System, Icra 3 (2009) 5. no. Fig. 1. [20] J. Blanco, Development of Scientific Applications with the Mobile Robot Programming Toolkit, 2008. [21] R. Li, M.a. Oskoei, H. Hu, Towards ROS based multi-robot architecture for ambient assisted living, in: 2013 IEEE Int. Conf. Syst. Man, Cybern, 2013, pp. 3458\u20133463. [22] J. Minguez, L. Montano, Nearness diagram (ND) navigation: Collision avoidance in troublesome scenarios, IEEE Trans. Robot. Autom. 20 (1) (2004) 45\u201359. [23] T. Lipp, S. Boyd, Minimum-time speed optimisation over a fixed path, Int" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0003257_1.5090680-Figure1-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0003257_1.5090680-Figure1-1.png", "caption": "FIGURE 1. The coordinate diagram of manipulator", "texts": [ " Double-arm rock drill rig is the key equipment for underground mining and tunnel excavation. It mainly completes the work through the automatic positioning and drilling of the manipulator, and it has four revolute joints and one prismatic joint. In order to analyze the transformation relationship between the base coordinate system and the endeffector coordinate system of the manipulator, the position relations of the joints of the manipulator are expressed by the simple graphical form of the coordinate system. As shown in Fig. 1. The kinematics model of the manipulator is established according to the MDH method, and the homogeneous transformation between the adjacent two-link coordinate systems is as shown in equation (1). 1 1 1 1 1( , , , , ) ( , ) ( , ) ( , ) ( , ) ( , )i i i i i i i i i i i iT a d Trans x a Rot x Trans z d Rot z Rot y 1 1 1 1 1 1 11 1 1 1 1 1 1 0 0 0 1 i i i i i i i i i i i i i i i i i i i ii i i i i i i i i i i i i i i i c c s c s a s c c s s c c s c s s c d s T s s c c s c s s s s c c d c (1) Here, 1i iT is the forward kinematic expression of the connecting rod i to the connecting rod 1i , c means cos , s means sin , ia is the length of the connecting rod i , id is the offset of the connecting rod i , i is the torsion angle of the connecting rod i , i is the rotation angle between the adjacent two joint axes 1iZ and iZ on the plane i i iX O Z , and i is the joint angle of the connecting rod i ,as shown in Fig" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0000499_6.2013-1701-Figure1-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0000499_6.2013-1701-Figure1-1.png", "caption": "Figure 1. Representation of the different analysis approaches used in this paper.", "texts": [ " Static analyses on isotropic30 and composite structures31 revealed the strength of LE models in dealing with open cross-sections, arbitrary boundary conditions and obtaining Layer-Wise descriptions of the 1D model. In this paper some new improvements have been introduced with respect the past works in order to address the problem of complex structures. A three-dimensional formulation based on the Carrera unified formulation is presented. Two different approaches are used in the analysis of aircraft structures: a singular non-uniform cross-section beam or multiple beam with non-uniform cross-section. In fig.1a is depicted the first approach that involved a single beam with a non-uniform cross section. In fig.1b is shown the second method, different beams are used to describe the different parts of the aircraft, they must be rotate and can be merged during the matrix assembling phase in different way: the first is to introduce new relation between the structural nodes, the second is to use two- or three- dimensional elements as node where the different beam are connected, a three-dimensional model is presented in terms of unified formulation. Both the method are employed and compared. 2 of 16 American Institute of Aeronautics and Astronautics D ow nl oa de d by U N IV E R SI T Y O F M IN N E SO T A o n Se pt em be r 21 , 2 01 3 | h ttp :// ar c" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0001609_0954406211427833-Figure3-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0001609_0954406211427833-Figure3-1.png", "caption": "Figure 3. Coordinate systems for simulation: reference frame, cutter home position, cutter Initial position and gear blank initial position.", "texts": [ "10 The solid models of cutter and work blanks are created. Cutter is at University of Ulster Library on May 14, 2015pic.sagepub.comDownloaded from modelled as a solid of revolution of the rack-cutter profile. The position and orientation of cutter solid and work blank are changed incrementally, maintaining the kinematic relationship, and the interfered material between the cutter and work blank is subtracted from the work solid by Boolean operation. As a result of progressive subtraction, the required tooth surface is generated on the work blank. Figure 3 shows the coordinate frames for CAD simulation. So:(Xo,Yo, Zo) is the common or reference frame. Sh:(Xh,Yh, Zh) is the cutter frame and a1, b1, c1 the cutter home position with respect to the reference frame. The rotation parameters are 1, \u20191, 1. Similarly, Sc:(Xc,Yc, Zc) is the frame for the instantaneous position of the cutter for which translation parameters are a2,b2, c2 and rotation parameters 2, \u20192, 2 from the reference frame coordinate. Sw:(Xw,Yw, Zw) is the work blank frame and a3, b3, c3the initial work blank position from the reference frame", " To start with, the cutter and the work blank solids are created at respective home positions, Oh and Ow. The specification of spherical gear pair considered in this study is presented in Table 1. The working of the spherical gears can be visualized with reference to Figure 1. During meshing of gear pair, the gears can rotate within 24 to 24 about the Y- and Z-axes independently, which are working axes in the assembled condition. The combination of rotation about these axes is also possible, as shown in the assembly. These working axes should not be confused with the notations for axes in Figure 3 used for CAD simulation. The simulation of the generation machining process is done according to the scheme presented in Figure 4. The different simulation parameters are presented in Table 2. The home position of the cutter is taken at 149.65mm. When the cutter touches the blank, its current position is 219.99mm. A total depth of 10.66mm (ha\u00fe hf\u00bc 5\u00fe 5.66mm from Table 1) is realized, when the cutter plunges to 230.65mm. In order to decide the oscillation angle from the centre position, it is determined from the following equation based on meshing condition" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0000437_s11223-013-9461-2-Figure3-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0000437_s11223-013-9461-2-Figure3-1.png", "caption": "Fig. 3. Skeleton curves of rigid rotor vibrations. (The line number corresponds to the NNM number.)", "texts": [ " Figure 2 presents the NNM dependence of the non-dimensional displacement of the shaft support A towards the x axis away from the independent phase coordinates p and q, which is determined from the first formula (11). Here, p is the displacement of the shaft support B towards the x axis. This surface of double curvature clearly demonstrates the degree of the system nonlinearity. In order to study the skeleton curves of rotor vibrations, one of the equations from (15), which describes the motion relatively to NNM, is solved using the describing function method [13]. Figure 3 shows the skeleton curves for all modes of vibrations of a rigid rotor. During vibrations in the first mode rotor makes mainly the longitudinal displacements. Since the axial stiffness of the ball bearing is smaller than the radial stiffness, the frequency of longitudinal vibrations is the lowest. The second and third frequencies are multiple, they are in consistence with rotor vibrations in two perpendicular planes, and moreover, both bearings undergo deformation in the same direction. The fourth and the fifth frequencies are consistent with rotor vibrations, under which support sections move in the opposite directions" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0003741_012077-Figure1-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0003741_012077-Figure1-1.png", "caption": "Figure 1. The first folding mechanism in: (a) compact configuration; (b) partially deployed configuration; (c) totally deployed configuration.", "texts": [ " The main goal of the research on bar mechanisms based structures is to find solutions of mechanism that can allow the structure to be folded into a compact bundle but also to get a maximum expanded configuration, with high rigidity, too. To do that, the first way is to find optimum structure of the mechanism, which must need a minimum number of actuators to unfold/fold it, and, then, the next step is to solve the dimensional synthesis problem. A first planar mechanism (figures 1) discussed here was proposed by Ten Fold Engineering in building different deployable structures. It is an eight bar mechanism that can have a very compact folded configuration (figure 1(a)) and, also, may be locked in a deployed configuration (see figure 1(c)). To find the degree of freedom of the mechanism, Grubler formula may be used, equation (1): 1213 gnF (1) where n is the number of links (including the frame) and g1 is the number of single revolute joints. According to this formula, for n = 8 and g1 = 10 (as the mechanism shown in figure 1), we will get: 1102183 F (2) It means that a single actuator is needed to actuate the mechanism. This actuator is suitable to be placed to A joint (the link 1 will be the driver link). Modern Technologies in Industrial Engineering VII, (ModTech2019) IOP Conf. Series: Materials Science and Engineering 591 (2019) 012077 IOP Publishing doi:10.1088/1757-899X/591/1/012077 If we consider 1 the rotational angle of the driver link, the area A1 of a wall formed by this mechanism in the deployed configuration will be equation (3): 1 2 1 sin2 AB lA (3) considering that ADAB ll ", "1088/1757-899X/591/1/012077 The first simulations have been done, considering the coordinates of the mechanism nodes in the fully extended configuration as illustrated in figure 5. We have to note that no dimensional synthesis of this mechanism has been realized till now. As consequence, the mechanism could not fold in the optimum compact configuration. The driving link 1 could only rotate with an angle 321 degree (see figure 6), starting from the extended configuration. It means that the simulated mechanism could only fold till the partially deployed configuration represented in figure 1(b). Also, the trajectory of the G and H nodes are not yet straight lines, as they should be, but as they are presented in figure 7, even if the link 7 has not a significant rotation around z axes, as seen in figure 8. The paths of the A, G and H nodes are shown in figure 9. Modern Technologies in Industrial Engineering VII, (ModTech2019) IOP Conf. Series: Materials Science and Engineering 591 (2019) 012077 IOP Publishing doi:10.1088/1757-899X/591/1/012077 Modern Technologies in Industrial Engineering VII, (ModTech2019) IOP Conf" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0003268_s12206-019-0107-6-Figure2-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0003268_s12206-019-0107-6-Figure2-1.png", "caption": "Fig. 2. Schematic diagram of rolling circle radius change of harmonic wear wheel.", "texts": [ " Wheel harmonic wear refers to the wheel profile wear formed one or more continuous out-of-round of harmonic forms [6], which can be expressed as a uniform wear wave depth of non-circular wear, hypothesis harmonic wear wheels rotating at angular velocity withw , expressing the out-ofround deviation of wheel rolling circle with RD as shown in Fig. 1, the cosine function can be described: 21 cos 0 2 20 . A t t nT TR nT t \u00ec \u00e6 \u00f6- \u00a3 \u00a3\u00ef \u00e7 \u00f7\u00ef \u00e8 \u00f8D = \u00ed \u00ef < \u00a3\u00ef\u00ee p p w (1) And, t R T L f = (2) where A is out-of-roundness wave depth; t is wheel rolling time;T is a harmonic wear period; n is harmonic order;w is rolling angular velocity; L is out-of-roundness wave length; f is phase angle. The rolling circle radius of harmonic wear wheel varies with the rolling angle ( )tf as shown in Fig. 2, which can be defined: 0( ) ( ) ( ) ( ) R R R t t t t f f f f w = + D\u00ec \u00ed + D = + D\u00ee (3) where 0R is nominal rolling circle radius of wheel;f is phase angle;w is rolling angular velocity; tD is time interval. Due to the limitations of field experiments, it is usually only at the position of the nominal rolling circle of the wheel profile that the out-of-roundness of wheel is tested [10]. However, the wheel profile of high-speed train caused by harmonic wear will change in both lateral and radial macroscopic dimensions, so the wheel instantaneous rolling circle deviation parameter is introduced as RD , and choosing the five special moments described in Fig" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0000442_s12555-012-0219-6-Figure2-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0000442_s12555-012-0219-6-Figure2-1.png", "caption": "Fig. 2. Structure of neural network.", "texts": [ " Structured of proposed multilayer back propagation neural network controller design The inputs of the NN are the desired system states, its derivatives, and the plant output. The multilayer back propagation NN is used in the present work. The multilayer back propagation NN is especially useful for this purpose, because of its inherent nonlinear mapping capabilities, which can deal effectively with real-time online computer control. The NN of the present work method has three layers: an input layer with n neurons, a hidden layer with n neurons and an output layer with one neuron as shown Fig. 2. Let xi be the input to the ith node in the input layer, zj be the input to the jth node in the hidden layer, y be the input to the node in the output layers. Furthermore Vij be the weight between the input layer and hidden layer Wj1 is the weight between the hidden layer and the output layer. 3.2. Learning of neural network The relations between inputs and output of NN is expressed as 1 , n inj oj i ij i Z V xV \u2212 = = +\u2211 (29) 01 11 , P ink j jj Y W z W \u2212 = = +\u2211 (30) _ ( ),j injZ F Z= (31) ( ), k ink Y F Y \u2212 = (32) where F(" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0000418_detc2011-47599-Figure6-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0000418_detc2011-47599-Figure6-1.png", "caption": "FIGURE 6. Geometric parameters of the planar two-3R cable-driven closed chain", "texts": [ " Hence the reciprocal screws have to be determined either analytically [18, 19] or numerically using the Gauss-Seidel elimination approach [20]. 4. For force-closure, a minimum of seven cables are required to form a convex polytope containing the origin. 5. W now represents a 6-D space of forces and moments, and force-closure analysis can be carried out by representing the solution for T = A#.W+N(A).\u03bb as linear inequalities. This section presents the case study examples for forceclosure analysis of a planar two-3R cable-driven closed chain with various cable routings. Figure 6 presents the geometric parameters and the external wrenches acting at the centre of each link. These wrenches are the weight of the individual links, i.e., 1 kgf. Figure 7 presents the planar two-3R cable-driven closed chain with various sets of active joints and cable routing configurations. Base Base X Y Chain 1 Chain 2 All dimensions are in millimeters. P = {xp, yp, } 100 100 50 50 50 50 50 50 50 50 5050 50 50 50 7575 1 kgf 1 kgf 1 kgf 1 kgf 1 kgf O Base Base X Y C ab le 1 Cable 2 Ca bl e 3 C able 4 Chain 1 Chain 2 Configuration D Chain 1 Chain 2 Configuration C Base Base X Y C ab le 1 Cable 2 Ca bl e 3 C able 4 FIGURE 7" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0000408_rm2012v067n02abeh004786-Figure10-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0000408_rm2012v067n02abeh004786-Figure10-1.png", "caption": "Figure 10. \u2018Adjacency with Whitney umbrella\u2019 singularity.", "texts": [], "surrounding_texts": [ "Here we mainly follow the papers [25]\u2013[28] in describing generic singularities of the boundary of the set of local transitivity of control systems on a plane and in three-dimensional space. The basic types of control systems are examined, namely, the cases where the indicatrices I(x) \u2282 TxM are: 260 A.A. Davydov and V.M. Zakalyukin \u2022 smooth curves or surfaces which are embedded (or immersed) in the linear spaces TxM of tangent vectors at points x \u2208 M ; \u2022 or smooth curves or surfaces with boundaries and corners embedded in the spaces TxM ; \u2022 or images of smooth maps of general form of a smooth closed manifold U with dimension at least two; \u2022 or, finally, images of smooth maps of a manifold U with boundary and corners. We recall that a corner on a manifold is a subset of it which is diffeomorphic to a quadrant of the arithmetic space of the corresponding dimension. The fact that the manifold U of controls has boundaries or corners can be interpreted as a simulation of constraints imposed on the control in the form of inequalities. This kind of control system is widely used in applications. Let us consider a point x of the manifold Mn as a parameter and fix some trivialization of the tangent bundle of M . We obtain a family of submanifolds I(x) embedded in Rn. Accordingly, by the boundary \u03a3 of the set of local controllability we shall mean the set of parameters x for which 0 \u2208 Rn belongs to the boundary of the convex hull of the corresponding submanifold I(x). We recall that the convex hull of a (closed) subset of an affine space is the intersection of all closed half-spaces containing this set. Note also that the boundaries of the convex hulls of smooth submanifolds in general position can have singularities. Their classification, which is closely related to the singularities of the boundaries of the zone of local transitivity, is also discussed below. 3.1. Formulation of results. We start with the simplest case of a control system on a two-dimensional manifold M . In the tangent plane at each point of the manifold let there be specified a set I(x) of admissible velocities which is a smooth closed curve. We recall [5] that the boundary of the convex hull of a generic smooth curve on the plane at any point is either the germ of a smooth curve (a convex curve or a straight line) or the germ of a curve which can be transformed via some diffeomorphism of the plane into the germ of the graph of the function y = f(x) at the origin, where f(x) = { 0 for x 6 0, x2 for x > 0. If the curve has endpoints, then the list mentioned above should also include the singularity that corresponds to the graph of the function g(x) = |x|. For a generic two-parameter family of smooth curves, the boundary \u03a3 of the set of local transitivity can have only one more singularity distinct from the singularities of the convex hull of an individual curve in general position [25], namely, the indicated singularity of the graph of the function y = |x| at the origin. It turns out that this list coincides exactly with the list of generic local singularities (up to diffeomorphisms of the plane) of the boundary of the zone of local transitivity for control systems on the plane whose indicatrices are smooth curves with endpoints. Moreover, it is the same as the list of singularities of convex hulls and boundaries of the zones of local transitivity in the case where the indicatrices are the images Controllability of non-linear systems 261 of a manifold U of arbitrary dimension under smooth maps which form families in general position. This being the case, it can be assumed that the manifold U has boundaries and corners. The techniques of the proofs of these (and some other) facts are discussed below. Let us turn to the main case of a three-dimensional manifold M . First we shall assume that the indicatrix is a C\u221e-smooth closed space curve. The list of generic local singularities of the convex hull of a curve in R3 up to diffeomorphisms of the ambient space was obtained in [29] and [6]. It includes the first six of the normal forms presented below in Theorem 1. In order to describe the singularities of the boundaries of the convex hulls of space curves with endpoints we introduce the following notation. A closed convex simple piecewise smooth curve \u03b3 that lies in the plane z = 1 of the space R3 = {(x, y, z)} is called a simple pivot curve if it consists of alternating segments of straight lines and strongly convex arcs and is C1-smooth at their common points. A curve \u03b3 is called a pivot curve with k corners if it contains the sides of k angles less than \u03c0 connected by convex pieces consisting of straight segments and arcs and is C1-smooth at all common endpoints except the vertices of the angles. Some sides may be common for adjacent angles. If we replace a straight-line side of an angle by a convex curve arc in this definition, then we obtain the definition of a pivot curve with k corners and a curved side. The conic surface formed by the segments that join the origin with the points of a pivot curve which is either simple (k = 0) or has k = 1, 2, . . . corners (possibly, with curved sides) will be called a k-cone. We note that the germs of these cones at the origin have functional invariants with respect to the action of diffeomorphisms. Indeed, the tangent vectors at the apex form the tangent cone and sweep out a convex curve on the sphere of directions. The tangent cone is subject only to linear transformations under the action of diffeomorphisms. Theorem 1 [27]. The germ at an arbitrary point of the convex hull boundary of a generic connected space curve with endpoints can be reduced by an appropriate diffeomorphism of R3 to the germ at the origin of the graph of one of the following functions z = fi(x, y), where i = 1, . . . , 7 (Figs. 2\u20135): 262 A.A. Davydov and V.M. Zakalyukin 1) f1 = 0 (the germ of a smooth surface); 2) f2(x, y) = |x| (edge); 3) f3(x, y) = { 0 for x 6 0, x2 for x > 0 (adjacency); 4) f4(x, y) = x2 for y 6 x, x > 0, y2 for y > 0, y > x, 0 for y 6 0, x 6 0 (bow); 5) f5(x, y) = 0 for y 6 0, x 6 0, x2 for y 6 \u2212x, x > 0, y2 for y > 0, y 6 \u2212x, 1 2 (x2 + y2)\u2212 y \u2212 x for x + y > 0 (stern); 6) f6(x, y) = minz\u2208R{z4 + xz2 + yz} (truncated swallowtail); 7) f7(x, y) = y2 + x for x > 0, y2 for y 6 0, x 6 0, (1\u2212 x)y2 for y > 0, x 6 0 (bend); or to the germ at the origin of a k-cone with k = 0, 1, 2. Remark 2. Generic singularities of the convex hull of a closed space curve (without endpoints) belong to the first six classes mentioned in this theorem. Remark 3. The smooth surface 1) may be strictly convex, ruled (Fig. 2, point B), or flat (Fig. 2, point A). The \u2018edge\u2019 2) arises at a generic point of the initial curve itself (Fig. 2, point C). The singularity 3) appears, in particular, at the points of adjacency of a ruled surface and a flat one (Fig. 2, point D). The germs 4) and 5) correspond to vertices E and F of the flat triangles which inevitably appear on the boundary of the convex hull of the curve. The cone apex coincides with the endpoint (Fig. 5, point K). The truncated swallowtail appears at a point of the curve such that the tangent line at this point also intersects the curve at some other point (Fig. 3, point J). Controllability of non-linear systems 263 The list of generic local singularities of the boundary of the transitivity zone for a control system on a three-dimensional manifold whose indicatrix is a closed space curve is given by the following theorem. Theorem 2. For a generic family of smooth curves rm : S1 \u2192 R3 depending on a three-dimensional parameter m = (x, y, z) \u2208 R3 the list of all local singularities on the boundary \u03a3 of the transitivity zone (considered up to diffeomorphisms of R3) consists of the germs 1)\u20137) mentioned in Theorem 1 and the germs 9), 12), 14) described in Theorem 3 below. For a space curve with endpoints the list of generic singularities of the transitivity zone is as follows. Theorem 3. For a generic three-parameter family of connected space curves with endpoints, the germ at any point of the boundary \u03a3 of the transitivity zone can be reduced by an appropriate diffeomorphism either to one of the germs 1)\u20137) of the boundaries of the convex hulls of generic curves in R3 listed in Theorem 1, or to the germ at the origin of the graph of one of the following functions : 8) f8(x, y) = 0 for x 6 0, y 6 0, y2 for y2 > x, y > 0, x for y 6 \u221a x , x > 0 (cut); 9) f9(x, y) = 0 for y 6 0, x + \u03b1y 6 0, (x + \u03b1y)2 1 + \u03b12 for y 6 \u03b1x, x + \u03b1y > 0, \u03b1 \u0338= 0, y2 for y > 0, x 6 0, x2 + y2 for x > 0, y > \u03b1x (adjacency of four surfaces, a cup); 10) f10(x, y) = \u2212y for y > 0, 0 for y 6 0, x > y, y(x\u2212 y)2 for y 6 0, x 6 y (helmet); 11) f11(x, y) = { |x| for y > 0, |x|+ y2 for y < 0 (butterfly); 264 A.A. Davydov and V.M. Zakalyukin or to the germ at the origin of the union of three surfaces with boundaries determined by the conditions 12) z = 0, y 6 0, y = x2, z 6 \u22124x2, z = \u2212t2, y = 1 4 z + tx, 0 6 t 2x 6 1 (adjacency with Whitney umbrella); or to the germ at the origin of the union of two surfaces with common boundary given by the conditions (book) 13) z = 0 for y, x > 0 or for x 6 0, y > 1 4 x2, z = 2t3 + xt2 for 3t2 + 2tx + y = 0, t > max { 0,\u2212x 2 } ; or to the germ at the apex of the lateral surface of a pyramid with n ridges for n = 3, 4, 5, that is, to the germ of the boundary of the domain given by the inequalities 14) x, y, z > 0 for n = 3, x, y, z > 0, z \u2212 x + y > 0 for n = 4, x, y, z > 0, z \u2212 x + y > 0, z \u2212 \u03b1x\u2212 \u03b2y > 0 for n = 5 (for n = 5 the normal form has two scalar invariants \u03b1 and \u03b2); or, finally, to the germ of a k-cone with k = 0, 1, 2. Controllability of non-linear systems 265 Remark 4. The boundaries a and b of the surfaces I and II of the germ 12) are given by the equations z = 0, y = 0 and y = x2, z = \u22124x2, respectively. These curves are tangent to each other at the origin, which is their unique common point. The surface III is a piece of a Whitney umbrella bounded by these curves and tangent to the surfaces I and II along a and b (Fig. 9). The boundary H(Ix) of the convex hull of a generic C\u221e-smooth surface Ix may have singularities which are only C1-smooth. Generic local singularities of H(Ix) [5] are either germs of a smooth surface or germs of a C1-smooth surface of \u2018adjacency\u2019, \u2018cup\u2019, or \u2018truncated swallowtail\u2019 type. For surfaces with boundary and corners, the following assertions hold. Theorem 4. The list of generic local singularities of the convex hull boundary of a smooth compact surface in R3 with a smooth boundary consists, up to diffeomorphisms, of the singularities 1)\u20137), 9), and 11). Theorem 5. Generic local singularities of convex hull boundaries for smooth compact surfaces in R3 with boundaries and corners are either the singularities 1)\u20137), 9), 10), or k-cones with arbitrary k. Theorem 6. The list of local singularities (up to diffeomorphisms of the parameter space) of the boundary of the set of local transitivity for a generic three-parameter family of smooth compact surfaces in R3 consists of the germs 1)\u20136), 9), 12), 14) of normal forms. Theorem 7. The list of local singularities (up to diffeomorphisms of the parameter space) of the boundary of the set of local transitivity for a three-parameter generic family of smooth compact surfaces with a smooth boundary consists of the singularities 1)\u201314). Theorem 8. The list of local singularities (up to diffeomorphisms of the parameter space) of the boundary of the zone of local transitivity for a generic three-parameter family of smooth compact surfaces with boundary and corners consists of all the singularities 1)\u201314) and k-cones with arbitrary k = 1, 2, . . . . Finally, let us consider the case where the indicatrices in three-dimensional space are the images of smooth maps of a given compact n-dimensional manifold Un of controls. 266 A.A. Davydov and V.M. Zakalyukin Theorem 9. The list of local singularities (up to diffeomorphisms of the target space) of the convex hull boundary of the image of a generic smooth map V : Un \u2192 R3 of a smooth n-dimensional compact manifold Un coincides with the list of local singularities of the convex hull boundary of a compact space curve for n = 1, or with the list of local singularities of the convex hull boundary of a smooth compact surface for n > 2, respectively. For n = 2 the list of local singularities of the convex hull boundary of a smooth closed surface also contains the \u2018adjacency with Whitney umbrella\u2019 singularity 12) described in Theorem 3. Theorem 10. The list of local singularities (up to diffeomorphisms of the parameter space) of the boundary of the zone of local transitivity for a generic threeparameter family of smooth maps V : Un \u2192 R3 of a smooth compact manifold Un of dimension n > 3 coincides with the list of germs 1)\u20136), 9), 12), and 14) of normal forms in Theorem 6. Theorem 11. The list of local singularities (up to diffeomorphisms of the target space) of the convex hull boundary of the image of a generic map V : Un \u2192 R3 of a smooth n-dimensional compact manifold Un with boundary and corners coincides with the list of local singularities of the convex hull boundary of a compact space curve with endpoints for n = 1, or with the list of local singularities of the convex hull boundary of a smooth compact surface with boundary and corners for n > 2, respectively. For n = 2 the list also contains the \u2018adjacency with Whitney umbrella\u2019 singularity 12) described in Theorem 3. Theorem 12. The list of local singularities (up to diffeomorphisms of the parameter space) of the boundary of the zone of local transitivity for a generic threeparameter family of smooth maps V : Un \u2192 R3 of a smooth n-dimensional compact manifold Un with boundary and corners coincides with the list of generic local singularities of the convex hull boundary of a smooth compact space curve with endpoints for n = 1, or with the list of generic local singularities of the convex hull boundary of a smooth compact surface with boundary and corners for n > 2, respectively. Remark 5. The lists described above show that generically the boundary of the convex hull and the boundary of the zone of local transitivity are Lipschitz. Apparently, this is true in any dimension. Remark 6. Moreover, in all cases we have the following statement, which provides an example of the principle [19] that \u2018good cases\u2019 dominate in many control-theory constructions, in contrast to Arnold\u2019s principle of \u2018fragility of the good\u2019, which is typical in singularity theory [30]. Consider a germ K of the generic transitivity zone whose base point corresponds to the origin being in a C1-smooth germ of the convex hull. Then either the boundary \u03a3 is smooth, or K is on the \u2018larger\u2019 side of the germ of the complement of \u03a3, which cannot be embedded in a half-space in any smooth local coordinates. Remark 7. The boundaries of the convex hulls of curves and surfaces contain ruled and flat pieces. These features are lost under a diffeomorphism: different arrangements of flat domains and straight lines on ruled pieces may correspond to diffeomorphic singularities. The statements of Theorems 1\u201312 contain only this rough Controllability of non-linear systems 267 classification up to diffeomorphisms. However, all possible cases of the affine structure of singularities can be found by the methods described below. 3.2. Legendre transformations and support hyperplanes. In this section we present the basic constructions and propositions ([25], [27]) which are involved in the proofs of the theorems formulated in the previous section and which may prove to be useful in the study of singularities in many other problems. First, let us note that if the indicatrix is given as the image of a map Rn \u2192 Rm with n > m, m = 2, 3, then, according to Thom\u2019s and Mather\u2019s classical results on the classification of singularities of generic maps (see, for instance, [17]), in the generic case only points of smoothness of the visible contour can appear on the boundary of the convex hull of the indicatrix. The only exception is the case of the \u2018Whitney umbrella\u2019 singularity R2 \u2192 R3. In all other cases it suffices to consider the indicatrices which are embedded curves and surfaces (possibly, with boundaries and corners), and their families. Surfaces with boundaries and corners, as well as curves with endpoints, can be treated uniformly as particular cases of stratified submanifolds embedded in R3. A collection J = {I1, . . . , Is} of closed embedded submanifolds I kj j (strata) of dimensions kj will be called a stratified submanifold if it contains a unique stratum of highest dimension, and any other stratum of lower dimension belongs entirely to some other stratum of higher dimension. For indicatrices in three-dimensional space we only deal with one-dimensional manifolds I1 with endpoints or with a smooth surface I2 in R3 on which there is a smooth curve I1 (a boundary) or two mutually transversal curves I1 1 and I1 2 (the sides of a corner) whose intersection point is treated as a separate stratum. However, the main constructions described below remain valid in the general multidimensional case. Thus, in the general case to a point q of a k-dimensional submanifold I we assign the set (diffeomorphic to the sphere S3\u2212k\u22121) of germs at q of all co-oriented planes tangent to the submanifold at this point. All such germs form a smooth Legendrian submanifold LI \u2248 I\u00d7S3\u2212k\u22121 in the space ST \u2217R3 of co-oriented contact elements. Forgetting the base point of the germ, we obtain the projection \u03c0\u2217 on the space of all co-oriented planes of this submanifold, which is a Legendre map of the submanifold LI to the dual space R\u03023. The image (wavefront) I\u0302 = \u03c0\u2217(LI) of this projection is called the Legendre transform of the initial submanifold I, or the dual surface to I. Denote by I\u0302A the germ of the dual surface to I at the set of all planes tangent to I at the point A. The Legendre transform for a stratified submanifold J is defined as the collection J\u0302 of the Legendre transforms \u03c0\u2217(LIj ) of all the strata Ij . A manifold Jc with boundary or corners is regarded as a subset of the stratified submanifold determined by the corresponding inequalities. Therefore, the corresponding Legendrian submanifold LJc is taken to be the subset of tangent elements at the points of Jc, and the Legendre transform J\u0302c is formed by the corresponding subsets of the dual surfaces of the strata of J . The space of all co-oriented planes in R\u03023 is fibred over the sphere of unit normals (with one-dimensional fibre of parallel planes). The boundary H(\u0393) of the convex hull of the compact subset \u0393 is determined by the set P (\u0393) of support planes of \u0393. Namely, for any unit normal we choose, among all the parallel planes intersecting 268 A.A. Davydov and V.M. Zakalyukin this hull, a plane P that corresponds to the maximum value of the coordinate whose gradient is directed along this normal. We note that a plane P \u2208 R\u03023 is a support plane of J if and only if it is tangent to some of the strata Ij at one or more points, or in other words, belongs to the image \u03c0\u2217LJ . For any support plane P we denote by SP = \u03b3\u22121(P ) the set of points at which this plane is tangent to \u0393. We call it the base of P and its points the base points. The number \u00b5P of distinct points of the base SP will be called its multiplicity. Thom\u2019s transversality theorem [17] for the space of multi-jets of maps or families of maps imposes some restrictions on the possible values of the multiplicity of support planes, types of Legendre maps \u03c0\u2217, and so on. Moreover, the condition that the origin O belongs to the support plane under consideration also imposes some constraints on the possible types of generic singularities of the Legendre map. For instance, for a generic surface \u0393 the multi-germs \u03c0\u2217 |SP are Legendre stable. Each germ is Legendre equivalent either to a germ with a singularity of A1 type (in a neighbourhood of such a point the surface is the graph of a Morse function and its quadratic form is negative definite) or to a germ with a singularity of A3 type (which corresponds to the Legendre map of the graph of the function h = x4 + (y\u2212 x2)2). The number of points in each base SP is at most three for singularities of A1 type and at most one for a singularity of A3 type. Thus, it is feasible to list all possible singularities of dual surfaces and to classify singularities of the convex hulls and the transitivity zones using the following advantageous properties of the Legendre transform. 1. The Legendre transformation of a generic hypersurface repeated twice is the identity map. Indeed, a Legendrian submanifold has two Legendre projections: \u03c0\u2217 and the standard projection \u03c0 : ST \u2217R3 \u2192 R of the fibre bundle. A smooth Legendrian submanifold LI is uniquely determined by its wavefront I\u0302 provided that the regular points of \u03c0\u2217 are dense in LI [31]. Hence the repeated Legendre transformation of the hypersurface I\u0302 yields the same Legendrian submanifold LI and the projection \u03c0, whose image \u03c0(LI) coincides with I. 2. The Legendrian submanifold LI2 of a smooth surface I2 has a regular projection on the surface itself: the tangent plane at each point is unique. In a neighbourhood of the convex hull of the base SP of a support plane P tangent to I2, the convex hull boundary H(I) is determined by the support planes which are close to P . The germ of the dual surface (I\u03022, P ), where P is the plane tangent to the generic surface I2 at a point q, is smooth provided that q is not a parabolic point. At a generic point of a parabolic line \u03b4, this germ is diffeomorphic to a semicubic cylinder (that is, to the bifurcation diagram of an A2 singularity). At isolated points of the line \u03b4, the Legendre transform has a singularity of A3 type, and the germ of (I\u03022, P ) is diffeomorphic to a swallowtail. We note that the interior points of A2 type cannot belong to the convex hull, since the tangent plane divides the surface in a neighbourhood of the point q. 3. The Legendrian submanifold LI1 for a smooth curve I1 is swept out by the circles Sq containing the tangent elements to I1 at its points q. The dual surface I\u03021 is ruled: in the affine chart the circles Sm correspond to straight lines. The set of support planes P tangent to the curve at a fixed point q forms a connected closed arc Eq on Sq. A support plane P passing through q \u2208 I1 can be rotated about the Controllability of non-linear systems 269 tangent to the curve until it touches the curve at some other point or becomes an osculating plane. The germ of the convex hull near the point q of a one-dimensional stratum I1 is determined by the germ Su(I1) on the arc Eq of support planes passing through q. On a generic curve some isolated points may have simple flattening (in a neighbourhood of such a point in the canonical Frenet frame the curve has the parametric form q1 = t, q2 = t2 + \u00b7 \u00b7 \u00b7 , q3 = t4 + \u00b7 \u00b7 \u00b7 ), while at all other regular points the germ of the curve has the canonical form q1 = t, q2 = t2 + \u00b7 \u00b7 \u00b7 , q3 = t3 + \u00b7 \u00b7 \u00b7 . In a neighbourhood of the osculating plane, which is given in both cases by the equation q3 = 0, the dual surface has a singularity of A3 type (swallowtail) in the flattening case and a singularity of A2 type (semicubic cylinder) in the regular case. The osculating plane at a regular interior point (one which is not an endpoint) cannot be a support plane: the curve is located on both sides of the plane (however, at an endpoint the osculating plane can be a support plane). 4. Any point q \u2208 R3 corresponds in the dual space to the plane q\u0302 consisting of all planes in R3 passing through q. The germ of the convex hull in a neighbourhood of a point q which is a zero-dimensional stratum of J (that is, a corner vertex or a curve endpoint) is determined by the germ of Su(J\u0302) on the convex domain Uq \u2282 q\u0302 consisting of all support planes passing through q. 5. The Legendre transforms I\u03021 and I\u03022 of two strata I1 \u2282 I2 are tangent along their intersection I\u03022 \u2229 I\u03021, which consists of the tangent planes to the stratum I\u03021 at the points of I\u03022. Indeed, the surfaces I\u0302i = \u03c0\u2217(LIi ), i = 1, 2, intersect along the set of common tangent planes at the points of the smaller stratum I1. Due to the involution property of the Legendre transformation, the planes tangent to \u03c0\u2217(LI) are points of the submanifold I itself. Therefore, the planes tangent to I\u03021 and I\u03022 at their intersection points correspond to the common points in I1 \u2229 I2 and hence coincide, as required. If the quadratic form of the surface I\u03022 is non-degenerate in the directions tangent to I2 and transversal to I1, then the tangency of I\u03021 and I\u03022 is of the first order. 6. A point Q \u2208 R3 corresponds in the dual space to the plane Q\u0302 consisting of all planes in R3 passing through Q. A point Q belongs to the convex hull boundary of a submanifold B if and only if the plane Q\u0302 is negatively supporting for the subset Su(B) of support planes of B: the open negative half-space of Q\u0302 does not contain points of Su(B), but the plane itself contains such points. In particular, the point O belongs to a convex surface X \u2282 R3 if and only if the dual plane O\u0302 is tangent to the dual surface X\u0302. 7. If O belongs to the convex hull of the base of the support plane for some generic three-parameter family of submanifolds B of dimension k, then the singularities of the Legendre transform at each base point are generic singularities of the Legendre transforms of k-dimensional submanifolds. Indeed, the dimension of the convex hull of the base consisting of l points is less than or equal to l\u2212 1. In order for the origin O to belong to the convex hull of the base, at least 3\u2212 l + 1 conditions must be satisfied. For l points to fall on the osculating plane, l conditions must hold. A degenerate point appears on the base if at least one condition holds. Thus, in this case more than three independent conditions must be satisfied, which is generically impossible for three parameters. 270 A.A. Davydov and V.M. Zakalyukin Now a classification of generic singularities of the boundary of the transitivity zone is obtained by listing all possible positions of the origin in the base of a support plane, describing the corresponding families of dual surfaces and their supporting subsets, and, finally, applying the following results in singularity theory. A diffeomorphism \u03a6: R\u03023 \u2192 R\u03023 mapping the plane O\u0302 onto itself and preserving its positive half-space is called admissible. If \u03a6 maps the support part Su(J1) of the Legendre transform of a stratified manifold J1 to the Legendre transform Su(J2) of another stratified manifold, then, clearly, the origin O belongs to the convex hull boundary H(J1) if and only if it belongs to H(J2). Thus, a family of admissible diffeomorphisms \u03a6m, fibred over a diffeomorphism of the parameter space, acts on the parameter-dependent families of surfaces I\u0302j and maps the respective transitivity zones to each other. Reducing the equations of the dual strata I\u0302j to normal forms via admissible diffeomorphisms, we obtain normal forms of the transitivity zone. For this purpose, consider the action of the diffeomorphism \u03a6m as a contact transformation on the product of the equations of all the components [32]. Denote by Pi(x), i = 0, . . . , n, the polynomials of degree ki in x \u2208 R of the form P0 = xk0 + \u2211k0\u22122 j=0 xja0j , Pi = xki + \u2211ki\u22121 j=0 xjaij , i = 1, . . . , n. Denote by a \u2208 RN the vector of coefficients of all these polynomials. We note that P0 is the standard miniversal deformation of the singularity Ak0\u22121, and Pi for i = 1, . . . , n are versal deformations of the singularities Aki\u22121. Lemma. Let gi(x, b), i = 0, . . . , n, be functions of x with parameters b \u2208 RN such that the values of gi(x, 0) and the values of their derivatives up to order ki \u2212 1 vanish at the origin (the function gi(x, 0) has singularity Aki\u22121 at the origin), and let g(x, b) = \u220fn i=0 gi(x, b). Then there is a contact equivalence consisting of a diffeomorphism (x, b) 7\u2192 ( X(x, b), B(b) ) and a non-zero function \u03d5(x, b) which reduces the function g to the form g(x, b) = \u03d5(x, b) \u220fn i=0 Pi(X, B). The assertion of the lemma is equivalent to the versality of the map described below with respect to a special group of equivalences. Consider the map G : Rk \u2192 Rn+1, G : x 7\u2192 ( g0(x), . . . , gn(x) ) , and in the target space consider the collection Y = \u22c3 {yi = 0} of coordinate hyperplanes. Let D be the group of diffeomor- phisms of Rn+1 which preserve Y ; that is, they have the form \u03b8 : (y0, . . . , yn) 7\u2192( h0(y)y0, . . . , hn(y)yn ) . Two maps G1 and G2 are called Y -contact equivalent if for some family of diffeomorphisms \u03b8x in the group D that depend on a parameter x we have G2(x) = \u03b8x \u25e6G1(X(x)) for some change of variables x 7\u2192 X(x). For a family Gm : Rk \u00d7 RN \u2192 Rn+1 of maps G depending on parameters a in RN , a parameter-dependent contact equivalence is defined in a natural way: the parameters a are replaced via a diffeomorphism by parameters b; a diffeomorphism of the form x 7\u2192 X(x, a) acts on the variables x; a family of diffeomorphisms \u03b8x,a : Rn+1 \u2192 Rn+1 acts on the target space Gm. The notions of versal and infinitesimally versal deformations of a map G are straightforward. Obviously, for this group D of transformations an analogue of the versality theorem holds, since the group is geometrical in the sense of J. Damon. We note that the map P\u0303 : (x, a) 7\u2192 (P0, . . . , Pn) is an infinitesimally versal deformation of the map x 7\u2192 (xk0 0 , . . . , xkn n ). The versality of the map P\u0303 implies the assertion of the lemma: under an equivalence in the group D each component Controllability of non-linear systems 271 is multiplied by a non-zero factor; hence the product is multiplied by a non-zero factor." ] }, { "image_filename": "designv11_62_0000130_j.physc.2013.04.021-Figure8-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0000130_j.physc.2013.04.021-Figure8-1.png", "caption": "Fig. 8. Scheme of magnetic field calculation.", "texts": [ " Due to its anisotropy, HTS coil is very sensitive to the external magnetic field, especially perpendicular field, which reduces the capacity. Fig. 7 shows the influence of external magnetic field perpendicular on critical current of HTS tape. Critical current decreases with perpendicular magnetic field. Therefore, the influence of the motor geometry on external magnetic field is researched. Then, a method to improve critical current of HTS coil is presented. First, we assume the coil takes only AC transport current. Fig. 8 is a simple and functional frame of this model. Treat the HTS coil as an infinitely long stack of tapes. Each tape carries the same transport current I and is insulated from each other. At the same time, the size of the slot is involved. Based on image method, the magnetic field caused by the image is treated as external magnetic field. The slot area can be classified into three parts based on the current distribution. Part II is the HTS coil. The perpendicular magnetic field effects on the critical current of conductor more significant than the parallel magnetic field does" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0002131_10402004.2017.1285970-Figure12-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0002131_10402004.2017.1285970-Figure12-1.png", "caption": "Figure 12: Scheme of the journal bearing geometry", "texts": [], "surrounding_texts": [ "1. Schirru, M., Mills, R., Dwyer-Joyce, R., Smith, O., and Sutton, M. (2015). Viscosity Measurement in a Lubricant Film Using an Ultrasonically Resonating Matching Layer. Tribology Letters, 60(3) pp. 1-11. 2. Mason, W.P., Baker W.O., McSkimin, H.J. et al. (1948). Measurement of shear elasticity and viscosity of liquids at ultrasonic frequencies. Physical Review. Vol. 75(6) pp. 936- 946 3. Bujard, M.R., (1989) Method of measuring the dynamic viscosity of a viscous fluid utilizing acoustic transducer. US Patent Number 04862384 4. Emmert, S.W., (1986) Apparatus and method for determining the viscosity of a fluid sample. US Patent Number 4721874 5. Greenwood, M.S,. and Bamberger, J.A., (2002) Measurement of viscosity and shear wave velocity of a liquid or slurry for on-line process control. Ultrasonics Vol. 39 pp. 623-630 6. Lamb, J., (1967) Physical properties of fluid lubricants: rheological and viscoelastic behaviour. Proceedings of the Institution of Mechanical Engineers, Conference Proceedings. Vol. 182 pp. 293-310 7. Barlow, A.J., and Lamb, J., (1959) The visco-elastic behaviour of lubricating oils under cyclic shearing stress. Proc. R. Soc. Lond. Vol 253 pp. 52-69 8. Dowson, D. G. R. A. V., Higginson, G. R., and Whitaker, A. V., (1962) \"Elasto- hydrodynamic lubrication: a survey of isothermal solutions.\" Journal of Mechanical Engineering Science 4.2 pp. 121-126. ACCEPTED MANUSCRIPT 27 9. Schirru, M., and Dwyer-Joyce, R.S., (2015) A model for the reflection of shear ultrasonic waves at a thin liquid film and its application to viscometry in journal bearings. Proc IMechE Part J: J Engineering Tribology, [online] accessible from: http://pij.sagepub.com/ 10. Collin, R.E., (1955) Theory and design of wide-band multisection quarter-wave transformers. Proceedings of the IRE Vol.36 pp. 621-629 11. Emmert, S.W., (1986) Apparatus and method for determining the viscosity of a fluid sample. US Patent Number 4721874 12. Bair, S., and Winer, W.O., (1980) Some observations on the relationship between lubricant mechanical and dielectric transitions under pressure, Trans. ASME, Journal of Lubrication Tech., 102, 2, pp. 229-235 13. Hutton, J.F., (1967) Viscoelastic relaxation spectra of lubricating oils and their component fractions. Proc. Roy. Soc. London Mathematical and Physical Sciences. Vol. 304, 1476, pp.65-80 14. Johnson, K. L., & Tevaarwerk, J. L. (1977). Shear behaviour of elastohydrodynamic oil films. Proceedings of the Royal Society of London A: Mathematical, Physical and Engineering Sciences Vol. 356, No. 1685, pp. 215-236 15. Kinsler, E., et al. (2000) Fundamentals of acoustics , 4 th Edition, Wiley, New York 16. Harrison, G., and Barlow, A.J.. (1981) Dynamic Viscosity Measurement. Methods of experimental physics 19 pp. 137-178. 17. Stachowiack, G.W., and Batchelor, A.W., (2001) Engineering Tribology. 2 nd ed. Butterworth Heinemann, pp. 204-210. ACCEPTED MANUSCRIPT 28 18. Raimondi, A.A., and Boyd, J., (1958) A solution for the finite journal bearing and its application to analysis and design. ASLE Transactions. Vol. 1 pp. 159-209 19. Juvinall, R.C., and Kurt, M.M., (2006) Fundamentals of machine component design. 5 th ed., John Wiley & Sons ACCEPTED MANUSCRIPT 32 ACCEPTED MANUSCRIPT 33 ACCEPTED MANUSCRIPT 34 ACCEPTED MANUSCRIPT 35 ACCEPTED MANUSCRIPT 36 ACCEPTED MANUSCRIPT 37 ACCEPTED MANUSCRIPT 38 ACCEPTED MANUSCRIPT 40 ACCEPTED MANUSCRIPT 41 ACCEPTED MANUSCRIPT 42 ACCEPTED MANUSCRIPT 43 ACCEPTED MANUSCRIPT 44 ACCEPTED MANUSCRIPT 45 ACCEPTED MANUSCRIPT 46 ACCEPTED MANUSCRIPT 47" ] }, { "image_filename": "designv11_62_0000217_j.oceaneng.2012.10.005-Figure3-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0000217_j.oceaneng.2012.10.005-Figure3-1.png", "caption": "Fig. 3. Body and Earth-fixed axis.", "texts": [ " To determine the position and the orientation of the USV, it is necessary to define six independent coordinates. Here, the first three coordinates and their time derivatives correspond to the position and translational motion along the x-, y-, and z-axes, and the last three coordinates and time derivatives correspond to orientation and rotational motion. The six different motion components for marine vehicles are usually defined as: surge, sway, heave, roll, pitch and yaw. When analyzing the motion of marine vehicles in 6 DOF it is usual to define two coordinate frames, as shown in Fig. 3. The moving coordinate frame xyz is fixed to the vehicle and it is usually referred as the body-fixed reference frame. The center of gravity (CG) is chosen to be the origin \u2018o\u2019 of the body-fixed frame. And CG is usually in the principal plane of symmetry. For marine vehicles, the body axes x, y and z coincide with the principal axes of inertia, where, x is usually defined as the longitudinal axis which is directed from aft to fore; y as transverse axis directed from port to starboard; z as normal axis directed from top to bottom" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0003741_012077-Figure5-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0003741_012077-Figure5-1.png", "caption": "Figure 5. Nodes coordinates of the simulated mechanism, in totally deployed configuration.", "texts": [ " The first example is a mobile house (see figure 3). Another application of the mechanism described here road barrier, represented in folded, partially deployed and totally deployed configurations (figure 4). Modern Technologies in Industrial Engineering VII, (ModTech2019) IOP Conf. Series: Materials Science and Engineering 591 (2019) 012077 IOP Publishing doi:10.1088/1757-899X/591/1/012077 The first simulations have been done, considering the coordinates of the mechanism nodes in the fully extended configuration as illustrated in figure 5. We have to note that no dimensional synthesis of this mechanism has been realized till now. As consequence, the mechanism could not fold in the optimum compact configuration. The driving link 1 could only rotate with an angle 321 degree (see figure 6), starting from the extended configuration. It means that the simulated mechanism could only fold till the partially deployed configuration represented in figure 1(b). Also, the trajectory of the G and H nodes are not yet straight lines, as they should be, but as they are presented in figure 7, even if the link 7 has not a significant rotation around z axes, as seen in figure 8", " Series: Materials Science and Engineering 591 (2019) 012077 IOP Publishing doi:10.1088/1757-899X/591/1/012077 Modern Technologies in Industrial Engineering VII, (ModTech2019) IOP Conf. Series: Materials Science and Engineering 591 (2019) 012077 IOP Publishing doi:10.1088/1757-899X/591/1/012077 Modern Technologies in Industrial Engineering VII, (ModTech2019) IOP Conf. Series: Materials Science and Engineering 591 (2019) 012077 IOP Publishing doi:10.1088/1757-899X/591/1/012077 The second mechanism discussed here is shown in figure 5. It has the total number of links, n = 8 and the number of rotary joints, g1 = 13. As we may see in the example of its application, this Modern Technologies in Industrial Engineering VII, (ModTech2019) IOP Conf. Series: Materials Science and Engineering 591 (2019) 012077 IOP Publishing doi:10.1088/1757-899X/591/1/012077 mechanism may allow to built deployable structures (mobile house, for example) with shading system (see figure 10). According to Grubler formula, for n = 10 and g1 = 13, the number of degrees of freedom of this mechanism is: 11321103 F " ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0001408_icuas.2013.6564790-Figure1-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0001408_icuas.2013.6564790-Figure1-1.png", "caption": "Fig. 1. Path following formulation problem", "texts": [ "00 \u00a92013 IEEE 1022 In this study we explore a path following application for a small fixed-wing UAV for which the directional control is of primary interest. Hence, this section provides a concise formulation of the problem of following a desired path and it introduces briefly the lateral equations of motion for a fixed-wing UAV in a non-steady atmosphere. Applications such as the one treated in this paper refer to the problem of steering a vehicle along a desired path, despite uncertainties in the system dynamics or external perturbations. A geometrical description of the problem is shown in Figure 1, where an airplane has a distance d relative to an inertial path. and it flies in a moving surrounding airmass W . Whether we refer to the movement of the aircraft relative to the surrounding atmosphere, or to the movement of the latter relative to the aircraft, the flight is possible due to the difference velocity of the airflow around the upper and the lower side of the aircraft wing. The direction of the airflow experienced by the aircraft is affected by environmental wind, which is any motion of the air mass relative to the Earth\u2019s surface. The wind affects both the longitudinal and the lateral variables of the airplane contributing to its nonlinear, coupled and complex dynamics. Therefore, knowing the wind is essential for the control system of the airplane. A simple kinematic model of the airplane depicted in Figure 1 was presented in [20] and it is given by p\u0307n = V cos\u03c8 + \u03c9 cos\u03c8\u03c9 (1a) p\u0307e = V sin\u03c8 + \u03c9 sin\u03c8\u03c9 (1b) where pn and pe represent the inertial position in the North and in the East axis, respectively, \u03c9 cos\u03c8\u03c9 = WN , \u03c9 sin\u03c8\u03c9 = WE , \u03c9 is the wind velocity and \u03c8\u03c9 describes the wind direction. The following assumptions are considered regarding the simplified model stated previously. The longitudinal variables of the airplane (airspeed, pitch and flight path angle) are stabilized by means of an inner autopilot and their time derivatives can be neglected in the flight dynamics", " The kinematics of a truly banked turn are given in terms of Euler angles as (see [21]) \u03c8\u0307 = g V tan\u03c6 (4) In what follows, we will assume that the inner autopilot of the airplane implements a bank-hold loop using ailerons and rudder as control input. The differential equation describing the resulting dynamics is expressed as \u03c6\u0307 = k\u03c6 (\u03c6 c \u2212 \u03c6) (5) where k\u03c6 is a positive constant and \u03c6c is the commanded bank angle. The advantage of using the line segment frame, Fs, is that the path following problem is formulated only in terms of cross-track error, d, as a regulation problem. This follows from the geometrical representation depicted in Figure 1, where the regulation of d involves Pe_s \u2192 0 in Fs, while Pn_s moves along the segment. To simplify further analysis, let us assume that the wind velocity and direction change slowly such that they can be considered quasi-constant. Notice from Eq. (3) that the deviation from the desired trajectory, parameterized by its orientation \u03c8s, depends on the airplane velocity and on the wind parameters. Thus, without loss of generality, the airplane dynamics for trajectory following purpose can be defined as d\u0307 \u2261 p\u0307es = V sin (\u03c8 \u2212 \u03c8s) + k\u03c9 (6) \u03c8\u0307 = g V tan\u03c6c (7) \u03c6\u0307 = k\u03c6 (\u03c6 c \u2212 \u03c6) (8) where k\u03c9 = \u03c9 sin (\u03c8\u03c9 \u2212 \u03c8s) is considered constant for control purpose, and d is the cross track error from the desired trajectory" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0003396_pedstc.2019.8697749-Figure3-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0003396_pedstc.2019.8697749-Figure3-1.png", "caption": "Fig. 3 Mesh plot for simulated SRM.", "texts": [ " In the Duty Cycle control algorithm a moving window is utilized to get the average of voltage in chopping instance in each sampling period and based on the obtained average avoltage in each instance, Duty Cycle for demagnetization will be calculated. Fig. 2 shows the proposed control algorithm in the block diagram format. In following sections a comparison between different methods is done to show how this control algorithm can reduce the vibration. MAXWELL 2D is used for simulation. Simulated motor is shown in Fig. 3 with the mesh grid plotted on and mesh properties are shown in table II. In this method three voltage levels of +VDC , 0 and \u2013VDC are imposed on the windings and conduction period for each phase is devided into three stages : magnetization , chopping and demagnetization. Also depending on the loading condition , rotation speed of shaft and level of current, the chopping instance may not occur or current waveform may be something different to what is considered. Here for illustrational purposes load torque, speed and current level are the same for all conditions to show how the control algorithm can influence the radial forces and induced vibration" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0002782_1350650117739764-Figure6-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0002782_1350650117739764-Figure6-1.png", "caption": "Figure 6. Performance of the lubricating test oils normalized to one, where a value of zero is excellent; a value of one would imply a bad performance.", "texts": [ " The properties of the wind turbine gear oils have been reviewed in many different laboratory scale tests. Every test provided considerable amounts of data, which here has been combined in order to identify the oils\u2019 performances as good as possible. For that purpose, an analysis is done by considering all results from the previous tests. To do this, a weighing scheme (Table 4) in combination with normalization (equation (4)) was used (see Appendix 1). Finally, the overall assessment considering all tests is found (Figure 6). The performance shows that PAG-I, PAO-I, and PAO-II do not inhabit any substantial changes when comparing their new and the used samples. This stability over the lifetime is a very good and aspired characteristic. The best examination regarding all tests was accomplished by PAO-III and MIN-II. However, the PAO-III shows a strong degradation towards the used oil. Especially, a depletion of the additives can cause those changes. The PAO-III has still better results in the used quality than the PAO-I and MIN-I" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0001685_978-3-642-33926-4_20-Figure1-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0001685_978-3-642-33926-4_20-Figure1-1.png", "caption": "Fig. 1. Structure of truck-trailer mobile robot system", "texts": [ " Based on the Parallel Distributed Compensation (PDC) concept [1-3, 8], a parameter-dependent quadratic Lyapunov function is used to for the synthesis of the discrete LPV T-S fuzzy systems. It can be found that the stability conditions derived in this paper can be transformed to the LMI formulas. Hence, the convex optimization technique [9] can be used to solve these LMI stability conditions. 2 LPV T-S Fuzzy Model of TTMR System The dynamic equations of nonlinear TTMR system have been introduced in [1-3]. The structure of the TTMR system can be referred to the Figure 1. According to [1-3], the simplified TTMR system can be described as follows. ( ) ( ) ( ) ( ) ( )1 1 2 1 1 1 k t k t x k x k u k L L \u03b8 \u03b8\u22c5\u0394 \u22c5\u0394 + = \u2212 + (1a) ( ) ( ) ( ) ( )2 1 2 2 1 k t x k x k x k L \u03b8 \u22c5 \u0394 + = + (1b) ( ) ( ) ( ) ( ) ( ) ( )3 1 2 3 2 1 2 k t x k k t sin x k x k x k L \u03b8 \u03b8 \u22c5\u0394 + = \u22c5\u0394 \u22c5 + + (1c) where ( )k1x is the angle difference between truck and trailer, ( )k2x is the angle of trailer, ( )k3x is the vertical position of rear end of trailer, ( )ku is the steering angle, 1L is the length of truck, 2L is the length of trailer, t\u0394 is the sampling time and ( )k\u03b8 is the speed of backward movement" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0000741_amr.668.495-Figure4-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0000741_amr.668.495-Figure4-1.png", "caption": "Fig. 4 Contact stress distribution.", "texts": [ " In the orthogonal experiment analysis, element type of modeling chooses Solid 185, which is used for the 3-D modeling of solid structures. The contact surface type of interference fit is surface-to-surface, using CONTA 174 to simulate area for contact [7]. The elastic modulus and Poisson's ratio for shaft and sleeve are the same, in which the elastic 111.2 EE = , Poisson's ratio 28.0=v . FEM Analysis and Calculation. In order to determine the primarily factors that influences contact stress, an orthogonal experiment was conducted to study the impacts of shaft diameter, wall thickness, mating length and interference on stress. Fig. 4 shows an example of the von Misses stress distributions for the shaft\u2013hub connection. According to Lame\u2019s solution for thick-walled cylinders, contact stress of Fig. 4 is calculated as Eq. 4. Results from the FEA analyses were compared with theoretical analysis on contact stress of contact surface. In addition, error between the FEA and the theoretical was also calculated using Eq.5. ( ) %56.5%100 336 33.317336 100% =\u00d7 \u2212 =\u00d7 \u2212 = FEM FEMTheory Error (5) An error not exceed 5% is considered acceptable. Even with that margin of error, it was clear that the theoretical results are close to the FEA model values. The analysis shows that stress concentration occurs obviously at the end edge of contact surface which is not predicted by Lam\u00e9 theory" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0000900_ever.2013.6521527-Figure6-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0000900_ever.2013.6521527-Figure6-1.png", "caption": "Fig. 6. Flux pattern in REL motor", "texts": [ " The ratio between the two inductances defines the saliency ratio of the REL motor, that is, \ufffd = Ld/ Lq\u2022 From the vector diagram, it is possible to obtain the follow equations: (14) where 0:; and o:\ufffd are the current vector and flux Iinkage vector angles, respectively, and 'Pi is the inner power factor angle. Since (15) the power factor angle is related to the current vector angle and the motor saliency. The motor torque can be computed from Maxwell's stress tensor or from the flux linkages as 3 Tre1 = \"2P (Adiq - Aqid) (16) The FE analysis is carried out, referring to the same size considered for the IM above. In particular, the same stator geometry is considered, together with the stator winding distribution. The flux lines are shown in Fig. 6 In each FE analysis the stator currents id and iq are fixed. They are transformed through Park's transformation in the actual stator currents ia, ib and ic, which are assigned within the slots, according to the winding distribution. From the magnetic field solution, the REL motor performance are computed, by means of relations defined from the vector diagram of Fig. 5. The d- and q-axis stator flux linkages versus d- and q-axis currents of the synchronous REL motor are shown in Fig. 7. Comparing with the results in Fig" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0003808_codit.2019.8820640-Figure1-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0003808_codit.2019.8820640-Figure1-1.png", "caption": "Fig. 1. Simplified cross-section diagram of the IPMSM.", "texts": [ " Therein, the q- and d-axes inductances and the permanent magnet flux linkage are updated online using the RELS algorithm to improve the accuracy of the optimal vector angle calculation. To track parameter variations in the IPMSM, the random walk and modified forgetting factor techniques are incorporated with the RELS algorithm. The effectiveness of the proposed adaptive MTPA control method is verified by simulation studies. 978-1-7281-0521-5/19/$31.00 \u00a92019 IEEE -397- II. MTPA OPERATION OF IPMSMS A three-phase IPMSM is considered in this paper, with the simplified cross-section diagram shown in Fig. 1. It is assumed that the machine has negligible eddy currents and hysteresis losses. The voltage equations of the IPMSM in the qd rotor reference frame are given by [16]: \u03bb \u03c9 \u03bb \u03bb \u03c9 \u03bb = + + = + \u2212 q q q e d d d d e q d v Ri dt d v Ri dt , (1) where vq,d, iq,d and \u03bbq,d are the stator voltages, stator currents, and flux linkages of the IPMSM, respectively; and \u03c9e denotes the electrical angular velocity of the machine. The relationship between the flux linkages and the stator currents can be obtained as [16]: \u03bb \u03bb \u03bb = = + q q q d d d f L i L i , (2) where \u03bbf is the flux linkage of the permanent magnet, and Lq,d are the stator inductances in qd reference frame" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0003356_978-981-13-5799-2_12-Figure12.21-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0003356_978-981-13-5799-2_12-Figure12.21-1.png", "caption": "Fig. 12.21 Transient tire model under the braking condition", "texts": [ " The value of n for the contact pressure distribution changes with speed from 3.66 to 4.89, while the value of f for the contact pressure distribution also changes with speed from 0.21 to 0.83. The maximum braking force decreases with speed. This is because the kinetic friction coefficient is governed by not only the slip ratio but also the sliding speed as discussed in Sect. 12.1.1-(4). (1) Tire deformation for fluctuating braking torque Araki and Sakai [6] studied the transient tire model in braking and driving conditions. Figure 12.21 shows the transient tire model in the braking condition where the road moves to the right at a speed VR in the coordinate system moving with the tire axle. When the fluctuating braking torque T is applied to the axle of a tire with a zero slip ratio, the braking force fx also fluctuates. Assuming that the speed of the belt VB fluctuates under the braking torque T, VB is expressed by where DV is the relative speed between the belt and road and is expressed by DV \u00bc a sinxt: \u00f012:39\u00de Here a and x are, respectively, the amplitude and frequency of the fluctuating velocity", " Differentiating Fx with respect to sB, we obtain dFx dsB \u00bc CFssB 2C2 Fs 3lsFz 2 ld ls sB \u00fe C3 Fs 9l2sF2 z 3 2 ld ls s2B \u00bc 0 ! CFs 3lsFz 3 2 ld ls sB 1 CFs 3lsFz sB 1 \u00bc 0 ! sB \u00bc 3lsFz CFs ls ld 3 ls ld 2 ; 3lsFz CFs : Note 12.3 Eqs. (12.51), (12.55) and (12.56) Equation (12.51) Note that Dx1 is the cumulative relative displacement between the belt and road from the leading edge of the contact patch to the position x1. Meanwhile, Dx0 is the relative displacement between the belt and road at some time. Figure 12.21 shows that the tire is under braking. The slip ratio is then positive and is given by s \u00bc VR \u00f0VB \u00feDV2\u00de VR \u00bc DV DV2 VR : Equations (12.55) and (12.56) Referring to Fig. 12.21, we have Dx2 \u00bc Fx Rx Dx1 \u00bc Fx Cxwl : Referring to Fig. 12.21, when torque is applied to the contact surface, we have Dx \u00bc Dx1 \u00feDx2 Gx \u00bc Fx Dx \u00bc Fx Fx Cxbl \u00fe Fx Rx \u00bc RxCxbl Rx \u00feCxbl : Referring to Eq. (12.9), CFs is given by CFs = Cxbl 2/2. Taking the limit x/VR ! 0, we obtain A \u00bc 1\u00fe Cxb Rx l VR x sin xl VR \u00bc 1\u00fe Cxb Rx l l sin xl VR = xl VR ! 1 B \u00bc CxbVR xRx cos xl VR 1 ! 0; ffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi VR x sin xl VR l 2 \u00fe V2 R x2 cos xl VR 1 2 s ffi ffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi l sin xl VR = xl VR l 2 \u00fe V2 R x2 x2l2 2V2 R 2 s ffi xl2 2VR ; Kd \u00bc CxbV x ffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi A2 \u00feB2 p ffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi VR x sin xl VR l 2 \u00fe V2 R x2 cos xl VR 1 2 s " ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0002302_b978-0-12-803581-8.10351-0-Figure9-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0002302_b978-0-12-803581-8.10351-0-Figure9-1.png", "caption": "Fig. 9 Transparent view of assembled monocoque showing cockpit stiffener outline. Reproduced from Radford, D.W., Fuqua, P.C., Weidner, L.R., 2004. Tooling development for a multi-shell monocoque chassis design. In: 36th International SAMPE Technical Conference, San Diego, CA, November 15\u201318, 2004, pp.1063\u20131077.", "texts": [ " In the multi-shell design approach, the cockpit rim stiffener is a major focus of the design strategy. Since the perimeter of the opening is continuous in the top shell, continuous fiber reinforcement will be utilized. Further, the assembly of the top and side shells results in a hollow bonded cockpit perimeter beam. The assembly of the two open geometries, again minimizing manufacturing complexity, creates this closed beam. Fig. 8 shows exploded views of the primary composite shells, illustrating the assembly of the monocoque. Fig. 9 outlines the hollow section around the cockpit opening that results from bonding the sides and top shells together, which directly follows the strategy summarized in the multi-shell pressure vessel discussion. To create the hollow section cockpit opening stiffener, the top shell incorporates three sides of the section and is bonded along two surfaces of each side shell as shown in Fig. 10. This hollow cross-section encircles the entire perimeter of the cockpit opening, but the dimensions of the cross-sectional geometry of the stiffener vary continuously" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0003394_s40995-019-00711-7-Figure2-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0003394_s40995-019-00711-7-Figure2-1.png", "caption": "Fig. 2 Silver-coated electrode (a) and prepared electrode (b)", "texts": [ " 2012; Yan et\u00a0al. 2010). In order to prepare nanocomposites of GO/polyaniline, it was dried using oven at 45\u00a0\u00b0C for 24\u00a0h. The steps of the formation of GO/PANI nanocomposite and functionalized GO with PANI and benzene derivative to induce extended conjugation in GO, and consequently, an increase in the electrical conductivity in GO/PANI nanocomposites is shown in Fig.\u00a01. The electrode designed for this study has special location for the placement of GO\u2013PANI nanocomposites along with tin oxide nanoparticles (Fig.\u00a02a). The design was fabricated as a silver-coated printed circuit board (in corporation with 1 3 Alpha Circuit Co.). In this electrode, the metal layer is 35 micron in thickness and the bottom of the metal fiber layer is made from fiberglass. To prepare the electrode, 2\u00a0mg of synthesized powder of GO\u2013PANI nanocomposite was added to 500\u00a0\u00b5l of distilled water and placed in an ultrasonic bath for half an hour. At specific location indicated on each electrode, 10\u00a0\u00b5l of the dispersant produced by the sampler was shown and the electrodes were placed in an oven at 45\u00a0\u00b0C temperature for 2\u00a0h. In this way, six electrodes were prepared (Fig.\u00a02b). In this study, six nanobiosensors were fabricated with the addition of various amounts of tin oxide nanoparticles to the prepared electrodes that are similar in appearance, but differ in the amount of added metal nanoparticles. The detection mechanism of ethanol gas, which is one of the VOCs, in each component of this nanobiosensor of graphene oxide/ metal oxide is as follows (Mori et\u00a0al. 2014; Zhu et\u00a0al. 2004). Equation\u00a0(1) is formulated for the oxygen absorption equation for metal oxides, and Eqs" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0000267_s11517-013-1052-7-Figure1-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0000267_s11517-013-1052-7-Figure1-1.png", "caption": "Fig. 1 a A typical structure carrying eccentric markers. b Alternative markers with pattern. c\u2013e A possible application of the method: the markers are attached to modified instruments used for eye surgery; a black stripe is painted over the tip marker. e The top of a eye surgery simulator; cameras and illuminators are inside the box of 400 mm side under the mask; surgical instruments are inserted through the holes on the mask eye", "texts": [ " Tracking surgical instruments is required also by simulation systems that are used to train surgeons inside a virtual environment, to increase their skills and reduce learning time [14, 16]. Marker-based tracking is by far the most used approach because of its reliability and accuracy (cf. Supplemental Material, Appendix C). Small spherical retro-reflective markers are attached to the instrument and two or more cameras survey the scene: a PC host computes the 3D position of the markers in real time and, from this, the six degrees of freedom (dofs) of the instrument are determined. A support to displace the markers off the instrument axis is used to measure axial rotation (Fig. 1a) [19], although this changes the mass distribution and it may interfere with surgeon movements. More recently, compact regular patterns attached to the instrument have been introduced [3, 17] (Fig. 1b), but tracking errors may still arise and the accuracy decreases when the pattern is viewed at an angle with respect to the optical axis of the cameras. We propose here an innovative methodology, based on structured markers, that combines the best of the above approaches, to compute in real time the six dofs of a surgical instrument. It uses two or more spherical markers inserted on the instrument\u2019s axis with a black stripe painted over the surface of one of them (Fig. 1c). The procedure is complemented by an innovative circle fitting procedure, which allows the marker center to be reliably estimated Electronic supplementary material The online version of this article (doi:10.1007/s11517-013-1052-7) contains supplementary material, which is available to authorized users. N. Alberto Borghese (&) I. Frosio Applied Intelligent Systems Laboratory, Department of Computer Science, University of Milano, Via Comelico 39, 20135 Milano, Italy e-mail: borghese@di.unimi.it URL: http://borghese", " Although it provides a general framework for robust tracking, its accuracy cannot achieve the same accuracy obtained here were a careful analysis of the shape allows defining a proper weight for each pixel of the structure. The method can be applied in a wide range of possible domains. It has been applied by us in a simulator of eye microsurgery. In such system, the method is used to track the motion of two surgical instruments inserted into the eye of a mannequin and moved to complete a surgery according to the vision of the virtual operating theatre (Fig. 1d\u2013e). The method allows minimizing the impact in the weight distribution over the instrument realizing a more compelling simulation and to realize a very small simulator (400 mm per side) that can be easily transported. It is finally to be noticed that, if the stripe is sufficiently thin, its projection is described by conics theory; in this case, methods like the one in [13] can be applied, but painting a regular, thin stripe onto a curved marker can be problematic especially for small markers. To conclude, the proposed method to compute the dofs of a surgical instrument is based on use of structured markers combined with a rigorous analysis of accuracy" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0000718_ijcat.2011.043881-Figure3-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0000718_ijcat.2011.043881-Figure3-1.png", "caption": "Figure 3 Graphic description of the rotary crane", "texts": [ " This is done by minimising the sum of quadratic errors between optimal gains and estimated ones. The architecture of the NN is given in Figure 2. 4.1 Example 1 The first example concerns the optimal position control of a rotary crane. The fundamental motion of a rotary crane is the rotation and the load hoisting (Sakawa et al., 1981; Sakawa and Shindo, 1982). For simplicity, we assume that the container load can be regarded as a material point and that frictional torques which may exist in torque-transfer mechanisms can be neglected. The following notations, which are shown in Figure 3, will be used: \u03b81, rotation angle of the trolley drive motor; J1, total moment of inertia of trolley drive motor, a brake, a drum and a set of reduction gears; b1, equivalent radius of the drum of trolley drive motor which is reduced to the motor side; \u03b82, rotation angle of the hoist motor; J2, total moment of inertia of hoist motor, a brake, a drum and a set of reduction gears; b2, equivalent radius of the drum of the hoist motor which is reduced to the motor side; \u03c6, the load swing angle; m, total mass of the trolley and operator\u2019s cab; M, total mass of the container load, the spreader and the attached equipments; T1, driving torque generated by the trolley drive motor; T2, driving torque generated by the hoist motor; and g, the gravity acceleration. The state variables are x1 = b2\u03b82 \u2013 h, x3 = b1\u03b81, x5 = \u03c6 and their derivatives with respect to time (see Figure 3). Euler approach is used as a discretisation method of the system with a sample time T = 0.01 s. The choice of the sampling period is done to validate, later, the approximation described by equation (37). So, the process, to be controlled, can be described by a sixth-order nonlinear discrete system: 1, 1 1, 2, 2, 1 2, 1, 3, 1 3, 4, 4, 1 4, 2, 1 5, 5, 1 5, 6, 6, 1 6, 2, 2 5, 2, 6, 1, = = = = ( ) = = ( 2 ) ( ) k k k k k k k k k k k k k k k k k k k k k k k x x Tx x x Tu x x Tx x x T u x x x Tx Tx x u x x x x h \u03b1 \u03b1 + + + + + + +\u23a7 \u23aa +\u23aa \u23aa + \u23aa + +\u23a8 \u23aa +\u23aa \u23aa \u2212 + +\u23aa +\u23a9 (43) with \u03b11 = 66 m/s2 and \u03b12 = 76 m/s2" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0003666_iemdc.2019.8785290-Figure6-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0003666_iemdc.2019.8785290-Figure6-1.png", "caption": "Fig. 6. Flux tubes indicating 3D flux paths of the fringing air-gap flux density in a FE simulation", "texts": [ " It features a volumetric field winding carrying a fixed current and a virtually infinitely permeable (\u00b5r = 10000) stator core, rotor core and clamping plate. Clamping bolts and the associated holes through the clamping plate as well as non-magnetic pressure fingers (\u00b5r \u2248 1) were not modeled, as exploratory transient FE calculations showed them not to have a meaningful impact on the clamping plate field. Given the chosen assumptions, the FE models do not account for saturation or eddy current effects. Incorporating these effects into the field calculation is planned to be part of future work. V. 3D FLUX PATHS Fig. 6 shows 3D flux paths in a FEM simulation. Despite setting the flux tube starting points almost dead center on a pole face, only flux tubes originating on the pole shoe actually enter the stator core or clamping plate. Flux tubes originating further down (in radial direction) move directly from one pole to the neighboring pole. This is in strong contrast to the behavior observed in the fringing air-gap flux density as calculated using a Schwarz-Christoffel mapping presented in fig. 4. Here, all flux lines originating on the rotor surface meet the stator core or clamping plate" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0002200_978-3-319-51222-8-Figure3.46-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0002200_978-3-319-51222-8-Figure3.46-1.png", "caption": "Fig. 3.46 Plot of the nodal vectors gi of the functional J (uh) = \u03c3xx , the horizontal stress in the plate close to the opening", "texts": [], "surrounding_texts": [ "After the system K u = f has been solved, the FE-program computes the nodal forces at fixed nodes as follows: \u2022 It expands the vector u back to full size by incorporating the previously canceled degrees of freedom ui = 0 at the fixed nodes, u \u2192 uG , 202 3 Finite Elements \u2022 It multiplies the non-reduced global stiffness matrix KG with the vector uG ; \u2022 the entries fi in the vector f G = KG uG , which belong to the fixed degrees of freedom, are the nodal forces at the fixed nodes without that portion of the load which wanders directly into the fixed nodes. The latter support reactions we call Rd (see Fig. 3.48) fi (complete) = fi + Rd = RFE + Rd . (3.196) \u2022 If the support is an elastic support, then this last maneuver is not necessary, then fi = RFE is the full support reaction. How this split of the support reaction comes about we study by looking at a beam. 3.27 Support Reactions 203 204 3 Finite Elements" ] }, { "image_filename": "designv11_62_0000444_j.proeng.2012.01.592-Figure4-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0000444_j.proeng.2012.01.592-Figure4-1.png", "caption": "Fig 4. Real test field Fig5. Virtual environment Fig7. UGV Tracking results Fig6. UGV telemetry", "texts": [], "surrounding_texts": [ "First calibrate the UGV movement (moving and turning), in this case calibrate the real signal comes from encoder on UGV with the parameter in simulink model. For example when UGV moved forward one meter how much signal read by encoder in simulink or when UGV rotate how much signal that represents 90o degrees rotation in reality. From there the accuracy of encoder to the real condition revealed and can be setup to simulink model. This UGV use only relative localization for knowing its local position. Odometry method were used here, where encoder act as a sensor to solve it. Then place the real UGV and UGV model in the same coordinate for starter. After that, turn on the UGV and also run the simulation. Drive UGV through obstacle from the field to the finish spot without touching the field wall. After the UGV reach the finish spot, repeat it for different field and after the experiment finish, stop the software simulation and turn off the UGV. 6. Result and discussion monitoring kind of field that used for this test in virtual reality environment and a UGV model also in virtual.Fig.6 shown the condition monitoring from UGV during the test based on the slope inclination from the pendulum and the orientation from the encoder. The comparison of tracking result from UGV movement based on encoder reading in reality performance and in simulink during exploration field A and field B are shown in Fig.7. There Red dot represents as a UGV start point and purple dot as a finish point. The blue lines represent the real trajectory of UGV and the orange lines represent the movement of UGV in simulink model. Based on Fig.7 the blue line shapes smoother than the orange line. It means in reality the movement is smoother than in simulink. It caused by the sensitivity of the joystick in simulink while the test running and also from the friction and slip that may occur during the movement in reality." ] }, { "image_filename": "designv11_62_0000253_iciea.2013.6566377-Figure10-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0000253_iciea.2013.6566377-Figure10-1.png", "caption": "Figure 10. Inner-race fault with window size 1024 for 64 segments", "texts": [], "surrounding_texts": [ "The future of the diagnostic techniques will be self-learning and adaptive. Biologically inspired ANN classifiers with diverse learning capabilities are among the best candidates for such future approaches. A neural network for fault classification can be achieved efficiently due to the fact that neural network is a nonlinear empirical model which can capture the nonlinear system dynamics and do not require knowledge of particular system parameters [7]. Neural network design includes input layer, hidden layers and output layer as shown in Fig. 6. Number of input layer neurons is equal to the input features (7) and the number of output neurons is equal to number of classes. There can be one or more than one hidden layers with different number of neurons. Most of the time increasing the number of hidden layers neurons guarantees good learning but it requires more training time and computational cost. Therefore minimum number of the hidden layers and neurons meeting certain classification accuracy is preferred. Neural Network pattern classifier learns system dynamics in the form of weighted links between the neurons during the training phase. Depending upon problem-nature training can be of the supervised or unsupervised type [4, 9]. In unsupervised learning the targets for input features are unknown while for supervised learning targets are known. We are using supervised learning because for each spectral features training input we have associated targets. For output classification we are using one versus all classifier. Before starting the training the training data is randomly shuffled and divided into training, validation and test sets with different percentages. MSE (Mean Square Error) is used to measure the accuracy of the training, 2013 IEEE 8th Conference on Industrial Electronics and Applications (ICIEA) 263 validation and test sets. Validation set is used as check to stop learning and testing data set is used to measure over fitting of the classifier. Neurons in different layers are connected through weighted links. These weights are tuned during the training of the ANN as shown in Fig. 7. Weights are randomly initialized before starting the training. During training, with random weights, output of the network is used to calculate the MSE using target values. If the calculated MSE is of acceptable value then training can be stopped, otherwise based on the MSE, weights are tuned using back propagation algorithm. Thus, after training, ANN can be used for classification. IV. RESULTS AND DISCUSSION Methodology developed in the previous sections will now be tested practically. Fig. 8 illustrates the experimental setup for recording the actual vibration data sets [10]. Experiments were conducted with four types of bearings including one normal and three faulty bearings with faults in inner-race, ball and outer-race. Faults in the bearing were created by electro discharge machining. Faulty bearings are supporting the shaft of the motor and the load is 2HP with a speed of 1750 r/min. The data have been collected through accelerometers using a 16-channel digital-audio-tape recorder and sampled at the rate of 12000 samples per second. Time vibrations recorded and converted to spectral features for these four signals are shown in the Figs. 5, 9, 10 and 11 respectively with a window size of 1024 for 64 segments. 264 2013 IEEE 8th Conference on Industrial Electronics and Applications (ICIEA) In this paper we have used only one hidden layer with different number of neurons to experimentally check the best training that can be achieved with minimum number of neurons. Number of input features, 513, is equal to spectral features of the time segment. Output layer of the ANN contains four neurons because of four classes as shown in Fig. 12. For training ANN, the spectral contents from different signals with different window size are grouped into a set with respective class (8): ($ , % ) = {($ , % ), ($ , % ), ($ , % ), \u2026 , ($&, %&)} (8) Window size used in this experiment are 256, 512, 1024 and 2048 samples. In (8) p' represents input spectral pattern for any of the four classes with any specific window size. t' represents target class or output for this particular input. Four of the target classes used in this experiment are: *- 1; % = [1 0 0 0] Normal signal *- 2; % = [0 1 0 0] Inner-race fault *- 3; % = [0 0 1 0] Bearing fault *- 4; % = [0 0 0 1] Outer-race fault This grouped data was randomly shuffled and then divided into training, validation and test sets with a respective proportion of 60%, 20% and 20% of the total grouped contents given by (8). With random weight initialization, the network was trained using feed forward back propagation algorithm with different number of hidden layers neurons. Training the network, more than hundred times, for different number of hidden layer neurons, resulted in almost similar minimum MSE. Minimum number of hidden layer neurons that gave acceptable accuracy was 2. Accuracy achieved with two neurons in the hidden layer and more than two neurons are comparable. But using higher number of hidden layer neurons increases the computational time. Fig. 13 shows, as the number of hidden layer neuron increases the learning time increases almost exponentially. Thus 2 hidden layer neurons are best selection for this particular scenario. The appropriate ANN architecture selected is shown in the Fig. 12. As the vibration signals from bearing are quasi-stationary therefore the problem of optimum window selection has been addressed by using multi-sized time-domain segmentation-window for augmented feature selection. Classifier trained with these augmented spectral features has shown 100% accuracy complying with the supposition that smaller windows will capture the signatures appearing for short duration or higher frequencies and larger windows will address the low frequencies contents. The trained network is tested with all the window sizes, 256, 512, 1024 and 2048 2013 IEEE 8th Conference on Industrial Electronics and Applications (ICIEA) 265 with classification accuracy of 100%. Thus multi-window approach offers a range of windows to fit a variety of transients appearing in the vibration signal providing a comprehensive data set for training and testing of the real time rotary machines vibration. V. CONCLUSIONS In this paper, multi-size-window time segmentation based spectral features augmentation for neural network bearing fault classification has been presented. Augmented spectral features of vibration signal, in rotary machines, calculated using multi-size time-segmentation-window have been used to train and test the neural network classifier. Classification results have shown that classifier, with multisize-window time-segmented spectral features, has learned the dynamics of quasi-stationary vibration signals efficiently for real time scenarios with 100% accuracy." ] }, { "image_filename": "designv11_62_0002927_sbr-lars-r.2017.8215275-Figure3-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0002927_sbr-lars-r.2017.8215275-Figure3-1.png", "caption": "Fig. 3. Inertial frame and body frame with the angles roll, \u03c6, pitch, \u03b8 and yaw, \u03c8.", "texts": [ " Considering all these effects, the equations of motion for the AVALON aircraft are derived here as a combination of the kinematic and dynamic modelling. The model presented in the following subsections considers the aircraft as a rigid body with six degrees of freedom and assumes that the Centre of Gravity (CG) and the body-fixed frame origin coincide. The inertial frame is an Earth-fixed tangential frame in local-level and tangent to the gravity. The relationship between the relevant frames of reference can be seen in Fig. 3. The aircraft kinematic and dynamic is represented by (1)-(4), where pi is position in the inertial frame, ub = [U, V,W ]T is the velocity in the body frame, q is the quaternion representation of the attitude, \u03c9 = [P,Q,R]T is the angular velocity, Ci b is the rotation matrix from the body to the inertial frame, \u03a9\u03c9 is the skew symmetric form of \u03c9, f is the total applied force, t = [M,N,L]T is the total applied torque, J is the moment of inertia and m is the mass [16], [17]. p\u0307i = Cn b ub (1) u\u0307b = \u2212\u03c9 \u00d7 ub + 1 m f (2) q\u0307 = 1 2 \u03a9\u03c9q (3) \u03c9\u0307 = \u2212J\u22121 (\u2212\u03c9 \u00d7 J\u03c9) + J\u22121t (4) In our implementation the angular acceleration (4) is expanded into (5)-(7) with the introduction of the constants Ci with i = 0, " ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0000477_s00202-011-0204-8-Figure10-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0000477_s00202-011-0204-8-Figure10-1.png", "caption": "Fig. 10 Damaged rotor with Nb = 20 broken bars and 30 residual undamaged bars", "texts": [ " Due to the symmetry which characterises the considered typical undamaged rotor, the third structural eigenspace is degenerate (see discussion at the end of Sect. 2.2), containing four orthogonal vibration modes (last four rows of Table 5) at the same eigenfrequency (III1,2,3,4 in Table 4). These degenerate vibration modes imply only translations of the end-rings with respect to the rotor stack, whereas all the modal rotations are zero. Assuming now the presence of damage, consisting e.g. of 10 consecutive broken bars on the right end of the rotor, symmetrically placed with respect to x2 axis as shown in Fig. 10, the set of Eq. (6) must be employed for the dynamical anal- ysis. The computed eigenfrequencies of the damaged rotor are reported in Table 6 whereas the corresponding vibration modes are reported in Table 7. By comparing the vibration modes in Table 7 with those of the undamaged rotor in Table 5, it turns out that the presence of damage, degrading the symmetry characteristics of the rotor, yields a coupling between torsional and translational components, as it clearly appears in the first two rows of Table 7", " Accordingly, system (6) can be rewritten into the form (\u2212M\u03c92 + i\u039b\u03c9 + K)X = F (11) where X contains the lagrangian coordinate unknown amplitudes, M is the mass matrix, is the damping matrix, depending on \u03b6, K is the stiffness matrix and F is the load vector containing M. The expressions of these matrices is straightforward and is not here explicitly reported for the sake of brevity. Equation (11) is used to compute the harmonic response of rotors exhibiting an increasing number of broken bars on the right end of the rotor. The broken bars are assumed to be consecutive, as commonly experienced by the Ansaldobreda Company, and, without loss of generality, symmetrically disposed with respect to the x2 axis, as shown in Fig. 10. In Fig. 11, the amplitude of the total translation of the right end-ring is reported as a function of the excitation frequency \u03c9, whereas Fig. 12 shows the amplitude of the relative rotation \u03b8r \u2212 \u03b8s of the right end-ring with respect to the rotor stack. When Nb=0, i.e. no damage is present, the total translation amplitude is identically equal to zero and only a single resonance peak appears in the diagram of the relative rotation; in fact, in a perfect symmetric rotor, the torque acting on the rotor stack is able to excite only the second torsional eigenmode, which is the only eigenmode exhibiting a nonzero modal component \u03b8s (see Table 5)", " Due to the coupling between vibration modes, a low-frequency peak appears in Fig. 12, becoming larger and larger as the number of broken bars increases, whereas the peak relevant to the second torsional eigenmode is reduced. Moreover, three resonance peaks appear in the diagram of the total translation amplitude in Fig. 11, corresponding to the excited modes I, II and III, whose height increases with the increase of Nb. Finally, the Von Mises stress \u03c3id is computed at the interface between the bar indicated by an arrow in Fig. 10 and the end-ring r , in correspondence of the representative point P , located on a main diagonal of the bar cross section at a distance from the centroid equal to one quarter of the diagonal length, where both normal and shear stresses are present. The corresponding diagram, reported in Fig. 13, exhibits a behaviour similar to the one of the total translation and relative rotation previously described. In particular, the stress level due to the lowest frequency eigenmode greatly increases with the increase of Nb" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0002048_clei.2012.6427231-Figure3-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0002048_clei.2012.6427231-Figure3-1.png", "caption": "Figure 3: Robot i con direccio\u0301n tipo Ackermann, en posicio\u0301n (x, y) y direccio\u0301n en el a\u0301ngulo \u03b8. Tomado de [5].", "texts": [], "surrounding_texts": [ "El algoritmo propuesto esta\u0301 orientado espec\u0131\u0301ficamente para anomal\u0131\u0301as de gradiente, donde se toma como base el algoritmo bang-bang y se integra la te\u0301cnica cla\u0301sica de control Proporcional Integral Derivativo (PID)[21] sobre la direccio\u0301n del veh\u0131\u0301culo auto\u0301nomo con el fin de optimizar el movimiento al seguir la curva de nivel del gradiente. El control PID es una te\u0301cnica de control por retroalimentacio\u0301n cuyo objetivo es aplicar una accio\u0301n correctora de ajuste con base en los valores medidos en el tiempo. Este mecanismo ha sido ampliamente aplicado en sistemas de control de nivel en tanques de l\u0131\u0301quidos, presio\u0301n en tubos, y recientemente en control de direccio\u0301n de veh\u0131\u0301culos robo\u0301ticos auto\u0301nomos[22], [23]. En este algoritmo se toma ventaja del resultado de la diferencia del valor medido y un valor de referencia vref para controlar la direccio\u0301n del sensor robo\u0301tico; este valor resultante para cada iteracio\u0301n de control se denota como f (ecuacio\u0301n 4), que depende de la medicio\u0301n de la variable f\u0131\u0301sica a trave\u0301s del sensor, que se denota como s(k) y el valor de referencia vref , el cual determina el valor constante de curva de nivel que se va a seguir. f(k) = s(k) \u2212 vref (4) Al igual que en el algoritmo bang-bang, en cada iteracio\u0301n de control se calcula el a\u0301ngulo de la direccio\u0301n como lo describe la ecuacio\u0301n 2, pero en este caso, el ca\u0301lculo de u se realiza con base en la te\u0301cnica cla\u0301sica de control Proporcional Integral Derivativo para seguir la curva de nivel del gradiente. En la ecuacio\u0301n 5 se define el modelo de PID para el ca\u0301lculo del control de direccio\u0301n con base en la medicio\u0301n de la magnitud de la variable f\u0131\u0301sica. u(k) = \u03b1f(k) + \u03b2 d dk f(k) + \u03b3 \u222b f(k)dk (5) Donde \u03b1 es la constante de proporcionalidad que define que tanto se cambia la direccio\u0301n con respecto a f , \u03b2 es la constante derivativa y \u03b3 es la constante integradora del error acumulado en el tiempo. Para calcular la derivada de f(k) se tiene en cuenta la diferencia el valor de f(k \u2212 1) y para la integral simplemente se suman todos los errores medidos en cada iteracio\u0301n de control. V. IMPLEMENTACIO\u0301N Los algoritmos bang-bang, bang-bang mejorado y el propuesto en este art\u0131\u0301culo han sido implementados y ejecutados en el simulador de redes de sensores mo\u0301viles MobSim[24]. MobSim es una plataforma modular que reduce el acoplamiento entre mo\u0301dulos para facilitar su trabajo como simulador o interfaz de comunicacio\u0301n con robots f\u0131\u0301sicos. De esta forma permite agilizar el desarrollo y pruebas de algoritmos aplicados a redes de sensores robo\u0301ticos al ofrecer una plataforma intermedia y robusta que le ofrece al desarrollador programar robots f\u0131\u0301sicos o simulados a la vez. Dentro de su estructura, MobSim requiere herramientas de desarrollo como: el lenguaje de programacio\u0301n Java, la plataforma modular OSGi, el framework de agentes JADE y la librer\u0131\u0301a de gra\u0301ficos vectoriales Apache-Batik. Es destacable que sobre esta plataforma de desarrollo MobSim ya se han implementado varios algoritmos para deteccio\u0301n de per\u0131\u0301metros y te\u0301cnicas de control robo\u0301tico, ya que este simula condiciones ambientales parametrizables, donde por defecto la temperatura del ambiente son 25C y la anomal\u0131\u0301a se puede conformar por mu\u0301ltiples gradientes que van aumentando hasta llegar a un ma\u0301ximo de 40C. Bajo estas condiciones, la plataforma incluye robots mo\u0301viles con sistema de movimiento diferencial u omnidireccional, los cuales pueden medir la temperatura del ambiente y de los gradientes que hayan sido definidos previamente.[24]. Para evaluar el desempeo de los dos algoritmos bangbang y el algoritmo propuesto, se programan mo\u0301dulos de control distribuido, donde cada agente de software ejecuta un comportamiento c\u0131\u0301clico cada 100ms, que es encargado de controlar el movimiento del sensor robo\u0301tico. De esta forma, los agentes de software buscan seguir la curva de nivel de una anomal\u0131\u0301a de forma irregular conformada por mu\u0301ltiples gradientes en conjunto utilizando el sensor de temperatura simulado. El valor de la curva de nivel para seguir es de 30 y se corren los tres algoritmos de seguimiento sobre la misma anomal\u0131\u0301a para evaluar el comportamiento de un sensor robo\u0301tico, partiendo de la misma ubicacio\u0301n y finalizando cuando terminan de recorrer todo el entorno. En la figura 7 se presenta la trayectoria obtenida al correr (a) el algoritmo bang-bang y (b) el algoritmo bang-bang mejorado. La trayectoria del algoritmo propuesto se puede observar en la figura 8. Cada implementacio\u0301n de algoritmos de seguimiento fueron desarrollados en el lenguaje de programacio\u0301n Java usando la librer\u0131\u0301a de MobSim-API y cada desarrollo se integra como un mo\u0301dulo sobre la plataforma OSGi[25]." ] }, { "image_filename": "designv11_62_0002100_978-1-4419-9834-7_127-Figure1-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0002100_978-1-4419-9834-7_127-Figure1-1.png", "caption": "Figure 1. Flapping mechanism and its schematics.", "texts": [ " Computational analyses are performed on the same wing configuration using a combined nonlinear structural dynamics and Navier-Stokes solution. The main objectives of this paper are to: a) discuss the development of an integrated experimental and computational approach to analyze flapping wing configurations, and b) to show some preliminary comparisons of flow structures and wing deformation between experimental data and computational response. Experimental Setup A single-degree-of-freedom flapping mechanism is designed and built for this study, as shown in Figure 1 (the same mechanism as the one used in reference [18]; but the wings shown in the figure are different from the aluminum wings tested). The design is created based on a Maxon motor system that includes a 15 W brushless DC motor EC16, a 57/13 reduction ratio planetary gear head, a 256 counts-per-turn encoder and an EPOS 24 controller. This system provides precise control of the motor system: the sensor provides position and velocity feedback to the controller that actively regulates the motor. Utilizing the high precision pre-assembled planetary gear head rather than constructing a custom gear transmission is also advantageous. The final output range of the motor shaft is: speed 0 to 45 revolutions per second (RPS) and nominal torque 0 to 21 N\u2022mm. The rotation output from the motor is first transformed into a reciprocating motion with a crank-slider mechanism; then a bar linkage mechanism realizes the flapping motion at the wing mount. A detailed schematic description of the flapping kinematics is presented in Figure 1. The geometric relationship between motor rotation (angle ) and flap angle is expressed in the equations in the figure, where is the flap angle; is the motor rotational angle; x is the vertical displacement from the center point when the wings are horizontally positioned. The rest of the parameters are selected so that a \u00b121 \u00ba amplitude is maintained. The experiments are performed at 10 Hz flapping frequency. The isotropic wings tested in this study are made from 0.4 mm thick aluminum sheets. The wing planform is of a 7" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0002329_s40997-017-0082-4-Figure9-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0002329_s40997-017-0082-4-Figure9-1.png", "caption": "Fig. 9 a Assembled shaft and bush. b Mesh size of the shaft used for interference modeling", "texts": [ " Standard deviation of this fitting procedure was regulated on less than 1 lm. The shaft and bush were modeled using 10-node modified quadratic tetrahedron elements. Applied mesh size was adequately fine to obtain local stress gradients in the contact surface of the parts, properly. Independence of the results from the mesh size was investigated, too. The average number of elements to meshing the joint parts was 70,000. The minimum dimension of the elements of shaft surface was 0.03 mm. Assembled parts of joints and sample algorithm of the shaft mesh are shown in Fig. 9. As noticed, we have tried to model real geometry of the shaft, so mesh of the shaft is not symmetric. The added part to the end of the shaft was used to apply proper boundary conditions. Figure 10 represents effective von Mises (vM) stress variation on the shaft with and without defects. Considerable variations on the stress values can be seen in this figure, which shows stress alteration from zero in some points to about 350 MPa in others. As can be seen, defects change the local distribution of the stresses, while in perfect shaft, there are no considerable changes except in the edges" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0000450_esda2012-82563-Figure6-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0000450_esda2012-82563-Figure6-1.png", "caption": "FIGURE 6. Rotor model with limit conditions", "texts": [ " In this work, a high speed and high power milling spindle is studied: 100kW and 30000rpm. The rotor is mounted on two tandems of hybrid bearings (characteristics in Tab. 1). The tool-holder is considered rigidly linked to the rotor. Loads on bearings are axial and correspond to the half of the preload. Its value corresponds to the stiff preload recommended by the bearing manufacturer: Pr = 2440N. It is applied by springs on the rear tandem of bearing. That is why the preload can be considered as constant in the model. As shown in Fig. 6, each bearing is taken into account with 3 translational and 2 rotational springs. Stiffness values correspond to the diagonal terms of the stiffness matrix K. In this case, 3 Copyright \u00a9 2012 by ASME Downloaded From: http://proceedings.asmedigitalcollection.asme.org/ on 01/30/2016 Terms of Use: http://www.asme.org/about-asme/terms-of-use extra-diagonal term of the stiffness matrix are null or negligible but they could be taken into account for a multi-DOF loading as discussed in [9]. This model is transferred to a 3D Finite Elements and multi body modeling software : LMS VIRTUAL LAB" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0003574_icnsc.2019.8743340-Figure3-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0003574_icnsc.2019.8743340-Figure3-1.png", "caption": "Fig. 3. Schematic diagram of robot modelling and identification of threat zone.", "texts": [ " The radius is a reflection of the shape and size of the robot, and is denoted by r. This paper sets the virtual obstacle on the boundary of the circle, which regards the force exerted by the virtual obstacle on the robot as the action taken by the robot itself to get rid of the local minimum trap. In order to effectively make the calculation of the route point get rid of the local minimum trap, the line connecting the virtual obstacle and the current route point should be perpendicular to the line connecting the current route point and the goal. As shown in Fig.3, the position 1and 2 is the alternative locations of the virtual obstacle. ROn and RGn denote two vectors pointing from the robot to the obstacle and from the robot to the goal, respectively. When the calculation of the route point falls into the local minimum trap, the placement of the location of the virtual obstacle need to take the locations of the real obstacles into consideration. For the real obstacles to consider, the angle between their ROn and RGn are within range of (0 ,90 ] , which form the threat area", " Step2: Calculate the number of the obstacles of the clockwise threat area and the counter-clockwise threat area respectively, denoted by Ln and Rn respectively, and the shortest distance between the robot and the obstacles in these two threat areas respectively, denoted by _ minLd and _ minRd respectively. Step3: Determine the location of the virtual obstacle. Case 1: L Rn n> , place the virtual obstacle in position 1. Case 2: L Rn n< , place the virtual obstacle in position 2. Case 3: =L Rn n and _ min _ minL Rd d> , place the virtual obstacle in position 2. Case 4: =L Rn n and _ min _ minL Rd d\u2264 , place the virtual obstacle in position 1. This paper presents a specific method for calculating the location of the virtual obstacle point. As shown in Fig.3, the point of intersection of RGn and the circle is ( , )T a a aq x y= . Assume that the current position of the robot ( , )T robot R Rq x y= is the coordinate origin, and then RG a RG nq r n = \u00d7 . The position of the virtual obstacle is denoted by ( , )T vo vo voq x y= . If the virtual obstacle is placed in the position 1, then cos sin sin cos vo a R vo a R x x x y y y \u03b8 \u03b8 \u03b8 \u03b8 \u2212 = + (3) If the virtual obstacle is placed in the position 2, then cos sin sin cos vo a R vo a R x x x y y y \u03b8 \u03b8 \u03b8 \u03b8 = + \u2212 (4) where = 90 " ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0002176_j.procs.2017.01.179-Figure3-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0002176_j.procs.2017.01.179-Figure3-1.png", "caption": "Fig. 3. Structure of a rod.", "texts": [ " Section 2 describes the aspect of the prototype robot and the experimental result of basic motion control. Section 3 demonstrates the calculation method of the balancing internal force. Finally, Section 4 concludes the discussion. Fig. 2 shows the proposed prototype, which consists of three rods and nine wires. Although the lengths of the rods are fixed, the lengths and/or tension of the nine wires can be changed using the nine actuators in the three rods. The power supply is given through electric cables from the external environment. Figure 3 shows the structure of a rod, which is made of an aluminum frame (50 \u00d7 50 \u00d7 600 [mm], 950 [g]). Each rod has three actuators inside to reel three wires, and has six wires entirely: three actuated and three unactuated wires. Notably, each wire is shared among two rods. Two of the three actuated wires are assigned at one tip, while the other actuated wire is assigned at the other tip. The three unactuated wires are simply attached with the tips of each rod, the lengths of which can be changed by the actuators in the other rods. The actuator consists of a reeling pulley made of aluminum, a bevel gear pair with gear ratio is 2:1, and a DC servo motor with an encoder, as depicted in Fig. 3 (b). The wire lengths can be measured using the encoders in the actuators. As the wires\u2019 directions are variable depending on the positions and orientations of the three rods, the tip is equipped with three small guiding pulleys (10 [mm] in diameter) and a guiding ring (30 [mm] in diameter) to follow the wires\u2019 directions, as shown in Figs. 3 (b) and (c). A pilot experiment was conducted to control the shape of the prototype robot. This experiment employed a simple PD controller in the wire-length coordinates [8]", " After calculating the desired wire lengths corresponding to the required positions/orientations of the rods through inverse kinematics, the wire tension input is given as \u03b1 = KP ( qd \u2212 q ) \u2212 KV q\u0307 + v + vg, (1) where the vector \u03b1 \u2208 R9 indicates the actuator input as the wire tension. In addition, the vector qd \u2208 R9 is the desired wire length, and the vector q \u2208 R9 indicates the wire lengths measured by the encoders in real time. The matrices KP \u2208 R9\u00d79 and KV \u2208 R9\u00d79 are diagonally positive matrices, implying feedback gains. The term v \u2208 R9 implies the internal force applied to avoid slackening of the wires. The final term vg \u2208 R9 is the gravity compensation. Fig. 4 illustrates a resultant motion of the pilot experiment. Fig. 3 shows the behavior of the wire lengths q1, . . . , q9 measured by the encoders. In this pilot experiment, it was presumed that three wires and a rod were connected at a single fixed point at the rod tip to easily calculate the inverse kinematics. Moreover, an approximately same amount of tension was provided to each wire as the balancing internal force. Fig. 5 shows that the resultant trajectories slightly deviated from the required trajectories. This may be attributed to the errors caused by factors such as the inverse kinematics, gravity compensation, internal force, and friction" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0000319_isas.2011.5960923-Figure3-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0000319_isas.2011.5960923-Figure3-1.png", "caption": "Fig. 3. Virtual force against a voronoi neighbor.", "texts": [ " (3) In Fig. 2, db is the shortest distance from Pi to the boundary and Pb is the point which is symmetry with respect to the boundary. Suppose the distance (dij) between Pi and Pb is shorter than dave and applying (2) gives, IF;bl = dave - dib. 2 (4) Notice that dib = 2db then IFbl is given in same formula as (2). This implies that the virtual force near the boundary corresponds to the force from the virtual axisymmetrical node. Suppose the case where three nodes are initially deployed as shown in Fig.3. Fij whose magnitude is exerted on Pi from the node Pj whose distance is shorter than dave. The magnitude f F\ufffd . d -d I h\u00b7 \u00b71 . h h o ij IS avc2 'J . n tis case, Pi moves unt! It reac es t e perpendicular bisector of Pj and PH 1. III. PROPOSED ALGORITHM Suppose that three nodes, a leader robot, a relay robot and a base station are initially deployed as shown in Fig. 4. The leader robot is driven manually by an operator and the position of the base station is fixed. VEC algorithm is applied to the control of the relay robot" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0003741_012077-Figure10-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0003741_012077-Figure10-1.png", "caption": "Figure 10. The second folding mechanism in: (a) compact configuration; (b) partially deployed configuration; (c) totally deployed configuration.", "texts": [ "1088/1757-899X/591/1/012077 The second mechanism discussed here is shown in figure 5. It has the total number of links, n = 8 and the number of rotary joints, g1 = 13. As we may see in the example of its application, this Modern Technologies in Industrial Engineering VII, (ModTech2019) IOP Conf. Series: Materials Science and Engineering 591 (2019) 012077 IOP Publishing doi:10.1088/1757-899X/591/1/012077 mechanism may allow to built deployable structures (mobile house, for example) with shading system (see figure 10). According to Grubler formula, for n = 10 and g1 = 13, the number of degrees of freedom of this mechanism is: 11321103 F . (4) Again, a single actuator is needed (placed to A joint) to actuate the mechanism, which has as advantage a minimum number of actuators for the entire spatial structure, a smaller cost, a simpler Modern Technologies in Industrial Engineering VII, (ModTech2019) IOP Conf. Series: Materials Science and Engineering 591 (2019) 012077 IOP Publishing doi:10.1088/1757-899X/591/1/012077 control algorithm, and so on" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0003613_012066-Figure4-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0003613_012066-Figure4-1.png", "caption": "Figure 4. Contour of von-Mises stress on faultless coil spring given the minimum load.", "texts": [], "surrounding_texts": [ "Transport vehicles require a good suspension system to dampen vibration, swings and shocks received as they travel along bumpy, hollow, and uneven roads [1]. These conditions are very uncomfortable and may cause accidents. The suspension is also expected to hold the load during some common vehicle maneuvers such as acceleration, braking or deflection while on the road [2]. The coil spring is one of the main components for dampening vibrations and shocks to the load so as to provide comfort and security while the vehicle is in motion [3]. Depending on the condition of their application, coil springs often sustain fatigue failure. This indicates that the tension received below by the coil spring from the maximum stress of the material while sustaining a dynamic load causes fatigue failure [4-8]. The yield strength of the material is also a criterion of failure. Components of automotive suspension must be changed with a traveling distance of 73,500 km, or every five years [9]. The fault of 13.18 % of 24.2 million vehicle tests was recorded [10]. With the development of computing technology, the numerical analysis method has become particularly suitable for use because it will increase the calculation efficiency, the cost-effectiveness as well as save time. Various numerical analysis methods are widely available, but the finite element analysis (FEA) has proven to be reliable in solving problems in the field of continuum mechanics [11]." ] }, { "image_filename": "designv11_62_0002498_s12555-016-0112-9-Figure1-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0002498_s12555-016-0112-9-Figure1-1.png", "caption": "Fig. 1. Prototype of eight-rotor MAV.", "texts": [ " Then the number of effective singular can be determined according to its critical point, and the type-2 fuzzy model is constructed with rules located by TLS decomposition. This paper is structured as followings: In Section 2, the mathematical nonlinear model of the eight-rotor MAV is developed. The flight control algorithm is given in Section 3, which is also devoted to the stability analysis of the control scheme. The simulation is given in section 4. Platform description and some experiences are given in Section 5 and Section 6 contains concluding remarks. Fig. 1 shows the prototype of eight-rotor MAV. Equation (1) describes the kinematics of a generic 6 degree of freedom rigid-body: \u03b7\u0307 = [ R 03\u00d73 03\u00d73 T ] v, (1) where the \u03b7 means orientation and position of eight-rotor MAV with respect to inertial reference frame; while the v means the linear and angular velocities of orientation and position of eight-rotor MAV with respect to the bodyfixed frame,x, y, z represents the linear positions, and \u03d5 , \u03b8 , \u03c8 represents the roll, pitch and yaw angles in inertial reference frame respectively" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0003913_s00773-019-00675-8-Figure8-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0003913_s00773-019-00675-8-Figure8-1.png", "caption": "Fig. 8 Grid system for the turning simulation. a Mesh distribution in the computational domain. b Mesh refinement in the stern of the ship. c Over-set grid during rudder rotation", "texts": [ " In the first step, the turning behavior of the ship that determines the drift angle, ship speed, and propeller revolution speed is analyzed. These factors are used as input parameters to the second step where the ship self-propulsion in the oblique flow is simulated. The propeller can be modelled by a simple virtual disk in the first step, whereas the real propeller geometry is used in the second step. The virtual disk method is widely used in marine flow problems related to ship-hull interaction. The fluid within the region of the virtual disk is imposed a mass force in accordance with the predefined open water characteristic. Figure\u00a08 displays the computational mesh for the ship turning simulation. To capture the flow related to the propeller-rudder interaction, the mesh near the virtual disk and rudder is refined, as illustrated in Fig.\u00a08b. The ship motion and rudder rotation are handled using the over-set grid (Fig.\u00a08c). Figure\u00a09 depicts the computed time histories of the drift angle and ship speed during the turning motion with a right rudder angle of 35\u00b0. In the first turning phase, the drift angle increases rapidly. The slope of the drift angle curve begins to decrease after 40\u00a0s. The maximum drift angle is near 25.8\u00b0. Three typical phase times are considered in the oblique flow analysis, as presented in Table\u00a03. These phase times correspond to the drift angles of 10\u00b0, 20\u00b0, and near maximum. The real propeller geometry is used in the oblique flow condition" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0003021_imece2017-70403-Figure6-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0003021_imece2017-70403-Figure6-1.png", "caption": "Figure 6. Configuration of the compound planetary gear set", "texts": [ " (a) tooth width b=20mm It is indicated that impact of the tooth coupling stiffness to the mesh stiffness is more remarkable than that of the gear body coupling stiffness. For the paper length limitation, numerical results for time varying mesh stiffnesses by the biaxial coupling model are not given herein. NUMERICAL EXAMPLE AND DISCUSSION For trade-off studies, five optional layouts of the compound planetary gear train taken into consideration are for the specified full hybrid transmission, one of which is well matched in a middle sized passenger car. Basic configuration of the planetary gear train is shown in Fig. 6 and the five combinations of gears with different tooth parameters are given in Table 1. Table 1. Gear parameters of different layouts for the CPGT parameters layout1 layout2 layout3 layout4 layout5 mn (mm) 1.5 1.25 1.2 1.25 1.18 A (\u00b0) 17.5 19.25 19.25 19 21 (\u00b0) 20 17.15 23.75 23 23 ZS2 31 38 38 38 37 Zpb 18 23 23 23 24 Zpa 25 31 31 31 31 ZS1 23 28 28 28 32 5 Copyright \u00a9 2017 ASME Downloaded From: http://proceedings.asmedigitalcollection.asme.org/ on 01/16/2018 Terms of Use: http://www.asme.org/about-asme/terms-of-use For comparisons numerical results of mesh stiffnesses for helical gear pairs by SM and BM are produced and the BM is proved to be much accurate" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0003741_012077-Figure9-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0003741_012077-Figure9-1.png", "caption": "Figure 9. The paths of B, G and H nodes, during folding.", "texts": [ " The driving link 1 could only rotate with an angle 321 degree (see figure 6), starting from the extended configuration. It means that the simulated mechanism could only fold till the partially deployed configuration represented in figure 1(b). Also, the trajectory of the G and H nodes are not yet straight lines, as they should be, but as they are presented in figure 7, even if the link 7 has not a significant rotation around z axes, as seen in figure 8. The paths of the A, G and H nodes are shown in figure 9. Modern Technologies in Industrial Engineering VII, (ModTech2019) IOP Conf. Series: Materials Science and Engineering 591 (2019) 012077 IOP Publishing doi:10.1088/1757-899X/591/1/012077 Modern Technologies in Industrial Engineering VII, (ModTech2019) IOP Conf. Series: Materials Science and Engineering 591 (2019) 012077 IOP Publishing doi:10.1088/1757-899X/591/1/012077 Modern Technologies in Industrial Engineering VII, (ModTech2019) IOP Conf. Series: Materials Science and Engineering 591 (2019) 012077 IOP Publishing doi:10" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0002302_b978-0-12-803581-8.10351-0-Figure80-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0002302_b978-0-12-803581-8.10351-0-Figure80-1.png", "caption": "Fig. 80 Forward tool parting plane concept extended to side tool.", "texts": [ " In: 36th International SAMPE Technical Conference, San Diego, CA, November 15\u201318, 2004, pp.1063\u20131077. As described previously, if split lines are necessary for part or mold removal, then deciding on locations is not always an easy task. For any surface a line splits, complex or not, a subsequent backing plate must be constructed to match this surface exactly. Extending the concept tested on the demonstration tool, a flange on the forward (nose attachment surface) side shell tool surface would start to like the image of Fig. 80. To create this mold flange the tooling was modified with temporary tooling plates added to the vertical surface of the side shell tool as shown in Fig. 81. While this seemed like a simple solution, the fillet at the base of the tool, at the vacuum flange/tool wall intersection became the first complication. As seen in Fig. 81 this fillet meant that the simple tooling board could not come down to meet the mold vacuum flange and a solution was required. The solution to this gap between the tooling board and the mold vacuum surface, and ultimately to more complicated split line geometries, was the use of microcrystalline wax as had been used in forming the more complex mold flanges as described in the previous chapter" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0000704_ssp.165.359-Figure1-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0000704_ssp.165.359-Figure1-1.png", "caption": "Fig. 1. Stacker reclaimer with rail undercarriage", "texts": [ " (3) It is also possible to determine the ratio of relative systematic measurement errors of the sensor at left and right rail: + \u2212 == + + c c l r r l l r l r xb xb R R G G R R I I I I I I 2 2 1 1 \u03b4 \u03b4 . (4) From measurements in position III: ( ) ( ) ( ) + \u2212 + = + + + + = + =+ c c l r lr r l l r r r r lr rc xb xb R R RR R RR R GG G b y I I IIIIII III IIIIII III IIIIII III 2 2 11 1 2 1 \u03b4\u03b4 \u03b4 . (5) Final equations: b xb xb R R RR R y b RRRR RR x c c l r lr r c llrr lr c I I IIIIII III IIIIII III \u2212 + \u2212 + = \u2212 +++ + = 2 1 2 2 , 2 1 . (6) The above method was used to determine the centre of gravity of LZKS 1600 stacker reclaimer (Fig. 1) with the 800 Mg mass. 9 measurement series were conducted for superstructure position in relation to undercarriage described in Fig. 4 as well as for three positions of bucket wheel boom. In each series at least 6 measurements were carried out. The example run of measurement signal at the machine driving forward for position I with bucket wheel boom at lower position is shown in Fig. 8. Fig. 9 presents the positions of centres of gravity for different boom orientations in relation to pitch diameter of a slewing bearing" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0002462_acc.2017.7963286-Figure1-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0002462_acc.2017.7963286-Figure1-1.png", "caption": "Figure 1. Three-dimensional missile-target engagement geometry", "texts": [ " A sufficient condition for both FTB and IO-FTS of a class of nonlinear systems is introduced in section IV, and then the new anti-saturation guidance law design with finite time convergence is provided. In section V, simulation results illustrate the effectiveness of the new guidance law design. Some concluding remarks are given in Section VI. II. PROBLEM STATEMENT Consider the relative motion between a missile and a maneuvering target in the three-dimensional spherical coordinates , ,( )r with origin fixed at the missile\u2019s centroid as shown in Fig.1, where ( , , )re e e are the unit vectors whose directions are the same as , ,r . M and T are the 978-1-5090-5992-8/$31.00 \u00a92017 AACC 2243 centroids of the missile and target respectively, r denotes the relative range between the missile and target, and represent the azimuths of LOS. Assume that the acceleration components of the missile and target are ( , , )Mr M Ma a a and ( , , )Tr T Ta a a respectively. The relative motion can be described by Eq. (1) [17-18]: 2 2 2= cos Tr Mrr r r a a (1a) cos 2 cos 2 sin T Mr r r a a (1b) 22 sin cos T Mr r r a a (1c) The missile autopilot in the two directions of and is assumed to have the first-order dynamics [19]: 11 1M Ma a u (2a) 21 1M Ma a u (2b) where is the time constant, and 1 2, T u uu is the control input to the autopilot" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0002500_icra.2017.7989667-Figure2-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0002500_icra.2017.7989667-Figure2-1.png", "caption": "Fig. 2: The robot frames and the controllable variables [22].", "texts": [ " (4) An RGB-D camera is attached to the tip of the robot so the robot is able to sense the surroundings as it moves, called a tip camera in this paper. (5) The robot arm is initially positioned, by human assis- tance, outside of the unknown task space with its tip camera facing the target object. Some typical environments and tasks are shown in Fig. 6. A general n-section continuum manipulator is used. Each section seci, i=1, ..., n is characterized by its three controllable variables: length si, curvature \u03bai, and orientation \u03c6i, as illustrated in Fig. 2(b). Each seci can be positioned by specifying its base point pi\u22121. The tip point pi of seci is computed using its corresponding base point and controllable variables. Adjacent sections are connected tangentially as shown in Fig. 2(a). The arm configuration of an nsection continuum manipulator can be formulated as C = {(s1, \u03ba1, \u03c61), ...., (sn, \u03ban, \u03c6n)} [20], [23]. This kinematic model is widely used in the literature thanks to its generality [23]\u2013[26]. Note that the arm can in fact have infinite degrees of freedom because it can deform upon contact. However, for the task of object modeling, the arm is not planned to contact the target object or obstacles. The finite number of controllable variables are sufficient to describe the collisionfree motion of the arm" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0003959_j.conengprac.2019.104161-Figure1-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0003959_j.conengprac.2019.104161-Figure1-1.png", "caption": "Fig. 1. Surgical robot design.", "texts": [ " In this step, the design requirements include the selection of the DoF (2 for the mobile platform and 3 for each robotic manipulator), the determination of the mechanical dimensions (to define the working space for the RS), the class of actuators (direct current motors) and sensors (rotatory and reflexive distance sensors) to be used and the analysis of the materials (including the evaluation of mechanical stress, unitary displacements and torsional resistance). In the design of the RS, the number of joints was determined considering the regular surgical space. However, not all the possible movements exerted by a human surgeon were taken into account for designing purposes and simplifying the electronic instrumentation. Fig. 1 demonstrates the general configuration of the bi-arm RS. The left arm carries the medical scalpel while the right one plays the role of the holder for the target tissue. For the mechanical dimensions, the RS considered the working space that has been identified in previous medical robots such as DaVinci\u00ae, Phantom\u00ae and ROBODoc\u00ae (Li & Taylor, 2004; Taylor, Menciassi, Fichtinger, & Dario, 2008). The selected dimensions are shown in Fig. 1. Once the dimensions were fixed and the number of DoF was chosen, the selected construction material was poly-lactic acid (PLA), due to its physical characteristics (mainly low-density polymer) and the easiness of the manufacturing processes (Athanasiou, Niederauer, & Agrawal, 1996). The pieces for assembling the RS were manufactured by using three-dimensional printing technology, which permits the fabrication of pieces made in a computer assisted design (CAD) software with a lower cost than other comparative materials and manufacturing processes (Bogue, 2013) such as computer numerical control (CNC) metal pieces and plastic injection", " At the end of this stage, the RS moves all the way back to its home position. From there, the automatic controlled cutting steps can be started once again. This sequence defines the hybrid nature of the control problem statement. Considering that the proposed mechanical structure associated to the mobile robotic manipulator with the scalpel, let consider that \ud835\udc5e \u2208 R2+\ud835\udc5b with \ud835\udc5b = 4 corresponds to the number of cylindrical joints. Here \ud835\udc43 \u2208 R3 defines the position of the scalpel tip placed at the distal section of the manipulator (Fig. 1). The position of the distal tool \ud835\udc43 relates to the joint configuration \ud835\udc5e by a suitable nonlinear (dipheomorphism) function \ud835\udee4 (direct kinematics), namely \ud835\udc43 = \ud835\udee4 (\ud835\udc5e) where \ud835\udee4 \u2236 R5 \u2192 R3. The realization of the mathematical model in this study uses the approach presented in Murray (2017). Let consider \ud835\udc070\ud835\udc5d \u2208 \ud835\udc46\ud835\udc42(3) as the homogeneous transformation denoting the position and orientation of the RS base platform. Notice that orientation can change due to the lateral forces exerted by the tissue contact with the scalpel" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0002480_978-3-319-60399-5_7-Figure6-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0002480_978-3-319-60399-5_7-Figure6-1.png", "caption": "Fig. 6 Double-enveloping worm gear modified by the AU method", "texts": [ " If the deviation close to its natural modification is provided at cutting of the worm thread, the time of running-in is multiply reduced and the gear can operate under maximum load from the very beginning. Various methods for machine-tool modification of the double-enveloping worm thread during its cutting are known. The most common of them is the method in which the modification law that is closest to the assigned one is provided by a simultaneous deviation of the center distance and the gear ratio from nominal values to greater ones. Since both flanks of the thread in this case are modified at one machine-tool setting, it is called the \u201cAU double-flank correction-free method.\u201d Figure 6 shows the schemes of the machine-tool meshing for cutting the double-enveloping worm by the AU method and the curve line of the modification law, with its maximum value at the input DS and extreme value DS0 at the point u \u00bc us. When calculating modification by the AU method, the following parameters are determined: the tooth number of the generating gearwheel z20, the increase in the machine-tool center distance at worm cutting Da0 = a0 \u2212 aw, the pitch diameter of the generating gearwheel d0, and the diameter of its profile circle Dp0" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0003514_978-3-030-20131-9_324-Figure2.1-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0003514_978-3-030-20131-9_324-Figure2.1-1.png", "caption": "Fig. 2.1. The over-actuated compliant joint.", "texts": [], "surrounding_texts": [ "joint The internal articulated structure is designed according the \u201ccompliant mechanism\u201d concept so that the required mobility of the link (crank) (1) is obtained by means of elastic joint (fig. 1) [3]. A helical spring (2) can be taken as a compliant element and the required movements of the joint are then obtained by means of bending, expanding or contracting EAPs (3). The actuators are attached to the base (4) by their one end while by another to the crank through some circularly and symmetrically arranged levers (5), which allows using large number of actuators and control of the developed moment by adjusting the lengths of the levers. The symmetrical arrangement of actuators also simplifies calculations and makes the model more predictable [6]. Depending on the acting mode of EAPs, we can reach the necessary rotations of the joint in all three dimensions (frontal, sagittal, lateral) and one translation of the link in sagittal plane, which is needed for providing safety and regulation of the mechanism. The required capacities of actuators are determined by dynamical modeling of the compliant joint. The simplified dynamic model (damping and the inertial couplings between the links and the actuators are neglected) of the latter (Fig. 1) can be de- scribed by the following equations [7]: \ud835\udc3c\ud835\udc57(\ud835\udc5e\ud835\udc57)?\u0308?\ud835\udc57 + \ud835\udc50(\ud835\udc5e\ud835\udc57 , ?\u0307?\ud835\udc57) + \ud835\udc54(\ud835\udc5e\ud835\udc57) + \ud835\udc3e(\ud835\udc5e\ud835\udc57 \u2212 \ud835\udc5e\ud835\udc4e) = 0, (1) \ud835\udc3c\ud835\udc4e?\u0308?\ud835\udc4e + \ud835\udc3e(\ud835\udc5e\ud835\udc4e \u2212 \ud835\udc5e\ud835\udc57) = \ud835\udf0f, (2) where \ud835\udc5e\ud835\udc57 \u2208 \ud835\udc45 is the vector of joint angular positions, \ud835\udc5e\ud835\udc4e \u2208 \ud835\udc45 is the vector of actuators positions, \ud835\udc3c\ud835\udc57 is the inertia matrix of joints, \ud835\udc3c\ud835\udc4e is the inertia matrix of actuators, \ud835\udc50(\ud835\udc5e\ud835\udc57 , ?\u0307?\ud835\udc57) is the vector of Coriolis and centripetal torques, \ud835\udc54(\ud835\udc5e\ud835\udc57) is the vector of gravity torques, \ud835\udc3e is the stiffness matrix, \ud835\udf0f is the torque or vector of input signals. N. Zakaryan et al.1782 Solving (1) for \ud835\udc5e\ud835\udc5e\ud835\udc4e\ud835\udc4e and differentiating twice, we get an expression for ?\u0308?\ud835\udc5e\ud835\udc4e\ud835\udc4e. Adding (1) to (2), and inserting the expression for ?\u0308?\ud835\udc5e\ud835\udc4e\ud835\udc4e yields: \ud835\udf0f\ud835\udf0f = (\ud835\udc3c\ud835\udc3c\ud835\udc57\ud835\udc57(\ud835\udc5e\ud835\udc5e\ud835\udc57\ud835\udc57) + \ud835\udc3c\ud835\udc3c\ud835\udc4e\ud835\udc4e)?\u0308?\ud835\udc5e\ud835\udc57\ud835\udc57 + \ud835\udc50\ud835\udc50(\ud835\udc5e\ud835\udc5e\ud835\udc57\ud835\udc57, ?\u0307?\ud835\udc5e\ud835\udc57\ud835\udc57) + \ud835\udc54\ud835\udc54(\ud835\udc5e\ud835\udc5e\ud835\udc57\ud835\udc57) + \ud835\udc3c\ud835\udc3c\ud835\udc4e\ud835\udc4e\ud835\udc3e\ud835\udc3e\u22121\ud835\udc37\ud835\udc37\u22121 [\ud835\udc3c\ud835\udc3c?\u0308?\ud835\udc57(\ud835\udc5e\ud835\udc5e\ud835\udc57\ud835\udc57, ?\u0307?\ud835\udc5e\ud835\udc57\ud835\udc57, ?\u0308?\ud835\udc5e\ud835\udc57\ud835\udc57) + +2\ud835\udc3c\ud835\udc3c?\u0307?\ud835\udc57(\ud835\udc5e\ud835\udc5e\ud835\udc57\ud835\udc57, ?\u0307?\ud835\udc5e\ud835\udc57\ud835\udc57) \ud835\udc51\ud835\udc513\ud835\udc5e\ud835\udc5e\ud835\udc57\ud835\udc57 \ud835\udc51\ud835\udc51\ud835\udc61\ud835\udc613 + \ud835\udc3c\ud835\udc3c\ud835\udc57\ud835\udc57(\ud835\udc5e\ud835\udc5e\ud835\udc57\ud835\udc57) \ud835\udc51\ud835\udc514\ud835\udc5e\ud835\udc5e\ud835\udc57\ud835\udc57 \ud835\udc51\ud835\udc51\ud835\udc61\ud835\udc614 + ?\u0308?\ud835\udc50 (\ud835\udc5e\ud835\udc5e\ud835\udc57\ud835\udc57, ?\u0307?\ud835\udc5e\ud835\udc57\ud835\udc57, ?\u0308?\ud835\udc5e\ud835\udc57\ud835\udc57, \ud835\udc51\ud835\udc513\ud835\udc5e\ud835\udc5e\ud835\udc57\ud835\udc57 \ud835\udc51\ud835\udc51\ud835\udc61\ud835\udc613 ) + ?\u0307?\ud835\udc54(\ud835\udc5e\ud835\udc5e\ud835\udc57\ud835\udc57, ?\u0307?\ud835\udc5e\ud835\udc57\ud835\udc57, ?\u0308?\ud835\udc5e\ud835\udc57\ud835\udc57)], (3) Thus the vector of input signals can be expressed as \ud835\udf0f\ud835\udf0f = \ud835\udc3c\ud835\udc3c\ud835\udc4e\ud835\udc4e (\ud835\udc5e\ud835\udc5e\ud835\udc57\ud835\udc57, ?\u0307?\ud835\udc5e\ud835\udc57\ud835\udc57, ?\u0308?\ud835\udc5e\ud835\udc57\ud835\udc57, \ud835\udc51\ud835\udc513\ud835\udc5e\ud835\udc5e\ud835\udc57\ud835\udc57 \ud835\udc51\ud835\udc51\ud835\udc61\ud835\udc613 , \ud835\udc51\ud835\udc514\ud835\udc5e\ud835\udc5e\ud835\udc57\ud835\udc57 \ud835\udc51\ud835\udc51\ud835\udc61\ud835\udc614 ). (4) Finally, the joint stiffness \ud835\udc58\ud835\udc58\ud835\udc57\ud835\udc57 is defined as the partial derivative of the torque (\ud835\udf0f\ud835\udf0f) with respect to the joint angle (\ud835\udc5e\ud835\udc5e\ud835\udc57\ud835\udc57): \ud835\udc58\ud835\udc58\ud835\udc57\ud835\udc57 = \ud835\udf15\ud835\udf15\ud835\udf0f\ud835\udf0f \ud835\udf15\ud835\udf15\ud835\udc5e\ud835\udc5e\ud835\udc57\ud835\udc57 . As it is known, there is a dependency between the longitudinal deformation (\ud835\udf00\ud835\udf00\ud835\udc50\ud835\udc50) of a contracting/expanding EAP and the applied voltage (\ud835\udc48\ud835\udc48\ud835\udc50\ud835\udc50) [8]. \ud835\udc48\ud835\udc48\ud835\udc50\ud835\udc50 = \u00b1\u221a(\ud835\udf00\ud835\udf00\ud835\udc50\ud835\udc50 \u2212 \ud835\udf0e\ud835\udf0e\ud835\udc50\ud835\udc50 \ud835\udc38\ud835\udc38 ) \u2219 2\ud835\udc38\ud835\udc38 1 \u2212 2\ud835\udc63\ud835\udc63\ud835\udc50\ud835\udc50 , (5) where \ud835\udf08\ud835\udf08 denotes the Poisson's coefficient, \ud835\udc38\ud835\udc38 Yung module, \ud835\udc50\ud835\udc50 capacity, \ud835\udf0e\ud835\udf0e mechanical stress. It is known that electrically induced bending moment \ud835\udc40\ud835\udc40\ud835\udc4f\ud835\udc4f of the bending actuator is proportional to the input voltage \ud835\udc48\ud835\udc48\ud835\udc4f\ud835\udc4f. It is reasonable to assume that \ud835\udc40\ud835\udc40\ud835\udc4f\ud835\udc4f is proportional to the width of the sheet d [8]: \ud835\udc48\ud835\udc48\ud835\udc4f\ud835\udc4f = \ud835\udc40\ud835\udc40\ud835\udc4f\ud835\udc4f \ud835\udc3e\ud835\udc3e\ud835\udc50\ud835\udc50\ud835\udc51\ud835\udc51 , where \ud835\udc3e\ud835\udc3e\ud835\udc50\ud835\udc50 is a normalized electromechanical coupling. In summary, after determining of the required torques it will be easy to control actuators by the applied voltages, substituting (5) in (4). Necessary calculations are implemented by MSC ADAMS simulation for the joint rotations along three axes and the required characteristics of the actuators are defined. Figure 2 presents the digital model of the joint, where body 1 models the sum of the masses of the leg and other links of the mechanism, while EAPs are modeled by springs with dampers (2). Fig. 2. MSC ADAMS model of the over-actuated compliant joint. Dynamic modeling of a new over-actuated compliant joint mechanism for human\u2026 1783 Fig. 3. Force-deformation diagram for the one-actuator model during the movement in the sagittal plane (contraction/expansion). During modeling it's easy to evaluate the capacity of one actuator and then to determine it for each one analytically due to the structure symmetricity. Since EAPs are generally not so powerful, we need multiple actuators but with less power. Simple calculations show (Fig. 3) that for 100 mm of lever length we need 12 two-directional working mode actuators, if each one can develop about 444 N force (it is known that PPY-metal coil composite actuator can develop up to 500 N force). Fig. 4. Sum of the developed forces of the bending actuators. The required movement of the compliant joint can't be achieved by operating only contracting/expanding actuators, so it will be easy to calculate them in sequence: first contracting/expanding, then bending. After the first operation the link will rotate only 500, so the rest of the job will be done by bending with the appropriate actuators (Fig. 4). Force evaluation of each actuator can be done in the same manner, using again the symmetry of the structure." ] }, { "image_filename": "designv11_62_0000786_aero.2013.6496960-Figure6-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0000786_aero.2013.6496960-Figure6-1.png", "caption": "Figure 6 \u2013 Dual pantograph mechanism", "texts": [ " This figure displays the excavation depth and motor torque over an elapsed time. As seen, the excavation velocity was fast for the first min and then maintained at an almost constant average velocity until approximately 2.5 min, where it slowed to nearly zero shortly afterwards. The robot excavated to a depth of 247 mm. The maximum motor torque was approximately 18 Nm, so the propulsion unit needs to maintain its body position orientation against this torque. Development of a Propulsion Unit with Sensors Figure 6 shows the structure of a subunit. The subunit was equipped with two stepper motors, two ball screws and dual pantographs (Fig. 7). The stepper motors and ball screws controlled things like the contraction and extension of plates. As the subunit contracted, the dual pantograph extended in a radial direction, and as the subunit extended, the dual pantograph contracted in a radial direction (Fig. 8). The dual pantograph can push in the parallel direction against the wall. In addition, the expansion plates contained large, circular arc areas to maintain contact with the wall surface, which holds the body position against the rotation action of the EA" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0002927_sbr-lars-r.2017.8215275-Figure1-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0002927_sbr-lars-r.2017.8215275-Figure1-1.png", "caption": "Fig. 1. AVALON design in different perspectives: isometric view in the top right, top view in the top left, front view in the bottom left and right view in the bottom right.", "texts": [ " We introduced the characteristics and operation of AVALON and described its kinematic and dynamic modelling. In this context, to provide a complete UAV platform, a new autopilot structure that guarantees that AVALON follows a trajectory described by waypoints and maintains it in predefined flight conditions during all flight stages was presented. The simulation results were performed in MATLAB R\u00a9. AVALON is a small electric tailsitter UAV with fixedwings and a structure to keep it in the vertical position, as seen in Fig. 1. It has just one propeller, two ailerons, two elevators and one rudder, as a typical fixed-wing aircraft. The design is symmetrical in the horizontal plane and the main components aerofoil is the flat plate, chosen due to simplicity for construction. Table I shows other physical parameters of the aircraft, obtained through the SolidWorks design. As other tailsitters, AVALON operation is divided in five flight stages, respectively referenced in Fig. 2: takeoff, Transition from Vertical to Horizontal position (TVH), horizontal 978-1-5386-0956-9/17/$31" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0002422_icosc.2017.7958662-Figure1-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0002422_icosc.2017.7958662-Figure1-1.png", "caption": "Figure 1. Schematic of Pioneer 3DX under consideration.", "texts": [ " The remainder of this article is organized as follows: In section tow we outline the model of the mobile platform Pioneer 3DX robot which demonstrates that the system is flat. In section three, we present the fuzzy controller and the aim of this technique to adjust the different gains in real time where various kinds of membership functions are used and tested where the best one in term of accuracy is selected. Results are detailed in the fourth section and we terminate with a conclusions and future directions work. 978-1-5090-3960-9/17/$31.00 \u00a92017 IEEE 173 A simplified model of the Pioneer 3-DX under consideration depicted in Fig. 1. The generalized coordinate Tyxq ],,[ \u03b8= represents the final robot configuration, where ),( yx describes the position coordinates of the axial center and \u03b8 represents the orientation of the robot frame respectively. The motion of robot including the nonholonomic constraints describes as: cos sin 0y x\u03b8 \u03b8+ = (1) 0)( =qqA (2) Where ]0cossin[)( \u03b8\u03b8\u2212=qA The robot model shown in fig.1 including nonholonomic constraints can be rewritten in matrix form as follow: cos 0 sin 0 1 0 x v y w \u03b8 \u03b8 \u03b8 = (3) Where v represents the robot velocity w is the angular velocity. We observe that the model of (3) cannot be linearized in closed loop. So, we assume that the flat output as: ),(),( 21 yxFFF == (4) The derivative of (4) with respect to time, one considers (3), yields to: = = w v y x F F 0sin 0cos 2 1 \u03b8 \u03b8 (5) Where \u03b8cosvx = and \u03b8sinvy = The singularity problem between the system inputs and the flat outputs addressed to introduce the input prolongation of velocity v , and enlargement the system to: wzvvyvx ==== \u03b8\u03b8\u03b8 ,,sin,cos 1 (6) In which 1z is the new system input, the double time derivatives of the flat outputs gives: \u2212 = = w z v v y x F F 1 2 1 cossin sincos \u03b8\u03b8 \u03b8\u03b8 (7) Basing on the principle of differential flatness, we can proceed as follows [19]: 2 2 2 1 2121 2 2 2 1 2211 1 1 21 2 2 2 1 21 tan ),(),( FF FFFFw FF FFFFvz F F FFv FFyx + \u2212 == + + == = += = \u2212 \u03b8 \u03b8 (8) From the equation (8) we can deduce that the system represented by the equation (3) is flat" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0003937_ab49be-Figure1-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0003937_ab49be-Figure1-1.png", "caption": "Figure 1. Configuration of locking\u2013unlocking mechanism (LUM) actuated by SMA springs.", "texts": [ " In addition, a wire encoder is attached to the ratchet rotary shaft to measure FSAs\u2019 displacement length, and its feedback is used to control the position. This study designs, fabricates, and applies the LUM to SMA-based SFM and experimentally evaluates the reduction in energy consumption and the position control performance. 2. Locking\u2013unlocking mechanism (LUM) The FSA should contract in accordance with the contraction length of human muscle depending on the action and should be able to hold the contraction length without consuming additional power if a holding motion is required. Figure 1 shows the LUM constructed for this purpose. The LUM was fixed to the upper end of the FSA. It consists of a ratchet, a pawl, pawl-rotating actuators (PRA), a wire encoder, and a small magnetic encoder [17]. The ratchet can be freely rotated by the pawl in both directions in an unlocking state. The other end of the wire wound around the ratchet\u2019s rotary shaft is fixed to the FSA\u2019s lower end (figure 2(a)). When a load is applied to the FSA in the unlocking state, the FSA extends and its length increases" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0001467_9783527644117.ch5-Figure5.3-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0001467_9783527644117.ch5-Figure5.3-1.png", "caption": "Figure 5.3 Schematics of (a) direct electron transfer (DET) and (b) mediated electron transfer (MET) between an enzyme active site and an electrode to oxidize a substrate.", "texts": [ " The coenzyme can be either tightly bound within the enzyme structure (coenzyme) or released from the enzyme \u2019 s active site during the reaction (cofactor). Common coenzymes/cofactors include fl avin adenine dinucleotide ( FAD ), nicotinamide adenine dinucleotide ( NAD ), and pyrroloquinoline quinone ( PQQ ) (Figure 5.2 ). If the active site of the enzyme is located suffi ciently close to the electrode surface electrons can be transferred directly from the enzyme to the electrode as depicted in Figure 5.3 a. In the case of an anodic reaction, the electrode replaces the natural co - substrate (such as oxygen) as an electron acceptor. This process is known as direct electron transfer ( DET ), often categorized as \u201c third - generation \u201d enzyme electrodes in the biosensor literature, and is the most elegant and simplest method of bioelectrocatalysis between an enzyme active site and an electrode. FAD (oxidized form) FADH2 (reduced form) a) Although DET has been reported for a wide range of enzymes [13, 17 \u2013 19] , several challenges need to be overcome to achieve signifi cant rates of DET, leading to appreciable current densities, between active sites and solid electrode surfaces", " Finally, even when an electrode can approach suffi ciently close to an active site to achieve DET, usually taking the place of a redox substrate/co - substrate, this may not necessarily lead to generation of a bioelectrocatalytic current for substrate electrolysis, as the electrode can block access to the active site of the co - substrate/substrate. As an alternative to DET, small, artifi cial substrate/co - substrate electroactive molecules (mediators) can be used to shuttle electrons between the enzyme and the electrode (Figure 5.3 b). This involves a process in which the enzyme takes part in the fi rst redox reaction with the substrate and is re - oxidized or reduced by the mediator which in turn is regenerated, through a combination of physical diffusion and self - exchange, at the electrode surface. The mediator circulates continuously between the enzyme and the electrode, cycled between its oxidized and reduced forms, producing current. This process is known as mediated electron transfer ( MET ). In MET, the thermodynamic redox potentials of the enzyme and the mediator should be accurately matched" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0002621_aim.2017.8014122-Figure3-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0002621_aim.2017.8014122-Figure3-1.png", "caption": "Figure 3. Foot position put in front of stairs too close", "texts": [ " 2 T R, , \u03b8\u03b8\u03b8\u03b8\u03b8 f\u03b8\u03b8\u03b8\u03b8\u03b8\u03c4 humanhumanhuman rapapaptorque ghM JghM (1) 3R \u03b8\u03b8p JJ (2) 2T R , rededd humanapimpedance KDJJMJ gghM fpp\u03b8\u03b8 \u03b8\u03b8\u03b8\u03b8\u03b8\u03b8\u03c4 (3) Where M and h are total (apparatus and user\u2019s leg) values. The calculation model of the apparatus is shown as Fig. 2. Predefined walking trajectory is calculated with human\u2019s leg model by using direct kinematics with the angle variation data for hip, knee, and ankle joint. It is a line chart composed of 20 points with x and z coordinates of the foot compared to hip joint. As the user walks close to stairs and prepare to climb on by using a predefined trajectory, there will be two possible dangerous situations happened. As shown in Fig. 3, if the forward leg is placed in front of stairs too closely, on the one hand, it will collide at next step because the space is too small to move leg along with a fixed stair walking trajectory. On the other hand, if it is placed too far, after the next step, the foot cannot be put on stair safely so that the single leg is difficult to prevent it from falling, as shown in Fig. 4, the reaction force to the CoP create an enormous torque causing the falling. When healthy people are climbing on stairs, they can move fluently with their inertia force because they have more powerful and flexible muscle and relative faster walking velocity to support climbing up, but the elderly and the spinal cord injuries don\u2019t have enough muscle strength and inertia force to preserve their climbing motion continuously" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0002966_s10514-020-09949-2-Figure1-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0002966_s10514-020-09949-2-Figure1-1.png", "caption": "Fig. 1 Inertial and Body Frames. The inertial frame is centered at oI, and the body frame is fixed to the quadrotor at its center of mass oB", "texts": [ " , pn rc collision radius (m) ai, bi, ci inner ellipsoid\u2019s semi-principle lengths (m) ao, bo, co outer ellipsoid\u2019s semi-principle lengths (m) Ei inner ellipsoid Eo outer ellipsoid Umax maximum potential (m/s) U ellipsoidal potential function \u21c0 pg position of the destination relative to oB (m) pg \u21c0 pg resolved in I (m) sd desired speed (m/s) ds stopping distance (m) Vg goal velocity (m/s) ne number of points p1, . . . , pn contained in Eo Implementation and Experiment Ts sample time (s) pwi pulsewidth sent to i th ESC-motor pair (ms) vbat battery voltage (V) Let I be an inertial frame, that is, a frame in which Newton\u2019s second law is valid. The origin oI of I is any convenient point on the Earth\u2019s surface, and the orthogonal unit vectors of I are \u0131\u0302 , j\u0302 , and k\u0302. Let B be the body frame, which is fixed to the quadrotor\u2019s center of mass oB and has orthogonal unit vectors b\u03021, b\u03022, and b\u03023. Figure 1 shows the inertial and body frames. The position of oB relative to oI is \u21c0 p = x\u0131\u0302+ y j\u0302+zk\u0302. Let F \u02d9(\u00b7) and F \u00a8(\u00b7) denote the first and second time derivatives of a physical vector with respect to the frame F. Thus, the acceleration of oB relative to oI with respect to I is \u21c0 a I( \u21c0\u0308 p) = x\u0308 \u0131\u0302 + y\u0308 j\u0302 + z\u0308k\u0302. Let [ \u00b7 ]F denote a physical vector or tensor resolved in the frame F. Therefore, \u21c0 a resolved in I is a [\u21c0 a ] I = [ x\u0308 y\u0308 z\u0308 ]T. Let \u03c8 , \u03b8 , and \u03c6 be the yaw, pitch, and roll Euler angles defined by a 3-2-1 rotation sequence (e" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0001235_speedam.2010.5542396-Figure1-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0001235_speedam.2010.5542396-Figure1-1.png", "caption": "Fig. 1. Magnetic force on a whirling rotor.", "texts": [ " This means that the bending critical speeds of rotors will cause problems more and more often. There are relatively many studies on the unbalanced magnetic pull and its effects on the rotordynamics of induction machines [1-4]. Synchronous reluctance machines have been studied less frequently [5] although they are often proposed for high-speed applications in which the interaction of magnetic forces and rotordynamics is important. We shall study the force acting between the rotor and stator when the rotor is performing cylindrical circular whirling motion with respect to the stator, Figure 1. This means that the rotor remains aligned with the stator but the geometrical centerline of the rotor travels around the geometrical centerline of the stator in a circular orbit with a certain frequency, called whirling frequency, and with a certain radius, called whirling radius. The tangential magnetic force in Figure 1 points in the same direction as the velocity vector related to the whirling motion. The force can transfer energy between the magnetic field and mechanical system. This is a potential source for rotordynamic instability. Measurements and simulations on cage induction motors [3] have shown that the unbalanced magnetic pull strongly depends on the whirling frequency. This is related to the interaction of eccentricity harmonics with the rotor-cage currents. In a synchronous reluctance machine, there is no rotor winding" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0002171_ilt-01-2016-0005-Figure1-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0002171_ilt-01-2016-0005-Figure1-1.png", "caption": "Figure 1 Kinematic relationship of roller bearing", "texts": [ " Comparison analysis between uncrowned roller and partially crowned roller (i.e. intersecting arc trimming roller) has been conducted. The surface morphologies and the profile information of generatrix will be obtained to verify the surface asperity flattening effect in EHL, which contributes to anti-fatigue life of roller bearing with sliding. With the basis of research results in this paper, further works will be conducted in the next step. Kinematic relation sketch of roller bearing is shown in Figure 1(a). In the orthogonal experiments, speed of highspeed motorized spindle is r w . As the location of highspeed motorized spindle is fixed, actually the speed of low-speed motorized spindle w will be the difference value of inner ring i w and revolution speed of roller m w . i m w w w= \u2212 , i.e. i m w w w= + (1) set, \u03b3 is the ratio of r D to m D . r m D D \u03b3 = (2) If the slip fraction of roller and inner ring is zero, the speed of inner ring is i w , and the ideal speed of roller is rT w . But the speed of roller will be r w with the slide/rolling contacts" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0003190_978-3-030-05321-5_8-Figure8.13-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0003190_978-3-030-05321-5_8-Figure8.13-1.png", "caption": "Fig. 8.13 Vane packing", "texts": [ "11a, b show the time response of the operating pressures and the position of the pistons on the pushing side and the pulling side, respectively. Although the pressure increases near the stroke end point, the operating pressure achieved is <0.01MPa. Oscillating torque actuators that can be operated with 35MPa were also developed, and the output density reached higher values, as shown in Fig. 8.12. The developed motors were designed to drive with no load operated by<0.2MPa. The vane packing was molded out of a special resin (Fig. 8.13). The developed oscillating actuator, depicted in Fig. 8.14, has one vane, with a rotating angle of 270\u25e6, and an output torque of 160\u2013670Nm. A 360\u25e6 oscillating motor was also developed for the high requirements in robotic usage (Fig. 8.15). This oscillating motor can generate 600 N m of torque with an applied pressure of 35MPa. Using this type of hydraulic motor can enable a design with a smaller outer diameter, which is advantageous for a compact hydraulic robot design. Hydraulic systems require pumps as high-pressure sources that are driven by electric motors or engines to supply hydraulic power to hydraulic actuators" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0003106_robio.2017.8324671-Figure7-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0003106_robio.2017.8324671-Figure7-1.png", "caption": "Figure 7. The quadrotor test platform", "texts": [], "surrounding_texts": [ "According to the dynamic model and inputs, the quadrotor is an under-actuated system. Therefore, the roll and pitch angle are designed as the inner virtual control input for the velocity subsystem and the controlled outputs are { \u0307 , \u0307 , , } to satisfy the output of the trajectory tracking guidance. Based on this, the objective is to design a backstepping controller to track the desired altitude , yaw angle and velocity , , , . The backstepping approach is derived in five steps shown below with the inputs defined as: \u23a9\u23a8 \u23a7 = ( \u2212 \u2212 + ) = ( + \u2212 \u2212 ) = ( \u2212 + \u2212 )= + + + (16) where (i =1,2,3,4) is the force generated by each propeller mounted on the arm at a relative position \u2208 { } to the gravity center of the quadrotor. > 0 is known as the force-moment scaling factor. Then, dynamic model equations (15) can be rewritten as: \u0307 = + ( , ) ( ) \u0307 = ( , ) + ( , )\u0307 = ( , ) + \u0307 = ( , ) + ( )\u0307 = ( , ) + ( ) (17) where the state vectors are defined as: = \u0307\u0307 = \u0307 = = = = = (18) and the vector ( = 0,1,2,3,4) are: = = \u2212 = + = 0( \u2044 ) = \u2212 + (19) and the matrix ( = 0,1,2,3,4) are: = \u2212 = 10 = 1\u2044 00 1\u2044 = 1 00 \u2044 = \u2044 00 1\u2044 and the vector is: = (21) with = \u2212 \u00d7 (22) where [ ] is the control input. \u0394 , \u0394 , \u0394 are (20) represent the external translational drag force including the gyro effect, wind gust, blade flapping and induced drag. While, \u0394 , \u0394 , \u0394 are represent the external rotational drag torque including the gyro effect caused by the unmolded rotor and air friction. These six terms will be compensated through backstepping approach. Step 1 Consider the virtual system: \u0307 = + (23) with = (24) The first tracking error is: = \u2212 (25) where is the desired velocity. Introducing the observation term with its estimated error = \u2212 and considering the Lyapunov function positive definite and its time derivative negative semi-definite: = 12 + 12 (26) \u0307 = \u0307 \u2212 + + \u0307 + (27) It should be noted that the determinant of matrix is ( \u2044 ) and the input represents the total thrust which is generally nonzero and equal to during the hover state. So, is nonsingular under the general flight condition. The stabilization of can be obtained by the virtual control input as: = + \u0307 \u2212 \u0307 = \u2212 (28) where \u2208 \u211d \u00d7 and are two positive definite diagonal matrix respectively. Hence, the Lyapunov derivative equation (27) becomes \u0307 = \u2212 \u2264 0. Step 2 Consider the virtual system: \u0307 = + (29) The second tracking error is: = \u2212 = + \u0307 \u2212 \u2212 (30) then, the derivative of tracking error can be rewritten as: \u0307 = \u2212 + + (31) The Lyapunov function positive definite and its time derivative negative semi-definite: = + 12 (32) \u0307 = \u2212 + + \u0307 \u2212 ( + ) (33) where is the Jacobian matrix of derived as: = ( ) = 0\u2212 (34) It should be noted that the determinant of matrix is . So, is nonsingular under the condition of < < and < < , which has been predefined in the dynamic modeling. Besides, the determinant of matrix is , which is also nonsingular under the same condition. The stabilization of can be obtained by the virtual control input as: = ( ( + + \u0307 ) \u2212 ) (35) where \u2208 \u211d \u00d7 is a positive definite diagonal matrix. The Lyapunov derivative equation (33) is then \u0307 = \u2212 \u2212\u2264 0. Step 3 Consider the following real system: \u0307 = + + (36) with = (37) The third tracking error is: = \u2212 = ( ( + + \u0307 ) \u2212 ) \u2212 (38) then, the derivative of tracking error can be rewritten as: \u0307 = \u2212 \u2212 + (39) Introducing the observation term with its estimated error = \u2212 and considering the Lyapunov function positive definite and its time derivative negative semi-definite: = + 12 + 12 (40) \u0307 = \u2212 \u2212 + \u0307 ++ + \u0307 \u2212 \u2212 \u2212 (41) It should be noted that the determinant of matrix is 1\u2044 . Therefore, is nonsingular since the body inertial and are commonly nonzero. The stabilization of can be obtained by the real control input as: = + + \u0307 \u2212 \u2212 \u0307 = \u2212 (42) where \u2208 \u211d \u00d7 and are two positive definite diagonal matrix respectively. Hence, the Lyapunov derivative equation (41) becomes \u0307 = \u2212 \u2212 \u2212 \u2264 0. Step 4 Consider the virtual system: \u0307 = + (43) The first tracking error is: = \u2212 (44) where is the desired altitude and yaw angle. The Lyapunov function positive definite and its time derivative negative semi-definite: = 12 (45) \u0307 = \u0307 = ( \u0307 \u2212 + ) (46) It should be noted that the determinant of matrix is \u2044 . So, is nonsingular under the condition of < < , which has been predefined in the dynamic modeling. The stabilization of can be obtained by the virtual control input as: = ( + \u0307 \u2212 ) (47) where \u2208 \u211d \u00d7 is a positive definite diagonal matrix. The Lyapunov derivative equation (46) is then \u0307 = \u2212 \u2264 0. Step 5 Consider the following real system: \u0307 = + + (48) with = (49) The fifth tracking error is: = \u2212 = ( + \u0307 \u2212 ) \u2212 (50) then, the derivative of tracking error can be rewritten as: \u0307 = \u2212 \u2212 (51) Introducing the observation term with its estimated error = \u2212 and considering the Lyapunov function positive definite and its time derivative negative semi-definite: = + 12 + 12 (52) \u0307 = \u2212 + \u0307 ++ + \u0307 \u2212 \u2212 \u2212 (53) It should be noted that the determinant of matrix is \u2044 . Therefore, is nonsingular because the body mas and inertial are commonly nonzero. The stabilization of can be obtained by the virtual control input as: = + + \u0307 \u2212 \u2212 \u0307 = \u2212 (54) where \u2208 \u211d \u00d7 and are two positive definite diagonal matrix respectively. Hence, the Lyapunov derivative equation (53) becomes \u0307 = \u2212 \u2212 \u2264 0. According to the positive definite Lyapunov function in step 3 and 5, the whole system is asymptotically stable by using the virtual control , and and real control and as shown below: \u23a9\u23aa\u23aa\u23aa \u23aa\u23a8 \u23aa\u23aa\u23aa\u23aa \u23a7 = ( \u2212 ) + \u0307 \u2212 = ( ( ( \u2212 ) + ( \u2212 ) + \u0307 ) \u2212 ) = ( ( \u2212 ) + \u0307 \u2212 ) = ( \u2212 ) + ( \u2212 ) + \u0307 \u2212 \u2212= ( \u2212 ) + ( \u2212 ) + \u0307 \u2212 \u2212 \u0307 = \u2212 ( \u2212 ) \u0307 = \u2212 ( \u2212 ) \u0307 = \u2212 ( \u2212 ) Finally, the desired force should be produced by each rotor can be obtained from the inverse of the equation 33. IV. SIMULATION RESULT The derived trajectory tracking guidance and backstepping controller for quadrotors are verified through the simulation using Runge-Kutta method under Matlab environment. The physical parameters for a real quadrotor are identified through several experiments and shown below: = 0.77 = 0.12 = 9.81 / = [5.20, 5.20, 10.27] \u00d7 10 . / = 4.6 \u00d7 10 Initially, the quadrotor is in hover flight located at [0 , 0] with the start conditions that (0) = (0) = (0) = (0) =(0) = [0, 0] and [ ] = [0, 0, 0, ] . The trajectory is consist of six waypoints located at (6.25, 0), (10, 5), (0, 15), (2.5, 19.33), (7.5, 19.33) and (10, 15), representing an S-shape path. The desired altitude are preset to zero and ignored since the trajectory tracking guidance are not cover vertical direction. The desired yaw angle is also set to zero. The desired velocity is considered as 1 / and defined as: \u0307 = 2(1 \u2212 ) The parameters used for trajectory tracking guidance and backstepping controller are defined as: = 0.5 = 0.9 = [3,3] = = = = [6,6] = = = [3,3] Figure 3 shows the result positions and ground velocity of the quadrotor. It can be seen the good performances and tracking ability of the desired trajectory and velocity. Figure 4 and 5 present the output velocities and accelerations of the derived guidance system and backstepping controller along x-axis and y-axis with tracking errors. The continuous and smooth variation of velocities and accelerations reflects the acceptability and physical feasibility of the guidance system and backstepping controller. Moreover, Figure 6 shows the corresponding roll and pitch angle outputs of the quadrotor. The variation ranges are acceptable. (55) V. CONCLUSION In this paper, we derive a trajectory tracking guidance for quadrotor UAVs. This approach controls the position of the quadrotor to track the desired trajectory by force the velocity direction converge to the line defined by current position and reference point. The backstepping control algorithm is applied to follow both the desired velocity output from the guidance system and the desired altitude and yaw angle. Finally, the simulation results show the good performance and feasibility of the proposed trajectory tracking guidance system. For the future works, the system will be tested on a quadrotor platform as show in the figure 3. The attitude information is computed from onboard IMU sensors, while the position and velocity feedbacks can be obtained from the Vicon motion capture system. REFERENCE [1] J. Li and Y. Li, \"Dynamic Analysis and PID Control for a Quadrotor,\" IEEE International Conference on Mechatronics and Automation, Beijing, China, 2011, pp. 573-578. [2] S. Khatoon, D. Gupta and L. K. Das, \"PID & LQR control for a quadrotor: Modeling and simulation,\" 2014 International Conference on Advances in Computing, Communications and Informatics (ICACCI), New Delhi, 2014, pp. 796-802.. [3] S. Islam, P. X. Liu and A. El Saddik, \"Robust Control of Four-Rotor Unmanned Aerial Vehicle With Disturbance Uncertainty,\" in IEEE Transactions on Industrial Electronics, vol. 62, no. 3, pp. 1563-1571, March 2015. [4] T. Madani and A. Benallegue, \"Backstepping Control for a Quadrotor Helicopter,\" 2006 IEEE/RSJ International Conference on Intelligent Robots and Systems, Beijing, 2006, pp. 3255-3260. [5] Z. Fang and W. Gao, \"Adaptive integral backstepping control of a Micro-Quadrotor,\" 2011 2nd International Conference on Intelligent Control and Information Processing, Harbin, 2011, pp. 910-915. [6] H. Lee, S. Kim, T. Ryan and H. J. Kim, \"Backstepping Control on SE(3) of a Micro Quadrotor for Stable Trajectory Tracking,\" 2013 IEEE International Conference on Systems, Man, and Cybernetics, Manchester, 2013, pp. 4522-4527. [7] M. J. Reinoso, L. I. Minchala, P. Ortiz, D. F. Astudillo and D. Verdugo, \"Trajectory tracking of a quadrotor using sliding mode control,\" in IEEE Latin America Transactions, vol. 14, no. 5, pp. 2157-2166, May 2016. [8] D. Mellinger and V. Kumar, \"Minimum snap trajectory generation and control for quadrotors,\" 2011 IEEE International Conference on Robotics and Automation, Shanghai, 2011, pp. 2520-2525. [9] R. Mahony, V. Kumar and P. Corke, \"Multirotor Aerial Vehicles: Modeling, Estimation, and Control of Quadrotor,\" in IEEE Robotics & Automation Magazine, vol. 19, no. 3, pp. 20-32, Sept. 2012. [10] J. T. Jang et al., \"Trajectory generation with piecewise constant acceleration and tracking control of a quadcopter,\" 2015 IEEE International Conference on Industrial Technology (ICIT), Seville, 2015, pp. 530-535. [11] S. Park, J. Deyst, and J. P. How, \u201cA new nonlinear guidance logic for trajectory tracking,\u201d in Proc. of the AIAA Guidance, Navigation and Control Conf., Providence, RI, USA, pp. 941-956, August 2004. [12] P. Martin and E. Sala\u00fcn, \"The true role of accelerometer feedback in quadrotor control,\" 2010 IEEE International Conference on Robotics and Automation, Anchorage, AK, 2010, pp. 1623-1629." ] }, { "image_filename": "designv11_62_0002867_sice.2017.8105701-Figure2-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0002867_sice.2017.8105701-Figure2-1.png", "caption": "Fig. 2: D* planned paths from S to G. (a) Path without obstacles. (b) Path when knowledge of obstacle A is informed to robot. (c) Path when obstacle A and B are informed. The darker shades in the colormap represents proximity to goal (G).", "texts": [ " In case of receiving an obstacle removal message, a check if performed if a shorter path is available from the current location. Thus, robots communicate and inform each other about the new obstacles in the map, enabling efficient path planning. The proposed technique was tested in the simulation environment shown in Fig.3 using the Matlab software. D* algorithm [5][6] was used for path planning, however, any other path planning algorithm can also be used. A grid based navigation is chosen with one unit cost for forward, back, left, and right movement, whereas, for diagonal movement the cost is \u221a 2 units. In Fig.2, S and G represents the start and goal locations of the robot, respectively. Figure 2(a) shows the path of Robot R1 from location S to goal G, and it is the shortest path found by D* path planning algorithm. Figure 2(b) shows a new obstacle A found by a robot R2 and its information is shared with R1. Thus, R1 plans an optimal path considering the new obstacle A. It is the optimal path given the knowledge of obstacle A. Figure 2(c) shows a scenario with two obstacles A and B found by robots R2 and R3. Robot R1 has timely information in this case about the remote obstacles and it plans an optimal path considering both the obstacles shown in Fig.2(c). In traditional robot navigation without obstacle knowledge sharing, the robot R1 would have navigated the shortest path shown in Fig.2(a). Upon encountering obstacles A it would then re-plan another shortest path passing through obstacle B of which it has no information. It would then re-plan its path again near obstacle location B to the goal. Thus, it would require planning the path three times and results in more time spent in navigation and path planning which are avoided in the proposed scheme. Table 1 and Table 2 shows the time required in path planning and the total distance navigated by the robot R1 in the traditional vs the proposed scheme in case of the scenario of two obstacles shown in Fig.2(c). The proposed method takes only 33% of the planning time and 39% of the distance is navigated by the robot as compared to traditional navigation. The saving in navigation distance also directly translates to saving battery power of the robots. Notice that, the first robot to find the obstacle needs to re-plan its path. However, subsequently, other robots benefit from the proposed shared autonomy. This paper proposed an architecture in which robots can timely inform other robots about the new obstacles observed in the map" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0001227_tim.2010.2046600-Figure8-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0001227_tim.2010.2046600-Figure8-1.png", "caption": "Fig. 8. Photos of current transducer installation.", "texts": [ " All outputs of the signal processing units are connected to a PC with an RS485 bus. The diagram structure of the measuring system is shown in Fig. 7. In Fig. 7(a), each of the switchgear cabinets 1 \u223c N contains three transducers and one signal processing unit. They are used to measure and process the three-phase currents inside the cabinet. The signal processing unit mainly consists of a microcontrolled unit (MCU), and its simplified architecture is shown in Fig. 7(b). A photo of the current transducer installation is shown in Fig. 8. About 60 transducers were installed in a substation in the City of Xinxiang, Henan Province, China. One of the installation structures is shown in Fig. 9. During the field trial for more than four years, the current transducers have shown good performance and measured several types of fault currents. Two types of fault current waveforms are plotted in Fig. 10. The rated currents flowing through the three-phase busbars are less than 1000 A; in some cases, it is even as less as 50\u201380 A. In Fig. 10, the maximum currents measured by the transducers during the fault condition are over 2300 A" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0003190_978-3-030-05321-5_8-Figure8.43-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0003190_978-3-030-05321-5_8-Figure8.43-1.png", "caption": "Fig. 8.43 Hydraulic tough hand", "texts": [ " In an actual operational situation, the operating condition of the robotic hand is the same as that for the existing constructionmachines. Thismeans that the hand should be able to work in wet and/or dry mud, or gravel, and should be able to grasp rocks, timbers, etc., without experiencing technical faults. Considering these heavy-duty conditions, we highly prioritized simplicity in the design of not only the structure of the hand, but also the hydraulic and electric control systems. Finally, the design priority pertained to weight saving and simplification. Figure8.43 shows the appearance of the tough hand. This hand has four fingers. The two fingers on the outer side can rotate around their vertical axes at their roots. The root of each finger is supported by a pair of tapered-roller bearings mounted in overhang plates on the palm box. Each rotating finger is driven by a vane-type swing motor placed on the upper side of the upper overhanging plate. Between the upper and lower overhanging plates, a 4-port swivel joint is located, of which, only two ports are in this application" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0003985_icamechs.2019.8861668-Figure2-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0003985_icamechs.2019.8861668-Figure2-1.png", "caption": "Fig. 2. An overview of the previous robot", "texts": [], "surrounding_texts": [ "We propose a multi-legged robot with a flexible leg mechanism that can switch the stiffness of the legs during locomotion. The mechanism of the proposed robot is presented in Fig. 4. The robot has nine segments and each segment has two legs as shown in Fig. 4. Each segment has two motors; one for vertical movement and the other for horizontal movement. The same pattern of movement with a phase delay was repeated, and the locomotion pattern was realized as shown in Fig. 5 , Fig. 6. Repetition (1) and (2) of the simple pattern reduced amount of information which had to be processed to allow for the control of the multiple legs. However, typically in complex environments, the robot when moving must adapt to the terrain. In the proposed robot, the adaption during locomotion can be realized through the flexibility of the trunk and the legs. The robot can passively adapt to the rough terrain by utilizing its flexibility and through the repetition of a simple locomotion pattern. Details of each mechanism are outlined in the subsequent sections. The pattern of locomotion pattern is shown as equation (1) and (2). (1) ' (2) n: Leg number" ] }, { "image_filename": "designv11_62_0001329_ls.121-Figure9-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0001329_ls.121-Figure9-1.png", "caption": "Figure 9. The schematic view of a heavily loaded contact in over-critical lubrication regime.", "texts": [ " The values of this coeffi cient fi nalise the fi lm thickness formulas. In this section, we will briefl y consider the so-called over-critical regimes in application to the same as before problem, i.e. the plane problem of isothermal lubrication for heavily loaded elastic solids with incompressible Newtonian fl uid. The problem can be analysed using the methods of matched asymptotic expansions. The structure of the contact area is different from the case of pre-critical lubrication regimes, and it is given in Figure 9. The contact region possesses two boundary layers (two inlet zones) with different characteristic sizes at one side of the external (Hertzian) region and two boundary layers (two exit zones) at its opposite side. In each of the larger inlet and exit zones, as well as in the Hertzian region (central zone where pressure is of order 1), the asymptotically valid equations for the main terms of asymptotic expansions of the unknown quantities can be considered. Some structural asymptotic formulas for lubrication fi lm thickness will be considered", " Further, it can be shown that asymptotic expansions (equations (86) and (87)) cannot be matched with the asymptotic expansions of the external solution pext(x). It is caused by ignoring the presence of small inlet and exit zones (\u03b50 zones) with characteristic size of \u03b50 located around points x = a0 and x = c0. In the inlet and exit \u03b50 zones, the effects of elasticity and lubrication fl ow are of the same order of magnitude. The solutions in the inlet \u03b50 and \u03b5q zones must match, as well as the inner asymptotic of the external solution pext(x) and the solution in the inlet \u03b50 zone (see Figure 9). Similarly, the solutions in the exit \u03b50 and \u03b5q zones must match, as well as the inner asymptotic of the external solution pext(x) and the solution in the exit \u03b50 zone. Note that the absence of singularities in pext(x) at x = a0 and x = c0 leads to the estimate pext(x) = O(\u03b50 1/2) for x \u2212 a0 = O(\u03b50) and x \u2212 c0 = O(\u03b50), \u03c9 << 1. That allows for matching pext(x) with the solutions in the inlet and exit \u03b50 zones. Based on the described behaviour of the solution pext(x) in the inlet and exit \u03b50 zones the solution of the problem can be searched in the form p x q r o q r O r x a O( ) ( ) ( ), ( ) ( ), ( ), ,/ /= + = = \u2212 = <<\u03b5 \u03b5 \u03b5 \u03c90 1 2 0 0 0 1 2 0 0 0 0 0 1 1 1 (88) p x g s o g s O s x c O( ) ( ) ( ), ( ) ( ), ( ), ,/ /= + = = \u2212 = <<\u03b5 \u03b5 \u03b5 \u03c90 1 2 0 0 0 1 2 0 0 0 0 0 1 1 1 (89) where q0(r0) and g0(s0) are unknown functions that should be determined, and r0 and s0 are the local point coordinates in the inlet and exit \u03b50 zones, respectively" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0001595_1.3640469-Figure7-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0001595_1.3640469-Figure7-1.png", "caption": "Fig. 7 Mathieu plane with coding for stabil ity-threshold lines used in Fig. 8. Region within d a s h e d square is enlarged a n d used in Fig. 8. Equations for l ines are approximat ions for s m a l l e. Shaded a r e a s represent stable osci l lations as solutions to Mathieu's equation.", "texts": [], "surrounding_texts": [ "\u00a3 m i Si = \u2014-7=; = T)'< M =\nV ix m2 ( 2 0\nIPto t i e following equations\n? \u00ab 2 M\nt + 1 2 V M ^ ( J S \" )\n( X - (22)\ni k His VV^i\n>7) = a,- I'' a, ri' (i + l)m2 ju Eg. ( i + l)\u00bbi2\n(i + l)m2 \\VV\nThe potential function (23) is given in terms 7} so\nbu _ a\u00a3 bu by dx b!j dx brj dx\n( 3 0 )\nThus\na2^ dx2\n^ - 2 c o s 0 s i n 0 ^ - + c o s 2 ^ : (31) a\u00a32 a??2\nand\nwith i = 1,3. These equations can be considered as the equations of motion of a particle of unit mass in a two-dimensional force field and can be derived from a potential function of the form\ndu du bu \u2014 = cos 8 \u2014r + sin 8 \u2014 by a\u00a3 drj\nHence the equations of motion when referred to the modal direction y, and x, perpendicular to the modal direction, are\nb2u \u201e \u201e bhi \u201e a2!t\"l sin2 d \u2014 - 2 cos 8 sm 0 - r \u2014 + cos2 8 \u2014\na\u00a32 b^br] by2\nh - V ) (23) dtt i/ = cos 8 \u2014r + sm 8 \u2014\n(32)\nNow, under the assumption that the modal constant has been found, the co-ordinate axes TJ are rotated to the modal line. Xet the y-axis be taken in the direction of the modal line and the x-axis be perpendicular to it. The transformation equations between the co-ordinates ij and x, y are as follows [6]:\n\u00a3 = x sin 8 -(- y cos 6, x = \u00a3 sin 8 \u2014 -q cos 8 7] = y sin 8 \u2014 x cos 8, y = \u00a3 cos 8 + rj sin 8\n\"where\n= tan - \u2022H m2\nmi The equations of motion in the x and ^-directions are now\nbu bx (x, y)\nV = y) by\n+ 3! b3u bx3 + \u2022 \u2022 \u2022 (27)\nbu Then since \u2014 (x, y) = 0 along the modal line, the second term ax on the right of (27) is zero and the equations (26) can be written\nfb2u \\ x2 (dau \\\nx2 a /a2it| \\ x3 _a / a 3 \u00ab h 2! ay W L . J + 3l a^ \\ax= bu by\n(28)\n+ . \u2022 \u2022\nNeglecting higher-order terms in x since we are looldng at the stability of the modal oscillation, that is, motion near the y-axis, we get the following:\n(b2u \\ X = X\nW o j\nbu y = sr by\n(29)\nThe first of (32) being derived from the potential function (23) which, for this case, has only second and fourth-order terms, takes the form\nx = x(fe + fay2) (33)\n(24)\n(25)\n(26)\n(where ki and fa are constants), which is a Hill equation, or, with certain approximations, a Mat.hieu equation. The second of (32) takes the form\nV = fay + fay3 (34)\nSince the modal line is given by x = 0 we can expand the potential function (23) about x = 0 by means of a Taylor's series, thus\n\u00ab ( . , \u201e ) = \u00ab\u00ab , , \u201e ) + x ( S | J + % ( g | J\n(where fa and k\\ are constants), which is a Duffing equation. Hence it is seen that the modal stability can be studied by analyzing the stability of a Duffing equation in the ?/-clirection and a Mathieu equation in the x-direction. If a perturbation introduced in the y-co-ordiuate, combined with an assumed small displacement in the x-direction makes the Bystem unstable, the modes are unstable.\nThe analysis thus involves a study of the Mathieu equation for the motion if the Duffing equation is stable.\nExamples of This Stability Analysis Taking now the first example given in the preceding section, we find from (25) and (24)\nh 1 tan 8 = -\\/2, sin 8 = - y - , cos 8 =\nand\n\u00a3 = y cos 8, V \u2014 V sin i\nsince x is zero along the modal line. Equation (33) finally reduces to\n3 x = x I - if\n\\ 2\nand (34) becomes\nv = -4 v 13.5 36\n(35)\n(36)\nFor this case, then, the stability of the mode depends on Equation (35) which can be reduced to a Mathieu equation of the form\nd*x dz2 + x(S + \u00a3 cos z) \u2014 0 (37)\nJournal of Applied Mechanics M A R C H 1 9 6 1 / 7 5\nDownloaded From: http://appliedmechanics.asmedigitalcollection.asme.org/pdfaccess.ashx?url=/data/journals/jamcav/25604/ on 04/10/2017 Terms of Use: http://www.asme.org/about-asme/terms-of-use", ".by the substitutions\ny = A cos at\nThen with z = 2arf, Equation (37) becomes\n 45\u00b0 have a good surface quality" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0001522_icra.2011.5980118-Figure4-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0001522_icra.2011.5980118-Figure4-1.png", "caption": "Fig. 4. Schematic of the arm model", "texts": [ " These tests were carried out using a simulation and are discussed as follows: first, the setup is described, second the goal-directed behavior is specified and third the simulated results are presented. All results presented in this paper were obtained with an ensemble of 100 trials and random start and goal constellation tasks. One time step was chosen such that any joint angle could change up to 0.3rad in one time step, resulting in approximately 30 time steps needed to reach any goal constellation. 12 frames of reference were used (cf. II) to describe the arm state. A schema of the arm model is shown in Fig. 4: a 6DOF-arm consisting of three limbs, each with length 0.5. The joints could rotate unrestricted. However, the model is not limited to these settings in any way. The manipulator variance is given by LPMPi,t \u2208 R 2\u00d72. Following the simplification of independent elements (x)l, 1 \u2264 l \u2264 dim(F ) of a state in frame F (q.v. Sec. II), we modeled the covariance-matrix LPMPi,t as a diagonal matrix. Analogous, the sensor covariance-matrices SkSPi,t have diagonal form. Note, that the manipulator variance LPMPi,t, the sensor variance LPSPi, t and the state estimate variance LPPi,t are all different variances given in the same space" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0002302_b978-0-12-803581-8.10351-0-Figure41-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0002302_b978-0-12-803581-8.10351-0-Figure41-1.png", "caption": "Fig. 41 Left clip interface bracket.", "texts": [], "surrounding_texts": [ "The mono shock wedge is designed to react suspension loads from the spring/damper unit and support the suspension box rear angle wall. The design utilizes the same U-shape cross-section as the intermediate bulkhead as seen in Fig. 29. The wedge will be bonded to the chassis using the two angular bonding flanges and to the main bulkhead with the two vertical bonding flanges shown in Fig. 37(a). The wedge design and bonding flanges allow for transfer of the load from the spring/damper unit to the chassis and then into the main bulkhead and the main structure of the chassis as shown in Fig. 37(b). The layup of the wedge follows that of the intermediate bulkhead, [745, 0, 0\u201390, 90, 745, 0, 0\u201390, 90, 745, 0, 0\u201390, 90], from the outer surface to the inner surface. Ungrouped plies indicate the use of unidirectional prepreg, while grouped plies indicate the woven material." ] }, { "image_filename": "designv11_62_0003742_ecc.2019.8795690-Figure4-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0003742_ecc.2019.8795690-Figure4-1.png", "caption": "Fig. 4: The 3D depth image for the maze and Block Average for each cell. The Block Average calculates the average of this block of pixels.", "texts": [ " The camera will return a 2D pixel matrix M (M \u2208 Rnr\u00d7nc , where nr, nc denotes the numbers of pixels in row and column respectively the camera captured). The pixel matrix M with color value includes the information for the maze. To extract the wall information, we can use the following preprocessing approach: M\u0302(i, j) = { 1, i f |M(i, j) \u2212 g| 6 \u03b8, 0, else, (1) where i = 0, 1, ..., nr \u2212 1; j = 0, 1, ..., nc \u2212 1, g is a constant representing the gray of the wall and \u03b8 is an offset range. After preprocessing, the depth image can be expressed in 3D form, as shown in Fig. 4. For a given cell with the center coordination (x, y), its border coordination can be acquired by adding or subtracting the width of w 2 , as shown in Fig. 4. With Block Average algorithm [3], it can be known whether there is a wall. Taking the north wall as an example, the Block Average algorithm for the north wall can be expressed as follows: Wn = 1, i f d/2\u2211 i=\u2212d/2 d/2\u2211 j=\u2212d/2 M\u0302(x+i, y+ w 2 + j) > d, 0, else, (2) where d denotes the width of the wall. The Block Average algorithm, which is used to estimate whether a wall exists or not, can effectively eliminate the effect of noise in the image capturing. The exist of the wall in south, west and east direction can be expressed by the same method" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0000747_s1070363211090210-Figure4-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0000747_s1070363211090210-Figure4-1.png", "caption": "Fig. 4. Voltammograms for poly-H2msalpn-1,2 in various background solvents: (1) 0.1 \u041c Bu4NClO4/CH3CN, (2) 0.1 \u041c Bu4NBF4/CH3CN, and (3) 0.1 \u041c Bu4NPF6/CH3CN. Vp = 0.05 V s\u20131.", "texts": [ " The films obtained in potentiodynamic conditions obviously have a more organized and regular structure and are characterized by lower Ea value. A high Ea value attests in favor of the fact that the Dct value reflects the slow motion of the charge carriers in the poly-H2msalpn-1,3 bulk. It should be noted that at the limiting character of ionic diffusion in the polymer bulk the Ea value does not exceed 25\u201330 kJ mol\u20131 [22]. The nature of the anions of the supporting electrolyte has a significant influence on the parameters of voltammograms of poly-H2msalpn-1,3 (Fig. 4). Experiments showed that increasing the size of counter-ions1 (as compared with the perchlorate ion) is accompanied by a decrease in the currents of anodic and cathodic peaks and the displacement of their potentials in the positive region. The effect of size of the doping anions can be explained by the increase in the distance between the fragments of the neighboring polymer chains, reducing the degree of \u03c0\u2013\u03c0-interaction between the fragments of poly-H2msalpn-1,3 and efficiency of charge transport by jumping mechanism" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0002738_s11771-017-3608-4-Figure1-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0002738_s11771-017-3608-4-Figure1-1.png", "caption": "Fig. 1 Ball screw drive experimental setup", "texts": [ " The ball screw drive is a various system with matched and mismatched disturbances and uncertainties since the external disturbances and time-varying uncertainties. To counteract the influence of the disturbances and uncertainties, an AISMC (adaptive integral sliding mode controller) with NDO (nonlinear disturbance observer) is adopted in this work for the flexible ball screw drive. The experiments were conducted to verify the effectiveness of the overall control strategy in comparison with a P-PI (proportional-proportional integral) controller. The ball screw drive experimental setup is shown in Fig. 1. The table is driven by a high-precision ball screw of 20 mm pitch and 20 mm diameter and supported on precision linear guideways on the both sides. The experimental setup is powered by a 9 kW synchronous AC servomotor connected to one end of the ball screw with a coupling. In addition, the end closed to the motor is fixed-end; the other end is supported-end. Two incremental rotary encoders are installed on the both sides of the ball screw and an incremental linear encoder is installed on the table", " (32) gives with the switching gain 3 d( ) ,l c e the system states will be guaranteed to reach the sliding surface \u03c3*=0 in finite time. Therefore, the AISMC with NDO is asymptotically stable even if both matched and mismatched disturbances and uncertainties exist. According to the above subsections, the ultimate control law for flexible ball screw drive system is obtained and illustrated in Fig. 4. The proposed control law was evaluated in the tracking performance experiments conducted with the ball screw drive experimental setup illustrated in Fig. 1. By using the least squares identification techniques, the parameters of Eq. (2) aimed for the experimental setup [24] are listed in Table 1. As a comparison, the P-PI controller with velocity and acceleration feed-forward and friction compensation was applied to the tracking experiments which has a widely usage in the motion control industry. The velocity loop adopted PI (proportional-integral) control and the position loop J. Cent. South Univ. (2017) 24: 1992\u20132000 1997 relied on P (proportional) control" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0001259_sav-2010-0604-Figure8-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0001259_sav-2010-0604-Figure8-1.png", "caption": "Fig. 8. The two different boundary conditions of a panel with single and double side constraints.", "texts": [ " With the above finite element model of the BIW, the modal analyses of it are carried out based on Nastran where Block Lanczos method is adopted. The calculated natural frequencies and vibration shapes are obtained, listed in Table 4. Some typical vibration shapes are shown in Fig. 7. The material properties and the boundary conditions of the BIW will change the modes effectively, especially the welding connections. Take the roof panel as an example to illustrate the influence, in which the roof panel is modeled with single or double side constraints, as shown in Fig. 8. For the first case of the model with single side constraints, the calculated modes from 1st to 4th order are listed in Table 5. The calculated modes from 1st to 4th order are listed in Table 6 for the roof panel modeled with double side constraints. It can be seen that, from Tables 5 and 6, both the natural frequencies and vibration shapes of the panel are obviously changed with different boundary conditions. In addition, the vibration shapes of them do not change obviously with material property values" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0002595_j.mechatronics.2017.08.001-Figure1-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0002595_j.mechatronics.2017.08.001-Figure1-1.png", "caption": "Fig. 1. LCD glass-handling robot system: (a) experimental equipment and (b) physical mode.", "texts": [ " Finally, the energy balance equation FF is imlemented to identify the unknown parameters for the real LCD lass-handling robot system experimentally. From the experimenal results, it is found that all identified states converge well to- ard to the real states and the system parameters are successfully dentified. The contribution of this paper is to propose an energetcs FF to be implemented in system identification for the mecharonic system. It can be concluded that the more system states are mplemented in the FF, the more correct parameters are identified or system identification by using the RGA method. . Dynamic modeling Fig. 1 (a) shows the real equipment of an LCD glass-handling obot system. It consists of arms 1\u20133, gears 1\u20134, timing belts 1 nd 2. The PMSM drives gear 1, and arm 3 is the straight line utput response. In dynamic modeling of the LCD glass-handling obot system, there are two equations including mechanical and lectrical equations. In this section, the PMSM dynamic equation is ntroduced, and then the mechatronic equation of the LCD glass- andling robot driven by PMSM is formulated as a state-space marix form", " (7) The flux position in the d \u2212 q coordinates can be determined y the shaft-position sensor because the magnetic flux generated rom the rotor permanent magnetic is fixed in relation to the rotor haft position. In Eq. (5) , if i d = 0 , the D -axis flux linkage \u03bbd is xed, and since L md and I fd are constants for a surface-mounted MSM, the electromagnetic torque \u03c4 e is proportional to i q , which s determined by closed-loop control. With the implementation of he field-oriented control, the \u03c4 e of the PMSM drive system can be implified as e = K t i q , (8) t = 3 2 p L md I f d , (9) here K t is the motor torque constant. .2. State-space matrix form Fig. 1 (b) shows the physical mode of the LCD handling-robot ystem. l 1 is the length of O 1 O 2 , l 2 is the length of O 2 O 3 , l 3 is the ength of O 3 O 4 , and \u03c61 is the angular displacement of arm 1. Acording to Eq. (1) , the input voltage can be written as follows: q = R s i q + L q d i q /dt + p \u03bbd \u02d9 \u03c61 . (10) In the mechanical dynamic equation, the dynamic behavior of he LCD handling robot system driven by a PMSM can be described s follows: J 0 \u2212 J 1 cos ( n \u03c61 ) ] \u0308\u03c61 + \u03bc sin ( n \u03c61 ) \u02d9 \u03c6 2 1 = \u03c4m " ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0002200_978-3-319-51222-8-Figure3.22-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0002200_978-3-319-51222-8-Figure3.22-1.png", "caption": "Fig. 3.22 Local solutions = one-element influence functions when the edges are fixed", "texts": [ "12 Examples 173 j ei = \u222b b 0 \u03c3xy(\u03d5i ) dy = \u2212E 2 a (1 + \u03bd) \u00b7 [ b (u2 \u2212 u4) + a (u1 \u2212 u7)+ + x (\u2212u1 + u3 \u2212 u5 + u7) + b 2 (\u2212u2 + u4 \u2212 u6 + u8) ] , (3.116) where the x is the x-coordinate of the vertical cross section. To compute j e1 , we set u1 = 1 and all other ui = 0. For j e2 , we set u2 = 1 and all other ui = 0, etc. The Index e on j ei is to indicate that these are element contributions. The resulting nodal force at each node is the sum over all element contributions. The shape of the influence function on the clamped element, s. Fig. 3.22 and Fig. 3.23, is the local solution which is added to the FE-influence function. 174 3 Finite Elements For a test we compute the influence function for the shear force V , s. Fig. 3.24. The functionwL(x) on the left is clamped at x = 0 and it satisfies the static boundary conditions M(0) = \u2212 6 32 E I V (0) = \u221212 33 E I , (3.117) so thatwL(x) = a \u03d53(x) + b\u03d54(x) is the natural choice and the coefficients a = \u22121 and b = 0 are the solution of the system [ M3(0) M4(0) V3(0) V4(0) ] [ a b ] = E I [ \u22126/32 \u221212/33 ] (3" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0000287_dscc2013-3963-Figure2-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0000287_dscc2013-3963-Figure2-1.png", "caption": "Figure 2. Target tracking control schematic diagram", "texts": [ " In addition, with the strongly nonlinear functions fu(u,v) and fv(u,v), it is hard to apply existing tracking control methods developed for similar underactuated systems with two independent control inputs [33], [34]. We will present a novel hybrid control strategy using the backstepping technique. We make the following assumption: Assumption 1: For the system described by (3), u \u2265 v and u > 0 hold when \u03b1A\u03c9\u03b1 6= 0. This assumption says that the surge velocity of the robot is no less than the sway velocity and that the surge velocity is positive with active tail movement, which are reasonable under most practical scenarios. Fig. 2 shows the setup for the target-tracking control problem. Here, (xs,ys) and (x,y) denote the locations of the target and the robot\u2019s center of mass, respectively, relative to the inertial coordinate frame. Define r as the distance between (xs,ys) and (x,y), and \u03b8 as the orientation error, specified by the angle between the robotic fish heading direction and the line connecting the body center to the target. The pair (r,\u03b8) can be considered as the target location in the body-fixed polar coordinates" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0000130_j.physc.2013.04.021-Figure1-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0000130_j.physc.2013.04.021-Figure1-1.png", "caption": "Fig. 1. Scheme of high temperature superconductor axial flux induction motor.", "texts": [ " [6], Bird and Lipo [7], and Paudel and Bird [8]. Permanent magnets suspension wheel can produce levitation force without movement. However, through the conductors the rapid translational motion of magnetic fields will create an unavoidable magnetic drag force, which tends to create fierce vibration and limit the speed improvement. In this paper, a HTS axial flux induction motor (HTS AFIM) is proposed, for which the primary windings are made from HTS tapes and the secondary is conductive nonmagnetic aluminum conductor, as shown in Fig. 1. When the rotating magnetic field is created by three phase AC currents, eddy currents are induced that give rise to an opposing magnetic field. This opposing field interacts with the source field, creating levitation force and torque simultaneously. When four HTS AFIMs are settled at each corner of a train, as shown in Fig. 2, the total levitation force is large enough to lift a train. Furthermore, because the directions of torques are arranged, the torques can be treated as internal force, which would not output net torque or force", " After that, the magnetic field distribution in HTS coil is calculated by analytic method. Then, an effective way to reduce external magnetic field is presented. Indeed, the critical current of HTS coil can be improved. In addition, as the magnet field in motor will increase the AC losses of HTS windings and then influences the stability of the motor, AC losses of HTS windings in HTS axial flux induction motor are estimated and tested. The proposed axial flux induction motor is manufactured from high temperature superconductor windings, as shown in Fig. 1. In the slots of the primary core, three phase HTS windings are settled to produce the time varying magnetic field. The secondary is nonmagnetic aluminum conductor where the magnetic field will produce eddy current. Because there is no ferromagnetic material in the secondary, the normal force of the axial flux induction motor will perform as repulsive force, which can be treated as levitation force [6\u20138]. Using this normal force, the HTS AFIM can lift itself and some other loads. HTS AFIM can be used as maglev train, as shown in Fig" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0002027_978-3-319-00479-2-Figure14-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0002027_978-3-319-00479-2-Figure14-1.png", "caption": "Fig. 14 FE simulation model of a passenger car", "texts": [ " The FE simulation was done for the load case US-NCAP and the load case \u2018sitting animal\u2019 to calculate the vehicle acceleration, which was used for a multibody simulation to investigate the loads on the occupants. 72 W. Sinz et al. The third step was the investigation of the \u2018concept cage\u2019, which should reduce the deformations into the occupant compartment. For this the \u2018moving space\u2019 of the occupant within the US-NCAP load case was used to define a \u2018safety area\u2019, where no intrusions should occur. To recognize the vulnerability and to identify the mechanisms finite element simulations were performed. Therefore for this analysis a vehicle was used, where a suitable model and FE mesh (Fig. 14) were available. Due to the lack of information regarding statistical data a simulation matrix was created, which was derived from the PC-Crash simulation for the assessment of the accident type, with velocity of vehicle and animal and impact angle of the animal as variation parameter. For the animal the FE models of a camel and of a moose were used. The camel model has a mass of 450 kg (Fig. 15), left side), the moose model has a mass of 350 kg (Fig. 15), right side). A simulation matrix was established with a set of 48 different simulation configurations, with different configurations: \u2022 Fully/partially covered \u2022 Different obstacle (camel, moose) \u2022 Angle of the obstacle (0\u00b0, 45\u00b0, 90\u00b0) \u2022 Velocity of the obstacle (0 kph, 10 kph) \u2022 Velocity of the car (50 kph, 100 kph)" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0002879_sice.2017.8105752-Figure1-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0002879_sice.2017.8105752-Figure1-1.png", "caption": "FIG 1. Inverted pendulum on cart in circle track (left) and side view half-length center of mass (right).", "texts": [], "surrounding_texts": [ "Inverted pendulum is a classical problem in the control system for a long time. The main reason is that it is unstable and nonlinear system. In designing appropriate controller is a challenge for given the system stable and good enough response as we need. In application of this system that we often see is Segway for personal transportation and the others such as self-balancing robot, landing a rocket [5]. In this research, we analyse inverted pendulum moving on the cart in circle tracking. The objective of inverted pendulum is to balance the pendulum in upward vertical and stable position of the cart. It means that we will make the inverted pendulum rod always go to vertical upward and the cart always come to starting position when the pendulum was settled in 10 degree. In this work, the pendulum rod can swing on the cart which moves along a circle track. When comparing with the cart moving only x-axis, the cart on the circle track needs less working space. We derive mathematical model using Euler-Lagrange equation. We aim to build a control system with low-cost budget but still obtain a good performance. Typically, state feedback controllers of dynamical systems need to have enough state measurement but most of real systems do not have enough sensors to measure all states for feedback. The state estimate used for state feedback is part of controller design. State observer and state feedback can be separately designed. Therefore, the combination of state observer and state feedback become a powerful tool of linear-control system design [6]. Inverted pendulum has limited information of the state. In our study, we can measure angular position of pendulum and cart position, then estimate the other unmeasured states. Both measured output and estimated state are fed back to a centralized controller. We design a reduced order observer to handle this problem and then use the estimated state and measured state for feedback with appropriate gain to achieve good response [1]. In control design, we use observer-based controller to stabilize the pendulum and to control the cart position. Controller design consists of two parts. For the first part, we find proper state feedback gain using LQR technique. The latter part, we design reduced-order observer and combine with state feedback to be a reduced-order observer-based controller. In numerical results, we simulate the response of pendulum angle and cart position, then compare the output response between reduced-order and full-order observer-based controller. This paper is organized as follows. Section 2 describes a dynamical model of inverted pendulum on cart. Section 3 presents design of reduced-order observer. Design of LQR and observer based controller is given in section 4. Section 5 presents numerical results of design responses and comparison with the full order observer. Conclusions are given in section 6." ] }, { "image_filename": "designv11_62_0002128_el.2016.3776-Figure1-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0002128_el.2016.3776-Figure1-1.png", "caption": "Fig. 1 Coordinate system of articulated body", "texts": [ " The direct acceleration method uses the end-effector\u2019s velocity and acceleration in Cartesian space to derive the joint space accelerations [5, 6] q\u0308 = J\u22121 x\u0308\u2212 J\u0307q\u0307 ( ) (1) As shown in (1), to obtain the joint space acceleration Jacobian differentiation is necessary. When the Jacobian differentiation is numerically derived, it is possible to obtain results within the bounds of practical error, though this incurs computational costs that make it difficult to use this method for real-time control. In this Letter, to reduce the computational cost without approximations, an analytical method is proposed to differentiate the Jacobian matrix for derivatives of the joint accelerations. Derivation of algorithm: Fig. 1 shows a kinematic chain describing the joint motion of the articulated links. z0i\u22121 is the z-axis of the i\u2212 1( )th coordinate frame from the base frame. Along this axis, the ith link is rotated. The variable designating the joint is qi. d0i\u22121 is the origin vector of the i\u2212 1( )th coordinate frame with respect to the base frame, whereas d0i\u22121, n is the displacement from the i\u2212 1( )th coordinate frame to the nth coordinate frame. From this configuration, the associated Jacobian matrix is described as follows: Ji = z0i\u22121 \u00d7 d0i\u22121, n z0i\u22121 [ ] (2) RONICS LETTERS 16th March 2017 Vol" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0000365_iros.2010.5650226-Figure1-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0000365_iros.2010.5650226-Figure1-1.png", "caption": "Fig. 1. Formulation of the kicking task. The kick request is defined by (pb,vb), the target of the foot motion is denoted by (pf ,vf ), and ph is the hitting spot. rb and rf are the radius of ball and half width of foot respectively. The involved directional vectors(vb, vh and vf ) are unit vectors.", "texts": [ " We will discuss them more detailed in the following sections. In this section we present the basic structure of the kicking motion. Now we formulate the kicking task geometrically as follows: let pb \u2208 R3 be the kicking point, i.e., the point which should be moved by the kicking motion, e.g., the center of mass of the ball. Further, let vb \u2208 R3 with \u2016vb\u2016 = 1 be the intended direction of the movement after kicking, e.g., direction to the goal. A pair (pb,vb) consisting a kicking point and an intended direction is called a kick request. The Fig.1 visualizes a kick request for ball. The whole adaptive kick is proposed as follows: firstly, the target position of kicking foot pf is calculated from the kick request (pb,vb); then a motion planning for the kick is made, e.g., the retraction position of kicking foot pr is determined. As already mentioned, the actual kick motion is performed in two phases: retraction and execution: during the retraction the robot moves the kicking foot to pr in the preparation, and to pf in the execution. All the calculations are on line, i", " The hitting spot is the collision point between ball and foot while kicking, it determines the movement of ball after kicking. On one hand, in order to reach the target direction of ball movement vb, the foot movement vf should be as close as possible to vb. On another hand, in order to achieve powerful kicking, the foot should be retracted as much as possible. But in some cases these two targets can not be reached at the same time. Because the front of Nao\u2019s foot is round, the collision between ball and foot can be simplified as collision between two balls, see Fig.1. The hitting spot ph is calculated to reach the target direction. Thus, we can calculate ph = pb \u2212 vb \u00b7 rb (1) pf = ph \u2212 vb \u00b7 rf (2) vh = vb (3) Note that vf is not determined here, it will be determined by motion planning in the reachable space (See section IV for details). The reachable space of a humanoid robot is defined as the set of points that can be reached by its end effector, with respect to a reference frame of the robot. The reachable space is very important in the planning and control of motion and manipulation" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0001663_icmtma.2011.795-Figure1-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0001663_icmtma.2011.795-Figure1-1.png", "caption": "Figure 1 Dual planetary gear", "texts": [ " when it is under the dangerous condition, the system make the vehicle stability through intervening the steering which is independent of the driver, By means of changing the Angle that the drivers given actively to make the response of the vehicles and the ideal response of the vehicles consistent. II. THE COMPONENT, PRINCIPLE AND FUNCTION OF FRONT ACTIVE STEERING SYSTEM Mechanical active front steering system retained the mechanical components that belong to the traditional steering system, Such as steering wheel , steering column, Rack-and pinion steering mechanism, and steering tie rod, etc, In addition, the biggest characteristic is the steering column which is between the steering wheel and Rack-and pinion steering mechanism installed a set of dual planetary gears, As shown in figure 1,This mechanism is consist of two sets of planetary gears, They share a planet shelf to transmit dynamic. The sun gear on the left connect with the steering wheel, it make the steering Angle input from the steering wheel transfer to the planetary gear pair on the right side through the planetary gear frame, And the planetary gear pair on the right side has two degrees of freedom for steering input, One is the steering wheel Angle transmitted through the planet frame, And another is the input of the gear ring driven by the servo motor through a self-locking worm gear and worm, This is the stack angle" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0002200_978-3-319-51222-8-Figure3.5-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0002200_978-3-319-51222-8-Figure3.5-1.png", "caption": "Fig. 3.5 Local and global coordinate systems", "texts": [ " This curve is added to the polygonal FE-solution, and if this is done on each element, then the resulting shape is the exact curve in Fig. 3.4b. Remark 3.2 To be precise, this holds only true if E A and E I are constants because only then are the element unit displacements \u03d5e i homogeneous solutions of the differential equations. In all other cases, the FE-solution is an approximation, and then, the nodal values are not exact. Since it is important to fully understand thismaneuver, wewant to detail the single steps also in the case of the two-element frame in Fig. 3.5. In step #1, the FE-program assembles the global non-reduced stiffness matrix KG of size 9 \u00d7 9 from the two 6 \u00d7 6 element matrices K (i) e and it deletes the rows and columns of KG which correspond to fixed degrees of freedom so that KG reduces to a 5 \u00d7 5 matrix K . In step #2, the FE-program computes for each element e the six fixed-end forces pei = \u222b le 0 pe(x)\u03d5e i (x) dx . (3.20) 3.3 Adding the Local Solution 147 This is a somewhat symbolic notation because pe(x) actually can have two directions, in local xe or local ze direction, and the unit displacements \u03d5e i (x) are understood to be the corresponding displacements, two in xe direction and four in ze direction" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0001344_pedstc.2011.5742411-Figure1-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0001344_pedstc.2011.5742411-Figure1-1.png", "caption": "Figure 1. BDFM system and its excitations [10]", "texts": [], "surrounding_texts": [ "Keywords-Brushless Doubly-Fed Machine; 2-D finite element model; magnetodynamic; stator fault.\nI. INTRODUCTION The idea of having two stator windings for variable speed applications is proposed in [1-3]. Dual stator winding machines have been categorized as split-wound and selfcascaded [4]. The self-cascade machine or Brushless DoublyFed Machine (BDFM) was introduced by Hunt. In a BDFM stator windings configuration is similar to stator winding of an induction machine. One of these windings is connected directly to the grid, which is called power winding. The other is fed by a partially VA-rated (e.g. 30% of machine\u2019s rating) bidirectional converter [5], which is called control winding. These windings have different pole number to avoid having coupling between them and usually are excited at different frequencies [6].\nBDFM has three different modes of operation depend on the connection of the control winding. First, simple induction mode when the control winding is open-circuited. Second, cascade mode when the control winding is short-circuited. Third, synchronous mode when the control winding is supplied with a certain frequency regarding the rotor speed. The BDFM is expected to run more efficiently in synchronous mode. In this mode, the rotor speed is independent of the torque, in contrast to the two other modes.\nIt requires a special rotor structure that has some nested loops on the circumference of the rotor to incorporate the effects of cascade connection [7-9]. The BDFM and its rotor structure for 3 loops per nest are shown in Figs. 1 and 2, respectively.\nDuring stator fault, the air gap field may contain components having all possible pole pair numbers. In this situation the magnetic coupling between power and control windings is not equal to zero anymore. So the coils apply forces to each other. As a result, the effective performance of the machine will decreases and sever destruction may occur. Therefore detection of these faults is important.\nIn this paper, to analyze stator faults a finite element model of the machine is introduced then simulation results of the finite element model is compared to the dynamic model of the machine [11], which used generalized harmonic analysis through the coupled circuit [12], to verify the accuracy of the finite element model. The following assumptions are made [13]: \u2022 The stator and the rotor are modeled by two smooth\nconcentric cylinders of infinitely permeable iron. \u2022 The stator windings and the rotor bars are replaced by\nequivalent point conductors lying on the stator and rotor surfaces, respectively.\n\u2022 Flux crosses the air gap in radial lines.\n978-1-61284-421-3/11/$26.00 \u00a92011 IEEE 169", "II. FINITE ELEMENT MODEL A 2-D magnetodynamic finite element model is represented to verify the validation of the proposed dynamic model. This model allowing the simulation of the healthy and faulted machine with its special rotor bar configuration. The machine is modeled in a 2-D domain using the Maxwell equations to formulate the field behavior and the finite element method (FEM) to discrete the domain. The formulation uses the magnetic vector potential as unknown, the Galerkin method to obtain the set of equations to be solved numerically, the Euler recurrence method to discrete the temporal derivatives and the successive approximation and Newton-Raphson method to consider the nonlinear characteristic of magnetic material. The movement is taken into account by means of the moving band technique, the Maxwell stress tensor and the mechanical oscillation equation. The general field equation describing an electromagnetic system in 2-D domain when the magnetic potential and current density only have axial variation is as,\n0 (1)\nThe machine structure in 2-D cases can be divided into 4 different zones with the same equations for elements located in each one. The explanations and formulation of these zones for BDFM are mentioned in subsections A to D.\nThis zone includes the elements of the stator and rotor cores (Fig. 3). There is no independent voltage or current source in this region. The field equation of the elements in zone 1 is stated in (2).\n0 (2)\nBy applying the Galerkin and the Euler recurrence methods, the matrix form equation of (2) is obtained as (3).\n(3) where, is the magnetic potential vector of nodes and and\nmatrices are defined in (4) and (5), respectively [14].\n(4)\n1 0.5 0.5 0.5 1 0.5\n12 0.5 0.5 1\nDQ t \u03c3 \u23a1 \u23a4 \u23a2 \u23a5= \u23a2 \u23a5\u0394 \u23a2 \u23a5\u23a3 \u23a6\n(5)\nWhere,\n(6)\nand are the global co-ordinate of the ith node in a triangular element with three nodes.\nTo consider the nonlinear characteristic of the magnetic material of the stator and rotor core (e.g. dependency of to\n), the set of equations must be solved recursively. A desired convergence rate can be obtained by using the result of the successive approximation method after several iterations as the initial value in the Newton-Raphson method [14]. The Newton-Raphson recursive equation is stated in (7) for the elements of zone 1. \u2206 \u2206 , \u2206 (7)\nis the Jacobian matrix and its procedure to extract is explained in [14].\nThe elements located in the air gap and interior regions of the stator and rotor slots except the conductor regions, are in this zone. There are no independent source and induced currents and the magnetic permeability is constant. The matrix form field equation of the elements in this zone can be written as (8).\n\u2206 0 (8)\nThe special issue about this zone is the moving band locating in the air gap. Fig. 4 provides some details of the mesh, especially in the air gap region. Introducing the movingband technique involves splitting the air gap into three layers, where the middle one is the moving band. In order to minimize the element distortion during the rotation, only the elements on the moving band were remeshed at every step time [15]. The remeshing scheme during the rotor movement is explained in [14-16].\nThe elements in this zone discrete the rotor conductor's regions. The field equation and its matrix form are as (9) and (10), respectively.\n0 (9)\n\u2206 \u2206 (10) where," ] }, { "image_filename": "designv11_62_0000287_dscc2013-3963-Figure1-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0000287_dscc2013-3963-Figure1-1.png", "caption": "Figure 1. Schematic representation of the tail-actuated robotic fish in planar motion.", "texts": [ " Using Lyapunov analysis, we show that the robot will converge to the target location. Simulation and experiment results are presented to illustrate the effectiveness of the proposed control approach. The remainder of the paper is organized as follows. We first review the dynamic model and its averaging in Section II. The target-tracking hybrid control scheme is developed and analyzed in Section III. Simulation and experiment results are presented in Section IV. Finally, concluding remarks are provided in Section V. Fig. 1 shows a schematic of a tail-actuated robotic fish constrained to planar motions [7], [8]. In the figure, (x,y) and \u03c8 denote the position and orientation, respectively, of the robot\u2019s center of mass relative to the inertial coordinate frame XYZ. The surge, sway, and angular velocity, expressed in the bodyfixed coordinates UOV, are denoted by u, v and \u03c9, respectively. \u03b1 denotes the tail deflection angle with respect to the negative U\u2212axis, and \u03b2 denote the angle of attack, formed by the direction of \u21c0 V = (u,v) with respect to the positive U\u2212axis" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0001620_978-1-4614-3475-7_2-Figure2.30-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0001620_978-1-4614-3475-7_2-Figure2.30-1.png", "caption": "Fig. 2.30 Example 2.7", "texts": [ " The total area is given by A = A1 +A2 = \u03c0 ( a2 \u2212 r2) . The coordinate xC of the mass center is given by xC = xC1 A1 + xC2 A2 A1 +A2 = xC2 A2 A1 +A2 = xC2 A2 A = (a\u2212 r)\u03c0 a2 \u03c0 (a2 \u2212 r2) or xC ( a2 \u2212 r2)= a2 (a\u2212 r). If xC = r, the previous equation gives r2 + ar\u2212 a2 = 0, with the solutions r = \u2212a\u00b1 a \u221a 5 2 . Because r > 0, the correct solution is r = a (\u221a 5\u2212 1 ) 2 \u2248 0.62a = 0.62(2) = 1.24 m. Example 2.7. Locate the position of the mass center of the homogeneous volume of a hemisphere of radius R with respect to its base, as shown in Fig. 2.30. Solution The reference frame is selected as shown in Fig. 2.30a, and the z-axis is the symmetry axis for the body: xC = 0 and yC = 0. Using the spherical coordinates, z = \u03c1 sin\u03d5 , and the differential volume element is dV = \u03c12 cos\u03d5d\u03c1d\u03b8d\u03d5 . The z coordinate of the mass center is calculated from zC \u222b \u222b \u222b \u03c12 cos\u03d5d\u03c1d\u03b8d\u03d5 = \u222b \u222b \u222b \u03c13 sin \u03d5 cos\u03d5d\u03c1d\u03b8d\u03d5 , or zC \u222b R 0 \u03c12d\u03c1 \u222b 2\u03c0 0 d\u03b8 \u222b \u03c0/2 0 cos\u03d5d\u03d5 = \u222b R 0 \u03c13d\u03c1 \u222b 2\u03c0 0 d\u03b8 \u222b \u03c0/2 0 sin\u03d5 cos\u03d5d\u03d5 , or zC = \u222b R 0 \u03c13d\u03c1 \u222b 2\u03c0 0 d\u03b8 \u222b \u03c0/2 0 sin \u03d5 cos\u03d5d\u03d5 \u222b R 0 \u03c12d\u03c1 \u222b 2\u03c0 0 d\u03b8 \u222b \u03c0/2 0 cos\u03d5d\u03d5 . (2.64) From (2.64), after integration, it results zC = 3R 8 . Another way of calculating the position of the mass center zC is shown in Fig. 2.30b. The differential volume element is dV = \u03c0 y2 dz = \u03c0 (R2 \u2212 z2)dz, and the volume of the hemisphere of radius R is V = \u222b V dV= \u222b R 0 \u03c0 (R2 \u2212 z2)dz=\u03c0 ( R2 \u222b R 0 dz\u2212 \u222b R 0 z2 dz ) =\u03c0 ( R3\u2212R3 3 ) = 2\u03c0 R3 3 . The coordinate zC is calculated from the relation zC = \u222b V zdV V = \u03c0 ( R2 \u222b R 0 zdz\u2212 \u222b R 0 z3 dz ) V = \u03c0 V ( R2 R2 2 \u2212 R4 4 ) = \u03c0 R4 4V = \u03c0 R4 4 ( 3 2\u03c0 R3 ) = 3R 8 . syms rho theta phi R real % dV = rho\u02c62 cos(phi) drho dtheta dphi % 0 0 \ud835\udf12 if \ud835\udc5f = 0 and \ud835\udf19 = 2arctan( \ud835\udc5f \ud835\udc5f\ud835\udc51 ). From (5), it can be observed that when \ud835\udc5f \u2192 0, the desired relative course angle approaches the direction along the radius away from the origin, and when \ud835\udc5f \u2192 \u221e, the desired relative course directly points to the origin" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0000260_2506095.2506124-Figure1-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0000260_2506095.2506124-Figure1-1.png", "caption": "Figure 1. Schematic diagram of quadrotor", "texts": [ " Section III gives the basic idea of MPC and its mathematical formulation. In Section IV, we provide the simulation results and conclusions are given in Section V. The dynamical model of quadrotor can be obtained from the Euler-Lagrange equations with external forces as: d dt ( \u2202L \u2202q\u0307 ) \u2212 \u2202L \u2202q = F (1) Where, q = (x, y, z, \u03c8, \u03b8, \u03c6) \u2208 R6 represents generalised coordinates. Here (x, y, z) denotes position of center of mass of quadrotor with respect to a fixed frame of reference and (\u03c8, \u03b8, \u03c6) are the three Euler angles - Yaw, Pitch and Roll as shown in Figure 1 [3]. The Lagrangian L is the difference between total kinetic energy (rotational K.E. + translational K.E.) and potential energy. We partition the overall model into 1. translational coordinates- \u03b1 = (x, y, z) \u2208 R3 2. rotational coordinates- \u03b2 = (\u03c8, \u03b8, \u03c6) \u2208 R3 The Lagrangian is L(q, q\u0307) = m 2 \u03b1\u0307T \u03b1\u0307+ 1 2 \u03b2\u0307TJ\u03b2\u0307 \u2212mgz (2) The matrix J is the inertia matrix for the rotational kinetic energy of quadrotor. The total force F = (F\u03b1, \u03c4) consists of translational force F\u03b1 and the generalized torques \u03c4 on variables \u03b2 as \u03c4 = \u03c4\u03c8\u03c4\u03b8 \u03c4\u03c6 (3) We define u as the total thrust produced by four quadrotor motors : u = f1 + f2 + f3 + f4 (4) Solving the Euler-Lagrange equation we obtain, m\u03b1\u0308 = u \u2212 sin \u03b8 cos \u03b8 sin\u03c6 cos \u03b8 cos\u03c6 + 0 0 \u2212mg (5) J\u03b2\u0308 = \u2212C(\u03b2, \u03b2\u0307)\u03b2\u0307 + \u03c4 (6) Where C(\u03b2, \u03b2\u0307) is referred as Coriolis matrix which takes into account centrifugal and gyroscopic terms" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0002300_s11740-017-0739-2-Figure1-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0002300_s11740-017-0739-2-Figure1-1.png", "caption": "Fig. 1 Process forces and chip geometry", "texts": [ " The paper uses the generated data to develop an empirical calculation model of the process forces and analyze the influencing variables. The machining processes with end-milling cutters result in certain process forces. With regard to process design and modeling, it is important to gain an in-depth understanding of these forces and their associated mechanical tool loads [5\u20137]. Process forces directly influence the milling process. On one hand, an increased active force leads to tool deflection; on the other hand, process forces result in a mechanical load on the individual cutting edge [8]. Figure\u00a01 shows a cutting tool and the resulting process forces during the milling process. The extended cycloid path of the cutting edge determines the comma-shaped chip geometry, which has a decisive influence on the process forces. The uncut chip volume is dependent upon the following process parameters: feed per tooth fz, width of cut ae, and depth of cut ap. Therefore, the cross section of the undeformed chip \u201dA\u201d and the specific cutting force are critical, as the cutting force affects the mechanical load on the cutting edge [9]", " The colors of the planes indicate the effect of tool wear. The cutting force is represented as a function of the cutting speed vc and feed per tooth fz, as well as a function of the depth and width of cut ap and ae (refer Eq.\u00a02). This equation clearly describes the increasing cutting force Fc as being based on the increasing feed per tooth and width of cut. This correlation is due to the increasing thickness of the undeformed chip h, which is influenced by the process parameters fz and ae (see Fig.\u00a0 1). However, based on the representation of the cross section of the undeformed chip, the depth of cut shows a minor influence on the cutting force. This is reflected in the smaller coefficients for the depth of cut in Eq.\u00a02. In the case of increasing cutting speed vc, the graph shows declining cutting forces, as reflected in the negative coefficient for cutting speed. The influence of cutting speed on the machining process is highly relevant. On one hand, there is an increasing material removal rate, which leads to an increasing cutting force" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0001620_978-1-4614-3475-7_2-Figure2.49-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0001620_978-1-4614-3475-7_2-Figure2.49-1.png", "caption": "Fig. 2.49 Problem 2.9", "texts": [ " The half-circle has a radius, R, equal to 30 mm. The coordinate axes are aligned so that the origin is at the bottom of the object at the center point of the half-circle. This half-ring has an inner radius, r, of 20 mm, and an outer radius, R, of 30 mm. The extrusion height for the half-ring, h, is equal to 25 mm. The density of the object is uniform and will be denoted \u03c1 . Find the coordinates of the mass center. 2.9 The ring with the circular cross section has the dimensions a = 3 m and b = 5 m. Determine the surface area of the ring (Fig. 2.49). 2.10 The belt shown in Fig. 2.50 has the dimensions of the cross area: a = 27 mm, b = 66 mm, and c = 58 mm. The radius of the belt is r = 575 mm. Find the volume of the belt (Fig. 2.50). 2.11 Determine the moment of inertia about the x-axis of the shaded area shown in Fig. 2.51 where m = h/b and b = h = 2 m. Use integration. 2.12 Determine the moment of inertia about the y-axis of the area shown in Fig. 2.52 where a = 2 m and b = 6 m. Use integration. ba c a r Fig. 2.50 Problem 2.10 y = m x b y x h Fig" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0002604_eurocon.2017.8011204-Figure1-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0002604_eurocon.2017.8011204-Figure1-1.png", "caption": "Fig. 1: IPMG equivalent circuits", "texts": [], "surrounding_texts": [ "The reliable control of the IPMG requires an appropriate mathematical model. It means that the model has to be detailed enough to satisfy the successful control requirements, but on the other hand it has to be proper for fast calculation. For these requirements it is not possible to use IPMG model in abc reference frame, because it consists of three non-linear differential equations. To achieve the controllable model it is necessary to use Clark and Park transformations [18]. Voltage equations of the IPMG in dq reference frame can be written as: usd = Rsisd + Ld disd dt \u2212 \u03c9reLqisq, (1) usq = Rsisq + Lq disq dt + \u03c9reLdisd + \u03c9re\u03c8r, (2) where Rs is stator resistance, isd and isq are d and q-axis stator currents, Ld and Lq are d and q-axis inductances and \u03c9re is rotor speed. In Figs. 1a and 1b d and q-axis equivalent circuits of the IPM machine are shown. In steady state, when there is no transients and d dt \u2192 0, (1) and (2) become: usd0 = Rsisd0 \u2212 \u03c9re0Lqisq0, (3) usq0 = Rsisq0 + \u03c9re0Ldisd0 + \u03c9re0\u03c8r, (4) where vsd0, vsq0, isd0, isq0 and \u03c9re0 are voltages, currents and rotor speed values in the steady-state." ] }, { "image_filename": "designv11_62_0001428_ajpa.21280-Figure2-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0001428_ajpa.21280-Figure2-1.png", "caption": "Fig. 2. Marker positions. Infrared reflective markers were attached over three sites on the gibbon\u2019s right side: the anterior superior iliac spine (ASIS), the greater trochanter (GT), and the lateral epicondyle of the femur (LE). We considered excursion of the hip joint angle (y; ASIS-GT-LE angle in three-dimensional space) to reflect hip flexion/extension motion.", "texts": [ " We measured the right lower trunk and thigh of the gibbon during continuous-contact brachiation with an infrared detecting 3D motion analyzer (ELITE System; BTS, Milan, Italy) at sampling rate of 100 Hz. Three hemispherical markers, which were made of styrene foam and covered with infrared reflective sheets, were attached on the shaved skin of the gibbon at three sites: the anterior superior iliac spine (ASIS), lateral to the hip (close to the greater trochanter; GT), and the lateral epicondyle of the femur (LE). We regarded angle ASIS\u2013GT\u2013 LE in 3D space as the hip joint angle, which we used as an index of hip flexion/extension (Fig. 2). In addition, an infrared-reflective sheet was bound around the right wrist to measure hand contact and release, which were used to determine the brachiation gait cycle. The markers were 1 cm in diameter and weighed less than 1 g, so as not to disturb any of the gibbon\u2019s motions. We reconstructed 3D coordinates of each marker using the ELITE software (BTS, Milan, Italy) based on the two sets of two-dimensional coordinates collected using two charged-coupled cameras (Ferrigno and Pedotti, 1985)" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0001712_978-1-4614-3997-4_4-Figure1-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0001712_978-1-4614-3997-4_4-Figure1-1.png", "caption": "Fig. 1. Propulsion of two model swimming microorganisms: (a) a pusher (such as a bacterium) exerts a propulsive force near its tail, whereas (b) a puller (such as a microalga) exerts a thrust near its head", "texts": [ " In nature, numerous swimming mechanisms exist at low Reynolds numbers, which all rely on non-reciprocal shape deformations as prescribed by Purcell\u2019s famous scallop theorem [34]. Most microorganisms make use of flexible appendages named flagella, which are actuated in a non-reciprocal fashion, thereby exerting a net thrust on the surrounding fluid. This is the case of many types of bacteria such as the common Escherichia coli and Bacillus subtilis, which use a bundle of flagella for propulsion, and of some types of microphytes such as Chlamydomonas reinhardtii, which beats two flagella in a breaststroke-like fashion (Fig. 1). While the resulting propulsive force Fp will in general be time-dependent, we will assume here for simplicity that it is steady: its value may be interpreted as a time average over one beat cycle (an approximation that is not necessarily easy to justify as unsteady effects may also have an impact on hydrodynamic interactions). If gravitational effects can be neglected, i.e. if the microorganism and the fluid have nearly matching densities, the swimmer is force-free and must therefore exert an equal and opposite drag force Fd = \u2212Fp on the fluid: this drag force is likely to be exerted mostly by those parts of the body that do not contribute to propulsion (i.e. the cell body for bacteria and microalgae). Because the application points of Fp and Fd differ by a distance l (of the order of the organism size), the net leading-order effect on the surrounding fluid is that of a force dipole, whose sign may depend on the mechanism for swimming. In the case of a bacterium (Fig. 1a), the propulsive force is exerted near the rear of the particle, and such a swimmer will be called a pusher. Conversely, an alga swimming the breaststroke (Fig. 1b) exerts a thrust near its front, and will be called a puller. The force dipole exerted by a swimmer can be characterized by the socalled stresslet S, which is a second-order tensor defined as the symmetric first moment of the two forces: S = \u2212 \u2211 i [ 1 2 (xiFi + Fixi)\u2212 1 3 (xi \u00b7Fi)I ] , (1) where the sum is over the two forces Fp and Fd. In Eq. 1, xi is the point of application of force Fi, and the last term on the right-hand side involving the idem tensor I is added to make S traceless. In the case of the two swimmers illustrated in Fig. 1, and defining the director p as a unit vector pointing in the direction of swimming, it is straightforward to simplify this expression to: S = \u00b1Fl ( pp\u2212 I 3 ) , (2) with F = |Fp|, and where the minus sign corresponds to the case of a pusher and the plus sign is for a puller. In the following, we introduce the dipole strength \u03c30 = \u00b1Fl, with \u03c30 < 0 for a pusher and \u03c30 > 0 for a puller. Note that the magnitude of \u03c30 is also related to the swimming speed U0 of the particle. Indeed, a force balance on the body of the organism yields F \u221d \u03bcU0l where the proportionality constant depends on the exact shape, which leads to \u03c30 \u221d \u03bcU0l 2. In the following, it will be convenient to define a dimensionless stresslet strength as \u03b1 = \u03c30/\u03bcU0l 2, which is an O(1) constant of the same sign as \u03c30. Of course, the description of Fig. 1 in terms of two equal and opposite point forces is simplistic, and in reality the microorganism exerts a distribution of stresses over the entire surface of its body. The definition of the stresslet Eq. 1 is then easily generalized as S = \u2212 \u222b S [ 1 2 (xf + fx)\u2212 1 3 (x \u00b7 f)I ] dS, (3) where the integral is over the surface of body, and f(x) is the traction (force per unit area) at any point x on the body. For an axisymmetric microorganism, this expression must also simplify to S = \u03c30 ( pp\u2212 I 3 ) , (4) where the value of \u03c30 will depend on the details of the traction distribution" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0000358_iccse.2010.5593596-Figure9-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0000358_iccse.2010.5593596-Figure9-1.png", "caption": "Figure 9. Spur gear mechanism constructed in Pro/E", "texts": [ " So, gear mechanisms are used widely and the shape of gear teeth need to be specified. Generally speaking, according to the difference of gear tooth profiles, gears can be classified into two categories: involute tooth gears and cycloidal tooth gears. According to the difference of gear tooth structural style, gears can be classified into spur gears, helical gears, herringbone gears, bevel gears, and so on. [5] ADAMS specializes in the simulations of kinematics and dynamics. Complicated 3D models can be built through 3D CAD software such as Pro/E, and then imported into ADAMS. Figure 9 shows a meshing pair of spur gears assembled in Pro/E, which is a sort of powerful 3D CAD software. Parametric gear models in the same style can be built through its Parameter and Relation module. Thus, gear models can be built rapidly. Through the MECH/Pro module, the plug-in between Pro/E and ADAMS, Pro/E models can be totally transferred into ADAMS software as shown in Figure 10. Add a type of \u2018solid to solid\u2019 contact analysis between two gears, and put \u2018motion\u2019 on the joint of active gear (smaller one in Figure 10), and then mesh state of a pair of spur gear mechanism can be simulated" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0003235_rnc.4470-Figure2-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0003235_rnc.4470-Figure2-1.png", "caption": "FIGURE 2 Pendubot", "texts": [ " Then, the SM controller (31) stabilizes the estimated error e\u0302k (26), achieving a locally ultimately bounded solution at the real error system ek (24). With the ultimately bounded solution of the error system, an approximate OT is achieved in system (1), satisfying the condition (Tsm). In this section, to show the effectiveness of the proposed control scheme, the tracking problem for the Pendubot system is considered and a simulation is performed. The Pendubot is an NMP underactuated mechanical system (Figure 2). The model and parameters used for simulations were obtained from the work of Serrano-Heredia et al.23 The Pendubot continuous-time dynamics are described by the following Euler-Lagrange equation:[ D11(q) D12(q) D21(q) D22(q) ] [ q\u03081 q\u03082 ] + [ C1(q, q\u0307) C2(q, q\u0307) ] + [ G1(q) G2(q) ] + [ F1(q\u0307) F2(q\u0307) ] = [ \ud835\udf0f 0 ] , where q = [q1 q2]T are the angular positions of both links; q1 is the actuated link, while q2 is the unactuated one; D(q) = [ D11(q) D12(q) D21(q) D22(q) ] , C(q, q\u0307) = [ C1(q, q\u0307) C2(q, q\u0307) ] , G(q) = [ G1(q) G2(q) ] , and F(q\u0307) = [ F1(q\u0307) F2(q\u0307) ] are the inertia matrix, the vector of Coriolis and centripetal torques, the vector of gravitational terms, and the vector of viscous frictional terms, respectively" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0003913_s00773-019-00675-8-Figure7-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0003913_s00773-019-00675-8-Figure7-1.png", "caption": "Fig. 7 Propeller hydrodynamic forces and moments", "texts": [ " This approach is a combination of solution of Reynolds equation of inner loop and shaft alignment calculation of outer loop (See Fig.\u00a05). The ship investigated is a super handy-max bulk carrier designed by Shanghai Merchant Ship Design and Research Institute. A single pitch-fixed screw is utilized (Fig.\u00a06) for the geometries of the ship and propeller. Table\u00a02 summarizes the principle dimensions of the ship and propeller. 1 3 The propeller hydrodynamic forces and moments consist of the longitudinal, horizontal, and vertical components, as shown in Fig.\u00a07. The longitudinal forces and moments are the propulsive loads that are related to the propulsive efficiency, whereas the horizontal and vertical forces and moments are the lateral loads that determine the bearing loads. Research shows that lateral loads can be remarkably increased in an oblique flow condition, such as the turning maneuver process [9, 17, 18]. The lateral flow in the oblique flow condition significantly affects the propeller lateral loads, especially the horizontal and vertical moments, which have a large effect on the safety of the shaft system" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0002109_1464419317689946-Figure3-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0002109_1464419317689946-Figure3-1.png", "caption": "Figure 3. Vector from point P to point Q in the geometries of moving bodies.", "texts": [ " As in equation (2), the vector in equation (3) is a function of s1, s2 and the reference position and orientation coordinates of the moving body. Contact search. When the contact points, expected or actual, are found on the surfaces of the two moving bodies, these contact points are presumed to be relative extremes of the function that gives as output the length of the vector from a point P to a point Q, which are arbitrary points at geometries on the bodies. The function for the length of the vector PQ is calculated using the vector from point P to point Q, which are located, respectively, on the surfaces of bodies i and j, as in Figure 3, and can be described in the following way rPQ \u00bc Rj \u00fe uj,Q Ri \u00fe ui,P \u00f04\u00de Calculation of the extreme of the modulus of vector rPQ given in equation (6) provides four non-linear algebraic equations that can be solved to determine the surface parameters si,1, si,2, sj,1 and sj,2 associated with the contact points of the two bodies in contact. The resulting system of non-linear equations is given by27 nTi,Ptj,Q,1 \u00bc 0 nTj,Qti,P,1 \u00bc 0 rTPQtj,Q,2 \u00bc 0 rTPQti,P,2 \u00bc 0 9>>>>= >>>>; ) si,P,1 si,P,2 sj,Q,1 sj,Q,2 2 6664 3 7775 \u00f05\u00de Solving equation (4) gives the surface parameters si,P,1, si,P,2, sj,Q,1 and sj,Q,2, which in turn, using equation (5), can be used to obtain the precise location of the contact points" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0001078_epe.2013.6631851-Figure1-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0001078_epe.2013.6631851-Figure1-1.png", "caption": "Figure 1: Three-phase 8/6-pole SRG: (a) cross-sectional profile and (b) equivalent circuit of one phase winding.", "texts": [ " In addition, a piecewise quadratic equation is used, instead of a piecewise linear one, to improve the performance further by reducing the rate-of-change of the current reference. To find optimal parameters for the non-unity TSF at different operating speeds, a sequential quadratic programming method [9] is used. The effectiveness and resulting improvement in the performance of the proposed torque minimization technique will be demonstrated through simulation results. Before introducing a new torque minimization technique, a brief analysis of the SRG\u2019s electromechanics will be given in this section. Fig. 1(a) shows the cross-sectional profile of a conventional four phase 8/6- pole SRG and Fig. 1(b) shows its equivalent circuit of one phase winding. In Fig. 1(a), only one phase winding is shown for the sake of simplicity. Both the stator and the rotor of the SRG are usually made of steel laminations. While each salient pole on the stator has a concentrated winding, the rotor does not have any winding or permanent magnet, but a chunk of laminated steel shaped like a gear. This structural feature makes the SRG simple, robust, and low-cost. The equivalent circuit of one phase winding can be simplified with a phase resistance, Rs, connected to a phase inductance, L(\u03b8r, i) in series" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0000300_imece2012-87081-Figure3-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0000300_imece2012-87081-Figure3-1.png", "caption": "Fig. 3: Fabricated wrinkled patterns with cracks and dislocations. (a) Optical microscope image at 40X, (b) Atomic force microscope image and (c) Close-up view of dislocations.", "texts": [ " Herein, we have used polydimethylsiloxane (PDMS) as the elastomeric base and plasma oxidation for generating the top glassy layer. Thus, the top glassy layer acts as the \u2018plate\u2019 and the elastomeric PDMS layer acts as the \u2018foundation\u2019. A PDMS strip 20 mm long, 10 mm wide and 3 mm thick was loaded onto a linear stage and stretched by ~15% along its length. The entire stage was then loaded into a plasma oxidation chamber and the PDMS strip was exposed to air plasma for 45 minutes. Wrinkled patterns that were obtained on releasing the pre-stretch are shown in Fig. 3. The wavelength and amplitude of the fabricated wrinkles were measured via atomic force microscope (AFM). These are listed in Table 1. The AFM image of the wrinkles is shown in Fig. 3(b). In our experimental technique, the thickness of the top plate layer is controlled via exposure time of the plasma oxidation process. As techniques for direct measurement of the plate thickness are not available, we rely on analytical models to estimate the thickness. The wavelength (\u03bb) is related to the plate thickness (h) and the material properties by [13]: 3 1 2 2 13 1 2 pf fp E E h (1) In Eq. 1, subscript \u2018p\u2019 denotes the plate and subscript \u2018f\u2019 denotes the foundation, E is the Young\u2019s modulus and \u03bd is the Poisson\u2019s ratio", " Density of defects depends on the (i) magnitude of the applied pre-stretch and (ii) local imperfections such as voids or variations in material properties. Herein, we link these sources to the observed defect modes and identify the limits of the feasible operating region for a target pattern quality. There are two primary defect modes that are observed during 1-D wrinkling. These are (i) transverse cracks that propagate throughout the surface and (ii) local period-doubling followed by period-halving bifurcations that manifest as dislocations. Both of these defects are labeled and shown in Fig. 3. Transverse fractures Periodic fractures that run almost perpendicular to the wrinkles are observed in our experiments (Fig. 3). These cracks arise due to the transverse stress that is generated during compression of the top plate. Compression of the plates along their length generates a tensile stress along the width. Cracks are formed when the transverse tensile stress exceeds the fracture strength of the plate. The fracture of the top glassy Downloaded From: http://proceedings.asmedigitalcollection.asme.org/pdfaccess.ashx?url=/data/conferences/asmep/76593/ on 02/06/2017 Terms of Use: http://www.asme.org/about-asme/terms-of-use 4 Copyright \u00a9 2012 by ASME layer has been studied earlier by Rand et al" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0003606_012033-Figure1-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0003606_012033-Figure1-1.png", "caption": "Figure 1. Calculation scheme.", "texts": [ " The obtained results can be used in all mechanic engineering sites operating on small-sized thrust bearings. 3.1. Formulation of the problem The steady-state flow of an electrically conductive liquid lubricant in the working gap of a thrust sliding bearing with an inclined liner operating in the hydrodynamic lubrication mode with a porous layer on the guide surface under the action of an electromagnetic field is considered. It is assumed that the liner is stationary, and the way moves in the gap at a constant velocity u* (Figure 1). In the Cartesian coordinate system, the equations of the adapted and deformable contours 1, 2 and 3 and the guide with a porous layer on its surface will be written as: 01 : \u03b1- sin \u03c9 ( )\u0421 y h x tg a x H x \u2013 the equation of the adapted non-deforming contour of the slider. Let us look for the equation of a deformable adapted contour in the form: 2 0: tg\u03b1 sin \u03c9 \u03c6 ( ) x \u0421 y h x a x a H x L , (1) 3 1: tg\u03b1\u0421 y h x \u2013 equation of a non-deforming contour. MEACS2018 IOP Conf. Series: Materials Science and Engineering 560 (2019) 012033 IOP Publishing doi:10" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0000704_ssp.165.359-Figure4-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0000704_ssp.165.359-Figure4-1.png", "caption": "Fig. 4. Computational model of numerical calibration", "texts": [ " This kind of arrangement eliminates the influence of rail bending caused by the non-axial wheel position. Scheme of meter circuit is given in Fig. 3. The significant sensitivity of measurement system on the distance between the active strain gauges proceeds at the rail web from the rail head [5]. It is necessary to conduct sensors calibration for the qualitative and quantitative character of results of measurements. The experimental calibration for typical running rails was described in [5]. This calibration can also be accomplished by the use of numerical methods e.g. FEM [5, 6]. Fig. 4 illustrates the discrete model of part of rail that undergoes loading from the road wheel. The stresses and strains field was determined in the rail cross-section under the loading (Fig. 5). The strains distribution at the height of the rail web reveals significant non-uniformity. The indications of strain gauge are proportional to the integral of strains at its active length. As a result of differences in strain gauges location and the differences in the subgrade stiffness the different values of indications are obtained (Fig. 6). The measuring error may reach up to 15%. The separate problem is the significant sensitivity of meter circuit on the distance of active strain gauges on the rail web from the rail head [1, 2]. In order to eliminate this cause of failure and the systematic failure caused by the differences of particular strain gauges the measurements for three positions are conducted (Fig. 4). The simplest determination of centre of gravity of the machine is where the centre of gravity of undercarriage is situated in the rotation axis. From measurements in position I and II: ( ) ( ) ( ) ( ) ( ) ( )lrrlllrr lrrr lrllrr lrr lr rc IIII II II I II I GGGG GG GG G GG G b x \u03b4\u03b4\u03b4\u03b4 \u03b4\u03b4 \u03b4\u03b4\u03b4\u03b4 \u03b4\u03b4 +++++++ +++ = = +++++ ++ = + =+ 1111 11 )2()2( )2( 2 1 (1) where: xc, yc - location of machine centre of gravity, Gl, Gr - actual loading of particular drive sets (l \u2013 left, r \u2013 right), Rl, Rr - measured loading of particular drive sets (l \u2013 left, r \u2013 right), \u03b4l, \u03b4r - relative systematic measurement errors of the sensor at the left and right rail, b - rails spacing", " (4) From measurements in position III: ( ) ( ) ( ) + \u2212 + = + + + + = + =+ c c l r lr r l l r r r r lr rc xb xb R R RR R RR R GG G b y I I IIIIII III IIIIII III IIIIII III 2 2 11 1 2 1 \u03b4\u03b4 \u03b4 . (5) Final equations: b xb xb R R RR R y b RRRR RR x c c l r lr r c llrr lr c I I IIIIII III IIIIII III \u2212 + \u2212 + = \u2212 +++ + = 2 1 2 2 , 2 1 . (6) The above method was used to determine the centre of gravity of LZKS 1600 stacker reclaimer (Fig. 1) with the 800 Mg mass. 9 measurement series were conducted for superstructure position in relation to undercarriage described in Fig. 4 as well as for three positions of bucket wheel boom. In each series at least 6 measurements were carried out. The example run of measurement signal at the machine driving forward for position I with bucket wheel boom at lower position is shown in Fig. 8. Fig. 9 presents the positions of centres of gravity for different boom orientations in relation to pitch diameter of a slewing bearing. The confidence intervals were set to 95%. The confidence intervals were determined by the t-Student distribution" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0003985_icamechs.2019.8861668-Figure12-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0003985_icamechs.2019.8861668-Figure12-1.png", "caption": "Fig. 12. The developed multi-legged robot prototype", "texts": [ " Two wires were installed on both sides of the trunk and were connected to an active pulley with a servomotor as shown in Fig. 11. This allowed for the robot to turn by pulling the wires using the servomotor. The trunk was moved by wires, and this made it to flexible even when the wire was being pulled. Therefore, the robot could passively adapt to rough ground using its flexibility. & A prototype was developed to evaluate the applicability of the proposed leg mechanism. The prototype is shown in Fig. 12, and its specifications are given in Tables I and II. To demonstrate the effectiveness of the proposed robot, we developed a prototype and conducted experiments. Vertical stiffness of the legs was determined by applying a vertical force as shown in Fig. 13. This resulted in the joint of the leg flexing about 15 degrees at most, and could support 2.5 kgf which is double the stiffness required for it to support the weight of the robot. The horizontal stiffness was determined by flexing the leg in mid-air, and when in contact with the ground as shown in Fig" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0000244_icma.2012.6285106-Figure3-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0000244_icma.2012.6285106-Figure3-1.png", "caption": "Fig. 3 Main sensor in the joint", "texts": [ " To overcome joint flexibility, increase joint sensory capability and reliability, earch joint is equipped with multisensor as shown in Table I. To protect the mechanical structure of robot arm and measure the actually exerted torque to each joint, the torque sensor is used and placed between the output of harmonic drive gear and the link. The deformation of radial beams is measured by strain gauges. By using eight strain gauges and temperature sensor, transverse forces and temperature effects can be compensated. The structure of torque sensor and I-DEAS analysis is shown in Fig 3.a. In the presence of the elasticity and hysteresis of the transmission system, absolute joint angle sensor is needed. To reduce the joint weight and increase sensor\u2019s system integration, conventional contact potentiometer is used as absolute joint angle, which is shown in Fig.3.b. Torque sensor and potentiometer are integrated together, which is shown in Fig.3.c. Motor position information is fed by digital Hall sensor and magnetic encoder. Moreover, limit position protection can be realized by only one digital Hall sensor assembled in the potentiometer. At the same time, to avoid the system overheating, 1-wire digital thermometer is assembled in every board. To realize the proposed controller, the FPGA is used as joint controller, and the DSP chip is used as the top controller. The structure of the hardware control system is shown in Fig. 4. In order to minimize cabling and weight of the 4-DOF flexible joint manipulator, a fully mechatronic design methodology was introduced into developing the hardware system" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0000392_iros.2013.6696761-Figure1-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0000392_iros.2013.6696761-Figure1-1.png", "caption": "Fig. 1. A planar vehicle model", "texts": [ " A dynamic simulation of a similar vehicle model but having a stiff suspensions shows that despite the ultimate speeds, the two wheels maintain contact with the ground at all times. It 978-1-4673-6358-7/13/$31.00 \u00a92013 IEEE 2862 is shown that the all-wheel-drive model produces a larger set of admissible speeds and accelerations, and hence results in faster speeds and shorter motion times than the single drive (front or rear) model. II. VEHICLE MODEL This paper focuses on the planar vehicle shown in Figure 1. It consists of a rigid body of mass m and moment of inertia I that is driven by two actuated mass-less wheels of radius R. The forces acting on each wheel are the normal force Fn in the direction of the normal n to the ground and the traction force Ft in the direction of the tangent vector t at the point of contact. The traction force Ft = T R is in effect the driving force, produced by the wheel torque T . The vehicle\u2019s position is represented by the position (X ,Z) of its center of mass and orientation \u03b8 in the the inertial frame. The back and front contact points are r1 and r2, respectively, relative to the center of mass, as shown in Figure 1. For a given terrain profile f (X) \u2208 R and a given rear contact point X1 \u2208R, we need to compute the contact point, X2 \u2208R, of the front wheel. This in turn will yield the location of the center of mass x and the vehicle orientation \u03b8 . This is computed numerically by modeling the wheel center and the contact points as a closed kinematic chain [11]. By following a specified path, the vehicle has one degreeof-freedom, which can be represented by the arc length, s \u2208 R, of the trajectory followed by the center of mass" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0001752_detc2011-48226-Figure4-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0001752_detc2011-48226-Figure4-1.png", "caption": "FIGURE 4. A COUPLED MOTION OF THE 4-RUU PM.", "texts": [ "org/about-asme/terms-of-use On the other side, in a constraint singular configuration, if the moving platform rotates about an axis of direction mi, the revolute joints attached to the moving platform will no longer be directed along z. As a consequence, the constraint wrench of each limb becomes a wrench of finite pitch (a combination of a force and a moment). In that case, the moving platform has neither pure constraint moments nor pure constraint forces. Moreover, the limbs constrain neither a pure rotation nor a pure translation. Such a configuration is shown in Fig. 4 and corresponds to a coupled motion [28]. In this paper, the actuation singularities correspond to configurations in which Jm is rank deficient while the constraint wrench system does not degenerate. In such configurations, the motion of the moving platform becomes uncontrollable, namely, the actuators cannot control the motion of the moving platform. According to Eqn. (12), these singularities are related to the vanishing conditions of term B. In order to obtain geometric and vector conditions for actuation singularities, term B is expressed in a more compact form by considering the following bracket simplifications: 1" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0000986_j.proeng.2011.03.137-Figure12-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0000986_j.proeng.2011.03.137-Figure12-1.png", "caption": "Figure 12: The open chain associated to the loop.", "texts": [ " Let\u2019s suppose that all the elements are disjoined and become free in space, except one. Only one segmented frame remains fixed, the second becomes mobile (Figure 11). After segmentation and fictional motion in space (or in plane) the number of temporarily mobile elements becomes equal to the number of kinematic joints, i.e. four. In this phase, the number of degrees of freedom of the system is 6 4, for a spatial movement, or 3 4, for a planar movement. We rejoin the elements by rotational joints of V class, including the temporarily segmented frame (Figure 12) and the constraints of the joints are eliminated 5 4 and 2 4 respectively. The extreme element (0) of the open chain has the spatiality three: Rx, Ty, Tz (whether for a spatial or a planar movement); in order to compose the mechanism (Figure 10), three degrees of freedom will be eliminated. If we calculate the mobility of the mechanism by using Eq. (5), for a spatial temporary movement, we obtain the mobility 1 (Eq. (8)) and for a planar temporary movement we obtain the same result (Eq. (9)). M=6m-5p-b1= 134546 (8) M=3m-2p-b1= 134243 (9) The mobility of the mechanism is one" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0000086_amr.423.143-Figure11-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0000086_amr.423.143-Figure11-1.png", "caption": "Figure 11 : 3D real model of fibre", "texts": [ " For the method 1, the value of the section was calculated using the average diameters Di gotten for each fibre orientations (Eq.1). For the method 3, the value of the section was calculated using the average diameters Dix gotten for each fibre orientations (Eq.2). For the polygonal model, the value of the section was determined using an algorithm allowing to calculate the polygonal area. 2 5 cicular D i i 1 S 5 4 = \u03c0 = \u2211 (1) x 2 5 cicular D i i 1 S 5 4 = \u03c0 = \u2211 (2) The 3D model was generated using the CATIA V5 CAD software. The model was extruded from ten sections spaced out of 10mm from each ones (Fig.11). The sections are the result of imaging treatment showing previously (Fig 12). An inverse optimization method coupling to a numerical simulation by the finite elements method using ABAQUS solver with the experimental tensile test was employed to estimate the hemp fibre mechanical properties considering the real section (Fig.16). For the numerical simulation, the mechanical behavior of fibre was considered as elastic linear with one linear part (Method 5a) and two linear parts (Method 5b), until the crack time" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0001482_iros.2011.6095088-Figure9-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0001482_iros.2011.6095088-Figure9-1.png", "caption": "Fig. 9. A lateral offset was applied to the end load, and the resulting deformation was calculated.", "texts": [ " 8 show that the prediction of moment along the backbone from the kinematics of the solution, that is, S T E U , agrees with the prediction of moment based on elastic forces, S $V to within 0.35% of the normalized unitary moment, or a moment of W /$V for a beam with arbitrary length, loading force and stiffness. This means that despite considering only the modal balance of forces on the beam, the local balance of forces everywhere are more or less in equilibrium. B. Three-Dimensional Loading The same sheet, having an aspect ratio . \u2044 2, was then loaded by a force offset from the centerline of the hinge by a distance X, as shown in Fig. 9, so that twisting as well as bending motion resulted. Because the axis of bending is skewed by the three-dimensional motion of the sheet, the moment arm was calculated by finding the normal distance from the line of applied force, the elastic moment was computed including the shape factor, S $V # / . The two predicted moment profiles are plotted in Fig. 10 for a lateral offset X 2 , and the agreement between the two profiles is poorer than the planar case, especially at the two ends of the sheet" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0000185_cecnet.2011.5768595-Figure2-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0000185_cecnet.2011.5768595-Figure2-1.png", "caption": "Figure 2. The calculation graph of theory mining force of bucket hydraulic-cylinder", "texts": [ "It is known to us that the theoretical digging force of hydraulic cylinder counters bucket hydro-cylinder. When some influencing factors are not considered, such as self-weight of working device, bucket load, hydraulic system, efficiency of link mechanism and working back-pressure of hydraulic system, the bucket hydro-cylinder is used alone, then the acting force which caused by the structure parameters of working device and the theoretical digging force of hydraulic cylinder is called theoretical digging force of this hydraulic cylinder. As is shown in figure 2, when we use bucket hydrauliccylinder to excavate the soil, the computational formula of the theoretical digging force of bucket hydraulic-cylinder is: 4068 978-1-61284-459-6/11/$26.00 \u00a92011 IEEE 6 5 222 s s FiF S F =\u22c5= . (1) Where: F2 Theoretical thrust of bucket hydraulic-cylinder, pAF \u22c5= 22 ; 2A Area of sprinklers operation of the large chamber of bucket hydraulic cylinder; s5, s6 Pressure levers; 6s is the crow-fly distance from the cutting basket edge with hinged shaft of the bucket; 2i transmission ratio of bucket hydraulic-cylinder, 6 5 2 s s i = " ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0000047_s13272-013-0095-7-Figure22-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0000047_s13272-013-0095-7-Figure22-1.png", "caption": "Fig. 22 Helicopter coupled with slung load", "texts": [ " When all tests in the ground-based system simulator have successfully passed, the pilot assistance system HALAS will make its first flight in 2013. Appendix In this section, the differential equations of motion for the helicopter with a slung load are derived. The helicopter is modelled as a rigid body with six DOF and the load is modelled as a point mass with three DOF. Helicopter and load motion are coupled by the suspension cable, which is represented by a spring-damper system exerting the cable force as constraining force on helicopter and load (see Fig. 22). The formulation of the equations for the helicopter and the load are set up separately. The general nonlinear equation of motion of the heli- copter for the translational motion is _VH;b \u00bc _u _v _w 2 4 3 5 H;b \u00bc m 1 H X FH;b xH;b VH;b \u00f04\u00de with the sum of the external forces P FH;b, which is composed of the aerodynamic force FA H;b, gravitational force FG H;b and cable force FC b , each of them formulated in the body-fixed frame. X FH;b \u00bc FA H;b \u00fe FG H;b \u00fe FC b \u00bc TbaFA H;a \u00fe TbeFG H;e \u00fe TbeFC e : \u00f05\u00de Tba and Tbe are the transformation matrices from the aerodynamic system and from the earth-fixed system to the body-fixed system, respectively", " The forces acting on the load are as for the helicopter the aerodynamic force FA L;e, gravitational force FG L;e and the cable force FC e : X FL;e \u00bc FA L;e \u00fe FG L;e \u00fe FC e : \u00f09\u00de The aerodynamic force acting on the load is a drag-only force in direction of the local airflow velocity, FA L;e \u00bc q 2 VL;e VL;eSLcD \u00f010\u00de with the density q, surface SL and drag coefficient cD. Calculation of the constraining cable force In the following, the derivation of the constraining cable force which couples the helicopter and the load motion will be described. The cable is represented by a spring-damper system with the resultant cable force FC e (see Fig. 22). For the calculation of the cable force in the earth-fixed frame, the load position relative to the suspension point at the helicopter xC;e is determined: xC;e \u00bc xL;e xHSP;e \u00f011\u00de with the load position xL;e and the suspension point xHSP;e in the earth-fixed frame. The suspension point can be obtained as follows: xHSP;e \u00bc xH;e \u00fe TebrHSP;b: \u00f012\u00de Teb is the transformation matrix with the Euler angles that performs the transformation from the body-fixed to the earth-fixed frame. The actual cable length L is the absolute value of the relative load position: L \u00bc xC;e : \u00f013\u00de The cable force is the sum of the spring and damping force FC e \u00bc \u00f0Fspring \u00fe Fdamp\u00de xC;e L \u00f014\u00de where the spring force is proportional to the elongation, Fspring \u00bc ccabledL \u00f015\u00de with the spring constant ccable and the elongation dL \u00bc L L0 \u00f016\u00de where L0 is the unstretched cable length" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0000865_civemsa.2013.6617399-Figure5-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0000865_civemsa.2013.6617399-Figure5-1.png", "caption": "Figure 5: Force and torque in drilling", "texts": [ " The cutting power P can be estimated by the following equation [14]: P = Kp \u00b7MRR (3) where Kp is the unit power consumption and MRR is the material removal rate. The MRR is calculated simply by [15]: MRR = MR / T (4) MR is the material removed in a period of haptic cycle and T is the period of haptic cycle (typical value is 1 ms). The cutting power P can also be approximated by: P = Ft \u00b7f+Mr \u00b7\u03c9 (5) where Ft is the thrust force, f is the feed rate, Mr is the torque of roll and \u03c9 is the angular rotating velocity of the drill bit as shown in Figure 5. The torque Mr is calculated with: Mr = Fc \u00b7D/2 (6) where Fc is the cutting force and D is the diameter of the drill bit. The cutting force Fc is assumed to be proportional to thrust force Ft: Fc = ktc \u00b7Ft (7) where ktc is a constant related to geometrical parameters of the drill bit and the bone material property. From the above equations we can get the thrust force Ft and torque of roll Mr: p t tc K MR F = (f +0.5K D \u03c9)T \u2217 \u2217 \u2217 (8) tc p r tc K K MR D M = (2f + K D \u03c9)T \u2217 \u2217 \u2217 \u2217 \u2217 (9) Equations (8) and (9) can be further simplified in actual haptic rendering as parameters apart from MR and f can be predetermined before the simulation starts" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0003019_imece2017-72486-Figure1-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0003019_imece2017-72486-Figure1-1.png", "caption": "Figure 1. Basic configuration of four-point contact", "texts": [ " Insights emanating from the proposed analytical model, which could be very useful for design and analysis of four-pointcontact rolling-ball machine elements, are also discussed. Section 3 presents two case studies that demonstrate how the proposed model could shed light on the roles of misalignments, manufacturing errors and loading conditions on rolling/sliding behavior and friction. Finally, conclusions and future work are presented in Section 4. There are many different forms of four-point contact in machine components such as ball bearings, linear guides and ball screws. But four-point contact is similar in all these components. Without loss of generality, Fig. 1 shows the basic configuration where a ball is in four-point point contact with two linear rails. It can be generalized to ball bearing or ball screw applications with circular or helical grooves instead of linear rails. In the setup, the bottom groove is fixed and the top groove is moving at constant velocity of magnitude v. We assume that the contact area is represented as a point, on which normal contact force and frictional force are concentrated. Normal contact forces, as a result of external loading, are calculated a priori based on their own static equilibrium", " Frictional forces based on Coulomb friction law are assumed not to affect the normal contact forces. The basic idea is to establish static equilibrium of force and moment for the friction corresponding to the rolling/sliding behavior, and solve the assumed ball motion. 2.1. Rigid Body Kinematics Friction is a function of ball motion, and it is highly dependent on the rolling/sliding behavior of the ball. Here the kinematics of the four-point-contact ball is first introduced. Figure 2 depicts the cross section of the basic configuration shown in Fig. 1. Let us define a global coordinate system (CS={x, y, z}), fixed in space with its z-axis passing through ball center and its xy-plane parallel to the cross section. A ball, with radius RB, is in four-point contact with the two grooves at BL, BR, TR and TL (representing Bottom/Top and Left/Right) points. The four-point contact happens at the cross section with contact angles \u03b2BL, \u03b2BR, \u03b2TR and \u03b2TL respectively measured from \u00b1y-axis (see Fig. 2), so \u03b2BL, \u03b2BR, \u03b2TR and \u03b2TL\u2208(0, \u03c0/2). Local coordinate systems CSBL, CSBR, CSTR and CSTL are established for the corresponding contact points such that local z-axes are parallel to global z-axis and local y-axes point to the origin of global coordinate system CS as shown in Fig" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0001297_iciea.2011.5975726-Figure1-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0001297_iciea.2011.5975726-Figure1-1.png", "caption": "Fig. 1. Scheme of a 2-DOF robotic system.", "texts": [ " The purpose of this research is, taking the advantages of both the CCC and event-driven control, to develop a position domain control. The main objective of the position domain control is to improve the contour tracking performance using a simple and easy implementing way. In this paper, the masterslave synchronization principle is adopted to develop the new position domain PD control for contour tracking of robotic systems. II. INVERSE KINEMATICS AND RELATIVE DERIVATIVE In this paper, we use a simple 2-DOF serial robotic system as an example for the development of the position domain PD control. Fig.1 shows a scheme of the robotic system where q1 962978-1-4244-8756-1/11/$26.00 c\u00a92011 IEEE A. Inverse kinematic analysis For a contour tracking problem, the position and velocity of the end-effector in a robotic system are defined from the contour requirements. To obtain the position and velocity in the joint level for the purpose of control, an inverse kinematics analysis is needed. According to the configuration shown in Fig. 1, the position of the end-effector can be expressed as: 1 1 2 1 2 1 1 2 1 2 cos cos sin sin l q l q q x l q l q q y (1) Let\u2019s define: 2 2 2 2 1 2 1 22 x y l l D l l (2) From Eq. (1), the corresponding joint positions can be calculated as follows: 2 1 2 1 1 2 2 1 1 2 2 1 tan sin tan / tan cos D q D l q q y x l l q (3) Differentiating Eq. (1) with respect to time, the relationship between the joint velocity and the end-effect velocity can be obtained as follows: 1 1 1 2 1 2 1 2 1 1 1 2 1 2 1 2 sin sin cos cos x l q q l q q q q y l q q l q q q q (4) Solving Eq", " Similar with the definition of the partial derivative, we define the relative derivative of joint 2 with respect to joint 1 as follows: 2 2 2 1 1 dq q q dq q (6) Submitting Eq. (5) into Eq. (6), the relative derivative can be expressed as: 1 2 1 1 2 2 1 1 2 1 2 sin cos sin 1 sin cos sin l q q x q y q l q q q x q q y (7) From Eq. (7), one can see that the relative derivative of joint 2 is a function of the angular position and the velocity of the end-effector. III. DYNAMIC MODEL AND POSITION DOMAIN CONTROLLER A. Dynamic Model in Position Domain A two DOF serial robot system with revolute joints shown in Fig. 1 has the following dynamic equation [24]: 11 12 1 11 12 1 1 21 22 2 21 22 2 2 d d q c c q G T d d q c c q G (8) With 2 2 2 11 1 1 2 1 2 1 2 2 1 2 2 12 21 2 2 1 2 2 2 2 22 2 2 2 ( 2 cos ) ( cos ) d m r m l r l r q I I d d m r l r q I d m r I 11 2 12 1 2 21 1 22, ( ) , , 0c hq c h q q c hq c , 2 1 2 2sinh m l r q 1 1 1 2 1 1 2 2 1 2( ) cos cos( )G m r m l g q m r g q q , 2 2 2 1 2cos( )G m r g q q If link 1 is chosen as the master link, then q1 will be the reference joint angle. Assuming q1 increases monotonically for a defined contour tracking, or 1 0q and 1 0dq , then the dynamic model of link 2 can be rewritten as a function of the reference q1 in the position domain through a one-to-one mathematic transformation from time domain (t ) to position domain ( q1)" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0002743_s11071-017-3840-3-Figure4-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0002743_s11071-017-3840-3-Figure4-1.png", "caption": "Fig. 4 The impacting inverted pendulum [11]", "texts": [ " Similar discussion can be made for the next two mechanical models. 2.2 An inverted pendulum Next, wewill consider the inverted pendulum. It is used in the modeling of various engineering applications, such as rings, printers, machine tools, dynamics of rigid standing structures and rolling railway wheel set [10,11]. The model in [11] will be discussed which has a lateral obstacle for the chattering. The inverted pendulum has impact against the rigid flat wall with a constant restitution coefficient \u03bc. The mechanical model can be observed in Fig. 4. The dynamics of the inverted pendulum between the lateral walls is described by the equations x\u0308 + 2\u03b4 x\u0307 \u2212 x = \u03b3 sin(\u03c9t), |x | < 1, x\u0307 ||x |=1 = \u2212(1 + \u03bc)x\u0307, (3) where x = \u03b8/\u03b8max is the normalized angle (Fig. 4), \u03b4 is the viscous damping (0 < \u03b4 < 1), f (t) = \u03b3 sin(\u03c9t) is the harmonic excitation representing the horizontal acceleration of the base. During the motion of the impacting pendulum, we will take the wall at the position x = 1 as an impacting surface, 0 \u2264 x \u2264 1, \u03b3 = 0.001, \u03c9 = 5, \u03b4 = \u22120.005 and \u03bc = 0.9. Denote x = y, x\u0307 = z. Then, system (3) will be z\u0307 = 0.01z + y + 0.001 sin(5t), z|y=1 = \u2212(1 + \u03bc)z, y\u0307 = z, (4) where 0 \u2264 y \u2264 1. In Fig. 5, one can observe that the pendulum performs many strikes in finite time if the initial values are z(0, \u03bc) = 0, y(0, \u03bc) = 0" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0000976_amr.479-481.2343-Figure3-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0000976_amr.479-481.2343-Figure3-1.png", "caption": "Fig. 3 Section of Workpiece in End-face", "texts": [ ": F / F (e ,n ) 0 0.5*pi pi 1.5*pi 2*pi -1.5 -0.5 0.5 1.5 2.5 Polar Angle:\u03b8 /rad J .F .: F / F (e ,n ) 0 0.5*pi pi 1.5*pi 2*pi -1.5 -0.5 0.5 1.5 2.5 Polar Angle:\u03b8 /rad J .F .: F / F (e ,n ) 0 0.5*pi pi 1.5*pi 2*pi -1.5 -0.5 0.5 1.5 2.5 Polar Angle:\u03b8 /rad J .F .: F / F (e ,n ) e=0.6, n=1 e=0.2, n=2 e=0.1, n=3 e=0.15, n=4 a ) first order elliptic b ) second order elliptic c ) third order elliptic d ) fourth order elliptic Fig. 2 Judgment Function whether Hobbing Process can be Applied As shown in Fig. 3, the )(oxyzS is machine coordinate system, and the )( bbbbb zyxoS is cutting tool coordinate system, and the )(oxyzS is work piece coordinate system. Moreover, we build the polar coordinate system that the pole is co and the polar axis is cx . The pitch curves equation of higher-order elliptic gear is )(\u03b8rr = . A tools rack is attainable by projecting the hob onto the end face. The rotation b\u03c9 of the hob can implement the movement bv of the tools rack. Based on the principle of hobbing with tools rack [6], the motion between the pitch line of the tools rack and the pitch curves of the higher-order elliptic gear is pure rolling, and tangent both in the instant center P. The four linkage shaft is the rotation b\u03c9 of the hob, the rotation c\u03c9 of the work piece, the movement yv of the work piece and the movement zv of the hob. Four-axle three-linkage hobbing model for spur gear. The velocity zv is constant. The angle between the polar radius and the pitch curves of elliptic gear is \u00b5 , as shown in Fig. 3. From the differential geometry theory [5], we can draw the conclusion as follow: ( ) ( )\u03c0\u00b5\u03b8\u00b5 <= \u22640 /tan ddrr (5) From the Eq. 1, the Eq. 3 and the Eq. 5: ( ) ( ) ( ) ( ) ( ) ( ) ( ) ( ) ( ) ( ) \u2212++ \u2212 = \u2212++ \u2212 = \u2212 = \u03b8\u03b8\u03b8 \u03b8 \u00b5 \u03b8\u03b8\u03b8 \u03b8 \u00b5 \u03b8 \u03b8 \u00b5 nenenne nne nenenne ne nne ne cossincos sin cos cossincos cos sin sin cos arctan 21 21 1 1 22222 22222 (6) From the Fig. 3 and the Eq. 6: ( ) ( ) ( ) ( ) ( ) \u00d7 +\u2212+ +\u2212 = += dt d nennene nennenen dt d \u03b8 \u03b8\u03b8\u03b8 \u03b8\u03b8\u00b5 \u03c0\u00b5\u03d5\u03b8 22222 222222 21 2 2 sincoscos sincos -c (7) From the Eq. 7: ( ) ( ) ( ) ( ) ( ) ( ) ( ) ( ) \u03c9 \u03b8\u03b8\u03b8 \u03b8\u03b8\u03b8 \u03c9 \u00d7 +\u2212+ ++++ = nennene nennennen 22222 2222222 21 4121 sincoscos sincoscosc (8) Where \u03c9 is the polar angular velocity of the work piece, and dtd\u03b8\u03c9 = . When the polar angular is\u03b8 , there is an equivalent spur gear, the rotary center of which is co and the pitch radius of which is \u03c1 . The module nm and the pressure angle n\u03b1 are similar to those of the elliptic gear" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0000010_afrcon.2011.6072176-Figure2-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0000010_afrcon.2011.6072176-Figure2-1.png", "caption": "Figure 2: Wind Axis System [11]. Note that the X-Axis always points towards the total velocity vector.", "texts": [ " This research is sponsored by Tellumat (Pty) Ltd \u2013 Defence Division as part of its Unmanned Systems Programme. Website: www.tellumat.com 978-1-61284-993-5/11/$26.00 \u00a92011 IEEE 2) Wind Axis System \u2013 The orientation of the wind axis is defined with the Xw axis always pointing in the direction of the total velocity vector ( ), the Zw axis is perpendicular to the Xw axis and is on the aircraft\u2019s plane of symmetry, and the Yw axis is such that Xw-Yw-Zw forms a right handed system [7]. This is graphically shown in Fig. 2. Also note that the unit vectors defining the wind axis are iW, jW and kW for the Xw, Yw and Zw axes respectively and the origin is the Centre of Gravity of the aircraft.. 3) Euler Angle Attitude System \u2013 Numerous attitude representations are available to represent the attitude of a body in an inertial frame [7]. However, it has been shown that the Euler Angle Attitude (\u03b8, \u03c6, \u03c8) system provides a simple, yet accurate solution [8]. More specifically, the Euler 3-2-1 System will be used. The reader should note that this system does suffer from a singularity at a pitch angle of 90\u00b0, but in spite of this, the system is still the most viable as the UAV pitching to 90\u00b0 is a highly unlikely event" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0000672_kem.450.75-Figure1-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0000672_kem.450.75-Figure1-1.png", "caption": "Fig. 1 Pick arrangement on the cutting head", "texts": [ " The results of practical test and computer simulation show that the load of cutting head has obvious fluctuation, which results in the body vibration of roadheader, accelerates the invalidation of gears and bears, decreases working reliability of the whole roadheader, and affects the feeding of cutting head in the process of swing cutting[1,2]. The load fluctuation and energy consumption are not only decided by the physical properties of coal, but also associated with various parameters of cutting head. All rights reserved. No part of contents of this paper may be reproduced or transmitted in any form or by any means without the written permission of Trans Tech Publications, www.ttp.net. (ID: 141.211.4.224, University of Michigan Library, Media Union Library, Ann Arbor, USA-12/07/15,12:46:41) Choice of Design Variables. As shown in Fig. 1, pick arrangement on the cutting head depends on the cutting spacing of each pick, the circumferential distributing angle, and the number of picks of each transversal on the premise of outline dimensions of cutting head unchanged. Obviously, the number of picks of each transversal is integer variable. The cutting head applying one line-one pick arrangement can cuts more reasonable slotting shape and consume less energy [3]. Thus it is directly chosen one, not chosen as a design variable. If the cutting line spacing of each pick and the circumferential distributing angle are chosen as design variables, the picks of cutting head should have \u00d72 design variables" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0001660_isam.2011.5942364-Figure8-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0001660_isam.2011.5942364-Figure8-1.png", "caption": "Figure 8. a) Clamping of the \u03b2-titanium hollow shaft for longitudinal turning and b) the attribution of x- and z-coordinates to indentify different positions of the tool at the surface of the \u03b2-titanium hollow shaft", "texts": [ " After a cutting length of lc = 3000 m, an average flank wear of VBavg = 93 \u00b5m and a notch wear of VBnotch = 205 \u00b5m is measured. In further experiments on machining the formed hollow shaft, it is recommended to set the cutting speed to vc = 50 m/min to avoid the occurrence of tool wear. This will prevent that the obtained results on process forces and surface roughness are superposed by strong modification of the cutting edge of the tool. 1) Clamping situation For longitudinal turning of the hollow shaft, the clamping jaws within the CNC-lathe are adapted to the inner diameter of the hollow shaft. Fig. 8a presents the clamping situation and the approach of investigating the machinability of the hollow shaft. To prepare the hollow shaft for longitudinal turning tests, the rough surface layer induced by the previous forming operation is firstly removed. The plane workpiece surface is machined in several steps with a constant cutting speed of vc = 50 m/min, a feed rate of f = 0.1 mm and a depth of cut of ap = 0.3 mm. The plane area (Fig. 8b), which is stepwise removed, is segmented into x- and z-coordinates to attribute the recorded process forces to a x- and z-position of the tool at the workpiece surface. 2) Measured process forces In Fig. 9a the measured cutting forces are presented with regard to the position of the tool at the workpiece surface. While the wall thickness of the hollow shaft decreases by the prolonged cutting, maximum cutting forces of fc = 92.3 N are measured in the middle of the wall thickness of the hollow shaft" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0000376_iecon.2013.6699607-Figure1-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0000376_iecon.2013.6699607-Figure1-1.png", "caption": "Fig. 1 shows the principle of the radial force production of 1 2/8 BSRM. The main winding named Nm, which consists of four coils connected in series . The levitation winding named N\" consists of two coils in each direction. When the two different windings conduct the currents, the flux density in air gap 1 increases, whereas it decreases in air gap 2. As a result, an unbalanced magnetic force Fa is produced toward the positive direction in the a-axis. The radial force Fp in the p axis can also be produced in the same way [8] . Thus, the radial force in any desired direction can be obtained through its feedback to achieve the rotor' s rotation and levitation.", "texts": [ " And with the increase of rotor angle, the influence of phase B current on the total radial force is reduced gradually because of the variation of the inductance shown in Fig. 5 . IV. RADIAL FORCE COMPENSATION STRATEGY In order to achieve the rotor' s levitation with the short circuited levitation winding , the lack of radial force needs to be compensated by the previous phase . It can be seen from equation ( 1 8) that compensation can be achieved by conducting the main winding and one suspension winding, or conducting the main winding and two sets of levitation windings . Here the former is selected for the simulation. Fig. 1 1 shows the system control block diagram for the control strategy with short-circuited levitation winding of BSRM. As shown in Fig. 1 1 , the rotor position is detected by the photoelectric sensor and the radial displacement signals in the a and \ufffd direction are converted to voltage signals by the eddy current sensors, respectively. The given main current im * is obtained by the speed loop control . Then the PID controller outputs the desired radial forces Fa * and Fp* . Considering the effect of the radial force generated by the short-circuit current in a direction of the short-circuited phase, the desired radial force Fa * needs to subtract the radial force generated by the short-circuit current and then proceeds to the next step ", " The simulation type is selected as fixed-step with the sampling time l e-7, which is the reason that all simulation waveforms tend to be flat on the top. What' s more, the solver is ode3 (Bogacki-Shampine) in the simulation. Specifically, the given force was 1 50N in the simulation and the relevant mechanical parameters are listed in Table I . The currents of the main winding and short circuited levitation winding of phase A are shown in Fig. 12 , while the normal phase C current i s shown in Fig. 1 3 . By comparing these two figures, it can be seen that the short circuit current is significantly smaller than the normal levitation winding current, resulting in lack of radial force . Fig. 14 shows the simulation results without the compensation. It can be seen that the radial force FAa of short-circuited phase reduced to only 50N averagely, and became uncontrollable. Thus, the levitation of BSRM cannot be achieved. By extending the conduction interval of the previous phase B and considering the short-circuit current in phase A, the radial force can be compensated well. Fig. 1 5 shows the simulation results with compensation, and the currents of the main winding and levitation winding of phase B are shown in Fig. 1 6 . To be more clearly, Fig. 17 shows the total radial force provided by phases A and B, and thus the given radial force can be tracked well during the whole operation region. V. CONCLUSIONS In this paper, a new mathematical model is developed with one levitation winding short-circuited for 1 2/8 BSRMs. Firstly, an equation of short-circuit current is deduced, based on which the radial force model is established. Then the short-circuit current and radial force models are verified by the finite element analysis" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0002289_j.robot.2017.04.002-Figure3-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0002289_j.robot.2017.04.002-Figure3-1.png", "caption": "Fig. 3. Velocity vectors of an omnidirectional wheel (top view).", "texts": [ " First, the robot velocity vectors are calculated at each wheel contact points by the following equation: vA,i = vCoG \u2212 \u03c9 \u00d7 rA,i (1) where vA,i is the linear velocity of the ith wheel contact point, vCoG is the linear velocity of the CoG, \u03c9 is the angular velocity of the robot, and rA,i is the vector pointing fromCoG to the corresponding wheel contact point. The driving directional velocity vector (vdrive,i) can be calculated for each wheel as the perpendicular projection of the contact point\u2019s velocity vector to the driving direction (vdir,i). (See Fig. 3.) The cosine of the angle between two vectors can be calculated by their scalar multiplication. The drive speed of the wheel is as follows: vdrive,i = \u23d0\u23d0vA,i \u23d0\u23d0 vA,i \u00b7 vdir,i\u23d0\u23d0vA,i \u23d0\u23d0 \u00b7 \u23d0\u23d0vdir,i \u23d0\u23d0 . (2) In order to evaluate a feasible motion control, we have to calculate the maximum velocity of the robot that can be achieved without exceeding the maximum velocity limits of the drives. According to Assumption III, the reference velocity of the CoG is given. We propose a velocity reserve multiplier (\u03bb) to express the maximum velocity of the CoG from the reference velocity of the CoG", " Based on Assumption II, we use the force and torque equations of the rigid body in order to construct the dynamic model of the robot. According to Assumption III, the required linear and angular acceleration of the CoG is given. Fig. 6 explains the notations used in dynamic modeling: h is the height of CoG from the ground plane, rA1, rA2, rA3 are vectors in the robot\u2019s coordinate system, those pointing to the wheel\u2019s gripping points from CoG, F 1, F 2, F 3 are wheel forces, FCoG and \u03c4 CoG are the force and torque vectors related to CoG. The axial component (free rolling direction, see vfree on Fig. 3) of the wheel force vector is zero. Therefore, the wheel forces can be split to load and drive components using radial and tangential unit vectors respectively, see Fig. 7. F i = Fi,drive \u00b7 ei,tangential + Fi,load \u00b7 ei,radial (8) where ei,radial and ei,tangential are the unit vectors in the wheel coordinate system. The sum of load forces is equal to the vertical force caused by gravity, therefore the force and torque equations of the robot body: FCoG = [FCoG,x FCoG,y mg ] = F 1 + F 2 + F 3 (9) \u03c4 CoG = [ 0 0 \u03c4CoG,z ] = rA1 \u00d7 F 1 + rA2 \u00d7 F 2 + rA3 \u00d7 F 3 (10) where m is the mass of the rigid body", " In our case (and in most cases), the electrical time constant of the motor is smaller than the mechanical by two orders of magnitude, therefore the armature inductance can be neglected. We can get a close estimation of the maximum output torque of the motor at maximum armature voltage: \u03c4i,max(\u2126i) = ( Va,max \u2212 k\u03c6\u2126i ) k\u03c6 Ra (22) where k\u03c6 is the speed constant of the DCmotor and\u2126 is the angular velocity in rad/s of the armature shaft. Va,max is the maximum armature voltage and Ra is the armature resistance.We can express \u2126 from the drive directional velocity (Fig. 3) of the appropriate wheel: \u2126i = vi,dirve rwheel (23) where rwheel is the radius of the wheel. The maximum torque can be expressed with the maximum force of the wheel in the driving direction: \u03c4i,max = Fi,drive,max \u00b7 rwheel. (24) Substituting (23) and (24) into (22), we can get the maximum force of the wheel depending on the maximum velocity of the wheel: Fi,drive,max(vi,drive) = ( Va,max \u2212 k\u03c6vi,drive rwheel ) k\u03c6 Ra \u00b7 rwheel . (25) Dividing the available drive forces by the required drive forces, a motor torque reserve multiplier \u03c3t could be phased in", " In case of the velocity space, the velocity limiter algorithm was tested based on the kinematic model. For the acceleration space, a Matlab application is shared to visualize the vector spaces in case of different robot and environmental parameters. Fig. 11 shows the linear velocity of the robot (b) and the velocity of each wheel (a) during a linear movement while the robot was rotating continuously. During the discrete time simulation, the maximum wheel velocity in the drive direction (vdrive in Fig. 3) was 300mm/s. The reference linear velocity of the robot in the world coordinate system was 150 mm/s, and the reference angular velocity was 0.5 rad/s. As the diagrams show, when any of the wheel speeds saturates, the overall robot velocity is limited according to the characteristics described in Fig. 4 to avoid path deviation. Notice that the velocity of CoG performs only smooth changes with no discontinuities or higher slopes. This favorable characteristic was utilized later in robotic implementation section" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0001248_epe.2013.6634388-Figure16-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0001248_epe.2013.6634388-Figure16-1.png", "caption": "Fig. 16: Pole pieces of improved gear. Fig. 17: Comparison of efficiency of prototype gear.", "texts": [ " 10, the leakage flux passes through the aluminum, and then the eddy current is induced in it. The eddy current loss of the aluminum of the prototype magnetic gear is estimated by 3D\u2013 FEA. The electrical conductivity of the aluminum is 3.76\u00d7107 S/m. Fig. 15 shows the eddy current density of the aluminum on the r\u2013\u03b8 plane. The figure reveals that the eddy currents are induced around the pole pieces. Therefore, it is necessary to replace the aluminum with non\u2013magnetic and non\u2013 conducting material. Fig. 16 shows the pole pieces of an improved prototype magnetic gear. All the support parts are replaced by Bakelite which is one of the synthetic resins, namely, non\u2013conducting and non\u2013 magnetic material. Fig. 17 shows the comparison of the efficiency of the initial and the improved prototype gears when a load torque is 8.0 N\u2022m. It is clear that the efficiency of the prototype gear employing the Bakelite is improved in comparison with the one using aluminum alloy as a supporter. The maximum efficiency of the improved gear achieves up to about 91 % when the inner rotor rotates at 200 r/min" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0003666_iemdc.2019.8785290-Figure1-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0003666_iemdc.2019.8785290-Figure1-1.png", "caption": "Fig. 1. Rendering of a generic stator end region including the compression system", "texts": [ "ndex Terms\u2014Clamping system, Conformal mapping, End region losses, Hydroelectric power generation I. INTRODUCTION Large generators require a compression system to apply enough clamping force onto the stator core to prevent damage to the stator winding insulation from vibrating teeth [2]. Fig. 1 shows a generic compression system consisting of pressure fingers, a chamfered clamping plate, keybars and clamping bolts applying force to the clamping plate through large spring washers. Due to cost constraints, clamping plates are often made of magnetic ferritic steel instead of approximately 10 times more expensive non-magnetic austenitic steel (see table I), despite causing increased clamping plate losses. Clamping plate loss calculation in hydro and turbo generators is a domain dominated by 3D FEA [3]\u2013[11] due to the complex, three-dimensional nature of the field effects and loss mechanisms at work" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0002500_icra.2017.7989667-Figure10-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0002500_icra.2017.7989667-Figure10-1.png", "caption": "Fig. 10: Snapshots of the arm motion in Task 2-2.", "texts": [ " 9(b), the arm initially follows the red arc, which is then updated to be the yellow arc when another sensing is made. Eventually, the arm is able to maneuver through the concave side by following the green arc. Due to the section length limit, the arm stops after it covers the concave side, resulting in an incomplete coverage of the side surface of the coffee can. To enable the tip camera to sense the coffee can from the other direction, the initial configuration of the arm is reset as displayed in Task 2-2. As shown in Fig. 10, the arm is able to observe the unmodeled part of the coffee can by moving in the counter-clockwise direction. By combining the images captured in Task 2-1 and Task 2-2, the complete model (Fig. 11(b)) of the side surface is obtained, which also includes some portion of the top surface. Table I shows the number of images captured, the short distance \u0394s used, and the total time for planning the robot arm motion in the experiments. \u0394s is set to be both sufficiently small to ensure a sufficient overlap between the images captured in two consecutive sensings and also large enough to be efficient" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0003845_icarm.2019.8834286-Figure1-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0003845_icarm.2019.8834286-Figure1-1.png", "caption": "Fig. 1: Equivalent transformation of the two viscoelastic models", "texts": [ ") may not known ahead in most situations, a few reasonable assumptions are necessary to made here. Assumptions: \u2022 Boundedness: the time integration of external force F(t) is bounded, that is \u222b t 0 \u2016F(\u03c4)\u2016d\u03c4 < \u221e \u2022 Convergence: convergence to zero, F(t)\u2192 0, as t \u2192 \u221e; this means the external forces (contact or impact,etc.) are finally varnished. \u2022 Finite value: integration of external force is a finite value, \u222b \u221e 0 F(\u03c4)d\u03c4 = \u03bb I. Equivalent transformation Using the equivalent transformation shown as Fig.1, original series connection of Maxwell model is changed into parallel way, with coefficient of damper C transformed to KC\u22121M. However, the force integral term KdC\u22121 d \u222b Fdt can be viewed as the additional term of this equivalent transformation. Therefore, we take much effort on this force integral term by our decoupling strategy. II. External force effects decoupling Due to the fact that coefficients in closed-loop dynamic Equation (5) are constant, so the superposition principle is suitable. The nominal external force F and its integral term KdC\u22121 d \u222b Fdt are considered as disturbances" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0000238_00405000.2011.553811-Figure5-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0000238_00405000.2011.553811-Figure5-1.png", "caption": "Figure 5. Forces in z-direction.", "texts": [], "surrounding_texts": [ "In the condensing zone, a flow field is created under the influence of the underpressure. For the sake of easier calculation, uniform pressure on unit length is used to denote the influence of the underpressure; the uniform normal force on unit length is used to denote the supporting force from the lattice; fiber band\u2019s gravity and centripetal force produced by curvilinear motion are ignored as they are relatively faint. Suppose the length of the suction slot is s . When s = 0, the diameter of fiber band is supposed to be constant. Suppose the velocity of the fiber band in the condensing zone is fixed. Arbitrarily select a d s -long part in Fiber Band Part 3, as shown in Figures 1(b) and 2(a). Figure 2(b) illustrates the forces put on d s . The line vertical to the curved surface and going through point O is the z -axis; the line tangential to the curved surface and going through point O is the y -axis; the line vertical to plane yOz and going through point O is the x -axis. Figure 2. Illustration of forces put on a d s -long fiber band in Fiber Band Part 3." ] }, { "image_filename": "designv11_62_0002375_iea.2017.7939175-Figure5-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0002375_iea.2017.7939175-Figure5-1.png", "caption": "Figure 5. Picture of standard nozzle size 0.4mm fitted with the FDM extrusion set", "texts": [ " Before proceeding with experimentation, it was necessary to control the mass of the stainless steel PLA filament used for experimentation. In order to perform this experiment, the mass of the stainless steel PLA was calculated using Archimedes principle, are as follows: SSPPLA = m / V (1) m = SSPPLA \u00d7 V (2) where SSPPLA is the nominal density of the stainless steel PLA is given as 2.38 g/cm3, m is the mass of material (g) and V is the volume of material (cm3). The experiment was conducted by using standard nozzle size 0.4mm, fitted with the FDM Inventor-3D machine, as shown in Fig. 5. The specimen (Fig. 3) is modelled in software and exported as STL file. STL file imported to FDM Inventor-3D machine. Hence, control factor (Table I.) are set as per experiment plan, as shown in Table II. In experiment plan, study of two variable factor at four level requires number of experiments is 16 (24). After the parts were printed the used nozzle was then cut and investigated its cross section by milling machine and microscope, respectively. The used nozzles are cut by milling machine, then their cross-section are measured in order to investigate to variation, as shown in Fig" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0002354_1350650117710813-Figure1-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0002354_1350650117710813-Figure1-1.png", "caption": "Figure 1. Schematic diagram of hydrodynamic rotor-bearing system with dissimilar radial clearances.", "texts": [ " The static characteristics (such as minimum film-thickness and rotor posture) and dynamic characteristics (in terms of stiffness and damping coefficients) were analyzed intensively. The remaining parts of this paper are structured as the following: section \u2018Theoretical analysis\u2019 introduces the theoretical model and mechanical equilibrium analysis. Section \u2018Solution procedure\u2019 presents the detailed solution procedures, which is verified by comparing the results with previous literature. Results of the research are discussed in section \u2018\u2018Results and discussion\u2019. Section \u2018Conclusions\u2019 gives the conclusions. Figure 1(a) shows the geometry of a hydrodynamic rotor-bearing system with dissimilar radial clearances. R represents the radius of the rotor. Width of the two bearings is identical to LB. The two bearings are concentric and located at a distance Lg. The left one, named bearing A, is assumed to be perfectly manufactured with a radius RA b and its radial clearance hA0 remain unchanged during following analysis. The right one, named bearing B, has variational radii RB b . It results in a different radial clearance hB0 compared with that of bearing A. Three cases in which an external load W0 is applied on the left point OL (named Case Left in this paper to describe conveniently), center point OC (Case Center), or right point OR (Case Right) of the rotor will be taken into consideration. OA and OB are the cross-points of the rotor axis and middle sections of bearing A and B, respectively. Figure 1(b) shows the projection of the rotor on X\u2013Z section from the view of left end. Variables XC and ZC denote the displacement of the center point OC of the rotor and X and Z refer to the misalignment angles along X and Z direction. The system is assumed to be lubricated by incompressible and Newtonian fluid. Inertia and thermal effects are neglected. The governing equations can be expressed as equation (1). @ @x h3 12 @p @x \u00fe @ @y h3 12 @p @y \u00bc U 2 @h @x \u00fe @h @t \u00f01\u00de where x and y denote the circumferential and axial co-ordinates of oil film, h is the height of film thickness, represents the oil viscosity, p refers to the pressure, U is the velocity of the rotor, and t represents the time", " The pressure distributions in the oil film of each journal bearing can be obtained by solving equation (2) under appropriate boundary conditions. Then the static and dynamic performance characteristics can be calculated using the expressions given in Jang and Kim24 and Das et al.25 Load capacity. Using equations (6) to (10), load capacity of bearing A and B in X and Z direction can be obtained as25,26 Wk X \u00bc Z Z Sk p cos xd xd y \u00f0k \u00bc A or B\u00de \u00f06\u00de Wk Z \u00bc Z Z Sk p sin xd xd y \u00f0k \u00bc A or B\u00de \u00f07\u00de In the schematic diagram (Figure 1), x represents the circumferential direction, and the domain of x is [0, 2 R]. The y represents the axial direction. Since the origin of coordinates is set at center of the rotor (point O), for bearing A, the domain of y is [Lg/2, Lg/2\u00feLB]. And for bearing B, the domain of y is [ Lg/2 LB, Lg/2]. From Table 1, parameters of geometry are set as LB\u00bc 2R, Lg\u00bc 4R. Thus, for bearing A and B, the non-dimensional limits of integration for determining the load capacity are as follow: The resultant load capacity contains the combined interaction of bearing A and B: WX \u00bc WA X \u00fe WB X \u00f09\u00de Stiffness and damping coefficients" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0002302_b978-0-12-803581-8.10351-0-Figure42-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0002302_b978-0-12-803581-8.10351-0-Figure42-1.png", "caption": "Fig. 42 Right clip interface bracket.", "texts": [], "surrounding_texts": [ "The mono shock wedge is designed to react suspension loads from the spring/damper unit and support the suspension box rear angle wall. The design utilizes the same U-shape cross-section as the intermediate bulkhead as seen in Fig. 29. The wedge will be bonded to the chassis using the two angular bonding flanges and to the main bulkhead with the two vertical bonding flanges shown in Fig. 37(a). The wedge design and bonding flanges allow for transfer of the load from the spring/damper unit to the chassis and then into the main bulkhead and the main structure of the chassis as shown in Fig. 37(b). The layup of the wedge follows that of the intermediate bulkhead, [745, 0, 0\u201390, 90, 745, 0, 0\u201390, 90, 745, 0, 0\u201390, 90], from the outer surface to the inner surface. Ungrouped plies indicate the use of unidirectional prepreg, while grouped plies indicate the woven material." ] }, { "image_filename": "designv11_62_0002343_978-3-319-59480-4_17-Figure2-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0002343_978-3-319-59480-4_17-Figure2-1.png", "caption": "Fig. 2. The humanoid robot NAO and some of the main sensors involved in our soft somatosensory system.", "texts": [ " Indeed, this description based on a single direction of the pain, from nociceptors to cortex, hides an important function of modulation of the pain performed by the PAG (Peri-Aqueductal Gray) along the downward path of the pain. Endogenous substances like endorphins or exogenous substances like opiates can operate this modulation function. Thus, the final perception of the pain is always an algebraic sum of these two components. Following the model of the pain sketched above an artificial pain bio-inspired model is designed for a robot. Starting from signals coming from its hardware sensors as described in Fig. 2, a set of soft sensors [2] increase the semantic level of information. They are mathematical tools able to calculate or estimate quantities that are impossible to measure using real sensors. The soft sensors were employed in several real world application by authors and they alway used as stochastic estimators replacing the real sensor [3\u201310]. The soft sensors used for the pain estimation will be described in depth in the next section. They exploit a particular activation function, also bio-inspired, to reproduce the somatosensory cortex elaboration function", " And finally the function of activation based on the exponential functions. Functions of this nature get on and off the level similarly to what happens it the humans. The perception of the pain is just one of the characteristics, though one of the most important, of a somatosensory system that is our concluding goal. Thus, we are working to extend to other stimuli similar consideration faced in this work. More in details, we are considering to involve in the artificial somatosensory system measure regarding accelerations, energy, touch and more in according to Fig. 2, and for each of this measure design a particular soft sensor able to transduce basic values in high-level perception. 1. Dubin, A.E., Patapoutian, A.: Nociceptors: the sensors of the pain pathway. J. Clin. Invest. 120(11), 3760\u20133772 (2010). doi:10.1172/JCI42843 2. Fortuna, L., Graziani, S., Rizzo, A., Xibilia, M.G.: Soft Sensors for Monitoring and Control of Industrial Processes. Springer Science and Business Media, Heidelberg (2007) 3. Cipolla, E., Maniscalco, U., Rizzo, R., Stabile, D., Vella, F" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0002085_6.2017-0500-Figure6-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0002085_6.2017-0500-Figure6-1.png", "caption": "Figure 6. Pitching moment and thrust degradation of rotor in gust flow.", "texts": [ " In addition, crosswinds generate head-up pitching moments around the UAV\u2019s center of gravity. The rotor\u2019s thrust depends on the rotor blade\u2019s shape, collective pitch, and speed of the passing airflow. Crosswinds accelerate the passing flow speed and degrade thrust. This passing flow is called induced velocity and is expressed by advance ratio (relative flow velocity to rotor rotation speed) in rotor aerodynamics. The thrust is reduced as the induced flow accelerates because the local effective collective pitch angle of the blades decreases, as shown in Fig. 6. Therefore, in a crosswind, rotors lose their thrust. Thrust loss is most significant when the rotor plane is tilted in the direction of the main airstream. Moreover, pitch is generated in a crosswind. The rotating blades catch the flow in different directions, and the local flow speed to the blades changes for each rotation angle. When a blade moves forward, it experiences greater Hover Yawing Movement Cross Wind Horizontal Movement D ow nl oa de d by U N IV E R SI T Y O F C O L O R A D O o n Ja nu ar y 12 , 2 01 7 | h ttp :// ar c" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0001368_eeeic.2011.5874826-Figure3-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0001368_eeeic.2011.5874826-Figure3-1.png", "caption": "Fig. 3. Magnetic field lines", "texts": [ " Due to the special construction of the models, the level of static eccentricity can be modified with ease, allowing the analysis of results for different eccentricity levels. For a 15% eccentricity level, the frequency spectrum of the noise and the distribution of forces that cause stator vibrations were studied. The obtained results were corroborated with noise-to-frequency measurements performed on the studied motor. Fig. 2 shows the Finite Element mesh of the modeled motor created and refined using FLUX 2D. The mesh has 37937 nodes and 18932 surface elements with 95% excellent quality mesh elements. In Fig. 3 and Fig. 4 the magnetic field lines and the radial component of the flux density of the modeled machine in case of 15% eccentricity level are shown: The magnetic forces which act on the stator surface and the harmonic spectrum were simulated for two cases: a case in which the motor is functioning under ideal conditions, with no eccentricity, and a case for 15% eccentricity level. The simulation results are presented in Fig. 5 and Fig. 6. From Fig. 5 it can be seen that the maximum value of the force for the motor without eccentricity is around 60 N" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0003781_j.automatica.2019.108553-Figure1-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0003781_j.automatica.2019.108553-Figure1-1.png", "caption": "Fig. 1. Coordinate frames.", "texts": [ " The spacecraft attitude equations of motion are then given by q\u0307 = 1 2 B(q)\u03c9 = 1 2 [ PT (q) \u2212qv ]T \u03c9, (1) J\u03c9\u0307 = \u2212\u03c9\u00d7J\u03c9 + \u03c4u + \u03c4d, (2) where P(q) = q4I + q\u00d7 v , J \u2208 R3\u00d73 is the spacecraft inertia matrix, \u03c4u is the control torque and \u03c4d is the disturbance torque. Define the disturbance torque as \u03c4d = \u03c4g + \u03c4\u0304d, where \u03c4g = 3 \u00b5 \u2225R\u22255 (C(q)R)\u00d7JC(q)R is the gravity-gradient torque, all other external disturbances are \u03c4\u0304d and R \u2208 R3 is the inertial orbital position coordinates, and C(q) = (q24 \u2212 qTvqv)I + 2qvqTv \u2212 2q4q\u00d7 v , is the rotation matrix corresponding to quaternion, q. The system is outlined in Fig. 1. A nominal spacecraft inertia matrix of the form Jn = diag{Jx, Jy, Jz} is assumed to be known, such that the spacecraft inertia is J = Jn +\u2206J , where \u2206J \u2208 R3\u00d73 is the uncertainty in the inertia matrix. The spacecraft is only actuated about the roll and pitch axes, making the system an underactuated case, and \u03c4u is given by \u03c4u = JnHu, H = [ I2 02\u00d71 ]T , (3) where u = [ux, uy] T is the designed control input. Note that by re-ordering of the body-frame axes, the cases when the roll/yaw or pitch/yaw axes are the only actuated axes can be transformed to the roll/pitch axis case" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0001609_0954406211427833-Figure4-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0001609_0954406211427833-Figure4-1.png", "caption": "Figure 4. Kinematic relationship between cutter and gear blank during simulation.", "texts": [ " The working of the spherical gears can be visualized with reference to Figure 1. During meshing of gear pair, the gears can rotate within 24 to 24 about the Y- and Z-axes independently, which are working axes in the assembled condition. The combination of rotation about these axes is also possible, as shown in the assembly. These working axes should not be confused with the notations for axes in Figure 3 used for CAD simulation. The simulation of the generation machining process is done according to the scheme presented in Figure 4. The different simulation parameters are presented in Table 2. The home position of the cutter is taken at 149.65mm. When the cutter touches the blank, its current position is 219.99mm. A total depth of 10.66mm (ha\u00fe hf\u00bc 5\u00fe 5.66mm from Table 1) is realized, when the cutter plunges to 230.65mm. In order to decide the oscillation angle from the centre position, it is determined from the following equation based on meshing condition.6 3 \u00bc \u00f0 m=4\u00de sin o \u00fe \u00f0ha \u00fe hf \u00de= cos o ro sin o \u00f03\u00de Oscillation of 20" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0000010_afrcon.2011.6072176-Figure1-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0000010_afrcon.2011.6072176-Figure1-1.png", "caption": "Figure 1: NED Axis is the reference frame utilized in the dynamic model of the UAV [9]", "texts": [ " In the proceeding sub-sections, the applicable axis and attitude systems are defined as well as the mathematics governing the UAV dynamics. The forthcoming axis definitions will form the basis on which the mathematical models are developed. Thus, an adequate understanding of these is vital for the reader. 1) Inertial Axis System - This axis system provides a fixed reference frame whereby the aircraft motion can be described. This right hand orthogonal axis is fixed to a flat and non-rotating Earth, with its origin at some arbitrary reference point. The North, East and Down (NED) directions are shown graphically in Fig. 1. This research is sponsored by Tellumat (Pty) Ltd \u2013 Defence Division as part of its Unmanned Systems Programme. Website: www.tellumat.com 978-1-61284-993-5/11/$26.00 \u00a92011 IEEE 2) Wind Axis System \u2013 The orientation of the wind axis is defined with the Xw axis always pointing in the direction of the total velocity vector ( ), the Zw axis is perpendicular to the Xw axis and is on the aircraft\u2019s plane of symmetry, and the Yw axis is such that Xw-Yw-Zw forms a right handed system [7]. This is graphically shown in Fig" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0003715_physrevfluids.4.083102-Figure1-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0003715_physrevfluids.4.083102-Figure1-1.png", "caption": "FIG. 1. The experimental setup. (a) The helical swimmer. The dimensions of the device are: Dh = 3 mm; Lh = 16 mm; Lt = 16 mm; \u03bb = 5.3 mm; 2R = 3 mm; and the pitch angle \u03b8 = 45\u25e6. The thickness of the wire, d , was 0.9 mm. (b) The upper and (c) the lower fluids. The swimmer with (d) head and (e) tail firsts.", "texts": [ " The density of the fluid was measured with a 25-ml pichnometer. The fluid viscosity was measured with a viscometer (DV-III, Brokefield). The interfacial tension between the two fluids was not measured. A value of \u03c3I = 0.05 N/m was inferred using the well-known Antonoff\u2019s formula, considering tabulated values of the surface tension for water and silicone oil (\u03c3W = 0.07 N/m [44] and \u03c3S = 0.02 N/m, from the technical data from the manufacturer, Gelest Inc). 083102-3 The swimmer, depicted schematically in Fig. 1(a), was placed inside the container, initially in the lower fluid fixed by a release mechanism. The rotating magnetic coil and swimmer were arranged to make the robot swim in the vertical direction, moving perpendicularly to the interface between the two fluids, see Figs. 1(d) and 1(e). The average density of the swimmer was 1050 kg/m3, slightly above the light fluid but slightly below the heavy fluid. Note that the density of the swimmer is not uniform, since the tail is denser than the head. Although their weights were similar (the head and tail weights were 0.172 and 0.175 g, respectively). The dimensions of the swimmer are reported in the caption of Fig. 1. The motion of the swimmer and the penetration process were recorded with a high-resolution video camera (920 \u00d7 1080 pixels, Sony RX10II) at 60 frames per second. The position of the swimmer in time was obtained using the software Tracker. The swimming speed, U , was calculated using a central difference scheme. In all cases, the rotation speed of the swimmer was the same as that of the external magnetic field: in other words, the swimmer was operated below the step-out frequency of the arrangement [43]", " We can, therefore, write Fp = Fd + Fb + F\u03c3 , (3) where Fp is the propulsion force due to the rotation of the helical tail, Fd is the drag on the tail and head of the swimmer, Fb is the net buoyancy force, and F\u03c3 is the force exerted on the swimmer by the deformed interface. From the classical resistive force theory [20], we know that the propulsive force generated by the tail is Fp \u2248 (\u03b5\u22a5 \u2212 \u03b5\u2016)\u03bc RLt sin \u03b8 (4) and that the drag force on the head and helix is Fd \u2248 (\u03b5\u22a5 sin2 \u03b8 + \u03b5\u2016 cos2 \u03b8 ) cos \u03b8 \u03bcLtU + \u03b50\u03bcLhU, (5) where \u03b5\u22a5 and \u03b5\u2016 are the normal and tangential drag coefficients, is the rotational speed, R, Lt and \u03b8 are the dimensions of the helix (see Fig. 1), and \u03bc is the fluid viscosity. The head is characterized by its length Lh and drag coefficient, \u03b5h. Neglecting the buoyancy force, and far from the interface, the swimming speed is Uo = R(\u03be \u2212 1) f sin \u03b8, (6) where f is defined as f = cos \u03b8 1 + (\u03be \u2212 1) sin2 \u03b8 + \u03be0L\u2217 h cos \u03b8 , (7) where L\u2217 h = Lh/Lt , \u03be0 = \u03b5h/\u03b5\u2016, and \u03be = \u03b5\u22a5/\u03b5\u2016. The buoyancy force can be calculated as Fb = \u03c1V g, (8) where \u03c1 = \u03c1swimmer \u2212 \u03c1 is the density difference between the swimmer and the fluid. V is the volume of the swimmer and g is the gravitational acceleration", " (15) Equations (14) and (15) can be used calculate the speed of the swimmer at different instants during the penetration process since most parameters are known (U1/Uo, L\u2217, \u03be , \u03beo, \u03b8, \u03b5\u2016) for a given value of Ca. Note that for small \u03b8 , f is reduced to f = 2 1 + \u03beoL\u2217 h The values of the drag coefficients, \u03b5\u2016, \u03b5h, and \u03b5\u22a5 can be obtained from Ref. [50], resulting in: \u03b5h = 1.632\u03c0 , \u03b5\u2016 = 0.753\u03c0 and \u03b5\u22a5 = 1.093\u03c0 for which \u03be = 1.45 and \u03be0 = 2.167. From the geometry of the swimmer used in this study L\u2217 = 1, as shown in Fig. 1. The only unknown parameter is L\u2217 \u03c3 , which is the ratio of the radius of curvature of the deformed interface to the length of the tail. From the images in Figs. 4 and 6 we can argue that the curvature of the interface is proportional to the diameter of the head; therefore, we can define a normalized meniscus curvature as \u03ba = L\u03c3 Dh . (16) From this definition, we can write L\u2217 \u03c3 = \u03baDh/Lt . For the swimmer used in the study, Lt/Dh = 5.33. Hence, \u03ba is the only unknown parameter, which we can be easily fit to the data" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0001687_978-90-481-9689-0_68-Figure6-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0001687_978-90-481-9689-0_68-Figure6-1.png", "caption": "Fig. 6 Geometry 2.B of the 3-UPU TPM.", "texts": [], "surrounding_texts": [ "In this section, three manufacturing solutions are presented in order to avoid the leg collision in the Geometries of type B (crossed legs) of the 3-UPU TPM. Geometry 1.B is taken (for clarity) as an example of this type of 3-UPU TPM. The first manufacturing solution S1, is to rebuilt the platform of the manipulator. This is obtained by disconnecting the platform of this geometry from the legs and rotating it by a suitable angle \u03b1 about the z axis of Sb, then connecting again the legs to the platform still keeping the same base axis directions. This means to manufacture a platform with the revolute axis directions rotated of \u03b1 (clockwise in the example shown in Figure 8a) with respect to the Geometry 1.B. This makes it possible to avoid the leg collision. In Figure 8a, the universal joints on the base and on the platform are represented by points for clarity, and the prismatic ones are omitted. After manufacturing the new platform, the coordinates of the center of the universal joint on the platform Ai, i = 1,2,3, are given by: 601 A.H. Chebbi and V. Parenti-Castelli OpAi = cos\u03b1OpA\u2032 i + sin\u03b1OpA\u2032\u2032 i , with OpA\u2032\u2032\u22a5OpA\u2032; and \u2225\u2225OpA\u2032\u2032 \u2225\u2225 = \u2225\u2225OpA\u2032 \u2225\u2225 (2) where A\u2032 i, i = 1,2,3, are the centers of the universal joints on the platform of the Geometry 1.B. The second manufacturing solution S2, schematically shown in Figure 8b, is to rebuild both the base and the platform of the Geometry 1.B in order to have the coordinates of the centers of universal joints at the base and at the platform, respectively Bi and Ai, i = 1,2,3, see Figure 8b, given as follows: ObBi = ObB\u2032 i + eq1i, and OpAi = OpA\u2032 i + eq4i (3) where B\u2032 i and A\u2032 i, i = 1,2,3, are respectively the center of the universal joints in the base and in the platform of the original Geometry 1.B; q1i and q4i, i = 1,2,3, are respectively the unit vectors of the revolute joints on the base and on the platform, which maintain the same directions of the original Geometry 1.B; e is a given distance between the corresponding center of universal joints in the platform of the Geometry 1.B and the platform rebuilt. The third manufacturing solution S3, schematically shown in Figure 8(c), is to rebuilt the second and the third link of each leg of the Geometry 1.B in order to change the physical position of the prismatic pairs on each leg along EiFi, where the coordinate of points Ei and Fi, i = 1,2,3 are given by: ObEi = ObBi + dq2i, and OpFi = OpAi + dq3i (4) where Bi and Ai, i = 1,2,3, are respectively the centers of the universal joints in the base and in the platform of the Geometry 1.B; q2i and q3i, i = 1,2,3, are respectively the unit vectors of the intermediate revolute joints of the i-th leg; d is a given distance between the directions of the prismatic pairs for the Geometry 1.B and the manipulator geometry after rebuilding. 602 Geometric and Manufacturing Issues of the 3-UPU Pure Translational Manipulator" ] }, { "image_filename": "designv11_62_0001607_0021998312463455-Figure5-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0001607_0021998312463455-Figure5-1.png", "caption": "Figure 5. Yarn compression schematic for braided sleeve element.", "texts": [ " Thus, equilibrium equations at point M can be written as FT cos \u00fe FA \u00bc FT FN \u00bc FT sin \u00fe P Aint \u00f01\u00de where P is the internal pressure of the braided composite actuator and Aint is the overlapping area by two intersected yarns. The out-of-plane component FN is the compressive force for fiber compaction. The compressive forces are simplified to point forces and we assume that the compressive forces are distributed on a small parallelogram area, as shown in Figure 4. As the fibers rotate, the area of the parallelogram changes, which is included in this model. In the figure, d is the width of the fiber yarn and is the braiding angle by the fiber and the generator of the cylinder. As shown in Figure 5, the solid lines and the dashed lines are the yarns before and after compaction, respectively. The thickness of the sleeve is reduced by . If the fiber yarns are inextensible, the length of the element edge is increased. Because the element dimension is much smaller than the actuator\u2019s radius, the element can be treated as flat. Therefore the length of the element edge is increased by approximately 2b(cos 1 cos 0), where 0 and 1 are the yarn bending angles before and after compaction, respectively" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0003346_irc.2019.00068-Figure1-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0003346_irc.2019.00068-Figure1-1.png", "caption": "Fig. 1. The Baxter\u2019s arm is divided into two parts: arm and wrist.", "texts": [ " The ability of neural network (NN) approaches to represent non-linear relationships that map the input to the output data based on provided training examples make them one of the common approaches used to find a solution for the IK problem [6]. In order to simplify the solution of the IK problem for a high DOF robotic manipulator, we propose to divide the robotic manipulator into two parts. The first part (arm) includes the major joints of the robotic manipulator (for Baxter: shoulder roll, shoulder pitch, elbow roll, elbow pitch), while the second part (wrist) includes the minor joints of the robot structure (for Baxter: wrist roll, wrist pitch, wrist roll), as shown in Fig. 1. The intersection point J between the two parts is computed first, and then each one of the two parts is treated separately. The end-point of the first part (arm) should reach point J with any orientation. A NN is trained to find the appropriate arm joint variables to accomplish this. The required joint variables for the second part (wrist) are calculated with the aid of a coordinate transformation and vector analysis [7] to make the robot end-effector reach the desired point with the desired orientation and zero or small error" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0003385_j.apm.2019.04.020-Figure1-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0003385_j.apm.2019.04.020-Figure1-1.png", "caption": "Fig. 1. Soft robotic finger taps rough target surface.", "texts": [ " The absolute nodal coordinate formulation (ANCF) is an optional method, which can describe the large dynamic deformation well. Hu et al. [26] successfully use it to discretize the soft structure when studying the contact walking dynamics of soft quadrupedal robot. Their results show ANCF might have great potential to be applied to analyze the contact-impact event of soft finger. In this paper, considering both tangential and normal contact compliance, a hybrid computational model (HCM) is developed to analyze the frictional impact for a humanoid soft finger tapping against a rough target surface (see Fig. 1 ). Based on the proposed HCM, the structural stress, large deflection and contact forces during the contact-impact phase are calculated. And the normal and tangential relative motions (i.e. compression-restitution and stick\u2013slip) between the contact points are depicted. In Section 2 , the HCM is developed. The humanoid soft finger consisting of multiple phalanges and joints is modeled as a flexible multi-link in the HCM. The model is not hard to be degraded for solving the most developed soft fingers consisting of single phalanx with joint" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0003186_jifs-169871-Figure6-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0003186_jifs-169871-Figure6-1.png", "caption": "Fig. 6. Test rig of planetary gearbox transmission.", "texts": [ " 338 For the appearance of the non-integer period T , it can 339 be rounded to the nearest position. 340 7. Experimental signal analysis 341 7.1. Instruction on test rig 342 To verify the validity of the MOMEDA on signal 343 from planetary gearboxes. Experiments on a plane- 344 U nc or re ct ed A ut ho r P ro of Fig. 5. Transfer paths of planet gear fault induced vibration signal (a) Transfer path 1 (b) Transfer path 2 (c) Transfer path 3. tary gearboxes transmission system is carried. The345 test object is a NGW planetary gearbox mounted on346 the test rig, which is shown in Fig. 6. Parameters of the347 NGW planetary gearbox are shown in Table 1. The348 vibrations under the normal and abnormal conditions349 are picked up respectively. In the abnormal condition,350 a tooth root crack about 4 mm is cut artificially on a351 planet gear tooth, which is shown in Figs. 7 and 8.352 Three acceleration sensors (DH112 with a sensitivity353 of 5.20 pC/g) are mounted on the planetary gearbox354 for observing the vibration, which are shown in Fig. 6355 labeled with 3, 4, and 5. The sampling frequency used356 in the experiment is 51" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0000720_amm.307.304-Figure1-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0000720_amm.307.304-Figure1-1.png", "caption": "Figure 1 Virtual spur gear equal to bevel gear", "texts": [ " Nalluveettil and Muthurearappan introduced a procedure for the finite element analysis of load and stress distribution of straight bevel gears [2]. Elkholy and Elsharkawy studied of meshing tooth stiffness of bevel gears and calculated Hertzian stress in bevel gear [3]. Hence in this paper a three dimensional distribution for Hertzian stress was obtained and compared with Hertzian stress obtained from finite element analysis. According to Tredgold approximation, contact of bevel gear can be replaced with a pair of spur gear witch their center line will be laid on the axes of bevel gears as shown in Fig. 1. All rights reserved. No part of contents of this paper may be reproduced or transmitted in any form or by any means without the written permission of Trans Tech Publications, www.ttp.net. (ID: 132.174.254.159, Pennsylvania State University, University Park, USA-28/05/15,11:41:40) The number of teeth of virtual spur gears and the equal diameter of them can be found from the relations [4]: \u03b3cos bevel spur z z = (1) \u03b3cos bevel spur d d = (2) where: Zspur Number of teeth in virtual spur gear Zbevel Number of teeth in bevel gear dbevel Pitch diameter of bevel gear dspur Pitch diameter of virtual spur gear \u03b3 Pitch cone angle In this study, for higher accuracy face width of each tooth of bevel gear replaced with multiple pairs of spur gears such as shown in Fig" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0000224_kem.504-506.1305-Figure3-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0000224_kem.504-506.1305-Figure3-1.png", "caption": "Figure 3: VBB measurement procedure according with ISO 3685:1993 and adopted notation for measured forces.", "texts": [ " The design variables are DOC and En. Cutting speed, feed rate and removed volume were kept constant for each run. After machining, tool wear analysis was carried out on the tools evaluating the influence of the cutting conditions. More in detail, in according with the international standard normative ISO 3685:1993 \u201ctool-life testing with single-point turning tools, for the tool wear analysis\u201d, the average width of the flank wear land in the regular worn zone B (VBB) was measured using an optical microscope Leica - DM4000M. Fig. 3 shows: the procedure adopted for VBB measurement, the reference system and notation used to present the data about cutting forces. The experimental results obtained during the machining operation are reported in Table 2. Cutting forces. A typical forces trend obtained in the experimental test is reported in Fig. 4. The obtained results show a different trend for the cutting forces depending by the considered engage angle conditions. In fact, for engage angle of 0\u00b0, 45 to 90\u00b0 it is evident, at the beginning when the tool engages with the workpiece, a high compressive value before to reach the steady state condition (Fig", " A final remark deductible from this forces analysis is that in machining of Inconel 718 low values of DOC and En should be used in order to obtain low cutting and thrust force components; the same observation can be done for the initial value of cutting force component at the beginning of the cutting operation (Fx,eng): low values of DOC and En lead to reduce initial cutting force components. Tool wear. The experimental results reported in Fig. 6 show some of the tool wear comparison analysis conducted with the aim to better understand the influence of the machining process conditions on the selected tool. In accordance with the international standard normative ISO 3685:1993 the average width of the flank wear land in the regular worn zone B, VBB (Fig. 3), were measured and compared at the different considered levels of En and DOC by optical microscope analysis. It is important to highlight that all the experiments were carried out in conditions of constant removal volume for each run (Vrim=27403mm 3 ). As concern the influence of En parameter, it can be noted in Fig. 6 as highest value for VB is registered for En=45\u00b0. The reason of this behavior is due to the thrust and the cutting force components since for En =45\u00b0 the thrust force component and the cutting force component have comparable values, leading to maximize tool wear in terms of VB" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0001491_wosspa.2011.5931450-Figure1-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0001491_wosspa.2011.5931450-Figure1-1.png", "caption": "Figure 1. The athlete\u2019s attitude during jumping", "texts": [ " The reported data were obtained from an athlete who is familiar with performing a correct jump. The best jumping recording by the force platform are presented. During vertical jumping, the muscles about the hips, knees, and ankles act rapidly and with great force in an attempt to produce the greatest possible velocity for the body as it leaves the ground [10]. The jump height is ultimately determined by the takeoff velocity [14]. A schematic representation of a typical jumping on a platform is shown in Fig.1. In this figure the platform is reduced to a simple plate moving on a spring and damper, which is equivalent to the force sensor realized. Initially, the athlete is immobile on the platform, just at the beginning of his gesture, before he takeoff, the system records the dynamic response of the platform. During this time, the athlete executes a vertical jump and falls again 978-1-4577-0690-5/11/$26.00 \u00a92011 IEEE 17 on the platform and will take his initial position again. The recorded signal is relative to a counter movement jump protocol [15], [16], [17] (CMJ protocol) that exhibits the evolution of the vertical force during jumping" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0000339_cdc.2012.6426020-Figure1-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0000339_cdc.2012.6426020-Figure1-1.png", "caption": "Fig. 1. Trajectories of the three agents in the plane.", "texts": [ " Finally, notice that since the value of the xi(k+1) is unknown at time k, then the feedback should concern only the velocities and the position at time k. We solve an optimization problem with feedback gains Ki, j = \u03b1i, jK where K = [ 0 1 1 0 0 0 0 0 0 0 1 1 ] for (i, j) = (1,2) and (i, j) = (2,3). The optimization objective has been to compute the feedback gains satisfying the condition presented in Theorem 1 minimizing the euclidean measure of the error. A set of trajectories in the plane are shown in Figure 1, obtained for initial conditions x1(0) = 1.8, vx 1(0) = 0.9, x2(0) = 0.9, y2(0) =\u22120.9, vx 3(0) =\u22120.36, y3(0) = 0.18 and 0 for all the other states. Notice that, as the control objective is the minimization of the norm of the errors, the systems states mismatches converge to zero. Moreover, the algebraic constraint is satisfied along the whole trajectories, also at the first instants, which are the most critical ones as the system initial conditions are close to the boundary of the feasible region, see Figure 2" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0000960_cdc.2013.6760861-Figure6-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0000960_cdc.2013.6760861-Figure6-1.png", "caption": "Fig. 6: View of the AC-ROV with the reference frames (xiyizi: earth-fixed frame, xbybzb: body-fixed frame).", "texts": [ "2 PM 100 deg PM 92 deg PM 84 deg GM 7.9 dB GM 7.9 dB GM 7.9 dB TABLE IV: Values of the stability margins when the derivative gain is changed. Parameter Value Parameter Value Parameter Value KP 3 KP 3 KP 3 KD 0.1 KD 0.2 KD 0.3 PM 94 deg PM 100 deg PM 106 deg GM 6.9 dB GM 7.9 dB GM 9.1 dB TABLE V: Values of the stability margins when the adaptation gain is changed. Adaptation Gain Phase Margin Gain Margin 5000 8.9 dB 101 deg 10000 8.4 dB 101 deg 100000 8 dB 101 deg UNDERWATER VEHICLE The AC-ROV submarine (cf. Fig. 6) is an underactuated underwater vehicle. The propulsion system consists of six thrusters driven by DC motors controlling five degrees of freedom. Four horizontal thrusters control simultaneously translations along x and y axes and rotation around the z axis (yaw angle). The two horizontal thrusters denoted \u2019Thruster 1\u2019 and \u2019Thruster 2\u2019 on Fig. 6 control depth position and pitch angle. The roll angle is unactuated but remains naturally stable due to the relative position of buoyancy and gravity centers. The robot weighs 3 kg and has a rectangular shape with height 203 mm, length 152 mm and width 146 mm. It has been modified by the Laboratory of Informatics, Robotics and Microelectronics of Montpellier (LIRMM) to become computer controllable. The different hardware components of the modified vehicle\u2019s hardware are detailed in [10]. The experiments have been performed in a 5 m3 pool" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0001704_978-3-642-19373-6_6-Figure6.5-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0001704_978-3-642-19373-6_6-Figure6.5-1.png", "caption": "Fig. 6.5 Three ways of doing work on a polyelectrolyte gel. Mechanical loads are applied by hanging a field of weights to the network. Chemical loads are applied using a field of pumps to injecting solvent molecules or solute ions into the gel. Electric loads are applied by connecting a field of batteries to the gel.", "texts": [ " A large number of polyelectrolyte chains can form a three-dimensional network by crosslinks. The network can imbibe liquid solution and swell, resulting in a polyelectrolyte gel. The amount of swelling can be regulated by geometric constraints, mechanical forces, ionic concentrations, and electric fields. In this section, by studying the non-equilibrium thermodynamics of polyelectrolyte gels, we will extend the field theory of neutral gels developed in previous sections to couple electrochemistry and large deformation. As illustrated in Fig. 6.5, there are three ways for the external agents to do work upon a polyelectrolyte gel. Just as in a neutral gel, the work done by a field of mechanical forces is given by Eq. (6.1). To define the fields of chemical potentials of mobile species, we imagine attaching fields of pumps to the gel. The pumps are connected to idealized reservoirs, and prescribe a time-dependent field of chemical potential \u03bca(X, t) for mobile species a. Upon injection of \u03b4ra number of particles of species a, the corresponding pump does work \u03bca\u03b4ra" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0000359_1.4004623-Figure5-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0000359_1.4004623-Figure5-1.png", "caption": "Fig. 5 Meshed FEM model with bolts", "texts": [ " In the second model, the elements are attached and the bolts are included in order to evaluate their temperatures. Due to symmetry of both the geometry and the thermal loading, the two models include only 180 sector of the joint subjected to the temperatures as obtained for each pass using Eq. (5). To simplify the models, the tube sheet is represented by a circular plate with only a central hole having an internal diameter such that its material volume is equivalent to the tube sheet with holes. This simplification is acceptable for the heat transfer study. Figure 5 show a general view of the meshed model with bolts. In both models, the gasket was not modeled and the initial tightening of bolts and the internal pressure were not applied. Therefore, the temperature distributions and the deflections are the result of the heat flow from the hotter internal fluid. The temperature distribution in the bolted joint is three-dimensional as shown in Fig. 6. Furthermore, the temperature distribution of the bolted joint is complicated by the presence of the bolts. Nevertheless, Fig" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0003716_6.2019-4392-Figure18-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0003716_6.2019-4392-Figure18-1.png", "caption": "Fig. 18 Printed CAD Model of Pintle", "texts": [ " Therefore, the added material (0.127 mm) on the outer diameter under the flange provided the machinist with more options for cutting in increments to ensure the part meets tolerance requirements. For more information on center annulus tolerancing, refer to the part drawing in the Appendix. Pintle D ow nl oa de d by U N IV E R SI T Y O F G L A SG O W o n Se pt em be r 2, 2 01 9 | h ttp :// ar c. ai aa .o rg | D O I: 1 0. 25 14 /6 .2 01 9- 43 92 The pintle, like the center annulus, was a straightforward design for printing. Fig. 18 showcases the print direction in which it was printed. The pintle did not require any additional support material other than the 2 mm (0.079\u201d) added to its foundation for build plate removal as mentioned earlier. However, like the center annulus, the outer diameter was slightly increased by 0.127 mm (0.005\u201d) to provide the machinist with an increased safety factor to ensure the final dimension falls within tolerance requirements. As for the most crucial orifices at the tip of the pintle, two versions were printed via two different pintles" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0001619_icicip.2012.6391538-Figure6-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0001619_icicip.2012.6391538-Figure6-1.png", "caption": "Fig. 6. Torque vector decomposition during the commutation.", "texts": [ " Since the electromagnetic torque produced by one phase winding is determined by the phase magnetomotive force, stator pole\u2019s centerline of each phase is the stable position of electromagnetic torque vector of each phase. The torque vector of three-phase switched reluctance is shown in Fig.5. If one phase winding is energized during the commutation, there will be a great torque ripple. As a result, synthetic directions of magnetomotive force are added to the three basic directions, which forms six torque vector directions. In the practical control, the rotor sharing function is used to improve the torque control accuracy and to reduce the torque ripple as shown in Fig.6. In order to obtain constant synthetic torque T, the torque component of each phase winding is calculated by the rotor position \u03b8, which makes the rotating of synthetic torque smoothly. Take the two-phase commutation between A and B for example, suppose that the amplitude of reference torque is T and the rotor position angle is \u03b8, the torque vector can be decomposed into A direction and B direction respectively. According to the spatial relationships of the vectors, the instantaneous torque of phase A is 2 sin(120 ) 3A TT \u03b8= \u2212 (7) Similarly, the instantaneous torque of phase B is 2 s i n 3B At this point, the switch status of phase A and phase B is determined respectively by the reference torque and the instantaneous torque" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0000047_s13272-013-0095-7-Figure14-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0000047_s13272-013-0095-7-Figure14-1.png", "caption": "Fig. 14 Principle relation between suspension point location and helicopter-load coupling", "texts": [ " If the translational dynamics of the helicopter would be identical in both directions, e.g. by constraining the rotational dynamics and supposing that the direct damping derivatives Xu and Yv are equal, the two load modes would have the same frequency and damping. If the load is suspended outside of the helicopter CG, the load dynamics is directly coupled to the rotational movement of the helicopter. The longitudinal load mode is then influenced mainly by the helicopter pitch motion and the lateral load mode by the roll motion. This relation is depicted in Fig. 14 and it can be easily concluded that a change in attitude leads to an acceleration of the suspension point and equally of the suspended load. The order and direction of acceleration is defined by the location of the suspension point relative to the CG, respectively. With increasing distance of the suspension pointing to the CG the coupling of the load modes to the helicopter\u2019s roll and pitch motion gets more pronounced. The prevailing difference between the dynamics of these two axes leads to the diverged tendency of the load mode poles in Fig. 15. For the suspension below the CG the frequency and damping of the longitudinal load mode increase with distance, whereas the lateral load mode pole moves to lower frequency and damping values. Figure 14 indicates that the load suspension above the CG reverses the direction of the load acceleration, resulting in the change of load mode tendency in Fig. 15. While the frequency of the longitudinal load mode decreases, the frequency of the lateral load mode increases. Both load modes get more instable with increasing distance. The influence of the suspension point location on the load mode damping can be explained by the principle of stabilising a suspended load, which states that a swinging load can be stabilised by following the load motion", " In contrast, suspending the load below the CG causes that initial load deflection and roll damping effect with the following lateral motion act in the same direction as the load deflection, resulting in a stabilising effect. The decrease in damping of the lateral load mode with increasing distance below the CG can be explained by the increase in the dihedral derivative Lv, which has a destabilizing effect on the load mode damping. Within the project HALAS the effect of the lateral load suspension is of particular interest and was examined by varying the suspension point laterally to the right and left side of the helicopter, when looking in direction of its nose. From Fig. 14 it becomes clear that the lateral variation determines the longitudinal load acceleration as a result of the direct coupling of the load motion to the helicopter\u2019s yaw motion. The results of the variation are shown in Fig. 16. Obviously, the suspension on the left side leads to an unfavourable effect on the longitudinal load mode, i.e. decreasing damping. The lateral mode becomes more stable with increasing distance to the CG, regardless of the suspension side. The analysis of the suspension point variation is completed with the variation in longitudinal direction, i.e. forward and backward direction relative to the CG. The results in Fig. 17 reveal a main impact on the lateral load mode, which can be explained also by Fig. 14. In the same way, as the lateral variation affects the longitudinal load acceleration, the variation along the helicopter x-axis determines the lateral load acceleration. A suspension in backward direction relative to the CG has a stabilising effect in contrast to the variation in the opposite direction. It is noteworthy that the longitudinal load mode is affected only slightly. The reason for the different behaviour compared to the lateral variation, where the lateral load mode pole changed markedly, can be found in the mode shapes of the load modes at the starting point of the variations, i" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0001301_s12206-010-0326-3-Figure2-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0001301_s12206-010-0326-3-Figure2-1.png", "caption": "Fig. 2. Linear inverted dumbbell model with nonzero GH& .", "texts": [ " The radial distance of the COM of the dumbbell from point P is denoted by r, and the distance of each mass particle from the COM of the dumbbell is l. \u03c1 denotes the angular position of the dumbbell COM and \u03b8 denotes the attitude angle of the dumbbell axis with respect to the radial direction from point P to the dumbbell COM. The dumbbell is free to spin about its midpoint. A pin can be used to lock the dumbbell into the position shown in Fig. 1. When the pin is withdrawn, the dumbbell is free to rotate about its midpoint, as shown in Fig. 2. The sum of the moments of all external forces must equal the change of angular momentum of the dumbbell. Since the dumbbell is symmetric, the COM is at the support point. The angular momentum of the dumbbell with respect to point P is 2 2 1 1 ( ) ( ) 2 2p i i G i i G G G i i m mH r v r l v mr r H = = = \u00d7 = + = \u00d7 +\u2211 \u2211 & (1) where GH is the angular momentum of the dumbbell with respect to the COM G. The parallel axis theorem gives 2 1 ( ) 2G i G i i mH l l\u03b8 = = \u00d7 \u00d7\u2211 & (2) where 2 2a bm m m= = and G\u03b8& is the angular velocity of the dumbbell" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0000023_s11012-013-9808-6-Figure1-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0000023_s11012-013-9808-6-Figure1-1.png", "caption": "Fig. 1 The studied contacts. A\u2014The inlet zone; B\u2014The Hertzian contact zone", "texts": [ "com Keywords Hydrodynamic lubrication \u00b7 Boundary slippage \u00b7 Load-carrying capacity \u00b7 Friction coefficient \u00b7 Line contact Nomenclature b half Hertzian contact width E\u2032 equivalent Young\u2019s elastic modulus of two contact surfaces [1] fa,fb friction coefficients at the faster and slower moving contact surfaces in the present contact respectively fa,n, fb,n friction coefficients at the faster and slower moving contact surfaces in the conventional hydrodynamic lubricated line contact respectively Ff,a,Ff,b friction forces per unit contact length at the faster and slower moving contact surfaces in the present contact respectively Ff,a,n,Ff,b,n friction forces per unit contact length at the faster and slower moving contact surfaces in the conventional hydrodynamic lubricated line contact respectively G =\u03b1E\u2032 h film thickness hc film thickness at the location where dp/dx = 0 H =h/R Hc =hc/R Hc,N central film thickness in a hydrodynamic lubricated line contact when the fluid is Newtonian (with no interfacial slippage at both the contact surfaces) J1, J4, J5 integrations mw, rit dimensionless variables dependent on G, U and W [2] p film pressure ps interfacial solidification pressure in the present designed contacts ps,conv interfacial solidification pressure in a conventional contact pmax maximum pressure in the contact P =\u03b1p R radius of the cylinder or the equivalent radius of two contact surfaces S slide-roll ratio, 2(ua \u2212 ub)/(ua + ub) ua circumferential speed of the faster moving contact surface ub circumferential speed of the slower moving contact surface u\u0304a film velocity at the faster moving contact surface u\u0304b film velocity at the slower moving contact surface u rolling speed, (ua + ub)/2 U =u\u03b7a/(E \u2032R) Ua =ua\u03b7a/(E \u2032R) Ub =ub\u03b7a/(E \u2032R) w load per unit contact length of the contact W =w/(E\u2032R) x coordinate X =x/b \u03bb =G \u221a 8W/\u03c0 \u03b7 fluid viscosity \u03b7a fluid viscosity at ambient pressure \u03b1 fluid viscosity-pressure index \u03b1\u03c4 interfacial shear strength-pressure proportionality at high pressures in a conventional hydrodynamic lubricated line contact \u03b2\u03c4 interfacial shear strength-pressure proportionality at high pressures in the present designed contacts \u03c1 fluid density \u03c1c fluid density at the location where dp/dx = 0 \u03c4a shear stress at the faster moving contact surface \u03c4b shear stress at the slower moving contact surface \u03c4sa fluid-contact interfacial shear strength at the faster moving contact surface \u03c4sb fluid-contact interfacial shear strength at the slower moving contact surface \u03c4sb,1 fluid-contact interfacial shear strength at the stationary contact surface in the inlet zone in Fig. 1(c) \u03c4l0, \u03c4 \u2032 sa,0, \u03c4sa,0, \u03c4s0 constants ua interfacial slipping velocity at the faster moving contact surface ub interfacial slipping velocity at the slower moving contact surface \u03c4\u0304c,1 a dimensionless critical shear stress, Eq. (40) \u03c4\u0304c,2 a dimensionless critical shear stress, Eq. (38) \u03c4\u0304l0 = \u03c4l0/E \u2032 \u03c4\u0304sa = \u03c4sa/E \u2032 \u03c4\u0304sb = \u03c4sb/E \u2032 \u03c4\u0304s0 = \u03c4s0/E \u2032 \u03c4\u0304 \u2032 sa,0 = \u03c4 \u2032 sa,0/E \u2032 \u03c4\u0304sa,0 = \u03c4sa,0/E \u2032 \u03c4\u0304sb,1 = \u03c4sb,1/E \u2032 Conventional theories for hydrodynamics are based on the assumption of no fluid-contact interfacial slippage", " The obtained results are compared and discussed. The benefits of the application of the boundary slippage in a hydrodynamic lubricated line contact are concluded. Figures 1(a), (b) and (c) respectively show the hydrodynamic lubricated line contacts formed between cylinders and planes studied in the present paper, which are equivalent to real contacts. In these three figures, the upper contact surface is moving faster with the circumferential speed ua , while the lower contact surface is moving slower with the speed ub except in Fig. 1(c) it is stationary with ub = 0. In Fig. 1(a), the boundary slippage is artificially augmented at the whole lower contact surface with the low interfacial shear strength, while at the whole upper contact surface no boundary slippage occurs. In Fig. 1(b), the boundary slippage is artificially augmented at the upper contact surface in the Hertzian contact zone, while at the remaining parts of the upper contact surface and at the whole lower contact surface no boundary slippage occurs. In Fig. 1(c), at the upper contact surface the boundary slippage is only artificially augmented in the Hertzian contact zone, at the lower con- tact surface it is only artificially augmented in the inlet zone, at the remaining parts of the two contact surfaces no boundary slippage occurs. The artificial augmentation of the boundary slippage can be realized by coating the slipping contact surface with a certain kind of oil-phobic material for yielding the low fluidcontact interfacial shear strength. While for preventing the boundary slippage at a certain part of the contact surface, a certain kind of oil-philic material can be coated there for yielding the high fluid-contact interfacial shear strength. 3.1 For the contact in Fig. 1(a) In Fig. 1(a), when the interfacial shear strength at the lower contact surface is very low, the film slips at the whole lower contact surface, and the shear stress at the lower contact surface is \u03c4b = \u03c4sb , where \u03c4sb is the interfacial shear strength at the lower contact surface. The shear stress at the upper contact surface is \u03c4a = \u03c4sb + hdp/dx. Usually, the term hdp/dx is not large; In that contact, the value of \u03c4a can then be lower than \u03c4sa so that no film slippage occurs at the upper contact surface, where \u03c4sa is the interfacial shear strength at the upper contact surface. For large slideroll ratios, this is very important for improving the load-carrying capacity of the contact, as in a conventional contact at large slide-roll ratios the film slips at both the contact surfaces in the Hertzian contact zone due to the surface sliding for medium and heavy loads and it greatly reduces the load-carrying capacity of the contact [1, 21]. On the other hand, the friction coefficient of the contact in Fig. 1(a) will be very small because of the very low value of \u03c4sb in a rolling and sliding condition. The interfacial slipping velocity at the lower contact surface in Fig. 1(a) is [1]: ub = ua \u2212 ub \u2212 h2 2\u03b7 dp dx \u2212 \u03c4sbh \u03b7 (1) where h is the film thickness, p is the film pressure, and \u03b7 is the fluid viscosity. For medium and heavy loads, in the Hertzian contact zone \u03c4sbh/\u03b7 \u223c 0 and h2(dp/dx)/(2\u03b7) \u223c 0. In this condition, the interfacial slipping velocity at the lower contact surface in the Hertzian contact zone in Fig. 1(a) is: ub \u2248 ua \u2212 ub (2) The Reynolds equation in the Hertzian contact zone in Fig. 1(a) for medium and heavy loads is then: \u03c1h3 \u03b7 dp dx = 12ua(\u03c1h \u2212 \u03c1chc) (3) where \u03c1 is the fluid density, hc and \u03c1c are respectively the film thickness and the fluid density at the location where dp/dx = 0. Equation (3) can be re-expressed as: h2 ua\u03b7 dp dx = 12 ( 1 \u2212 \u03c1chc \u03c1h ) (4) For medium and heavy loads, in the Hertzian contact zone usually h2(dp/dx)/(ua\u03b7) < 1 \u00d7 10\u22122. In this condition, in the Hertzian contact zone thus \u03c1chc \u2248 \u03c1h. For incompressible fluids, in the Hertzian contact zone hc \u2248 h. Even for normal compressible fluids, in the Hertzian contact zone in Fig. 1(a) there is hc \u2248 h. In the Hertzian contact zone, when h \u2248 hc , the film pressure is very approximate to the Hertzian contact pressure in the dry contact in the elastic contact condition. An inlet zone analysis can then be taken for the contact in Fig. 1(a) for the condition of medium and heavy loads. Governing equation For medium and heavy loads, the dimensionless film thickness in the inlet zone in Fig. 1(a) is [22]: H(X) = Hc \u2212 4W \u03c0 { X \u221a X2 \u2212 1 + ln [\u2212X \u2212 \u221a X2 \u2212 1 ]} for X \u2264 \u22121 (5) where X = x/b, H = h/R, Hc = hc/R, W = w/(E\u2032R). Here, b is the half Hertzian contact width, R is the radius of the cylinder or the equivalent radius of two contact surfaces, w is the load per unit contact length of the contact, and E\u2032 is the equivalent Young\u2019s elastic modulus of two contact surfaces. When the interfacial shear strength at the lower contact surface is very low, the film slips at the whole lower contact surface. For this case, neglecting the fluid compressibility and assuming the fluid within the film as Newtonian, the dimensionless Reynolds equation in the inlet zone in Fig. 1(a) is: dP dX = 3\u03bbUae P H \u2212 Hc H 3 \u2212 3\u03bb\u03c4\u0304sb 2H (6) where P = \u03b1p, \u03bb = G \u221a 8W/\u03c0 , G = \u03b1E\u2032, Ua = ua\u03b7a/(E \u2032R), and \u03c4\u0304sb = \u03c4sb/E \u2032. Here, \u03b1 is the fluid viscosity-pressure index, and \u03b7a is the fluid viscosity at ambient pressure. For \u03c4\u0304sb = 0, Eq. (6) becomes: dP dX = 3\u03bbUae P H \u2212 Hc H 3 (7) Integrating Eq. (7) gives:\u222b P(\u22121) 0 e\u2212P dP = 3\u03bbUa \u222b \u22121 \u2212\u221e H \u2212 Hc H 3 dX (8) It is obtained from Eq. (8) that: 3\u03bbUa \u222b \u22121 \u2212\u221e H \u2212 Hc H 3 dX = 1 \u2212 e\u2212P(\u22121) (9) For medium and heavy loads, e\u2212P(\u22121) 1, thus Eq. (9) becomes: 3\u03bbUa \u222b \u22121 \u2212\u221e H \u2212 Hc H 3 dX \u2248 1 (10) From Eq", "2041 ( lg W Hc )2 \u2212 2.1124 lg W Hc \u2212 1.2725 for 0.01 \u2264 W Hc \u2264 200 (13) By substituting Eq. (13), it is then obtained from Eq. (11) that: \u22120.2041 ( lg W Hc )2 + 0.8876 lg W Hc + lg(6.0954GUa) \u2212 1.5 lgW \u2212 1.2725 = 0 for 0.01 \u2264 W Hc \u2264 200 (14) In a rolling and sliding condition, the friction force in a hydrodynamic lubricated line contact mainly arises from the Hertzian contact zone especially for medium and heavy loads. The friction forces per unit contact length at the lower and upper contact surfaces in Fig. 1(a) are respectively: Ff,b = \u222b b \u2212b \u03c4sbdx (15) and Ff,a = \u222b b \u2212b ( \u03c4sb + dp dx h ) dx = Ff,b + hc \u222b b \u2212b dp = Ff,b (16) When \u03c4sb approaches to zero, it is obtained from Eqs. (15) and (16) that Ff,a = Ff,b \u2192 0. Thus for this case, the friction coefficients at the lower and upper contact surfaces in Fig. 1(a) are fa = fb \u2192 0. 3.2 For the contact in Fig. 1(b) In Fig. 1(b), when the interfacial shear strength at the upper contact surface is lower than that at the lower contact surface, the film first slips at the upper contact surface in the Hertzian contact zone due to the surface sliding. In this case, the shear stress at the upper contact surface in the Hertzian contact zone is \u03c4a = \u03c4sa . The shear stress at the lower contact surface in the Hertzian contact zone is \u03c4b = \u03c4sa \u2212 hdp/dx. For hdp/dx not large, when \u03c4sa is limited, the magnitude of \u03c4b in the Hertzian contact zone can be limited to lower than the interfacial shear strength at the lower contact surface so that the film does not slip at the lower contact surface in the Hertzian contact zone. The film in the inlet zone does not slip at the upper contact surface because of the shear stress lower than the interfacial shear strength there, and it also not slip at the lower contact surface in the inlet zone because of the high interfacial shear strength at that surface. The interfacial slipping velocity at the upper contact surface in Fig. 1(b) is [1]: ua = ub \u2212 ua \u2212 h2 2\u03b7 dp dx + \u03c4sah \u03b7 (17) For a load not very light and a considerable surface sliding, in the Hertzian contact zone |\u03c4sah/\u03b7 \u2212 h2(dp/dx)/(2\u03b7)|/(ua \u2212 ub) 1. In this condition, the interfacial slipping velocity at the upper contact surface in the Hertzian contact zone in Fig. 1(b) is: ua \u2248 ub \u2212 ua (18) The film velocity at the upper contact surface in the Hertzian contact zone is then u\u0304a \u2248 ub . Based on Eq. (18), the Reynolds equation in the Hertzian contact zone in Fig. 1(b) is: \u03c1h3 \u03b7 dp dx = 12ub(\u03c1h \u2212 \u03c1chc) (19) Equation (19) can be re-expressed as: h2 12ub\u03b7 dp dx = 1 \u2212 \u03c1chc \u03c1h (20) For a load not very light and a low value of ub ( = 0), in the Hertzian contact zone, the pressure will be sufficiently high to give |h2(dp/dx)/(12ub\u03b7)| < 0.1. In this condition, it can thus be taken that \u03c1chc \u2248 \u03c1h in the Hertzian contact zone. Also for normal compressible fluids, in the Hertzian contact zone in Fig. 1(b) it can be taken that hc \u2248 h when \u03c1chc \u2248 \u03c1h. In the Hertzian contact zone, when h \u2248 hc, the film pressure is thus close to the Hertzian contact pressure in the dry contact in the elastic contact condition. An inlet zone analysis can then also be taken for the contact in Fig. 1(b) based on the Hertzian contact pressure assumption in the Hertzian contact zone. Governing equation Based on the Hertzian contact pressure assumption in the Hertzian contact zone, the dimensionless film thickness in the inlet zone in Fig. 1(b) is also expressed by Eq. (5). Based on Eq. (18), by neglecting the fluid compressibility and taking the fluid in the inlet zone as Newtonian, the dimensionless Reynolds equation in the inlet zone in Fig. 1(b) is: e\u2212P dP dX = 12\u03bbU H 2 \u2212 12\u03bbUb Hc H 3 (21) where U = u\u03b7a/(E \u2032R) and Ub = ub\u03b7a/(E \u2032R). Here, the rolling speed u is (ua + ub)/2. Integrating Eq. (21) yields:\u222b P(\u22121) 0 e\u2212P dP = 1 \u2212 e\u2212P(\u22121) = 12\u03bbU \u222b \u22121 \u2212\u221e 1 H 2 dX \u2212 12\u03bbUbHc \u222b \u22121 \u2212\u221e 1 H 3 dX (22) When e\u2212P(\u22121) 1, Eq. (22) can be expressed as: 12\u03bbUJ4 ( W Hc ) \u2212 12\u03bbUbJ5 ( W Hc ) \u2212 H 2 c = 0 (23) where J4(W/Hc) = \u222b \u22121 \u2212\u221e (Hc/H)2dX and J5(W/Hc) = \u222b \u22121 \u2212\u221e (Hc/H)3dX. Formulations of J4 and J5 For 0.01 \u2264 W/Hc \u2264 200, the functions J4 and J5 can be respectively numerically integrated by substituting Eq", " (24) and (25) are respectively compared with the numerical integration results of J4 and J5 in Fig. 2. It is shown that the regression accuracy of Eqs. (24) and (25) are fairly good. Film thickness equation Substituting Eqs. (24) and (25) into Eq. (23) and rearranging gives: 3.928618GUbW 1/2 2 \u2212 S ( W Hc )\u22120.210252 lg W Hc \u22121.297268 \u2212 0.673574GUbW 1/2 ( W Hc )\u22120.352672 lg W Hc \u22121.737197 \u2212 H 2 c = 0 for 0.01 \u2264 W/Hc \u2264 200 (26) where the slide-roll ratio S is 2(ua \u2212 ub)/(ua + ub). For given values of S, G, Ub and W , solving Eq. (26) gives the central film thickness in the contact in Fig. 1(b). The interfacial shear strength \u03c4sa at the upper contact surface in the Hertzian contact zone in Fig. 1(b) is predicted by the following equation [24]: \u03c4sa = { \u03c4 \u2032 sa,0 + \u03b2\u03c4p for p \u2265 ps \u03c4sa,0 for p < ps (27) where ps is the interfacial solidification pressure, \u03b2\u03c4 is the interfacial shear strength-pressure proportionality at high pressures, \u03c4 \u2032 sa,0 and \u03c4sa,0 are constants. In the contact in Fig. 1(b), \u03c4sa is so chosen that the film slippage occurs nearly at the whole upper contact surface in the Hertzian contact zone but it does not enter into the inlet zone. According to this requirement, the following equation is obtained [24]: 2SUbe 0.6G \u221a W/(2\u03c0) (2 \u2212 S)Hc + Hcmwrit 8 ( 1 \u2212 r2 it )\u22120.5 = \u23a7\u23a8 \u23a9 \u03c4\u0304 \u2032 sa,0 + 0.6\u03b2\u03c4 \u221a W 2\u03c0 for 0.6E\u2032\u221aW/(2\u03c0) \u2265 ps \u03c4\u0304sa,0 for 0.6E\u2032\u221aW/(2\u03c0) < ps (28) where \u03c4\u0304 \u2032 sa,0 = \u03c4 \u2032 sa,0/E \u2032, \u03c4\u0304sa,0 = \u03c4sa,0/E \u2032, and mw and rit are dimensionless variables dependent on G, U and W [2]. For achieving the effect of reducing the friction coefficient by the artificially augmented boundary slippage, in the contact in Fig. 1(b) the load should be relatively light to avoid too high shear stresses at the upper contact surface in the Hertzian contact zone, as can be seen from the left hand side of Eq. (28).When the maximum pressure in the Hertzian contact zone is lower than ps , \u03c4sa = \u03c4sa,0 in the whole Hertzian contact zone. In a rolling and sliding condition, the friction force per unit contact length at the upper contact surface in Fig. 1(b) is: Ff,a = \u222b b \u2212b \u03c4sadx = \u222b b \u2212b \u03c4sa,0dx = 2b\u03c4sa,0 (29) The friction coefficient at the upper contact surface is then: fa = Ff,a w = ( 8 \u03c0W )1/2[4SUbe 0.6G \u221a W/(2\u03c0) (2 \u2212 S)Hc + Hcmwrit (1 \u2212 r2 it ) \u22120.5 4 ] (30) The friction force per unit contact length at the lower contact surface is: Ff,b = \u222b b \u2212b ( \u03c4sa \u2212 dp dx h ) dx = Ff,a \u2212 hc \u222b b \u2212b dp = Ff,a (31) The friction coefficient at the lower contact surface is fb = fa . 3.3 For the contact in Fig. 1(c) In the contact in Fig. 1(c), in the Hertzian contact zone, when the fluid-contact interfacial shear strength \u03c4sa at the upper contact surface is low, the boundary slippage occurs at the upper contact surface because of the interfacial shear stress exceeding the interfacial shear strength due to the surface sliding. While, at the lower contact surface in the Hertzian contact zone the shear stress is \u03c4b = \u03c4sa \u2212 (dp/dx)h; For modest loads the magnitude of this shear stress can easily be less than the fluid-contact interfacial shear strength at the lower contact surface in the Hertzian contact zone so that no boundary slippage occurs at the lower contact surface in the Hertzian contact zone. The interfacial slipping velocity at the upper contact surface in the Hertzian contact zone is still expressed by Eq. (17), where ub = 0. For the load heavy enough, it is obtained from Eq. (17) that ua \u2248 \u2212ua due to the large viscosity \u03b7 of the fluid under high pressures in the Hertzian contact zone. For this case, the fluid film velocity at the upper contact surface in the Hertzian contact zone is u\u0304a \u2248 0. For the contact in Fig. 1(b), when ub > 0 and the load is heavy enough, the pressure in the Hertzian contact zone is close to the Hertzian contact pressure in dry contact and the film thickness in the Hertzian contact zone is nearly constant. For ub \u2192 0, this conclusion also stands. It can thus be obtained that in the contact in Fig. 1(c), when ub = 0, the above conclusion also stands for a load heavy enough. An inlet zone analysis can thus be taken for the contact in Fig. 1(c) based on this conclusion. Governing equation Based on the Hertzian contact pressure assumption in the Hertzian contact zone, the dimensionless film thickness in the inlet zone in the contact in Fig. 1(c) is still expressed by Eq. (5). When the interfacial shear strength \u03c4sb,1 at the lower contact surface in the inlet zone is very low, the film slips at the whole lower contact surface in the inlet zone in Fig. 1(c). For this case, neglecting the fluid compressibility and assuming the fluid within the film as Newtonian, the dimensionless Reynolds equation in the inlet zone in Fig. 1(c) is: dP dX = 3\u03bbUae P H 2 \u2212 3\u03bb\u03c4\u0304sb,1 2H (32) where \u03c4\u0304sb,1 = \u03c4sb,1/E \u2032. For \u03c4\u0304sb,1 = 0, Eq. (32) becomes: dP dX = 3\u03bbUae P H 2 (33) Integrating Eq. (33) gives: 3\u03bbUa \u222b \u22121 \u2212\u221e 1 H 2 dX = 1 \u2212 e\u2212P(\u22121) (34) For the load heavy enough, e\u2212P(\u22121) 1, thus Eq. (34) becomes: 3\u03bbUaJ4(W/Hc) H 2 c \u2248 1 (35) Substituting Eq. (24) into Eq. (35) and rearranging gives the following film thickness equation for the contact in Fig. 1(c): \u22120.2103 ( lg W Hc )2 + 0.7027 lg W Hc + lg ( 0.4911GUaW \u2212 3 2 ) = 0 for 0.01 \u2264 W Hc \u2264 200 (36) In a simple sliding, the friction force in the contact in Fig. 1(c) mainly arises from the Hertzian contact zone. If pmax \u2264 ps , where pmax is the maximum pressure in the Hertzian contact zone, \u03c4sa = \u03c4sa,0. According to this, the friction coefficient at the upper contact surface in the contact in Fig. 1(c) is: fa = \u222b b \u2212b \u03c4sadx w = 2\u03c4\u0304sa,0 \u221a 8 \u03c0W (37) Because of the symmetry of the pressure distribution in the Hertzian contact zone, the friction coefficient at the lower contact surface in the contact in Fig. 1(c) is: fb = fa . For generating the boundary slippage at the upper contact surface in the whole Hertzian contact zone, it should be satisfied that \u03c4\u0304sa \u2264 \u03c4\u0304c,2 [1, 24], where \u03c4\u0304sa = \u03c4sa/E \u2032 and \u03c4\u0304c,2 = Uae 0.6G \u221a W/(2\u03c0) Hc + 1 8 mwHcrit ( 1 \u2212 r2 it )\u22120.5 (38) The maximum shear stress at the upper contact surface in the inlet zone is [25]: \u03c4c,1 = \u03c4sb,1 + Max ( dp dx ) hc = \u03c4sb,1 + 1 4 E\u2032Hcmwrit ( 1 \u2212 r2 it )\u22120.5 (39) For avoiding the boundary slippage occurrence at the upper contact surface in the inlet zone, it should be satisfied that \u03c4\u0304sa \u2265 \u03c4\u0304c,1, where \u03c4\u0304c,1 = \u03c4\u0304sb,1 + 1 4 Hcmwrit ( 1 \u2212 r2 it )\u22120.5 (40) The interfacial shear strength condition at the upper contact surface in Fig. 1(c) is \u03c4\u0304c,1 \u2264 \u03c4\u0304sa \u2264 \u03c4\u0304c,2. If \u03c4\u0304sa = \u03c4\u0304c,1, the achievable lowest friction coefficient in the contact in Fig. 1(c) for \u03c4\u0304sb,1 \u2192 0 is: fb = fa = 1 2 Hcmwrit ( 1 \u2212 r2 it )\u22120.5 \u221a 8 \u03c0W (41) In a conventional rolling and sliding hydrodynamic lubricated line contact (with the same interfacial properties at both the contact surfaces), the film thickness at the Hertzian contact center is reduced due to the interfacial slippage occurring at both the contact surfaces [1]. Due to the interfacial slippage effect, for 0 \u2264 S \u2264 2, the film thickness at the Hertzian contact center in a conventional hydrodynamic lubricated line contact is expressed as [24]: H v\u2212p c (S) = { Hc,N for pmax \u2264 0", "572\u03c0(ps,conv/E \u2032)2 2b\u03c4s0 for W < 2\u03c0(ps,conv/E \u2032)2 (44) where \u03b1\u03c4 is the interfacial shear strength-pressure proportionality at high pressures in a conventional hydrodynamic lubricated line contact, ps,conv is the interfacial solidification pressure of a conventional hydrodynamic lubricated line contact, and \u03c4l0 and \u03c4s0 are constants [24]. According to Eq. (44), the friction coefficients at the lower and upper contact surfaces in a conventional hydrodynamic lubricated line contact are: fa,n = fb,n = \u23a7\u23aa\u23aa\u23aa\u23aa\u23aa\u23aa\u23a8 \u23aa\u23aa\u23aa\u23aa\u23aa\u23aa\u23a9 4 \u221a 2 \u03c0W \u03c4\u0304l0 + \u03b1\u03c4 for W \u2265 5.572\u03c0(ps,conv/E \u2032)2 2\u03c4\u0304s0 \u221a 8 \u03c0W for W < 2\u03c0(ps,conv/E \u2032)2 (45) where \u03c4\u0304l0 = \u03c4l0/E \u2032 and \u03c4\u0304s0 = \u03c4s0/E \u2032. When 2\u03c0(ps,conv/E \u2032)2 \u2264 W < 5.572\u03c0(ps,conv/E \u2032)2, the friction coefficients fa,n, fb,n are linearly interpolated according to Eq. (45). 5.1 For the contact in Fig. 1(a) Figure 3(a) plots the central film thickness Hc in the contact in Fig. 1(a) against the slide-roll ratio S respectively for W = 5 \u00d7 10\u22125 and W = 5 \u00d7 10\u22124 when Ua = 1 \u00d7 10\u22129 and G = 4500. It is compared with the value of Hc in the conventional contact for the same operating condition. The central film thickness in the contact in Fig. 1(a) is independent of the slideroll ratio, while the central film thickness in the conventional contact is drastically reduced with the slideroll ratio increase for the given operating conditions. For W = 5 \u00d7 10\u22125, the value of Hc in the contact in Fig. 1(a) is greater than that in the conventional contact when S is over 1.1. The difference between these two Hc values is large for the slide-roll ratios over 1.5. For the heavier load W = 5 \u00d7 10\u22124, the value of Hc in the contact in Fig. 1(a) is greater than that in the conventional contact for the same operating condition when S is over 1.65. These show the benefit and application value of the designed contact in Fig. 1(a) for improving the load-carrying capacity for large slideroll ratios in the condition of medium and heavy loads. Figure 3(b) plots the similar phenomena as in Fig. 3(a) when Ua = 1 \u00d7 10\u22128 and G = 4500. The comparison between Figs. 3(a) and (b) shows that for a higher value of Ua the benefit of the contact in Fig. 1(a) in improving the load-carrying capacity is more pronounced for large slide-roll ratios. Figure 4 plots the friction coefficients fa and fb of the contact in Fig. 1(a) against the slide-roll ratio when \u03c4sb \u2192 0. The friction coefficients fa,n and fb,n of the conventional contact are also plotted in Fig. 4 respectively for W = 5 \u00d7 10\u22125 and W = 5 \u00d7 10\u22124 when \u03c4\u0304l0 = 9.57 \u00d7 10\u22125 and \u03b1\u03c4 = 0.036. It is shown that for a very low value of \u03c4sb the friction coefficient of the contact in Fig. 1(a) is very low in a rolling and sliding condition. The comparisons between fa (fb) and fa,n (fb,n) in Fig. 4 shows that the designed contact in Fig. 1(a) has the benefit of hugely reducing the friction coefficient as compared to the conventional contact when the slide-roll ratio is large and \u03c4sb is low. This shows the great application value of the designed contact in Fig. 1(a) in practice for reducing the friction force, energy loss and temperature rise in a contact. 5.2 For the contact in Fig. 1(b) Figure 5(a) plots the central film thickness in the contact in Fig. 1(b) against the slide-roll ratio respectively numerically calculated from Eq. (23) and predicted from Eq. (26) for different loads when Ub = 1\u00d710\u221212 and G = 4500. They are compared with the corresponding central film thickness in the conventional hydrodynamic lubricated line contact respectively numerically calculated (for Hc,N ) [23] and predicted from Eq. (42) (for Hc,N predicted from Eq. (43)). It is shown that the prediction by Eq. (26) is satisfactory especially for relatively light loads and large slide-roll ratios. The numerical calculations show that the central film thickness in the contact in Fig. 1(b) is greater than that in the conventional hydrodynamic lubricated line contact for the same operating condition. It shows that the designed contact in Fig. 1(b) with the augmented boundary slippage can improve the load-carrying capacity of the contact. This effect is very significant for large slide-roll ratios. It is noticed that the central film thickness in the contact in Fig. 1(b) predicted from Eq. (26) may be lower than that in the conventional hydrodynamic lubricated line contact predicted from Eq. (42) for the same operating condition at small slide-roll ratios and relatively heavy loads. This is entirely caused by the prediction errors of these two equations for the central film thickness, and can not cover the beneficial effect of the designed contact in Fig. 1(b) in increasing the load-carrying capacity. Figure 5(b) plots the similar phenomena as in Fig. 5(a) for heavier loads and the higher value of Ub when G = 4500. The numerical calculations consistently show that the designed contact in Fig. 1(b) improves the load-carrying capacity as compared to the conventional hydrodynamic lubricated line contact for the same operating condition. This effect is however reduced when the load is relatively heavy. For achiev- ing the effects of increasing the load-carrying capacity but reducing the friction coefficient by the boundary slippage in the contact in Fig. 1(b), the suitable slideroll ratio is significantly reduced with increasing Ub , as compared with Fig. 5(a). The prediction accuracy of Eq. (26) in Fig. 5(b) is satisfactory. However, the central film thickness predicted from Eq. (26) in the contact in Fig. 1(b) may still be lower than that predicted from Eq. (42) in the conventional hydrodynamic lubricated line contact for the same operating condition at small slide-roll ratios and increasingly heavy loads. Figure 5(c) plots the central film thickness against the slide-roll ratio as similarly done in Figs. 5(a) and (b) for heavier loads and a higher value of Ub than in the fore two figures when G = 4500. In Fig. 5(c), for W = 1 \u00d7 10\u22124, the Hc \u2212 S curve numerically calculated for the contact in Fig. 1(b) is overlaid with that numerically calculated for the conventional hydrodynamic lubricated line contact. It is shown in Fig. 5(c) that when the load is relatively heavy, the suitable slide-roll ratio for the contact in Fig. 1(b) is small and in this case the effect of improving the load-carrying capacity by the contact in Fig. 1(b) is minor. The comparison between Figs. 5(b) and (c) for the case W = 3 \u00d7 10\u22125 shows that for a given load and slideroll ratio, the effect of improving the load-carrying capacity by the boundary slippage in the contact in Fig. 1(b) is stronger with a higher value of Ub , as compared with the load-carrying capacity in the conventional hydrodynamic lubricated line contact for the same operating condition. Figure 6(a) plots the friction coefficients fa and fb in the contact in Fig. 1(b) against the slide-roll ratio respectively for the cases Ub = 5 \u00d7 10\u221211, W = 5 \u00d7 10\u22125 and Ub = 1 \u00d7 10\u221212, W = 1 \u00d7 10\u22125. These friction coefficients are respectively compared with the friction coefficients (fa,n and fb,n) in the conventional hydrodynamic lubricated line contact for the same load (i.e. for the same operating condition) when \u03c4\u0304s0 = 1.435 \u00d7 10\u22124. The comparison shows that the boundary slippage designed in the contact in Fig. 1(b) can greatly reduce the friction coefficient of the contact especially for relatively light loads. Figure 6(b) plots the friction coefficients in the contact in Fig. 1(b) against the slide-roll ratio respectively for Ub = 1 \u00d7 10\u221211and Ub = 5 \u00d7 10\u221211 when W = 3 \u00d7 10\u22125. They are compared with the friction coefficients in the conventional hydrodynamic lubricated line contact for the same load when \u03c4\u0304s0 = 1.435 \u00d7 10\u22124. The comparison also shows that the friction coefficient of the contact is greatly reduced by the designed boundary slippage in Fig. 1(b). 5.3 For the contact in Fig. 1(c) Figure 7(a) plots the dimensionless central film thickness Hc against the dimensionless load W respectively for the contact in Fig. 1(c), the conventional contact and the contact in Fig. 1(b) when Ua = 1.0 \u00d7 10\u221210 and G = 4500. It is shown that for the same operating condition the contact in Fig. 1(c) has a considerably higher lubricating film thickness i.e. loadcarrying capacity than the contact in Fig. 1(b), especially for lighter loads. However for relatively heavy loads, the contact in Fig. 1(c) seems to have no special advantage in the load-carrying capacity compared to the contact in Fig. 1(b). For medium loads the film thicknesses in the contact in Fig. 1(c) and the contact in Fig. 1(b) both are much higher than that in the con- ventional hydrodynamic lubricated line contact, which approaches to vanishing, for the same operating condition. However for light loads these advantages are not existent. Figure 7(b) plots the curves of Hc against W similar as in Fig. 7(a) when Ua = 1.0 \u00d7 10\u22129 and G = 4500. It is shown that from medium loads to comparatively heavy loads, in the load-carrying capacity, the contact in Fig. 1(c) is the most advantageous among these three contacts for the same operating condition. This advantage of the contact in Fig. 1(c) is obvious com- pared to the conventional contact. It is also obvious compared to the contact in Fig. 1(b) for medium loads; However it is very modest for comparatively heavy loads. Figure 7(c) plots the similar Hc versus W curves as in Figs. 7(a) and (b) when Ua = 1.0 \u00d7 10\u22128 and G = 4500. For a higher sliding speed Ua , the film thickness in the contact in Fig. 1(c) is significantly higher than those in the other two contacts when the load is medium or heavy. The comparisons among Figs. 7(a), (b) and (c) suggest that from the viewpoint of the load-carrying capacity the contact in Fig. 1(c) has an obvious advantage than the contact in Fig. 1(b) at a high sliding speed Ua . Figure 8(a) plots the friction coefficients against W respectively for the contact in Fig. 1(c), the conventional contact and the contact in Fig. 1(b) when Ua =1.0\u00d710\u221210, G = 4500, \u03c4\u0304sb,1 \u2192 0, and \u03c4\u0304sa = \u03c4\u0304c,1. In calculation for the conventional contact, it is taken that \u03c4\u0304l0 = 9.57 \u00d7 10\u22125 and \u03b1\u03c4 = 0.036. Compared to the friction coefficient of the conventional contact, the friction coefficients in the contact in Fig. 1(c) and in the contact in Fig. 1(b) both are much lower for the same operating condition. For the given operating condition, the friction coefficient in the contact in Fig. 1(c) is very low. Figure 8(b) plots the friction coefficients against W respectively for these three contacts similar as in Fig. 8(a) when Ua = 1.0 \u00d7 10\u22129, G = 4500, \u03c4\u0304sb,1 \u2192 0, and \u03c4\u0304sa = \u03c4\u0304c,1. It is also shown that for the given operating condition the contact in Fig. 1(c) as well as the contact in Fig. 1(b) is obviously advantageous in reducing the friction coefficient, compared to the conventional contact. It is noticed that at light loads the friction coefficient in the contact in Fig. 1(c) is considerably increased compared to that at relatively heavy loads. Figure 8(c) plots the friction coefficients against W respectively for these three contacts when Ua = 1.0 \u00d7 10\u22128, G = 4500, \u03c4\u0304sb,1 \u2192 0, and \u03c4\u0304sa = \u03c4\u0304c,1. It is shown that for the high sliding speed Ua the friction coefficient in the contact in Fig. 1(c) is the highest among these three contacts for the same operating condition and it can even reach nearly 0.15. However, for a high value of Ua , the contact in Fig. 1(b) still reduces the friction coefficient to vanishing, with a prominent advantage. The figure shows that the contact in Fig. 1(c) may be not suitable for the high sliding speed from the viewpoint of reducing the friction coefficient. The present paper proposes three modes of hydrodynamic lubrications in line contacts respectively artificially augmented with the boundary slippage for improving the load-carrying capacity but reducing the friction coefficient. In one mode of the lubrication, the boundary slippage is augmented at the whole slower moving contact surface, while at the faster moving contact surface the boundary slippage is absent" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0003254_s00170-019-03312-1-Figure2-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0003254_s00170-019-03312-1-Figure2-1.png", "caption": "Fig. 2 (a) Surface characteristics and terminology\u2014note the \u201clay\u201d or the dominant pattern, and (b) standard measurement pattern symbols [3]", "texts": [ " It is recommended that AM specific standards and simulation based assessment approaches for surface roughness quantification be developed based on the issues highlighted in this work. The surface topology or surface roughness typically has two structures associated with it: waviness and roughness [3]. The roughness is characterized by closely spaced \u201cwaves\u201d and is directly related to the fabrication process (i.e., cutting tool marks); whereas, the waviness element occurs at a lower frequency, and the causes may be related to structural elements within the system (i.e., clearances and vibrations within a machine) (Fig. 2). There are many measurement directions or \u201clay\u201d variants that can be applied to collect the surface roughness data. Surface roughness implies the texture and flaws are considered, but tolerance/form variations are not. The issues related to shrinkage, warpage, and distortion are not reflected for this research, but are issues within the AM process family. Two standard formulations for surface roughness are: Ra and Rq, where Ra is the average deviation from the profile, and Rq is the root mean square deviation" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0000637_s10846-013-9966-8-Figure2-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0000637_s10846-013-9966-8-Figure2-1.png", "caption": "Fig. 2 Quadrotor platform: 3D visualization (left) and experimental platform (right)", "texts": [ " 13 is a function of the state matrix A, that is Am = A \u2212 BK, where K is the gain matrix obtained applying the pole placement theory. The control matrix B is evaluated as B = [Bm Bum]. The state vector is x = [u v w p q r \u03c6 \u03d1 \u03c8]T , the control vector is u = [ 1 2 3 4]T where i is the rotational speed of the ith driving motor, defined by the control scheme of Table 3. The output vector (for the unmatched case) is y = [w p q r \u03c6 \u03d1]T . In Table 3 Ri for i = 1, . . . , 4 is the relative driving motor (see Fig. 2), H is the hover rotational speed (about 580 rad/s obtained from flight tests) and is the variation of the rotational speed with respect to the hover condition. The rotational speed of the ith driving motor is i = H + i. As detailed in [8], for a quadrotor UAV the longitudinal and lateral-directional planes cannot be decoupled, thus the cross-coupling derivatives are evaluated and included in the state space matrices. Uusing the theory presented in [23, 25] and the experimental results in [24], the state matrix can be defined as A = \u23a1 \u23a2\u23a2\u23a2\u23a2\u23a2\u23a2\u23a2\u23a2\u23a2\u23a2\u23a2\u23a2\u23a2\u23a2\u23a2\u23a2\u23a3 Xu Xv 0 Xp 0 0 0 \u2212mg 0 Yu Yv 0 Yp Yq 0 mg 0 0 Zu 0 Zw 0 0 0 0 0 0 Lu Lv 0 Lp Lq 0 0 0 0 Mu Mv Mw Mp Mq 0 0 0 0 0 0 Nw 0 0 Nr 0 0 0 0 0 0 1 H 0 0 0 0 0 0 0 0 0 1 0 0 0 0 0 0 0 0 0 1 H 0 0 0 \u23a4 \u23a5\u23a5\u23a5\u23a5\u23a5\u23a5\u23a5\u23a5\u23a5\u23a5\u23a5\u23a5\u23a5\u23a5\u23a5\u23a5\u23a6 " ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0003991_rusautocon.2019.8867688-Figure1-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0003991_rusautocon.2019.8867688-Figure1-1.png", "caption": "Fig. 1. The pattern of magnetic and electric fields for linear current", "texts": [ " The receiver system may contain measuring transducers to register the components or the induction module of the magnetic field, as well as other equivalent parameters of electromagnetic field strength. The choice a specific type of the emitting system, as well as the type and the orientation of the measuring axes of the transducers depends on the cable location, tracking and checkup tasks. Irrespective of the method of current excitation in the cable, secondary variable electromagnetic field is generated in the surrounding environment. The distribution of the instantaneous amplitude values of the electric and magnetic vectors of the field is shown in figure 1. The strength vector of the electric field has one component EX parallel to the cable axis, and the induction vector of the magnetic field has one component B\u03b8 tangent to circles that have cable axis as their centre. The physical basis of these methods is the inductive coupling of the generator and the cable line. The inductive coupling is used when it is necessary to check a specific area for cable lines or metal pipelines - for example before excavation works or when it is impossible to connect the generator to the cable line directly", " Due to the attenuation of the electromagnetic field of the emitting antenna in the conducting medium, the source of EMF appears on a limited cable section, and the current is transmitted incessantly. In real-life situations, the current along the cable will gradually shrink depending on the distance from the induction EMF excitation point due to the presence of the external conductor medium and imperfection of insulation. Since the effective cable length is far longer than the distance from the cable to the avoidance tool, the calculation model shown in figure 1 can be used to calculate the secondary magnetic and electric fields. Since formulae (1) and (2) contain scalar products of two vectors, the structure of magnetic and electric fields (fig. 1) provides for the maximum EMF value to be achieved when the magnetic momentum M of the magnetic dipole is orthogonal to the cable axis, if the current is induced in the cable by the magnetic dipole. If the direction of magnetic dipole momentum M is parallel to the cable axis, the EMF equals to zero. Hereof it follows that: 1) to identify the location of the cable, the dipole momentum of the emitting magnetic antenna must be directed either along the cable avoidance movement direction or vertically; 2) in order to use the cable route tracking mode, the dipole moment of the emitting magnetic antenna must be oriented transversely (to the cable avoidance movement direction) or vertically [7]: )( 2 1 0 rK I B c c \u03b3 \u03c0 \u03b3\u00b5 =\u03b8 ; )( 2 0 0 rK Ii E cx \u03b3 \u03c0 \u03c9\u00b5 \u2212= , (3) where cc i \u03c3\u03c9\u00b5=\u03b3 0 is the electric field medium propagation constant; \u03c3\u0441 is the medium specific conductance; K0, K1 are modified Bessel functions of the first kind, zero and first orders" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0002636_icma.2017.8015910-Figure4-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0002636_icma.2017.8015910-Figure4-1.png", "caption": "Fig. 4 Annular Plate Transversely Load 2 3 3", "texts": [ " Equivalent radius of curvature is ( ) 1 2 1 2Ei i iR \u03c7 \u2212= \u0398 (11) The deformation predicted by Hertz theory is 1 2 3 2*2 9 cos cos 16H Ei P F E R \u03b4 \u03b2 \u03b1 = (12) * 1 2 1 1 1 E E E = + (13) where P is the load, F2 is the correction factor of the displacement, E1 and E2 are the Young\u2019s modulus of the two contact objects[17]. The Thread Stiffness. The thread could be seen as an annular plate fixed on cylinder. The deflection of annular plate can be considered as the delfection of threads. Warren C. Young [18] showed a model to predict the deflection of an annular plate transversely loaded at an arbitrary point. The bending deflection sketch is shown as Fig. 4. The thread is assumed to be uniform thickness.The thickness of pitch diameter of thread is assumed as the thickness of thread. The deflection equation is written as: 1 2 3 3TB b b rb b r r r y rF M F Q F G D D D \u03b4 \u03b8 \u03c9= + + + \u2212 (14) For the condition of a thread contact, it can be be viewed as a plate with its outer edge free and inner edge fixed. So under this condition, there are some simplified assumption. ( ) 3 0 9 0 9 2 8 0, 0, 0, 0, 0, , , 12 1 b b a ra a rb b y y M Q r C ra E t M L Q D C b b \u03b8 \u03c9\u03c9 \u03bd = = = = = = \u2212 \u2212 = = \u2212 So the deflection equation of the thread can be simplified as : 2 3 3 0 9 0 9 2 3 3 8 TB r C ra r r r L F F G C b D b D D \u03c9\u03c9\u03b4 \u03c9 =\u2212 \u2212 + \u2212 (15) where \u03c9 is the load; r0 radial location of unit line loading or start of a distributed load; t is the thickness of the plate; a is the outer radius; b is the inner radius for annular plate; r is radial location of quantity being evaluated; C8,C9, L9 ,F2,F3,G3 are constants given in Appendix A" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0001298_eeeic.2011.5874589-Figure4-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0001298_eeeic.2011.5874589-Figure4-1.png", "caption": "Figure 4. New geometry of the induction machine", "texts": [ " SHORT-CIRCUIT FAULTS SIMULATION The components of electrical circuit are assigned to the magnetic circuit (Fig.3). Magnetization of the machine to the initial state is made by a resolution magneto-dynamic. The latter is independent of time. Once the machine has been magnetized, we launched the resolution magnetoevolutionary. The machine model previously presented does not allow considering this type of fault. In order to study the inter- turn short-circuit in the stator slot, we have brought changes in the original geometry and in the electrical system of the stator. The new geometry (Fig.4) of the induction machine is characterized by the division of each stator slot into two distinct regions. The first region is the side of airgap, the second is located at the bottom of the slot. There is no separation between the two magnetic regions of the notch. The new geometry of the machine remains the same as the real machine and does not change the electrical and magnetic characteristics. The advantage of the new geometry is that it allows to consider the short-circuits in the slot. This permit to choose the coil and the number of short-circuited turns In the stator circuit which has been modified (Fig" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0002302_b978-0-12-803581-8.10351-0-Figure45-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0002302_b978-0-12-803581-8.10351-0-Figure45-1.png", "caption": "Fig. 45 Backboard model with fuel cell access.", "texts": [ " This ergonomically contoured driver backboard also ensures that the driver\u2019s head will be more correctly positioned. In addition to lowering the drivers head this molded rear backboard moves the driver\u2019s shoulders forward 1.500, allowing adequate room for the driver\u2019s helmet in front of the required foam padding in the head restraint area, which ensures proper maneuverability of the driver\u2019s helmet to provide a full range of visibility. The preliminary design for the ergonomically contoured backboard can be seen in Fig. 45. From this figure it can be seen that the design for the backboard is curved both horizontally and vertically for the drivers back, but in fact it is actually molded somewhat oversize to enable custom seat inserts for various drivers. In addition to supporting the driver, the molded backboard provides an upper \u201cshelf\u201d as a location to mount the shoulder safety harness. The molded backboard is shown in the monocoque in Fig. 46. This design was extensively modeled to ensure structural integrity, especially in the safety harness mounting locations" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0002480_978-3-319-60399-5_7-Figure2-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0002480_978-3-319-60399-5_7-Figure2-1.png", "caption": "Fig. 2 Profiles of relieved surfaces of the hob teeth in the case of localized contact", "texts": [ " The authors found the relation between values of fmax and fmin determined by expressions (1) and (3) with tolerances fPt, fPx, fzz0, fhs for different degrees of accuracy, according to the Russian standard GOST 3675-81. It is stated that the required value of the reduced concavity f should correspond to the range fhs f 2:0fhs: \u00f04\u00de The reduced concavity of the hob-generating surface with respect to the worm thread profile is formed due to the concavity of the front surface of teeth obtained during sharpening (Fig. 1) and the concavity of the profile of the relieved tooth flanks of the hob with respect to the profile of the ground worm thread (Fig. 2). Figure 1 presents the front surface profile of the hob with the following parameters: m is the axial (or normal) module of the hob; rF is the radius of the pitch cylinder of the generating worm; ra is the radius of the outer cylinder of the active profile of the hob tooth; hm = 2 m is the height of the active profile of the hob tooth; fvi max is the maximum deviation of the generatrix from the radial straight line; hE = 2hm/3 is the distance from the boundary of the active segment of the profile to the point of maximum concavity of the generating line of the front surface. Figure 2a shows the profile of the relieved surface for which the helix angle kr is greater than the helix angle k for the thread of the generating surface (the right flank of the tooth of the right-hand hob). In this case, the distance from the internal cylinder of the active flank of the thread to the maximum value of the concavity is hE = 2hm/3. In Fig. 2b, the helix angle kr of the shank is less than the angle k (the left flank of the tooth of the right-hand hob), in this case, hEL 0.5774hm. The concavity of the profile of flanks for the hob teeth is provided by adjusting the set-up parameters for the radial-axial relief [11]. In this case, different adjusting methods are used for two flanks of the tooth, which is why coordinates of maximums for the concavity of profiles of the front and relieved surfaces do not coincide and are separated by distance: 0" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0000076_1.4003502-Figure1-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0000076_1.4003502-Figure1-1.png", "caption": "Fig. 1 Bolted joint model", "texts": [ " The interfacial pressure distribution between the clamped plates was also investigated. They proposed an analytical model for determining the variation in the joint clamp load and bolt tension under a separating tensile force. In this study, new closed form formulas are developed for the clamp load loss in a bolted joint under separating load that is placed at different distances from the bolt center. The effect of the eccentricity value of the tensile load is investigated. 2 Modeling of the Bolted Joint Figure 1 shows the bolted joint model used in this study. Prior to subjecting the joint to the separating force Fe, the bolt preload Fb and the joint clamp load Fc are both equal to some initial value Fi. The application of the separating force increases the fastener tension and reduces the clamp force, simultaneously. As long as APRIL 2011, Vol. 133 / 021206-111 by ASME hx?url=/data/journals/jpvtas/28543/ on 02/16/2017 Terms of Use: http://www.asme.org/about-asme/terms-of-use t a a e f v 3 t s b s f r o b t d w s F i 0 Downloaded Fr he separating force Fe is acting on the joint, the fastener tension nd the clamping force will not be equal", " 12 showed that the tensile stiffness Kc was larger than the compressive stiffness Kc of the joint used by Nassar and Matin 9\u201311 . In their model, they assumed that the external tensile force Fe was applied at the center of the bolt; hence, both the compressive and tensile stiffness of the joint, respectively, Kc and Kc were the same since the clamped plates will not bend. However, the model used in this study uses an external tensile force Fe that is not at the center of the bolt; hence, the clamped plates will bend, as illustrated in Fig. 1. The joint behavior under Fe is determined by an effective joint stiffness Kc that should take into account the bending of the clamped plate as well the classical compressive stiffness Kc of the clamped plates and the bolt stiffness Kb, as it is outlined in a later section. The bolt line ABCD is described by the function F= f x , which represents the relationship between the fastener tension F and the fastener elongation x in the elastic-plastic region. Point B represents the fastener elastic limit, and point C represents the initial tightening condition, where the fastener is elongated by the amount xi that creates an initial tension equal to Fi", " The clamp load loss ignificantly decreases as the eccentricity 2b increases; this means hat the corresponding increase in bolt tension is also significantly educed by increasing the eccentricity. Finally, this study shows hat the strain hardening rate of the fastener material plays a sigificant role in the amount of clamp load loss; it increases as the train hardening rate decreases. omenclature Fb fastener tension Fc clamping force Fi initial fastener tension preload Fe separating force b fastener elongation c joint compression 2b eccentricity of the tensile load, as shown in Fig. 1 Kb fastener stiffness Kc compressive joint stiffness Kc tensile joint stiffness when the load is applied off-center on the joint , it increases as the eccentricity 2b increases, and Kc =Kc when load is applied on the bolt axis Kplate stiffness of the plate Kbending bending stiffness load location factor that depends on the location of the separating load, i.e., the eccentricity ig. 17 Effect of eccentricity 2b on clamp load loss for various evels of preload 21206-8 / Vol. 133, APRIL 2011 om: http://pressurevesseltech" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0003700_iemdc.2019.8785080-Figure1-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0003700_iemdc.2019.8785080-Figure1-1.png", "caption": "Fig. 1. Flux focusing type DSSR AFPM.", "texts": [ " In section IV, the magnetization characteristics of the conventional stacked lamination and tapewound lamination are analyzed, and the influence of the physical dimensions, heat- treatment and gluing the lamination is investigated. In section V, the core loss measurements of the SMC stator, and glued and heat-treated (GHT) type tape-wound electrical steel stator are provided. The research work is concluded in section VI. II. FLUX FOCUSING TYPE DSSR AFPM The permanent magnets (PMs) are inserted in the rotor disc and the rotor is sandwiched between two identical slotted stators of the flux focusing type DSSR AFPM, as shown in Fig. 1. To avoid mechanical problems, the PMs are inserted in the solid electrical steel without lamination. The slotted stators are developed by milling the tape-wound electrical steel and the 978-1-5386-9350-6/19/$31.00 \u00a92019 IEEE 1061 SMC ring. The design specifications of the flux focusing type DSSR AFPM are given in Table I. The double-layer tooth coil windings are wound on the stator teeth, having reduced endwinding length. In the flux focusing type DSSR AFPM, the air-gap flux flows in the axial direction, therefore, a tape-wound lamination is required, unlike with the conventional stacked lamination employed in radial flux machines [3]", " Therefore, the iron losses increase, if \ud835\udc5f increases, and the magnetic field no-uniformity increases. The heat-treatment, as well as, gluing the lamination reduces the iron losses due to stress relief, therefore, NT tape-wound lamination has higher iron losses than the HT and GHT typewound lamination. The GHT type tape-wound lamination is preferred for the stator of the flux focusing type DSSR AFPM. The HT3 and GHT3 type tape-wound electrical steel have the same dimensions, as required by the flux focusing type DSSR AFPM prototype of Fig. 1, however, the SMC has reduced physical dimensions, as shown in Table II. As there is a significant influence of the physical dimensions on the iron losses, therefore, the comparison is not rational. However, it provides a fair understanding of the magnetization characteristics of the SMC and tape-wound steel ring. The magnetization characteristics of the HT3 and GHT3 tape-wound and SMC ring are shown in Fig. 8. At higher frequency, the iron losses increases rapidly in the tape-wound ring compared to the SMC ring, therefore, SMC is preferred for higher frequency applications", " Additionally, due to the rare-earth PMs placed in the rotor disc having flux focusing behavior, the materials with high permeability and flux saturation level are preferred. Therefore, for the prototype, the SMC material is not preferable, because, it has lower permeability and also higher iron losses than GHT and HT tape-wound electrical steel at a frequency of 150 Hz, as shown in Fig. 9. By increasing either the frequency or the flux density, losses in the HT3 tape-wound electrical steel increases rapidly compared to the GHT3 tapewound ring, as shown in Fig. 8(b) and 9. Therefore, for the flux focusing type DSSR AFPM of Fig. 1, the GHT3 tape-wound electrical steel is selected. V. CORE LOSS MEASUREMENT OF THE AFPM STATOR Principally, the toroid test setup is for the magnetization characteristics and core loss measurement of the material in an exact toroid shape. Due to the ring structure, the magnetic flux has circular close path and it is almost uniformly distributed. The voltage form factor is controlled. As the stator is slotted, therefore, the magnetic length will not be exactly circular, and the flexibility in the form factor of the voltage waveform is considered" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0000679_20120403-3-de-3010.00065-Figure1-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0000679_20120403-3-de-3010.00065-Figure1-1.png", "caption": "Fig. 1. GLMAV Operational regimes", "texts": [ " Likewise, we study the transition stage providing the modeling and control strategies. The dynamics models of the rotorcrafts are deduced using the Newton-Euler approach. In terms of control algorithms, we are using a hierarchical control law. The ultimate objective of the project is to launch a vehicle about 500 meters with a height of 100 meters, at this point the vehicle collects valuable visual information of the target zone through a pointing-downwards camera. To achieve this goal, let us consider two rotorcrafts MAVs, a GLMAV (Fig. 1) and the ALMAV (Fig. 2). Despite that some mechanical similarities are evident in the vehicles design, the actuators nature and distribution separates the flight control profile. 2 The term \u201dremote\u201d is employed to denote a zone that is beyond of the aerial range of a mini rotorcraft 2.1 GLMAV The GLMAV is a rotary-wing MAV based on a coaxial rotor system, that is to say a dual motor driving two counterrotating blades, of which only one is fully controlled by a swashplate. During the ballistic phase, (BP) the rotors are folded to fit in the projectile shell" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0000492_s10035-013-0439-3-Figure6-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0000492_s10035-013-0439-3-Figure6-1.png", "caption": "Fig. 6 Nadai\u2019s idealization of the orientations of major compression \u03c31 and minor compression \u03c33 along any circle traced around the ridge of a valley as the center where all \u03c31 converge toward the pole of \u03c31 situated at the bottom of the circle while all \u03c33 converge toward the pole of \u03c33 situated at the top of the circle", "texts": [ " The problem of a semi-infinite loose planar valley inclined at an angle of repose is illustrated in Fig. 5, where the z axis is chosen to pass through the central plane, symmetrically dividing the valley into left-hand and right-hand sides. In the same manner as in a wedge problem, the shear stress on plane 1 represents anti-clockwise rotation and shear stress on plane 2 represents clockwise rotation. The limiting state of stress along both planes can be visualized by superimposing these planes onto Mohr-Coulomb envelopes as shown in Fig. 2. According to Fig. 6, Nadai considered the following boundary conditions for the valley problem, where the major axis of compressive stress lies at a right angle to the direction of gravity or parallel to the x axis. Furthermore, the angle of major compressive stress along the surface with sand sliding down on the left-hand side with the same magnitude as the wedge problem but with a minus sign due to clockwise rotation. c = |\u03b8=\u03b8c =0 where \u03b8c = \u03c0/2 (56) f = |\u03b8=\u03b8 f =\u2212 (\u03c0/4+\u03c6/2) where \u03b8 f = \u2212\u03c6 (57) Nadai idealized the problem in the same manner as the wedge problem by treating an angle of the major principal stress as a linear function of an angular coordinate \u03b8 based on two extreme conditions given by Eqs" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0000986_j.proeng.2011.03.137-Figure13-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0000986_j.proeng.2011.03.137-Figure13-1.png", "caption": "Figure 13: The trajectory of the point E.", "texts": [], "surrounding_texts": [ "6. An example of mobility calculus of a mechanism with one passive element\nWe consider the mechanism shown in figure 10, where AB=DC, AB//DC, BC=AD, BC//AD.\nLet\u2019s suppose that all the elements are disjoined and become free in space, except one. Only one segmented frame remains fixed, the second becomes mobile (Figure 11). After segmentation and fictional motion in space (or in plane) the number of temporarily mobile elements becomes equal to the number of kinematic joints, i.e. four.", "In this phase, the number of degrees of freedom of the system is 6 4, for a spatial movement, or 3 4, for a planar movement.\nWe rejoin the elements by rotational joints of V class, including the temporarily segmented frame (Figure 12) and\nthe constraints of the joints are eliminated 5 4 and 2 4 respectively.\nThe extreme element (0) of the open chain has the spatiality three: Rx, Ty, Tz (whether for a spatial or a planar movement); in order to compose the mechanism (Figure 10), three degrees of freedom will be eliminated. If we calculate the mobility of the mechanism by using Eq. (5), for a spatial temporary movement, we obtain the mobility 1 (Eq. (8)) and for a planar temporary movement we obtain the same result (Eq. (9)). M=6m-5p-b1= 134546 (8) M=3m-2p-b1= 134243 (9) The mobility of the mechanism is one. A point E, situated in the middle of the element BC, describes a circular trajectory, with the dimension of the radius equal to the dimension of the element AB.\nWe join an element EF on the point E (Figure 14). We present two situations:", "The length of the element EF is equal to the length of the element AB and the element EF is parallel to the segment AB (Figure 14).\nThe length of the element EF is different from the length of the AB element (Figure15). For the situation shown in figure 14 we calculate the mobility of the mechanism. For that purpose, we cut twice the frame, so that the number of the joints becomes equal to the number of the\nelements, i.e. 6 (Figure 16).\nThe mobility number of the element 6 is b1=3 (Ty, Tz, Rx), like in the previous example. Let\u2019s calculate the mobility number of the element 5. With this aim in view, we check the possible independent\nmovements of the extreme element 5 relative to the frame 0 (Figure 16).\nThe point E is situated on the elements 2 and 4, at the same time. For this reason, in the case of an infinitesimal displacement Ty of the element 5, from the point F1 to the point F2, the final point E3 must be at the intersection of the circular trajectory T1 (the circle with the dimension of the radius equal to the dimension of the element AB, and with the centre of curvature in the point F1) with the circular trajectory T2 (the circle with the dimension of the" ] }, { "image_filename": "designv11_62_0002440_978-3-319-60867-9_14-Figure1-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0002440_978-3-319-60867-9_14-Figure1-1.png", "caption": "Fig. 1. Active Ankle prototype Fig. 2. ASPM architecture", "texts": [ "eywords: Parallel manipulator \u00b7 Kinematic analysis \u00b7 Direct kinematics \u00b7 Algebraic geometry A novel, almost-spherical parallel manipulator (ASPM) Active Ankle (Fig. 1) and its comparison with similar mechanisms like Agile Eye has recently been introduced in [5,6]. Due to its unique, simple and compact 3[R 2 [SS]] design (topological equivalent of Delta robot), the constraint of moving the endeffector about an exact center (of rotation) in case of spherical parallel manipulators (SPM) is relaxed to almost spherical motions that includes a shift of the end effector about a tolerated, very small domain. Due to the presence of a closed c\u00a9 Springer International Publishing AG 2018 S" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0002762_1.4038165-Figure3-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0002762_1.4038165-Figure3-1.png", "caption": "Fig. 3 A two-link manipulator robot system 11 1 12 2 2 1 1 2 2 1 1( + ) ,H q H q hq q h q q q g (15a) 2", "texts": [ " Finally, the block diagram MFASMTC based on a DOB is illustrated in Fig. 2. 4 Design of a MFASMTC based on a DOB for the MIMO mechanical system: A two-link manipulator robot system A two-link manipulator robot system with unknown time-varying disturbance demonstrates the robust control performance by the proposed MFASMTC based on the DOB. The uncertainties of the robot system including the moment of inertia for the links, Coriolis, centrifuge and gravity loading forces are considered in the dynamic model. 4.1 Dynamic model Figure 3 shows the configuration of the two-link manipulator robot system, and the motion of the robot system is in the vertical plane. 1q and 2q are the angular displacement of links 1 and 2, respectively. 1m and 2m are the mass of links 1 and 2, respectively. 1 and 2 are the control torque in joint 1 and 2, respectively. 1l and 2l are the length of links 1 and 2, respectively. 1cl and 2cl are the length from the joints to the center of mass of links 1 and 2, respectively. 1I and 2I are the moment of inertia for the links 1 and 2, respectively", " Acc ep te d Man us cr ip t N ot C op ye ite d Journal of Dynamic Systems, Measurement and Control. Received May 07, 2017; Accepted manuscript posted October 10, 2017. doi:10.1115/1.4038165 Copyright (c) 2017 by ASME Downloaded From: http://dynamicsystems.asmedigitalcollection.asme.org/ on 11/08/2017 Terms of Use: http://www.asme.org/about-asme/terms-of-use Page 1 of 31 Figure Captions Fig. 1 Mass-spring-damper system Fig. 2 Block diagram of the MFASMTC based on a DOB for the mass-spring-damper system Fig. 3 A two-link manipulator robot system Fig. 4 Block diagram of the MFASMTC based on a DOB for the two-link manipulator robot system Fig. 5 Responses comparisons of the MFSMTC by using the fixed control gains 10 and 35 ; (a) displacement 1x vs. time, (b) velocity 2x vs. time, (c) control input u vs. time and (d) sliding surface function s vs. time. Fig. 6 Responses comparisons of the MFSMTC among the fixed control gains 10, 35 and real control gain ; (a) comparisons between 10 and , (b) comparisons between 35 and ", " 2 Block diagram of the MFASMTC based on a DOB for the mass-spring-damper system Acc ep te d Man us cr ip t N ot C op ye di te d Journal of Dynamic Systems, Measurement and Control. Received May 07, 2017; Accepted manuscript posted October 10, 2017. doi:10.1115/1.4038165 Copyright (c) 2017 by ASME Downloaded From: http://dynamicsystems.asmedigitalcollection.asme.org/ on 11/08/2017 Terms of Use: http://www.asme.org/about-asme/terms-of-use Page 4 of 31 Y X1q 2q 1 2 1m g 2m g 1l 2l 1cl 2cl 1I 2I Joint 1 Joint 2 Fig. 3 A two-link manipulator robot system Acc ep te d Man us cr ip t N ot C op ye di te d Journal of Dynamic Systems, Measurement and Control. Received May 07, 2017; Accepted manuscript posted October 10, 2017. doi:10.1115/1.4038165 Copyright (c) 2017 by ASME Downloaded From: http://dynamicsystems.asmedigitalcollection.asme.org/ on 11/08/2017 Terms of Use: http://www.asme.org/about-asme/terms-of-use Page 5 of 31 0 0 \u02c6 \u02c6 \u02c6 \u02c6 Q A Q G U \u0394 KE 0 0 Q A Q G U \u0394 * * * 0 0 Q A Q G U * \u02c6U \u03a6\u03a8 0 \u02c6 \u02c6 E PE E\u0302 0K E\u0302 Q\u0302 Two-Link Manipulator Robot System Desired System Model *U 0\u0394\u0302 \u0394\u0302 \u0394\u0302 Q E\u0302 U *Q ,\u03b1 \u03b2 s Q + - + - - + Disturbances and states observer Fig" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0001150_aero.2013.6497416-Figure6-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0001150_aero.2013.6497416-Figure6-1.png", "caption": "Figure 6. 6-DoF frame of the Pelican quadrotor, with rotations, translations and thrust vectors.", "texts": [ " Problem Statement Hovering control consists in properly adjust the amount of thrust to each of the four motors of the MAV individually so that the quadrotor remains still at a desired position in the 3D space. It turns out that, the precise combination of perpendicular thrust forces, obtained by adjusting the spin rate of each propellers (\u21261..4) acting over the quadrotor, produces a disturbing torque, which is responsible for the 6-DoF movement of such aircraft. The Pelican MAV is composed by two pairs of rotating propellers (M1, M3) and (M2, M4), which spin in opposite senses, as one may see in Fig. 6. This is a very well established and intentional resource that allows torque cancellation, which makes the quadrotor configuration relatively simple and also possible. Still in the figure, attitude angles (\u03c6, \u03b8, \u03c8) as well as translations (x, y, z) are achieved by a combination of varying speeds of rotation of each propellers. The complete Newton-Euler based model that describes the Pelican aircraft is discussed in the next section. MAV Dynamics The availability of the states of the aircraft discussed in Section 2 - thanks to the emulation of a complete IMU (Inertial Measurement Unity), altimeters and gyroscopes - makes the closed loop control possible" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0003356_978-981-13-5799-2_12-Figure12.3-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0003356_978-981-13-5799-2_12-Figure12.3-1.png", "caption": "Fig. 12.3 Contributions of shear stress in the adhesion region and the sliding force in the sliding region to the total braking force (reproduced from Ref. [2] with the permission of JSAE)", "texts": [ " We here define the critical clip ratio scritical as that when the whole contact area becomes the sliding region. Substituting lh = 0 into Eq. (12.8), scritical is obtained as scritical \u00bc 3lsFz=CFs: \u00f012:17\u00de speak decreases with increasing X = ls/ld. In the particular case that ls/ld = 1, speak = scritical is satisfied. When the whole contact area becomes the sliding region (lh = 0), the term of the adhesion region in Eq. (12.13) is zero. The braking force Fx is then given by Fx \u00bc ldFz: \u00f012:18\u00de The right figure of Fig. 12.3 shows the contribution of forces in the adhesion and sliding regions to the total braking force calculated using Eq. (12.16). The vertical axis of the figure is normalized by the load. There are both adhesion and sliding regions in the contact patch in the case that 0 s < scritical while there is only the sliding region in the contact patch in the case that scritical s 1. The force in the adhesion region is dominant at a small slip ratio while the force in the sliding region is dominant at a large slip ratio. Note that Fig. 12.3 is similar to Fig. 11.7. 2Note 12.2. 3See Footnote 2. (3) Comparison between experiment and calculation Yamazaki et al. [3] compared the calculation made using Eq. (12.15) with experimental results. The tire size was 185/70R13, the inflation pressure was 190 kPa, and the drum tester was covered with a safety walk (#600). Parameters used in the calculations were Cx = 0.133 MPa/mm, b = 101 mm, l = 129 mm, l0d = 1.3, ld = 1.0 at V 0 = 30 km/h and Fz = 1.96, 2.94, 3.92, 4.9 kN. Figure 12.4 shows that the calculation is in good agreement with the experimental results" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0002469_cnsa.2017.7973934-Figure1-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0002469_cnsa.2017.7973934-Figure1-1.png", "caption": "Fig. 1. Basic parameters of the quadrocopter", "texts": [ " Understanding these recommendations and dependencies, we use the PID controller parameters to fine tune the control system. If we need a smooth picture from the camera, then one PID settings are used, if it is necessary to maneuver sharply in the limited space and to resist external disturbances, then others. The modeling results are confirmed by numerical experiments and correspond to previous researches of authors and other scientists [9]\u2013[17]. The basic parameters of the quadrocopter are shown in Fig. 1. Here F1, F2, F3, F4 are the engines forces; FT is the gravitational force; the point M is the center of mass of the apparatus (it coincides with the geometric center); l is the distance between point M and the center of the screw; \u03d5, \u03b8, \u03c8 are the angles of bank, pitch, and yaw respectively; \u03a91, \u03a92, \u03a93, \u03a94 are the angle rates of screws. The equation of the PID controller output signal has the form [4]: u(t)=P + I +D=Kpe(t) +Ki \u222b t 0 e(\u03c4)d\u03c4 +Kd de dt , (1) where Kp, Ki, Kd are the amplification factors of the proportional, integrating, and differentiating components of the regulator" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0002416_j.crme.2017.05.014-Figure1-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0002416_j.crme.2017.05.014-Figure1-1.png", "caption": "Fig. 1. The rail vehicle model.", "texts": [ " Then, the authors show that the determinist optimal solutions can be seriously altered by DP uncertainties. In Section 4, a novel algorithm is developed and used in multi-objective robust optimization. The results are discussed and compared with literature results. Finally, some concluding remarks are presented in section 5. The RV system is made of a rigid car body C , bogies Ck and wheelsets Ski . The connection between these components is represented by the secondary and the primary suspensions. Each suspension is formed by a system of linear springs and dampers, which work in three directions (Fig. 1) [1,11]. The longitudinal symmetry of the RV system leads to the decoupling of lateral, vertical and longitudinal motions [1,11]. In this paper, we focus on the lateral dynamic behavior of the RV system. To simplify the analysis without reducing the accuracy of the model, we will consider only a quarter model of the RV (Fig. 1) [1,11]. Hence, the RV system has only eight degrees of freedom represented by the generalized coordinate vector q: q = [ y\u0304, \u03b1\u0304, y1,\u03b11, y11,\u03b111, y12,\u03b112]T (1) In this study, the rail is assumed to be smooth and rigid. Moreover, due to the fact that the rail curve radius and the RV speed are constant, damping forces were found not to be important, compared to the elastic ones [1,11]. We may find the dynamic model of the RV system by applying the Lagrange method: d dt ( \u2202L \u2202q\u0307i ) \u2212 \u2202L \u2202qi = Q i (2) L is the Lagrangian function and Q i represents the generalized forces applied to the system" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0001286_iros.2011.6094587-Figure4-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0001286_iros.2011.6094587-Figure4-1.png", "caption": "Fig. 4. Representation of the geometrical invariant point Oh and implementation of invariants identifications with markers pij . The invariant Oh is traced as the intersection of axes ni resulting from the interpolation of circular trajectories centered in Ci and obtained moving the markers \u2200j poses along \u2200i trajectories. Due to the choice of poses in the algorithm, t{r}|i \u2261 t{r} \u2200i.", "texts": [ " Rather, it reverses the geometrical meaning of the hand-eye problem, which is to find the projection of the same unknown rototranslations between the two {c} and {r} coordinate frames. Such projections, instead, can be derived when notable geometries are fiducially constant between the frames. Then specific geometrical projections are purposefully searched instead of being the result of random poses. The main idea is therefore to constrain all the generated robot poses to trace constant manifolds and points. In this way, the coordinate frame {c} and {r} display a fixed rototranslational relationship (represented by Z) along the entire procedure (see Fig. 4). X is derived as the only remaining unknown in the kinematical relationship of moving frames. The sampling noise is therefore bounded to the estimation of the manifolds. In particular, the estimation of a single manifold (e.g. a circle) employ a relatively large number of poses. In this way, any noise in samples propagates throughout the computational process to a limited extent, improving the overall accuracy of calibration. The key procedural issue of the method is to generate abstract entities (the manifolds), traced by robot trajectories, that can be used to identify a constant point in both {c} and {r}, as Oh. For simplicity, Oh is w.l.o.g. purposefully made overlap the robot shoulder: the constant point is therefore known (i.e. programmed) {r} and to be interpolated in {c} as the abstract intersection of the axes of the generated manifolds. In Fig. 4, the manifolds are circle whose normal axes ni are estimated and intersect in Oh according to {c} coordinates. Since ni are known in {r} and interpolated in {c}, their rotational relationship is ni{c} = RZni{r} (2) and is valid for all the manifolds. Therefore, concatenating all the N manifolds, two 3\u00d7N matrices5 of manifold axes can be populated in {c} and {r} Uc = [n1{c}, . . . ,nN{c}] Ur = [n1{r}, . . . ,nN{r}] and provide the rotational part of Z: Uc = RZUr \u2192 RZ = UcU + r (3) where + denotes the Moore-Penrose pseudoinverse" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0002200_978-3-319-51222-8-Figure2.50-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0002200_978-3-319-51222-8-Figure2.50-1.png", "caption": "Fig. 2.50 Two-way grid", "texts": [ "124) and so we find an 3 in the end deflection produced by a point load P w( ) = P 3 3EI , (2.125) and if the beam carries a single moment M = \u2212EI w\u2032\u2032, then we see an 2 w( ) = M 2 2EI . (2.126) 2.14 Influence Functions Integrate 135 In the case of a distributed load p, the end deflection is w( ) = p 4 8EI (2.127) where the 4 matches the fourth order of the differential equation EI wIV = p. We find the l3 of (2.125) also in the equation Pa Pb = l3b l3a . (2.128) which splits a point load P = Pa + Pb on a two-way grid structure (see Fig. 2.50) into two parts. A stiffness matrix, K u = f , instead differentiates, and so we find the \u2018inverse factor\u2019 EI/l3 up front in a beam matrix and EA/l in a bar matrix. The governing equation of second-order beam theory is EI wIV (x) + P w\u2032\u2032(x) = p(x) (2.129) whereP is the compressive force in the beam and p(x) the lateral load (see Fig. 2.51a). This is a linear self-adjoint fourth-order differential equation with constant coefficients, but the problem in this case is that the coefficient P is load case dependent and so also the shape of the influence functions depends on P" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0000986_j.proeng.2011.03.137-Figure14-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0000986_j.proeng.2011.03.137-Figure14-1.png", "caption": "Figure 14 The mechanism with AB equal to EF.", "texts": [ " The extreme element (0) of the open chain has the spatiality three: Rx, Ty, Tz (whether for a spatial or a planar movement); in order to compose the mechanism (Figure 10), three degrees of freedom will be eliminated. If we calculate the mobility of the mechanism by using Eq. (5), for a spatial temporary movement, we obtain the mobility 1 (Eq. (8)) and for a planar temporary movement we obtain the same result (Eq. (9)). M=6m-5p-b1= 134546 (8) M=3m-2p-b1= 134243 (9) The mobility of the mechanism is one. A point E, situated in the middle of the element BC, describes a circular trajectory, with the dimension of the radius equal to the dimension of the element AB. We join an element EF on the point E (Figure 14). We present two situations: The length of the element EF is equal to the length of the element AB and the element EF is parallel to the segment AB (Figure 14). The length of the element EF is different from the length of the AB element (Figure15). For the situation shown in figure 14 we calculate the mobility of the mechanism. For that purpose, we cut twice the frame, so that the number of the joints becomes equal to the number of the elements, i.e. 6 (Figure 16). The mobility number of the element 6 is b1=3 (Ty, Tz, Rx), like in the previous example. Let\u2019s calculate the mobility number of the element 5. With this aim in view, we check the possible independent movements of the extreme element 5 relative to the frame 0 (Figure 16). The point E is situated on the elements 2 and 4, at the same time" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0002642_978-3-319-66866-6_2-Figure7-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0002642_978-3-319-66866-6_2-Figure7-1.png", "caption": "Fig. 7. Stress distribution while dimensioning (a) Adapted cross-sectional area (b) Stressoptimized model", "texts": [ " Starting from the concept model with a constant cross-sectional area (A:A), various preliminary designs are examined. As a result, a hollow profile with 45\u00b0-surfaces is provided in the outer areas. The less stressed area is substituted by a thin layer without cavities. Due to this modification, material can be saved and manufacturability can be improved by reducing overhangs. After adapting both, the outer shape and the cross-section area, the stress distribution of the preliminary design can be improved, as depicted in Fig. 7-a. However, the maximum stress is still critical. Against this background, the areas near the bearing has to be optimized. After evaluating different strategies for the material distribution, an adapted \u201ccore\u201d is built up, shown in Fig. 7-b. Holes are used for maximum weight reduction. The optimization of the preliminary design results in a homogeneous stress distribution with maximum stresses (r 240 N/mm2) below the allowable values. The preliminary design is finally evaluated in comparison to the design guidelines. Therefore, the orientation and position of the model in the process chamber has to be defined. After weighing the criteria for production time, accuracy, loading capacity due to anisotropy, post-process effort and avoiding damage by the coater, the orientation and position is set as depicted in Fig" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0001590_6.2013-1503-Figure7-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0001590_6.2013-1503-Figure7-1.png", "caption": "Figure 7. Student Preliminary Design Aerodynamic Analysis", "texts": [ " Again, the familiarity of Google Hangout and Yammer (since it\u2019s very much akin to Facebook) to the students aided immensely in this regard. After analysis of three concepts \u2013 a 214 ft span raked wingtip, a 214 ft span blended winglet, and a 232 ft span folding wingtip \u2013 the blended wingtip emerged as the front runner. Whereas the raked wingtip lacked the performance to meet the requirements, the folding wingtip was discarded as the students lacked the tools to effectively analyze the complications arising from the folding mechanism. Figure 7 shows some of the aerodynamic analysis carried out by the students. Reintegrating the level 1 analysis results into FLOPS for performance predictions, the students\u2019 concept offered a nominal performance increase of 25% less fuel per seat (or per pound of payload) compared to the 777-200ER at a unit cost of $283 million. D ow nl oa de d by R O K E T SA N M IS SL E S IN C . o n Fe br ua ry 4 , 2 01 5 | h ttp :// ar c. ai aa .o rg | D O I: 1 0. 25 14 /6 .2 01 3- 15 03 American Institute of Aeronautics and Astronautics 9 Overall, RFP 2 met three distinct learning objectives: 1) Students learned how a complex system decomposes into subsystems and constituent components for preliminary and detailed design analysis and evaluation successive applications of functional deployments" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0001248_epe.2013.6634388-Figure10-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0001248_epe.2013.6634388-Figure10-1.png", "caption": "Fig. 10: Flux density distribution around a pole piece. Fig. 11: Relationship between stack length of pole pieces and transmission torque.", "texts": [ " It is understood that the magnetic gear has essentially overload protection function which mechanical gear does not have. The measured maximum torque of 9.40 N\u2022m has the difference with the one of 12.5 N\u2022m obtained from 3D\u2013FEA show in Fig. 3. The reason for the difference between the calculated and measured torques is the difference of the stack length of the pole pieces. The stack length of the pole pieces of the prototype magnetic gear is 16 mm in order to be easily fixed them, while the ones of the inner and outer rotors are 10 mm. Fig. 10 shows the flux density distribution on the r\u2013z plane around a pole piece. The figure reveals that the flux leaks to air around the overhang part of the pole piece. Fig. 11 shows the stack length of the pole pieces versus the maximum transmission torque of the magnetic gear when stack lengths of the inner and the outer rotors are 10 mm. It is clear that the torque of the magnetic gear is maximized when a stack length of the pole pieces is 10 mm as the same stack length of both rotors. The measured maximum torque of 9", " 13 shows the mechanical input of the prototype magnetic gear at no\u2013load. It is understood that the no\u2013load input obtained from the experiment is significantly larger than the one obtained from 3D\u2013FEA, and that the no\u2013load input increases nonlinearly as the inner rotor speed increases. The main cause is the eddy current loss in the stator housing made from an aluminum alloy to hold the pole pieces. Fig. 14 shows the expanded view of the prototype gear around a pole piece. The pole pieces are held by the aluminum alloy. Therefore, as shown in Fig. 10, the leakage flux passes through the aluminum, and then the eddy current is induced in it. The eddy current loss of the aluminum of the prototype magnetic gear is estimated by 3D\u2013 FEA. The electrical conductivity of the aluminum is 3.76\u00d7107 S/m. Fig. 15 shows the eddy current density of the aluminum on the r\u2013\u03b8 plane. The figure reveals that the eddy currents are induced around the pole pieces. Therefore, it is necessary to replace the aluminum with non\u2013magnetic and non\u2013 conducting material. Fig" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0001321_icma.2012.6283381-Figure3-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0001321_icma.2012.6283381-Figure3-1.png", "caption": "Fig. 3 The model of the digital hyperextension in the IP. a) Kendall strip tape peeling model[18]. b) Phalanx II peels while Phalanx I stays still. c) Phalanx I peels after Phalanx II stops.", "texts": [ " A tape of elastic material as tissues of the digit was hollowed to hold the linkbars inside, part of which was adhesive on the plantar surface from the distalmost point to the middle of Phalanx I. The bars and the flexible material were stuck together as the bones and muscles in biological tissues, meaning there\u2019s no longitudinal sliding between them. B. Viscoelastic Analysis of the Inherent Motion Pattern The dynamic analysis was based on Kendall\u2019s Theory [18], in which the adhesive pad was treated as a strip tape. Fig.3 shows the peeling process by hyperextension in the IP. The force applied to the digit was taken as forces in x-axis and y-axis acting on Joint I. The longitudinal peeling force of the pad can be estimated as ( ) + (1 cos ) = 0 (1) where is the longitudinal peeling force in the pad, b is the width of the pad, E is the Young\u2019s modulus of the tissue, is the peeling angle, R is the adhesive energy between the pad and the surface. R is a velocity-dependent parameter. Here we assumed that R remained constant during the process and peeling angle was equal to the angular displament of the joints, provided that Joint II rotated first and Joint I rotated right after II stopped" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0001687_978-90-481-9689-0_68-Figure1-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0001687_978-90-481-9689-0_68-Figure1-1.png", "caption": "Fig. 1 The 3-UPU TPM.", "texts": [ "ey words: parallel manipulators, geometry, singularity, leg collision avoidance Great attention has been devoted to parallel manipulators (PMs) for their complementary characteristics with respect to the serial ones. The topology of parallel manipulators features a fixed base connected to a moveable platform by a number of serial chains (leg). Since many applications do not necessarily need six degrees of freedom (DOFs), manipulators with less than six DOF are also very interesting, in particular three-DOF PMs which provide the platform with a pure translation, a pure rotation or a mixed of translation and rotation have been studied in the literature [2, 3, 6]. The 3-UPU parallel manipulator [5] (Figure 1) has been extensively studied both as a three-DoF pure translational and a pure rotational manipulator (U and P are for Universal and Prismatic joint respectively). However, the influence the manipulator geometric parameters have on the manipulator performances still deserves attention. Indeed, new geometries can be devised which provide mechanisms with appealing kinematic and static features. This paper will focus on the 3-UPU pure translational manipulator, hereafter called 3-UPU TPM. The influence of both the directions of the base/platform revolute axes and the leg position is further investigated and some new geometries of the mechanism are presented, which exhibit interesting performances", " Parenti-Castelli Moreover, for the cases where the three legs of the manipulator intersect, some design solutions are proposed for the leg collision avoidance. The paper is organized as follows. In Section 2, two known 3-UPU TPM geometries are recalled to show the main properties of the 3-UPU TPM. Section 3 presents four new geometries that show the influence of the directions of base/platform revolute axes and of the leg location on the singularity loci of the manipulator. Section 4 presents three design solutions that allow the leg collision avoidance. Finally some conclusions are reported. A schematic of a 3-UPU parallel manipulator is shown in Figure 1. It features a translational platform connected to a fixed base by three extensible legs of type UPU. The universal pair U comprises two revolute pairs with intersecting and perpendicular axes, centred at point Bi, i = 1,2,3 in the base and at point Ai, i = 1,2,3 in the platform. In order to prevent possible rotations of the platform, two conditions must be fulfilled for each leg [1, 5]: \u2022 the axes of the two intermediate revolute pairs are parallel to each other; \u2022 the axes of the two ending revolute pairs are parallel to each other" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0000731_iros.2010.5651815-Figure2-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0000731_iros.2010.5651815-Figure2-1.png", "caption": "Fig. 2. Controlling humanoid robot with virtual spring-damper elements: (a) the virtual spring-damper elements on the upper body of humanoid and human, (b) the detailed configuration of hand", "texts": [ " VIRTUAL SPRING-DAMPER CONTROL The virtual spring-damper has been used widely as a motion control framework in the field of robotics [8], [14]- [16]. The virtual components create virtual forces when the virtual components interact with a robot system. This strategy is conceptually compact and requires relatively small amount of computation. We have implemented this method by applying ten virtual spring-damper elements on the upper body of the humanoid robot to mimic the human-like motion with real-time conversion (see Fig. 2). The movement of the robot is calculated by following dynamic equations of the motion: M(q)q\u0308 +N(q, q\u0307) +G(q) = \u03c4 + JTFex (1) where q is the joint angle vector, M(q) is the inertia matrix, N(q, q\u0307) and G(q) represent the Coriolis and Centrifugal force and gravity term, respectively. \u03c4 is the joint torque vector, J is the Jacobian matrix and Fex is virtual external force. We divided the virtual spring-damper controller by the form of two terms as{ Fex = k\u2206p\u2212 \u03b6 \u221a kp\u0307c \u03c4 = G\u0302(q)\u2212 C0q\u0307 + \u03c4v (2) where k denotes the spring stiffness coefficient, \u2206p is a differences between the target and current position vector, \u03b6 \u221a k is the damping coefficient and p\u0307c is the current velocity vector from eq. (2). G\u0302(q) is the compensated gravitation term which is ideally equal to G(q) in eq. (1). C0 and \u03c4v is the damping coefficient matrix and limit reaction torque which will be dealt in the next subsection. A. Virtual Force of Spring-Damper We suggest the virtual external force of spring-damper in eq. (2) corresponds to the recursive forward dynamics algorithm equations. As shown in fig. 2, we applied translational force on three points denoted as thumb, middle and pinky for the avoidance of representation singularity of rotational matrix. Likewise, for the motion expression of the waist rotation and for the human-like movements, two other translation forces are applied on both shoulder and elbow position; therefore, total ten virtual external forces are applied to control the upper body of the humanoid robot by the following form JTFex = 10\u2211 j=1 Ad\u2217Ti,j Fj +Ad\u2217TF T FFT (3) Fj = k(pd,j \u2212 pc,j)\u2212 \u03b6 \u221a kp\u0307c,j (4) where j could be the frame of thumb, middle, pinky, elbow and shoulder, Ti,j \u2208 SE(3) expresses frame j with respect to i which is the reference frame where the force applied (i could be the frame of hand, elbow and shoulder) and pd,j and pc,j are target and current position vectors of the frame j, respectively" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0003623_1.4044296-Figure1-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0003623_1.4044296-Figure1-1.png", "caption": "Fig. 1 (a) Section view of the ball bearing turbocharger computer-aided design (CAD) model and (b) turbocharger test rig labeled to show locations of components and connections", "texts": [ " Floating ring bearing TCs use two radial floating ring bearings and a bi-directional thrust bearing to support the shaft. The floating ring bearing consists of a journal bearing within another journal bearing or \u201cfloating ring.\u201d In some configurations, the \u201cfloating ring\u201d is prevented from rotating by an anti-rotation (AR) pin; this creates a squeeze film damper [1,10]. The thrust bearing tends to be failure-prone, especially, if exposed to frequent rotordynamic instability or unstable flow conditions through the compressor [9]. It is also responsible for a large portion of low-speed TC friction [5]. Figure 1(a) illustrates a section view of a ball bearing TC that contains a single bearing cartridge. The ball bearing cartridge consists of an opposed pair of angular contact ball bearings (ACBBs) which can support both radial and thrust loads [4]. Ball bearing TCs have been shown to reduce TC lag and friction losses compared with floating ring bearing TCs [4,11]. ACBBs operating at high speeds experience high centrifugal loads and increased heating due to pivoting or spinning on the non-controlling race [12,13]", " The TC was minimally modified to accept two axial load sensors at either end of the cartridge. Then, the AR pin was replaced with a load sensor to measure the force required to restrain the cartridge from rotating. All three sensors were necessary to fully characterize the loads applied to the bearing cartridge and understand how bearing loads relate to shaft whirl. The TTR developed for this investigation is a cold-gas TC test stand. The TTR was instrumented to measure shaft whirl as well as thrust and frictional loads on the TC under various operating conditions. Figure 1(b) depicts the TC integrated into the TTR. The TC bearing cartridge followed the standard configuration of two opposed ACBBs with steel races and ceramic balls. Approximate bearing dimensions are given in Table 1. 111101-2 / Vol. 141, NOVEMBER 2019 Transactions of the ASME Downloaded From: https://tribology.asmedigitalcollection.asme.org on 08/04/2019 Terms of Use: http://www.asme.org/about-asme/terms-of-use General Setup. In a TC, the turbine, center, and compressor housings are each separate parts to allow for assembly and installation", " A Jaquet Apollo speed probe, calibrated for titanium, was also installed on the compressor side of TC and was used to measure the shaft speed using the compressor wheel blades. Axial Load Sensors. In this ball bearing TC design, there are both axial and radial clearances around the outer race to allow for the proper functioning of the squeeze film damper. This allows the outer race to make contact with either the turbine side or the backplate on the compressor side of the housing (indicated in Fig. 1(a)) depending on the direction of the axial load. Thus, the measurement of axial load required a sensor on each side of the bearing. Compressor Side Sensor. In order to measure the thrust loads on the compressor side, a pocket was machined into the backplate to accept an instrumented beam insert. Figure 2(a) illustrates a CAD model of the modified backplate and insert. The insert was composed of three triangular cantilever beams oriented at 120 deg angles from each other. Figure 2(b) depicts how each beam on the insert was instrumented with a strain gage to measure load" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0000047_s13272-013-0095-7-Figure10-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0000047_s13272-013-0095-7-Figure10-1.png", "caption": "Fig. 10 Optical marker", "texts": [ " iMAR\u2019s advanced so called \u2018\u2018No Aiding for Attitude Algorithm\u2019\u2019 (NoA2 algorithm) is implemented, which has been initially developed for applications where no redundant aiding information from an external source (such as a GNSS) is available, and provides good attitude performance both under static and dynamic conditions. This algorithm is based on an Extended Kalman Filter (EKF), which internally uses very general motion constraints [38]. iMAR\u2019s vertical reference unit (VRU) provides the helicopter\u2019s accelerations and angular rates as well as the helicopter\u2019s attitude and heading. Attitude accuracy is\\1 with velocity aiding. The qualification of the system with respect to vibration, shock and EMC/EMI according to DO-160E is in progress. The optical marker (see Fig. 10), also developed by iMAR Navigation GmbH, is used to mark the load position. Energy is delivered by accumulators continuously for one complete flight mission. In order to save energy, the power is switched off when it has a physical contact with the hoist (the latter not being operated). Moreover, triggering of the optical marker is possible and currently under development, yielding a reduced power consumption. A special mechanism was designed to fix the optical marker over the bumper of the cargo hook on the cable" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0003656_wamicon.2019.8765445-Figure1-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0003656_wamicon.2019.8765445-Figure1-1.png", "caption": "Fig. 1. A potential application of suspended interconnects - isometric view of the additively manufactured interconnects over a chip placed in a cavity.", "texts": [ " The method uses a combination of additive manufacturing processes and laser machining to fabricate the proposed interconnects. Specifically, 3D printed flexible and lowloss suspended CPWs are fabricated utilizing fused 978-1-5386-95975/19/$31.00 \u00a92019 IEEE deposition modeling (FDM), and micro-dispensing technologies along with laser machining. These suspended interconnects could enable direct integration of multiple active or passive device chips with low coupling in SoP implementations. As illustrated in Fig. 1, using the proposed suspended CPW scheme, multiple interconnects could be fabricated in a vertical configuration above device chips that may be embedded in a 3D printed substrate. It is also possible to make direct connections to embedded devices. The following sections present the design, fabrication and preliminary characterization data for the suspended CPWs. The measured attenuation of the lines is approximately 0.26 dB/mm at 30 GHz. To the best of the authors\u2019 knowledge, this is the first demonstration of an mm-wave, suspended transmission line that is fabricated using an additive manufacturing approach" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0002865_we.2149-Figure12-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0002865_we.2149-Figure12-1.png", "caption": "FIGURE 12 Stresses in (A) base design and (B) optimized design (consider different scaling of the legends)", "texts": [ " Since buckling, which is strongly dependent on the thickness of the springs, was found to be a critical factor, the parameters defined first were the length and the number of springs, which have less influence on the buckling effects. The spring thickness could then be iteratively decreased to find a low value while still avoiding buckling. Although longer springs relieve the hinge, the overall size of the cage hinge should be fairly small. The length was therefore set to 2 m. With a chosen number of 28 springs, it was possible to achieve a thickness of about 17.5 mm without any buckling effects. Figure 12A,B shows the maximum von Mises stresses of the initial design (left) and the new design (right). The stresses are significantly reduced from about 720 to about 390 MPa while the required torque is still at a realizable level of about 500 kNm. To summarize, the results of the dimensioning process demonstrate that the elastic angular range of \u00b1 3\u25e6 , which was defined in Section 2, can be achieved with the chosen flexure topology. With our dimensioning method, we were able to adapt an initial design to meet our specific requirements, although the asymmetric stresses resulting from the high root bending moment did complicate the design process" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0002678_chicc.2017.8028765-Figure1-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0002678_chicc.2017.8028765-Figure1-1.png", "caption": "Fig. 1: The double pendulum system", "texts": [ " We begin our investigation with a well-known example in classical mechanics. The double pendulum is a simple physical system that exhibits rich dynamic behavior with a strong sensitivity to initial conditions. [9] The motion of a double pendulum is governed by a set of coupled second order ordinary differential equations and is chaotic. By Theorem 1 this system can be controlled through a decentralized PID controller under some conditions without knowing explicit system functions. To explicate this we consider the double pendulum system shown in Fig.1 where l1, l2 denote the length of the rods and m1,m2 denote the mass of the two bobs respectively. Let \u03b81, \u03b82 denote the angle subtended by the rods and the vertical axis. We assume that the rods are rigid and the double pendulum is free to swing in the vertical plane. Obviously, the configuration space of this system is S1 \u00d7 S1, which is not homeomorphic to R2. [10] We can write the equation of motion in the tangential di- rection as\u23a7\u23aa\u23aa\u23a8 \u23aa\u23aa\u23a9 \u03b8\u03071 = v1 \u03b8\u03072 = v2 m1l1v\u03071 = f1(\u03b81, \u03b82, v1, v2) +m1l1u1 m2l2v\u03072 = f2(\u03b81, \u03b82, v1, v2) +m2l2u2, (19) where mj ljuj(j = 1, 2) is the control input acting on the bob j and the force is in the tangential direction" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0000207_12.843999-Figure3-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0000207_12.843999-Figure3-1.png", "caption": "Figure 3: 3-D mapping catheter model initialization from two views. Figure 3(a): This general case shows two possible (dual) solutions when reconstructing a 3-D ellipse from bi-plane 2-D ellipses. The correct solution can be found by using prior knowledge, e.g., of the diameter of the circumferential mapping catheter. Figure 3(b): This special case reconstructs a 3-D ellipse from one 2-D ellipse in one X-ray view and a line in the other.", "texts": [ " The projections of the CMC into 2-D (X-Ray) images are approximated as 2-D ellipses. Proc. of SPIE Vol. 7625 762507-2 Downloaded From: http://proceedings.spiedigitallibrary.org/ on 06/05/2015 Terms of Use: http://spiedl.org/terms For the last assumption, we also accept the special case of a 2-D line as an ellipse with one half-axis being 0. The case arises under some special C-arm viewing directions, as shown in Figure 5. For the model generation, we differentiate between the regular case and the special case as shown in Figure 3. Note that we do not consider the case that the 3-D mapping catheter becomes a line in both X-ray projections as this is a very undesirable case in a clinical environment and can be omitted during clinical procedures, by adjusting one of the C-arms. For the reconstruction it is essential to know the 2-D ellipse parameters. To get the projection of the circumferential mapping catheter on the imaging plane, the catheter is first extracted by manual clicking followed by fast marching in one frame of the fluoroscopy sequence, as explained in [15]", " Given two projection images of an ellipse in 3-D on both A-plane view and B-plane view of a bi-plane system, respectively, we get two 3-D cones denoted QA for imaging plane A and QB for imaging plane B. The representations of QA and QB are quadrics of rank 3. The reconstructed ellipse is then computed by calculating a \u03bb such that the quadric statisfies [19] rank (QA + \u03bb \u00b7 QB) = 2. (4) A quadric of rank 2 describes two intersecting planes, which are extracted by an eigenvalue and eigenvector analysis. Each of these planes contains a valid ellipse in 3-D that projects onto the ellipses in the imaging planes, see Figure 3. As we require only one solution for our tracking approach, we utilize our prior knowledge about the anatomy of the pulmonary veins and select the result that is more circular, because the CMC inserted into a PV resembles a circle more closely than an ellipse in normal human anatomy. The circularity is determined by: \u03ba = |\u03c6\u2212 \u03c8| (5) with the axes \u03c6 and \u03c8 of an ellipse. To obtain the more circular solution, the ellipse with the smaller value for \u03ba is used. For the special case where the circumferential mapping catheter is projected close to being a line in one view, the method in [19] is not stable, as this method requires for QA and QB to be of rank 3" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0003416_j.automatica.2019.04.002-Figure1-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0003416_j.automatica.2019.04.002-Figure1-1.png", "caption": "Fig. 1. Guaranteed basin of attraction of the continuous-time controller (10) (cyan line) and of its hybrid implementation (magenta line). Trajectory-based estimate of the basin for the latter controller (yellow region) and trajectories of the system (9)\u2013(10). (For interpretation of the references to color in this figure legend, the reader is referred to the web version of this article.)", "texts": [ " Since cos(\u00b1\u03c0/2) = 0 and, correspondingly, x\u0307|x=\u00b1\u03c0/2 = \u00b1\u03c0/2 for any u \u2208 R, the largest achievable basin of attraction for the zero-equilibrium coincides with the interval (\u2212\u03c0/2, \u03c0/2). Suppose that a dynamic controller inspired by LQG strategies has been designed \u2013 based on the linearized system \u2013 to satisfy Assumption 1, namely \u03be\u0307 = (1 \u2212 \u2113\u221e)\u03be + \u2113\u221ex + u , u = \u2212p\u221e\u03be , (10) with p\u221e = \u2113\u221e = 1 + \u221a 2 obtained by solving the underlying Riccati (control and filtering) equations. The function V associated to the linearized system then provides an estimate of the basin of attraction of the zero equilibrium for the closed-loop extended system (9)\u2013(10) as shown by the cyan line in Fig. 1, since the union of the red and green regions depicts the subset of the state-space in which V\u0307 (x, \u03be ) \u2a7e 0. By relying on the arguments in Corollaries 1 and 2 and the discussion Please cite this article as: T. Mylvaganam, C. Possieri andM. Sassano, Global stabilization of nonlinear systems via hybrid implementation of dynamic continuous-time local controllers. Automatica (2019), https://doi.org/10.1016/j.automatica.2019.04.002. in Remark 2, the conditions (HR) are checked in the region where V\u0307 (x, \u03be ) \u2a7e 0" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0003613_012066-Figure8-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0003613_012066-Figure8-1.png", "caption": "Figure 8. Contour of von-Mises stress on a coil spring with a defect of 3.5 mm given the minimum load.", "texts": [], "surrounding_texts": [ "Transport vehicles require a good suspension system to dampen vibration, swings and shocks received as they travel along bumpy, hollow, and uneven roads [1]. These conditions are very uncomfortable and may cause accidents. The suspension is also expected to hold the load during some common vehicle maneuvers such as acceleration, braking or deflection while on the road [2]. The coil spring is one of the main components for dampening vibrations and shocks to the load so as to provide comfort and security while the vehicle is in motion [3]. Depending on the condition of their application, coil springs often sustain fatigue failure. This indicates that the tension received below by the coil spring from the maximum stress of the material while sustaining a dynamic load causes fatigue failure [4-8]. The yield strength of the material is also a criterion of failure. Components of automotive suspension must be changed with a traveling distance of 73,500 km, or every five years [9]. The fault of 13.18 % of 24.2 million vehicle tests was recorded [10]. With the development of computing technology, the numerical analysis method has become particularly suitable for use because it will increase the calculation efficiency, the cost-effectiveness as well as save time. Various numerical analysis methods are widely available, but the finite element analysis (FEA) has proven to be reliable in solving problems in the field of continuum mechanics [11]." ] }, { "image_filename": "designv11_62_0003043_mawe.201700160-Figure2-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0003043_mawe.201700160-Figure2-1.png", "caption": "Figure 2. Schematic representation of the experimental setup, showing the weld test piece (bottom), placed on the sample manipulator, the laser optics (top) and the infrared camera (on the left).", "texts": [ " Low-transformation-temperature filler materials are designed to transform to martensite at low temperatures in order to be effective in mitigating residual stresses [9]. The principal alloying elements for this purpose are chromium and nickel. The process parameters for each of the welds remain constant, Table 2. During welding, the quasi-stationary temperature field in the vicinity of the heat source is captured using infrared thermography. This is achieved by placing an infrared camera so that it is focussed on the point at which the laser beam is incident upon the plate, Figure 2. The plate is then translated during welding so that the camera remains focussed on the region in the vicinity of the molten pool. Images are captured at a frequency of 1 kHz during welding. It is possible to adjust the range over which temperatures are measured to improve resolution. The accuracy of the temperature measurements is 2 % of the measurement range under consideration, for temperatures >100 8C. During welding, the emitted radiation changes due to physical changes within the material, and due to factors such as the viewing angle and surface quality" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0001610_icisa.2012.6220936-Figure4-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0001610_icisa.2012.6220936-Figure4-1.png", "caption": "Figure 4. The sub-graphic instance:topology stable.", "texts": [ " Thus, the node 5 decides that node 2 is the upstream node. The Upstream_Node state of node 5 set into node 2 form NULL state. Then, node 5 broadcast the upstream_change message to neighbors. Base on upstream update process, the node 5 needs to consider the information of node 2 and node 3. Therefore, the probability that the sink can reach node 5 by a broadcast \u03bbG(v) = [1-(1-0.855)(1-0.68)]=0.9536. Besides, node 6 and 7 also compute the upstream reachable probability and node reachable probability by using topology construction algorithm. In Fig. 4, node 5 broadcasts the beacon message with information of itself to neighbors. The upstream node of node 7 is node 3. When the node 6 and node 7 receive this message, node 7 find the upstream reachable probability of node 5 (\u03bbup(5)(7) = 0.9 0.989 = 0.89) higher than node 3 (\u03bbup(3)(7) = 0.85 0.6=0.51). Thus, the upstream node of node 7 switches to node 5 from node 3. Then, the node reachable probability of node 7 need to update (\u03bbG(7) = 0.983 0.9=0.93). After the node 6 also receives this beacon packet, it finds the new route via node 5" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0000047_s13272-013-0095-7-Figure6-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0000047_s13272-013-0095-7-Figure6-1.png", "caption": "Fig. 6 Optical-inertial sensor system", "texts": [ " On top of the bumper of the cargo hook an optical marker (see Sect. 2.3.2) is placed to serve as reference point for the optical load position measurement. The optical-inertial sensor system delivers the longitudinal and lateral cable angle and rates in respect to the helicopter\u2019s body coordinate system. The information of the actual helicopter position used for the automatic load positioning system (ALPS) is determined by the DGPS System of the ACT/FHS. The optical-inertial sensor system (see Fig. 6) consists of a smart camera, an inertial measurement system and a control unit. All these components are integrated in a compact, robust housing. The optical-inertial sensor needs a 28 V DC DO-160 compliant power supply and a CANBUS connection for data transfer. The smart camera is equipped with wide-angle lens, band-pass filter, CCD sensor with a resolution of 640 9 480 pixels and a 1 GHz digital signal processor (DSP). Daylight may reduce the contrast between the optical marker (see Sect. 2.3" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0002185_i2017-11352-9-Figure11-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0002185_i2017-11352-9-Figure11-1.png", "caption": "Fig. 11. The coordinate position and angles of the cross section of the constraint cylinder. (a) Elliptical cross section. (b) Circular cross section.", "texts": [ " [26], x1 = H\u0303, x3 = 1 can be determined. Similar to analyses in the literature [23,26], three boundary conditions satisfied by eq. (1) can be obtained: I1 = (\u2212H\u0303 sin \u03b3 + x2 cos \u03b3)r\u0303 + m2x6 = const (A.3) I2 = 1 2 (1 + \u03c1)(x4 sin \u03c6 \u2212 x5 cos \u03c6 cos \u03b8 + x6 cos \u03c6 sin \u03b8)2 + 1 2 (x4 cos \u03c6 + x5 sin \u03c6 cos \u03b8 \u2212 x6 sin\u03c6 sin \u03b8)2 + 1 2 (1 + \u03bd)(x5 sin \u03b8 + x6 cos \u03b8)2 + cos \u03b8 m2 + x2 sin \u03b8 m2 = const (A.4) I3 = x5 sin \u03b8 + x6 cos \u03b8 = const. (A.5) In a circular cross section of a constrained cylinder, \u03c1 = 0, \u03b3 = 0, r = const. (A.6) As shown in fig. 11(a), point P and point P0 on the elliptical cross section coincide. The radial and circumferential axes \u2212d20 and d10 from point P coincide with the radial and circumferential axes er1 and e\u03c80 from point P0, respectively. They can be recorded as the radial and circumferential axes er1 and e\u03c81 from point P , as shown in fig. 11(b). By consolidating eqs. (A.3)\u2013(A.5), the original system can be further simplified as a system with one degree of freedom. The analytic expressions of the four heteroclinic solutions x (appendix B.1\u2013B.4) can be obtained by direct application of the undetermined Pade\u0301 approximation method [27\u201330]: xL1 = QPA4 = (1 \u2212 0.56656e6.77255t \u2212 1.73933e13.5451t + 0.0738418e20.31765t + 0.0248361e27.0902t)/(1 + 1.08004e6.77255t \u2212 3.42051e13.5451t \u2212 1.43855e20.3177t \u2212 0.0630491e27.0902t), (B.1) xL2 = QPA4 = (\u22120" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0002581_icuas.2017.7991368-Figure5-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0002581_icuas.2017.7991368-Figure5-1.png", "caption": "Fig. 5. EXPERIMENTAL DEVICE", "texts": [ " Apply control signal ua (t) to the first control level. The response Yc (t) is generated and serves as a reference to the second control level. 4. The second level of control calculates Y\u0302 (t+ j|t) with horizons Np2 and Nu2 using the model (16). 5. The unconstrained cost function (19) is minimized to obtain the control sequence u (t). 6. The first signal component of u (t) is sent to the real quadcopter. 7. From u (t), (10) is calculated and attitude control is computed with (22) 8. Return to step 1. Figure 5 shows the infrastructure employed to perform a real-time implementation of the proposed control strategy. The Optitrack photogrammetry system measures the quadrotor position using markers on the quadcopter. The quadrotor position arrives at a personal computer where the two MPC control loops run. MPC algorithms were programmed in C++ language where the quadratic programming code obtained from [3] was used. The personal computer sends, using a modem via the Wifi protocol, the control actions ux, uy and uz to the quadcopter" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0000299_s1990478912010127-Figure2-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0000299_s1990478912010127-Figure2-1.png", "caption": "Fig. 2. The ball on the beam", "texts": [ " The zero solution of the system x\u03071 = x2, x\u03072 = u2(t) = \u2212x1(t \u2212 2r) \u2212 2x2(t \u2212 r) is asymptotically stable for r < r0 = 1 2(1 + \u221a 5) \u221a 2 \u2212 \u221a 2 10(2 + \u221a 2) \u2248 0.02, and, for all t \u2265 0, the solution x(t) = (x1(t), x2(t)) with the initial point (0, \u03d50) admits the estimate |x(t)| < \u2016\u03d50\u2016 \u221a 2 + \u221a 2 2 \u2212 \u221a 2 (1 + \u221a 5r) er\u2212\u03b3(t\u2212r)/2, \u03b3 = r0 \u2212 r r0(2 + \u221a 2) \u2212 r(r0\u2212r) ln r0\u2212ln r > 0. The function V (x) = 3x2 1 + 2x1x2 + x2 2 satisfies the conditions of Theorem 1 for the system in question. Example 1. Consider a model of the elementary mechanical system consisting of a rotating homogeneous beam and a homogeneous ball rolling along it (see Fig. 2). Various problems of control concerning this system and the history of their study can be found in [19]. Assume that the rotation axis of the beam goes through its center of inertia. The dynamics of such a system is described by the couple of equations m [ r\u0308 ( 1 + J1/mR2 1 ) + \u03b8\u0308(\u2212d \u2212 R1 \u2212 J1/mR1) \u2212 r\u03b8\u03072 + g sin \u03b8 ] + C1r\u0307 = 0, JOURNAL OF APPLIED AND INDUSTRIAL MATHEMATICS Vol. 6 No. 1 2012 \u2212 mr\u0308(d + r1 + J1/mR1) + \u03b8\u0308[J2 + m((d + R1)2 + r2 + J1/m)] + 2mrr\u0307\u03b8\u0307 \u2212 mg(sin \u03b8(R1 + d) \u2212 r cos \u03b8) + C2\u03b8\u0307 = \u03c4, where m is the mass of the ball, J1 and J2 are the moments of inertia of the ball and the beam correspondingly, C1 and C2 are the coefficients of viscous friction between the ball and the beam and that in the rotating mechanism correspondingly, \u03c4 is the torque applied to the beam; and other designations are clear in Fig. 2. Denote x1 = r, x2 = x\u03071, x3 = \u03b8, x4 = x\u03073, and u = \u03b8\u0308. Thus, the problem is reduced to the stabilization of the zero solution of the system x\u03071 = x2, x\u03072 = k21u + k22x1x 2 4 \u2212 k23 sin x3 \u2212 k24x2, x\u03073 = x4, x\u03074 = u, where k21 = R1 + dR2 1 R2 1 + J1/m , k22 = R2 1 R2 1 + J1/m , k23 = k22g, k24 = C1R 2 1 mR2 1 + J1 . This is a triangular system in which x(t) = (x1(t), x2(t)) and y(t) = (x3(t), x4(t)); and if the control u does not depend on the values of coordinates x1 and x2 then the vector-function g(t, xt, yt) equals ( 0, k22x1(t)x2 4(t) \u2212 k23 sin x3(t) ) , and the correspondent subsystem (6) has the form x\u03071 = x2, x\u03072 = \u2212k24x2" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0003707_icuas.2019.8798229-Figure2-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0003707_icuas.2019.8798229-Figure2-1.png", "caption": "Fig. 2. Drawing of the Skywalker X8 including body-fixed axes {xb, yb, zb} and wind axes {xw, yw, zw}, as well as the angle of attack \u03b1 and side-slip angle \u03b2. The relative velocity vector vr is aligned with xw . With courtesy of [19].", "texts": [ " Development towards a real-time applicable solution in which model inaccuracies will be taken into acount will be part of future work. A a real-time iteration scheme together with a sequential quadratic programming approach (see e.g. [25]) will be the most likely modifications to the presented algorithm. In this section we will discuss implementation of the designed controller and compare its performance to that of a set of conventional PID controllers in a simulation study. A. Implementation The vehicle model for the simulation study is the Skywalker X8 shown previously in Fig. 2. The structure of the model matches the structure of the aerodynamic equations discussed in Sec. II and the parameters of the aerodynamic model are given in [19]. The parameters for the thrust model are taken from [26]. Both the NMPC and the forward simulation are based on the same set of parameters. The dynamic equations are integrated at 100 Hertz using an explicit fourth-order Runge-Kutta integrator. For the NMPC, the time horizon T = 10 s is chosen which is divided into N = 40 control intervals" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0000308_cdc.2012.6426165-Figure1-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0000308_cdc.2012.6426165-Figure1-1.png", "caption": "Fig. 1. Illustrative example in which the discontinuous Lyapunov function does not allow to prove stability.", "texts": [ " In order to have some convexity and concavity properties at the same time we ask here ourselves the following question: is it possible to find two different reparametrizations, one that convexifies and another that concavifies the nonlinear parametrization? Is this property useful to prove stability? This idea is illustrated in [7] with three precise examples when dealing with the problem of identification of nonlinear regressions. However, to carry out the stability analysis, discontinuous Lyapunov functions need to be used, which turns out to be a main obstacle in the proof of stability. Figure 1 illustrates this situation. Indeed, due to the discontinuity of V , the level curves are not closed and the trajectories can escape to infinity. We show in the next section that stability can be achieved using a judiciously chosen non-certainty equiva- lence switched based control structure. This structure allows the use of a standard continuous Lyapunov function to prove stability. Then, in a certain way, non-certainty equivalence replaces the need for a switched Lyapunov function. III. CONTROL OF NON-LINEARLY - MULTILINEAR TYPE - PARAMETERIZED SYSTEMS WITH A PARAMETERIZATION DEPENDENT ON THE STATE We consider the class of plants in (1) where Xp \u2208 IRn is the plant state, u(t) \u2208 IR is the control input, \u03b8j , j = 1, \u00b7 \u00b7 \u00b7 , l are scalar unknown parameters, Ap \u2208 IRn\u00d7n is unknown and b \u2208 IRn is known with (Ap, b) controllable, and \u03c6i are nonlinear functions of Xp" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0002475_icempe.2017.7982133-Figure1-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0002475_icempe.2017.7982133-Figure1-1.png", "caption": "Fig. 1. Force conditions of spherical free conducting particle", "texts": [ "00 \u00a92017 IEEE electric field can be neglected in most of the area within the cavity which far away from the basin insulator and pipe interface, thus, the radial electric field could be considered when we deal with force analysis of spherical free conducting particle. Before particle lift off, the gravitational force G, Coulomb force Fq and normal force FN mainly acts at the particle; after particle lift off, the gravitational force G, Coulomb force FQ, normal force FN, electrical gradient force Fgrad and gas viscous force Fvisc mainly acts at the particle. The force conditions are shown in Fig. 1 (a) and (b). In the following discussion, the voltage of high voltage conductor is taken as a positive voltage, and the vertical direction is taken as a positive direction. The gravitational force G always acts on the spherical particle can be expressed as = \u2212 (1) Where a denotes the particle radius, g the gravitational acceleration, \u03c1 the particle density. The Coulomb force Fq magnitude depends on the electric field of particle position, this force can be expressed as = \u2212 (2) Where k denotes the correction factor caused by the electronimage force, q the charge carried by the particle, and E the intensity of the electric field at the particle" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0002799_icinfa.2017.8079003-Figure2-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0002799_icinfa.2017.8079003-Figure2-1.png", "caption": "Fig. 2 L1Guidance Law", "texts": [ " In the last section of the paper, numerical simulations verify the correctness of the control law design and the performance of the controller. II. L1 NONLINEAR GUIDANCE LAW The L1 guidance law was first proposed by Amidi for ground vehicles in [13], and then improved by park and used in fixed wing aircrafts. The L1 guidance law is briefly described below as derived in [12]. The basic train of thought of the L1 guidance law is that a reference point is selected on the target path, and a lateral acceleration is commanded on the basis of the position of the reference point and the aircraft. Fig. 2 shows the principle the L1 guidance law. The symbol of V is a horizontal velocity vector in the inertial reference system. The angle between the velocity vector V and L1 segment is \u03b7 . D is an arc with a radius of R, starting at the current position of the aircraft and is tangent to the velocity vector of the aircraft. 978-1-5386-3154-6/17/$31.00 \u00a92017 IEEE Proceedings of the 2017 IEEE International Conference on Information and Automation (ICIA) Macau SAR, China, July 2017 737 The L1 segment in Fig. 2 always starts from the position of the aircraft and terminates on the desired route. The starting and end points of the L1 segment are moving forward constantly. The arc D connects both ends of the L1 segment and is uniquely determined by the two endpoints and the radius R. According to the geometric relation, the center angle of the arc is 2\u03b7 . We connect the center of the circle and the midpoint of the L1 segment and then 1 2sin L R \u03b7 = (1) In order to maintain the centripetal force required by the aircraft to complete the circumferential motion, the centripetal acceleration required is: 2 c V a R = (2) Substituting (1) into (2) produces the expected lateral acceleration: 2 1 2 sincmd V a L \u03b7= (3) The direction of the command acceleration is perpendicular to the velocity vector in the inertial reference system and is directed toward the centre of the circle" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0002509_msec2017-2841-Figure2-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0002509_msec2017-2841-Figure2-1.png", "caption": "FIGURE 2 - SCHEMATIC REPRESENTATION OF TRACING PATH OF SURFACE PROFILER TIP FOR WEAR TRACK PROFILE MEASUREMENT", "texts": [ " For wear test 12mm\u00d712mm samples were cut and top surface of the clad-layers were polished with a P400 grade silicon carbide paper until the undulations due to track overlaps were smoothened. Then ball-on- disk wear test on the polished top surface of the two samples having two different clad-layers were performed with the help of a tribometer (DUCOM, TR-20M42) against 5 mm diameter WC ball at 300 RPM with 2 kg load for 20 minutes duration. Wear track diameter was around 6 mm. Wear track profiles were measured with the help of a surface profiler (Taylor Hobson, Form Talysurf 50) by tracing the profiler tip along the diameter of the circular wear track as shown schematically in Figure 2. Thus, single movement of profiler tip recorded the profile of wear track at two positions. From two pairs of profiles of each clad-track, average wear track cross-sectional area and corresponding wear volume were calculated. Wear track Potentiodynamic polarization tests were also conducted to compare the corrosion resistance of the clad-layers of the two Stellite alloys using an AUTOLAB Potentiostat in 3.5 % NaCl solution at room temperature with standard calomel reference electrode, a platinum rod as counter electrode and the polished clad-layer of samples as work-electrode" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0003163_b978-0-12-394400-9.00007-1-Figure7.25-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0003163_b978-0-12-394400-9.00007-1-Figure7.25-1.png", "caption": "Figure 7.25 (a) For Fm3m and Fd3m, we show parts of the page from the ITA indicating the plane groups for special projections. (b) The projection onto the (110) face of the fcc structure. The (001) view of the three-dimensional structure is also shown to help in understanding the projection along [110]. The atoms at height z\u00bc\u00bd are shown dashed. (c) Similar to (b), but for the diamond structure. (d) The atoms on the (110) face of an fcc structure. (e) Similar to (d), but for the diamond structure.", "texts": [ " Consider any structure that has either the Fm3m\u00f0O5 h\u00de or Fd3m\u00f0O7 h\u00de space groups. Cu with the fcc structure in Fig. 7.1a and Si with the diamond structure in Fig. 7.2f serve as useful models. The ITA gives the plane groups for projections in a few special directions. For example, when projected along a [110] direction (in the cubic system, [110] is perpendicular to the (110) plane), any structure with either of these space groups will have the symmetry of the c2mm plane group with b0 \u00bc c and a0 \u00bc\u00bd( a\u00fe b). Figure 7.25a shows part of the page from the ITAwith the relevant information for these two space groups. Projections for the fcc structure and diamond structures are shown in Fig. 7.25b and c, respectively, where for both structures we also show views of the three-dimensional structures that clarify how the projections along [110] are obtained. For the projection of the fcc structure (Fig. 7.25b), the centered cell is obvious for the c2mm plane group; the plane group cell has dimensions b0 and a0, or a and ffiffiffi 2 p a/2, where a is the edge length of the conventional three-dimensional unit cell. The plane group unit cell is indicated by the dashed line. For the projection of the diamond structure, the same plane group and unit cell is obtained (dashed line), but the result is more complicated. This arises because the basis in the three-dimensional structure is more involved, being composed of atoms at (0, 0, 0) and (\u00bc, \u00bc, \u00bc). As a result, for the two-dimensional projection, the basis has atoms at (0, 0) and (\u00bd, \u00bc) with respect to the new unit cell whose dimensions are a0 b0 (Fig. 7.25c). The projection of the three-dimensional structure along the directions studied by surface scientists is not typically what is of interest. Instead, they are interested in the atomic arrangements that lie just on the surface of the crystal (i.e. in a single atomic layer). For example, again consider the fcc structure just discussed. The arrangement of atoms that occur on a single (110) plane is shown in Fig. 7.25d. Note how it differs from Fig. 7.25b, and that it has the p2mm plane group symmetry, rather than c2mm. The reason for this should be clear from Fig. 7.25b; in the projection of the structure, the atoms that are projected onto the centered positions are not those on the actual (110) surface that we have chosen. Similarly, for the diamond structure, the arrangement of atoms on a single (110) plane is shown in Fig. 7.25e. This arrangement also has p2mm plane group symmetry with a basis of Si at (0, 0) and (\u00bd, \u00bc). The atoms on or very near a clean surface of a crystal need not maintain the atomic positions found in the bulk. This is because atoms on the surface (taken as the xy-plane) interact with atoms below the surface plane but not above. Deviations from the bulk atomic positions give rise to two types of changes. The first type is relaxation in which the symmetry of the atoms on or near the surface is maintained, but the distances change between the successive planes of atoms parallel to the xy-plane (so that each atomic layer has the same unit cell but the surface layers translate with respect to the bulk)", "26d), the unit cell is again doubled in the x-direction, but the plane group is pm because the two atoms on the surface are at different heights; the plane group would be p1 if these two atoms were to move in the y-direction with respect to each other. Consideration of the different plane group symmetries could play a role in deducing the best model. S on Ir(110): We have already discussed the reconstruction of a surface with no impurities. Let us consider the Ir(110) surface reconstruction as well as the arrangement of sulphur (S) atoms that are adsorbed onto this reconstructed surface. For the fcc structure, a projection along [110] yields the c2mm plane group (Fig. 7.25b), while the plane group for the topmost layer of atoms is p2mm. Figure 7.27a again shows the (110) plane of an fcc structure with the unit cell outlined; this has dimensions a a ffiffiffi 2 p . When the (110) surfaces of the free metals Ir, Pt and Au are studied by LEED intensity analysis, a 1 2 reconstructed missing-row model for the surface is found. This reconstruction is shown in Fig. 7.27b. Parallel to the [100] direction, a row of atoms from the topmost layer is removed. The unit cell for this layer is the same size in the x-direction as for the unreconstructed surface" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0001597_iros.2012.6386081-Figure10-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0001597_iros.2012.6386081-Figure10-1.png", "caption": "Fig. 10. Detailed drawing of wire structure of FRT-type thumb exoskeleton", "texts": [ " Both a thumb of a wearer and an environment contact part of the exoskeleton touch the environment. A nonslip sponge is installed on the thumb contact part so that it prevents a wearer\u2019s thumb from moving at the thumb contact part when grasping an object. The distribution factor of the assistive force is determined by a horizontal position of the intermediate point. It becomes 1:1.2(=cos\u03c82 : cos\u03c81) when an angle inside the exoskeleton, \u03c81, is 26\u25e6 and an angle outside the exoskeleton, \u03c82, is 43\u25e6 as shown in Fig.10. A bioelectric potential is measured by surface electrodes for grasping force estimation. Our developed active electrode that includes an impedance transfer for artifact reduction, amplifier (\u00d75000 - 20000), and a band-pass filter is attached over the first dorsal interosseous (FDI) via two Ag/AgCl gel sheets. The dimensions of the active electrode is 25 [mm] long, 34 [mm] wide, and 8.5 [mm] high and its weight is 6 [g]. The active electrode is shown in Fig.11. A human hand has very wide range with regards to finger position and grasping force" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0002551_iemdc.2017.8002063-Figure6-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0002551_iemdc.2017.8002063-Figure6-1.png", "caption": "Fig. 6. Fluxes of the machine when the current Irp achieves its peak value", "texts": [ " When a machine experiences a transient like a sudden stator short-circuit, currents can reach several times their nominal values, leakage inductances can no longer be treated as constant if one wants to assure a good accuracy of dynamic simulations [7]. During a regular operation, leakage flux is very hard to evaluate, because flux leakage is very small when compared to the common flux. In the simulation presented in this article, saturation of the leakage fluxes can be observed. In order to see the saturation of leakage fluxes, we can look at two particular instants of short-circuit during which the common flux is weak compared to the leakage fluxes. First, we take the instant when the field current reaches its peak value, Fig. 6, most of the stator flux linkage is leakage flux. Pole tips and stator teeth are strongly saturated. Main flux in the stator is very weak. Second during the short-circuit steady state, Fig. 7, most of the flux linkages constitute also leakage fluxes. This time, the magnetic circuit is not saturated. This type of saturation would not be taken into consideration by most models, which assume the linearity of the leakage inductances. If the fluxes in the saturated state are inaccurate, it has a direct impact on the machine torque, power and power factor" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0002302_b978-0-12-803581-8.10351-0-Figure21-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0002302_b978-0-12-803581-8.10351-0-Figure21-1.png", "caption": "Fig. 21 Roll structures within monocoque.", "texts": [ " Inside the shells that form the monocoque, bulkheads are required to carry loads across the chassis and the loads from the necessary rollover structures. There is a front bulkhead made from composite that supports the front crush zone, an intermediate bulkhead also made from composite for suspension mounting, a main forward bulkhead that includes composite panels, which surround and integrate a steel roll hoop mandated by certain competition rules, and a rear composite bulkhead, which is the structural location for attachment of the engine and drive assembly (Fig. 21). The composite bulkheads incorporate molded flanges around the perimeter, which follow the inner monocoque profiles, enabling enhanced structural bond areas. The purpose of the bulkheads is to distribute localized loads, such as the suspension loads, to the main monocoque shells and to provide additional stiffness to the overall structure. Bulkheads commonly serve as crossframes, transferring loads around and across the driver volume within the monocoque. The front bulkhead is considered a part of the rollover structure by the rules for this class of chassis and must withstand a specified rollover force in order to be in compliance" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0001082_upec.2013.6715001-Figure3-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0001082_upec.2013.6715001-Figure3-1.png", "caption": "Fig. 3. Example of the distribution of magnetic energy density (J/m3), at the middle of the winding height, for the 3D model.", "texts": [ " Exact values of inductances for the 2D and 3D models If the total magnetic energy (EMAG) of the analyzed region is obtained from the FEM software [16], then the inductances (L) can be easily computed: L2D = 2 EMAG-2D / I2 (1) L3D = 2 EMAG-3D /(3I2) (2) Subscripts 3D and 2D correspond to simulation of cases of Fig. 1 and Fig. 2, respectively. The current (I) in the windings is assumed to be the same for the three legs, in case 3D, because it is a zero-sequence condition. The reactances (X) are simply computed from these values of inductance, and the active part of Z0M is negligible (thus, X \u2248 Z0M). C. Magnetic energy density distribution for the 2D and 3D models Fig. 3 and Fig. 4 show an example of the distribution of magnetic energy density, at the middle of the winding height, for the 2D and 3D models, respectively (for the same transformer, and for the same conditions of current). Highest values of magnetic energy density are near to the windings. This fact suggested the proposed 2D approximate method. D. 2D approximate method The 2D approximate method is based on the separation of the magnetic energy in two parts (Fig. 2, Fig. 4): a) E1, from the symmetry axis until DLC/2; b) E2, from DLC/2 until the end of the analyzed space" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0000972_physreve.83.056315-Figure1-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0000972_physreve.83.056315-Figure1-1.png", "caption": "FIG. 1. Sketch of a planar sheet moving in the \u2212z direction due to a plane wave current density, polarized in the y direction, and running in the z direction with phase velocity c = \u03c9/k. The torque density in the ferrofluid, which generates the motion, is in the y direction.", "texts": [ "10) \u03c9p2y(x) = \u00b1We\u2212\u03ba|x| \u00b1 C \u03be e\u22122k|x|, for x \u2277 0, where \u03ba = \u221a 4\u03b7\u03b6 \u03b7\u2032(\u03b7 + \u03b6 ) , \u03be = 4\u03b6 ( 1 \u2212 4k2 \u03ba2 ) . (3.11) The coefficients U and W can be determined from the boundary conditions. From the no-slip condition for v and from Eq. (2.14) for \u03c9p one finds U2 = C 2(\u03b7 + \u03b6 )(2k + \u03ba)(1 + \u03ba\u03bbs) , (3.12) W = C\u03ba2(1 + 2k\u03bbs) 4\u03b6 (4k2 \u2212 \u03ba2)(1 + \u03ba\u03bbs) . The flow velocity v2z(x) is even in x and the particle rotational velocity \u03c9p2y(x) is odd in x. At large distance |x| the flow velocity tends to U2. This implies that in the laboratory frame the sheet moves in the \u2212z direction with velocity U = \u2212U2ez. In Fig. 1 we show a sketch of the geometry. To second order in K0 the dissipation in the system is purely magnetic. From the linear relaxation equation \u2202 M1 \u2202t = \u2212\u03b3 (M1 \u2212 \u03c70 H1), (3.13) one derives \u2202 \u2202t ( \u03bc0 2\u03c70 M2 1 ) = \u03bc0 H1 \u00b7 \u2202 M1 \u2202t \u2212 \u03bc0\u03b3 \u03c70 (M1 \u2212 \u03c70 H1)2. (3.14) The left-hand side is the rate of change of the secondorder magnetization energy density, the first term on the right represents the work done by the magnetic field as the magnetization varies, and the second term on the right represents the local rate of dissipation: m2 = \u03bc0\u03b3 \u03c70 (M1 \u2212 \u03c70 H1)2" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0002200_978-3-319-51222-8-Figure1.16-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0002200_978-3-319-51222-8-Figure1.16-1.png", "caption": "Fig. 1.16 Influence function for the moment in a rafter", "texts": [ " A standard application of the principle of virtual displacements is the calculation of the support reactions of statically determinate structures, e.g., the support reaction RA of the beam in Fig. 1.15. We start by removing the left support so that we can rotate the beam about its right end, \u03b4w(x) = 1 \u2212 x/l, G (w, \u03b4w) = \u222b l 0 p(x) \u03b4w(x) dx \u2212 V (0) \u03b4w(0) = 0 , (1.133) and we solve this identity for RA = V (0) Ra \u00b7 1 = \u222b l 0 p \u00b7 (1 \u2212 x l ) dx . (1.134) In the case of a slanted beam, a rafter, as in Fig. 1.16, only the part of \u03b4w(x) which points in the direction of the traveling load counts. We could apply the principle of virtual displacements also to compute internal forces, for example, the shear force of the beam in Fig. 1.15c. In this case, we would install a shear hinge at the source point and spread it by one unit apart. But this is the same as if we would apply Betti\u2019s theorem B (w1, w2) = W1,2 \u2212 W2,1 = W1,2 = 0 (W2,1 = 0) . (1.135) 30 1 Basics Fig. 1.15 One-span beam and influence functions (a) (b) (c) (d) The function w2 is the rigid body motion, the rotation of the beam about its support or the spread of the shear hinge, and w1 is the deflection curve of the beam" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0000480_s00170-013-4974-1-Figure6-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0000480_s00170-013-4974-1-Figure6-1.png", "caption": "Fig. 6 The drawing of the date corrector pinion", "texts": [ " The system is controlled by two identical multiaxis controllers (manufacturer: Googol Technology Ltd., model: GT400-SV) [http://www.googoltech.com]. The two controllers are installed on an industrial personal computer (IPC) through a peripheral component interconnection slot. All motion control programs, including the polar coordinate programming modules, are developed in the IPC. A large number of machining tests have been conducted. A typical example of the polar coordinate programming module is to machine the part shown in Fig. 6. Note that the part is only about 2.2 mm in size. This part is difficult to machine because of its complex shape and small size. In the part, the cylinders with diameters of 0.8 and 0.5 mm are machined by the turning operations, and the lines in the right view of Fig. 7. Lines AB, CD, and the cylinders with diameters of 1.43 and 2.2 mm are machined by the milling operation based on the developed programming module based on the polar coordinate. In order to use the programming module, C1 axis should be changed as position mode to generate a rotation movement", " The experimental results are shown in Table 1, from which it can be concluded that the maximum contouring errors of line AB, circle CD, and line EF are smaller by using the developed CNC program module. Theoretically, the Cartesian coordinate system can get better contouring accuracy in line interpolation, but the contouring accuracy in circular interpolation is low, and the body structure of the machine center becomes too complex, so the paper does not use this method. Five key dimensions of ten random samples are measured to analyze the machining errors of the parts. The measured dimensions are marked in Fig. 6, and the results are shown in Table 2. From the table, it is concluded that the programming modules are effective and precise. Comparing to the existing commercial CNC turn-mill machining centers, the presented polar coordinate interpolation modules have a number of advantages. First, with the same interpolation steps, the contour accuracy is significantly higher. Second, the programming module is easy to use with only a few G code lines. Third, machining cost is lower because of the reduced machining time" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0000632_amr.490-495.589-Figure4-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0000632_amr.490-495.589-Figure4-1.png", "caption": "Fig. 4 shows a robot with a grinder. {0} is the robot base frame. Og is the grinder head\u2019s center of gravity. The gravity compensation of the grinder head is calculated by Equations (3), (4)", "texts": [], "surrounding_texts": [ "In order to implement the force control scheme into the system a simulation is carried out. In the simulation, the matlab R2008a and robotics toolbox 8 are used. Fig. 5 shows the simulation model. In this simulation, we suppose that the robot is in the following state: (a) the grinder of the robot contacts with the workpiece; (b) the force between the grinder and the workpiece is 0; (c) the Z direction position is 0. If the robot moves the tool in Z direction, the force will be detected by the force sensor. The SIASUN RH6 robot moves the tool in a circle of radius 0.05m in X and Y directions, centered at the point (0.5, 0, 0). The desired force is 2N in Z direction. The robot is modeled by an integrator as a simple velocity servo. Fig. 6 shows the actual tool path of the robot in X and Y directions. We can see that the tracking path in X and Y directions is the desired circle without the influence of the force. Fig. 7 shows the actual force between the grinder and the workpiece. The robot tracks the desired force quickly. In order to remove the frequency interference, a low-pass filter whose cutoff frequency is 5Hz is used. Fig. 8 shows the force data without filter. Fig. 9 shows the force data with a low-pass filter. We can see that the filter can remove the frequency interference effectively. 0.45 0.5 0.55 -0.05 0 0.05 X/m Y /m 0 2 4 6 8 1 2 3 T/S F /N Fig. 6 The tracking result for a circle path Fig.7 Force between the grinder and the workpiece" ] }, { "image_filename": "designv11_62_0003071_iscid.2017.115-Figure3-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0003071_iscid.2017.115-Figure3-1.png", "caption": "Figure 3. Path planning", "texts": [ " The relationships between , and are shown as follows:8/++ (4) The following path planning for automatic parallel vehicle based on the minimum size of the parking space, so and , then we can get the target point of the vehicle as follows: (5) IV. PATH PLANNING In most cases, the initial direction of the vehicle axis is not parallel to the parking space, so we propose an algorithm for automatic parallel parking from arbitrary initial angle. The parallel parking path planning is to generate a geometric path composed of two tangential arcs of different radii. Given a barrier-free environment, the path of the vehicle can be expressed in Fig.3. The initial coordinate is , the coordinate of the intersection of the two tangential arcs is , and the target coordinate is . The center of two arcs are and respectively. The radii of two arcs are and . By Eq. (3) (4) (5), and can be obtained. By the laser sensors and coordinate conversion, the initial coordinate can be determined that and . Then the center , , and the coordinate of the intersection can be calculated as follows: (6) (7) (8) (9) Considering the vehicle performance constraint, the radius of the first arcs must satisfy that , so when the initial angle , the initial y coordinate must satisfy that " ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0001081_qr2mse.2013.6625567-Figure1-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0001081_qr2mse.2013.6625567-Figure1-1.png", "caption": "Figure 1. The structure of the 1.5MW wind power gearbox.", "texts": [ " DYNAMIC RELIABILITY ANALYSIS OF THE WIND POWER GEARBOX COMPONENTS P: planet wheel; r: internal gear (fixed); c: planet carrier; s: sun wheel; 4: level 1 active helical gear; 5: level 1 driven helical gear; 6: secondary active bevel gear; 7: secondary driven helical gear; Tin: input torque; Tout: output torque; gi: the code of rolling bearings; i: on behalf of the bearing Numbers, i=1, 2, \u2026 , 8. In this paper, we take the sun wheel of 1.5MW horizontal axis wind power gearbox as example and then dynamic reliability and failure rate are performed. The gearbox adopts a level NGW planet gear and two levels stage helical gear transmission. Its structure is given in Fig.1. The planetary gears are the low-speed level; the helical gear pairs 4-5 are the middle-speed level and helical gear pairs 6-7 are the higherspeed level in this transmission system. The sun wheel\u2019s revolving speed s 78.93r/minn , the diameter of reference circle s 351mmd , the transmitted power 1.5385 0.0234MWP , tooth width 370mmb , number of teeth s 27z , modulus 13mmm . The 20CrMnTi are selected as the sun wheel\u2019s material and the surface is treated by carburizing and quenching technology" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0001115_acc.2010.5530641-Figure1-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0001115_acc.2010.5530641-Figure1-1.png", "caption": "Fig. 1. A network of 11 integrators and the applied decentralization scheme.", "texts": [ "1, for a given set D of admissible demands, the set of parameters \u03a0 for which there exists a decentralized RCI set may be empty. To resolve this, one may introduce additional parameters for lower and upper bounds in the constraints on the state vector x, thus permitting the variation of the storage capacities in order to cope with the given demands d \u2208 D. This extension is direct and the resulting extended set of parameters is again, using observations in Lemma 4.1, a polyhedral set. Consider a network of integrators defined by the graph shown in Figure 1. The upper bounds on the states x(i), inputs u(i) and the disturbances d(i) are given in Table I. Lower bounds on xi and di and ui are 0 for all i \u2208 N[1,11]. Considered network decomposition, as indicated in Figure 1, is \u2206 = {S1,S2,S3,S4}, where S1 = {1, 3, 6}, S2 = {2, 4, 5}, S3 = {7, 8, 9} and S4 = {10, 11}. The goal is to find a set of edge capacities uij and uij connecting the nodes such that the decentralized robust control invariant set exists for the network. The upper and lower bounds u(i) and u(i) are kept fixed, while the edge capacities uij and uij , (i, j) \u2208 E , are optimized by minimizing the cost: \u2016uij\u20161 + \u2016uij\u20161 subject to feasibility constraints defined by (21) and the bounds: \u22122 \u2264 uij \u2264 uij \u2264 2, for (i, j) \u2208 E " ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0003254_s00170-019-03312-1-Figure12-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0003254_s00170-019-03312-1-Figure12-1.png", "caption": "Fig. 12 (a) Obround representations for material extruded beads (b) undercuts and possible simplified models: horizontal boundary trimming and introduction of a 90\u00b0 shoulder, and (c) visual region models", "texts": [ " l\u2212calculated \u00bc bheight sin\u03b8 \u00f06\u00de Large triangle : Area calculated \u00bc b2height 8 * cos\u03b8 sin\u03b8 \u00f07\u00de Small triangle : Area calculated \u00bc R2 1 2 or Area calculated \u00bc R2 2 2 \u00f08\u00de Chord area \u00bc Area calculated \u00bc R2 1 2 \u03c0 2 \u22121 or Area calculated \u00bc R2 2 2 \u03c0 2 \u22121 \u00f09\u00de Case 1 : Ra \u00bc Large triangle\u2212small triangle\u00fe chord area\u00f0 \u00deR1\u00fe Large triangle\u2212small triangle\u00fe chord area\u00f0 \u00deR2 l calculated critical angle \u00bc tan\u22121 bheight 2*R1 ortan\u22121 bheight R1 \u00fe R2 \u00f010\u00de Case 2 : Ra \u00bc Large triangle\u2212small triangle\u00fe chord area\u00f0 \u00deR1\u00fe Large triangle\u00fe small triangle\u2212chord area\u00f0 \u00deR2 l calculated If R1 \u00bc R2;Ra \u00bc 1 4 * bheight cos\u03b8\u00f0 \u00de \u00f011\u00de Case 2 critical angle\u2013two bead base line case\u00f0 \u00de : Large triangle\u2212small triangle\u00fe chord area\u00f0 \u00deR1\u00fe Large triangle\u00fe small triangle\u2212chord area\u00f0 \u00deR2\u00fe Ra \u00bc 2* small triangle\u2212chord area\u00f0 \u00deR1 l calculated \u00f012\u00de Here, it is assumed that the side-to-side beads are parallel to the boundary, and this is not a sharp corner condition. The obround geometry introduces blended curves into the bead shape, and is an accurate reflection for AM material extrusion processes that apply an external force to modify the deposited bead shape (i. e., LSAM system). For the obround bead stacking, shown in Fig. 12(a), the surface roughness can be calculated analytically using circle and rectangle/triangle based geometric relationships, but at a critical angle (Eq. 13.), complexity is introduced into the analytical model. As shown in Fig. 12(b), an undercut region can occur. The physical validity of this situation is process-heatmaterial dependent. where \u03b8critical \u00bc invtan h w\u22121 2h ! \u00f013\u00de where h is the bead height, w is the bead width and \u03b8 is the surface angle from the horizontal. Unlike the rectangle plus a fillet set\u2013based models, there are multiple options to calculate the area, mean line, and surface roughness, which are: (1) Employ a rigorous mathematical definition for the boundaries and bounded areas (i.e., trace all continuous boundaries, determine the mean lines and areas, and calculate Ra), (2) Introduce a modified version of the bead boundary contour, which would represent some heat-melt-cool geometric blending, or the characteristics of the profilometer measuring stylus (90\u00b0 measuring angle) (Fig. 12(b)), or (3) Consider a \u201cvisual\u201d region perpendicular to the inclination angle, which a laser scanner would perceive (Fig. 12(c)). Option 1 is modeled with CAD/CAM tools and with Rhino-Grasshopper\u00ae (R-G model). Option 2 is modeled via extending the base bead surface until it intersects the upper bead, and tracing the boundary curve, and by introducing a 90\u00b0 shoulder using CAD/CAM tools. Option 3 is modeled using R-G tools. A detailed flow chart illustrating the R-G approach is presented in the Appendix Fig. 23. The influence of these options is presented in Fig. 13. The noticeable plateau for the three data sets is due to the undercut-void region that can occur once the critical angle has been reached" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0000865_civemsa.2013.6617399-Figure4-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0000865_civemsa.2013.6617399-Figure4-1.png", "caption": "Figure 4: Volumetric bone model with heterogeneous properties: (a) Sketch of a femur bone structure; (b) single out a cylindrical bone volume.", "texts": [ " When a drill bit has contact with the bone, a cylindrical volume aligned with the drill bit is scooped out as in Figure 3(b). The rest of the bone model is then converted into a mesh model as in Figure 3(c) for efficient visualization (the visualization of a voxel model is slow and the surface does not look smooth). In constructing the cylindrical bone volume, bone heterogeneity is also considered since thrust force difference is significant in drilling compact bone and spongy bone. The difference helps the modeling of the sense of drill-through. The composition of a sample bone structure is illustrated in Figure 4(a). The compact bone (the bone shell) is a hard material structure while the spongy bone is a soft spongy structure. The voxel representation of the different materials can be seen in the different colors in Figure 4(b) where the spongy bone is represented as green color and the compact bone as golden color. III. HAPTIC RENDERING Developing a mathematical model of drill-bone interaction is difficult since the force largely depends on drill bit geometry and various other parameters in bone drilling. In earlier research [11, 12, 13], it was found that important parameters for a realistic force model were related to the drill speed, types of drills, feed rate and the material properties of the bone. Wiggins and Malkin [11] investigated the interrelationships between thrust pressure, feed rate, torque, and specific cutting energy (the energy per unit volume required to cut the material) for three types of drill bits" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0000491_ecce.2013.6647287-Figure1-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0000491_ecce.2013.6647287-Figure1-1.png", "caption": "Fig. 1. (a) 6pole-9slot PMSM motor cross-sectional figure, (b) series connected winding configuration with turn short fault, (c) parallel connected winding configuration with turn short fault.", "texts": [ " PMSMs are modeled for the series and parallel winding connection, respectively. The proposed model is developed under the consideration of the flux couplings between the fault winding and other healthy windings in the same phase. This flux coupling effect has to be considered especially on high pole PMSMs since it has many different flux path affecting the flux linkage between the same phase winding. The FEM simulation is conducted to validate the proposed model. II. PMSM MODEL WITH SERIES CONNECTED WINDING Fig. 1 shows (a) the figure of the 6pole-9slot PMSM motor with concentric winding and (b) the winding configuration of series connection with the inter turn short fault. a1, a2, a3, b1, b2, b3, c1, c2, and c3 are the winding number of each phase. ia, ib, ic, va, vb, and vc denote the phase current and voltage of a, b, and c-phase winding. The insulation failures do not provide a zero resistance path [17]. Hence, the failed spot is described by a fault resistance. We denote it by Rf and the turn short fault current by if. Here, it is assumed that a-phase winding has a turn short fault. In Fig. 1, the turn short fault winding makes another fault circuit loop which is composed of the fault resistance Rf, the fault winding inductance, and the fault winding flux linkage. Every winding of the motor is not only electrically connected to each other, but coupled magnetically also. Hence, the cross coupling effects should be considered in the analysis. Since the turn number of the a-phase is reduced by the shorted turn number, the PMSM motor model changes depending on the faulted turn number. To establish the fault inductance model, the magnetic flux path should be considered", " By substituting (10) into (4) and assuming the steady state, the positive and the negative sequence voltage equations are derived: Px LLLLI LLRRRI LLLLI LLRRRI V V m ssssd sssssq ssssq sssssd qe de /)1(23 22 2 22 2 3 1 12232211 241412142211 23221211 241421142211 (11) Px LLLLI LLRRRI LLLLI LLRRRI V V m ssssd sssssq ssssq sssssd qe de /)1(2 22 22 3 1 23221211 241412142211 23221211 241421142211 (12) where qe de qe de c b a V V V V v v v J\u03b8J\u03b8 TeTe 2/32/1 2/32/1 01 T , qe de V V cossin sincosJ\u03b8e Tqede VV and Tqede VV denote the positive and negative sequence voltage. By multiplication of the rotor flux linkage terms and the winding current, the torque equation is obtained 2 2sinsin)1( 22 3 1 2 2 d qmmq IIxIPT (13) The first term of (13) is the nominal torque equation and the second term is induced by inter turn short faulting. When the PMSM has no turn short fault (x=1), the second term is zero. III. PMSM MODEL WITH PARALLEL CONNECTED WINDING A. Dynamic equation of Parallel Connection PMSM with turn short fault Fig. 1 (c) shows the parallel connected winding configuration with turn short fault. ip1 and ip2 denote the remained healthy a1 winding current and other winding current. The sum of ip1 and ip2 is ia (ia=ip1+ip2). Without an inter turn short fault, ip1 and ip2 is ia/n and ia(n-1)/n. However, since a1 winding has lower impedance by an inter turn short fault, ip1 and ip2 have different current magnitudes and phases. Hence, for the parallel connected winding has to be considered not only the fault current but also the healthy winding\u2019s current distribution" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0003835_ccdc.2019.8832418-Figure2-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0003835_ccdc.2019.8832418-Figure2-1.png", "caption": "Fig 2. The simplified Model of YuMi manipulator", "texts": [ " On the basis, the RRT* algorithm is modified by the minimum distance index. For further research, we selected YuMi left arm as the object of study as shown in Fig 1. After that a corresponding simulation was provided and at last a summary is made to conclude this article. Key Words: obstacle avoidance; robot kinematics; oriented bounding box; minimum distance; path planning 4184978-1-7281-0106-4/19/$31.00 c\u00a92019 IEEE To analyze the kinematics of manipulator efficiently, the simplified model of YuMi is shown in Fig 2. Unlike the SRS structure of manipulator, the non-SRS manipulator has offsets at shoulder(S), elbow(E) and wrist(W). 2.1 Forward Kinematics The end-effector of manipulator performs the task in Cartesian space while the control is in the joint space. It is convenient to use MD-H [9] (Modified Denavit-Hartenberg) model to solve the problem with homogeneous transformation matrix. Fig 3 shows the MD-H model of the robot mechanism, and the MD-H parameters are given in Table 1. The homogeneous transformation matrix is given by: 1 1 1 0 1 i i i i i i T R P T \u2212 \u2212 \u2212 = (1) 1 1 1 1 1 1 1 cos sin 0 sin cos cos cos sin sin sin cos sin cos i i i i i i i i i i i i R \u03b8 \u03b8 \u03b8 \u03b1 \u03b8 \u03b1 \u03b1 \u03b8 \u03b1 \u03b8 \u03b1 \u03b1 \u2212 \u2212 \u2212 \u2212 \u2212 \u2212 \u2212 \u2212 = \u22c5 \u22c5 \u2212 \u22c5 \u22c5 (2) 1 1 1 1 sin cos i i i i i i i a P d d \u03b1 \u03b1 \u2212 \u2212 \u2212 \u2212 = \u2212 \u22c5 \u22c5 (3) where ia , id and 1i\u03b1 \u2212 are MD-H parameters, 1i iR \u2212 and 1i iP \u2212 give the transformation from the (i-1) axis to the i axis" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0003538_s1063780x19060011-Figure2-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0003538_s1063780x19060011-Figure2-1.png", "caption": "Fig. 2. Phase portrait of a system of M macroparticles at the time of instability saturation (\u03c4 = 1570).", "texts": [ " Such a phenomenon was called self-trapping [38]. As noted in [38], it is the effect of electron self-trapping by the beam wave that leads to the chaotization of the electron beam and disappearance of regular oscillations at the resonance frequency. Such behavior illustrated in Fig. 1 by the dependence . The effect of self-trapping of the beam electrons and their subsequent chaotization is demonstrated by the phase portraits of macroparticles shown in Figs. 2\u20134 for = 0.5 cm. It can be seen from Fig. 2 that a group of reflected macroparticles (marked with the letter R) appears in the stage of instability saturation. The macroparticles are reflected from the humps of the beam potential wave. It is worth noting that the beam electrons are self-trapped by the beam wave in the nonlinear stage of instability of a quasi-monochromatic initial perturbation under the conditions of the collective Vavilov\u2013 Cherenkov effect [38]. Figures 3 and 4 show phase portraits of macroparticles for the times \u03c4 = 2550 and 5000, corresponding to chaotic oscillations of the slowly varying amplitude" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0001478_case.2011.6042402-Figure1-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0001478_case.2011.6042402-Figure1-1.png", "caption": "Fig. 1 Positive convex function )(, ii uS \u0393\u2212\u03c3", "texts": [ " (15) For asymptotic stability, the following lemma is introduced: Lemma 1: Assume that the following relations hold: 00 >V , 00 >W , 0>JV , 0>JW . (16) Then, 00 >+= JVVV and 00 <+= JVVV &&& are guaranteed, implying that V is a Lyapunov function. Therefore, the closed-loop system is asymptotically stable. For input saturation (2), the following property is utilized: Property 1: There exists a positive convex function )(, ii uS \u0393\u2212\u03c3 such that the following relation is satisfied: ii i ii u du udS \u0393\u2212=\u0393\u2212 )( )(, \u03c3\u03c3 . (17) This function )(, ii uS \u0393\u2212\u03c3 is constructed from two straight lines and a quadratic curve as shown in Fig. 1. Property 2: The summation of the convex function \u2211 \u0393\u2212 i ii uS )(,\u03c3 is given by )}()()(){( 2 1)(, u\u03c8u\u03c8\u0393u\u0393u TT i ii uS \u2212\u2212\u2212=\u2211 \u0393\u2212\u03c3 . (18) Indeed, taking the inner product of u&\u2212 and \u03c3J and using relation (3) to replace )(u\u03c3 gives )}(){( u\u03c8\u0393uuJu \u2212\u2212=\u2212 TT && \u03c3 , (19) and using relation (17), the left-hand side term can be represented as \u2211 \u0393\u2212=\u2212 i ii T uS )(,\u03c3\u03c3 && Ju . (20) Remark 1: The convex function (18) is a candidate for term \u03c3V in (14) that makes up part of the Lyapunov function for the input saturated system" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0001660_isam.2011.5942364-Figure1-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0001660_isam.2011.5942364-Figure1-1.png", "caption": "Figure 1. Components made of \u03b2-titanium alloys: a) landing gears, b) rotor head, c) piping and d) springs", "texts": [ " In addition to the excellent corrosion resistance, \u03b2-titanium alloys are comparatively insensitive to embrittlement due to hydrogen generation. The combination of all properties of \u03b2-titanium alloys can be employed for the engineering of lightweight components featuring smaller dimensions and often increased mechanical and chemical properties. \u03b2-titanium alloys are used in the aerospace sector for the production of forged landing gears, rotor heads and fasteners. The oil and gas production utilizes \u03b2-alloys for elements of the drill string and piping, and the automobile industry applies a low-cost-alloy for springs, Fig. 1 [3]. New applications could be realized in engine and powertransmission components of automobiles or in helicopters if hollow \u03b2-titanium-shafts are integrated instead of solid rods. The decrease of moving translatory or rotatory masses will result in a higher overall system efficiency. The intention of this work is to investigate the machinability of a \u03b2-titanium alloy used as the raw material in the manufacturing sequence of a hollow shaft. II. SEQUENCE TO MANUFACTURE A HOLLOW SHAFT MADE OF \u03b2-TITANIUM TI-10V-2FE-3AL Ti-10V-2Fe-3Al is a metastable \u03b2-titanium alloy developed in the 1970s especially for improved forgeability" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0002302_b978-0-12-803581-8.10351-0-Figure17-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0002302_b978-0-12-803581-8.10351-0-Figure17-1.png", "caption": "Fig. 17 Top shell.", "texts": [ " The layup is the basic stacking sequence given above, but the strake is filled with a preshaped structural foam shape and capped with additional layers of prepreg on the inside surface during side shell manufacture, creating a stiff beam structure in the region shaded in red in Fig. 16. The monocoque has two added unidirectional plies in for the suspension box region (shaded in blue in Fig. 16) with fibers running fore and aft, for enhanced front crush protection. This region of the monocoque consists of five layers of plain weave, and two additional layers of unidirectional prepreg surrounding 6 mm (\u00bc00) thick Nomex honeycomb core. 745;745;0; 0290; core; 0290;0;745\u00bd Fig. 17 shows the top shell of the multi-shell monocoque design, which completes the structure of the cockpit rim stiffener, includes the structural sidepod and head restraint structure, and provides the contour for the nonstructural suspension box cover. This shell, except for the suspension box cover, is permanently bonded in place during monocoque assembly. In the regions shaded in red, the bonded sides and top generate a highly reinforced hollow beam stiffener around the cockpit opening as described in previous sections", " 0290; 745\u00bd The rim stiffener recovers strength and torsional rigidity lost due to the hole that is cutout for the driver. The rim stiffener, shaded in red, consists of 11 layers around the cockpit opening. The 0\u201390 and 745 are plain weave and the 0 is unidirectional. 0290;745;0;0;0;0; 0; 0;0;0;0\u00bd In areas where the top does not bond to the side shells of the monocoque, honeycomb core is added to increase the flexural stiffness. The top shell consists of four layers of plain weave surrounding a 6 mm (\u00bc00) thick Nomex core in the region shaded in blue in Fig. 17. This region defines the top portion of the structural sidepod. 0290;745; core; 745;0290\u00bd Fig. 18 shows the bottom shell. For the most part the bottom shell is used to carry the outside skin loads from the left side shell to the right side shell. The region that is not in direct contact with the side shells, creates the lower portion of the structural sidepods and acts as the entry to the underbody diffuser. The bottom shell gives added strength where it is bonded to the monocoque and consists of two layers of plain weave that is permanently bonded to the left and right shells of the monocoque" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0001517_0731684409348345-Figure2-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0001517_0731684409348345-Figure2-1.png", "caption": "Figure 2. Coordinate setting of the novel twin-screw kneader.", "texts": [ " The key to improving the overall performance of the kneader lies in the design, analysis and optimization of the tooth profile of screw rotors. Therefore, geometric analysis, optimization, and tooth profiling of screw rotors of the kneader reactor are discussed in following sections. DESIGN THEORIES OF ROTOR PROFILES The profiles of the novel twin-screw kneader are designed using the design method of twin-screw pump and spatial engaging theory. Generally, the profiles were designed with one or several arcs and cycloids as well as lines. The primitive profiles and coordinate systems of the novel twin-screw kneader are shown in Figure 2. The properties of primitive profiles of the kneader are listed in Table 1. The fixed frames are: ra(oa xa,ya,za), rb(ob xb,yb,zb). at UNIV OF CALIFORNIA SANTA CRUZ on November 25, 2014jrp.sagepub.comDownloaded from Based on gear engagement theory [8], the correct engagement condition between female and male rotor is: n * v *\u00f0 fm\u00de \u00bc 0, \u00f01\u00de where n * is the normal vector at contact point and v *\u00f0 fm\u00de is the relative velocity vector at contact point. The equation of engagement in moving frame rf between female and male rotor can be induced as: f \u00f0uf, f, f \u00de \u00bc x00f cos\u00f0 f f \u00de \u00fe y00f sin\u00f0 f f \u00de kmf Aimf \u00f0x0fx 0 0f \u00fe y0fy 0 0f\u00de \u00bc 0, \u00f02\u00de at UNIV OF CALIFORNIA SANTA CRUZ on November 25, 2014jrp" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0002642_978-3-319-66866-6_2-Figure6-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0002642_978-3-319-66866-6_2-Figure6-1.png", "caption": "Fig. 6. Optimizing the topology for stress reduction (a) Adapting the outer shape (b) Adapting the cross-sectional area", "texts": [ " Due to the high material effort, the concept model \u2018truss structure\u2019 (1) shows a higher component weight compared to the initial model. \u2018Honeycombs\u2019 and \u2018cuboids\u2019 show improved weight savings. However, the stresses partly increase significantly (rmax >> rallowable). The concept model \u2018bamboo\u2019 already shows a synthesis between maximum stresses and component weight and thus is selected for the further optimization. To define a preliminary design, the topology of the concept model \u2018bamboo\u2019 undergoes a rough adaption to the asymmetrical shape originating from the topology optimization (see Fig. 6-a). It can be seen that high stresses occur in the area of the bearing (r 280 N/mm2), which exceed the allowable value rallowable 245 N/mm2. Furthermore, low stresses occur along the neutral axis. In a first step, the areas with low stresses are optimized by removing material according to the optimization results (see Fig. 6-b). Starting from the concept model with a constant cross-sectional area (A:A), various preliminary designs are examined. As a result, a hollow profile with 45\u00b0-surfaces is provided in the outer areas. The less stressed area is substituted by a thin layer without cavities. Due to this modification, material can be saved and manufacturability can be improved by reducing overhangs. After adapting both, the outer shape and the cross-section area, the stress distribution of the preliminary design can be improved, as depicted in Fig" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0000491_ecce.2013.6647287-Figure2-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0000491_ecce.2013.6647287-Figure2-1.png", "caption": "Fig. 2. (a) Equivalent magnetic circuit for inductance calculation, (b) magnetic flux flow generated by the a1 swinding.", "texts": [ " 1, the turn short fault winding makes another fault circuit loop which is composed of the fault resistance Rf, the fault winding inductance, and the fault winding flux linkage. Every winding of the motor is not only electrically connected to each other, but coupled magnetically also. Hence, the cross coupling effects should be considered in the analysis. Since the turn number of the a-phase is reduced by the shorted turn number, the PMSM motor model changes depending on the faulted turn number. To establish the fault inductance model, the magnetic flux path should be considered. Fig. 2 (a) shows an equivalent magnetic circuit of the PMSM, where Rg, Ry1, and Ry2 denote the magnetic reluctance of air-gap and slot, stator back iron, and rotor back iron, respectively. N denotes the turn number of one winding. Since the iron relative permeability is larger than the air relative permeability, Ry1 and Ry2 are much less than Rg. By assuming that Ry1 and Ry2 are zero, the turn short fault model was derived in [18]. But, they are not zero and obstruct the flux flow. Due to these stator and rotor back iron reluctance values, the flux flows by one typical winding would be decrease as the length of the flux path increases. As the PMSM pole number increases, the difference between the maximum and minimum flux path reluctances increases. Hence, the model derived in [18] is not accurate for the PMSM with a high pole number. Fig. 2 (b) shows the magnetic flux flow by a1 winding. Since c3 and b1 are adjacent winding to a1, magnetic flux flows are large. b2 and c2 are the farthest winding from a1. Their magnetic flux flows are small. Since a1 winding has the turn short fault, it is composed of the healthy and faulted windings. The healthy winding has the same current magnitude but the faulted winding has a different current magnitude. Hence, the flux coupling model between a1, a2 and a3 winding has to consider the different current magnitude. The flux coupling of the same phase windings depends on the structure as shown in Fig. 2(b). To define the coupling flux linkage between the same phase winding, we define 12 and 13 which represent the winding coupling factor of a1-a2 and a1-a3 windings in Fig 2 (b). For n windings per each phase, the magnetic flux linkage of the a1, a2\u2026 and an winding can be described as, )1( 1 1 1 )( 2 12/312/3 12/312/3 12/312/3 2 1 21 121 112 2 1 cb m an a a slsm sm n sm n sm nslsm sm sm n sm slsm an a a ii n L i i i LL P L P L P LLL P L P L P LLL where 1a , 2a \u2026, and an denote the flux linkage of a1, a2\u2026, and an winding excluding the rotor flux. ia1, ia2\u2026, and ian denote the winding current, respectively" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0003356_978-981-13-5799-2_12-Figure12.62-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0003356_978-981-13-5799-2_12-Figure12.62-1.png", "caption": "Fig. 12.62 Predicted deformation of snow and the shear stress distribution in the contact patch [84]", "texts": [ " Sipes and the friction coefficient between the tire and snow surfaces are ignored because this simulation is focused on predicting the shear force of snow in a void. The appearance of the predicted snow surface and a comparison with the snow-covered road surface are shown in Fig. 12.61. The track of the tread pattern is observed on the predicted snow surface, which is in good qualitative agreement with a photograph. One advantage of numerical simulation is that we can make visible what we cannot observe in an experiment. For example, the predicted shear stress distribution of a tire on snow can be predicted as shown in Fig. 12.62. The shear stress is mainly generated at the lateral grooves of the tread pattern. A comparison of the shear stress distributions of tires with various pattern designs shows that the new tread pattern can be effectively designed by simulation. (3) Validation of predictability The predictability of simulation was validated by comparing experimental results with the prediction for several basic tread patterns. The tire size was 195/65R15, the inflation pressure was 200 kPa, the load was 4.0 kN, the velocity was 60 km/h, and the slip ratio was 30%" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0000860_icma.2012.6282812-Figure2-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0000860_icma.2012.6282812-Figure2-1.png", "caption": "Fig. 2 Six error components of rotary joint", "texts": [ " So this vision-based kinematic calibration can be done through two steps: firstly, the rotational DOF must be calibrated based on up-vision to make the pose of chip almost be the same after alignment; secondly, the translation DOFs of X-axis and Y-axis must be calibrated based on down-vision to compensate the deviations of pad. A. Kinematic modeling for rotational DOF of pick-andplace robot For a rotational DOF or rotary joint, there are six geometric error components known as three translational errors\u2014two radial errors (perpendicular to the rotation axis) and one axial error (along the rotation axis) and three angular errors\u2014two tilt errors and one angular position error [14]. Shown as Fig. 2, translational errors are denoted by x( ), y( ) and z( ) and angular errors are noted as x( ), y( ) and z( ),where \u201c \u201d and \u201c \u201d represent the translational error and angular error respectively, subscript is the error direction and a letter in bracket represents the position coordinate. As shown in Fig. 2, OXYZ is the absolute reference coordinate frame which is embedded in the stator of rotary joint and C0 is short for it; O1X1Y1Z1 is embedded in the rotor of rotary joint and C1 is short for it. The HTM 0T1 which represents the transformation of C1 with respect to C0 considering the effect of six error components can be derived by multiplying Te and Ti shown as (1)~(3), here Te and Ti are the HTMs represent the transformation of C1 with respect to C0 in the case of no error existing and only erroneous movement existing, respectively" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0002791_b978-0-12-812138-2.00003-9-Figure3.4-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0002791_b978-0-12-812138-2.00003-9-Figure3.4-1.png", "caption": "FIGURE 3.4", "texts": [ " It should be noted that the three-phase stator windings perform the roles of both the armature and field windings of a DC motor. These three-phase windings are placed in the slots that are axially cut along the inner periphery of the iron core as shown in Fig. 3.3. They are displaced from each other by 120 electrical degrees along the periphery and are typically connected in delta for low-supply voltage or in wye for high-supply voltage. All turns in each winding are continuously distributed in the numerous slots spread around the periphery, such that the winding density can be sinusoidal as shown in Fig. 3.4. The purpose of this arrangement is to establish a sinusoidal flux distribution in the air gap when currents flow through them. This winding Structure of a typical induction motor. Stator windings. type is called distributed winding. The distributed winding configuration increases the utilization of the iron core and reduces magnetomotive force (mmf) space harmonics, resulting in a lower torque ripple compared to that of the concentrated winding, in which all coils of the phase winding are placed in one slot under a pole" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0000039_nano.2010.5697860-Figure1-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0000039_nano.2010.5697860-Figure1-1.png", "caption": "Fig. 1. Schematic of the polar flagellar motor of V. alginolyticus. (OM; outer membrane, PG; peptidoglycan, IM; inner membrane.)", "texts": [ " Its diameter of the tip can be controlled by the heating power, load, pulling velocity, etc. Generally, the minimum inner diameter of the tip of the glass nanopipette is ~10 nm [11]. Bacterial flagellum is the locomotive organ in liquid. Each flagellum consists of the helical filament that acts as a propeller extending from the cell body, the basal body embedded in the cell surface, and the flexible hook that connects them [12-15]. More than 20 structural proteins are required for this organelle (Fig. 1). The bacterial flagellar motor is a molecular machine that converts ion-motive force into mechanical force; the energy source of the rotation is the electrochemical gradient of Na+ or H+ ions across the cytoplasmic membrane. So, the rotational speed of flagellar motor can be controlled by controlling the environmental Na+ or H+ ion concentration [9]. In Vibrio alginolyticus, two types of flagellar systems, polar flagella and lateral flagella are used for movement in the same cell, depending on environmental conditions [8]" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0000418_detc2011-47599-Figure1-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0000418_detc2011-47599-Figure1-1.png", "caption": "FIGURE 1. Cable-driven chains with various backbone structures", "texts": [ " setup costs and greater possibilities of cable interference. These efforts were mainly focused on single rigid-bodied cable-driven platforms [1] - [7]. There have also been efforts in force-closure analysis of cable-driven multi-bodied chains. Rezazadeh [8, 9] analyzed a special structure of cable-driven open chains, where each cable can only be attached to one link. The analysis did not take into consideration the cable force coupling effect when cables are routed to the distal segments through the proximal segments (see Fig. 1(a)). As such, the authors recently proposed a novel and systematic methodology based on the reciprocal screw theory to analyze force-closure of cable-driven open chain systems with arbitrary cable routings [10]. This paper will extend this methodology to analyze force-closure of cable-driven closed chains (see Fig. 1(b)). To the best knowledge of the authors, force-closure of cable-driven closed chains has not been addressed in literature. 1 Copyright c\u00a9 2011 by ASME Downloaded From: http://proceedings.asmedigitalcollection.asme.org/ on 02/01/2016 Terms of Use: http://www.asme.org/about-asme/terms-of-use The major contributions of this paper are as follows: (i) a novel and systematic methodology based on the reciprocal screw theory to analyze force-closure of a cable-driven closed chain system, (ii) minimum cable requirement for the general case to achieve force-closure, and (iii) addressing arbitrary cable routing, i" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0000858_s00542-012-1593-y-Figure1-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0000858_s00542-012-1593-y-Figure1-1.png", "caption": "Fig. 1 Scanning electron microscope (SEM) image of bit pattern media (Young 2010)", "texts": [ " Talke Center for Magnetic Recording Research, University of California, San Diego, USA Contact between a slider and a disk was also studied experimentally by a number of researchers. Liew et al. (2000) performed nano-indentation and scratch testing of the recording media to simulate head\u2013disk impacts, and to study magnetic degradation. Liu et al. (2000) showed that erasure can occur even if temperature is below the Curie temperature using laser heating of a recorded magnetic film. They found that the high density magnetic transitions are of high likelihood of being affected by the flash temperature. Bit patterned media (BPM) (Fig. 1), is being investigated presently as a promising new approach to increase the magnetic recording density beyond 1 Tb/in2 (Li and Talke 2009; Li et al. 2010). Since the recording surface of BPM consists of discrete, nanometer-sized bit cells, intermittent contact between bit patterned media and a slider shown in Fig. 2 is more challenging from the point of view of tribological failure than contact between a slider and a conventional \u2018\u2018smooth\u2019\u2019 disk (Ye and Komvopoulos 2003; Gong and Komvopoulos 2004)" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0000608_j.triboint.2010.11.012-Figure3-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0000608_j.triboint.2010.11.012-Figure3-1.png", "caption": "Fig. 3. Perturbed whirl of rigid rotor\u2013aerostatic bearing model.", "texts": [ " Based on the assumptions of compressible, isothermal laminar flow in bearing clearance and that air is a perfect gas, the non-dimensional Reynolds equation of this air film that is formed in the bearing clearance, which is derived from Navier\u2013Stokes and continuity equations can be expressed in two dimensional Cartesian coordinates as @ @y h 3 @P 2 @y \" # \u00fe D L 2 @ @z h 3 @P 2 @z \" # \u00bc 2L @ @y \u00f0Ph\u00de\u00fe4L @\u00f0Ph\u00de @t \u00f01\u00de where D and L are bearing diameter and length; P and h are the nondimensional pressure and the thickness of the film; y and z are the angular and axial coordinates of the bearing, respectively; t\u00bcot is the non-dimensional time, and L is the bearing number. If the journal center leave its equilibrium position (e0,f0) because of a small perturbation, as shown in Fig. 3, the whirl with radial and tangential components can be represented by Re(epest) and Re(e0fpest), respectively. This small journal whirl can induce small perturbations in the film pressure and film thickness, which can also be decomposed into static and dynamic parts. The non-dimensional film thickness and pressure distributions can be expressed as h\u00bc h0\u00feepest cosy\u00fee0fpest siny \u00f02\u00de and P\u00bc P0\u00feepestPe\u00fee0fpestPf \u00f03\u00de where h0 is the static film thickness, which is defined by h0\u00f0y, z\u00de \u00bc 1\u00fee0 cos\u00f0y f0\u00de \u00f04\u00de e0\u00bce0/c and f0 are the eccentric ratio and the attitude angle of the center of the journal at static equilibrium; ep\u00bcep/c is the perturbed eccentric ratio and fp is the perturbed angle" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0000954_s1001-6058(10)60054-6-Figure1-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0000954_s1001-6058(10)60054-6-Figure1-1.png", "caption": "Fig. 1 Computational Grids", "texts": [ " The approach is parallelized using domain-decomposition and message-passing-interface strategies for the platform on PC clusters with 64-bit XEON-E5520. Description: Unsteady flows past RLG, dominated by massive separation and interactioons among different components, are calculated using URANS based on SST model. In addition, the geometry of RLG is also very complex. Then, it\u2019s very difficult to accurately calculate the unsteady flows and airframe noise using CFD and CAA methods. The computational grids contain 55 blocks and the overall cells are 10,772,708, shown in figure 1. The Reynolds number based on the wheel diameter (D, which is equal to 0.4064m) and reference velocity (U0 = 40 m/s) is equal to 106. The unsteady flows are calculated by solving URANS equations with SST turbulence model. Normalized time step \u0394t*=\u0394t\u00d7U/D =0.005, which is corresponding to 5.08\u00d710-5 s for each step in physical time. Then, no instantaneous results, such as instantaneous pressure fluctuations, frequency on some samples, are presented using URANS. Boundary Condition: Uniform inflow is applied to the upstream boundary; Outflow is taken downstream boundary with zero pressure gradients" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0000492_s10035-013-0439-3-Figure2-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0000492_s10035-013-0439-3-Figure2-1.png", "caption": "Fig. 2 State of admissible stress confined within Coulomb\u2019s failure envelope, with Mohr\u2019s stress circle describing the mobilized state of stress along the symmetrical sliding planes (plane 1 and plane 2), whose inclination sloping at the angle of repose is taken to be parallel to the Mohr-Coulomb envelope slanted by the angle of friction", "texts": [ " Depending on the rotating direction of \u03c4r\u03b8 , \u03c8 f and f are found to be: \u03c8 f = \u00b1 (\u03c0/4 \u2212 \u03c6/2) (22) f = \u00b1 (\u03c0/4 + \u03c6/2) (23) It can be figured out that the magnitude of f exactly represents a half-way between the direction of gravity and the sliding direction. States of stress at the occurrence of slippage for material moving to the left-hand side can be drawn on plane 1 in the upper-half of Mohr\u2019s stress circle and those moving to the right-hand side can be drawn on plane 2 in the lower-half of Mohr\u2019s stress circle, as illustrated in Fig. 2. Marais [24] pointed out that though all tractions vanish at the slope surface as given in Eq. (24), the angle (or \u03c8) is not indeterminate and has to be derived from the stress distribution in a nearby thin layer. For the condition of an infinite slope in which the yielding condition is entirely saturated, is wholly taken to be f [23] because the inequality shown in Eq. (16) changes to equality. However, for a sand heap, Eq. (18) implies that traction along the slope surfaces saturates the incipient failure, but that stress in the bulk remains stable", "1 Solution for wedge problem The problem of a semi-infinite loose planar wedge inclined at an angle of repose is illustrated in Fig. 3, where the z axis is chosen to pass through the central plane, symmetrically dividing the wedge into left-hand and right-hand sides. Shear stress on plane 1 represents anti-clockwise rotation and shear stress on plane 2 represents clockwise rotation. The limiting state of stress along both planes can be visualized by superimposing these planes onto Mohr-Coulomb envelopes as shown in Fig. 2. According to Fig. 4, Nadai considered the following boundary conditions for the wedge problem where subscript c denotes the location along the central plane where the major compressive stress lies parallel to the direction of gravity or the z axis. c = |\u03b8=\u03b8c = \u03c0/2 where \u03b8c = \u03c0/2 (34) f = |\u03b8=\u03b8 f = \u03c0/4 + \u03c6/2 where \u03b8 f = \u03c6 (35) Nadai idealized the problem by assuming an angle of the major principal stress direction to be a linear function of an angular coordinate \u03b8 based on two extreme conditions given by Eqs", " The problem of a semi-infinite loose planar valley inclined at an angle of repose is illustrated in Fig. 5, where the z axis is chosen to pass through the central plane, symmetrically dividing the valley into left-hand and right-hand sides. In the same manner as in a wedge problem, the shear stress on plane 1 represents anti-clockwise rotation and shear stress on plane 2 represents clockwise rotation. The limiting state of stress along both planes can be visualized by superimposing these planes onto Mohr-Coulomb envelopes as shown in Fig. 2. According to Fig. 6, Nadai considered the following boundary conditions for the valley problem, where the major axis of compressive stress lies at a right angle to the direction of gravity or parallel to the x axis. Furthermore, the angle of major compressive stress along the surface with sand sliding down on the left-hand side with the same magnitude as the wedge problem but with a minus sign due to clockwise rotation. c = |\u03b8=\u03b8c =0 where \u03b8c = \u03c0/2 (56) f = |\u03b8=\u03b8 f =\u2212 (\u03c0/4+\u03c6/2) where \u03b8 f = \u2212\u03c6 (57) Nadai idealized the problem in the same manner as the wedge problem by treating an angle of the major principal stress as a linear function of an angular coordinate \u03b8 based on two extreme conditions given by Eqs" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0002302_b978-0-12-803581-8.10351-0-Figure1-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0002302_b978-0-12-803581-8.10351-0-Figure1-1.png", "caption": "Fig. 1 Monocoque design evolution.", "texts": [ " The design studies for a successor to this earlier monocoque were focused on two key factors, improved manufacturability and geometric changes to enhance future aerodynamics development. The critical design elements were determined to be: \u2022 improve manufacturability \u2022 design for composites \u2022 maximize torsional stiffness \u2022 increase side impact safety structure \u2022 allow for implementation of advanced aerodynamic features \u2022 lower the center of gravity Based on these design elements, chassis design concepts were developed, evaluated and upgraded. Fig. 1 shows a portion of the early concept development, evolving from upper left to lower right. These designs of Fig. 1 show a number of common features including: \u2022 manufacturability improved with a new multi-shell design concept, \u2022 \u201cbucket\u201d structure behind drivers head \u2013 gives preferential fiber orientation for torsional stiffness around the rear of the cockpit opening, which had been an issue in the previous monocoque configuration, \u2022 side impact improved with widened structure at the base of the sides of the driver area, \u2022 a lowered driver position for improved center of gravity, and \u2022 raised nose allows for improved aerodynamics implementation", " Some of the noteworthy features to be incorporated included structural sidepods, to improve both stiffness and driver side impact safety, and strakes ahead of the sidepods, which were to aid airflow to the cooling system, the underbody aero-features and yield an added degree of side impact safety, as indicated in Fig. 7. The concepts of spatial complexity are useful in any design for manufacture process. The shape complexity classification for a one piece, complete chassis, as shown in Fig. 7, is U7.7 This is the most complex classification for a manufactured part, meaning that it is undercut-U and irregular-7. For comparison, the idealized chassis (Fig. 1) has half the shape complexity rating, T4 (tube, closed one end). Undercut, or reentrant, geometries are very difficult to create using cost-competitive composites tooling and layup techniques. Features in the top (such as the desired single piece cockpit rim stiffener) and side details of the chassis (strakes) drove the design to the use of a top, bottom, and sides. Like the pressure vessel, the top and bottom would overlap the sides to create the total structure (Fig. 8). However, unlike the pressure vessel design, the multi-shell chassis top and bottom do not meet fully half way up the sides, but do retain a large lap-bond area" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0000245_j.jbiomech.2011.06.013-Figure1-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0000245_j.jbiomech.2011.06.013-Figure1-1.png", "caption": "Fig. 1. Definition of angles for a two link system. The inertial parameters for this system are l1\u00bc0.87 m, l2\u00bc0.83 m, lcm1\u00bc0.642 m, lcm2\u00bc0.23 m, m1\u00bc22.94 kg, m2\u00bc53.06 kg, I1\u00bc2.86 kg m2, and I2\u00bc3.26 kg m2. Note center of mass locations are measured from each segments distal joint.", "texts": [ " In the present context the singular values can be considered as the gains of the matrix MIAI, and thus provide an index of the potential for induced accelerations. A two link, two joint system is used here to illustrate the use of the IAI. This two link system is used here to represent the human body during quiet standing. While simple this representation is more complex than other models adopted for examining this task (e.g., Peterka, 2000; Winter et al., 1998). The first joint is the ankle joint and the second joint the hip joint (see Fig. 1). If we examine the effect of Joint 1 on Joint 2 then IAI1,2 \u00bc 1\u00fe m2l1lcm2 cos\u00f0y2\u00de m2lcm2 2\u00fe I2 \u00f07\u00de where, m2 is the mass of Segment 2, l1 is the length of Segment 1, lcm2 is the length from proximal end of Segment 2 to its center of mass, y2 is the orientation of Segment 2 and I2 is the moment of inertia of Segment 2 about its center of mass. For this system the acceleration produced at Joint 2 due to the muscular action at Joint 1 is only dependent on one factor due to segment 1: its length. The relationship between the two joints is more complex when the reverse case is considered; the effect of Joint 2 on Joint 1" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0003399_pierc19010804-Figure17-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0003399_pierc19010804-Figure17-1.png", "caption": "Figure 17. PM and massive conductor in (a) x-direction, (b) y-direction.", "texts": [ " Jc is the current density in bar-shaped volume, and B is the field created by a cylindrical magnet. By representing the coil as four overlapping straight beams (Fig. 16(b)), the magnetic force has been obtained analytically. The force is derived for the bar-shaped volume shown in Fig. 16(c), with volume current density J and dimension (2A\u00d7 2B \u00d7 2C). The coil moves in translation along the z direction. The interaction force between the coil and cylindrical magnet is given by Eqs. (22)\u2013(25). The current in x-direction (Fig. 17(a)) Fy = \u2212 I CB B 4\u03c0 M\u2211 t=1 \u2211 i \u2211 j \u2211 k \u2211 l \u2211 p \u2211 q (\u22121)i+j+k+l+p+q \u03d5n (22) Fz = I CB B 4\u03c0 M\u2211 t=1 \u2211 i \u2211 j \u2211 k \u2211 l \u2211 p \u2211 q (\u22121)i+j+k+l+p+q\u03c8n (23) The current in y-direction (Fig. 17(b)) Fx = \u2212 I CA B 4\u03c0 M\u2211 t=1 \u2211 i \u2211 j \u2211 k \u2211 l \u2211 p \u2211 q (\u22121)i+j+k+l+p+q\u03d5n (24) Fz = I CA B 4\u03c0 M\u2211 t=1 \u2211 i \u2211 j \u2211 k \u2211 l \u2211 p \u2211 q (\u22121)i+j+k+l+p+q\u03c6n (25) The forces are obtained with the intermediate variables which are given by \u03d5n = \u222b\u222b\u222b arctg ( V U WR ) dXdY dZ = R 6 ( U2 + V 2 \u2212W 2 ) + 6UVWarctg ( V U WR ) \u22123U ( V 2 \u2212W 2 ) arctgh ( R V ) \u2212 3V ( U2 \u2212W 2 ) arctgh ( R V )) (26) \u03c6n = \u222b\u222b\u222b log (\u2212V +R) dXdY dZ = 1 36 ( \u221224U3arctg ( V U ) + 12V WR\u2212 18U2W log(R+ V ) \u221218U2V log (R+W ) + 36UV W log (R\u2212 U) + 18UW 2arctg ( UV WR ) + 18UV 2arctg ( UW V R ) +6W 3 log (R+ V ) + 6V 3 log (R+W ) + 24U2V \u2212 6U3arctg ( VW RU ) + 18UV 2arctg ( W V ) +36UW 2arctg ( V W ) + 18UW 2arctg ( W V ) \u2212 54UVW \u2212 2V 3 ) (27) \u03c8n = \u222b\u222b\u222b log (\u2212U +R) dXdY dZ = 1 36 ( \u221224V 3arctg ( U V ) + 12UWR \u2212 18V 2W log(R+ U) \u221218V 2U log (R+W ) + 36UV W log (R\u2212 V ) + 18V W 2arctg ( UV WR ) + 18V U2arctg ( VW UR ) +6W 3 log (R+ U) + 6U3 log (R+W ) + 24V 2U \u2212 6V 3arctg ( UW RV ) + 18V V U2arctg ( W U ) +36V W 2arctg ( U W ) + 18V W 2arctg ( W U ) \u2212 54UVW \u2212 2U3 ) (28) with U = \u03b1+ (\u22121)iA\u2212 (a (t+ 1) \u2212 2j\u0394x1t (t)) V = \u03b2 + (\u22121)k B\u2212 (\u22121)lb (t) W = \u03b3 + (\u22121)pC\u2212 (\u22121)qc \u0394x1t (t)) = ((a (t+ 1) \u2212 a (t))/2 (29) The force acting on the rectangular coil is then equal to the sum of the Lorentz force acting on each of the four volumes" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0000091_icems.2011.6073540-Figure1-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0000091_icems.2011.6073540-Figure1-1.png", "caption": "Fig. 1 Motor drive system and on-board charging system", "texts": [ " PHEV shares the characteristics of both a conventional hybrid electric vehicle, having an electric motor and an internal combustion engine; and all of electric vehicle, also having a plug to connect to the electrical grid. So it takes more advantage from this integrated structure. Due to the battery capacity is not high, the on-board battery charger can satisfy charge requirement of PHEV. The on-board battery charger and the motor driving are the key technologies in PHEV application[2]. Motor drive system and on-board charging system are shown in Fig. 1. The on-board battery charger can reduce the volume, the weight, and is easily to carry with the vehicle. Generally, it has two types: isolated type and non-isolated type. The isolated type is used the transformer to isolate the primary side and secondary side. The non-isolated type means the primary side and secondary share the ground. A boost converter is classical non-isolated type. The integrated converter technology which is embedded components of motor drive system to construct the battery charger circuit is one of key technology for the trend of PHEV development" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0003211_j.apacoust.2018.12.039-Figure15-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0003211_j.apacoust.2018.12.039-Figure15-1.png", "caption": "Fig. 15. Bevel gear measuring cell for the investigation of excitation and noise behavior.", "texts": [ " Both effects were predicted theoretically in the simulation and confirmed accordingly in the experiment. The dominant non-gear mesh amplitudes in the low order range are qualitatively matching as well to the test. The amplitudes of the low order range of the R02 variant are at a comparable level to the first gear mesh amplitude and all below 45 mrad. In contrast for variant N10, the amplitudes below order 46 are higher than the first gear mesh amplitude and exceed 50 mrad in test and simulation. The bevel gear measurement cell used for the investigation at operational conditions is shown in Fig. 15. As in [4], the fixture consists of a modular frame structure on which the bearing plates are assembled. An adjustment of the hypoid offset is enabled by additional bearing plates for the pinion side. The basis is built by the stiff frame structure which is milled out of solid material. To achieve a high positioning accuracy within the range of automotive applications, the bearing plates are ground in assembled and bolted condition. The measurement of the differential rotary acceleration is realized by an acceleration measurement systemwith a telemetric system on the pinion and the gear shaft which is available at the WZL as used in [4]" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0001140_s1672-6529(13)60226-7-Figure3-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0001140_s1672-6529(13)60226-7-Figure3-1.png", "caption": "Fig. 3 Mass center and body-centered coordinate system of the Chinese mitten crab. L represents the rows of the leading legs. T represents the rows of the trailing legs. The dashed arrows point to the direction of the velocity.", "texts": [ " This signal was also used as the starting point to analyze the video images in the motion analysis system. The effective area was calibrated using a 0.6 m \u00d7 0.4 m \u00d7 0.2 m calibration frame with 16 non-colinear points that approximately filled the overlapping region of the four cameras before the experiments. The calibration frame was recorded by the four cameras and was then removed from the effective area. The global coordinate system X-Y-Z was defined based on the result of the calibration analyzed through 3D motion analysis (see Fig. 2). Fig. 3 shows the body-centered coordinate system x-y-z located at the MC, the x-axis along the body symmetrical axis toward the heads of the crab, the y-axis laterally to the left, and the z-axis perpendicular to the xy-plane. The location of the MC on the dorsal carapace for each crab was determined by suspending the crab from strings attached to two chelas[2,18]. We recorded the crab while suspended from each location toward the dorsal carapace and defined the extended line of the string in each image using the Adobe Photoshop SC5 Extended software. The MC lies at the intersection of the extended lines from the overlapped areas. A high-contrast white marker was placed at the location of the MC to facilitate digitization of the video recordings (Fig. 3). The shutters of the four high-speed cameras were pressed by four trained technicians after the calibration to ensure that the cameras would begin recording at roughly the same time. A Chinese mitten crab was then released from the edge of the effective area and was allowed to move freely around the effective area. Once the crab moved in a straight path and performed an even sideways gait in the effective area, the flashlight was triggered to produce a flash signal which was captured simultaneously by the four high-speed cameras" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0000413_j.tsf.2012.05.006-Figure3-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0000413_j.tsf.2012.05.006-Figure3-1.png", "caption": "Fig. 3. Schematic of the experimental set up used for the determination of the specific conductivity.", "texts": [ " Specific conductivity measurements We obtained specific conductivity calculations by basing on voltage/ current (V/I) characteristics measured with an electrometer Keitley model 6517, driven by computer. For thesemeasurements, we deposited 1, 10, and 30monolayers of nanocomposites onto glass substrates in both undoped and doped formsbyusing themethods of deposition previously described and then contacting the fabricated device to the electrometer by means of silver wires and silver paint. We carried out measurements of current by applying a potential ranging between \u221210 V and 10 V. Fig. 3 illustrates the experimental set up used for the determination of specific conductivities. The analysis of UV\u2013vis spectra, aswill be described in the next section, showed the syntheses issued nanocomposite containing conducting polymers in the EB form. It is important to point out that during the polymerization process (NH4)2S2O8 was able to oxidize only the monomer, since the temperature and the strength of the oxidizing agent were not capable to operate an oxidation of the CNT surface, as widely demonstrated in our previous work [16]" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0000397_00368791111140495-Figure4-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0000397_00368791111140495-Figure4-1.png", "caption": "Figure 4 Illustration of bearing type", "texts": [ " The stability condition of rotor whirl is dependent on nondimensional mass, which is stable as M is smaller than Mcr . On the other hand, when a rotor rotates with a speed v, self-excited vibration with the frequency Vc \u00bc gcv can occur as M of this rotor is larger than Mcr . 5. Results and discussion This study will analyze the static, dynamic characteristics, and whirl instability of a rigid rotor supported by two identical aerostatic bearings with double-array entry holes as shown in Figure 1. The ratio of bearing length to diameter L/D is 1.0 and land width of axial flow a equals L/4. As shown in Figure 4, each array comprises six or three entry holes which are arranged symmetrically around circumference surface of the bearing. For the number of n \u00bc 6, the entry holes are located at ur \u00bc 08, 608, 1208, 1808, 2408 and separately equally. Two types of arrangement of entry holes for number of three are taken into consideration, type 1 (one hole on Inherent restriction on stability of rotor-aerostatic bearing system Cheng-Hsien Chen et al. Industrial Lubrication and Tribology Volume 63 \u00b7 Number 4 \u00b7 2011 \u00b7 277\u2013292 the top) is arranged by locating three holes at ur \u00bc 08, 1208 and 2408, and type 2 (one hole on the bottom) is arranged by locating three holes at ur \u00bc 608, 1808 and 3008, respectively" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0001609_0954406211427833-Figure8-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0001609_0954406211427833-Figure8-1.png", "caption": "Figure 8. Comparison of analytical and simulated tooth profiles for a concave gear: (a) superimposed profiles and (b) deviation of the simulated profile with reference to the analytical one.", "texts": [ " For machining the teeth which are at the outer ring, the work solid is tilted by 24 about the X-axis with respect to the sphere centre (Figure 6(c)). Similarly, for simulating the other teeth at the outer ring, initially, the work solid is first indexed by 30 about the Z-axis and then tilted by 24 about the X-axis. The same steps are followed for the rest of the teeth at the outer ring. CAD model after simulation for the centre tooth of the concave gear is shown in Figure 7. By highlighting the tooth surface, the curvature on it can be seen distinctly. Figure 8(a) shows the CAD-simulated profile obtained by taking an axial section of the centre tooth along with the superimposed profile obtained analytically following the procedure outlined in \u2018Analytical method of tooth surface generation\u2019 section. Since the difference cannot be seen clearly, the deviations in the X-coordinate values are plotted along different Y-coordinate values, taking the analytical profile as a reference. Therefore, the analytical profile is represented as a vertical line and difference in the X-coordinate values are plotted as in Figure 8(b). It may be seen that the maximum deviation of 17.1mm occurs at the top. This is not to be treated as a limitation of CAD approach, as the simulation in this study is carried out using 320 iterations with simultaneous motions. Generation of tooth on convex gear Initially, the cutter is placed in such a way that the pitch line of the cutter lies at a position tangential to the pitch at University of Ulster Library on May 14, 2015pic.sagepub.comDownloaded from circle diameter of the gear blank and the gear blank axis makes 6 with the cutter axis, as shown in Figure 9(a)" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0003716_6.2019-4392-Figure8-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0003716_6.2019-4392-Figure8-1.png", "caption": "Fig. 8 Printed CAD Model of Combustion Chamber", "texts": [ " To find out more about the build plate removal methods used on the Balerion engine, refer to Section VI Subsection C Build Plate Removal. Additionally, for all parts, text engraving was used for not only organizing the nomenclature of all sensor ports and their measuring purposes (pressure, temperature, etc.), but also to mark alignment features for added reliability during engine assembly. Reference Section VII, Subsection B Engine Assembly for more on the assembly aspects. Specifics about each component is broken up in the sub-sections below. The chamber was printed in flange up direction, as seen in Fig. 8, due to two primary reasons. One was to avoid having the internal geometry of the chamber need any support material as that would be difficult to remove and the other was the design of the injector. The fuel annulus was pre-designed into the chamber and serves as the transfer point for the fuel traveling up the regenerative channels to go into the combustion chamber to mix with the oxidizer and be ignited. Since the fuel travels from the constant velocity manifold through 150 separate channels, it needs to be re-combined into a centralized manifold before being injected into the chamber to ensure an equalized sheet of fuel is hitting the oxidizer for optimal combustion" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0003190_978-3-030-05321-5_8-Figure8.29-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0003190_978-3-030-05321-5_8-Figure8.29-1.png", "caption": "Fig. 8.29 Cross-sectional view of the proposed three-way valve consisting of two orifice plates oscillated by piezoelectric transducers", "texts": [ " A three-way valve is a component that has an inlet and outlet for a fluidic actuator. A three-way valve to control hydraulic artificial muscle type actuators has been designed and fabricated. As the valve is small, it can be built into the actuator unit easily. In this three-way valve, the supply and drain of the working fluid is switched by using two piezoelectric transducers installed in each port [16]. The cross-sectional view of the prototype of the proposed three-way valve is illustrated in Fig. 8.29. The actuator port is connected with a fluidic actuator. The inlet and outlet ports have orifice plates for the supply and drain of the working fluid. In each port, particles are excited by each orifice plate to control the flow rate of the port. As the two orifice plates are of different sizes, they have different natural frequencies. As schematic of a transducer is shown in Fig. 8.30, where the structure is based on a bolt-clamped type piezoelectric vibrator. Ring-typePZTplates are used for the oscillation, and the electrodes are made of copper" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0000604_amm.397-400.126-Figure5-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0000604_amm.397-400.126-Figure5-1.png", "caption": "Fig. 5 Temperature field distribution of inner ring Fig. 6 Temperature field distribution of outer ring", "texts": [ " Temperature field simulation and analysis of the angular contact ball bearing Thermal load and boundary conditions are shown in table 1. Surrounding medium around angular contact ball bearing was air, and its temperature was the inlet temperature of compressed air, 1 20 a T = \u2103. The loads and convective heat transfer coefficient were applied. Fig.3 shows the finite element model of the bearing. Bearing temperature field distribution is shown in Fig.4. Fig.3 Finite element model of bearing Fig. 4 Temperature field distribution of bearing The temperature field distribution of bearing inner ring is shown in Fig.5. The temperature field distribution of bearing outer ring is shown in Fig.6. From Fig. 4, Fig.5 and Fig. 6, the friction can be created between ceramic ball and inside and outside the ring of ball at high-speed, which leads to repeated heating. The ceramic ball is inside bearing, and the contact area of ceramic ball connected with lubricating oil is small, so the convective temperature is 47.259\u2103. The convection heat transfer coefficient of lubricating oil is bigger due to the larger cooling area of inner ring end face, and the temperature of inner ring end face is the lower. The cross section temperature curve of the bearing inner ring is shown in Fig" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0002967_iecon.2017.8216659-Figure1-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0002967_iecon.2017.8216659-Figure1-1.png", "caption": "Fig. 1. Sectional view of 12/9-pole unit DSEM motor", "texts": [ " When one phase open-circuit fault happens, the bilateral switch connected to the faulty phase is controlled to switch on and the FTCA of the remaining three phases are adjusted to produce the same excitation torque as that of normal operation. The stator copper loss expression of FTCA is provided. The relationship between average output torque, output torque ripple and FTCA are analyzed through co-simulations of Maxwell and Simplorer. At last, we obtain an optimal FTCA range, which ensures high output torque, low output torque ripple and low stator copper loss. II. FOUR-PHASE 24/18-POLE DSEM MOTOR Four-phase 24/18-pole fault-tolerant DSEM motor has the same structure as 12/9-pole unit DSEM motor, whose sectional view is shown in Fig.1. Counterclockwise is defined as the positive rotation direction. The stator and This work is sponsored by the National Natural Science Foundation of China(51477075); Jiangsu Province Science and Technology Supporting Project for Industrial Applications(BE2014136). rotor are both in the shape of salient poles. The stator has concentrated windings on poles. Armature windings of phase B are shown in Fig.1 as an illustration. Phase A, C and D has the same winding structure as B. Each excitation winding spans three stator poles. The rotor has a simple structure without any windings or permanent magnets. We define rotor position angle to be zero degree of the position in Fig.1. Phase B has the maximum and minimum flux linkage when =0\u00b0 and =20\u00b0, respectively, which are shown in Fig.2(a) and Fig.2(b). We suppose that magnetic saturation effect and edge effect is neglected. Then the self-inductance Lp (p=a, b, c, d) of phase P (P=A, B, C, D) and mutual-inductance Lpf of phase P and excitation winding are considered to be piecewise linear functions of rotor position [1] which are shown in Fig.3. According to the variation law of four phase inductances, one electrical period is divided into four stages, which are S1, S2, S3 and S4 in Fig" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0000449_amr.805-806.911-Figure1-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0000449_amr.805-806.911-Figure1-1.png", "caption": "Fig. 1 Three models of people under HVDC Lines", "texts": [ " To analyze this phenomenon, a simulation circuit model is set up. The simulation results agreed well with measured values and the simulation circuit model is verified. Models of transient electric shock People or conductors with good insulation in the ionized electric field will gain a certain charging potential due to ion implantation. The potential depends on the grounding resistances, for the people, that is the resistance of the shoes. The equivalent circuit of people under the HVDC transmission lines is shown in Fig.1(a), where Rp is the insulation resistance. The insulation resistance of the shoes is usually between 3M\u2126 and 100M\u2126, hardly more than 10 10 \u2126 [3]. Cp is human-body-to-ground capacitance, typically 100pF. Up stands for the charging voltage. The severity of the transient electric shock could be evaluated by the energy stored in the capacitance. Transient electric shock will emerge when people touch the grounding objects (Fig.1 (b)) or people touch the objects with large grounding resistances (Fig.1 (c)). Relation between electric potential on human body and electrostatic stoking level was listed in [8].When Up=3kV, there is slightly hurt with a feeling of acupuncture. When Up reaches as much as 12kV, there is a quite uncomfortable feeling with the whole hand has a strong strike. All rights reserved. No part of contents of this paper may be reproduced or transmitted in any form or by any means without the written permission of Trans Tech Publications, www.ttp.net. (ID: 130.88.90.140, The University of Manchester, Manchester, United Kingdom-15/05/15,19:18:09) Experiment of the transient electric shock The measurement devices of the transient current waveform are shown in Fig" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0001687_978-90-481-9689-0_68-Figure3-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0001687_978-90-481-9689-0_68-Figure3-1.png", "caption": "Fig. 3 Singularity loci for the Geometry 1.B of the 3-UPU TPM.", "texts": [ " Analytically, it can be proved that a singularity locus is a right cylinder \u0393 [5], with circular directrix \u03b3 and axis coincident with the z axis of Sb. Therefore, conversely, once defined the points O\u2032 p, O\u2032\u2032 p and O\u2032\u2032\u2032 p , the circle \u03b3 is defined and the cylinder \u0393 is defined too. The three points can be easily found by geometrical inspections thus representing a simple and efficient method to easily find the cylinder \u0393 . This cylinder has radius r = 2(b\u2212 p). In order to investigate the influence of the legs location on the singularity of the manipulator, a second geometry defined as Geometry 1.B, is shown in Figure 3. This geometry is obtained by disconnecting the platform of the Geometry 1.A from the legs and rotating it 180\u25e6 about the z axis of Sb, which is defined as in the previous geometry, then connecting again the legs to the same corresponding platform revolute pairs. This makes the three legs intersect at one point, which is a practical drawback since the legs cannot penetrate each other. However, suitable manufacturing solutions can overcome it. Indeed, some efficient manufacturing solutions are presented in the next section" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0003525_s42417-019-00149-6-Figure2-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0003525_s42417-019-00149-6-Figure2-1.png", "caption": "Fig. 2 Bearing design parameters [11]", "texts": [ " These frequencies are in an arbitrary order, so the optimization algorithm contemplates a step to accommodate them from lower to higher to find the difference between consecutive frequencies. Design Variables The design variables correspond to those parameters that the optimization algorithm will vary to find the best distribution of the excitation frequencies. According to (2)\u2013(5), the excitation frequencies of the bearings are functions of their design (2) ir = N 2 [ 1 + d D cos ( ) ] i, (3) or = N 2 [ 1 \u2212 d D cos ( ) ] i, (4) c = 1 2 [ 1 \u2212 d D cos ( ) ] i, (5) re = D d [ 1 \u2212 ( d D cos ( ) )2 ] i. (6) 1 = , (7) 2 = 1 r , (8) gm = 1 \u00d7 N1 = 2 \u00d7 N2. parameters (Fig.\u00a02). However, to obtain a real design these parameters are taken from a database built from Timken ball bearing catalog [14]. The data of the bearings were classified according to their catalog number, which is identified as b . Thus, the input bearings (pinion) are denoted as b1 and the output bearings (gear) as b2 , and each one associates the four design parameters. For example, if the optimization algorithm selects the fourth element of the database for the bearing b1 , it corresponds to the design parameters described in Table\u00a01" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0003594_j.ergon.2019.06.010-Figure1-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0003594_j.ergon.2019.06.010-Figure1-1.png", "caption": "Fig. 1. Workspace arrangement in the experiment.", "texts": [ " Participants with the experience for playing this baseball game in a Nintendo 3DS were not allowed to join the experiment. The convenience sampling was used for recruiting participants. A TOPCON ACP7 vision tester and Standard Pseudo Isochromatic charts were used to test participants\u2019 visual acuity and colour vision. A Nintendo 3DS LL handheld console (Nintendo Co., LTD., Kyoto, Japan) used a parallax barrier with glasses-free technology for 3D imagery. Konami \u9b42\u30d7\u30ed\u91ce\u7403\u30b9\u30d4\u30ea\u30c3\u30c4 2011 was the game cassette used for the console. The experimental task arrangement is shown in Fig. 1. The handheld console was positioned on a 73-cm-tall table. The front edge of the table was 20 cm from the display centre. The inclination angle of the mobile display was 105\u00b0 from the vertical axis (Turville et al., 1998). The viewing distance was 25 cm, and the subject's head was restrained by a chinrest 15 cm above the table. Before conducting the experiment, participants could adjust the height of seat to make themselves comfortable. The dependent variables include hit, miss, false alarm, and correct rejection rates, \u03b2, d\u2019, ROC space, the Simulator Sickness Questionnaire (SSQ), and the iGroup Presence Questionnaire (IPQ), as detailed below: (1) Hit rate: The pitcher pitches a strike (left side of Fig" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0000022_j.jappmathmech.2012.05.003-Figure3-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0000022_j.jappmathmech.2012.05.003-Figure3-1.png", "caption": "Fig. 3.", "texts": [ ", the constraint eaction is specified uniquely, only in the first of these cases. In cases b and d the positive energy is unbounded below, indicating a shock or any initial values of the reaction. Such a shock is also possible in cases c and e along with the stable value of the reaction in the potential ell determined from formula (4.6) or (4.7), respectively. xample. 2,10Two heavy point masses M1 and M2 with unit masses are joined by a rigid weightless rod of length l. The point M1 can slide n a rough tube OX with a coefficient of friction (Fig. 3). As generalized coordinates we take the x and y coordinates of the point M1 (the OY 152 A.P. Ivanov / Journal of Applied Mathematics and Mechanics 76 (2012) 142\u2013 153 a o m k W a a e b d xis is directed vertically downward) and the angle between the rod and the OX axis (which is measured clockwise). When the equations f motion are written, the system is released from the constraint y = 0, which is replaced by the sum of the normal reaction N (which is easured in the downward direction) and the friction force F" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0000900_ever.2013.6521527-FigureI-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0000900_ever.2013.6521527-FigureI-1.png", "caption": "Fig. I. Rotor slot angle. the first three slots are considered. It is cq = ar /2, al = 3ar /2 and al = 5ar /2, where ar = 27r / Qr is the slot pitch angle in mechanical degrees.", "texts": [], "surrounding_texts": [ "All FE simulations are carried out in the rotor ftux reference frame, imposing both rotor and stator currents. The rotor currents are adjusted iteratively so as to achieve the rotor ftux linkage only along the d-axis. In the simulation of no-load operations, only d-axis stator current isd is imposed (isq = 0), and therefore, from (1) and (4), there are no currents in the rotor. The corresponding stator and rotor ftux linkages have only the d-axis component. The ftux lines, achieved from the FE simulations, are symmetrical and do not cross the d-axis. The resulting ftux linkage vector --c+ A s is along the d-axis as weIl. In a similar way, the stator ftux along the q-axis is com puted, imposing only the stator current component isq (i.e. isd = 0). The rotor current component isq is achieved from (4). Then, since Asd = 0 the ftux Iinkage As = Asq is achieved from (6). The d- and q-axis stator ftux linkages versus the d- and q-axis stator current are reported in Fig. 2. In the simulation under load, a the stator q-axis current isq is forced in the stator winding besides the d-axis current isd. A corresponding rotor current irq has to be assigned in the rotor. A preliminary estimation of such a current is achieved by me ans of (4), assuming a perfect coupling, i.e., LrLs ':':' LXi' The actual field solution is achieved iteratively. At each step, the rotor flux Iinkage is computed from the field solution. If the q-axis component Arq is different from zero, the q-axis rotor current irq is modified in such a way to satisfy the relationship Arq = 0 (i.e., the FO condition). The convergence process is quite rapid, requiring only a few iterations (typically three) , since the magnetic circuit is linear along the q-axis (the flux approaches zero). The flux Iines under load are shown in Fig. 3. Let us remark that, since both stator and rotor currents are imposed as the field sources, only a magnetostatic FE analysis is necessary also for the simulation under load, so as the saturation effects are careful taken into account. Once the field problem is solved, the motor performance can be computed. The stator flux Iinkage components, Asd and Asq, are computed by means of the magnetic vector potential. The inductances are computed from the rotor flux equation, and the rotor resistance is computed from the rotor Joule losses. From the stator flux Iinkage, the terminal voItage is determined as a function of the operating rotor speed, adding the 3-D effects: end-winding leakage inductances and stator resistance. Fig. 4 shows the vector diagram of the IM. The motor torque is computed from Maxwell's stress tensor integrated along the rotor periphery, directly from the field solution. Alternatively, the torque can be computed from the flux linkages as 3 Tim = 2PArdirq (11 ) which yields a satisfactory prediction of the average motor torque [20]. III. SYNCHRONOUS RELUCTANCE MOTOR In the REL machine, the d - q reference frame is determined by the rotor geometry. The d-axis commonly corresponds to the higher permeance path. The vector diagram of the REL motor is show in Fig. 5. The stator resistance and the 3-D effects are not included. According to the d - q axis currents id and iq, the flux Iinkages are expressed as: Ad = Ldid Aq = Lqiq (12) (13) where Ld is the d-axis inductance and Lq is the q-axis inductance. The ratio between the two inductances defines the saliency ratio of the REL motor, that is, \ufffd = Ld/ Lq\u2022 From the vector diagram, it is possible to obtain the follow equations: (14) where 0:; and o:\ufffd are the current vector and flux Iinkage vector angles, respectively, and 'Pi is the inner power factor angle. Since (15) the power factor angle is related to the current vector angle and the motor saliency. The motor torque can be computed from Maxwell's stress tensor or from the flux linkages as 3 Tre1 = \"2P (Adiq - Aqid) (16) The FE analysis is carried out, referring to the same size considered for the IM above. In particular, the same stator geometry is considered, together with the stator winding distribution. The flux lines are shown in Fig. 6 In each FE analysis the stator currents id and iq are fixed. They are transformed through Park's transformation in the actual stator currents ia, ib and ic, which are assigned within the slots, according to the winding distribution. From the magnetic field solution, the REL motor performance are computed, by means of relations defined from the vector diagram of Fig. 5. The d- and q-axis stator flux linkages versus d- and q-axis currents of the synchronous REL motor are shown in Fig. 7. Comparing with the results in Fig. 2, it is worth noticing that the behaviour of the main flux linkage (i.e., the d-axis flux Iinkage) remains the same. IV. PM ASSISTED RELUCTANCE MOTOR According to the axis notation given for the REL motor, a permanent magnet (PM) is added so as to produce a PM flux Iinkage Am in the negative q-axis. Therefore, the flux Iinkages become: Ad Ldid Aq Lqiq - Am The corresponding vector diagram is shown in Fig. 8. (17) (18) The advantage of using a PM in the rotor is twofold. At first, a part of the PM flux saturates the iron bridges of the rotor. Secondly, the PM flux linkage tends to reduce the angular distance between the current vector angle and the voltage vector angle, that is, it tends to reduce the power factor angle. As a consequence, there is a slightly higher torque for the same current, with a corresponding increase of the motor efficiency. In addition, a noticeable increase of the power factor is achieved, that implies a reduction of the inverter volt ampere ratings. Fig. 10 shows the flux Iinkages versus current. Two PMAREL motors are considered with different volume of PM. V. COMPARISON" ] }, { "image_filename": "designv11_62_0000786_aero.2013.6496960-Figure7-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0000786_aero.2013.6496960-Figure7-1.png", "caption": "Figure 7 \u2013 Subunit inner structure", "texts": [ " As seen, the excavation velocity was fast for the first min and then maintained at an almost constant average velocity until approximately 2.5 min, where it slowed to nearly zero shortly afterwards. The robot excavated to a depth of 247 mm. The maximum motor torque was approximately 18 Nm, so the propulsion unit needs to maintain its body position orientation against this torque. Development of a Propulsion Unit with Sensors Figure 6 shows the structure of a subunit. The subunit was equipped with two stepper motors, two ball screws and dual pantographs (Fig. 7). The stepper motors and ball screws controlled things like the contraction and extension of plates. As the subunit contracted, the dual pantograph extended in a radial direction, and as the subunit extended, the dual pantograph contracted in a radial direction (Fig. 8). The dual pantograph can push in the parallel direction against the wall. In addition, the expansion plates contained large, circular arc areas to maintain contact with the wall surface, which holds the body position against the rotation action of the EA" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0000789_ast.77.124-Figure4-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0000789_ast.77.124-Figure4-1.png", "caption": "Fig. 4: Thermotropic switching by phase transition in core/shell particles.", "texts": [ " A more systematic investigation on phase transition systems with adequate switching described the preparation of several thermotropic films by embedding of a thermotropic additive in UV cured resin [18]. Only moderate switching ranges of the normal-hemispherical solar transmittance (\u2206Tnh < 9 %) were reported. Optimized materials exhibited an improved performance with switching ranges of up to 16 % [8]. Further improvements could be achieved by the utilization of thermotropic additives as core/shell particles (Fig. 4) [9]. This approach is characterized by the following advantages: - All added particles contribute to the thermotropic behavior. - The amount of thermotropic additive can be reduced in comparison to the application of unprotected domains. - The size of the particles and their distribution can be adjusted to result in optimized scattering behavior and improved performance of the thermotropic sun protecting assembly. - A polar modified shell prevents the diffusion of the core materials. As an overall result, the thermotropic performance and the long-term stability is essentially enhanced. The challenge of the approach consists in the preparation of core/shell particles possessing at all temperatures below the switching point a matching of the refractive indices of polymer matrix, particle shell and particle core to ensure a high transmittance in the off-state of the thermotropic layer (Fig. 4). Thus, the switching results only in a significant reduction of the refractive index of the particle core, and the incident light is predominantly scattered at the core/shell interface. Accordingly, the core diameter of the particles determines the scattering characteristics. Exploiting the appearing refractive index difference of ~ 0.07 between particle and matrix at elevated temperatures, the largest difference in the normal-hemispherical transmittance (\u2206Tnh ~ 28 %) can be found with a particle amount of 6 % and a median scattering domain diameter of ~ 380 nm" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0001248_epe.2013.6634388-Figure14-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0001248_epe.2013.6634388-Figure14-1.png", "caption": "Fig. 14: Expanded view of prototype gear. Fig. 15. Eddy current loss density distribution on the aluminum.", "texts": [ " The efficiency of the prototype gear is lower than the calculated value obtained from FEA. In order to discuss the reason, the no\u2013load loss is indicated below. Fig. 13 shows the mechanical input of the prototype magnetic gear at no\u2013load. It is understood that the no\u2013load input obtained from the experiment is significantly larger than the one obtained from 3D\u2013FEA, and that the no\u2013load input increases nonlinearly as the inner rotor speed increases. The main cause is the eddy current loss in the stator housing made from an aluminum alloy to hold the pole pieces. Fig. 14 shows the expanded view of the prototype gear around a pole piece. The pole pieces are held by the aluminum alloy. Therefore, as shown in Fig. 10, the leakage flux passes through the aluminum, and then the eddy current is induced in it. The eddy current loss of the aluminum of the prototype magnetic gear is estimated by 3D\u2013 FEA. The electrical conductivity of the aluminum is 3.76\u00d7107 S/m. Fig. 15 shows the eddy current density of the aluminum on the r\u2013\u03b8 plane. The figure reveals that the eddy currents are induced around the pole pieces" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0000873_vppc.2010.5729109-Figure1-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0000873_vppc.2010.5729109-Figure1-1.png", "caption": "Fig. 1. example flux dispersion of electromagnetic machines with rotating magnetic field, here asynchronous machine", "texts": [ " The main difference is the air-gap rotor flux: In case of the ISC for asynchronous machines, the air-gap rotor flux results from the air-gap stator flux and rotor currents, induced by the relative rotation of the rotor with respect to the air-gap stator flux. In contrast the air-gap rotor flux of the permanent-magnet synchronous machine (PMSM) is defined by the rotor permanent magnets. The principle of flux dispersion of all electromagnetic machines with rotating magnetic field can be displayed based on the example of an asynchronous machine (cp. fig. 1). This physical similarity gives the basis for a common control principle for such machines. The difference between the airgap flux of the rotor \u03a8\u03b4,rotor and the air-gap flux of the stator \u03a8\u03b4,stator defines the actual torque as well as the field weakening of the machine. In the special case of a PMSM the direction of the main component of the rotor flux is equal to the actual orientation angle of the rotor. The idea of the flux-based control of the PMSM is to directly compute the necessary stator flux of the machine to realise the desired torque and (if needed) the field weakening" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0002601_gt2017-64123-Figure19-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0002601_gt2017-64123-Figure19-1.png", "caption": "Figure 19. Distribution of fatigue life in the optimized impeller", "texts": [ " Gas-dynamic and strength characteristics of the optimized impeller Figure 16 shows the velocity field distribution for the middle diameter of an optimized impeller. Figure 17 shows the equivalent von Mises stress distribution in the optimized structure, normalized relative to the maximum stress in the original design. The distribution of the long-term strength factors in the optimized full-length and short blades is shown in Figure 18. The fatigue life distribution of the optimized structure is shown in Figure 19. A Campbell diagram of the optimized structure is shown in Figure 20. The natural frequencies and harmonics values are normalized relative to the first natural frequency. 8 Copyright \u00a9 2017 ASME Downloaded From: http://proceedings.asmedigitalcollection.asme.org/ on 08/20/2017 Terms of Use: http://www.asme.org/about-asme/terms-of-use The optimized impeller flow around the blade leading edge is acceptable. At the same time, the vortex occurring near the impeller outlet is reduced and shifted to the periphery, which benefits the useful work of the impeller" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0002176_j.procs.2017.01.179-Figure6-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0002176_j.procs.2017.01.179-Figure6-1.png", "caption": "Fig. 6. Arrangement of wires and rods (Rod 1 is fixed on the reference frame).", "texts": [ " The following presumptions are made: 1. The target system is under the nongravity condition, and the friction among mechanical parts can be ignored. 2. The twist around the longitudinal direction of a rod can be ignored. Therefore, each rod has 5 degrees-of-freedom (DOF): 3 DOF (x, y, z) for a translation motion and 2 DOF (\u03b8, \u03c6) for a rotational motion. 3. One of the three rods is fixed at the reference frame \u03a30. The system can control the positions and orientations of the other two rods; there are totally 10 DOF (See Fig. 6(a)). 4. Thicknesses of the wires and rods are ignored. Three wires are ideally connected at each tip, as shown in Fig. 6. Any collisions among the wires and rods are not considered. Owing to the redundancy of a wire-driven system [8], the active controllable number of DOF is eight with the use of the nine wires: 9 \u2212 1 = 8. Based on the abovementioned assumptions to fix rod 1, this system can control 10 DOF motions of rods 2 and 3. Form the viewpoint of the number of actuators, this system with nine actuators can actively control the 8 DOF of rods 2 and 3. Note that the remaining 2 DOF cannot be controlled actively but can be determined, depending on the actively controlled 8 DOF. 3.2. Balancing internal force This section presents the numerical calculation method for the balancing internal force. Let us consider the tensegrity robot, as shown in Fig. 6. The position and orientation of the i-th rod are defined as ix = ( i x, iy, iz, i\u03b8, i\u03c6)T. As rod 1 is fixed at the reference frame, 1x = const. at any time. The position and orientation vector of the three rods, x, is defined as x = ( 1xT, 2xT, 3xT)T. In addition, the force and moment vector on the i-th rod i f \u2208 R5, which is generated by the attached wires is defined as, i f = iw1(x)i\u03b11 + iw2(x)i\u03b12 + . . . + iw6(x)i\u03b16, (2) where i\u03b1 j ( j = 1, . . . , 6) indicates the values of the j-th wire tension on the i-th rod; the value is positive when the wire transmits the tension", " STEP 6: After changing the values of dependent variables (3\u03b8\u0302d and 3\u03c6\u0302d) and ik discretely within the target ranges, repeat from STEP 4. Note that the solution of the internal force and dependent variables is not necessarily unique for a set of independent values. In other words, a set of independent variables can possess plural solutions of the internal force according to 3\u03b8\u0302d, 3\u03c6\u0302d and ik. To verify the proposed method, an example is demonstrated in which each rod is 2 [m] long. Furthermore, the local frames \u03a3i are fixed on the centers of gravity on the rods, as shown in Fig. 6. The position of a rod is defined as the position of the center of gravity. Rod 1 is fixed on the reference frame; thus, its position and orientation are unchangeable. The rod angles i\u03b8d and i\u03c6d are defined as Roll and Yaw angles. In this case, each vector ik is constant Table 2. Example of solutions of balancing internal force. v1 v2 v3 v4 v5 v6 v7 v8 v9 0.43 0.43 0.43 0.43 0.43 0.43 1.0 1.0 1.0 Table 3. Example of resultant values of independent angles. \u03b8\u0302d [deg] \u03c6\u0302d [deg] Rod 3 315 50 A B C D E F Rod 1 Rod 2 Rod 3 (a) Isometric view" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0001433_pi-a.1962.0130-Figure3-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0001433_pi-a.1962.0130-Figure3-1.png", "caption": "Fig. 3.\u2014General arrangement of cascade-type generator.", "texts": [ " For example, missiles FORD: BRUSHLESS GENERATORS FOR AIRCRAFT\u2014A REVIEW OF CURRENT DEVELOPMENTS 441 require a robust generator to supply a substantially constant load for a short period. This generator is usually coupled, together with a hydraulic pump, to a small high-speed turbine and it is convenient and economical to use the generator as a motor for checking and testing purposes on the ground. A polyphase induction machine is practically ideal for this dual role. (6.3) Cascade Induction Generator Fig. 3 shows a typical general arrangement of this type of machine. Essentially it comprises two ordinary induction machines arranged in a common housing with their rotor cores mounted on a common shaft. The rotor windings are connected together so that the m.m.f.s due to the rotor currents rotate, with respect to the shaft, either in the same direction or in opposite directions (see Fig. 4). A scheme has been described3 whereby both the stator windings of a machine of this kind, when connected to common busbars, would operate as a variable-speed constant-frequency generator without the inherent limitations of a single induction machine" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0003674_rpj-01-2019-0025-Figure6-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0003674_rpj-01-2019-0025-Figure6-1.png", "caption": "Figure 6 Analysis of internal structure following 3-pt bending and measurement of external surface roughness", "texts": [ " Albeit these figures are very encouraging, further improvements to the samples mechanical strength could be achieved by conducting secondary post-processing such as polishing and chamfering of the test specimens. Post-sintering, the density of the components was determined following the same protocol used in production by Morgan Advanced Materials. The density determination apparatus, in conjunction with a 0.1mg microbalance enabled the volume of the manufactured components to be calculated. The results showed that the components had <0.5 per cent porosity. Further analysis of the sintered components revealed densities of 99.7 per cent, which was validated by analysis under SEM. Figure 6 (a) shows half of a bend test sample that was used to obtain the surface roughness data interlayer region Ceramic components Jack Hinton et al. Rapid Prototyping Journal Volume 25 \u00b7 Number 6 \u00b7 2019 \u00b7 1061\u20131068 sinter fully to yield a monolithic component. Figure 6(b) shows the cross-section of the bend test sample along the break interface. The interface shows a number of 10-20 mm pores attributable to insufficient degassing during the packing of the material. Figure 6 (bottom row) shows the results of the surface roughness assessment on a bend test sample. Figure 6(c) shows the surface profile of the top surface that has been machined using an end milling operation. Figure 6(d) shows the surface profile for the edge of the component that has been machined using climb milling. Figure 6(e) shows the underside of the test samples that has not been machined. The markings are a result of dispensing the ceramic directly onto a PLA substrate. This paper has presented the development of a new digitally driven, hybrid manufacturing system, which has been validated by the production of high-density ceramic components. The synergistic application of additive and subtractive manufacturing processes facilitates personalised production of components with Ceramic components Jack Hinton et al" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0001259_sav-2010-0604-Figure11-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0001259_sav-2010-0604-Figure11-1.png", "caption": "Fig. 11. The 1st-4th order vibration shapes.", "texts": [ " After a set of signals are obtained, the coherent function values of them are firstly checked and only those bigger than 0.8 are kept for calculating transfer functions next. An obtained frequency response function is shown in Fig. 10 (a), and its coherent function is shown in Fig. 10 (b). As shown in Fig. 10 (b), because the coherent function is almost equal to 1, this means that the input and output signals have a good linear relation. For the BIW, the testing results of the natural frequencies and corresponding vibration shapes are shown in Table 7. Some vibration shapes are also shown in Fig. 11(a)\u2013(d). The signal coherences and the obtained vibration shapes show that the modal results are reasonable. Comparing the results obtained from FE calculated modes and the testing results of the BIW, it is shown that the natural frequencies and vibration shapes correspond to each other, just as shown in Table 8. But the calculated frequencies are a little bit greater than those of the tested. The reason for this difference concerns the structure damping assumption. For the natural frequency of a structure with damping, it is defined as follows" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0003660_iemdc.2019.8785146-Figure8-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0003660_iemdc.2019.8785146-Figure8-1.png", "caption": "Fig. 8 Analysis models of multi-three-phase SPMSMs", "texts": [ " SIMULATION OF STATIC ECCENTRICITY MODEL In this chapter, based on the theory described in Chapter 2, we verify the detectability of static eccentricity on surface permanent magnet motors (SPMSM) and interior permanent magnet motors (IPMSM) by extraction of no-load induced voltage. In order to inspect the possibility of detecting static eccentricity, we analyze and consider models in which static eccentricity is applied to each motor with a multi-three-phase and SPMSM structure. The models are shown in Fig. 8. For verification purposes, the models have ring-shaped magnets through which the gap magnetic flux density waveform becomes a square wave, making it easy to detect the harmonics. The analysis conditions are the same as in Table 1. Each model has a common stator and only the rotor is different. The number of groups in the model is 2, 3, and 6 in order and, in each model, the U phase, V phase, and W phase are arranged clockwise around each group. The no-load stator voltage waveforms after FEA are shown in Fig" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0001699_cbo9780511780509.003-Figure2.5-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0001699_cbo9780511780509.003-Figure2.5-1.png", "caption": "Figure 2.5. Single degree of freedom torsional example.", "texts": [ " Thus, using the notation of this application, fs = kx. The damping coefficient, c, for a viscous damper is defined as the force required to obtain a unit velocity across the damper. Hence, we have fd = cx\u0307. Substituting for fs and fd in Equation (2.10) and rearranging gives mx\u0308 + cx\u0307 + kx = f (t) (2.11) Dividing this ordinary differential equation (ODE) by m and letting c/m = 2\u03b6\u03c9n and k/m = \u03c92 n, we have x\u0308 + 2\u03b6\u03c9nx\u0307 + \u03c92 nx = f (t)/m (2.12) Before solving Equation (2.12), we consider the system in Figure 2.5, which consists of a disk connected to a shaft and a torsional damper. An external torque, \u03c4 (t), is applied to the disk. The motion \u03b8(t) is such that the shaft is twisted and behaves as a torsional spring. The shaft and torsional damper are assumed to have a negligible polar moment of inertia compared to the disk. The free-body diagram for the disk is shown in Figure 2.6. In the case of rotation about a fixed axis, Newton\u2019s second law may be stated as follows: The sum of the torques acting on a body rotating about a fixed axis is equal to the rate of change of angular momentum", "24) in terms of the initial displacement and velocity, x0 and x\u03070. If m = 1 kg, c = 0, k = 9 N/m, x0 = 1 mm, and x\u03070 = 0, calculate the resulting transient response. Calculate the response again if the damping is c = 1 Ns/m. Sketch the response in both cases. 2.2 A steel disk of 400 mm diameter and 10 mm thickness is mounted on one end of a steel shaft of 10 mm diameter and 150 mm length, whose other end is fixed. Assuming that the system may be modeled as a single degree of freedom torsional system, as shown in Figure 2.5, calculate the natural frequency in Hz. The torsional stiffness of the shaft is k = GJ/ , where G is the material shear modulus, J = \u03c0d4/32 is the second moment of area of the shaft, and and d are the shaft length and diameter. The polar moment of inertia of the disk is I = \u03c1\u03c0hD4/32, where D and h are the disk diameter and thickness and \u03c1 is the material density. For steel, assume G = 80 GPa and \u03c1 = 7, 800 kg/m3. 2.3 A car may be crudely modeled as a single degree of freedom system, in which the tires and suspension are assumed to act collectively as a single spring" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0001597_iros.2012.6386081-Figure8-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0001597_iros.2012.6386081-Figure8-1.png", "caption": "Fig. 8. Structure of FRT-type thumb exoskeleton", "texts": [ " The environment contact part and the finger contact part are connected each other with wires at both right and left sides. An assistive force from the main exoskeleton acts on an intermediate point of the wire. The distribution factor of the assistive force is determined by a horizontal position of the intermediate point. It becomes 1:1.2(=cos\u03c62 : cos\u03c61) when an angle inside the exoskeleton, \u03c61, is 28\u25e6 and an angle outside the exoskeleton, \u03c62, is 42\u25e6 in Fig.7. The FRT-type thumb exoskeleton is shown in Fig.8. It also has three components: a main exoskeleton, a thumb contact part and an environment contact part. In the same way as the index finger exoskeleton, the right and left environment contact parts are fixed each other with a solid arched components which is silver color in Fig.9. Both a thumb of a wearer and an environment contact part of the exoskeleton touch the environment. A nonslip sponge is installed on the thumb contact part so that it prevents a wearer\u2019s thumb from moving at the thumb contact part when grasping an object" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0002552_978-3-319-65298-6_21-Figure4-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0002552_978-3-319-65298-6_21-Figure4-1.png", "caption": "Fig. 4. Tennis-pick robot", "texts": [ " In 2011, the twelfth \u201cChallenge Cup\u201d national college students\u2019 extracurricular academic and technological works contest also develop a tennis-picking robot research. The works include embedded intelligent tennis-picking robot, tennis-picking intelligent robot prototype based on smart car and camera recognition technology, and smart tennispicking robot [7\u20139] (Fig. 3). In 2012, Southeast University, Beijing University of Posts and Telecommunications, and Alvin Tan Wei, University of Bradford, UK [10, 11], develop kinds of drum-type, roller-type and spring-type tennis-picking robot (Fig. 4) which mainly relies on machine vision, remote control or GPS planning path to complete the process of picking up. In 2013, Lanzhou University of Technology designs the intelligent cleaning-type tennis recycling robot based on the visual recognition and infrared sensor obstacle avoidance [12]. It is composed of image acquisition system, obstacle avoidance system, power system, pickup counting system and information processing system. The data display function is realized between the wireless communication network and the host computer" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0000756_msna.2012.6324516-Figure1-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0000756_msna.2012.6324516-Figure1-1.png", "caption": "Figure 1. Ball bearing geometry", "texts": [ " This impulse will motivate the nature frequency of inner, outer race and rolling elements [9]. For a particular bearing geometry, inner race, outer race and rolling element faults generate vibration spectra with unique frequency components. These frequencies, known as the defect frequencies, are functions of the running speed of the motor and the pitch diameter to ball diameter ratio of the bearing. Outer and inner race frequencies are also linear functions on the number of balls in the bearing. Given the geometry of the bearing in Fig. 1, for an angular contact ball bearing in which the inner race rotates and the outer race is stationary, the four characteristic frequencies is presented in Table 1. Where the outer race is fixed, if is the rotation frequency of shaft in hertz, D is the pitch diameter, d is the ball diameter, \u03b1 is the contact angle, and Z is the number of balls. Assume the contact between balls and inner race and outer race is pure rolling contact [2]. On line vibration measurement and analysis instrument is one important tool for rolling bearing faults identification" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0001985_978-1-62703-550-7_18-Figure3-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0001985_978-1-62703-550-7_18-Figure3-1.png", "caption": "Fig. 3 Transesterification of t-butyl alcohol with tributyrin under solvent free conditions. In the control freeze dried preparation of the enzyme was used. 1 mL reaction medium containing tertiary alcohol (1 M) in excess tributyrin [29] was incubated with 10 mg lipase formulations at 20 \u00b0C under constant shaking at 250 rpm for 72 h. Initial rates were based upon conversions obtained by gas chromatography analysis. The reactions were carried out with 0 % water added (w/w enzyme) (black bars) and 3 % water added (w/w enzyme) (grey bars) [18]", "texts": [ " Again, the comparison with lyophilized powders shows the advantage of using EPRPs (see Table 2, Scheme 1). The third protocol (see Subheading 3.3) describes a somewhat different application and one which needs wider exploration. Transformation of t-alcohols and their esters by most of the lipases is extremely slow or not possible [24, 25]. Only few lipases, with space available in their active sites to accommodate the t-alkyl groups, can carry it out. Protocol described in Subheading 3.3 shows that a simple precipitation with a t-alcohol Ipsita Roy et al. 277 helps (see Fig. 3). It also probably involves some \u201cimprinting\u201d effects [26]. The protocol involves employing one of the substrates, tributyrin (taken in excess), as the reaction medium. Such media are sometimes called solvent free media and certain advantages are associated with this approach [27]. High Activity Preparations of Lipases and Proteases for Catalysis in Low Water\u2026 278 A few papers (unfortunately literature comparing catalytic efficiency of various formulations is scarce) show that other formulations may be better for catalyzing some biotransformation [15, 28]" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0001609_0954406211427833-Figure1-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0001609_0954406211427833-Figure1-1.png", "caption": "Figure 1. Assembly model of the spherical gear pair mechanism mounted on a frame.6", "texts": [ " Based on the envelope theory of two-parameter family of rackcutter surfaces, mathematical and geometrical models of spherical gear with discrete ring-involute teeth were derived. When the rack cutter performs a reciprocating motion, the spherical gear performs two rotary motions. Later on, a two-dimensional rack-cutter profile was used instead of a rack-cutter surface to derive the spherical gear with discrete ring-involute teeth in a different approach. Using the envelope theory of oneparameter family, each tooth surface was generated individually.6 An assembly model with discrete ringinvolute teeth for convex and concave gear is shown in Figure 1. The spherical gear can rotate about either Y- or Z-axis. When the convex gear rotates about the Z1-axis, the mating concave gear rotates about Z2-axis, keeping the Y1-and Y2-axes fixed. To demonstrate the double degree of freedom motion of the spherical gear mechanism, actual model with discrete ring-involute teeth spherical gear pair has also been manufactured using rapid prototyping technique.6 Yang7 proposed a spherical gear with ring-involute teeth where gear teeth are continuously distributed on the segment of spherical surface", " Similarly, Sc:(Xc,Yc, Zc) is the frame for the instantaneous position of the cutter for which translation parameters are a2,b2, c2 and rotation parameters 2, \u20192, 2 from the reference frame coordinate. Sw:(Xw,Yw, Zw) is the work blank frame and a3, b3, c3the initial work blank position from the reference frame. the rotation parameters are 3, \u20193, 3. To start with, the cutter and the work blank solids are created at respective home positions, Oh and Ow. The specification of spherical gear pair considered in this study is presented in Table 1. The working of the spherical gears can be visualized with reference to Figure 1. During meshing of gear pair, the gears can rotate within 24 to 24 about the Y- and Z-axes independently, which are working axes in the assembled condition. The combination of rotation about these axes is also possible, as shown in the assembly. These working axes should not be confused with the notations for axes in Figure 3 used for CAD simulation. The simulation of the generation machining process is done according to the scheme presented in Figure 4. The different simulation parameters are presented in Table 2" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0000540_mspct.2011.6150487-Figure2-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0000540_mspct.2011.6150487-Figure2-1.png", "caption": "Fig. 2. RMF and synchro based IAS measurement setup", "texts": [ " The frequency of the induced emf in the rotor circuit is given by 120 120 120 2 120 3 Since the supply frequency (fs) and the number of stator poles (P) are constant, therefore, the frequency of the induced emf in the rotor circuit varies linearly with the variation of the rotor speed, nr. For this purpose, a three-phase balanced voltage supply is realized with the help of a sine wave oscillator and a single-phase to three-phase converter using three audio amplifiers as shown in Fig. 1. These voltage signals are applied to the stator winding of a synchro (Fig. 2). When the rotor of the synchro (rotating member) is standstill, the synchro acts as a transformer. Therefore the frequency of rotor emf (fr), is same as that of stator (fr = fs). When the rotor rotates, the frequency is proportional to the slip (ns \u2213 nr), where ns is constant. Figure 2 shows the set up for the realization of IAS measurement of low speed rotating machines. To obtain low instantaneous angular speed within a range of 0 to 2 rpm, a three-phase induction motor is used as a prime mover. The speed of this motor is first reduced to a maximum limit of about 10 rpm with the help of adjustable frequency ac drive system. To reduce the speed further, a pulley and wheel arrangement is used, which is mechanically coupled to the rotor of synchro. A sine wave signal of 40 Hz (18 V peak to peak) from a stable arbitrary function generator is supplied to a centretapped transformer" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0000458_amr.591-593.1879-Figure8-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0000458_amr.591-593.1879-Figure8-1.png", "caption": "Fig. 8 Total supporting stiffness of bearing 3Y. Fig. 9 Finite element model of the rotor-bearing system.", "texts": [ " Some peaks are found in the frequency range of interest, indicating that at these frequencies the casing mode is excited. The casing support dynamic stiffness at this bearing and direction is calculated by dividing the excitation force (1,000 N) by the displacements shown in Fig. 6. The dynamic stiffness of bearing 3Y is shown in Fig. 7. The total support stiffness then can be obtained according to Eq.1. The bearing stiffness bK is given as a constant value 1\u00d710 7 N/m. The total support stiffness of bearing 3Y is shown in Fig. 8. The total support stiffness of other bearings and directions are calculated in the same way. Rotordynamic Analysis with Foundation Effects. The total support stiffness data calculated in the previous section is retrieved and used to define parameter tables which describe the support stiffness varying against frequency. ANSYS bearing element COMBIN214 takes the parameter tables as its real constants and in this way the support dynamic stiffness characters are incorporated into the rotor-bearing model" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0000065_s1068371210080092-Figure1-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0000065_s1068371210080092-Figure1-1.png", "caption": "Fig. 1. Block diagram of two step vibroexciter.", "texts": [ " The processes that occur in electromechanical systems, which are a collection of electrical and mechanical devices, can be investigated using electromechanical analogy. The investigation of the processes that take place in the mechanical system can be replaced by investigations of the processes that occur in an electric circuit [3]. Let us investigate the electromechanical system of a two step vibroexciter for generating a mathematical model. A block diagram of the two step vibroexiter is presented in Fig. 1 and its electric circuit is given in Fig. 2. A two step vibroexiter consists of bed 1; cores 2, 3; winding systems 4, 5; springs 6, 7, 8; stiff element 10; tips 11, 12; ropes 13, 14, and vessels 15. If the core windings are connected to the ac source, diode D1 operates for one semi period and diode D2 operates for the other semi period. If the current passes through winding 4, the stiff element 9 is attracted to the core 2 and, after that, the tip 12 is lowed by ropes 13, 14 and the weight 11 is raised" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0003385_j.apm.2019.04.020-Figure7-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0003385_j.apm.2019.04.020-Figure7-1.png", "caption": "Fig. 7. Contact forces: (a) hard finger tapping system; (b) soft finger tapping system.", "texts": [ "3 , the movement posture, stick\u2013slip motion state, deflection of phalanx and structural stress of a soft finger tapping system driven by external moments are discussed. In the computation based on HCM, the event driven scheme given in Section 3.1 is adopted to determine the stick\u2013 slip and compression\u2013restitution transitions. The contact forces are calculated by LPM. The generalize- \u03b1 method is used to integrate the governing equation of system, and the computational parameters are chosen as \u03be = 0.8 and tol = 1 \u00d7 10 \u20135 . Meanwhile, the size of time step is set as t = 1ms for the non-impact phase as well as t = 0.1\u03bcs for the contact-impact Fig. 7 is the contact forces under different N . From Fig. 7 , it can be found that the contact forces under N = 15 have a good agreement with those under N = 60. It shows that the HCM has a good convergence to the density of mesh. Hence, N is chosen as 60 in the following calculation. To validate the accuracy and efficiency of the HCM, the HCM solution is compared with LS-DYNA solution. A threedimensional finite element model (see Fig. 8 ) is built by L S-DYNA software. The L S-DYNA model\u2019s physical parameters including Young\u2019s modulus, density, Poisson\u2019s ratio, geometric size, gravity acceleration and coefficient of friction are the same as those of HCM" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0003348_j.jmmm.2019.04.011-Figure2-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0003348_j.jmmm.2019.04.011-Figure2-1.png", "caption": "Fig. 2. Schematic illustration of three-dimensional magnetic field generated by the Helmholtz coils.", "texts": [], "surrounding_texts": [ "The flaky-shaped diatomite was employed as a template to fabricate magnetic microdisks by electroless nickel plating. The detailed process used for preparation can be found in our previous reports [13, 23-24]. The cleaned diatomite was first briefly activated using a colloidal palladium solution, followed by a peptization process for chemical activation. The activated wet diatomite (1g) were placed in a 50 ml coating solution for electroless nickel plating at 60 \u00b0C using a water bath. Vigorous mechanical stirring was adopted during this process to avoid agglomeration of the diatomite and to facilitate uniform nickel coating. The nickel coating solution was prepared in advance using 30 g/L NiSO4\u00b76H2O, 30 g/L NaH2PO2\u00b7H2O, 25 g/L C6H5O7Na3\u00b72H2O and 80 ml/L NH3\u00b7H2O. The nickel-coated diatomite was collected with a stainless cell scribble after reaction completion and was washed with deionized water. The products were further filtrated using a 350 mesh sieve to eliminate broken diatomite and residual nickel particles, followed by drying in an oven at 80 \u00b0C for 4 h." ] }, { "image_filename": "designv11_62_0003359_wje-05-2018-0167-Figure1-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0003359_wje-05-2018-0167-Figure1-1.png", "caption": "Figure 1 Rigid rotor model", "texts": [ " In Section 4, a discrete optimization algorithm based on Jaya is proposed, while a numerical example based on a balancing of the rigid rotor is demonstrated in Section 5. Finally, conclusions are outlined in Section 6. The balancing of the rotor using the numbers of masses at a corresponding angular position is investigated in this section. The rigid rotor requires two planes for its balancing, while flexible rotor requires multiple planes for its balancing. Further, reaction forces acting on the bearings are calculated using Newton\u2013Euler equation. A rigid rotor of mass (M) is mounted on bearings P and Q, as shown in Figure 1. A coordinate system (x y, z) is fixed coordinate system, while rotating coordinate system (xr, yr, zr) is attached to the shaft, rotating at a constant angular velocity v about the z-axis. The origin of the coordinate system is denoted by the point \u201cO.\u201d Two balance planes are considered, which are centered at points c1 and c2, respectively. The center of mass of the rotor \u201cG\u201d is eccentric at a distance of \u201ce\u201d from the axis of rotation owing to unbalance of the rotor. FP and FQ are the unbalance reaction forces acting at angles u P and uQ on the bearings P and Q, respectively. lP and lQ are lengths from \u201cO\u201d to the bearings P and Q, respectively. Numbers of discrete masses mij (j = 1 to Ni) placed at radius Ri and angular positions aij measured from x-axis, as shown in Figure 1 on balance plane i(i = 1,2) for rigid rotor, while (i = 1,\u00b7\u00b7\u00b7, p) for flexible rotor. Where p represents the number of planes. The position of the center of mass G is given as: OG ex; ey\u00f0 \u00dewhere ex \u00bc ecos u 1 v t\u00f0 \u00de and ey \u00bc esin u 1 v t\u00f0 \u00de Newton\u2013Euler equations of motion (Chaudhary and Saha, 2009) for the rigid body are written as: Mv\u0307 \u00bc fO (1) IGv\u0307 1 ~xIGx \u00bc MO (2) where fO is the resultant of external forces acting at supports, v\u0307 is linear acceleration of center of mass, x\u0307 is angular acceleration andMO is the resultant of all the external moments about point O" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0001672_978-1-4614-0222-0_34-Figure5-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0001672_978-1-4614-0222-0_34-Figure5-1.png", "caption": "Fig. 5 The first four out of plane mode shapes", "texts": [], "surrounding_texts": [ "The modal testing of the SRT model was recently studied [24]. Overall, a satisfactory agreement was found between the experimentally measured frequencies and those predicted in the finite element approach, Tables 9 and 10. The results indicate that in the out of plane direction, the FE models are slightly stiffer than the experimental model. While in the in-plane direction, except for the first mode, the experimental model is stiffer than the FE model with the lumped mass and is softer than the model with flap. A few sources of discrepancies could be indentified in this approach explaining the aforementioned differences between experimental and predicted frequencies. In the modal testing, since maintaining the internal pressure was done manually, it is less likely to achieve the same pressure at all times. Variable pressure contributes to variable stiffness which alters the modal parameter of the structure. Owing to presence of very small punctures in the SRT, the air leakage made the distribution of pressure non-uniform in the torus. In the finite element model, the pressure was assumed to be constant in the entire torus. Owing to the extremely small values of the film thickness, any increase in the thickness significantly affects the stiffness and mass of the structure and eventually the eigenvalues. For example, the average seams\u2019 thickness was above six times the shell film thickness. Furthermore, the method of fabrication, as well as the extensive repair that was done on the flaps, resulted in variations of the thickness of the joined regions. This contributed to non-uniform stiffness and mass distribution in both the torus and the flaps, affecting the modal parameters of the SRT. The flaps were not completely flat and had wrinkles at various locations. In the finite element study, the flaps were assumed to be flat with a uniform thickness equal to the average thickness of the inner and outer flaps." ] }, { "image_filename": "designv11_62_0001132_icnc.2013.6817972-Figure1-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0001132_icnc.2013.6817972-Figure1-1.png", "caption": "Figure 1. Notations and frame definitions for ship motion descriptions", "texts": [ " For marine vessels moving in the horizontal plain, the North-East-Down coordinate system e e ex y z (NED frame) and the body-fixed reference frame b b bx y z (B frame) is necessary to determine the position and the orientation. In the NED frame, the x -axis points towards the true North, and the y -axis points towards the East, and the z -axis points downwards normal to the Earth surface. The B frame is a moving coordinate frame which is fixed to the vessel. The position and orientation of the vessel are described relative to the NED frame while the linear and angular velocities of the vessel are expressed in the B frame, see Fig. 1. The horizontal motion of a ship is described by the motion components in surge, sway and yaw. Therefore, we choose [ ]u v r \u03a4=\u03bd and [ ]n e \u03c8 \u03a4=\u03b7 for the horizontal motion of surface ships. For the horizontal motion of a vessel, the kinematic equations takes the form as follows: cos sin 0 ( ) sin cos 0 0 0 1 R \u03c8 \u03c8 \u03c8 \u03c8 \u03c8 \u2212\u23a1 \u23a4 \u23a2 \u23a5= \u23a2 \u23a5 \u23a2 \u23a5\u23a3 \u23a6 (1) The vessel kinematics and the dynamics are conventionally described as in [2]: ( )R \u03c8=\u03b7 \u03bd (2) ( ) ( ) ( ) (t)NLM C D D R \u03c8\u03a4= \u2212 \u2212 \u2212 + + +\u03bd \u03bd \u03bd \u03bd \u03bd \u03bd b \u03c4 w (3) where [ ]n e\u03c8 \u03a4= \u03b7 denotes the position and orientation of the vessel in the NED frame e e ex y z (see Fig. 1), and [ ]u v r \u03a4=\u03bd denotes the translation velocities and angular rate in the B frame b b bx y z , and \u03c4 denotes the control forces and moments as system inputs. ( )R \u03c8 is a state-dependent transformation matrix, as in (1). 0M M \u03a4= > is the mass and inertial matrix. ( )C \u03bd is the Coriolis and centripetal matrix. D is the damping matrix. b is a bias term representing slowly varying environmental forces and moment. 0 0 0 0 u v g r g z r m X M m Y mx Y mx Y I N \u23a1 \u23a4\u2212 \u23a2 \u23a5= \u2212 \u2212\u23a2 \u23a5 \u23a2 \u23a5\u2212 \u2212\u23a3 \u23a6 (4) 13 23 13 23 0 0 ( ) 0 0 0 c C c c c \u03bd \u23a1 \u23a4 \u23a2 \u23a5= \u23a2 \u23a5 \u23a2 \u23a5\u2212 \u2212\u23a3 \u23a6 (5) 0 0 0 0 u v r v r X D Y Y N N \u2212\u23a1 \u23a4 \u23a2 \u23a5= \u2212 \u2212\u23a2 \u23a5 \u23a2 \u23a5\u2212 \u2212\u23a3 \u23a6 (6) 2 ( ) 0 0 0 0 NL uuuu u v v r v v r r r v v r v v r r r D X u X u Y v Y r Y v Y r N v N r N v N r = \u23a1 \u23a4\u2212 \u2212 \u23a2 \u23a5 \u23a2 \u23a5\u2212 \u2212 \u2212 \u2212 \u23a2 \u23a5 \u2212 \u2212 \u2212 \u2212\u23a2 \u23a5\u23a3 \u23a6 \u03bd (7) where 13 ( ) ( )v g rc m Y v mx Y r= \u2212 \u2212 \u2212 \u2212 and 23 ( )uc m X u= \u2212 " ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0002041_isr.2013.6695683-FigureI-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0002041_isr.2013.6695683-FigureI-1.png", "caption": "Fig. I. The exoskeleton and human subject to external forces.", "texts": [], "surrounding_texts": [ "foot of the exoskeleton is in contact with the ground for locomotion. The foot in contact with the ground exerts the force for supporting and moving its' body on the ground. The Ground Reaction Force (GRF) is reaction of that force. Otherwise, the GRF is a required force of the leg of the exoskeleton for locomotion. Based on this fact, we propose method that estimates desired GRF for locomotion of the exoskeleton and introduce control method of the exoskeleton using it. Then, we describe way and results that have a simulation in order to verify the effectiveness of the control method.\nIndex Terms-Exoskeleton, ground reaction force (GRF),\nzero moment point (ZMP), joint torque control.\nI. INTRODUCTION\nThere has been considerable interest in a exoskeleton for performance augmentation. The exoskeleton in the field of industry is expected to increase working efficiency of workers and prevent joint disease of them. Also, the exoskeleton in the medical field is anticipated for an opportunity to improve the mobility of the elderly and hemiplegic patients. But to commercialize the exoskeleton, the exoskeleton has to track the wearer's movement while carrying a payload. Therefore, the exoskeleton has been studied extensively to solve it.\nHybrid Assistive Leg (HAL) that has been developed at the University of Tsukuba is the exoskeleton for performance augmentation and rehabilitation [1, 2]. HAL's Cybernics Voluntary Control (CVC) is torque control method using the EMG sensors to measure muscle activity of the user. It is possible to predict the required joint torque for the wearer. But the control method using the EMG sensor is a serious problem. It is that data of the EMG sensor is extremely influenced by the given detection condition: it depends on subjects and even day to day measures of the same muscle site [3].\nHelper that has been developed at the Tohoku University is the exoskeleton for performance augmentation. Helper's control method is Model-based control algorithm for Anti-Gravity Muscles Assist with Ground Reaction Force feedback, reffered to as MAGMA-GRF [4, 5]. It is possible to calculate joint support moments by approximating human body to the kinematic model. But the GRF feedback system is prone to failure because of impact by frequent contacts with the ground.\nIn this paper, we propose new control method without using EMG sensor and GRF sensor. It is the way to estimate the desired GRF using accelerometer attached to human's body and transform for the desired GRF into joint torque. In the following part of this paper, we first explain the proposed control method using GRF estimation. Then, by applying the control method to the exoskeleton model, simulation results demonstrate the effectiveness of it.\n.. .. - -xexo = xh F:xo = Rexo - g exo\n(1)\n(2)", "In the single support phase, the only one leg is in contact with the ground so that we estimate the desired GRF acing on the leg supporting the exoskeleton. Figure 2 shows forces acting on the exoskeleton in the single support phase, and the sum of them is as follows:\nI F = mZi mg - R = mZi\n(3)\nIn the equation (3), m is mass, R is ground reaction force at the center of pressure (CoP), CoP is the contact\npoint on the ground, g is the acceleration of gravity and\na is the acceleration at the center of mass (CoM). Since\nwe can determine the acceleration of the exoskeleton the acceleration of the exoskeleton by measuring the acceleration of the wearer, the desired GRF can be estimated like equation (4).\nThen, we transform the desired GRF into the joint torque. The joint torque is obtained like equation (5) by using Jacobian.\n(5)\nIn the double support phase, two legs are in contact with the ground so that we estimate the desired GRF acting on the each leg. If we know CoP and the desired GRF acting on it, moment equation at the CoM of the exoskeleton is as follows:\nFig. 3.\nt F ,lit I R I R,\nThe exoskeleton subject to external forces in the double support phase.\n(6)\nIn the equation (6), x is a distance between the CoP\nand the CoM, Ry is the vertical component of the GRF,\nXl and x2 are distance between point of application of\nthe GRF and the CoM, and both FR and FJ. are the\nvertical components of the GRF at the contact point of feet. Like equation (7), we can estimate the desired GRF acting on each foot.\n(7)\nIn the equation (7), we can find the desired GRF by using the CoP. The CoP is the point on the supporting surface where resultant force of the GRFs is exerted [6]. Therefore, the CoP exists between the contact points of the right foot and left foot on the ground. It can be calculated as follows:\nx = Xl FH. + X2F'r\" FR+FL\nBut we can not measure FR and FJ.\n(8)\nin current\nsystem because the exoskeleton have not GRF sensing system. Therefore, we can not calculate the CoP. In order to solve this problem, we use zero moment point (ZMP). The ZMP is the point where the sum of all moments of active forces is equal to zero [7]. If the ZMP is within the support area, the ZMP coincides with the CoP [8]. ZMP is obtained as follows:" ] }, { "image_filename": "designv11_62_0001520_aim.2010.5695789-Figure3-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0001520_aim.2010.5695789-Figure3-1.png", "caption": "Fig. 3. The two methods to cut off the route for the fluid: a)bend a flat tube; b)pinch a tube with two rollers", "texts": [ " Moreover, drawing out two or more tubes arranged face to face as shown in Fig. 2 makes the restraint of the generation of the sliding friction possible. 978-1-4244-8030-2/10/$26.00 \u00a92010 IEEE 1368 One possible method to generate the force to draw out the folded tube is to pressurize it and to use the fluid energy inside it. If the route for the fluid is cut off, the fluid energy is transformed into the driving force at the cut off point. In order to cut off the route, two methods can be thought. One is to bend a tube flattened by heat-treatment as in Fig. 3a which has been developed by the Authors [4], and another is to pinch a tube with two rollers, named \u201cpinch roller\u201d as in Fig. 3b proposed by the Authors [5]. In this case, the robot consists of a head unit equipped with a small camera and a microphone, and of a tube to generate the driving force. The tube is bent inside the head unit, and it is fixed in such a way to not separate from the head unit. Thus, the driving force generated at the bending point of the tube is transmitted to the head unit, and the advancing motion of the robot is achieved. On the other hand, the retreating motion is achieved by supplying pressure into the opposite side of the tube" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0003803_s0025654419010023-Figure13-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0003803_s0025654419010023-Figure13-1.png", "caption": "Fig. 13.", "texts": [ " Let us consider the case of \u03b51 =0 (r =0) and construct the phase portrait of system (4.26) on the\u03c9z, vy plane. On the straight line (4.19), which meets the condition of non-slip of the front wheel of the bicycle model in the transverse direction A1y1, the tangents to the phase paths are horizontal. Neglecting O(\u03b531) terms, we obtain the equation of a curve on which the tangents to the phase trajectories are vertical vy = \u2212\u03c9za ( 1 + \u03b521c 2\u03b2\u03b3 a2 ) + vx\u03b4. (4.27) The phase portrait of the system (4.26) for the case \u03b4 < 0 (\u0394 < 0) is shown in Fig. 13. The phase portrait for \u03b4 0 (\u0394 > 0) is obtained from this phase portrait by parallel transfer of straight lines (4.19), (4.27) to 2vx|\u03b4| up along the vy axis. The analysis of phase portraits shows that, both in the framework of the multicomponent dry friction model and in the framework of the Coulomb friction model, the rotation of the front wheels affects the skidding of both axes of the device with locked or slipping rear wheels. For any initial conditions corresponding to (4.9) with respect to the variables vy, \u03c9z we have vy \u2192 vx\u03b4, \u03c9z \u2192 0, then from (4" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0002302_b978-0-12-803581-8.10351-0-Figure103-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0002302_b978-0-12-803581-8.10351-0-Figure103-1.png", "caption": "Fig. 103 Honeycomb cored mold stiffeners. Reproduced from Radford, D.W., Fuqua, P.C., Weidner, L.R., 2004. Tooling development for a multishell monocoque chassis design. In: 36th International SAMPE Technical Conference, San Diego, CA, November 15\u201318, 2004, pp.1063\u20131077.", "texts": [ " Based on the distortion after postcure of the flat test panels, described in a previous section, the decision was made to incorporate a grid stiffener system to the molds prior to postcure. Stiffeners were waterjet cut to CAD generated files, from wet layup panels produced from one layer of carbon cloth on each side of 12.7-mm-thick Nomex honeycomb core material. The same resin was used in the stiffeners as had been used in the molds. The stiffeners were designed, manufactured, assembled, and bonded to the completed molds prior to postcure so that the stiffened molds could go through freestanding postcure as a unit. Fig. 103 shows the CAD image of the interlocking grid stiffener system for the top shell mold on the left and the actual stiffener being bonded to the top shell mold. The stiffeners were designed to also function as mold stands once inverted, for preparation of the composite multi-shell components, as can be seen by the flat plane described by the top of the grid stiffener assembly. To prepare the stiffened unit, the molds were placed back on the appropriate tools prior to the addition of the stiffeners to help ensure the accurate shape retention" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0001059_1.4001772-Figure14-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0001059_1.4001772-Figure14-1.png", "caption": "Fig. 14 Special case for the second order, three-DOF planar motion of a rigid body when the angular velocity of the body is zero", "texts": [], "surrounding_texts": [ "p o i p\np s u s s a i t o v m d t t u p t a c o a f w W m T p w l s l\nA\nT o c f\ni\n0\nDownloaded Fr\ngeometry but depends on the kinematic properties translational and rotational velocities of the rigid body motion. 3. In general, the IC of any order follows two properties: directionality and proportionality. That is every motion property such as velocity, acceleration, jerk, etc. of every point E in the rigid body is at a constant angle k to the radius vector k that joins the point with the corresponding IC Ik. In case of first order velocity , the angle is 90 deg. The magnitude of the motion property is proportional to the radius vector, k. 4. The general equation for kth order motion property of any point in a rigid body under a planar motion can be expressed as follows: mk= Mk /cos k k ei k+ k , where Mk / cos k k is the magnitude while ei k+ k is the direction. Then the proportionality can be analytically described as mk k. Also the directionality comes from the fact that the angle between two vectors mk and k is k. 5. Study of special cases of motion provides valuable physical insight into the motion of the mobile platform.\nThe principal purpose of this paper was to create the most comlete analytical description of planar dynamic motion up to the kth rder as represented by K, mK, and K for each associated IK. It s understood that considerable development is still required to resent a uniform approach to mobile platform motion synthesis.\nThe dilemma with motion programming is to be able to obtain hysical meaning for the numerical values used in the motion pecification. Without this physical meaning, the programmer will nnecessarily have to guess at useful values. The next paper to be ubmitted will give an in-depth description of a full collection of pecial cases for the four orders of motion analysis derived here long with their physical meaning. For example, all points on the nflection circle travel along inflections in their motion. Given that 2, 3, and 4, are all 90 deg ensures that there are long inflecions in their paths. We will generalize the time motion synthesis f each wheel parameter in a future paper, which will lead to a ery direct and simplified physical understanding of the platform otion. This paper will use these physical meanings to provide esired motions for the platforms that meet the operator\u2019s needs in erms of those physical meanings. Furthermore, it will be shown hat the previous forward analysis approach in literature leads to nnecessary complexity, uncertainty of a solution, and virtually no hysical understanding of the specification meaning. Furthermore, he inverse approach suggested here will always have a known nd desired solution since the inverse is always parallel in its omputations, i.e., all wheel motions can be computed separately nce their attachment point motions are known, which this paper ccomplishes. This inverse formulation result is always the case or all parallel systems. Note that for serial manipulators, the forard analysis is always direct while the inverse is truly complex. hat has been achieved here is a synthesis of the mobile platform otion at all inputs by a simple and direct inverse calculation. his result should rapidly advance the motion planning of mobile latforms such that the really important results of dynamic forces, heel slippage, efficiency, turning effectiveness, mobile manipuation operations, etc., can all now be treated more directly, in the ame manner that was used to create similar results for manipuator operation 1,2 .\nppendix\n1 First Order Instant Center Formulation: Special Cases. he location of IC yields significant information about the motion f the platform. The following are some of the special cases that an be used to better understand the physical nature of the platorm motion.\n1.1 Case: X\u0307P= Y\u0307P=0 . When the velocity of point P is zero, \u02d9 \u02d9\n.e., when XP=YP=0 and 0, Eq. 5 becomes\n31015-10 / Vol. 2, AUGUST 2010\nom: http://mechanismsrobotics.asmedigitalcollection.asme.org/pdfaccess.\nXI1 = XP A1 YI1 = YP\nThus, the IC is located at the reference point P. This means that the body is under an instantaneous rotation around the axis passing through point P. This condition is shown in Fig. 10.\n1.2 Case: =0 . If the body is under pure translation motion, i.e., =0, Eq. 5 reduces to\nXI1 \u2192 A2 YI1 \u2192\nWith the IC at infinity, the velocity of a general point E cannot be defined using the Eq. 8 . However using Eq. 7 , we can see\nthat X\u0307E= X\u0307P and Y\u0307E= Y\u0307P. Also, we know that the velocity of E is orthogonal to the radius vector between E and IC. Thus, for every point in the body, the IC is located at infinity along the vector orthogonal to the velocity of the point.\nThus, when the angular velocity of the body is zero, there are an infinite number of ICs at infinity and every point in the body moves with the same linear velocity equal to the velocity of P, as shown in Fig. 11.\n1.3 Other Cases. Other platform configuration specific cases may constrain the location of the velocity IC and we can quickly evaluate if the motion requirement can be achieved or not. For example, in case of a fixed wheel, we know that the IC always lies on the wheel rolling axis. Thus, if the IC is away from the axis, we realize that the motion cannot be achieved without skidding. An example is a two wheeled platform such as shown in Fig. 12. The IC is constrained to lie along the wheel rolling axis.\nUsing the same platform configuration, we can consider another special case. When the velocity IC is coincident with one of the\nTransactions of the ASME\nashx?url=/data/journals/jmroa6/27999/ on 03/15/2017 Terms of Use: http://www.asme.org/about-asme/terms-of-use", "w v c p\nC .\nz\nv p\ns T a F\nf\nF f\nF f r\nJ\nDownloaded Fr\nheel attachment points, we can see that the whole platform reolves around that wheel, as shown in Fig. 13. This information an then be used to evaluate the feasibility of the motion by comuting wheel-ground interaction forces.\n2 Second Order Instant Center Formulation: Special ases\n2.1 Case: =0 . When the angular velocity of the platform is ero, the acceleration IC Eq. 16 is located at\nXI2 = XP \u2212 Y\u0308P\nA3\nY12 = YP + X\u0308P\nNotice that this expression is a similar to the expression for elocity IC given by Eq. 5 . The total acceleration for point E Eq. 20 also takes a form identical to the velocity of a general oint Eq. 8 as follows:\nX\u0308E = \u2212 Y 2 \u00b7\nA4 Y\u0308E = X 2 \u00b7\nThus, the location of acceleration IC and total acceleration asume a form identical to the first order formulation when =0. his means that when the angular velocity of the body is zero, the cceleration of a point is orthogonal to radius vector, as shown in ig. 14.\n2.2 Case: =0 . When the angular acceleration of the platorm is zero, the acceleration of point E Eq. 20 reduces to\nX\u0308E = \u2212 2Y 2 A5 Y\u0308E = \u2212 2X 2\nig. 13 The velocity IC coincident with one of the wheel center\nVP\nP\nI1\nVRVL\u03c9\nig. 12 The location of the velocity IC for a two wheeled diferentially driven platform is constrained to the axis of wheel otation\nor a two wheeled differentially driven platform\nournal of Mechanisms and Robotics\nom: http://mechanismsrobotics.asmedigitalcollection.asme.org/pdfaccess.\nHere, the only nonzero component of acceleration is the centripetal acceleration, which is always directed toward the center. Thus when angular acceleration of the body is zero, the total acceleration of every point on the body is directed toward the acceleration IC, as shown in Fig. 15. In this case, 2=180 and the first two ICs collide, i.e., I1= I2.\n2.3 Case: X\u0308P= Y\u0308P=0 . When the acceleration of point P is\nzero, i.e., when X\u0308P=Y\u0308P=0, Eq. 20 becomes\nXI2 = XP A6 YI2 = YP\nThus, the IC is located at the reference point P. This means that the body is under pure rotation up to the second order around an axis passing through the point P. This condition is shown in Fig. 16. The acceleration of a point on the body is dependent on its distance from the point P and is at an angle 2 with the radius vector.\n2.4 Case: = =0 . When both the angular velocity and angular acceleration of the mobile platform are zero, the IC for acceleration goes to infinity Eq. 16 . The acceleration of point E cannot be computed using the acceleration IC i.e., using Eq. 20 . However, using Eq. 17 , we can see that the acceleration of point E is instantaneously equal to the acceleration of the point of reference P. It means that all the points in the body move with same linear velocity up to the second order, as shown in Fig. 17.\nFig. 15 Special case for the second order, three-DOF planar\nAUGUST 2010, Vol. 2 / 031015-11\nashx?url=/data/journals/jmroa6/27999/ on 03/15/2017 Terms of Use: http://www.asme.org/about-asme/terms-of-use", "R F m\n0\nDownloaded Fr\neferences 1 Tisius, M., Pryor, M., Kapoor, C., and Tesar, D., 2009, \u201cAn Empirical Ap-\nproach to Performance Criteria for Manipulation,\u201d ASME J. Mech. Rob., 1 3 , p. 031002. 2 Kapoor, C., and Tesar, D., 2006, \u201cIntegrated Teleoperation and Automation for Nuclear Facility Cleanup,\u201d Ind. Robot, 33 6 , pp. 469\u2013484. 3 Uicker, J., Pennock, G., and Shigley, J., 2003, Theory of Machines and Mechanisms, Oxford University Press, New York, NY. 4 Cowie, A., 1961, Kinematics and Design of Mechanisms, International Textbook Company, Scrantom. 5 Hirschhorn, J., 1962, Kinematics and Design of Plane Mechanisms, McGrawHill, New York. 6 Myklebust, A., and Tesar, D., 1975, \u201cThe Analytical Synthesis of Complex Mechanisms for Combinations of Specified Geometric or Time Derivatives up to the Fourth Order,\u201d ASME J. Eng. Ind., 96, pp. 714\u2013722. 7 Oleska, S., and Tesar, D., 1971, \u201cMultiply Separated Position Design of the Geared Five-Bar Function Generator,\u201d ASME J. Eng. Ind., 92, pp. 74\u201384. 8 Lorenc, S., Stanisic, M., and Hall, A., 1995, \u201cApplication of Instantaneous Invariants to the Path Tracking Control Problem of Planar Two Degree-ofFreedom Systems: A Singularity-Free Mapping of Trajectory Geometry,\u201d Mech. Mach. Theory, 30 6 , pp. 883\u2013896. 9 Goehler, C., Stanisic, M., and Perez, V., 2004, \u201cA Generalized Parameterization of T1 Motion and Its Application to the Synthesis of Planar Mechanisms,\u201d Mech. Mach. Theory, 39, pp. 1223\u20131244. 10 Ambike, S., and Schmiedeler, J., 2008, \u201cA Methodology for Implementing the Curvature Theory Approach to Path Tracking With Planar Robots,\u201d Mech. Mach. Theory, 43, pp. 1225\u20131235. 11 Pennock, G., 2008, \u201cCurvature Theory for a Two-Degree-of-Freedom Planar Linkage,\u201d Mech. Mach. Theory, 43, pp. 525\u2013548. 12 Soh, G., and McCarthy, J., 2008, \u201cParametric Design of a Spherical Eight-Bar Linkage based on a Spherical Parallel Manipulator,\u201d ASME J. Mech. Rob., 1 1 , p. 011104. 13 Sreenivasan, S. V., and Nanua, P., 1999, \u201cKinematic Geometry of Wheeled Vehicle Systems,\u201d ASME J. Mech. Des., 121 11 , pp 50\u201356. 14 Bottema, O., 1961, \u201cSome Remarks on Theoretical Kinematics I. On Instantaneous Invariants,\u201d Proceedings of the International Conference on Teachers of Mechanisms, pp. 159\u2013164. 15 Bottema, O., and Roth, B., 1979, Theoretical Kinematics, North Holland, Amsterdam. 16 Veldkamp, G., 1963, Curvature Theory in Plane Kinematics, Walters, Groningen, Delft. 17 Veldkamp, G., 1967, \u201cCanonical Systems and Instantaneous Invariants in Spatial Kinematics,\u201d Mechanisms, 3, pp. 329\u2013388. 18 Veldkamp, G. R., 1969, \u201cAcceleration Axes and Acceleration Distribution in Spatial Motion,\u201d ASME J. Eng. Ind., 89, pp. 147\u2013151. 19 Alleivi, L., 1895, Cinematica Della Biella Piana, Tipografia Francesco Giannini and Figli, Naples. 20 Mueller R., 1960, \u201cPapers on Geometrical Theory of Motion,\u201d Special Report No. 21, Kansas Engineering Experiment Station.\nig. 16 Special case for the second order, three-DOF planar otion of a rigid body when the acceleration of point p is zero\n31015-12 / Vol. 2, AUGUST 2010\nom: http://mechanismsrobotics.asmedigitalcollection.asme.org/pdfaccess.\n21 Tesar, D., 1967, \u201cThe Generalized Concept of Three Multiply Separated Positions in Coplanar Motions,\u201d J. Mech., 2, pp. 461\u2013474. 22 Tesar, D., 1968, \u201cThe Generalized Concept of Four Multiply Separated Positions in Coplanar Motions,\u201d J. Mech., 3, pp. 11\u201323. 23 Tesar, D., and Sparks, J., 1968, \u201cThe Generalized Concept of Five Multiply Separated Positions in Coplanar Motions,\u201d J. Mech., 3, pp. 25\u201333. 24 Ridley, P., Bokelberg, E., and Hunt, K., 1992, \u201cSpatial Motion\u2014II. Acceleration and the Differential Geometry of Screws,\u201d Mech. Mach. Theory, 27, pp. 17\u201335. 25 Wang, L., Liu, J., and Xiao, Z., 1997, \u201cKinematic Differential Geometry of a Rigid Body in Spatial Motion\u2014III. Distribution of Characteristic Lines in the Moving Body in Spatial Motion,\u201d Mech. Mach. Theory, 32, pp. 445\u2013457. 26 Martinez, J., and Duffy, J., 1998, \u201cDetermination of the Acceleration Center of a Rigid Body in Spatial Motion,\u201d Eur. J. Mech. A/Solids, 17 6 , pp. 969\u2013977. 27 Denavit, J., and Hartenberg, R., 1955, \u201cA Kinematic Notation for Lower Pair Mechanisms Based on Matrices,\u201d ASME J. Appl. Mech., 22, pp. 139\u2013144. 28 Muir, P., and Newman, C., 1987, \u201cKinematic Modeling of Wheeled Mobile Robots,\u201d J. Rob. Syst., 4 2 , pp. 281\u2013340. 29 Campion, G., Bastin, G., and D\u2019Andrea-Novel, B., 1996, \u201cStructural Properties and Classification of Kinematic and Dynamic Models of Wheeled Mobile Robots,\u201d IEEE Trans. Rob. Autom., 12 1 , pp. 47\u201362. 30 Agullo, J., Cardona, S., and Vivancos, J., 1987, \u201cKinematics of Vehicles With Directional Sliding Wheels,\u201d Mech. Mach. Theory, 22 4 , pp. 295\u2013301. 31 Oetomo, D., and Ang, M., 2008, \u201cSingularity-Free Joint Actuation in Omnidirectional Mobile Platforms With Powered Offset Caster Wheels,\u201d ASME J. Mech. Des., 130 5 , p. 054501. 32 Low, K., Loh, W., Wang, H., and Angeles, J., 2005, \u201cMotion Study of an Omni-Directional Rover for Step Climbing,\u201d Proceedings of the IEEE International Conference on Robotics and Automation, pp. 1585\u20131590. 33 Alexander, J., and Maddocks, J., 1989, \u201cOn the Kinematics of Wheeled Mobile Robots,\u201d Int. J. Robot. Res., 8 5 , pp. 15\u201327. 34 Saha, K., Angeles, J., and Darcovich, J., 1995, \u201cThe Design of Kinematically Isotropic Rolling Robots With Omnidirectional Wheels,\u201d Mech. Mach. Theory, 30 8 , pp. 1127\u2013137. 35 Low, K., and Leow, Y., 2005, \u201cKinematic Modeling, Mobility Analysis and design of Wheeled Mobile Robots,\u201d Adv. Rob., 19, pp. 73\u201399. 36 Gracia, L., and Tornero, J., 2008, \u201cKinematic Models and Isotropy Analysis of Wheeled Mobile Robots,\u201d Robotica, 26, pp. 587\u2013599. 37 Yi, B. J., and Kim, W., 2002, \u201cThe Kinematics for Redundantly Actuated Omni-Directional Mobile Robots,\u201d J. Rob. Syst., 19 6 , pp. 255\u2013267. 38 Borenstein, J., 1995, \u201cControl and Kinematic Design of Multi-Degree-ofFreedom Mobile Robots With Compliant Linkage,\u201d IEEE Trans. Rob. Autom., 11 1 , pp. 21\u201335. 39 Reister, D., 1991, \u201cA New Wheel Control System for the Omnidirectional HERMIES-III Robot,\u201d Proceedings of the IEEE International Conference on Robotics and Automation, pp. 232\u2013237. 40 Beggs, J., 1983, Kinematics, Hemisphere, Newport, Australia. 41 Gallardo-Alvarado, J., and Rico-Martinez, J., 2001, \u201cJerk Influence Coeffi-\ncients, via Screw Theory, of Closed Chains,\u201d Meccanica, 36 2 , pp. 213\u2013228.\nTransactions of the ASME\nashx?url=/data/journals/jmroa6/27999/ on 03/15/2017 Terms of Use: http://www.asme.org/about-asme/terms-of-use" ] }, { "image_filename": "designv11_62_0001752_detc2011-48226-Figure5-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0001752_detc2011-48226-Figure5-1.png", "caption": "FIGURE 5. AN ACTUATION SINGULARITY: THE TWO ACTUATION FORCES F1 AND F2 ARE PARALLEL.", "texts": [ " However, a parallel singularity occurs iff the linear variety span(F1, F2, F3, F4, M1, M2, M3, M4) becomes of dimension lower than 6. Here, the degeneracy of the latter linear variety, for the singularity conditions revealed by GCA in Table 2, is examined. (a) (fi \u2016 f j). Without loss of generality, let us consider that f1 \u2016 f2, i.e., b\u2261 d, as depicted in Fig. 7(a). In that case, 1. From (2b1), L lies in the planar pencil at infinity Consequently, the linear variety spanned by M1, M2 M3 M4 F1 and F2\u2014which is generally of dimension 4\u2014 becomes of dimension 3. Such a configuration (f1 \u2016 f2), exemplified in Fig. 5, corresponds to condition (3b) of GG. (b) (ukl i j \u2016 z). In such a case, one has: Ti j \u2261Tkl and (Pi\u2229P j)\u2261 (Pk \u2229Pl) as illustrated in Fig. 7(b). As a result, the four planes Pi (i= 1, . . . ,4) of normal vectors (z\u00d7fi)=mi intersect at a common projective line Ti j \u2261 Tkl passing through point j and coplanar with each of the four actuation forces. Consequently, the latter line crosses the four constraint moments (the planar pencil at infinity through point j) as well as the four actuation forces. Thus, such a configuration (ukl i j \u2016 z) corresponds to a singular complex, i" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0003383_s11041-019-00349-7-Figure1-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0003383_s11041-019-00349-7-Figure1-1.png", "caption": "Fig. 1. Facility for laser cladding (a) and scheme of the process (b ).", "texts": [ " Thus, it becomes possible to change the structure and the properties of the alloy in individual regions of a part by varying the process parameters, i.e., to create a structurally graded material [13 \u2013 15]. The aim of the present work was to study the structural features of a gradient material obtained by laser cladding. We fabricated structurally graded specimens using a facility for laser cladding comprising the following components: a source of laser radiation with maximum powder 3 kW and characteristic wavelength 1070 nm, a cladding head, a powder feeder, and a manipulator (Fig. 1). The cladding head is the principal executive device of the laser cladding facility, which combines many systems, i.e., a system for laser focusing, an internal system for cooling the lenses and the nozzle, a system for feeding the material, a control Metal Science and Heat Treatment, Vol. 60, Nos. 11 \u2013 12, March, 2019 (Russian Original Nos. 11 \u2013 12, November \u2013 December, 2018) 739 0026-0673/19/1112-0739 \u00a9 2019 Springer Science+Business Media, LLC 1 National Technology Initiative \u201cNew Production Technologies\u201d Center at Peter the Great St" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0000515_s11661-013-2093-0-Figure1-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0000515_s11661-013-2093-0-Figure1-1.png", "caption": "Fig. 1\u2014Schematic of ECAP die setup showing processing parameters, (Die and Ram were made of Tool Steel H13).", "texts": [ " For SPD of the samples, an ECAP die with two equal cross section channels (30 mm in diameter) was used. The angle between two inner channels was 90 deg while the outer corner angle was 20 deg. Ram speed was 1 mm s 1 and the process was carried out at room temperature. In order to reduce friction, MoS2 was used as a lubricant. Samples were subjected to ECAP up to 8 passes in route A. The amount of strain at each pass is approximately about 1.[18,19] The process was carried out using a hydraulic press with the load increment increasing from 28 up to 45 t with each pass. Figure 1 shows the schematic of the ECAP die used in this research. Samples were machined from specimens treated by ECAP (10910910 mm). They were cut from the deformed regions located in the central part of the samples. All the microstructures were observed longitudinal to the ECAP direction defined as the y direction. To observe the microstructural evolution, a reheating procedure was performed on the samples with different levels of induced strain in a resistance furnace with Argon atmosphere and then the samples were quickly quenched in the water" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0000794_j.proeng.2012.04.015-Figure1-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0000794_j.proeng.2012.04.015-Figure1-1.png", "caption": "Fig. 1. The global coordinate system Fig. 2. The disc fixed coordinate system", "texts": [ " Crowther and Potts [1] developed a spinning disc wing and simulated straight flight. Based on previous work and in order to understand the phenomenon such as the steep change of the traveling direction of the disc, we validated the accuracy of the simulation based on the wind tunnel test data in this paper. Two coordinate systems were used in this research. The first was a global Cartesian coordinate system whose Y axis is along with the throwing direction and Z axis is in the upward direction (Fig. 1). The second was defined on the disc (Fig. 2). This disc coordinate system is also a Cartesian coordinate system. The x- axis and the y-axis were fixed on the surface of the Disc while the z-axis was perpendicular to them. The direction of each axis was fixed on the disc. The origin of this system was fixed on the center of the disc. The disc axes were written in small letters to distinguish them from those of the global coordinates. The transformation of the coordinate systems is performed using a quaternion" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0001461_scored.2010.5704057-Figure2-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0001461_scored.2010.5704057-Figure2-1.png", "caption": "Figure 2: (a) Current Sensor Board (b) Torque and Speed Measurement", "texts": [ "2 A, rated speed: 1000 rpm, related torque:7.8 3 Nm. The number of rotor bars is 28. Current, voltage, torque and speed were measured for a healthy motor and faulty motor at different levels of load. Experimental works were conducted at different conditions of no load and at 51 % and 80% of the motor related torque. The motor was coupled to a AC generator which acts as a load. The current signals were measured under a normal operation of motor. The sensor board is comprised of Hall Effect current sensors, LA25P by LEM (Figure.2.a). Torque and speed were used to ensure the motors are in the same situation in each test(Figure.2. b). The rotor bar breakage was forced in the laboratory by opening the motors and drilling the bar artificially (Figure.3) .. The data were obtained using a high speed data acquisition system, namely NI-PCI 6052E and SCXI 1125&1140 (National Instrument). In this research, the rate of frequency sampling is 20 kHz and the time of sampling is 5 second. Data analysis was performed using MA TLAB software. V. RESULT AND DISCUSSION This study is intended to investigate the accuracy of the frequency domain analysis of the stator current during its normal operation for fault diagnosis in 1M" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0000649_1.4001727-Figure3-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0000649_1.4001727-Figure3-1.png", "caption": "Fig. 3 Common perpendiculars between the screw axis and two homologous lines", "texts": [ " In the two-dimensional Reuleaux method, the angle of rotation of the rigid body around the pole is the angle between the line connecting any point of the body at the first position to the pole, and the line that connects the corresponding point to the pole after the rotation see Fig. 1 . Following the same line of thinking, we can see that in the three-dimensional case, the angle of rotation of the rigid body around the screw axis will be the angle between the common perpendicular of any line of the rigid body before displacement and the screw axis, and the common perpendicular of the corresponding line and the screw axis after the displacement see Fig. 3 . Let G be the common perpendicular between any line of the body at the first position and the screw axis, and G be the common perpendicular between the corresponding line and the screw axis after the helical motion. Both lines will intersect the screw axis at a perpendicular angle. The angle of rotation of the rigid body around the screw axis will be the angle between G and G , Transactions of the ASME 016 Terms of Use: http://www.asme.org/about-asme/terms-of-use L b T t t 4 d h o r t s t o p a p t t I m J Downloaded Fr For the special case that La is parallel to La, and Lb is parallel to b, i" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0003891_ilt-06-2019-0213-Figure2-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0003891_ilt-06-2019-0213-Figure2-1.png", "caption": "Figure 2 Static equilibrium position of rotor", "texts": [ " The following dimensionless parameters are introduced: h \u00bc h0H; P \u00bc PaP ; K \u00bc 6vhR2=Pah20; X \u00bc 12hR2=Pah20; The dimensionless fluid lubrication equation can be written as: @ @u rH3 @P @u 1 sinw @ @w r sinwH3 @P @w ! \u00bc Ksin2w @\u00f0rH\u00de @u 1Xsin2w @\u00f0rH\u00de @t (6) where H is dimensionless oil film thickness P is dimensionless hydrodynamic pressure. As the spherical bearing is subjected to its own gravity and external load, the spindle exists eccentric as a result of the uneven oil film thickness. It is assumed that rotor deviates from the static equilibrium position O1 in x, y and z directions, as shown in Figure 2. The ex, ey and ez represent the eccentricities of the rotor along the x, y and z directions: Analysis of spherical hybrid sliding bearings JianWang, Jing Feng Shen and YaWen Fan Industrial Lubrication and Tribology Volume 72 \u00b7 Number 1 \u00b7 2020 \u00b7 93\u2013100 where \u00ab x, \u00aby and \u00ab z represent the eccentricity ratios along the x, y and z directions, respectively. The expression of the oil film thickness is: h \u00bc h0\u00f011 \u00ab xcosu sinw 1 \u00ab ysinu sinw 1 \u00ab zcosw\u00de (8) The expression of dimensionless oil film thickness can be written as: H \u00bc 11 \u00ab xcosu sinw 1 \u00ab ysinu sinw 1 \u00ab zcosw (9) Considering that the liquid is an incompressible fluid, the density and viscosity of the liquid do not change with changing of pressure and temperature" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0000826_kem.516.203-Figure2-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0000826_kem.516.203-Figure2-1.png", "caption": "Fig. 2 Schematic illustration of sintered test model Fig. 3 SEM image of metallic powder 100\u00b5m", "texts": [ " Subsequently, the laser beam was irradiated on the layer of the deposited metallic powder according to the CAD data. After forming a layer of sintered material, these processes were repeated as the laser scanning direction was varied by 90\u00b0 until a complete model was created. The laser sintering process was performed in a nitrogen atmosphere at room temperature to prevent oxidization. Test Model. To determine how residual stress develops in sintered material, a beam shaped test model as shown in Fig. 2 was proposed. The specification of the test model is summarized in Table 1. The below part of the test model was the base plate and the upper part was the sintered material. The base plate was sandblasted with #35 of average grain size to improve the wetting property of melted powder [5]. The sintered material was made from a mixture of metallic powder as shown in Fig. 3. The mixture consisted of 70 % chromium molybdenum steel powder, 20 % copper alloy powder and 10 % nickel powder in weight with a mean diameter of 25 \u00b5m" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0000685_coase.2012.6386315-Figure2-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0000685_coase.2012.6386315-Figure2-1.png", "caption": "Fig. 2. Experimental setup for measuring hysteresis of a pneumatic artificial muscle.", "texts": [ " Nor is it necessary for each client to have a highperformance processor, because the server processes data in a database. This results in the reduction of costs when adding more clients. \u2022 A system in which some clients (and robots for rehabilitation or nursing) are remotely monitored by one therapist through the Internet, as described above, could contribute to the care and treatment of a future aging society. Prior to constructing an actual networked JIT control system, the JIT modeling performance for PAMs has to be tested. Fig. 2 illustrates the experimental setup used for determining the actual relationship between the contraction and the internal pressure of a PAM. The investigated PAM has a diameter of 20 mm and a nominal length of 400 mm. As shown in Fig. 2, this muscle is mounted vertically, and a weight is suspended on its bottom flange. The internal pressure, p, is detected by a pressure transducer, and its signal is sent to a PC through an analog-to-digital (A/D) interface. A proportional-integral (PI) control algorithm for regulating the internal pressure is implemented on the PC, which produces the operating voltage of a proportional control valve to adjust the air flow rate into the muscle. Applying the voltage through a D/A interface to the valve allows the internal pressure to be set to a reference pressure" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0001838_ma302559w-Figure4-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0001838_ma302559w-Figure4-1.png", "caption": "Figure 4. Definition of strain phase \u03d5 of oscillatory shear strain \u03b3 (solid line) and rate of shear strain \u03b3 \u0307 (dotted line).", "texts": [ " Figure 3c reflects a weak preferential orientation of the BCC lattice as evidenced by a weak \u03bc dependence of the first-order maximum at qm. The two peaks at \u03bc = 90\u00b0 and 270\u00b0 are the local maximum with a lower intensity, while the four peaks at \u03bc = 35\u00b0, 145\u00b0, 215\u00b0, and 325\u00b0 are the local maximum with a higher intensity, which may be a memory reflecting both a weak preferred orientation of hex\u2212cyl in the as-cast film30,31 and the OOT process from hex\u2212cyl to BCC-spheres. The details will be discussed later in section IV-2. III-2. Strain-Phase Resolved SR-DSAXS Studies during LAOS. Figure 4 shows the definition of strain phase \u03d5 of oscillatory strain \u03b3 (solid line). The numbers 1\u22124 in the figure designate the points corresponding to \u03d5 = 0, \u03c0/2, \u03c0, and 3\u03c0/2 rad, respectively, while \u03b3 \u0307 is the rate of shear strain (dotted line). Figure 5 shows the SR-DSAXS patterns in the parameter space of strain phase \u03d5 and strain cycle N. The patterns in rows (a), (b), and (c) are obtained at N = 1, 5, and 20, respectively, corresponding to the points 2, 3, and 4 in Figure 1b, respectively. The X-ray exposure time to record each pattern is 1 s, corresponding to the strain-phase interval of 0" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0003901_3351917.3351930-Figure2-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0003901_3351917.3351930-Figure2-1.png", "caption": "Figure 2: The relative motion relationship in horizontal plane.", "texts": [ " The orientation angle i is that between i v and the X-axis in the inertial system. The real-time position of each quadrotor can be uniquely determined by its velocity and orientation angle. To simplify the problem in actual analysis, we first assume that the orientation angle of each quadrotor UAV is 0 i , that is, without considering the coordinate system conversion problem, it can be directly analyzed in the inertial system. Taking the two-quadrotor UAV case as an example, the relative motion relationship in the horizontal plane in the inertial system is shown in Figure 2. Based on Active Disturbances Rejection Control CACRE, July 19-21, 2019, Shenzhen, China Based on the virtual leader-follower method, the desired position { , , } Fd Fd Fd x y z of each follower can be determined by the actual position { , , } L L L x y z of the virtual leader and their expected distance d l and angle d as: cos sin Fd L d d Fd L d d Fd L x x l y y l z z (2) Then, in subsection 2.3, the formation keeping controller can obtain the desired velocity commands for the followers by using such position information" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0001212_tia.2010.2070052-Figure6-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0001212_tia.2010.2070052-Figure6-1.png", "caption": "Fig. 6. Arrangement of thee-phase suspension windings.", "texts": [ " An inductor, capacitor, resistor (LCR) resonant circuit is constructed to enhance voltage variation. The capacitor voltages Vc(AB) and Vc(CD) are the inputs of a differential amplifier. Then, the output signal of the differential amplifier is multiplied by a phaseshifted carrier signal, which is the output of all-pass-filter. DC component of an output signal of the multiplier is proportional to the rotor displacement. The low-pass-filter eliminates highfrequency components. V. INFLUENCE OF SUSPENSION WINDINGS Fig. 6 shows a winding conductor arrangement of twopole three-phase suspension windings. Because search coils and suspension windings are set up to the same stator core, interference between suspension windings and search coils occurred. Flux of suspension windings links to the search coils and induced voltage occurs. The induced voltage provides bad influence to the sensor output. The interference between suspension windings and search coils is investigated, and search coil connections are compared. To evaluate the interference, a mutual inductance between search coil AB and U-phase suspension winding is calculated" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0001189_jtam-2013-0001-Figure1-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0001189_jtam-2013-0001-Figure1-1.png", "caption": "Fig. 1. Principle scheme of band saw", "texts": [ " The centre of mass of the disc is displaced from the axis of rotation of the distance e (eccentricity) and the axis of the disk makes an angle \u03b1 with the axis of rotation. In this paper, the dynamic reactions in the bearings of the basic shaft, which drives the band saw machines, are analyzed. These reactions are caused by the external loading and the kinematics and mass characteristics of the rotating disk. The expressions for the full dynamic reactions are obtained. These expressions allow the parameters of the machines to be chosen in such a way that the loading in the shaft and the bearings to be minimal. Figure 1 shows scheme of the band saws [1, 2, 3, 4, 5]. We define the following symbols: 1, 2, 5, 6 \u2013 belt pulleys, E \u2013 electric motor, 3 and 4 \u2013 feeding wheels, A \u2013 band-saw blade, 7 and 8 \u2013 chain-wheels. The dynamic model, shown in Fig. 2 is used for solving the problems. The feeding wheel 3 and the belt pulleys 2 and 5 perform rotation with a constant angular velocity \u03c9 about the axis of rotation AB. In this case, the mechanical system (the feeding wheel, the belt pulleys and the basic shaft) describes an angle \u03d5 = \u03c9t" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0000299_s1990478912010127-Figure3-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0000299_s1990478912010127-Figure3-1.png", "caption": "Fig. 3. An inverted pendulum on the cart", "texts": [ " Next, the zero solution of the subsystem x\u03073(t) = x4(t), x\u03074(t) = u(t) is stabilized by the control u(t) = \u2212x3(t \u2212 2r) \u2212 2x4(t \u2212 r); and, by Lemma 3, this control is globally stabilizing for rather small r \u2265 0, and the solutions of the correspondent closed system satisfy the condition 5 of Theorem 3. Note also that the function g satisfies the condition 4 of Theorem 3. Thus, for the considered system closed by the control u(t) = \u2212x3(t\u2212 2r)\u2212 2x4(t\u2212 r), all conditions of Theorem 3 are fulfilled; and, therefore, the specified control provides the global stability of the zero solution for rather small r > 0. Example 2. Consider a model of the inverted pendulum placed on a linearly moving cart (see Fig. 3). Let the pendulum and the cart move in one plane. Denote the mass and the length of the pendulum by m and l respectively where M is the mass of the cart, and F is the force applied to the cart. Some issues concerning the control of this system and corresponding references are discussed e.g. in [20\u201322]. Consider the issue of stabilizing the pendulum equilibrium. Let x be the relation of the cart displacement to the length of the pendulum, \u03b8 be the angular displacement of the pendulum, u = F/(mg), and \u03b4 = M/m" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0002200_978-3-319-51222-8-Figure2.17-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0002200_978-3-319-51222-8-Figure2.17-1.png", "caption": "Fig. 2.17 Rules for pole-plans", "texts": [ "4 Statically Determinate Structures 95 EI or EA changes in a beam, then the associated forces f + balance (see Chap. 5), and because equilibrium forces are orthogonal to all rigid body modes, that is, kinematic chains (the influence functions), N, M, and V do not change. To draw the displaced shape of a mechanism requires the knowledge of the instant centers of rotation or poles of the single segments. We call the plan of all these poles pole-plan. The following rules apply to the construction of pole-plans (see Fig. 2.17): 1. Each fixed hinge is the main pole of the attached segment. 2. Each hinge forms the secondary pole of all the segments connected with the segment. 3. The line orthogonal to a roller support forms the location of the main pole of the segment attached to the support. 96 2 Betti\u2019s Theorem 4. The secondary pole of two segments which are connected by a sliding hinge (normal force hinge or shear force hinge) lies on the line orthogonal to the sliding movement at infinity. 5. Themain poles of two segments and their common secondary pole lie on a straight line: (i) \u2212 (i, j) \u2212 (j), e" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0002551_iemdc.2017.8002063-Figure7-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0002551_iemdc.2017.8002063-Figure7-1.png", "caption": "Fig. 7. Fluxes of the machine in the steady state after short-circuit", "texts": [ " In the simulation presented in this article, saturation of the leakage fluxes can be observed. In order to see the saturation of leakage fluxes, we can look at two particular instants of short-circuit during which the common flux is weak compared to the leakage fluxes. First, we take the instant when the field current reaches its peak value, Fig. 6, most of the stator flux linkage is leakage flux. Pole tips and stator teeth are strongly saturated. Main flux in the stator is very weak. Second during the short-circuit steady state, Fig. 7, most of the flux linkages constitute also leakage fluxes. This time, the magnetic circuit is not saturated. This type of saturation would not be taken into consideration by most models, which assume the linearity of the leakage inductances. If the fluxes in the saturated state are inaccurate, it has a direct impact on the machine torque, power and power factor. The most realistic model of flux linkage in the machine depends on all currents in its windings. The model presented here considers that every flux linkages d, q and f depends on stator currents id, iq and rotor current if" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0001123_gt2012-68510-Figure4-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0001123_gt2012-68510-Figure4-1.png", "caption": "Figure 4: Recuperator layers with and without sacrificial support material.", "texts": [ " As seen in Figure 3, each segment was equivalent to a 10 degree section of the recuperator. Therefore, the entire recuperator consists of 36 segments with each segment measuring 257 layers tall. The 257 layer tall segments are comprised of 64 four-layer wafers plus a single \u201cend cap\u201d layer. In the green state (meaning before sintering), the material was very flexible and was prone to break during handling. Adding sacrificial material around the exterior and interior supports made the layers more durable and helped maintain straight channels throughout fabrication. Figure 4 shows the actual geometry of the heat exchanger layers before and after the supports and sacrificial material were added. Internal supports located on alternating layers were removed by simply applying sufficient pressure to break them. A good understanding of each component\u2019s physical properties was essential to determine a proper fabrication process. Table 1 lists all the components of the liquid mixture, known as the slip, used to make the tapecast sheets. The table includes mass percentages and a brief description of the purpose of each constituent" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0003742_ecc.2019.8795690-Figure3-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0003742_ecc.2019.8795690-Figure3-1.png", "caption": "Fig. 3: The image capturing process for a maze using a quadrotor with a camera at bottom. The camera can distinguish different colors in its camera plane, for example the gray represents the wall.", "texts": [ " Image Processing forMaze Reconstruction The maze environment is generated in a simulator based on Unreal Engine 4 and Aisrim [13], [14]. The simulator is an open-source 3D robotics simulator, which integrates a variety of development libraries such as Open Physics Engine library. as depicted in Fig. 2. To reconstruct the maze, several steps are implemented, as depicted in Fig. 2. The quadrotor with a camera at its bottom is manipulated to fly to the right above of the maze center C0(x0, y0), as shown in Fig. 3. When it arrives above the maze, it will rise vertically to a certain height in order to scan the whole maze. The camera will return a 2D pixel matrix M (M \u2208 Rnr\u00d7nc , where nr, nc denotes the numbers of pixels in row and column respectively the camera captured). The pixel matrix M with color value includes the information for the maze. To extract the wall information, we can use the following preprocessing approach: M\u0302(i, j) = { 1, i f |M(i, j) \u2212 g| 6 \u03b8, 0, else, (1) where i = 0, 1, ..., nr \u2212 1; j = 0, 1, " ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0000678_j.mechmachtheory.2012.09.003-Figure3-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0000678_j.mechmachtheory.2012.09.003-Figure3-1.png", "caption": "Fig. 3. The non-hybrid six-bar NH1.", "texts": [ " [4], corresponding link-lengths a1, a2, b1, b2, g1, g2, d1, d2 are defined in standard fashion, namely, d2s\u03b11 \u00bc a1s\u03b42 b1s\u03b12 \u00bc a2s\u03b21 b1 \u00fe b2\u00f0 \u00des \u03b11 \u00fe \u03b12\u00f0 \u00de \u00bc a1 \u00fe a2\u00f0 \u00des \u03b21 \u00fe \u03b22\u00f0 \u00de g1s\u03b22 \u00bc b2s\u03b31 g2s\u03b41 \u00bc d1s\u03b32 b1\u2212d2\u00f0 \u00des \u03b32\u2212\u03b11\u00f0 \u00de \u00bc g2\u2212a1\u00f0 \u00des \u03b21\u2212\u03b42\u00f0 \u00de ) ; \u00f03\u00de in which the letter s abbreviates sine. It is immediate that only two of the link-lengths can be freely selected. As a consequence, 12 hybrid six-bars of Waldron's [7] type may be assembled, one at each revolute; corresponding chains in the plane or on the sphere would have mobility 3. Alternatively, the network might be perceived as an assemblage of four non-hybrid six-bars, such as that illustrated in Fig. 3, called NH1 in Ref. [5]. Now, whether the various loops are displayed as open or crossed, the relevant closure equations remain valid; possible reversal of a joint axis direction is easily accommodated. Taking advantage of this fact, the skew network encapsulated in Fig. 2 is now portrayed schematically by Fig. 4. Although interchangeability between any of the non-hybrid six-bars and each of the apposite \u201csyncopated contiguous hybrids\u201d (SCH), \u201csyncopated opposed hybrids\u201d (SOH) and \u201caugmented internal hybrids\u201d (AIH)was demonstrated in Refs", " In the same way, for isograms 3\u20134\u20139\u20138\u2013 and 6\u20131\u201312\u201310\u2013 we have t \u03b832 t \u03b812 \u00bc t \u03b412 \u00fe t \u03b32 2 t \u03b412 \u2212t \u03b32 2 \u00bc t \u03b12 2 \u00fe t \u03b21 2 t \u03b12 2 \u2212t \u03b21 2 ; \u00f06\u00de in accordance with the first of Eqs. (2), and showing the loops to be similar. On account of symmetry, the remaining two loops would also be kinematically similar, but Eqs. (5) and (6) may be combined to yield t \u03b82 2 t \u03b83 2 \u00bc t \u03b422 \u00fe t \u03b11 2 t \u03b422 \u2212t \u03b11 2 ! t \u03b412 \u00fe t \u03b32 2 t \u03b412 \u2212t \u03b32 2 ! ; while, directly from chains 2\u20133\u201311\u201312\u2013 and 5\u20136\u20137\u20139\u2013, we see that t \u03b82 2 t \u03b83 2 \u00bc t \u03b31\u00fe\u03b32 2 \u00fe t \u03b41\u00fe\u03b42 2 t \u03b31\u00fe\u03b32 2 \u2212t \u03b41\u00fe\u03b42 2 \u00bc t \u03b21\u2212\u03b42 2 \u00fe t \u03b12\u2212\u03b31 2 t \u03b21\u2212\u03b42 2 \u2212t \u03b12\u2212\u03b31 2 : With regard to Fig. 3 and Eq. (4), it is revealed that \u03b87 \u00bc \u03b82\u2212\u03c0; \u03b88 \u00bc \u03c0 \u00fe \u03b81: Thus, the skew network of eight links comprises a nested colligation of three pairs of similar isograms. As a spatial assembly without special geometrical properties, its mobility would be reckoned as\u221218. Even the planar and spherical counterparts have a nominal mobility of \u22123. It is collapsible by reason of the contributing four-bars simultaneously adopting in-line configurations during their cycles of motion. In principle, too, the assembly is capable of infinite extension, whereby it would function as one cell among others" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0002177_3029798.3038422-Figure3-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0002177_3029798.3038422-Figure3-1.png", "caption": "Fig. 3. Leap motion. Fig. 4. Coordinate systems.", "texts": [ " In the authors\u2019 previous research, a touch panel display was used as a brush position measuring device, however, it caused an extension and a complexity of the system. Therefore, a new selftraining system for calligraphy is developed in this paper; Leapmotion sensor is newly introduced as a brush position measurement device. The possibility of measurements of the brush is confirmed, and pilot experiments are carried out for the evaluation of the developed system. The developed calligraphy training system consists of a pressure presentation device, Leapmotion sensor and a computer (see Figure 2). Leapmotion (see Figure 3) is an input device for a computer that is dealt by Leapmotion Company. Leapmotion has two infrared cameras, and it can measure the position and the orientation of a stick object like a finger or a pen in space of 50cm radius with 0.01mm precision up to 295Hz. For example, a user can operate a computer by gestures of the hand with Leapmotion. In this research, the tip position and the orientation of the brush are measured by Leapmotion at 20Hz sampling, and the brush point pXYZ is calculated by (1) Permission to make digital or hard copies of part or all of this work for personal or classroom use is granted without fee provided that copies are not made or distributed for profit or commercial advantage and that copies bear this notice and the full citation on the first page" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0001235_speedam.2010.5542396-Figure3-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0001235_speedam.2010.5542396-Figure3-1.png", "caption": "Figure 3. Cross-sectional geometry of the machine studied.", "texts": [ "00 \u00a92010 IEEE SPEEDAM 2010 International Symposium on Power Electronics, Electrical Drives, Automation and Motion ( ) ( ) ( ) ( ) 1 s w 1 0 1 s w 1 0 \u02c6 cos 1 2 \u02c6 cos 1 2 p p p p p p B B p t B B p t \u03b5 \u03c6 \u03c9 \u03c9 \u03b1 \u03b4 \u03b5 \u03c6 \u03c9 \u03c9 \u03b1 \u03b4 + + \u2212 \u2212 = + \u2212 + + = \u2212 \u2212 \u2212 + (1) where \u03b5 is the relative eccentricity, \u03b40 the average radial air gap, Bp the peak value of fundamental flux density, p the number of pole pairs and \u03c9w the whirling frequency. The method of analysis has been validated earlier by modeling and measuring a 15 kW cage induction motor [3]. Figure 2 shows a comparison between the measured and computed forces. The whirling motion was obtained by equipping the machine with active magnetic bearings and controlling them properly. A 40 kW axially laminated four-pole synchronous reluctance motor is used as an example machine. The parameters of the machine are given in Table I and crosssectional geometry in Figure 3. The machine is supplied from a three-phase sinusoidal voltage source at 50 Hz. There can be one, two or four parallel branches in a double-layer stator winding. The two parallel branches can be made in different ways. Of these, we study the ones that are first wound over separate pole pairs and then connected in parallel. From the electromechanical interaction point of view, it is important to know how the eccentricity harmonics (1) are coupled with the stator winding. This is described by the winding factors of the branches for the eccentricity harmonics" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0002587_icca.2017.8003129-Figure1-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0002587_icca.2017.8003129-Figure1-1.png", "caption": "Fig. 1. The relationship between xi, \u03bei, \u03b8i, \u03d5i, \u03d5\u0304i.", "texts": [ " , 1]T (hereafter this denotation will be simplified just as 1 whose dimension determined from the context) Suppose there are n targets with unknown positions \u03bei(t) \u2208 R2 at time t. There are also n agents with known trajectories xi(s) \u2208 R2 for s \u2264 t. We assume that each agent i\u2019s motion obeys a single integrator model x\u0307i = ui, i = 1, 2, . . . , n, (1) where xi(t) represents the agent i\u2019s position, and ui(t) is agent i\u2019s velocity control input. Assume that each agent i can only obtain the knowledge of the bearing angle with respect to target i, denoted by \u03b8i(s) for s \u2264 t. The case s = t is depicted in Fig. 1. The unit directed vector from xi(t) to \u03bei(t) is denoted by \u03d5i(t) = [ cos\u03b8i(t) sin\u03b8i(t) ] , and can be defined as \u03d5i(t) = \u03bei(t)\u2212 xi(t) \u2016\u03bei(t)\u2212 xi(t)\u2016 := \u03bei(t)\u2212 xi(t) \u03c1i(t) . (2) We denote by \u03d5\u0304i(t) \u2208 R2 the unit vector perpendicular to \u03d5i(t) obtained by \u03c0/2 clockwise rotation of \u03d5i(t). The center of the targets is defined by \u03be\u2217(t) = 1 n \u2211 j \u03bej(t). (3) Our aim is to design a center estimator that cooperatively estimate the unknown target center \u03be\u2217(t) using bearing measurements up to time t and a controller that makes each agent i move toward and then on a circle with desired radius \u03c1di centered at the target center \u03be\u2217(t) such that all targets in the circle" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0000900_ever.2013.6521527-Figure3-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0000900_ever.2013.6521527-Figure3-1.png", "caption": "Fig. 3. Flux lines in TM motor", "texts": [ ", LrLs ':':' LXi' The actual field solution is achieved iteratively. At each step, the rotor flux Iinkage is computed from the field solution. If the q-axis component Arq is different from zero, the q-axis rotor current irq is modified in such a way to satisfy the relationship Arq = 0 (i.e., the FO condition). The convergence process is quite rapid, requiring only a few iterations (typically three) , since the magnetic circuit is linear along the q-axis (the flux approaches zero). The flux Iines under load are shown in Fig. 3. Let us remark that, since both stator and rotor currents are imposed as the field sources, only a magnetostatic FE analysis is necessary also for the simulation under load, so as the saturation effects are careful taken into account. Once the field problem is solved, the motor performance can be computed. The stator flux Iinkage components, Asd and Asq, are computed by means of the magnetic vector potential. The inductances are computed from the rotor flux equation, and the rotor resistance is computed from the rotor Joule losses" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0000786_aero.2013.6496960-Figure10-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0000786_aero.2013.6496960-Figure10-1.png", "caption": "Figure 10 \u2013 Mechanical model of a subunit", "texts": [ " The sensor (FlexiForce, Nitta) is attached between the link and the expansion plate (Figure 9). Each link has a sensor, and an expansion plate has two sensors. Each subunit has a total of eight sensors, each of which measure the vertical push force against the wall. The output voltage from a sensor is amplified with an amplified circuit, and it is measured by a computer through an A/D converter. Pushing Force Model of Expansion Plate inside Cylindrical Shape The pushing force of the expansion plates is modeled here. Figure 10 shows the inner structure of the dual pantograph. A propulsion subunit expands and contracts in a radial direction along with the contraction and extension of plates in the axial direction. The contraction force of the plates is converted to an expansion force. Two links connect an expansion plate. The pushing force of the two links is given using contraction force W from experiments (see the contraction force in Table 1) and arm angle in the following equation. tan21 W FF (1) Figure 11 shows the top view of one of the expansion plates" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0003413_3312714.3312724-Figure2-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0003413_3312714.3312724-Figure2-1.png", "caption": "Figure 2. Surfaces of the Lyapunov function levels", "texts": [ " Such a change of variables transforms the original coordinate system into a new system and corresponds to an inhomogeneous linear transformation, which means the displacement of the origin and the parallel transfer of the coordinate axes. Investigation of the stability of any steady state refers to the study of perturbations stability with respect to the origin. ...);,...,1(),( 2121 CCniCXXV i trajectory, velocity vector and gradient vector gradV in the new mobile coordinate system It is obvious that if the derivative along the phase trajectory is everywhere negative, the trajectory of motion tends to the origin, i.e. the system is stable with respect to the origin ( ) in the original coordinate system (fig. 2). Otherwise, when the derivative is positive, the trajectory goes beyond the level lines ( ) ( ), i.e. the system is unstable. In this case, in the mobile coordinate system, the total derivative of the Lyapunov function ( ) constructed for perturbations ( ( ) ( )) in time will be equal to ( ) . Here, the first vector represents the gradient of the Lyapunov function ( ) and it is always directed to the side of the greatest growth of the function ( ). If the state equations of the system are defined in deviations ( ) and the origin corresponds to a stationary state , then the function ( ) increases as it moves away from the origin; i" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0003716_6.2019-4392-Figure13-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0003716_6.2019-4392-Figure13-1.png", "caption": "Fig. 13 Printed CAD Model of Nozzle", "texts": [ " D ow nl oa de d by U N IV E R SI T Y O F G L A SG O W o n Se pt em be r 2, 2 01 9 | h ttp :// ar c. ai aa .o rg | D O I: 1 0. 25 14 /6 .2 01 9- 43 92 Fig. 12 Section cut of printed CAD model of combustion chamber, showcasing the unique shape of the film cooling orifices designed for more reliable printing in the set direction. The nozzle in terms of print considerations was a lot less concerning since it was open on both ends. Primarily, this meant that the flange could be printed downward as demonstrated in Fig. 13. This drastically reduced the need for the heavy support material that was placed onto the chamber, and thus reduced the overall cost to print the part since less material was needed. However, now with the set print direction there was a concern within the regenerative channels that a large area would not be cooled due to no fluid having access to the area. The solution that was implemented involved creating a cavity zone that would allow the fluid to circulate while transferring between the nozzle and chamber" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0002455_icosc.2017.7958660-Figure1-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0002455_icosc.2017.7958660-Figure1-1.png", "caption": "Fig. 1. Diagram of the 6-degree-of-freedom gear dynamic model.", "texts": [ " The signal has the property that all the negative frequencies of x(t)have been filtered. The amplitude modulation (AM), the 978-1-5090-3960-9/17/$31.00 \u00a92017 IEEE 422 phase modulation(PM)and the frequency modulation (FM) are obtained as follows: A(t) = |xas(t)|= \u221a x2(t)+ j y2(t) (5) \u03b8(t) = tan\u22121 ( y(t) x(t) ) (6) \u03c9(t) = \u2202 (\u03b8(t)) \u2202 t (7) Where tan\u22121 ( y(t) x(t) ) gives the phase values in the range of [\u2212\u03c0 : +\u03c0]. The developed method of this research work uses Eq5to estimate the envelope signal analysis. The dynamic model of the electromechanical system is given in Fig. 1. We assume that the wheels are mounted on flexible bearings and are connected to each other by means of flexible teeth. The dynamic model is developed to study the bearing defect which influences on the dynamic behavior of the system. In this purpose, we propose a novel approach to model the dynamical behavior of bearing defect. The system consists of two parameters model, involving stiffness and damping, with torsional and lateral vibration and it has 6 degrees of freedom.Where \u03b8m the motor is rotational angle,\u03b81and \u03b82 are the pinion and the wheel rotational angles respectively", " The function of the observed stiffness ks(t) is given in Fig.3, during rotation. Since the rotation is constant, the apparent rotating stiffness of the system varies periodically over time. The crack closes for t \u2208 [0 : 0.015] the stiffness is that of the healthy system ks(t) = K0 and t \u2208 [0.015 : 0.03] The crack opens indicate a strong loss of rigidity so ks(t) \u227a ko The crack is said to be breathable. 978-1-5090-3960-9/17/$31.00 \u00a92017 IEEE 423 After some mathematical manipulations using equations of the electromechanical system given in Fig.1, to solve the matrix equation with periodic coefficients, we used the Newmark integration scheme. This method is useful to provide the dynamic response in each of the freedom degrees of the transmission in time domain. This approach was first developed by [14], . The frequency responses conversion information can be interpreted much easier. Our aim is to provide better understanding of defects inherent to gear transmissions of the electromechanical system used in industry and their influence on their dynamical behavior" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0001324_fuzzy.2011.6007664-Figure2-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0001324_fuzzy.2011.6007664-Figure2-1.png", "caption": "Fig. 2. Arm robot actuated by a DC motor", "texts": [ "2 \u2212 \ud835\udefe2\ud835\udf0e\u2223\ud835\udf0e\u2223( 1 2 )\ud835\udf09(\ud835\udf0e) ) +\ud835\udc37\ud835\udf0e (44) If we choose the following adaptation laws: \ud835\udf031 = \ud835\udefe1\ud835\udf0e\ud835\udf09(\ud835\udf0e)\ud835\udc61\ud835\udc50 (45) \ud835\udf032 = \ud835\udefe2\ud835\udf0e\u2223\ud835\udf0e\u2223( 1 2 )\ud835\udf09(\ud835\udf0e) (46) we obtain ?\u0307? = \u2223\ud835\udf0e\u2223\ud835\udc37 \u2212 ( \ud835\udf06\u22171\ud835\udc61\ud835\udc50 + \ud835\udf06 \u2217 2\u2223\ud835\udf0e\u2223( 1 2 ) ) \u2223\ud835\udf0e\u2223 \u2264 0 (47) Thus, the proposed control law guaranties both robustness and stability of the closed loop system. Moreover its simplicity allows an implementation for a real-time control. VI. ILLUSTRATIVE EXAMPLE In this section, the proposed control is used to control an arm robot with single joint, of mass \ud835\udc5a and languor \ud835\udc59. This arm is actuated by a DC motor as shown in figure 2. \ud835\udefc means the angle rotation of the arm relative to the vertical. The dynamic of the robot is given by the 3\ud835\udc5f\ud835\udc51 order equation: \ud835\udefc(3) = \ud835\udc530(\ud835\udefc, ?\u0307?) + \ud835\udc540(\ud835\udefc, ?\u0307?)\ud835\udc62+\ud835\udc37 (48)\u23a7\u23a8 \u23a9 \ud835\udc530(\ud835\udefc, ?\u0307?) = \u2212\ud835\udc45 \ud835\udc3f \ud835\udf03 \u2212 ( \ud835\udc3e\ud835\udc4f\ud835\udc41 2\ud835\udc3e\ud835\udc61 \ud835\udc5a\ud835\udc592\ud835\udc3f + \ud835\udc54 \ud835\udc59 \ud835\udc50\ud835\udc5c\ud835\udc60\ud835\udefc ) ?\u0307?\u2212 \ud835\udc45\ud835\udc54 \ud835\udc59\ud835\udc3f \ud835\udc60\ud835\udc56\ud835\udc5b\ud835\udefc \ud835\udc540(\ud835\udefc, ?\u0307?) = \ud835\udc3e\ud835\udc61\ud835\udc41 \ud835\udc5a\ud835\udc592\ud835\udc3f where \ud835\udefc(3), ?\u0308? and ?\u0307? are the time derivatives of the angle \ud835\udefc. \ud835\udc54, \ud835\udc3f, \ud835\udc45, \ud835\udc41 , \ud835\udc3e\ud835\udc4f and \ud835\udc3e\ud835\udc61 are respectively the gravity and the motor parameters whose signification is given in Table (1). \ud835\udc37 = \ud835\udc51 + \u0394\ud835\udc53 + \u0394\ud835\udc54 means the sum of unknown external disturbances and model uncertainties" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0001248_epe.2013.6634388-Figure5-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0001248_epe.2013.6634388-Figure5-1.png", "caption": "Fig. 5: Eddy current loss density distribution.", "texts": [ " The magnetic permeability of vacuum is \u03bc0, and the relative permeability of conductor is \u03bcr, respectively. The frequencies of the flux in the inner and outer rotors are given by Eq. (4). The \u03bcr and \u03c3 of the Nd\u2013Fe\u2013B magnet used in the magnetic gear are 1.037 and 6.67\u00d7105 S/m, respectively. Thus, the skin depths of the inner and outer magnets are 14.7 mm and 47.1 mm, respectively. The radial and axial lengths of the inner and outer magnets are short enough in comparison with the skin depths. Therefore, the eddy currents are expected to be induced in the whole magnets. Fig. 5 shows the eddy current loss density of the rotor magnets calculated by the 3D\u2013FEA. It is understood that the eddy currents are induced in the whole outer magnet and on the surface of the inner magnet. The efficiency of the magnetic gear is calculated by using the torque and losses described above when the gear operates as a reduction gear. The efficiency \u03b7 is expressed as follows: (%)100\u00d7 ++ = eddyironout out WWP P\u03b7 , (7) where the sum of the iron losses of the pole pieces and the rotor yokes is Wiron, the eddy current loss of the magnets is Weddy, and the mechanical output of the magnetic gear is Pout, which is given by lloutP \u03c4\u03c9= , (8) where the angler velocity and the average torque of the outer rotor is \u03c9l and \u03c4l, respectively" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0001421_cdc.2010.5718183-Figure1-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0001421_cdc.2010.5718183-Figure1-1.png", "caption": "Fig. 1. Inverted rotary pendulum", "texts": [ " Furthermore, an additional error, due to the actuator time constant \u00b5 will cause an error of order O(\u00b5) ([9]). Now, the stabilization of the x` coordinate via nested backward compensation will be x\u0307` (t) = As`x` (t)+ \u03b5 (23) where As` is Hurwitz and \u03b5 = O(\u00b5)+O(\u03b4 \u03b1\u2212`). Thus we do not stabilize the origin, but we can achieve an ultimate bound that depends on the system time constants. Whenever is possible, this bound can be improved by reducing \u00b5, \u03b4 or by increasing the order of the differentiator. Let us consider an inverted rotary pendulum Fig. 1. An L-shaped arm, or hub, is connected to the DC motor shaft and pivots between \u00b1180 degrees. At the end of the arm, there is a suspended pendulum attached. The system state equations with x3 = x\u03071, and x4 = x\u03072, linearized along the point x = [ r \u03c0 0 0 ] are x\u0307 = 0 0 1 0 0 0 0 1 0 82.94 \u22121.31 0 0 56.81 \u22120.37 0 \ufe38 \ufe37\ufe37 \ufe38 A x+ 0 0 46.75 13.2 \ufe38 \ufe37\ufe37 \ufe38 u + 0 0 59.41 26.85 \ufe38 \ufe37\ufe37 \ufe38 D w (24) Let us consider a smooth unknown input w = .3sin4t + cos24t +0.3. The above system do not satisfy the condition span(D)\u2282 span(B)" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0000508_embc.2012.6346038-Figure4-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0000508_embc.2012.6346038-Figure4-1.png", "caption": "Figure 4. (a) Encapsulation Test Unit. (b) Details of Test Section.", "texts": [ " This system, shown in Figure 3, is capable of accelerated characterization of 100 test units simultaneously during periods of extended soaking in hot saline solution at variable temperatures. The test units for this system are uniquely designed to simulate the topographical, thermal, mechanical and electrical stresses, as well as the fabrication procedures, fully implantable active neurosensors. With the test units we are capable of monitoring continuously th tween interdigitated conductors on the substrate surface as well as leakage current through the encapsulation material. A picture of these test units is shown in Figure 4. These units are currently being empl ultiple types of polymeric encapsulation mat ses and co ations. It is our objective to build a trustworthy dat of material properties to share with the scientific comm e, LCP in oyed to test m erials, proces nfigur abase unity. Work with PDMS (but also Polyimide, Parylen , and SU8 among others) is underway, as shown gure 5 for three of the many PDMS encapsulated test units (here showing development of ionic leakage over period of H pa PDMS and re in-vivo system val Interfacing with External World \u2013 Progress in plantable Microelectronic Neuroengineering Devices,\u201d Proc" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0001158_074683410x480230-Figure1-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0001158_074683410x480230-Figure1-1.png", "caption": "Figure 1. Rolling the parabola y = x2 on the x-axis.", "texts": [ " They may also know that if the chain is subjected to a uniform load, then the chain assumes the form of a parabola; the shape of a suspension bridge, in which the bridge deck provides the load, is another example [2]. One connection between these two curves is that they arise under the similar physical conditions just mentioned. In this paper we establish a different connection. Suppose one were to roll the parabola y = x2 along the x-axis without slipping. How does its focus move? In other words, what is the locus of this focus? It turns out, as we will demonstrate, that the locus is a catenary! Suppose we roll the parabola y = x2 along the x-axis as shown in Figure 1. We assume that the parabola does not slip as it rolls and we wish to determine the path followed by the focus F(0, 1 4) of the parabola. To solve this problem we introduce variables as indicated in Figure 2. Here, \u03b81 = \u03b81(t) is the angle between the tangent line to the parabola at P(t, t2) and the x-axis; \u03b82 = \u03b82(t) is the angle between line FP and the x-axis; \u03b1(t) = \u03b81 \u2212 \u03b82 is the angle VOL. 41, NO. 2, MARCH 2010 THE COLLEGE MATHEMATICS JOURNAL 129 F(0,1/4) x y and the slope of line FP is tan \u03b82 = t2 \u2212 1 4 t " ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0002601_gt2017-64123-Figure8-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0002601_gt2017-64123-Figure8-1.png", "caption": "Figure 8. The parameterized solid model", "texts": [ " The developed parameterized model contains a minimum of simplifications and is equivalent to models used in standard gas-dynamic engineering analysis and in determining strength characteristics of centrifugal compressors impellers. Therefore, the impeller structure obtained as a result of optimization will not require further structural redesign. Parameterized solid sub-models description A parameterized 3D solid model, which includes two sub-models, is developed to automate the optimization process. The solid model is shown in Figure 8. One submodel describes the air and includes an intake part, the airgas channel of the impeller, and the vaneless diffuser outlet (see Figure 9). The intake configuration is not modified during the optimization. The other sub-model describes a centrifugal compressor impeller with its blades and disc (see Figure 10). Each sub-model includes a cyclic-symmetric sector with one short and one full-sized blade. The sub-models are coordinated with each other, so a modification of one submodel leads to corresponding changes in the other submodel" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0002605_s10015-017-0385-y-Figure2-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0002605_s10015-017-0385-y-Figure2-1.png", "caption": "Fig. 2 Grinding robot model", "texts": [ " The joint angles and angular velocities can be detected easily but the frictional force and frictional grinding coefficient that influence the contacting force control results are difficult to measure correctly. In this paper, the grinding resistance coefficient is obtained by experiments and it is confirmed that appropriate grinding control has been performed by compensating the influences from grinding resistance force to contacting force of the grinder. A photo of the experiment device is shown in Fig.\u00a01. A concept of grinding robot of constrained motion is shown in Fig.\u00a02. Constraint condition C is a scalar function of the constraint, and is expressed as an algebraic equation of constraints as where r(m \u00d7 1) is the position vector from origin of coordinates to tip of grinding wheel and q(n \u00d7 1) is joint angles. The grinder set at the robot\u2019s hand is in contact with the material that is to be ground. The equation of motion of grinding robot is modeled as following Eq. (2) [9\u201311]: where M is a n \u00d7 n matrix, h is centrifugal and coriolis force vector, D is viscous friction coefficient matrix, g is gravity vector" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0000169_j.robot.2010.03.012-Figure2-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0000169_j.robot.2010.03.012-Figure2-1.png", "caption": "Fig. 2. Omnidirectional vision system.", "texts": [ " In particular, the data processingmethoddescribedhere is responsible for filtering out the inherently noisy reading, obtained from the sensors, and it provides to the cooperative controlmethod stable and accurate estimates for the Self Evaluation. In the following, we detail how this process is realised using an omnidirectional vision system. We represent the robot team as a set of moving points {R1, R2, . . .} on a plane surface. Now, let us define the objects each robot perceives in the operating field, considering that the main sensor system is based on an omnidirectional camera where each object appears reflected on a conic surfacewith an angle \u03b8 , referred to the forward direction as appears in Fig. 2. So, \u03b8 varies on a 2\u03c0 range, namely, \u2212\u03c0 <= \u03b8 <= \u03c0 , and the object is positioned at a distance r from the origin of the frame of reference centered on the robot itself. We introduce on the arena a fixed frame of reference, and we consider the positions \u3008xi, yi\u3009 and \u3008xj, yj\u3009 of the robots Ri and Rj, respectively. The distance d between the points can be easily computed by means of the well-known Euclidean formula d2 = (xi \u2212 xj)2 + (yi \u2212 yj)2 = r2i + r 2 j \u2212 2rirj cos(\u03d5i \u2212 \u03d5j) = (ri \u2212 rj)2 + 2rirj(1\u2212 cos(\u03d5i \u2212 \u03d5j)) = (ri \u2212 rj)2 + rirj(\u03d5i \u2212 \u03d5j)2 where we have used both Cartesian and polar coordinates, referring to the standard translation formula, and the first term approximation of the cosine in the last equality" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0003211_j.apacoust.2018.12.039-Figure2-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0003211_j.apacoust.2018.12.039-Figure2-1.png", "caption": "Fig. 2. Single-flank test \u2013 measurement principle and evaluation [2\u20134].", "texts": [ " Furthermore, the manufacturing and the assembly process influence the operating gear geometry and gear set position and, hence, the excitation of the gear set. The transmission error is a quasi-static measure which summarizes the excitation mechanisms. The definition is given as the deviation from uniform relative angular motion of the driving and the driven gear considering the transmission ratio. Its appearance at the beginning and at the end of the contact mesh shows a parabolic shape in the time domain. The scheme for a transmission error measurement is illustrated in Fig. 2 and pictorially explains the definition of the transmission error. The rotational positions of the gears are determined by high-resolution angle encoders. The analog signal is transferred into a counter signal and finally calculated into the actual rotational angle of the gears. The calculation is defined according to Eq. (2-1). By multiplying the rotational angle u1 of the driving gear with the transmission ratio i, the theoretical rotational angle of the driven gear u2;calc is derived. Subtraction of the theoretical angle u2;calcand actual angle u2 amounts to the transmission error Du", " The transmission error is analyzed in the time domain in the evaluation specification of the single flank test. There are four different characteristic values defined in DIN 3965 [13] which evaluate the manufacturing quality of the topography and the pitch and run-out deviation in the time domain. The analysis in the frequency domain is recommended for the SFT of one hunting tooth, which is when each pinion tooth meshes with each gear tooth, amounting to a number of pinion revolutions equal to the number of gear teeth [11]. Typical graphs are displayed at the bottom of Fig. 2. 2.1. Manufacturing related excitation behavior The accuracy of the hard finishing process, influencing the gear geometry, is a major difference which determines the excitation behavior. The lapping process is a hard finishing process not only with undefined cutting edges but also with varying tool geometries provided by the meshing gear partners. This leads to a strong dependency of the final geometry on the geometry after the heat treatment. For a good excitation behavior, an incoming quality regarding pitch and run-out of DIN 7 is required" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0000165_jjap.49.04dk16-Figure2-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0000165_jjap.49.04dk16-Figure2-1.png", "caption": "Fig. 2. (Color online) Schematic cross-sectional structure of membrane switches in its off (a) and on (b) states.", "texts": [ " In this study, we succeeded in the further optimization of the fabrication process and device structure to achieve 16 16 matrix array that can be driven at a lower voltage. The results showed that the concept of the membrane switch can be extended easily for larger panels sizes that can be driven at even lower voltages. 2. Experimental Procedure 2.1 Device structure Figure 1 shows a schematic diagram of the display device proposed in this work. The device consisted of a membrane switch directly attached to an EPD panel. The EPD panel was commercially obtained from E Ink Co., while the membrane switch was newly developed in this study. Figure 2 shows the structure of the membrane switch in its off (a) and on (b) states. The membrane switch consisted of two polyimide (PI) films having Cu electrodes that were attached face-to-face with a wall spacer in between. One of the PI films was flexible and the other was fixed. The PI films had two sets of electrodes: electrostatic electrodes for controlling the on/off states, and signal electrodes for driving external devices such as EPDs. By applying voltage between the electrostatic electrodes, a flexible PI film could be electrostatically actuated close to the signal electrodes, as shown in Fig" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0003015_2017-36-0413-Figure6-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0003015_2017-36-0413-Figure6-1.png", "caption": "Figure 6. Rear bumper of AISI 1020 carbon steel.", "texts": [ " The results demonstrate the equivalent stresses for the steel and composite bumpers. Element chose and length definition The simulation was conducted to the limit of the computational program of the academic version of ANSYS. The type of element used was the SHELL281 and the bumper along with the vertical support showed 28,432 nodes. The tensile analysis at this degree of refinement was conducted for both the metallic material and the laminated composites. For the steel bumper, the deflection value found was 0.91mm. The equivalent tension was 112.4 MPa, as shown in Figure 6. The value lies within the material flow limit, defined as 420MPa. For the epoxy/fiberglass composite bumper, two simulations were performed using unidirectional laminate with 90\u00b0 and 0\u00b0 orientation for the fibers. In the first simulation, deflection results and equivalent stresses were obtained for the composite at 90\u00b0. The value obtained for deflection was 1.2 mm. The obtained equivalent tension was smaller in relation to the value of the limit of rupture for laminates in this orientation obtained in the experimental analysis of the author" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0003913_s00773-019-00675-8-Figure3-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0003913_s00773-019-00675-8-Figure3-1.png", "caption": "Fig. 3 Geometric relation for journal bearing", "texts": [ " Given that the after stern tube bearing is long and the centerline of the journal has a deflection, the after tube bearing is divided into several bearing segments, as illustrated in Fig.\u00a02. The midpoint of the journal centerline of the bearing segment is used as the node of the beam element in the calculation model of shafting alignment. To calculate the oil film force of each bearing segment, the nodes are grouped in axial order, and each bearing segment corresponds to a group. Thus, the Y-component Fyi and z-component Fzi of the oil film force for each bearing segment can be obtained. Figure\u00a03 displays that under the action of oil film forces Fx and Fy, journal center position O2 works in an eccentric position relative to bearing center O1. The eccentric distance 1 3 is e, the attitude angle is , the bearing inner radius is R1, the journal outer radius is R2, the clearance is c = R1 \u2212 R2 , and the center line is O1O2 . One end is the greatest clearance of oil film hmax = c + e , and the other end is the smallest clearance of oil film h min = c \u2212 e . The hydrodynamic lubrication film is formed by the relative motion between the inner surface of the bearing and the outer surface of the journal. The lubrication state generally belongs to hydrodynamic lubrication. Reynolds equation can be used to describe the viscous flow phenomena in this narrow gap. If the elastic deformation of the surface for the bearing and journal is ignored, then the cylindrical coordinates of the Reynolds equation for incompressible hydrodynamic lubricated bearings are as follows [15]: where h is the film thickness, p is the film pressure, and x is the length in the axial coordinate. Figure\u00a03 presents the other symbols. Let h = cH , p = 2\ud835\udf02\ud835\udf14 \ud835\udf132 p\u0304 , x = L 2 , z\u0307 = c\ud835\udf14z\ufffd , y\u0307 = c\ud835\udf14y\ufffd , e\u0307 = c\ud835\udf14y\ufffd , and ?\u0307? = \ud835\udf14\ud835\udf19\ufffd , where L is the effective length of the bearing and is the rotation angle from the y axis clockwise to this position. The dimensionless form of Eq.\u00a03 is as follows [15]: By using the small parameter method and omitting the high order, then [15]: (3) 1 r2 \ud835\udf15 \ud835\udf15\ud835\udf19 ( h3 \ud835\udf02 \ud835\udf15p \ud835\udf15\ud835\udf19 ) + \ud835\udf15 \ud835\udf15x ( h3 \ud835\udf02 \ud835\udf15p \ud835\udf15x ) = 6\ud835\udf14 \ud835\udf15h \ud835\udf15\ud835\udf19 + 12(y\u0307cos\ud835\udf19 + z\u0307sin\ud835\udf19) , (4) \ud835\udf15 \ud835\udf15\ud835\udf19 ( H3 \ud835\udf15p\u0304 \ud835\udf15\ud835\udf19 ) + ( D L )2 \ud835\udf15 \ud835\udf15\ud835\udf06 ( H3 \ud835\udf15p\u0304 \ud835\udf15\ud835\udf06 ) = 3 \ud835\udf15H \ud835\udf15\ud835\udf19 + 6 ( y\ufffdcos\ud835\udf19 + z\ufffdsin\ud835\udf19 )" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0003976_demped.2019.8864881-Figure1-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0003976_demped.2019.8864881-Figure1-1.png", "caption": "Fig. 1. Six-phase PMSM with a spatial shifting of /6 between three-phase windings.", "texts": [ " The faulty harmonic components in the emf are estimated through an analysis of the voltage references generated by the current regulators of the fieldoriented control. The feasibility and the effectiveness of the developed method are confirmed by some preliminary experimental results. The machine analyzed in this paper is an asymmetrical six-phase PMSM, which has two independent three-phase start-connected stator windings A1, A2, A3 and B1, B2, B3, rotated by 30\u00b0 degrees relative to each other (Fig. 1). The model of such a machine can be found by means of the Vector Space Decomposition (VSD), which is a linear transformation of the electromagnetic quantities that allows representing the machine as a set of decoupled equations in separated sub-spaces, or planes. A. Vector Space Decomposition The th space vector of the stator currents iSAk and iSBk (k=1,2,3) is defined as follows: 4 2113 1 SASBSAS iiii 9 3 8 3 5 2 SBSASB iii (1) where 6 j e . (2) Equation (1) has three notable properties. First, it is a periodic function of SS ii )12( ", " If the effect of the slot openings is neglected, the magnetic field produced by the rotor magnets can be written as a function of R as follows: odd j ReRR RehH ,1 (6) where 212 1 12 jjM R eeHh (7) AdM RM M BH 0 (8) and is the air gap length, M is the magnet thickness, BR is the residual flux density of the magnetic material and d is the differential permeability of the recoil curve. The magnetic field generated by the stator currents can be calculated in the stator reference frame as a function of the angular position S, expressed in electrical radians and whose origin is aligned with the magnetic axis A1, shown in Fig. 1: odd j SeSS SehH ,1 (9) where S wSS S i p KN h 2 1 1 3 (10) and NS is the number of series-connected conductors per phase, is the total air gap length, equal to A + M, p the number of pole pairs, KwS is the th winding coefficient. The relationship between the rotor angular coordinate R and the stator angular coordinate S is S = R + (11) where is the rotor position in electrical radians. Once the spatial distribution of the magnetic field generated by the stator current and the rotor magnets is known by means of (6)-(10), it is possible to calculate the flux linkages with the stator phases" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0002043_isr.2013.6695712-Figure1-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0002043_isr.2013.6695712-Figure1-1.png", "caption": "Fig. 1. Reference coordinates rotated on the basis of the angle of inclination.", "texts": [], "surrounding_texts": [ "The robot trajectory for the inclined surface is generated on the basis of reference coordinates that are determined by the slope. In F ig.1, coordinate frame {A} is established on the horizontal surface and coordinate frame {B} is established on the inclined surface. Thus, the expressions of gravity force are different depending on the selected reference coordinate frame. For example, the gravitational force acts on the mass in the negative Z- axis of reference coordinate {A}. But, it acts in the Z and X-axes of reference coordinate frame {B} . In Fig. (1), M is the total mass of the robot assumed to be one particle, zl is the position vector from the supporting point to the COM of the robot based on coordinate frame {B} , p\ufffdp is the position vector of the ZMP, pll is the vector from the ZMP to the COM of the robot, and e indicates the angle of inclination. y The gravitational force expressed with coordinate frame {B} is (1) where R: denotes the rotation matrix that defines the orientation between coordinate frames {A} and {B}. To generate the walking trajectory, the angular momentum equation about the ZMP is used. where -B -B -B P =q - PZMP' Z/=[Xyzt, -Il ] P 7MI' = [x7MI' Y 7MI' z7MI'] , gA=[OO - gy . From Eq. (2), .. 2 2 X = W X - w XZMP - g \u00b7siney \u2022\u2022 2 2 Y= w Y- w Y7MI' where w= g\u00b7cosey +io (Zo -ZZMP) (2) (3a) (3b) (4) If the robot moves straight forward, it can be assumed that Z(t) = Zo' where constant height. So, i(t) = io = 0 . Furthermore, Z7MI' is zero. In this paper, the desired ZMP trajectory in the X- and Y -directions during a double support phase is assumed to be a 3fd-order polynomial while it is fixed at the supporting foot during a single support phase . X7MI' = at3 +bt2 +ct+d (T, < t <0; T, + \ufffd) (Sa) YZMP=at3+bt2+ct+d (r; 0 and it determines the trade-off between the flatness of function and the amount up to which deviations larger than e are tolerated. This situation is shown graphically in Figure 2. The deviation of points outside the shaded region is penalized to contribute in a linear fashion. The dual formulation is as follows. min , 1 2 \u00f0 \u00de TQ \u00f0 \u00de \u00fe \"eT \u00fe \u00f0 \u00de yT \u00f0 \u00de \u00f019\u00de Subject to eT \u00fe \u00f0 \u00de \u00bc 0 04 i, i 4C, i \u00bc 1, . . . , n where i i is the dual coefficient which holds the dif- ference, Q is n by n positive semi-definite matrix, e is the vector of all ones, C4 0 is the upper bound, Qij K xi, xj \u00bc xi\u00f0 \u00de T xj is the kernel. The higher dimensional space function used for mapping training vectors" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0001487_6.2010-8418-Figure3-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0001487_6.2010-8418-Figure3-1.png", "caption": "Figure 3: Flight MAV with avionics (left) and wind tunnel MAV with mount for force balance (right)", "texts": [ " Force and moment measurements are conducted using the CU-1 balance, a six component internal balance custom built by Modern Machine and Tool Co. This balance was designed to measure MAV-scale loads with a maximum force of 3 lbs on the normal force axis and a maximum moment of 5 in-lbs on the pitching moment axis. Calibration of the CU-1 indicated maximum errors in the order of less than 0.5% of full scale loads. The MPS and CU-1 interfaced with the test section are shown in Figure 2. Figure 2: Model Positioning System mounted under the test section The CU MAV (shown in Figure 3) was designed using an iterative procedure based upon initial inviscid models and modifications derived from flight test results. It employs a cambered/reflexed airfoil and an inverse Zimmerman planform to enhance lift and two vertical tail surfaces to improve lateral stability. The root chord of the vehicle is 6.8 in and the maximum span is 8 in. The MAV flies at approximately 39 ft/s at a flight weight of 2.28 oz. It utilizes a front mounted propeller for thrust, although wind tunnel models were tested with only the airframe to assess the passive stability characteristics of the vehicle" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0003015_2017-36-0413-Figure8-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0003015_2017-36-0413-Figure8-1.png", "caption": "Figure 8. Composite rear bumper at 0\u00b0.", "texts": [], "surrounding_texts": [ "The analysis of the bumper, both for composite orthotropic material and for isotropic steel material, were performed under the same contour conditions and force application. The results demonstrate the equivalent stresses for the steel and composite bumpers." ] }, { "image_filename": "designv11_62_0003970_s11370-019-00294-7-Figure4-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0003970_s11370-019-00294-7-Figure4-1.png", "caption": "Fig. 4 Serial-to-RPR parallel conversion a serial manipulator, b derived parallel manipulator RPR", "texts": [ " 3 Inter\u2011kinematics Employing the characteristics observed in Sect.\u00a02.1.3 and inverse of matrix having axis coordinates (\u201cAppendix\u201d), the inter-kinematics between serial manipulator and parallel manipulator will be investigated. What to declare in this section is summarized as follows: (i) Geometry of serial manipulator is expressed by axis coordinate (ii) Geometry of its inversion is found as the geometry of parallel manipulator expressed in ray coordinate We start with the geometry of a 3-DOF serial manipulator given in Fig.\u00a04a. When the input magnitudes of 1, 2 and 3 are known, we can compute the output twist \u2322T . The forward kinematics for the serial manipulator is given by where and (22)S\u0302 ij = \u23a1 \u23a2\u23a2\u23a2\u23a3 yij \u2212xij 1 \u23a4 \u23a5\u23a5\u23a5\u23a6 . (23) |||s\u0302 T k S\u0302ij ||| = |||S\u0302 T ij s\u0302k ||| = qk (24)T\u0302 = J?\u0307?, (25)T\u0302 = \u239b\u239c\u239c\u239d vox voy \ud835\udf14z \u239e\u239f\u239f\u23a0 , (26)?\u0307? = \u239b\u239c\u239c\u239d \ud835\udf141 \ud835\udf142 \ud835\udf143 \u239e\u239f\u239f\u23a0 (27)J = \ufffd S\u03021 S\u03022 S\u03023 \ufffd = \u23a1\u23a2\u23a2\u23a3 Y1 Y2 Y3 \u2212X1 \u2212X2 \u2212X3 1 1 1 \u23a4\u23a5\u23a5\u23a6 . S\u03021 , S\u03022 and S\u03023 are the Pl\u00fccker coordinates of the lines that pass through the joint axes. They are the lines represented in axis coordinates", " = \u23a1 \u23a2\u23a2\u23a2\u23a2\u23a2\u23a2\u23a2\u23a2\u23a2\u23a3 s\u0302T 23 s\u0302T 23 S\u03021 s\u0302T 31 s\u0302T 31 S\u03022 s\u0302T 12 s\u0302T 12 S\u03023 \u23a4 \u23a5\u23a5\u23a5\u23a5\u23a5\u23a5\u23a5\u23a5\u23a5\u23a6 T\u0302 . \u21d2 Tu = \ufffd s\u0302 23 s\u0302T 23 S\u03021 s\u0302 31 s\u0302T 31 S\u03022 s\u0302 12 s\u0302T 12 S\u03023 \ufffd T\ud835\udf19, (29) v1 = \ud835\udf141s\u0302 T 23 S\u03021 v2 = \ud835\udf142s\u0302 T 31 S\u03022 v3 = \ud835\udf143s\u0302 T 12 S\u03023, (30)v = jT T\u0302 , (31)v = \u239b\u239c\u239c\u239d v1 v2 v3 \u239e\u239f\u239f\u23a0 1 3 It is remarked that each row of (32) denotes a vector connecting two joints of serial manipulator, which is actually ray coordinate representation. Resultantly, the inverse geometry of the serial manipulator can be represented as geometry of RPR parallel manipulator given in Fig.\u00a04b. In summary, when the instantaneous twist T\u0302 of parallel platform is specified, the actuator speeds v1, v2, v3 can be computed as (30). The geometry of parallel manipulator converted from serial manipulator can be diverse. Here, RRR-type parallel manipulator is considered. The linear velocities v1, v2, v3 along the direction of each second link can be obtained from the angular velocities ( 1)1, ( 1)2, and ( 1)3 of the first joint of each chain of the 3-DOF RRR parallel manipulator shown in Fig.\u00a05" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0001609_0954406211427833-Figure10-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0001609_0954406211427833-Figure10-1.png", "caption": "Figure 10. Simulated centre tooth of a convex gear.", "texts": [ " While simulating the machining of convex gear, gear blank axis of revolution tilts at 6 with respect to the cutter axis. For centre tooth, the surrounding material of the tooth is removed with gear blank axis of revolution in this position. For machining the teeth at the inner and outer rings, the work solid is tilted by 18 and 30 about the X-axis (Figure 9(b) and (c)). Indexing for each tooth is done for simulation machining of other teeth at the inner and outer rings. CAD model after simulation for the centre tooth of the convex gear is shown in Figure 10. The crests are clearly seen on the tooth surface as well as on the surface of the cavity surrounding the tooth. Figure 11 shows the X-deviation of the simulated tooth profile from the analytical profile and the maximum deviation of 37 mm is observed near the bottom portion. Generation machining of spherical gear pair For machining conventional gears, standard gear cutting machines and cutters are available. Though spherical gears have a standard involute profile, it cannot be machined using the conventional gear cutting machines" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0000199_icma.2012.6283230-Figure1-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0000199_icma.2012.6283230-Figure1-1.png", "caption": "Fig. 1. A 2-DOF rigid-link robotic manipulator", "texts": [ " = \ud835\udc34\ud835\udc65+\ud835\udc35\ud835\udc63 \ud835\udc66 = \ud835\udc36\ud835\udc65 (28) where the new input \ud835\udc63 is given by \ud835\udc63 = ( \ud835\udc531(\ud835\udc65, \ud835\udc61) \ud835\udc532(\ud835\udc65, \ud835\udc61) ) + ( \ud835\udc541(\ud835\udc65, \ud835\udc61) \ud835\udc542(\ud835\udc65, \ud835\udc61) ) \ud835\udc62+ ( \ud835\udc511 \ud835\udc512 ) (29) The robotic model of Eq. (28) can be written in discrete-time form after applying common discretization methods, and next state estimation can be performed using the standard Kalman Filter recursion, as described in Eq. (2) and Eq. (3). The performance of the proposed derivative-free nonlinear Kalman Filter was tested in the benchmark problem of state estimation-based control for a 2-DOF rigid-link robotic manipulator (Fig. 1). The differentially flat model of the robot and its transformation to the Brunovksy form has been analyzed in Section IV. It was assumed that only measurements of the angle of the robot\u2019s joints could be obtained through the robot\u2019s encoders. The reference set-point for each joint was a sinusoidal signal of amplitude 1.0 and period \ud835\udc47 = 10\ud835\udc60\ud835\udc52\ud835\udc50. At the beginning of the second half of the simulation time an additive sinusoidal disturbance of amplitude \ud835\udc34 = 0.5 and period \ud835\udc47 = 10sec was applied to the system" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0003124_itec-india.2017.8333719-Figure3-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0003124_itec-india.2017.8333719-Figure3-1.png", "caption": "Fig. 3. Relative position of space phasors", "texts": [ " So space phasor of back-emf can be represented as leading \u03a8s by 90\u25e6 in the direction of motor rotation. Assume R-Y-B winding be distributed in space as shown in Fig. 2 to determine the space phasor diagram relating line-line back emf vector and flux space vector. Line voltage between R-phase and Y-phase (vry) leads R-phase voltage (vr) by 30\u25e6 in space, irrespective of direction of motor rotation. Thus vll phasor leads vph by 30\u25e6 in space. Relative position of line voltage and flux space phasors when motor is rotating in anti-clockwise and clockwise direction is shown in Fig. 3. Thus by measuring line voltage of the motor, rotor position of the PMSM can be located exactly. Let \u03b8ll be the position of vll. When the motor is rotating in anti-clockwise direction, rotor flux vector lags vll by 120\u25e6. Sine and cosine components of rotor flux position obtained from measuring back-emf (sin \u03b8emf , cos \u03b8emf ) in terms of line-line voltage space vector position (\u03b8ll) are given by sin(\u03b8emf ) = sin(\u03b8ll \u2212 120\u25e6) = \u22120.5 (sin \u03b8ll + \u221a 3 cos \u03b8ll) (18) cos(\u03b8emf ) = cos(\u03b8ll \u2212 120\u25e6) = 0.5 ( \u221a 3 sin \u03b8ll \u2212 cos \u03b8ll) (19) Similarly, when the motor is rotating in clockwise direction, rotor flux vector position can be computed as sin(\u03b8emf ) = sin(\u03b8ll + 60\u25e6) = 0" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0002029_978-3-319-02609-1-Figure3.1-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0002029_978-3-319-02609-1-Figure3.1-1.png", "caption": "Fig. 3.1 (a) The Alice 2002 mobile robot has a size of 21 mm\u00d721 mm\u00d720 mm and is equipped with four infrared sensors (rear sensor not visible on the picture) for crude proximity sensing and communication. (b) Simulated counterpart of the Alice robot implemented in Webots (see Section 3.2.1).", "texts": [ " In this particular case study, the ground truth is provided by realistic simulations implemented in Webots. Case Study II This case study is similar to the previous one, with the exception of the geometry of the aggregates. Indeed, rather than allowing for an uncontrolled aggregation, the robots may now decide not to aggregate with each other if the resulting bond does not satisfy some geometric constraints that are evaluated from sensory data. In particular, they require a strong activation of their front proximity sensor as well as a relatively low activation of their two side sensors (Figure 3.1a), thereby leading to the formation of chains of variable length. One of the research questions tackled by this case study is whether one can control the size of these chains, and what is the best approach to do so. First, we consider a baseline controller in which the robots, once aggregated, never disaggregate (Figure 4.1). As discussed later (Section 10.2), this behavior invariably leads to an exponential-like distribution of chain sizes, which we call the trivial distribution. In order to tune the system in a non-trivial fashion, we propose two orthogonal approaches: (1) a deterministic controller where robots communicate with each other in order to determine the size of each chain, and adapt their behavior accordingly, and (2) a probabilistic controller where chain size is controlled by the probability that a robot will leave an aggregate" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0001520_aim.2010.5695789-Figure16-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0001520_aim.2010.5695789-Figure16-1.png", "caption": "Fig. 16. Structure of the prototype of the robot", "texts": [ "2 MPa, it is thought that the tube that has three layers generates losses in transforming the expansion of the tube into the axial stretch. This is because there is the space between the second layer and the third layer. Therefore, the expansion of the tube does not transmit to the third layer efficiently. However, it is thought that the pressure level of 0.3 MPa is within the realistic range. Therefore, the film made of the thermoplastic is adopted as the second layer. In this case, the outer diameter of this tube actuator with three layers is 12mm. Fig. 15 shows the overall view of the prototype of the robot, and Table I and Fig. 16 show its specification and its structure. In this robot, the pinch rollers are installed in the head unit since cutting off the route for the fluid by them is better in the case of the tube which its shape is similar to a pipe. In addition, the guide box is installed in order to lead the tube into the pinch roller, and a small heater is installed in its internal perimeters. In order to shorten the time that is needed to steer the robot in the desired direction, a small cartridge heater is adopted" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0002163_s10483-017-2182-6-Figure5-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0002163_s10483-017-2182-6-Figure5-1.png", "caption": "Fig. 5 Elastic ribbon with elliptical cross section", "texts": [ " Figure 3 also shows that all curves intersect at a point (1, 35\u25e6). This is because that, the pitch angle \u03d5 is independent of Poisson\u2019s ratio when \u03b7 = 1, as indicated by Eq. (24). For a given elastic ribbon, the intrinsic twisting \u03c90 3 is fixed. Therefore, we discuss the pitch angle changing with the twist rate of the ribbon (see Fig. 4). The result indicates that the pitch angle increases with an increase in the twist rate, which is consistent with the helical spring[45]. If the cross section is an ellipse with the major axis b and the minor axis a (see Fig. 5), the bending rigidity and twisting rigidity of the ribbon can be written as[46] B = 1 4 E\u03c0ab3, C = G\u03c0 a3b3 a2 + b2 . Let b = ka, where k is the shape factor characterizing the axial ratio of a symmetrical cross section, and Eq. (23) be rewritten as follows: \u03c6 = arctan ( 2 + 2 (1 + \u03b7)2 \u2212 4 (1 + k2)(1 + \u03bd) )\u2212 1 2 . (25) Clearly, the cross section is circular if k = 1. Take \u03bd = 0.23. Then, the pitch angle \u03d5 is completely determined by the values of \u03b7 and k. When k is given, the change of \u03d5 with \u03b7 is shown in Fig" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0001374_eeeic.2011.5874721-Figure3-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0001374_eeeic.2011.5874721-Figure3-1.png", "caption": "Figure 3. Flux distribution and density shadow due to excitation of phase a in case of the noneccentric air-gap.", "texts": [ " To investigate the effect of dynamic eccentricity on the 6/4 switched reluctance behaviour, the motor is simulated utilizing 2-D finite element analysis by Finite Element Method Magnetics (FEMM) package1. Using finite elements method is a priori appealing for solving complex problems with a better accuracy. Iron material was used in the structure of the stator and rotor cores with the following static B-H curve shown in Appendix. Number of turns in each phase equal to 120 and the winding of phase was excited with a current magnitude of 2A. Fig.3 and fig.4 , reveals the magnetic flux distribution and density shadow in the case of the healthy machine and in the case of 40% dynamic eccentricity, respectively. From the comparing of the healthy and faulty case we observe that the flux density has increased with increasing the relative dynamic eccentricity. The main data of the studied switched reluctance machine are given in Table I. The static torque and the magnetizing flux linkages are obtained at different rotor positions from 0 to 90\u00b0 taking rotational steps of 5\u00b0 where the rotor moves from unaligned to fully aligned position" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0003679_j.mechmachtheory.2019.05.026-Figure6-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0003679_j.mechmachtheory.2019.05.026-Figure6-1.png", "caption": "Fig. 6. Illustration of relation between neighboring involution curves.", "texts": [ " When using non-standard auxiliary rack cutters to cut the same gear, three parameters must be adjusted, namely (i) the pitch radius, (ii) the pitch position of the auxiliary rack cutter, and (iii) the tooth width of the auxiliary rack cutter. When cutting a specific gear, the pitch radius r \u2032 of the non-standard auxiliary rack cutter is determined as r \u2032 = r \u2032\u2032 ( C\u03b1\u2032\u2032 C\u03b1\u2032 ) = m \u2032\u2032 n g 2 ( C\u03b1\u2032\u2032 C\u03b1\u2032 ) , (19) where r \u2032\u2032 = m \u2032\u2032 n g / 2 is the standard pitch radius, and the displacement of the pitch line of the non-standard rack cutter from the standard pitch line is given by r \u2032 = r \u2032\u2032 \u2212 r \u2032 = m \u2032\u2032 n g 2 ( 1 \u2212 C\u03b1\u2032\u2032 C\u03b1\u2032 ) . (20) Referring to Fig. 6 , the angle \u03c6b can be expressed as \u03c6b = \u03c0 n g \u2212 2 \u00b7 inv \u03b1\u2032\u2032 , (21) where inv is the involute function, and is defined as inv \u03b1 = tan \u03b1 \u2212 \u03b1. (22) The interval angle between neighboring teeth is denoted as \u03c6\u2032 pitch and is given as \u03c6\u2032 pitch = \u03c6b + 2 \u00b7 inv \u03b1\u2032 . (23) Hence, the tooth thickness of the rack t \u2032 rack is obtained as t \u2032 rack = r \u2032 \u03c6\u2032 pitch . (24) Substituting Eqs. (20) to (23) into Eq. (24) , the tooth thickness of the rack t \u2032 rack can be further determined as t \u2032 rack = m \u2032\u2032 n g ( C\u03b1\u2032\u2032 C\u03b1\u2032 )( \u03c0 2 n g \u2212 inv \u03b1\u2032 + inv \u03b1\u2032\u2032 ) " ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0002200_978-3-319-51222-8-Figure6.17-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0002200_978-3-319-51222-8-Figure6.17-1.png", "caption": "Fig. 6.17 Stress distribution (\u03c3xx ) at the fixed edge when the corners are rounded out", "texts": [], "surrounding_texts": [ "To generate in an FE-model the influence function for \u03c3xx , we would apply the stresses \u03c3xx of the shape functions at the corner point as equivalent nodal forces. Since only the shape functions of the corner element itself have nonzero stresses at the corner point, only the four nodes of this element will carry nodal forces fi and these are proportional to E/h where E is the modulus of elasticity, E = 2.1 \u00b7 105 N/mm2 (steel), and h is the element length. Numerical tests confirmed indeed an exponential increase in the displacements when the element size h tends to zero, h \u2192 0, see Fig. 6.14. Imagine that we replace the main stresses by pairs of orthogonal arrows (\u2018stream lines\u2019), see Fig. 6.15a, b. In each cross section, the vector sum of these arrowsmust be equal to the applied load. Now, the more these streamlines approach the left edge, the more they flatten because the edge is fixed in vertical direction, and consequently, the 302 6 Singularites" ] }, { "image_filename": "designv11_62_0002311_ls.1385-Figure10-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0002311_ls.1385-Figure10-1.png", "caption": "FIGURE 10 Analysis diagram of shaft under dynamic load", "texts": [ " When the eccentricity is constant, the oil film damping is deduced by the influence of velocity slip. And the more extent of velocity slip, the more reduce ratio of oil film damping. The effect law of Cxx influenced by velocity slip is comparatively complex. When the eccentricity is different, the effect law is different accordingly. The calculated oil film pressure is used in the description of rotor motion behavior, so the analysis of fluid\u2010structure coupling of hydrostatic bearing\u2010rotor system is realized. In Figure 10, O is the center of the bearing position, O' is the center of the rotor position. The dynamic load acting on the rotor in x and y directions are Qx and Qy, respectively; the corresponding oil film force in x and y directions are Fx(\u03c90t) and ur figure can be viewed at wileyonlinelibrary.com] Fy(\u03c90t), respectively. Mg is the gravity of the rotor. Then the motion equation of the rotor can be written as M x\u2022\u2022 \u00bc Fx \u03c90t\u00f0 \u00de \u00fe Qx; (32) M y\u2022\u2022 \u00bc Fy \u03c90t\u00f0 \u00de \u00fe Qy \u00feMg: (33) Assumed that the rotor is influenced by single dynamic load effect of eccentric mass, eg is the eccentric quality distance of the rotor, so unbalance load acting on rotor are Qx \u00bc Meg \u03c90 2 sin \u03c90t\u00f0 \u00de; (34) Qy \u00bc Meg \u03c90 2 cos \u03c90t\u00f0 \u00de: (35) According to the Formulas 19 and 20, the oil film force Fx(\u03c90t) and Fy (\u03c90t) can be replaced by oil film dynamic stiffness and damping coefficient, so the motion equation of the rotor can be written as M x\u2022\u2022\u00fekxxx\u00fe kyyy\u00fe cxx x \u2022\u00fecxy y \u2022 \u00bc Meg \u03c90 2 sin w0t\u00f0 \u00de; (36) \u2022 \u00bc Meg \u03c90 2 cos w0t\u00f0 \u00de \u00feMg: (37) The dimensionless form are M0 X \u2022\u2022\u00feKxx X \u00fe Kxy Y \u00fe Cxx X \u2022 \u00feCxy Y \u2022 \u00bc M0\u03b5g sin\u03c4; (38) M0 Y \u2022\u2022 \u00fe Kyx X \u00fe Kyy Y \u00fe Cyx X \u2022 \u00fe Cyy Y \u2022 \u00bc M0\u03b5g cos\u03c4 \u00feMg \u2014 ; (39) where equivalent qualityM0 \u00bc M\u03c90 3\u03b7L h0 R 3 ; eccentric quality in dimensionless form is \u03b5g \u00bc eg h0 ; equivalent gravity of the rotor is M \u2014 g \u00bc Mg 3\u03b7LR\u03c90 h0 R 2 " ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0000470_systol.2010.5675998-Figure1-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0000470_systol.2010.5675998-Figure1-1.png", "caption": "Fig. 1: Components of the Twin Rotor MIMO System", "texts": [ " Description of Twin-Rotor MIMO System The TRMS is a laboratory setup developed by Feedback Instruments Limited for control experiments. The system is perceived as a challenging engineering problem due to its high non-linearity, cross-coupling between its two axes, and inaccessibility of some of its states through measurements. The TRMS mechanical unit has two rotors placed on a beam together with a counterbalance whose arm, with a weight at its end, is fixed to the beam at the pivot and it determines a stable equilibrium position (Fig.1). The TRMS can rotate freely both in the horizontal and vertical planes by changing the input voltage of the DC motors that drive the rotational speed of the (tail and main) rotors. The system input vector is uk = [Uh,k, Uv,k]T where Uh and Uv are the input voltages of the tail and main motor, respectively. The system states are xk = [iah,k, \u03c9h,k, \u2126h,k, \u03b8h,k, iav,k, \u03c9v,k, \u2126v,k, \u03b8v,k]T with iah/av is the armature current of tail/main rotor, \u03c9h/v is the rotational velocity of the tail/main rotor, \u2126h/v is the angular velocity around the horizontal/vertical axis and \u03b8h/v is the azimuth/pitch angle of beam" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0002176_j.procs.2017.01.179-Figure1-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0002176_j.procs.2017.01.179-Figure1-1.png", "caption": "Fig. 1. Examples of tensegrity structure.", "texts": [ " Furthermore, a calculation method for balancing internal forces among wires is proposed. c\u00a9 2016 The Authors. Published by Elsevier B.V. Peer-review under responsibility of organizing committee of the 2016 IEEE International Symposium on Robotics and Intelligent Sensors (IRIS 2016). Keywords: tensegrity structure, locomotion, wire driven, cable 1. Introduction The tensegrity structure consists of combinations of multiple rods and tensile materials (e.g. wires), and maintains its shape through a balance among the rods and tensile materials (See Fig. 1). The tensegrity structure was originally studied for applications in the architectural field because it can easily have high stiffness despite of being lightweight [1,2,3]. In recent years, studies have reported tensegrity applications in robotics. Generally, these tensegrity robots can drastically change their shape by changing the lengths of rods and/or wires [4]. For example, Paul et al. [5] developed a tensegrity robot by using rubber elastic cables as tensile materials. Sabelhaus et al. [6] developed a tensegrity robot by using struts, actuated spring-cables, and passive spring-cables for rolling locomotion", " rg/licenses/by-nc-nd/4.0/). Peer-review under responsibility of organizing committee of the 2016 IEEE International Symposium on Robotics and Intelligent Sensors(IRIS 2016). prevent slackening of the wires and maintain system stiffness. Therefore, the calculation of the balancing internal force for a desired shape is necessary to control the shape precisely. The contributions of this study are as follows: (i) the designing of a basic motion experiment of the prototype with three rods and nine wires, as shown in Fig. 1(a); (ii) proposal and verification of a numerical calculation method of the balancing internal force among wires. The remainder of this paper is organized as follows. Section 2 describes the aspect of the prototype robot and the experimental result of basic motion control. Section 3 demonstrates the calculation method of the balancing internal force. Finally, Section 4 concludes the discussion. Fig. 2 shows the proposed prototype, which consists of three rods and nine wires. Although the lengths of the rods are fixed, the lengths and/or tension of the nine wires can be changed using the nine actuators in the three rods" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0000302_28001-Figure1-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0000302_28001-Figure1-1.png", "caption": "Fig. 1: (Colour on-line) (a) Illustration of a multi-filament configuration showing the interaction of the i-th microtubule (marked in red) with all other microtubules in motor contact. Respective forces acting on this filament are sketched as arrows. (b) Motor-mediated, binary microtubule interaction: motors attach at the rods\u2019 intersection point, zip across the rods causing alignment, and detach.", "texts": [ " Model of filament organization. \u2013 A semi-dilute solution of microtubules (or short actin filament bundles) interacting via molecular motors is modeled as a collection of N stiff rods of fixed length L. For the sake of simplicity (and since most experiments are carried out in a quasi\u2013two-dimensional geometry) we restrict our modeling to two spatial dimensions; the orientation of filament i can be described by a unit vector ni = (cos\u03d5i, sin\u03d5i), or by the angle \u03d5i (with respect to the x-axis, see fig. 1). The position of the center of mass of the i-th filament is denoted by ri. For a system of i= 1 . . . N interacting filaments, we can write the equations of motion in the following Langevin-type form: \u03d5\u0307i = \u03b6 \u22121 r [ lmni\u00d7 \u2211 j,i\u2229 j nb\u22a5ij F m ij + \u03be r i ] , (1) r\u0307i =M(\u03d5i) \u03b6 \u22121M(\u2212\u03d5i) [ \u2211 j,i\u2229 j nb\u22a5ij F m ij + \u03bei ] . (2) Equations (1) and (2) represent torque and force balances, respectively; Fmij is the magnitude of the force due to motor action on the filament pair (i, j) and is specified below. The force is assumed to be perpendicular to the bisecting line of two intersecting filaments (i, j); i", " For the motor force we write Fmij = \u03c3ij ( \u03ba lm + \u03be m j ) sin(\u03c8ij), (3) with \u03c8ij = \u03d5i\u2212\u03d5j 2 the angle between a filament and the bisecting line of the pair. \u03c3ij is the number of motors spanning the filament pair (i, j). The first term in brackets is the average (i.e. deterministic) strength of the motor force and the stochastic term \u03bemj describes fluctuations around the average. These fluctuations, unlike the thermal noise represented by \u03beri and \u03bei, are intrinsically multiplicative. For simplicity, we have assumed that every intersection of two filaments is symmetric; i.e. that motors act perpendicular to the bisecting line, see fig. 1b). For filament pair interactions, such configurations are rapidly achieved (see the discussions in refs. [20,25,28].) The number of motors \u03c3ij of the pair (i, j) includes an additional exponential form, \u03c3ij = \u03c30 exp(\u2212Est/kBT ), where \u03c30 is the average number of motors per microtubule. The argument of the exponential represents the ratio of the motor stretching energy, Est = \u03baal| sin(\u03d5i\u2212\u03d5j2 )|, to the thermal energy. This form is motivated by experimental studies [29] and was developed in ref. [30]" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0002172_j.proeng.2017.02.183-Figure1-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0002172_j.proeng.2017.02.183-Figure1-1.png", "caption": "Fig. 1. Kinematic diagram of the worm (1) \u2013 wormwheel (2) technological gear with the axis-wise motion of the worm (tool action surface).", "texts": [ " In the machining process, after the mill has been \"cut in\" to the full depth by the tangential or radial method, the tool (its action surface) and the wormwheel being machined form a worm\u2013wormwheel technological gear [6-9]. To simplify further discussion, the worm\u2013wormwheel gear (both constructional and technological), in which rotational motion occurs around the axis of both gear links, can be substituted in theoretical discussion with its equivalent gear, in which axis-wise worm motion takes place without rotation \u2013 Figure 1 [10, 11]. In order to define the wormwheel tooth surface as the envelope of tool action surface, transition should be made from the worm {1} (tool) coordinate system to the wormwheel {2} coordinate system \u2013 Figure 1, which can be generally written with the following equation T 12 p,z1,aa,3,1i,3 xx (1) where - relative tool and wormwheel turning motion parameter; - angle between the tool and wormwheel rotation axes; p - tool helical action surface parameter; a - tool and wormwheel axis distance for machining by the tangential method; a - axis distance error with the tangential method or the quantity allowing for the tool cutting into the machined wormwheel with machining by the radial method; z, - tool positioning errors; i - worm gear transmission ratio" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0001838_ma302559w-Figure15-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0001838_ma302559w-Figure15-1.png", "caption": "Figure 15. Schematic diagram showing the sheared BCC lattice with the lattice strain \u03b3lattice = +0.4 for A-sphere (part a) and A\u2032-sphere (part b), where blue dotted lines and red solid lines represent the lattice before and after the shear deformation. The parts (c) and (d) present simplified lattice for A- and A\u2032-sphere, respectively, and schematic illustration of the conformations of the bcp chains in the matrix subjected to the shear deformation. The shear deformation of the spheres are ignored for the sake of simplicity.", "texts": [ " This unequal lattice deformation causes the asymmetric relaxation of the deformed lattice at strain phase 4, which in turn causes A\u2032-sphere be relaxed and visible, as shown by the unhatched A\u2032-sphere in parts (a-4) and (b-4), but Asphere be under a stressed state with some lattice distortions and invisible as schematically shown by the hatched A-sphere. This accounts for the large stress level shown by the point P4 in Figure 7d. IV-6. Physical Factors Controlling the Unequal Lattice Deformation for A-Sphere and A\u2032-Sphere in A/A\u2032-Twin. We can estimate the lattice strain imposed on {110} and {200} lattice spacing, \u03b5110 and \u03b5200, respectively, when A- and A\u2032sphere are subjected to the \u03b3lattice = +0.4, as shown in Figure 15, where \u03b5 = \u2212d d d( )/110 110,d 110,u 110,u \u03b5 = \u2212d d d( )/200 200,d 200,u 200,u and dhkl,u and dhkl,d are the lattice spacing of (hkl) plane before and after imposing the given lattice shear strain \u03b3lattice. Figures 15a and 15b show the BCC unit cells for A and A\u2032-sphere, respectively, before (blue broken lines) and after the given shear strain on the respective lattice (red solid lines). The estimated values for \u03b5110 and \u03b5200 are shown in Table 1, where the positive and negative values indicate the expansion and contraction of the lattice spacing", " In the calculation of \u03b51 and \u03b52, we ignored deformation of the spheres for the sake of simplicity. The results are also shown in Table 1 where the positive and negative values indicate the expansion and contraction, respectively. The values \u03b51 and \u03b52 intimately involve the deformation of the corona PI block chains emanating from the PS spheres and hence are related to the conformational entropy loss and entropy elasticity of the corona chains as shown schematically in the simplified A-sphere lattice (Figure 15c) and A\u2032-sphere lattice (Figure 15d). The results in Table 1 elucidate the following pieces of important information. Under the given lattice shear strain, \u03b3lattice = +0.4, imposed on the lattice for A- and A\u2032-sphere, A- sphere is subjected to smaller strains on \u03b5110 and \u03b5200 than A\u2032sphere but larger strains on \u03b51 and \u03b52. Thus, in the case when \u03b3lattice > 0, A-sphere is favored from the viewpoint of the elastic energy contribution to the lattice deformation. However, Asphere is favored from the viewpoint of the entropic contribution to the deformation", " In Block Copolymer Science and Technology; Meier, D. J., Ed.; MMI Press Symposium Series; Harwood Academic Pub.: London, 1983; Vol. 3, pp 63\u2212108. (31) Sakurai, S.; Momii, T.; Taie, K.; Shibayama, M.; Nomura, S.; Hashimoto, T. Macromolecules 1993, 26, 485\u2212491. (32) Free energy (F0) of a single bridge chain of PI block emanated from the interfaces of the nearest-neighbor PS spheres is given by F0 \u221d H0 2, where H0 is the end-to-end distance of the bridge chain. When H0 changes to H by the shear strain on the lattice, \u03b3lattice, as shown in Figure 15, the corresponding free energy changes from F0 to F, where F \u221d H2. Thus, the relative free energy change due to \u03b3lattice defined by \u0394F/F0 \u2261 (F \u2212 F0)/F0 is given by (H2 \u2212 H0 2)/H0 2 = |\u03b5|(|\u03b5| + 2), where \u03b5 \u2261 (H \u2212 H0)/H0. The value of \u03b5 along the direction 1 (defined by \u03b51,K) and that along the direction 2 (defined by \u03b52,K) for K-sphere (K = A or A\u2032) is given by Table 1, where the directions 1 and 2 are defined dx.doi.org/10.1021/ma302559w | Macromolecules 2013, 46, 1549\u221215621561 in Figure 15. Thus, the ratio of the relative free energy change for Asphere [defined by (\u0394F/F0)A] and A\u2032-sphere [defined by (\u0394F/F0)A\u2019] with respect to the lattice deformation is given by (\u0394F/F0)A/(\u0394F/ F0)A\u2019 = [2|\u03b52,A|(|\u03b52,A| + 2) + |\u03b51,A|(|\u03b51,A| + 2)]/[2|\u03b51,A\u2032|(|\u03b51,A\u2032| + 2) + |\u03b52,A\u2032| (|\u03b52,A\u2032| + 2)] = 1.4. Thus, the free energy change of the bridging chains due to the lattice deformation for A\u2032-sphere is smaller than that for the A-sphere. dx.doi.org/10.1021/ma302559w | Macromolecules 2013, 46, 1549\u22121562156" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0002791_b978-0-12-812138-2.00003-9-Figure3.60-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0002791_b978-0-12-812138-2.00003-9-Figure3.60-1.png", "caption": "FIGURE 3.60", "texts": [ " The d and q axes synchronous reactances are given by Xd 5Xad 1Xal (3.83) Xq 5Xaq 1Xal (3.84) where Xal is the armature leakage reactance. The salient pole synchronous motor always Xd .Xq, normally Xq 5 \u00f00:5B0:8\u00deXd. Considering the d and q axes synchronous reactances, and referring to Eq. (3.75), the voltage equation of the salient pole rotor synchronous motor can be expressed by Vs 5 IsRs 1 Id jXd 1 Iq jXq 1Ef (3.85) From this equation, the equivalent circuit of the salient pole rotor synchronous motor is shown in Fig. 3.60. Now, we will derive the torque of a salient pole rotor synchronous motor. In addition to the torque of a cylindrical rotor, it can be expected that a salient pole rotor synchronous motor has a reluctance torque due to the saliency of the rotor as already discussed in Chapter 1. Salient pole rotor synchronous motor. phasor diagram can be expressed as Fig. 3.61. From this phasor diagram, we have Vscos \u03b45Ef 1 IdXd (3.87) Vssin \u03b45 IqXq (3.88) Combining Eqs. (3.86) (3.88), the output power and the torque are given by P5 3 VsEf Xs sin \u03b41 3 V2 s \u00f0Xd 2Xq\u00de 2XdXq sin 2\u03b4 (3" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0001606_s11044-010-9211-1-Figure7-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0001606_s11044-010-9211-1-Figure7-1.png", "caption": "Fig. 7 The Tippe\u2013Top geometry", "texts": [ " The tangent forces model being represented here has been verified in two stages: (a) for the case of circular contact; (b) for the case of elliptic noncircular contact. The known Tippe\u2013 Top dynamical model was investigated as an example of the first case. All parameters and initial conditions are exactly the same as in paper [2] whose authors obtained their data in turn from the work [23]. Our case differs only in that the depth of penetration and normal force are computed dynamically during the simulation run thus forming an unrestricted contact problem. The top body, is assumed geometrically rigid, composed of two balls, Fig. 7, one of larger radius R = 1.5 \u00d7 10\u22122 m, and another, smaller, one of radius r = 0.5 \u00d7 10\u22122 m. The top mass center location is assumed to lie \u201cunder\u201d the larger ball geometric center on its axis of symmetry at a distance of a0 = 3 \u00d7 10\u22123 m and at a distance a1 = 16 \u00d7 10\u22123 m from the smaller ball center. The top mass is equal to m = 6 \u00d7 10\u22123 kg. The top body is assumed dynamically symmetric, and the central principal moments of inertia are the following: an equatorial moment equals to 8\u00d710\u22127 kg m2, and a polar one has the value of 7\u00d710\u22127 kg m2" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0003356_978-981-13-5799-2_12-Figure12.46-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0003356_978-981-13-5799-2_12-Figure12.46-1.png", "caption": "Fig. 12.46 New pattern geometry for a formula one tire [44]", "texts": [], "surrounding_texts": [ "(1) Application of CFD simulation to hydroplaning CFD simulation was first applied to analyze hydroplaning using a two-dimensional Reynolds equation including the fluid/structure interaction [38, 39]. Threedimensional CFD simulation has since been applied to analyze hydroplaning using the Navier\u2013Stokes equation including the free surface and a turbulence model [40, 41]. Three requirements in simulating the hydroplaning of a tire are the fluid/structure interaction in a three-dimensional model, the analysis of a rolling tire and the modeling of a practical tread pattern geometry. Although the fluid/structure interaction has been considered in the previous hydroplaning simulations [38\u201340], the other two requirements have not been considered. Seta and Nakajima et al. [42\u201344] inflation pressure (kPa) contact length (mm) contact width (mm) 210 250 170 210 250 130 98 100 91 160 103 130 103 100 103 R oa d re te nt io n ra te : K Velocity: V (km/h) 0 0.5 1.0 0 50 100 Fig. 12.39 Road retention rates for various inflation pressures and speeds (reproduced from Ref. [19] with the permission of Guranpuri-Shuppan) first established a new numerical procedure for hydroplaning in which the three requirements are considered and applied the procedure in the design of a tire pattern. They used the commercial explicit code MSC.Dytran whereby the tire was analyzed through FEA with a Lagrangian formulation and the fluid was analyzed using a FVM with an Eulerian formulation [45]. Furthermore, the interface between the tire and fluid was modeled by adopting general coupling. Because dynamic hydroplaning that occurs in region A of Fig. 12.26 is mainly studied, they ignored the viscosity of water. Using the same procedure proposed by Seta and Nakajima et al., the patent of which was granted for Bridgestone in 2002 [46], studies were conducted on the critical hydroplaning speed [47, 48], the effect of the tire pattern on hydroplaning [49] and the braking distance on a wet road [50, 51]. Osawa and Nakajima of Bridgestone [52] developed the riblet wall technology that suppresses the turbulence inside the main groove by CFD and validated to improve the hydroplaning of a tire. (2) Validation of hydroplaning simulation The hydroplaning simulation of a rolling tire with a practical tread pattern is shown in Fig. 12.40, where water drains into the two circumferential grooves and lateral grooves. The velocity of the tire is 60 km/h, the tire size is 205/55R16, the inflation pressure is 220 kPa, and the load is 4.5 kN. Prescribed velocities are applied to the tire model in the horizontal and rotational directions. The coefficient of friction between the tire and road is zero. The depth of water is 10 mm, and flow boundaries around the water pool are wall boundaries through which no water can flow. An explicit FEA/FVM scheme is applied in this simulation, and water flow is evaluated after the hydrodynamic force becomes stable. Figure 12.41 compares the prediction and measurement of water flow. The photograph taken through a glass plate shows the water flow of a rolling tire. The water depth is 10 mm, and the velocity is 60 km/h in the prediction and experiment. Although the viscosity of water is ignored in this simulation, the prediction of water flow is in good qualitative agreement with the measurement. (3) Global\u2013local analysis for hydroplaning simulation The hydroplaning analysis of a rolling tire requires much computational time, particularly for a tire with a practical pattern for the whole tread. A new procedure based on global\u2013local analysis was therefore developed to predict the water flow around a practical tread pattern [42]. In the scheme of global\u2013local analysis as shown in Fig. 12.42, a rolling tire with a blank tread is first analyzed in global analysis considering the fluid/structure interaction. The history of the displacement of the belt is then obtained in the global analysis. After the practical pattern model is glued to the belt model, the prescribed velocities calculated from the displacement of the global analysis are applied to the belt in local analysis. The local analysis is run as a separate analysis. Because the local model is more precise than the global model, the nodal coordinates in the local model are not necessarily the same as those in the global model. The prescribed velocities in the local model are thus determined using element interpolation functions. The local analysis also considers the fluid/structure interaction. Because the local model is independent of the global model, the effect of a small change in the tread pattern design on hydroplaning can be analyzed by changing only the local model. The computational time can thus be reduced in this global\u2013local analysis. (4) Development of a new pattern design through hydroplaning simulation Global\u2013local hydroplaning simulation is conducted to develop new pattern technology. To control water flow around the tread pattern, the three-dimensional design of the block shape is studied. The water flow around the block tip can be made smooth by a sloped block tip, which is shown in Fig. 12.43. Typical dimensions of the block tip are given in the figure, and the depth of the tread pattern is 8 mm. Two rolling tires are simulated with and without the sloped block to study the effect on hydroplaning. The water flow of the tire with the sloped block becomes smoother than that of the tire without the sloped block as shown in Fig. 12.44, indicating that the sloped block avoids an increase in the hydrodynamic pressure around the block tip. The measurement of the tire hydroplaning velocity shows a 1-km/h improvement due to the sloped block, demonstrating that the sloped block is effective in improving the hydroplaning performance. A new wet pattern for a Formula One tire is designed through hydroplaning simulation. Hydroplaning readily occurs in motorsport owing to the high speeds of racing, but tread patterns have been designed through trial and error according to the designer\u2019s experience. The procedure of pattern design begins with the simulation of a tire with a blank tread. The predicted streamline of a front tire is asymmetric owing to the camber angle, while the predicted streamline of a rear tire is symmetric as shown in Fig. 12.45. The velocity is 200 km/h, and the water depth is 2 mm in this calculation. The groove geometry on the tread is then designed according to the predicted streamline. Because the streamline indicates where water tends to flow out of the contact patch, water can be effectively drained by grooving voids having the same geometry as the streamline on the blank tread as shown in Figs. 12.45 and 12.46. The alignment of grooves of front and rear tires gradually changes from the circumferential direction in the center area to the lateral direction in the shoulder area. The groove configuration becomes asymmetric for the front tire because the centerline of the contact area is not located at the center of the tire geometry owing to the camber angle. The hydrodynamic pressure of the control pattern and newly designed pattern is shown in Fig. 12.47. The numbers in parentheses are hydrodynamic force indexes, which show that the hydroplaning performance is better for the newly designed pattern. The lap time of a tire with the new pattern was shorter in a field test. Finally, the groove geometry on the block of the new wet pattern for a Formula One tire is designed. The water drainage on the block can be further improved by adding a small void called a sipe. Global\u2013local analysis is conducted to predict the water drainage on the block as shown in Fig. 12.48. The hydrodynamic pressure is largely reduced by adding the sipe, and the lap time in a field test was shortened by adding sipes. (5) Riblet wall to improve hydroplaning Osawa and Nakajima [52] developed the riblet wall technology inside main grooves that makes the near-wall turbulence structure smooth as shown in Fig. 12.49. Using the CFD, the vorticity on a flat surface is found to be larger than that on a surface with riblets (the riblet wall). Because the turbulent flow contacts with the vertices of riblet on the riblet wall, the vorticity is suppressed on the riblet wall. The riblet wall technology was applied to main grooves of a passenger tire and validated to improve the hydroplaning performance in the proving ground. New pattern (65)Control pattern (100)" ] }, { "image_filename": "designv11_62_0001509_1350650111404114-Figure3-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0001509_1350650111404114-Figure3-1.png", "caption": "Fig. 3 Model of gas journal bearing system", "texts": [ " To keep the stability of the calculation scheme, the following condition should be considered: the calculation nodes between arbitrary two orifices should not be less than two orifices; the step time should be less than 0.01. During the solving procedure, the over-relaxation method is employed to form and update the triangular matrix and the discrete equations are solved. By integrating the transient gas film pressure P\u00f0 , , \u00de acting on the rotor, the transient gas film forces can be obtained as follows Fx \u00bc PaR2 Z2 0 ZL=R 0 P\u00f0 , , \u00de cos d d \u00f09\u00de Fy \u00bc PaR2 Z2 0 ZL=R 0 P\u00f0 , , \u00de sin d d \u00f010\u00de The model of journal gas-bearing system is presented in Fig. 3. The rotor system is formed by discrete disc, elastic shaft segment with distributed mass. The numbering scheme is shown in Fig. 3, where the journal gas bearing locates on node 2, 6 and the unbalance mass locates on node 4. In the simulation procedure, it is the same disc that is mounted once without static eccentricity and the other time with unbalance. The finite element method is employed to form the stiffness and mass matrixes. The Timoshenko beam element is used. Then, the equation of motion is given as follows M total \u20acuf g \u00fe C total _uf g \u00fe K total uf g \u00bc F totalf g \u00f011\u00de where uf g is the displacement and angular vector of each node for the rotor system, given as x1, y1, y1, x1, x2, y2, y2 x2 xN , yN , yN , xN " ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0002606_jfm.2017.535-Figure1-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0002606_jfm.2017.535-Figure1-1.png", "caption": "FIGURE 1. On stability of a cylindrical blob: (a) problem set-up, (b,c) qualitative velocity fields of the base state without (b) and with (c) the lateral acceleration g. The vector lengths show the relative velocity magnitudes.", "texts": [ " The problem of breakup of a liquid jet is often studied in the incompressible inviscid potential approximation, starting with the work of Rayleigh (1878) and continuing with more recent studies (Agbaglah et al. 2013), which allows one to explain the breakup phenomena robustly. Adopting this formulation, the governing equations for a fluid of density \u03c1 and surface tension \u03c3 reduce to the Laplace equation for the velocity potential \u03c6, the Cauchy\u2013Lagrange integral for the pressure p and the kinematic boundary condition at the cylinder interface r = f (t, z, \u03d1) (cf. figure 1a): 1\u03c6 = 0, for r 6 f (t, z, \u03d1), (2.1a) \u2202\u03c6 \u2202t + 1 2 |\u2207\u03c6|2 =\u2212 1 \u03c1 p+ gy+C(t), for r 6 f (t, z, \u03d1), (2.1b) \u2202f \u2202t = \u2202\u03c6 \u2202r \u2212 1 r2 \u2202\u03c6 \u2202\u03d1 \u2202f \u2202\u03d1 \u2212 \u2202\u03c6 \u2202z \u2202f \u2202z , at r= f (t, z, \u03d1). (2.1c) Here, C(t) is an arbitrary function of time which can be added to \u03c6 without changing the velocity field v = \u2207\u03c6. The cylinder interface S = r \u2212 f (t, z, \u03d1) is characterized by the unit normal vector n = \u2207S/|\u2207S| and the interfacial curvature, which simplifies to \u2207 \u00b7 n ' r\u22121 \u2212 fzz \u2212 r\u22122f\u03d1\u03d1 for small departures from a circular cylinder shape", "o rg /c or e/ te rm s. that this is possible only if there is no flow in the z-direction and g= 0. Since we are going to consider the case Bo= \u03c1gF2/\u03c3 1, then deviation from this base state due to acceleration can be treated as a perturbation (\u00a7 4.1). The Laplace equation (2.1a) admits a non-trivial solution of the above type only when there is a line source at the axis of symmetry providing a mass flux Q, treated here as constant, \u03a6(t, r)= (Q/2\u03c0) ln r, (2.2) producing an axisymmetric flow field; cf. figure 1(b). The logarithmic singularity can be removed by considering a mass source on the cylinder axis of a small radius \u03b5 > 0. The time-dependent cylinder radius F(t) is found from the kinematic boundary condition (2.1c), yielding F(t) = \u221a F2 0 +Qt/\u03c0, the same as the rim growth on a retracting soap film in the Taylor\u2013Culick theory (Taylor 1959; Culick 1960). We will focus, without loss of generality, on the asymptotic case, F(t)' \u221a Qt/\u03c0, valid for t \u03c0F2 0/Q\u2261 tc. (2.3) For a shrinking cylinder, i.e. when t\u2032= tc\u2212 t tc, F(t)' \u221a Qt\u2032/\u03c0, and since \u2202t\u2192\u2212\u2202t\u2032 , the stability results would be the exact opposite of the case considered below", " If one naively applies the stability results for a time-independent cylinder to a time-dependent one, it seems to be intuitive to hypothesize that the most unstable wavenumber should still scale as k \u223c F\u22121, now being time-dependent. Indeed, after analysing equations (3.1) with \u03ba = const., we conclude that for large time, t\u0303\u2192\u221e, the solution behaves as f\u0302 \u20320 \u223c t\u0303\u2212a\u0302(\u03ba)/4\u2212(1/8)e4 \u221a c t\u0303 1/4 , c= b\u03022(\u03ba)(1\u2212 \u03ba2)(t\u2217/t0) 3/2. (3.2a,b) For truly t\u0303 \u2192 +\u221e, the growth is dominated by the exponential only, so that the maximum growth rate is achieved when c(\u03ba) is at maximum, which occurs at some finite \u03ba = O(1) similar to the classical dispersion relation (3.5); cf. also figure 1.5 of Drazin & Reid (2004). For finite times, the optimum \u03ba is affected by a(\u03ba) as well. The problem with the above hypothesis k\u223cF\u22121 is that it implicitly requires the Fourier transform with time-dependent wavenumber to commute with the operation of the time derivative, whereas they clearly do not. Hence, it can be surmised that the hypothesis is not viable. In fact, by solving an initial value problem for (3.1a), 827 R3-5 ht tp s: // do i.o rg /1 0. 10 17 /jf m .2 01 7. 53 5 D ow nl oa de d fr om h tt ps :// w w w ", " 827 R3-11 ht tp s: // do i.o rg /1 0. 10 17 /jf m .2 01 7. 53 5 D ow nl oa de d fr om h tt ps :// w w w .c am br id ge .o rg /c or e. C ol um bi a U ni ve rs ity L ib ra ri es , o n 22 A ug 2 01 7 at 1 3: 41 :4 2, s ub je ct to th e Ca m br id ge C or e te rm s of u se , a va ila bl e at h tt ps :// w w w .c am br id ge .o rg /c or e/ te rm s. 4.1. Base state correction From (2.8), it is obvious that non-zero acceleration g affects mode n = 1 only and thus induces a non-axisymmetric flow field in the blob such as in figure 1(c), Lf\u0302 \u20321 = gF/2, (4.1) which after the transformation u1 = f\u0302 \u20321 exp( \u222b a dt/2) reduces to u\u2032\u20321 \u2212 s2(t)u1 =Ct1/2gF(t)\u2261 h(t), where s2(t)= t\u22122 [4\u22121 \u2212 (t/t2) 1/2k2F2 ]. (4.2) Looking for a long-time correction, we can simplify s2(t) to s2(t)=\u2212t\u22121/2t\u22121/2 2 t\u22121 < 0, provided that t 42/3t2/3t1/3 2 . The solution for the homogeneous part of (4.2) is u1(t)= \u221a tZ2/3 ( 4 3 t\u22121/2 1 t\u22121/4 2 t3/4 ) , (4.3) where Z\u03bd(z)= C1J\u03bd(z)+ C2J\u2212\u03bd(z), z= \u03b3 t3/4, \u03b3 = (4/3)t\u22121/2 1 t\u22121/4 2 and \u03bd = 2/3. If u1 = C1v1+C2v2, with v1= t1/2J\u03bd(z) and v2= t1/2J\u2212\u03bd(z), then the Wronskian is w= v1v \u2032 2\u2212 v2v \u2032 1 and the solution to the inhomogeneous problem (4" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0002799_icinfa.2017.8079003-Figure5-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0002799_icinfa.2017.8079003-Figure5-1.png", "caption": "Fig. 5 The following of circle path. After Fig.8 in [12]", "texts": [ " 3) The Control Laws of Pitch Channel: 0 ( ) ( ) ( ) t r V r V r V r d kp V V ki V V dt kd V V dt \u03b8 = \u2212 \u2212 \u2212 \u2212 \u2212 \u2212 (17) Where rV is the expected speed, and Vkp , Vki and Vkd is respectively the coefficients of the PID terms. In summary, a path following controller may be represented as 1 0 0 1 0 tan ( ) ( ) ( ) ( ) ( ) ( ) k ( ) cmd cmd t r V r V r V r r t r r p r d i r a g d kp V V ki V V dt kd V V dt d R k k dt dt\u03c8 \u03c8 \u03c8 \u03c6 \u03b8 \u03c8 \u03c8 \u03b7 \u03c8 \u03c8\u03c8 \u03c8 \u03c8 \u03c8 \u2212 = = \u2212 \u2212 \u2212 \u2212 \u2212 \u2212 = + \u2212 = \u2212 \u2212 + \u2212 (18) As shown in Fig. 5, 1\u03b7 represents the angle between L1 segment and the chord corresponding to arc, 2\u03b7 represents the angle between the velocity vector and the tangent at the circle radius determined by the current position. 3\u03b7 represents the angle between the tangent and chord. \u03b7 represents the angle between V and L1 segment. \u03b1 represents the angle between the circle radius determined by the current position and the horizontal line. \u03b8 represents the angle between V and the horizontal line. A derivation of circle path presented in [12] is adapted to Fig 5. It is assumed that 1\u03b7 and 2\u03b7 are small initially. Then 1 0\u03b7 \u2248 , 2 0\u03b7 \u2248 (19) According to the previous derivation, there is also a relational expression 1 3sin 2 L R \u03b7 \u2248 (20) So 2 1 3cos 1 2 L R \u03b7 \u2248 \u2212 (21) Let 3cosm \u03b7= (22) According to the small angle hypothesis, we know { } { } 2 2 2 1 2 3 1 2 3 1 2 3 1 1 1 2 1 3 2 3 3 1 2 sin 2 sin( ) 2 sin( )cos cos( )sin 2 cos cos sin V V V L L L V L \u03b7 \u03b7 \u03b7 \u03b7 \u03b7 \u03b7 \u03b7 \u03b7 \u03b7 \u03b7 \u03b7 \u03b7 \u03b7 \u03b7 \u03b7 = + + = + + + \u2248 + + (23) And on the basis of the geometric relationship in Fig. 5 1 3 1 cos d L \u03b7 \u03b7\u2248 (24) 2 d V \u03b7 \u2248 (25) Substituting (24) and (25) into (23) results in 2 2 22 1 1 1 2 sin 2 2 V V m V m V d d L L L R \u03b7 = + + (26) From Fig. 5 we can obtain lateral offset velocity 2 2sind V V\u03b7 \u03b7= \u2248 (27) According to the angle relationship: 22 \u03c0\u03b8 \u03b1 \u03b7\u2212 = \u2212 (28) Therefore 2( )cmda V V V V d\u03c9 \u03b8 \u03b1 \u03b7 \u03b1= = = \u2212 = \u2212 (29) So 2 cmd V a V d d R \u03b1= \u2212 = \u2212 (30) Combining (26) and (30), it is concluded that 22 0n nd d d\u03c2\u03c9 \u03c9+ + = (31) With 0.707\u03c2 = , 1 2 n V m L \u03c9 = (32) Because 0 1\u03c2< < , the system is the under damped second order state and its time response has oscillatory characteristics. When 0d \u2192 , t \u2192 \u221e . That is, the lateral deviation error can converge to zero" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0001620_978-1-4614-3475-7_2-Figure2.26-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0001620_978-1-4614-3475-7_2-Figure2.26-1.png", "caption": "Fig. 2.26 Example 2.3", "texts": [ "8) xc \u222b \u03b1 \u2212\u03b1 Rd\u03b8 = \u222b \u03b1 \u2212\u03b1 xRd\u03b8 . (2.56) Because x = Rcos(\u03b8 ) and with (2.56), one can write xc \u222b \u03b1 \u2212\u03b1 Rd\u03b8 = \u222b \u03b1 \u2212\u03b1 R2 cos(\u03b8 )d\u03b8 . (2.57) From (5.19), after integration, it results 2xc\u03b1R = 2R2 sin (\u03b1) , or xc = Rsin(\u03b1) \u03b1 . For a semicircular arc when \u03b1 = \u03c0 2 , the position of the centroid is xc = 2R \u03c0 , and for the quarter-circular \u03b1 = \u03c0 4 , xc = 2 \u221a 2 \u03c0 R. Example 2.3. Find the position of the mass center for the area of a circular sector. The center angle is 2\u03b1 radians, and the radius is R as shown in Fig. 2.26. Solution The origin O is the vertex of the circular sector. The x-axis is chosen as the axis of symmetry and yC =0. Let MNPQ be a surface differential element with the area dA = \u03c1d\u03c1d\u03b8 . The mass center of the surface differential element has the abscissa x = ( \u03c1 + d\u03c1 2 ) cos ( \u03b8 + d\u03b8 2 ) \u2248 \u03c1 cos(\u03b8 ). (2.58) Using the first moment of area formula with respect to Oy, (2.15), the mass center abscissa xC is calculated as xC \u222b \u222b \u03c1d\u03c1d\u03b8 = \u222b \u222b x\u03c1d\u03c1d\u03b8 . (2.59) Equations (2.58) and (2.59) give xC \u222b \u222b \u03c1d\u03c1d\u03b8 = \u222b \u222b \u03c12 cos(\u03b8 )d\u03c1d\u03b8 , or xC \u222b R 0 \u03c1d\u03c1 \u222b \u03b1 \u2212\u03b1 d\u03b8 = \u222b R 0 \u03c12d\u03c1 \u222b \u03b1 \u2212\u03b1 cos(\u03b8 )d\u03b8 " ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0001529_msf.701.77-Figure1-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0001529_msf.701.77-Figure1-1.png", "caption": "Figure 1 \u2013 3D model of the bench mark", "texts": [ " No part of contents of this paper may be reproduced or transmitted in any form or by any means without the written permission of Trans Tech Publications, www.ttp.net. (ID: 130.113.86.233, McMaster University, Hamilton, Canada-02/04/15,18:13:42) The first step in this study was to select the shape of the product [7].The bench mark selected was the spanner \u2013 a representative of the hand tool industry. Hand tool industry is a vibrant and developing industry which can derive benefit from the RP techniques. CAD model (Refer figure 1) of the bench mark was made on UNIGRAPHICS software. The 3D CAD model was converted into the STL format which was fed into the computer attached to the FDM machine. The machine was cleaned and the benchmark was set in various orientations. After setting of orientations the component was sliced, layer by layer construction was done and thereafter the components were cleaned. Build orientation is important for several reasons. First, properties of rapid prototypes vary from one coordinate direction to another" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0000608_j.triboint.2010.11.012-Figure2-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0000608_j.triboint.2010.11.012-Figure2-1.png", "caption": "Fig. 2. Configuration of an aerostatic bearing with a double-array of feeding holes: (a) sectional side view and (b) cross-section view.", "texts": [ " This study investigates the rotor\u2013aerostatic bearing system with a double-array of feeding holes, to determine the static and dynamic characteristics of such an aerostatic bearing and the stability thresholds of a rotorbearing system. Both the critical inertial force and the critical whirl ratio at various rotor speeds, journal eccentricities and restriction parameters are determined and compared with those obtained by separate inherence and orifice restrictions. 2. Perturbation analysis For an aerostatic journal bearing, like that shown in Fig. 2, air supplied from an externally pressurized source passes through feeding holes with an orifice or an inherence compensation, or with double compensations by both restrictions in series. These holes are located in double-rows symmetrically on a plane and distributed uniformly around the circumference of the bearing. Based on the assumptions of compressible, isothermal laminar flow in bearing clearance and that air is a perfect gas, the non-dimensional Reynolds equation of this air film that is formed in the bearing clearance, which is derived from Navier\u2013Stokes and continuity equations can be expressed in two dimensional Cartesian coordinates as @ @y h 3 @P 2 @y \" # \u00fe D L 2 @ @z h 3 @P 2 @z \" # \u00bc 2L @ @y \u00f0Ph\u00de\u00fe4L @\u00f0Ph\u00de @t \u00f01\u00de where D and L are bearing diameter and length; P and h are the nondimensional pressure and the thickness of the film; y and z are the angular and axial coordinates of the bearing, respectively; t\u00bcot is the non-dimensional time, and L is the bearing number", " (19) and (20) must be solved iteratively by adding a(gc gc,old) to gc until 9gc gc,old9=gc,old converges to an allowance (gc,old andgc are the last and newest values in an iteration, and 0oar1 is an under-relaxation factor). The details of the solution scheme that is used in this study can be found in Yang [22]. 5. Results and discussion This study analyzes the static and dynamic load capacities and stability thresholds of a rigid rotor that rotates at L\u00bc5.0 and1.0, representing high and low speeds, respectively. As shown in Fig. 2, this rotor is supported by two identical aerostatic bearings, which are installed a double-array of feeding holes with a ratio of bearing length to diameter L/D of 1.0. Two arrangements of feeding holes; each array comprises six or three feeding holes that are arranged symmetrically around circumferential surface of the bearing. All the orifices that are installed in the feeding holes have the same values of the restriction parameter. When the number of feeding holes n\u00bc6, the holes are located equi-angularly at yr\u00bc01, 601, 1201, 1801, 2401 and 3001" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0001597_iros.2012.6386081-Figure7-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0001597_iros.2012.6386081-Figure7-1.png", "caption": "Fig. 7. Detailed drawing of wire structure of FRT-type index finger exoskeleton", "texts": [ " The right and left environment contact parts are fixed each other with a solid link that is a black-colored arched part in Fig.6. The environment contact part and the finger contact part are connected each other with wires at both right and left sides. An assistive force from the main exoskeleton acts on an intermediate point of the wire. The distribution factor of the assistive force is determined by a horizontal position of the intermediate point. It becomes 1:1.2(=cos\u03c62 : cos\u03c61) when an angle inside the exoskeleton, \u03c61, is 28\u25e6 and an angle outside the exoskeleton, \u03c62, is 42\u25e6 in Fig.7. The FRT-type thumb exoskeleton is shown in Fig.8. It also has three components: a main exoskeleton, a thumb contact part and an environment contact part. In the same way as the index finger exoskeleton, the right and left environment contact parts are fixed each other with a solid arched components which is silver color in Fig.9. Both a thumb of a wearer and an environment contact part of the exoskeleton touch the environment. A nonslip sponge is installed on the thumb contact part so that it prevents a wearer\u2019s thumb from moving at the thumb contact part when grasping an object" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0003140_tapenergy.2017.8397379-Figure1-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0003140_tapenergy.2017.8397379-Figure1-1.png", "caption": "Fig. 1. Free body diagram of 3-DOF Laboratory Helicopter", "texts": [ " Since here a new model is developed economically with the exact helicopter structure and movements. It is very economical for researchers to analyse the different advanced control system effectiveness. This paper presents a study on the indigenously developed laboratory helicopter prototype with three degrees of freedom. The developed model is based on the Quanser laboratory helicopter. The complete structure and mechanism of the Quanser model 3-DOF laboratory helicopter given in [10-12]. From Fig. 1, one front dc motor and one back dc motor are placed at the two extremes of the frame of helicopter to control the different motions. A long arm is used in the structure, and the frame is connected at one end and other end carries a counter weight. Long arm results in the 2 DOFs for the system to elevate and travel. Two propellers of the model are driven by DC motors mounted on the two ends of a rectangular frame. When the voltages on the each motor are equal, it moves in elevated direction. When unequal voltages are applied, it will pitch", " A limitation of the system is that there are three degrees of freedom but only two inputs (under actuated system). This makes it is impossible to design a controller capable of keeping all three angles at arbitrary reference values. The system has three outputs, pitch, elevation, and travel angles. In this work the control scheme designed to control the travel angle. The proposed controller objective is to track the reference for the travel channels.The free body diagram of the system is given in Fig. 1.The distance from front and back motor to centre of the system is Lx1. The distance from helicopter body centre to the centre shaft is represented as Ly1. The distance from helicopter centre shaft to the counter weight is Lz1. The three outputs of system are pitch, elevation, and travel angles. In this work the control scheme designed to stabilize the system by controlling the travel angle. The proposed attitude controller objective is to track the reference for the travel channel. The mathematical model of the system is obtained from [10]" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0000321_016918610x512631-Figure7-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0000321_016918610x512631-Figure7-1.png", "caption": "Figure 7. One-link manipulator using each approximate model for simulation.", "texts": [ " To evaluate these models, we simulated the one-link manipulators equipped with the approximate CEEs, which consisted of two wires, two actuators and two CEEs, using software available on the market (NASTRAN4D; MSC Software) as a dynamics simulator. The specification of the computer used was CPU: Pentium 4 (3.2 GHz), memory: 1.5 GB and OS: Windows XP. NASTRAN4D has a user-friendly graphical interface so that it is easy to create the mechanical structure and easy to analyze the dynamics motion. However, it have a disadvantage of requiring a long time for calculation. Figure 7 presents an overview of the simulation systems. In the simulation, the same bias wire tension was loaded to the two wires first, then the joint stiffness K was estimated numerically by measuring the minute dis- D ow nl oa de d by [ U ni ve rs ity o f C on ne ct ic ut ] at 1 1: 45 1 1 Ja nu ar y 20 15 1646 H. Kino, D. Nakiri / Advanced Robotics 24 (2010) 1639\u20131660 Figure 8. Experimental setup to measure joint stiffness of a one-link manipulator equipped with two CEEs. Figure 9. Photograph of the one-link tendon-driven manipulator equipped with two CEEs" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0000662_detc2013-13305-Figure4-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0000662_detc2013-13305-Figure4-1.png", "caption": "FIGURE 4. Serial chain with nonholonomic rolling constraint", "texts": [], "surrounding_texts": [ "The system modeled in this test case is shown in figure (4). It consists of four bodies that are connected together and to the inertial frame by pin or revolute joints to form a serial chain topology. The axis of rotation of these joints is perpendicular to the plane of the paper. The mass and moments of inertia of each body is set to 1kg and 1kgm2 respectively. The reference frame used to model the system is shown in the figure and gravity of magnitude 9.81m/s2 acts along the negative y0 axis. The body 4 is a sphere of radius 0.5m and it is subject to the nonholonomic rolling constraint i.e it rolls on the inertial surface shown as the horizontal line while other bodies have a length of 1m. The initial condition is chosen such that all bodies are aligned and the first body makes an angle of 30 degrees with the ground. At initial instant, the bodies all have zero angular and translational velocity. The effective point of rolling is explicitly calculated from the geometry of the problem. A ten second simulation is conducted using the algorithm. The temporal results i.e. the joint angles, the relative speed and relative accelerations are compared against MSC ADAMS. Figure (5) shows the position, velocity and acceleration of the disk center in the horizontal direction while figure(6) shows the joint variables. Similarly, the constraint loads arising from the imposition of the nonholonomic constraint are also compared agains MSC ADAMS and shown in figure (7). As can be seen, there is excellent agreement between the results obtained from the two sources. Finally, the nonholonomic constraint violation is shown in figure (8) at the velocity and acceleration level respectively. As observed, the acceleration level constraint is satisfied to the 10\u221216 and velocity level at 10\u221210. This is without any explicit constraint stabilization or satisfaction procedure. The agreement of the results with those obtained from MSC ADAMS and excellent constraint satisfaction at both velocity and acceleration level demonstrate the verification of the algorithm for serial chains on inertial surfaces. 7 Copyright \u00a9 2013 by ASME Downloaded From: http://proceedings.asmedigitalcollection.asme.org/ on 01/31/2016 Terms of Use: http://www.asme.org/about-asme/terms-of-use" ] }, { "image_filename": "designv11_62_0001460_j.sna.2010.05.007-Figure2-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0001460_j.sna.2010.05.007-Figure2-1.png", "caption": "Fig. 2. Driving principle of Micro USM I.", "texts": [ " The input signal to Phase (for x-direction vibration) and Phase B (for y-direction vibration) re, VA = A sin(2 ft) VB = A sin(2 ft + ) (1) here, VA and VB are input signal to Phases A and B, A is the ampliude, f is the frequency (corresponding to the natural frequency of he first bending vibration mode), and is the phase between VA nd VB ( /2 or \u2212 /2 according to the rotational direction). When driving, the first bending vibration modes of the stator (cf. ig. 1) in the x-/y-direction are combined with a phase difference f /2, resulting in the head of stator vibrate as shown in Fig. 2. ccording to this vibration, the rotor being in contact with the staor head rotates along the z-axis by the frictional force between he stator and the rotor. In case the first bending vibration modes re combined with a phase difference of \u2212 /2, the rotational direcion is reversed. This ultrasonic motor could be categorized in the raveling wave ultrasonic motor, because a traveling wave with ne wavelength is generated in the circumferential direction of the tator head. 3. Concept As obvious from the driving principle of traveling wave ultrasonic motor shown in Section 2, the motor requires two AC signal inputs to drive" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0001433_pi-a.1962.0130-Figure21-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0001433_pi-a.1962.0130-Figure21-1.png", "caption": "Fig. 21.\u2014General arrangement of Lundell-type alternator.", "texts": [ " With either of these methods the cooling-fluid (air or liquid) inlet temperature could be low enough for the rotating-rectifier type of generator. In these circumstances the high-temperature capability of the solid-rotor generator would be of less importance than its inherent simplicity and robustness; these merits alone would then have to be weighed against its rather greater weight. (11.2) The LundeU-Type Generator The construction of this type of generator is generally similar to that described in Reference 11 and is illustrated in Fig. 21. The stator core and windings, which are of conventional construction, are mounted together with the stationary field coil inside a cylindrical yoke of magnetic material. The rotor con- sists essentially of two pole groups mounted on and magnetically insulated from the shaft. Each pole group comprises all the poles of one polarity extending into the stator bore from a continuous end-ring. Each end-ring runs inside a magnetic end-bracket, the small radial clearance between them constituting the auxiliary airgap. The magnetic flux path is shown by the dotted line in Fig. 21; it encircles the stationary circular field coils mounted inside the magnetic yoke. These two coils together develop the m.m.f. required to excite all the poles and require less power than the field winding of a conventional salient-pole machine. The m.m.f.s of the two field coils are not axially opposed and the bearings may be damaged by eddy currents if they are not magnetically isolated. An examination of Fig. 21 will show that the general arrangement of this machine is highly favourable for direct cooling either by blast air or by liquid. The rotating field structure will readily permit air to flow through it and scrub the inner bore surface of the rotor, and air passages of any required cross-section can be provided between the stator core and the yoke. By simple modification of the yoke, passages can be provided for liquid coolant to flow* in close proximity to all the sources of heat losses. The main design problem concerns the rotating field structure, in which adequate mechanical strength must be obtained without including magnetically superfluous material", " Experimentally, rotatingrectifier generators have operated satisfactorily with cooling air at an inlet temperature of 150\u00b0C, but more extensive trials would be required to prove them for regular flying service at this temperature. For variable-speed constant-frequency power generation, generators of several different types have been proposed in a variety of schemes. None of these has been so far developed as to give a clear guide to the most favourable generator construction. To Mr. Davies.\u2014With the single-Lundell construction shown in Fig. 21 there is theoretically no resultant end-thrust; any which might result from imperfections of manufacture is not likely to be greater than in any other type of generator, with the exception of axial-field machines. The effect of a.c. waveform on rectification phenomena does not appear to have been analysed mathematically. A general inspection of the problem seems to indicate that flat-topped voltage waves would give less ripple on no-load but rather more ripple on full-load than would sinusoidal waves, and that consequently there is little overall advantage in designing machines with special waveforms for brushless d" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0000139_1.3656894-Figure1-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0000139_1.3656894-Figure1-1.png", "caption": "Fig. 1 Diagram of journal bearing", "texts": [], "surrounding_texts": [ "The net shear force now follows directly by integration\nir^D2L Fi = 2c {.No - N,. A \\ 2c 6 ~ / ( 1 7 )\nStability of Small Oscillations The origin to which motion is referred is taken to be that where the journal and bearing axes are coincident. The system will be said to be asymptotically stable with respect to this origin if as time increases indefinitely the two axes return to coincidence after any small displacement. On the contrary, if after any small disturbance the distance between the two axes becomes unbounded with time the system will be called unstable. Other possible motions of bounded amplitude, such as steady oscillation or whirl about the origin or any position within the clearance circle of the bearing, could in a sense be considered stable. However, these motions, while possibly not leading to bearing failure, are sources of noise and for this reason the stability criteria given in this section are formulated to insure asymptotic stability about the origin.\nThe equations of motion are\n4>e + = j-FT M\nn =0 CO CO\n- FT = e \u00a3 C\u201e(A* - A\u201e*)\u00bb - I J 2 D \" ( A * - A*\u00b0>\" + e S A \u201e _ n n\n(23) n = 0 n = 0\nThe above form lends itself easily to a treatment of small oscillations of A* about the value Ao* because, as is customary with small oscillations, one may neglect higher powers of (A* \u2014 A0*).\nIn order to simplify the equations of motion somewhat it is helpful to change to new dependent variables defined by\ne = exp (Q0t + f Qdt)\n

now have the form\nQ + (Bo + 2Q0)Q + [Ai + B,Q\u00bb + 2fc2(A - Ao*)]

} + \" a'a3) { \u0302 } = \u00b0\n(30)\n(31)\n(32)\nFrom the definitions of Qo, Q, and ip it is apparent that the necessar}' and sufficient conditions for asymptotic stability are\nQo < 0\noi \u2014 03 > 0\ndid J \u2014 dl(l3 > 0\n(33)\n(34)\n(35)\n(18)\n(19)\nwhere the two forces FK and F.R contain the radial and transverse components of the pressure and viscous forces. Continuing with the conventions previously adopted these two forces are given bj'\nFR = Re{Fi(A*)e + F.( A*)E + F5} (20)\nFT = Im{F,(A*)e + F2(A*)e + F3} (21)\nAs it has been shown these forces depend on e, e, and in a com/ 2 A plicated fashion upon A* = A ( 1 \u2014 \u2014 ). A convenient form \\ N / for analytical solution is ^ cc aj\n- FR = - e X ) - Ao*) \" - I YJ BN(A* - A 0 * ) \" ( 22 )\nThe general procedure for numerical evaluation of the stability criteria for a given bearing is as follows: (a) Pick a value of A (proportion to the operating speed); (6) solve equations (28) and (29) for Qo and A0* (a trial-and-error approach seems necessary except for some special cases); (c) with a knowledge of Qo and Ao*, calculate the constants Oi, a?, as, and ai. It would, of course, be more desirable to be able to predict the stability characteristics directly from a knowledge of the bearing parameters. Fortunately, this daes become possible for a special case-\u2014low speed operation.\nIt is possible by eliminating Qo from equations (28) and (29) to show that as A approaches zero so also does A0* or, in other words, the whirl frequency approaches zero along with the operating speed. This provides for considerable simplification in the form of the fluid film forces. With the aid of equations (11) through (14) it follows that for small Ao*\n1 - 3 L L \u2014 sh2 \u2014 - \u2014 2 D P A\u201e*2\nBo =\nCo = J\nsh2 - 1 - D\n( c h 2 f + l )\n1 - I K + O J A\u201e*\n(36)\n(37)\n(38)\nDo =\nL L 3 s h 2 - - -\n(39)\nEquations (36) through (39) now permit an analytical solution of equations (28) and (29) for Qo and the whirl frequency\nN2 (40)\nD E C E M B E R 1 9 6 3 / 5 1 5\nDownloaded From: http://fluidsengineering.asmedigitalcollection.asme.org/ on 01/29/2016 Terms of Use: http://www.asme.org/about-asme/terms-of-use", "to = 1 + S - 2Q0\ni \u00b1 At* + G M Qo ^T (41)\nFurthermore, the stability criteria reduce to just one inequality for stable low-speed operation\n1 -\n3 L, L 2 s h 2 D - ~ L 7\n(Ch2|- + 1) sh2\n1 - D\ni K + ' ) J\nExample As an illustrative example let us consider a fixed bearing with a rotating journal which is either stable in the conical mode or else constrained such that conical motions are not allowed. The translational mode of oscillations will now be investigated for stability.\nThe first requirement is that the bearing be initially stable; that is, equation (42) must be fulfilled. Within the limits imposed by this inequality the bearing geometry and operating conditions may be adjusted to suit whatever the practical requirements may be. Let us say then, just by way of example, that the bearing shall be constructed as follows:\nc = 50 microin.\nD = 3 in.\nL = 9 in.\nM = 2 1b\nEquation (42) agrees with the experimental evidence which indicates that small length to diameter ratio bearings are particularly prone to instability and that with increasing L/D ratio a point of gradually diminishing stability is also reached. The general observation regarding small clearance ratios is seen to play a particularly important part too. Incidentally, this equation indicates why some investigators have considered all vertical bearings to be unstable; equation (42) is ordinarily not satisfied. It should be mentioned also, for reemphasis, that this low-speed stability criteria does not provide information as to what, if any, the ultimate stable operating speed may be. For this information and the characteristics of the small oscillations it is required to use the more general procedure and stability criteria given previously.\nL / 0 = I\nDYNAMIC BEARING NUMBER i f -A ( ' - (DIMENSIONLESS)\nTransverse pressure force characferislics arising from a displace-\nL / 0 ' 3\nDYNAMIC BEARING NUMBER A* \u00bb A (I - ^ MDIMENSIONLESS)\nFig. 2 Radial pressure force characteristics arising from a displacement e\n5 1 6 / D E C E M B E R 1 9 6 3\nTransactions of the A S M E\nFig. 4 Radial pressure force characteristics arising from a velocity\nDownloaded From: http://fluidsengineering.asmedigitalcollection.asme.org/ on 01/29/2016 Terms of Use: http://www.asme.org/about-asme/terms-of-use", ".1\n.08\n: .06 ,\no '{ ,04\n.02\nO~--~~~----~----__ - o 10 20 -,\nN 110 rod lsec\n30\nAs it has been aeen, the two quantities Qo and Ao\u00b7 completely determine the essent.ial characteristics of the motion. With a knowledge of the behavior of Qo and Ao\u00b7 t.he stability criteria given by equations (33), (34), (35) cnn be investigated for any operating speed N and, what is equivalent, the nature of the motion can be determined. To aid in the solution of equations (28), (29) for Qo and Ao\u00b7 Figs. 2, 3, 4, and 5 are recommended. These .simplify the computations somewhat when !\\ trial-and error approach is used. The results of these computations are\nshown in Figs. 6, 7, 8, and 9 where the behavior of Qo, (1 _ 2;0).\nal - aI, und a2 0 would imply an undamped or nominally unstable ships in the sway direction . Assumption 2. The ocean currents are assumed to be timevarying, irrotational, and bounded by V > 0 such that \u2016Vc\u2016 \u2264 V . Meanwhile, the terms of bu , bv , and br are lumped uncertainties satisfying |bi |i=u ,v ,r \u2264 bi , where bi is a constant . 3.2. Problem formulation As illustrated in Fig. 1, P = [x , y]T is the current location and Pd is denoted as Pd = [xd (\u03b8 ), yd (\u03b8 )]T in the desired path, meanwhile, defining a local reference frame: the path parallel reference frame {pp}. For the sake of simplicity, it is assumed that the USV is moved at the constant surge velocity ur = ud > 0 and supposes that Pd is sufficiently smooth. It is clear that the USV model is a nonlinear control system with two-inputs and four-outputs. The maneuvering control problem is concerned with designing the surge force and yaw moment to steer a USV to follow a geometric path without temporal constraint" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0000986_j.proeng.2011.03.137-Figure16-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0000986_j.proeng.2011.03.137-Figure16-1.png", "caption": "Figure 16 The scheme for the impossible independent movement Ty.", "texts": [ " We join an element EF on the point E (Figure 14). We present two situations: The length of the element EF is equal to the length of the element AB and the element EF is parallel to the segment AB (Figure 14). The length of the element EF is different from the length of the AB element (Figure15). For the situation shown in figure 14 we calculate the mobility of the mechanism. For that purpose, we cut twice the frame, so that the number of the joints becomes equal to the number of the elements, i.e. 6 (Figure 16). The mobility number of the element 6 is b1=3 (Ty, Tz, Rx), like in the previous example. Let\u2019s calculate the mobility number of the element 5. With this aim in view, we check the possible independent movements of the extreme element 5 relative to the frame 0 (Figure 16). The point E is situated on the elements 2 and 4, at the same time. For this reason, in the case of an infinitesimal displacement Ty of the element 5, from the point F1 to the point F2, the final point E3 must be at the intersection of the circular trajectory T1 (the circle with the dimension of the radius equal to the dimension of the element AB, and with the centre of curvature in the point F1) with the circular trajectory T2 (the circle with the dimension of the radius equal to the dimension of the element EF, and with the centre of the curvature in the point F2), that means between the point E1 and the point E2 (Figure 16, Figure 17). Tybeing very small, the point E3 must be very close to the point E1 (because we displace the point F from F1 to F2), but at the same time it must be very close to the point E, which is the initial position (Figure 17). But this is not possible and the independent movement Ty of the extreme element 5 is not possible, too. The same procedure is applied to the independent movement Tz, and the result is identical. But the movements Ty(Tz) and Rx are possible. That means the spatiality of the element 5 is: b=2" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0000744_isie.2013.6563755-Figure4-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0000744_isie.2013.6563755-Figure4-1.png", "caption": "Fig. 4. Backlash when bending", "texts": [ " This allows the inner wire to bend inside the outer tube when the wire is pushed. Since the inner wire consists of twisting several strands, the length of the inner wire is slightly shorter than that of it when the wire is pulled by applying a tension. Also, wire strands become more loosen and the wire becomes shorter when the wire is pushed. This phenomenon is illustrated in the Fig. 3. If a bending of the thrust wire is concerned, the inner wire will be in different positions inside the outer tube depending on the applied force. This is shown in Fig. 4. Above mentioned problems causes the backlash of the thrust wire. When the backlash is present, it affects the transmission of the position information from one end to the other. This greatly affects if the application considered is related to the haptics or the transmission of tactile information. Especially when the thrust wire movement direction changes the presence of the backlash causes a considerable position error. Therefore a compensation for the backlash is necessary. Even if the backlash is presented on a thrust wire, it does not affect much if the wire is used to a continuous action of either pulling or pushing of an object" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0003691_iemdc.2019.8785136-Figure1-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0003691_iemdc.2019.8785136-Figure1-1.png", "caption": "Fig. 1. Structure of PMSLM.", "texts": [ " Second, demagnetization degrees and remanence of five PM experiment samples were measured in stove at temperatures varying from room temperature to 300 \u00b0C to obtain the real data for next-step modeling. Third, ELM was introduced to map the nonlinear relationships between temperature and demagnetization degrees of PM and build the demagnetization model. Finally, comparison experiments between linear, polynomial, and ELM can certify the effectiveness and advancement of this proposed method. II. PERMANNET MAGNET CHARACTERISTICS PRESENTATION AND FINITE ELEMENT ANALYSIS The structure of PMSLM is shown in Fig. 1, this PMSLM is mainly consists of PMs, coils and Back-iron. The PMs used in PMSLM are the Ndfeb35 material, the remanence of this PM is 1230mT. The length, width and thickness of this PM are 40mm, 15mm, and 3mm, respectively. The FEA simulation results for normal PM (Br=1230mT) and demagnetization PM (Br=615mT) are shown in Fig. 2, this FEA simulation is by manual adjustment of the Br. But In the real working conditions, the PM often occurs demagnetization with temperature increasing, it is important to know the relationships between the temperature and remanence and build the accuracy model" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0002200_978-3-319-51222-8-Figure2.31-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0002200_978-3-319-51222-8-Figure2.31-1.png", "caption": "Fig. 2.31 Rotational springs coupled with beam ends", "texts": [ " The reciprocal of the displacement is kS . Piers basically act like springs (see Fig. 2.30). If the pier is soft, then a unit spread \u22121 traveled by the base of the pier is consumed to a large part by the pier and the beam hardly notices the spread at the base of the pier, which means that the influence function has a very low profile in the beam. Vice versa if the pier is quite stiff, then nearly the full signal will reach the beam, and consequently, the pier will carry a large part of the moving load. In rotational springs as in Fig. 2.31, the moment is coupled to the rotation \u03d5 of the spring via the rotational stiffness k\u03d5 M = k\u03d5 tan \u03d5 . (2.73) To determine the joined rotational stiffness k\u03a6 of the spring and the beam end(s), we let two moments X = \u00b11 that act on the two sides of the hinge determine the rotational stiffness of the structure (S) kS \u03d5 = 1 tan \u03d5l + tan \u03d5r , (2.74) and so the stiffness k\u03a6 is 112 2 Betti\u2019s Theorem 1 k\u03a6 = 1 k\u03d5 + 1 kS . (2.75) A side effect of the equation EI wIV = \u2212V \u2032(x) = p is that in a continuous beam, the distance between the zeros of the shear force is related to the magnitude of the support reaction (see Fig" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0001240_s12206-012-0612-3-Figure1-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0001240_s12206-012-0612-3-Figure1-1.png", "caption": "Fig. 1. Schematic figure of laser induction hybrid cladding process.", "texts": [ ": +86 27 87282120, Fax.: +86 27 87384670 E-mail address: ysf@mail.hzau.edu.cn \u2020 Recommended by Editor Sung-Lim Ko \u00a9 KSME & Springer 2012 used for single cladding experiments was mild steel cylinder with size of 400 mm (length) \u00d7 \u0424108 mm (diameter) \u00d7 5 mm (thickness). The chemical composition (wt.\uff05)of the powder material and substrate is 0.8C-17Cr-3.5B-4Si-5Fe-bal.Ni and 0.15C-0.20Si -0.47Mn-0.021S-0.017P-bal.Fe, respectively. The laser induction hybrid cladding process is illustrated schematically in Fig. 1. As shown, the screwy clad track was formed through the rotation and longitudinal direction line movement of the cylinder substrate, while the laser beam and induction coil were kept fixed. The off-axial powder nozzle was set at an angle of 45\u00b0 to the axis of laser beam, while the laser beam defocused on the surface of substrate perpendicularly. Argon was used as delivering and shield gas. The laser power PL was 4 KW and laser spot size D on the substrate was 5 mm. The frequency of induction f was selected as 30 KHz, and the induction specific power P0 was selected as 2 W/mm2 on the surface of the substrate" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0002601_gt2017-64123-Figure15-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0002601_gt2017-64123-Figure15-1.png", "caption": "Figure 15. The original (left) and optimized (right) impeller", "texts": [ " This thickening affects the blade and disc stiffness in this area without any significant increase in weight, which allows a reduction of the blade tension at the exit conjugation radiuses. Thus, during the optimization, 33 parameters vary: 4 parameters of the air-gas channel; 9 parameters of short blade; 11 parameters of the full-size blade; 9 disk parameters. The time of the optimization was 4 weeks. During this period, 300 configurations were analyzed. The original impeller models and the structure obtained as a result of the optimization are shown in Figure 15. Gas-dynamic and strength characteristics of the optimized impeller Figure 16 shows the velocity field distribution for the middle diameter of an optimized impeller. Figure 17 shows the equivalent von Mises stress distribution in the optimized structure, normalized relative to the maximum stress in the original design. The distribution of the long-term strength factors in the optimized full-length and short blades is shown in Figure 18. The fatigue life distribution of the optimized structure is shown in Figure 19" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0001590_6.2013-1503-Figure5-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0001590_6.2013-1503-Figure5-1.png", "caption": "Figure 5. Student Conceptual Design 3-View", "texts": [ " The IPPD methodology relies on a six step process to make a design decision: establish the need, define the problem, establish value, generate feasible alternatives, evaluate alternatives, and make a decision. It combines top-down design, systems engineering, and quality engineering into one common framework. 20 All analysis work feeds into the IPPD methodology which was used as the backbone of all projects. The final results consisted of three conventional up-scaled 777 based models, and one concept resembling an MD-11. The major analytical tools employed by the teams consisted of \u201clevel 0\u201d conventional empirically based parametric design equations and FLOPS. Figure 5 and Figure 6 show one of the student designs. Throughout the process, the students consistently demonstrated the ability to learn new material from virtual resources, and more importantly, collaborate with their geographically dispersed peers to effectively deliver the requirements of this phase. Overall, RFP 1 met three distinct learning objectives: 1) Students learned to conduct conceptual tradeoffs for a new or derivative commercial transport in a collaborative environment with dispersed team members" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0001258_icra.2013.6631391-Figure1-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0001258_icra.2013.6631391-Figure1-1.png", "caption": "Fig. 1. Under-actuated biped model with knees", "texts": [ " To avoid this unexpected falling, it would be necessary to mathematically understand how the walkers get into the unstable motion. About this problem, we make a hypothesis that limit cycle walkers often fall down because the rear leg does not leave the ground even if the fore leg lands for some reason. Throughout the gait analysis, we discuss the falling mechanism from the viewpoint of the emergence of DLS motion. This paper deals with the model of a planar biped model with active knees and semicircular feet shown in Fig. 1. \u2022 Let m1, m2 and mH be the masses of the lower leg frames, the thigh ones, and the hip. The inertia moment of each link can be neglected. \u2022 We call the stance leg Leg 1 and swing leg Leg 2. \u2022 Let (x, z) be the end position of Leg 1,and this is identical to the attachment position of semicircular foot. \u2022 Let u1 be the control torque on the knee joint of the stance leg, u2 be that on the hip joint, and u3 that on the knee joint of the swing leg. The clockwise direction is set to be the positive rotation", " 5 (b), we can see that \u03bbI4 tends to increase and become positive where m2 is sufficiently large. The value of m2 where \u03bbI4 reaches zero from negative becomes smaller as mH decreases. As in the previous result, it is thought to be causally related to speeding-up of the generated gait but the details are still unclear. More investigations are necessary. Here, we briefly report the analysis results of the effects of an upper body. Fig. 6 shows the model; we add an upper body to the underactuated bipedal model of Fig. 1 incorporating the joint torque, u4, between the upper body and the swing leg. We apply an output PD control to the upper body angle, \u03b85, for maintaining it at the desired one, \u03b85d [rad]. The walker can generate stable level gaits by choosing the parameters of the upper body and PD gains. The physical parameters except the upper body ones were chosen as listed in Table I. Fig. 7 shows the simulation results of level dynamic walking where mt = 10 [kg], Lt = 1.0 [m], T1 = 0.05, T2 = 0.10 [s] and \u03b2 = 0 [rad]" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0003031_icciautom.2017.8258701-Figure1-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0003031_icciautom.2017.8258701-Figure1-1.png", "caption": "Figure 1. workspace area enhancement of redundant mechanism relative to nonredundant mechanism when b 80mm=", "texts": [ " Also, Qu and Guo [13] proposed a novel planar kinematically redundant mechanism and the kinematics and workspace of the mechanism were studied, the statics simulation was performed and the stress and displacement distribution was obtained. In this work, a new 3- PRPR parallel mechanism is presented and a broad kinematic analysis of the mechanism is explained. The workspace of the mechanism is obtained and is compared with the same generation of non-redundant one. In the last part, the effect of base circle radius and moving plate\u2019s dimension on the workspace of the mechanism is studied. The structure of the suggested mechanism is depicted in Fig. 1. The mechanism has three distal prismatic actuators and three active joints move on a circular path with radius r. Each kinematic chain of the mechanism consists of an active curvilinear prismatic actuator, a passive revolute joint, a linear prismatic actuator and a passive revolute joint. As depicted in Fig. 1, two set of coordinate systems are used; R-frame as reference frame and M-frame as moving frame. The R-frame is attached to the geometric center of the moving plate where z-axis is perpendicular to the moving plat and y-axis passes through node 2B . Also, M-frame is attached to the center of circular path as shown in Fig. 1. Position vector of the geometric center of the moving plate is specified by [ ]Tx y=p and its orientation is denoted by \u03c6 . The angular position of the prismatic actuators on the curvilinear guide, is specified by i\u03b8 , with respect to R-frame, given as Ai i Ai= ,r R u (1) in which [ ]T Ai r 0 ,=u (2) i i i i i . cos sin = sin cos \u03b8 \u2212 \u03b8 \u03b8 \u03b8 R (3) The moving frame is considered to be an equilateral triangle with node pair distance, iPB equal to b . The position vector of node iB can be written in two different ways given as: 1) i iOA B chain: i iBi Ai= + ,r r s (4) where, i and is are the length of lateral prismatic actuators and the unit vector from node iA to iB , respectively" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0001369_icemi.2011.6037890-Figure6-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0001369_icemi.2011.6037890-Figure6-1.png", "caption": "Fig. 6 Compatible Matching of ADC Driver", "texts": [ " Corresponding capacitive matching shall be done at the link of high-speed ADC driver, which can reduce bandwidth loss substantially. High-speed ADC adopted in the system design is of differential input type, with an input impedance of INR =100 , a capacitive reactance of D_INC =0.08 pF, and an equivalent earth capacitance of individual signal input pin of G_INC =2.2 pF[8]. In addition, capacitive matching shall also be done for output impedance of operational amplifier of the driver as shown in C1 and C2 of Fig. 6 by taking into consideration of distributed capacitance generated by PCB wiring, etc. Magnitude of matching capacitance is basically equal to the equivalent input capacitance of ADC, value of which can be taken as 2 pF or so. Since there are too many aspects to be considered on signaling links, it is very hard to achieve an ideal status for overall system frequency response, with problems such as insufficiency of bandwidth allowance or uneven frequency response curve, etc. Therefore, frequency response calibration technique is adopted here, in which digital signal algorithm processing is performed on collected waveform data, thus relatively ideal frequency response is realized" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0000735_icca.2013.6564939-Figure2-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0000735_icca.2013.6564939-Figure2-1.png", "caption": "Fig. 2. The schematic drawing robotic manipulator with kinematics parameters.", "texts": [], "surrounding_texts": [ "The results obtained from the numerical simulation of proposed control scheme on a two-link RLFJ manipulator are shown in this section. The dynamic parameters of robot in equation (1) are given by: M(q)= [ (m1 +m2)l 2 1 +m2l 2 2 + 2m2l1l2 cos(q2) M12 M21 m2l 2 2 ] , g(q)= [ \u2212(m1 +m2)gl1 sin(q1)\u2212m2gl2 sin(q1 + q2) \u2212m2gl2 sin(q1 + q2)) ] , C= [ \u2212m2l1l2 sin(q2)(2q\u03071q\u03072 + q\u030722) \u2212m2l1l2 sin(q2)q\u03071q\u03072, ] , M12=M21 = m2l 2 2 +m2l1l2 cos(q2) Km = diag{10, 10} Num/rad,Mm = diag{0.5, 0.5} kg\u00b7m2 and Fm = diag{4, 0.5} Nm\u00b7sec/rad. Table I shows the kinematic parameters of the manipulator. For i = 1, 2, mi and li denote mass and length of the ith link. In the simulation, we assume that the end-effector is observed by a fixed camera whose delayed measurements are used directly to calculate the joint position. The delayed time of slow measurement is 0.05s. The process and measurement noise are chosen as Q = 0.01I and R = 0.1I . The initial position value of joints (joint1,joint2) for tracking control is given by (0.5rad, 0.5rad), and for fusion estimator is given by (0.6rad, 0.45rad). The initial velocity value for tracking control is given by (0, 0), and for estimation is (0,\u22120.25)rad/s. The parameters are chosen as: h = diag{2, 2}, \u03bb = diag{1.5, 1.5}, T = 0.001s, l = 5 and \u03b5 = diag{25, 25}. The estimate errors of position and velocity are plotted in Fig. 3. The position tracking using the proposed method is shown in Fig. 4 and Fig. 5 where the comparative result of SMC without fusion estimator is also plotted. In conclusion, the simulation results clearly indicate that the proposed 0 0.5 1 1.5 2 2.5 3 3.5 4 4.5 \u22120.3 \u22120.2 \u22120.1 0 0.1 0.2 0.3 0.4 time(s) e st im a ti o n e rr o r velocity estimation error of link1 position estimation error of link1 velocity estimation error of link2 position estimation error of link2 Fig. 4. The estimation errors for q1, q\u03071, q2, q\u03072 and \u03b8. 0 0.5 1 1.5 2 2.5 3 3.5 4 4.5 \u22121 \u22120.8 \u22120.6 \u22120.4 \u22120.2 0 0.2 0.4 0.6 0.8 1 time/s p o si ti o n t ra ck in g o f jo in t 1 /r a d reference trajectory proposed method SMC Fig. 5. Desired trajectories and real position of joints. approach guarantees the convergence of tracking errors and has the better tracking accuracy." ] }, { "image_filename": "designv11_62_0001689_978-94-007-4620-6_40-Figure1-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0001689_978-94-007-4620-6_40-Figure1-1.png", "caption": "Fig. 1 Planar RRRP linkage.", "texts": [ " The algorithm is being adapted for synthesis of spatial motion platforms. The motion of the coupler link in a four-bar planar mechanism is determined by the relative displacements of all links in the kinematic chain. The relative displacement of two rigid bodies in the plane can be considered as the displacement of a Cartesian reference coordinate frame E attached to one of the bodies with respect to a Cartesian reference coordinate frame \u03a3 attached to the other. Without loss of generality, \u03a3 may be considered fixed with E free to move, see Figure 1. The homogeneous coordinates of points represented in E are given by the ratios (x : y : z). Those of the same points represented in \u03a3 are given by the ratios (X : Y : Z). The mapping between the coordinates of points expressed in the two reference frames is given by the homogeneous coordinate transformation \u23a1 \u23a3 X Y Z \u23a4 \u23a6 = \u23a1 \u23a3 cos\u03b8 \u2212sin\u03b8 a sin\u03b8 cos\u03b8 b 0 0 1 \u23a4 \u23a6 \u23a1 \u23a3 x y z \u23a4 \u23a6 , (1) where (a,b) are the (X Z , Y Z ) Cartesian coordinates of the origin of E with respect to \u03a3 , and \u03b8 is the orientation of E relative to \u03a3 . Any point (x : y : z) in E can be mapped to (X : Y : Z) in \u03a3 using this transformation. For rigid body guidance, each pose is defined by the position and orientation of E with respect to \u03a3 , which is specified by the ordered triple (a,b,\u03b8). Dyads are connected through the coupler link at the coupler attachment points M1 and M2, see Figure 1. There is a specific type of constrained motion corresponding to each one of the four types of planar lower-pair dyad. The ungrounded R pair in an RR dyad is constrained to move on a circle with a fixed centre. Because of this they are denoted circular constraints. Linear constraints result when PR and RP dyads are employed because the R pair attachment point is constrained to move on a line defined by the P pair translation direction. The PP dyad represents a planar constraint: the line of one P pair direction is constrained to translate on the direction line of the other", " When K0 = 1 Equation (2) represents the implicit equation of points on a circle, and when K0 = 0 the equation becomes that of a line. Nonetheless, K0 is still an homogenizing parameter whose value is arbitrary. The Ki can be normalized by K0, but only when K0 is nonzero. Equations (2), (3), and (4) are used to integrate type and approximate dimensional synthesis of planar for-bar mechanisms for rigid-body guidance. Constructing the required synthesis matrix C based on the prescribed poses is done by relating the position of the two rigid body attachment points M1 and M2 in both reference frames E and \u03a3 , see Figure 1. Reference frames \u03a3 and E are correlated in two ways: 1. Points M1 and M2 move on circles or lines in \u03a3 ; 2. Points M1 and M2 have constant coordinates in E. Let (x,y) be the coordinates expressed in E of one of the coupler attachment points, M , and (X,Y ) be the coordinates of the same point expressed in \u03a3 . Carrying out the matrix multiplication in Equation (1) yields X = x cos\u03b8 \u2212y sin\u03b8 +az, Y = x sin\u03b8 +y cos\u03b8 +bz, Z = z. (5) Ignoring infinitely distant coupler attachment points, it is reasonable to set z = 1 in Equation (5) and substituting the result into Equation (2), with j \u2208 {1,2, " ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0003256_00207721.2019.1567864-Figure1-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0003256_00207721.2019.1567864-Figure1-1.png", "caption": "Figure 1. CE 150 laboratory helicopter.", "texts": [ " Then, we have \u23a1 \u23a2\u23a2\u23a2\u23a2\u23a2\u23a2\u23a2\u23a3 P\u03021ijkl(\u03bc(t, t + 1)) \u2217 \u2217 GT 2 P\u0302 2 ijkl(\u03bc(t, t + 1)) GT 2 P\u0302 3 ijkl(\u03bc(t, t + 1))G2 \u2217 A(\u03bc(t + 1))G1 + B(\u03bc(t + 1))F(\u03bc(t + 1))G1 A(\u03bc(t + 1)) G1 + GT 1 \u2212 P1 0 GT 2A(\u03bc(t + 1)) + GT 2L(\u03bc(t + 1))C(\u03bc(t + 1)) I \u2212 Q2 W1/2C(\u03bc(t + 1))G1 W1/2C(\u03bc(t + 1)) 0 R1/2F(\u03bc(t + 1))G1 0 0 \u2217 GT 1C T(\u03bc(t + 1))W 1 2 GT 1F T(\u03bc(t + 1))R1/2 \u2217 CT(\u03bc(t + 1))W1/2 \u2217 \u2217 \u2217 \u2217 GT 2 + G2 \u2212 Q3 \u2217 \u2217 0 I \u2217 0 0 I \u23a4 \u23a5\u23a5\u23a5\u23a5\u23a5\u23a5\u23a5\u23a6 > 0 (30) \u23a1 \u23a2\u23a2\u23a2\u23a2\u23a2\u23a2\u23a2\u23a3 P1 \u2217 \u2217 Q2 Q3 \u2217 A(\u03bc(t))G1 + B(\u03bc(t))F(\u03bc(t))G1 A(\u03bc(t)) G1 + GT 1 \u2212 P\u03021ijkl(\u03bc(t, t + 1)) 0 GT 2A(\u03bc(t)) + GT 2L(\u03bc(t))C(\u03bc(t)) I \u2212 GT 2 P\u0302 2 ijkl(\u03bc(t, t + 1)) W1/2C(\u03bc(t))G1 W1/2C(\u03bc(t)) 0 R1/2F(\u03bc(t))G1 0 0 \u2217 GT 1C T(\u03bc(t))W 1 2 GT 1F T(\u03bc(t))R1/2 \u2217 CT(\u03bc(t))W1/2 \u2217 \u2217 \u2217 \u2217 G2 + GT 2 \u2212 GT 2 P\u0302 3 ijkl(\u03bc(t, t + 1))G2 \u2217 \u2217 0 I \u2217 0 0 I \u23a4 \u23a5\u23a5\u23a5\u23a5\u23a5\u23a5\u23a5\u23a6 > 0 (31) By defining the following matrices Ni = FiG1, Mi = GT 2Li, Q2 ijkl = GT 2P 2 ijkl, Q3 ijkl = GT 2P 3 ijklG2 (32) and by substituting (22) and (3) in (30) and (31), kl ij and ij kl can be defined as (17). Considering the following properties r\u2211 i=1 r\u2211 j=1 r\u2211 k=1 r\u2211 l=1 \u03bci(t + 1)\u03bcj(t + 1)\u03bck(t)\u03bcl(t) kl ij = 1 2 r\u2211 i=1 r\u2211 j=1 r\u2211 k=1 r\u2211 l=1 \u03bcijkl( kl ij + kl ji ) > 0 (33) r\u2211 i=1 r\u2211 j=1 r\u2211 k=1 r\u2211 l=1 \u03bci(t + 1)\u03bcj(t + 1)\u03bck(t)\u03bcl(t) ij kl = 1 2 r\u2211 i=1 r\u2211 j=1 r\u2211 k=1 r\u2211 l=1 \u03bcijkl( ij kl + ij lk) > 0 (34) it is sufficient that the following inequalities hold in order to satisfy (33) and (34) : kl ij + kl ji > 0 (35) ij kl + ij lk > 0 (36) It is clear from (15) that the upper bound of cost function (9) is depended on the initial condition; in order to remove this dependency, we define matrix such that the following inequality holds:[ \u2217 I G + GT \u2212 P ] > 0 (37) where = 1 \u2217 2 3 ", " If there exists the positive definite matrices P1 and Q3 and matrices G1,G2,Ni,Mi and Q2 for every i, j \u2208 L such that ij + ji > 0 (44) where ij = \u23a1 \u23a2\u23a2\u23a2\u23a2\u23a2\u23a2\u23a3 P1 \u2217 \u2217 Q2 Q3 \u2217 AiG1 + BiNj Ai G1 + GT 1 \u2212 P1 0 GT 2Ai + MjCi I \u2212 Q2 W1/2CiG1 W1/2Ci 0 R1/2Nj 0 0 \u2217 GT 1C T i W 1/2 NT j R 1/2 \u2217 CT i W 1/2 \u2217 \u2217 \u2217 \u2217 GT 2 + G2 \u2212 Q3 \u2217 \u2217 0 I \u2217 0 0 I \u23a4 \u23a5\u23a5\u23a5\u23a5\u23a5\u23a5\u23a6 (45) then fuzzy system (7) is globally asymtotically stable and upper bound (43) can be minimised for cost function (9). The proof is presented in Appendix. The Humusoft CE150 twin-rotor helicopter, shown in Figure 1, is a laboratory scale helicopter which is selected to experimentally apply the controllers. The apparatus consists of two DC motors that drive the propellers. The main motor provides an ability for elevation angle movement in the vertical plane, and the side motor provides an ability for azimuth angle movement in the horizontal plane. Thus, the voltages to these two motors are the plant inputs, and the measured azimuth and elevation angles are the outputs of this multivariable dynamic plant. The plant is essentially nonlinear and unstable, and all inputs and outputs are coupled", " (A2) By substituting (18) with respect to (7) in (A1)[ Acl(\u03bc(t))x\u0303cl(t) ]TG\u2212TPG\u22121[Acl(\u03bc(t))x\u0303cl(t) ] \u2212 x\u0303Tcl(\u03bc(t))G\u2212TPG\u22121x\u0303cl(\u03bc(t)) + [ Ccl(t)x\u0303cl(\u03bc(t)) ]TW[ Ccl(t)x\u0303cl(\u03bc(t)) ] + [ Fcl(\u03bc(t))x\u0303cl(\u03bc(t)) ]TR[ Fcl(\u03bc(t))x\u0303cl(t) ] < 0 (A3) And by rewriting (A3) \u2217 = x\u0303Tcl(t)G \u2212T [ GTAT cl(\u03bc(t))G\u2212TPG\u22121Acl(\u03bc(t))G \u2212 P + GTCT cl(\u03bc(t))WCcl(\u03bc(t))G + GTFTcl(\u03bc(t))RFcl(\u03bc(t))G ] G\u22121x\u0303cl(t) < 0 (A4) In other words for all x \u2208 Rn, x(t) = 0, \u2217 < 0 should hold, which means GTAT cl(\u03bc(t))G\u2212TPG\u22121Acl(\u03bc(t))G \u2212 P + GTCT cl(\u03bc(t))WCcl(\u03bc(t))G + GTFTcl(\u03bc(t))RFcl(\u03bc(t))G < 0 (A5) (A5) can be converted to (A6) by Schur complement\u23a1 \u23a2\u23a2\u23a3 P \u2217 Acl(\u03bc(t))G GP\u22121GT W1/2Ccl(\u03bc(t))G 0 R1/2Fcl(\u03bc(t))G 0 GTCT cl(\u03bc(t))W1/2 GTFcl(\u03bc(t))R1/2 \u2217 \u2217 I \u2217 0 I \u23a4 \u23a5\u23a5\u23a6 > 0 (A6) Substitution of (19) and (8) in (A6) results in \u23a1 \u23a2\u23a2\u23a2\u23a2\u23a2\u23a2\u23a3 P1 \u2217 \u2217 P2 P3 \u2217 A(\u03bc(t))G1 + B(\u03bc(t))F(\u03bc(t))G1 A(\u03bc(t))G\u22121 2 G1 + GT 1 \u2212 P1 0 A(\u03bc(t))G\u22121 2 + L(\u03bc(t))C(\u03bc(t))G\u22121 2 G\u2212T 2 \u2212 P2 W1/2C(\u03bc(t))G1 W1/2C(\u03bc(t))G\u22121 2 0 R1/2F(\u03bc(t))G1 0 0 \u2217 GT 1C T(\u03bc(t))W1/2 GT 1F T(\u03bc(t))R1/2 \u2217 G\u2212T 2 CT(\u03bc(t))W1/2 \u2217 \u2217 \u2217 \u2217 G\u22121 2 + G\u2212T 2 \u2212 P3 \u2217 \u2217 0 I \u2217 0 0 I \u23a4 \u23a5\u23a5\u23a5\u23a5\u23a5\u23a5\u23a6 > 0 (A7) By defining T1 = diag{I,G2, I,G2, I, I} and Q2 = GT 2P 2,Q3 = GT 2P 3G2, multiplying before and after (A7) by TT 1 and T1. Then we have \u23a1 \u23a2\u23a2\u23a2\u23a2\u23a2\u23a2\u23a3 P1 \u2217 \u2217 Q2 Q3 \u2217 A(\u03bc(t))G1 + B(\u03bc(t))F(\u03bc(t))G1 A(\u03bc(t)) G1 + GT 1 \u2212 P1 0 GT 2A(\u03bc(t)) + GT 2L(\u03bc(t))C(\u03bc(t)) I \u2212 Q2 W1/2C(\u03bc(t))G1 W1/2C(\u03bc(t)) 0 R1/2F(\u03bc(t))G1 0 0 \u2217 GT 1C T(\u03bc(t))W1/2 GT 1F T(\u03bc(t))R1/2 \u2217 CT(\u03bc(t))W1/2 \u2217 \u2217 \u2217 \u2217 GT 2 + G2 \u2212 Q3 \u2217 \u2217 0 I \u2217 0 0 I \u23a4 \u23a5\u23a5\u23a5\u23a5\u23a5\u23a5\u23a6 > 0 (A8) We define Ni = FiG1 and Mi = GT 2Li then substitute in (A8) with respect to (4) ij = \u23a1 \u23a2\u23a2\u23a2\u23a2\u23a2\u23a2\u23a3 P1 \u2217 \u2217 Q2 Q3 \u2217 AiG1 + BiNj Ai G1 + GT 1 \u2212 P1 0 GT 2Ai + MjCi I \u2212 Q2 W1/2CiG1 W1/2Ci 0 R1/2Nj 0 0 \u2217 GT 1CiW1/2 NjR1/2 \u2217 CiW1/2 \u2217 \u2217 \u2217 \u2217 GT 2 + G2 \u2212 Q3 \u2217 \u2217 0 I \u2217 0 0 I \u23a4 \u23a5\u23a5\u23a5\u23a5\u23a5\u23a6 (A9) Finally we can rewrite (A9) as follows: r\u2211 i=1 r\u2211 j=1 \u03bci(t)\u03bcj(t) ij = 1 2 r\u2211 i=1 r\u2211 j=1 \u03bci(t)\u03bcj(t)( ij + ji) > 0 (A10) So Lyapunov stability condition V(x\u0303cl(t + 1)) \u2212 V(x\u0303cl(t)) < 0 holds if the following inequality holds ij + ji > 0 (A11) The upper bound of the cost function can be obtained same as the proof of the section main result for the non-monotonic Lyapunov function-based controller design which is given in the form of (43)" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0002311_ls.1385-Figure7-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0002311_ls.1385-Figure7-1.png", "caption": "FIGURE 7 Analysis diagram of shaft under micro disturbance", "texts": [ " When the rotor has a tiny disturbance along the negative direction of axis Y, and a displacement \u0394yoccurs, the bearing center O moves to O3; at this time, the oil film force in x and y directions are Fx3 and Fy3;When the rotor has a tiny disturbance along the positive direction of axis Y, the shaft center O' moves to O4; at this time, the oil film force in x and y directions are Fx4 and Fy4. According to the calculation formula of oil film dynamic stiffness, oil film stiffness can be expressed as Kxy \u00bc Fx3\u2212Fx4 2\u0394y ; (23) Kyy \u00bc Fy3\u2212Fy4 2\u0394y : (24) According to Figure 7, the rotor under external load balances in position of O'.Under the perturbations of (\u0394e,\u0394\u03b8), axis position moves to O1. At this time, the oil film forces along \u0394e direction and perpendicular to the \u0394e direction are Fe and F\u03b8, respectively: Fe \u00bc \u222b 1 2L \u2212 1 2L \u222b 2\u03c0 0 p cos \u03c6Rd \u03c6dz; (25) F\u03b8 \u00bc \u222b 1 2L \u2212 1 2L \u222b 2\u03c0 0 p sin \u03c6Rd \u03c6dz: (26) According to the oil film damping definition in Section 3.1, the oil film damping along \u0394e direction and perpendicular to \u0394e direction, the damping can be defined as cij \u00bc \u2202Fi \u2202 j \u2022 " ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0001661_tasc.2011.2109050-Figure3-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0001661_tasc.2011.2109050-Figure3-1.png", "caption": "Fig. 3. Magnetization method.", "texts": [ " 2, the rail is composed of two magnets facing to each other. Then, the magnetic poles face to each other. Yoke material is inserted in between the magnets to increase a magnetic flux density over the central part of the magnetic rail. The induction motor is composed of a primary coils and an induction plate with three phases to drive the levitated conveyor. The specifications of the linear induction motor are shown in Table II. The illustration of magnetically levitated conveyor with electromagnets for pulse-field magnetization is shown in Fig. 3. By installing bigger electromagnets over the levitated conveyor, 1051-8223/$26.00 \u00a9 2011 IEEE pulse-field magnetization is easily applied to superconductors in the levitated conveyor and larger pulse-field magnetization is easily obtained. Since the magnetic flux density distributions in superconductors are uniform along the magnetic rails, it is necessary for the electromagnets to produce magnetic flux densities similar to those from the magnetic rails. The specifications of electromagnets are shown in Table III" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0000806_6.2010-1206-Figure3-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0000806_6.2010-1206-Figure3-1.png", "caption": "Figure 3. CAD representation of the hybrid propulsion concept", "texts": [], "surrounding_texts": [ "IV. Full Scale design solution In order to compare what could be achived with alternative propulsion systems, the Lancair Legacy was chosen to be a baseline as a representative high performance sport airplane. The Legacy was also used for benchmarking the sizing program and validate the performance estimation and the weight estimation.\nThe work conceptual design work was reduced to 4 different versions to be studied. Any kind of trade off from the base line was permitted, the goal being to produce a configuration having the less compromise from the base line. The goal is to achieve performances as close as possible to the baseline. Some key parameters were set to be driving the design; those parameters were decided to be:\n Range\n Endurance\n Cruise speed\nOther parameters such as respecting FAR-23 rules were also imposed. Other similitude with the Lancair Legacy was up to each group to decide. In order to reduced the number of propulsion combination the following were chose to be studied:\nAmerican Institute of Aeronautics and Astronautics\n4", " Fuel cells\n Piston engine + batteries\n Fuel cells + solar cells\nA. Hybrid Propulsion The main challenge with an hybrid propulsion is to find a suitable way to connect the two\ndifferent engines and to study the best way to combine the both engine. The hybrid propulsion is directly inspired from the current development in the automotive industry.\nAmerican Institute of Aeronautics and Astronautics\n5", "B. Solar Panels and Batteries Propulsion Solar powered aircraft have been fascinating man kind for some decades[6-8], and several design have been presented from the first solar powered manned aircraft in 1979 by Solar Riser8, to the recent Solar Impulse project1, design for a round the world trip. The common issues for all those configurations have been the small available amount of power and the need for extreme low weight.\nAmerican Institute of Aeronautics and Astronautics\n6" ] }, { "image_filename": "designv11_62_0003282_s11041-019-00329-x-Figure1-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0003282_s11041-019-00329-x-Figure1-1.png", "caption": "Fig. 1. Scheme of non-vacuum electron-beam treatment of materials (the scanning range corresponds to the width of the treated plate): 1 ) electron beam; 2 ) electromagnetic deviating system; 3 ) treated billet; 4 ) treated layer.", "texts": [ " Prior to the tests, we milled the surface layer of the plates to remove the oxide film and degreased the surface with acetone. The electron-beam treatment was conducted at the Budker Institute of Nuclear Physics of the Siberian Branch of the Russian Academy of Sciences (Novosibirsk) with the help of an ELV-6 electron accelerator. The ELV-6 direct-action accelerator was equipped with a system emitting a concentrated electron beam with energy 1.4 MeV into air atmosphere. The scheme of treatment of a plate with a non-vacuum electron beam is presented in Fig. 1. We chose the modes for non-vacuum electron-beam treatment basing ourselves on the data presented in [9]. The treatment parameters are given in Table 1. The specimens for studying the structure and measuring the hardness were prepared by a method involving grinding and polishing of the surface. The final polishing was conducted using a suspension composed of 73 vol.% colloidal SiO 2 , 18 vol.% H 2 O 2 (40% solution) and 9 vol.% Kroll\u2019s reagent. The Kroll\u2019s reagent consisted of 92 vol.% H 2 O, 6 vol" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0003622_j.promfg.2019.06.133-Figure1-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0003622_j.promfg.2019.06.133-Figure1-1.png", "caption": "Fig. 1. Structured light scan of pre-welded blade (blue) and post-welded blade (grey). Not enough weld is deposited on convex side of the blade (red dashed oval).", "texts": [ " [8] established a morphing algorithm to non-rigidly register the ideal blade geometry to the deformed blade geometry so to accurately interpolate the tip geometry underneath the weld, which accounts for angular distortion, lean, and wear of the parent blade and to mitigate inconsistency in fixturing. Morphological deformation of the serviceable blade is important in adaptive hybrid blade repair because each blade exhibits unique deformation after it is extracted from service. Further, the accuracy of registration determines the final repaired product. In the additive process, the registered 3D tip geometry provides a template to guide the additive toolpath to not only increase material efficiency, but also prevent the problem of having insufficient weld to carry out the repair, as in Fig. 1. Once the tip is welded, machining is used to blend the weld to the parent material, with the ideal result being kept within the specified limit suited to customer requirement [9].. While this method was successful in adaptive repair of the blade geometries, it did not consider the required digitization of the blade model as well as optimization of probing and geometric reconstruction strategies to guide the repair process. The present study establishes an optimized probing strategy to implement in a HM-based repair process so to generate crosssection profiles for reconstructing the 3D blade geometry" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0002581_icuas.2017.7991368-Figure1-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0002581_icuas.2017.7991368-Figure1-1.png", "caption": "Fig. 1. Frameworks", "texts": [ " In the case of a control problem with constraints, it is necessary to solve a quadratic programming problem at each sample step to get the optimal control sequence. There is not an analytical expression for the control policy. The constraints can define limits on the control signals, the desired trajectory, the maximal overshoot, among other characteristics of the response of the system. In this case, it is a problem of quadratic optimization of the cost function (2) subject to constraints of the form Rru \u2264 c (5) To describe the dynamic model of the quadcopter it is necessary to define two frames of reference (see Figure 1), the body reference frame (xb, yb, zb), and the inertial reference frame (xe, ye, ze). The rotational quadcopter dynamics is given by J\u2126\u0307 + \u2126\u00d7 J\u2126 = Me (6) R\u0307 = R\u2126\u0302 (7) where \u2126 = [ p q r ] is the vector of angular velocity around the axes xb, yb and zb. The map \u2227 \u00b7 : R3 \u2192 so (3) is defined as x\u0302 y = x\u00d7 y. R is the rotation matrix from the inertial reference frame (xe, ye, ze) to the body reference frame (xb, yb, zb). The translational dynamics, in the inertial frame of reference, is described by the following equation X\u0308 = ge3 \u2212 1 ma TTRe3 (8) with X = [ x y z ]T the translational position vector, e3 the unit vector in the direction of the zb axis, ma the mass of the quadcopter, g the gravity constant, TT the total thrust given by TT = 4 \u2211 i=1 Ti (9) The control objective is stated as follows", " For the generation of the model that is used for the generation of constrained references, the closed-loop system shown in the Figure (3) is used, which is already controlled by the unrestricted GPC. The input of this model is defined as Ua = [ uax uay uaz ]T and the output as Yc = [ xc yc zc ]T . The Figure (4) shows the step responses of the translational positions x, y, and z, obtained experimentally. There is a change on the sign for the path along the z with respect to the coordinate frame shown in Figure 1 to display a graph with positive values. From the step responses, using the peak time tp and the maximum overshot Mp, three second-order linear systems model the closed loop dynamics are obtained. Table I shows the transfer functions for each second order system. The discretization of the transfer functions employed the zero-order hold method with a sampling period of 0.16 seconds. The path constraints along the axes x, y and z are included as described in Section I. The prediction of the output signals xc(t), yc(t) and zc(t) is obtained according to the standard GPC process by solving a Diophantine equation [7] with the discrete transfer functions shown in table II" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0000922_amr.586.259-Figure3-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0000922_amr.586.259-Figure3-1.png", "caption": "Fig. 3. Scheme of the combined process of asymmetric rolling and plastic bending", "texts": [ "224, University of Michigan Library, Media Union Library, Ann Arbor, USA-19/04/15,02:30:10) Fig.1 Scheme of deformation zone and its loads in asymmetric rolling where Owing to this approach it became possible to verify the distribution of normal contact strain on contact arches (Fig. 2). Fig.2 Search for normal contact stresses New technology has been developed to produce parts of bulky bodies of rotations with preassigned camber on the basis of combined processes of vertical asymmetric plate rolling and plastic bending (Fig. 3). The combined process is supposed to have three stages: 1) vertically asymmetric rolling when the front end of the plate does not contact the unbending roller; 2) not adjusted combined process of vertically asymmetric rolling and plastic bend (it starts when the front end of the plate contacts the unbending roller); 3) adjusted combined process of vertically asymmetric rolling and plastic bend (it starts when the front end of the plate comes off the unbending roller). Theoretical and experimental research showed serious drawbacks in the rolling technology caused by the growth of dynamic loads arising at the moment of plate contact with the bending roller (Fig" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0001600_tasc.2010.2100801-Figure2-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0001600_tasc.2010.2100801-Figure2-1.png", "caption": "Fig. 2. Coil shift along the X axis.", "texts": [ " Geometrical displacements and deformations can occur in fabrication of the coil due to imperfection of materials and manufacturing tolerances. An original technique [2], [3] allows evaluation of coil deformation. The technique is based on numerical reconstruction of possible geometrical distortions and misalignments of the coil using measured data on a spatial coil field. The technique is quite universal and can be applied to any coils. Standard distortions of the PF1 coil related to the accepted manufacture/assembly tolerances include [4]: 1) Linear shift along the X and Y axes, , (Fig. 2). 2) Tilt about the horizontal plane, ; at low tilt angles , the coil centerline is tilted with , where R is the coil radius (see Fig. 3). 1051-8223/$26.00 \u00a9 2011 IEEE 3) In-plane ellipticity, , , that is a deformation of the coil shape in the plane XY defined in terms of two elliptic half-axes , and an angle of rotation around Z-axis . Fig. 4 illustrates ellipticity at , , . 4) Vertical displacement of the coil centerline, or so called warping, occurred at winding, characterized by an amplitude and angle , where corresponds to the maximum deviation of the coil centerline relative to the horizontal plane XY; is the toroidal angle corresponding to this maximum deviation, " ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0000251_1743284713y.0000000367-Figure8-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0000251_1743284713y.0000000367-Figure8-1.png", "caption": "Fig. 8a and b, the laser remelting gave rise to a martensitic transformation, which is considered to be a hard phase in the microstructure.25 At the grain boundaries, the ultrafine carbide network or the eutectic was crossed. Superrefined carbides inside the crystal grains of these cellular structures were found, which were much finer and more uniform than those in the fatigue material. These results are consistent with the cases reported by others26,27 and can be explained by a very high cooling rate in the laser remelting process.", "texts": [ " Note that the constitutional phase in the fatigue substrate is pearlite with a mass of large size carbides, as compared with the original substrate. This can be attributed to the precipitation of alloy carbides as a function of thermal fatigue during the process of thermal cycling.23,24 The XRD patterns (Fig. 7b) indicate that the precipitate is in the form of M7C3 and M23C6 (M5Cr, Fe). In comparison, the very fine microstructure of the strengthening units mainly consists of martensite as shown in Fig. 8a. The XRD analysis better identified the phases formed in the microstructure 3 Sketch of laser surface remelting morphology and sam- ple dimensions 4 Sketch of thermal fatigue test machine Microhardness and microstructure after thermal fatigue tests Figure 9 gives the microhardness of samples as a function of thermal fatigue cycles. After remelting, the unit has an average microhardness of ,680 HV0?2, while the substrate hardness is ,270 HV0?2. The high hardness of the unit could be attributed to the formation of martensite and ultrafine carbides, along with microstructure refinement in the remelting process" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0002494_med.2017.7984152-Figure1-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0002494_med.2017.7984152-Figure1-1.png", "caption": "Fig. 1: Body frame of a hexacopter with PNPNPN configuration, where P denote that a rotor rotates clockwise, and N denotes that a rotor rotates counter-clockwise.", "texts": [], "surrounding_texts": [ "I. INTRODUCTION\nVertical take-off and landing Unmanned Aerial Vehicles (UAVs) have been used in a wide range of applications nowadays, for example, bridge inspection, package delivery, precision agriculture, etc. They have several advantages compared with conventional helicopters or planes, such as the mechanical simpleness [1], and quite cheap a prototype implementation. However, there are still some barriers slowing down their integration into our civil airspace, especially the poor reliability and low robustness. For example, abrupt mechanical fault during a flight mission may cause vehicle damage, even personal injury [2].\nA multicopter UAV is propelled by a group of actuators, each actuator is constituted of a motor, a rotor, and a propeller. Experience have indicated that actuators used in multicopter UAVs may suffer from some faults like rotor fast aging and wear, motor/propoller sudden damage, etc [3]. These faults may lead to a deterioration of the aerodynamic efficiency, or even urgent dangerous status. Recently, with growing demands of reliability and an increasing awareness about the system malfunction, safety has become an essential concern of safety-critical systems [4], [5]. Therefore, to reduce the impact of the foreseeable actuator fault problem, it is necessary to carefully design fault tolerant methods and strategies for multicopter UAVs.\n\u2020 the State Key Lab. of Industrial Control Technology, Zhejiang University, Hangzhou, \u2021 the Dept. of Automation, Shanghai Jiao Tong University, Shanghai, China. e-mail: (jiesu.zju@gmail.com, jphe@sjtu.edu.cn, {pcheng, jmchen}@iipc.zju.edu.cn).\nThis work is partially supported by NSFC under grant U140120094, National Key R&D Program under grant 2016YFB0800204, and the Fundamental Research Funds for the Central Universities.\nTo ensure a safe flight mission, different Fault Tolerant Control (FTC) mechanisms have been applied to multicopter UAV in the last decade. FTC can be classified into active and passive methods. Passive Fault Tolerant Control (PFTC) schemes treat faults as bounded model disturbances without online detection, and aim at synthesizing one robust controller, which can make the closed-loop system (nearly) insensitive to certain faults [6]. Various PFTC approaches tackling with actuator fault problem of multicopter have been proposed in [7]\u2013[9]. These applicability are restricted to faults that have a small effect on the behavior of the system.\nDifferent with PFTC approaches, Active Fault Tolerant Control (AFTC) schemes integrate a Fault Detection, Isolation and reconstruction (FDIR) module with a particular control allocation mechanism which can achieve fault tolerant effect [10]. The control allocation methods aim to map the control command to the over-actuated actuator group [11]. For a multicopter, the control allocation methods include pseudo inverse, cascading generalized inverse [12], multi-parametric programming [3] and LPV based control allocation method [13], etc. These methods rely on the fault location and efficiency loss information, which is of great importance. Thau observer [14] and least-square regression [15] are adopted as the FDIR method, both of which tend to be sensitive but demand huge computational resources. Sliding Mode Observer (SMO) techniques are utilized in [1], [16]\u2013[18] to generate residuals for FDIR purpose. Such an approach is based on the fact that the actuator fault will yield a change of the attitude dynamics. However, SMO has a complex architecture, and the super twisting parameters of SMO are hard to design and tune. Meanwhile, due to the chattering effect of SMO, the fault reconstruction performance may lead to an extravagant control allocation.\nConceptually, the direct implemented FDIR scheme based on parity relationship is more straightforward than the observer-based approach [19]. Such a method is typically called Analytical Redundancy (AR), which exploits the nullspace of the state-space observability matrix to allow the creation of a set of test residuals to learn about the system behavior [20]. In [21], the linear AR is extended into nonlinear realm, named as Non-linear Analytical Redundancy (NLAR), in which a novel nonliner model based residual generation approach is proposed. Both of SMO and NLAR approaches are applied to an autonomous electric vehicle with FDIR purpose in [22], whose results show that the smoothness and signal reconstruction performance of NLAR residuals are better than SMO residuals. Inspired by the results of [22], we apply the NLAR techniques to a hexacopter, to construct\n978-1-5090-4533-4/17/$31.00 \u00a92017 IEEE 413", "a novel FDIR approach. However, the attitude dynamics of a hexacopter is different with an autonomous electric vehicle, such an application is challenging. Proceeding with residuals derivation from the attitude dynamics of the hexacopter, we can choose the model related ones to detect and isolate the faulty actuator. Using low pass filter, the generated residuals can be reconstructed to the additive fault. Main contributions of this paper are summarized as follows:\n1) Utilizing the NLAR algorithm, we derive the residuals\nof the hexacopter attitude dynamics based on the state space model. The derived NLAR residuals are highly model-relevant, simple and straightforward. The generated residuals can be applied to multicopters with the other frame configurations, such as star frame configured quadrotor or octorotor, etc. 2) Three model related residuals are chosen to detect\nand isolate the faulty actuator. Due to the existance of gyroscope measurement noise, we determine the bilateral threshold for the residuals. Then a first order butterworth filter is applied to the chosen residuals for reconstructing the fault signal. 3) We validate and compare the isolation effectiveness\nand reconstruction performance of our approach and SMO based approach [1], [16], [17] via a real experiment, respectively.\nA star configuration of the hexacopter is shown in Fig. (1), which is considered as a single 6-Degree of Freedom (DoF) rigid body. A body-fixed reference frame is named as RB , and xB , yB , zB axis are originating at the mass center of the vehicle. xB , yB , zB points to the head, the right side and the down side of the rigid body relative to the earth frame, respectively. Deriving from the Euler-Lagrange approach, we obtain the equations demonstrating the motion of the hexacopter [23],\u23a7\u23aa\u23aa\u23aa\u23aa\u23aa\u23aa\u23aa\u23aa\u23a8 \u23aa\u23aa\u23aa\u23aa\u23aa\u23aa\u23aa\u23aa\u23a9 P\u0308x = (cos\u03c6 sin \u03b8 cos\u03c8 + sin\u03c6 sin\u03c8) \u2217 T m , P\u0308y = (cos\u03c6 sin \u03b8 sin\u03c8 \u2212 sin\u03c6 cos\u03c8) \u2217 T m , P\u0308z = cos\u03c6 cos \u03b8 \u2217 T m \u2212 g, \u03c9\u0307x = Iyy\u2212Izz Ixx \u03c9y\u03c9z \u2212 Jr Ixx \u03c9y\u03c9\u0393 + \u03c4\u03c6 Ixx , \u03c9\u0307y = Izz\u2212Ixx Iyy \u03c9x\u03c9z + Jr Iyy \u03c9x\u03c9\u0393 + \u03c4\u03b8 Iyy ,\n\u03c9\u0307z = Ixx\u2212Iyy Izz \u03c9x\u03c9y + \u03c4\u03c8 Izz ,\n(1)\nwhere Px, Py and Pz denote the position of the vehicle relative to the earth frame, and \u03c6, \u03b8 and \u03c8 are roll, pitch, yaw angles of the vehicle body relative to the earth frame. \u03c9x, \u03c9y and \u03c9z represent the angular velocities of the vehicle relative to the body frame. Ixx, Iyy, Izz are the moments of inertia of the hexacopter on the three axes relative to the bodyfixed reference frame, respectively. Jr is the rotor inertia. \u03c9\u0393 represents the disturbance depending on the rotor rotation speed, which is described by \u03c9\u0393 = \u03c91 \u2212 \u03c92 + \u03c93 \u2212 \u03c94 + \u03c95\u2212\u03c96, where \u03c9i, i = 1, 2, 3, 4, 5, 6, represents the rotation speed of the i\u2212th rotor. T represents the thrust, \u03c4\u03c6, \u03c4\u03b8 and\n\u03c4\u03c8 denote the roll, pitch, yaw torque, respectively. Given the hexacopter frame as shown in Fig. (1), these variables are calculated by,\u23a7\u23aa\u23aa\u23aa\u23aa\u23aa\u23aa\u23a8 \u23aa\u23aa\u23aa\u23aa\u23aa\u23aa\u23a9 T = Kf \u22116 i=1 \u03c9i 2, \u03c4\u03c6 = Kfd(\u2212\u03c912 + \u03c92 2) \u2212Kf d 2 (\u03c93 2 \u2212 \u03c942 \u2212 \u03c952 + \u03c96 2), \u03c4\u03b8 = Kf \u221a 3 2 d(\u03c93 2 \u2212 \u03c942 + \u03c95 2 \u2212 \u03c962),\n\u03c4\u03c8 = Kd(\u03c91 2 \u2212 \u03c922 + \u03c93 2 \u2212 \u03c942 + \u03c95 2 \u2212 \u03c962).\n(2)\nwhere fi = Kf\u2217\u03c9i2, i = 1, 2, 3, 4, 5, 6, is the thrust produced by the i-th propeller, Kf denotes the corresponding thrust factor, and d represents the arm length. \u03c4i = Kd \u2217 \u03c9i2, i = 1, 2, 3, 4, 5, 6, is the counter-torque of the i-th rotor, Kd is the corresponding counter-torque coefficient. The equation transforming the angular velocities from the body frame to the earth frame is written as,\u23a1 \u23a3\u03c6\u0307\u03b8\u0307 \u03c8\u0307 \u23a4 \u23a6 = \u23a1 \u23a31 sin\u03c6 tan \u03b8 cos\u03c6 tan \u03b8 0 cos\u03c6 \u2212 sin\u03c6 0 sin\u03c6/ cos \u03b8 cos\u03c6/ cos \u03b8 \u23a4 \u23a6 \u23a1 \u23a3\u03c9x\u03c9y \u03c9z \u23a4 \u23a6 .\nB. IMU Complementary Filter\nSince the transformation matrix suffers the gimbal lock problem, \u03b8 = (2k\u22121)\u03c0/2, we utilize Direct Cosines Matrix (DCM) as the representation of the attitude dynamics, to circumvent the singularities associated with the computation\nof Euler angles. The DCM is written as WT = [ rx ry rz ] , where the coordinate vector ri, (i = x, y, z), represents the coordinates of the i-th basis vector of the earth frame as measured in the body-fixed frame. Therefore, R is a denotation of the relative orientation of the body frame from the earth reference frame. The kinematic equation for the body rotation is,\nW\u0307 = W\u03a9\u00d7,\nwhere \u03a9\u00d7 conducts to the skew symmetric calculation of the measured rotation rate vector from the gyroscope, satisfying\n\u03a9\u00d7 = \u23a1 \u23a3 0 \u2212\u03c9x \u03c9y \u03c9z 0 \u2212\u03c9x \u2212\u03c9y \u03c9x 0 \u23a4 \u23a6.", "Utilizing the passive complementary filter provided in\n[24], we obtain that,\n\u02d9\u0302 W = W\u0302\u03a9\u00d7 + kP e+ kI\n\u222b t+\u03b4t\nt\ne(\u03c4)d\u03c4,\ne = a\u00d7 d, d = \u2212W\u0302T ez,\nwhere W\u0302 is the DCM estimation, kP and kI represent the proportional and integral gain of a PI controller for correcting the rotation rate vector, respectively. e denotes the error vector, conducted by the cross product of the normalized accelerometer measurement a and the determined gravity vector d. We use the current attitude estimation W\u0302 and the current z axis error vector of body frame. The conversion from DCM to Euler angle relative to the earth frame is, \u03c6 = \u2212Asin(rz(1)), \u03b8 = Atan2(rz(2), rz(3)), \u03c8 = Atan2(ry(1), rx(1)), where ri(j), i = x, y, z, j = 1, 2, 3, denotes column j of the i-th row vector.\nTo stabilize the attitude dynamics of the hexacopter, we uitlize a PID controller to correct the error obtained from the IMU complementary filter. The PID controller is given by,\nU = KP e+KD e\u0307+KI\n\u222b t\n0\ne(\u03c4)d\u03c4,\nwhere KP ,KI ,KD denote the controller\u2019s gains, and e is difference between the angles and the desired ones, respectively. U represents the system input, which includes four\nvariables, U = [ T \u03c4\u03c6 \u03c4\u03b8 \u03c4\u03c8 ]T . From equation (2), we extract the control allocation matrix from the control law to the rotation speed of each actuator, which is described by\nB = \u23a1 \u23a2\u23a2\u23a3 Kf Kf Kf Kf Kf Kf \u2212Kfd Kfd \u2212Kf d 2 Kf d 2 Kf d 2 \u2212Kf d 2\n0 0 Kf \u221a 3 2 d \u2212Kf \u221a 3 2 d Kf \u221a 3 2 d \u2212Kf \u221a 3 2 d\nKd \u2212Kd Kd \u2212Kd Kd \u2212Kd\n\u23a4 \u23a5\u23a5\u23a6 .\nDenote \u0393 = [ \u03c91 2 \u03c92 2 \u03c93 2 \u03c94 2 \u03c95 2 \u03c96 2 ]T as the control allocation output vector. Since U = B \u00b7 \u0393, utilizing pseudo inverse method, we obtain that\n\u0393 = BT \u00b7 (B \u00b7BT )\u22121 \u00b7 U. (3)\nThe fault of the actuators are modeled as a control\neffectiveness loss percentage [16], which is described by,\n\u03b7 = [\u03b71, \u00b7 \u00b7 \u00b7 , \u03b76]T , (4)\nwhere \u03b7i \u2208 [0, 1] indicates the actuator faulty severity. Define Bf as the faulty control allocation matrix, it can be calculated by a multiply of equation (3) and equation (4), namely,\nBf = B \u00b7 \u03b7. Utilizing Bf , we can reconfigure rotation speed of each actuator to adapt the actuator fault. However, it depends on obtaining \u03b7 beforehand, revealing the fault isolation and\nreconstruction result. This result relies on the generated residuals, which needs burdensome calculations. Meanwhile, since not all of the residuals can be utilized, it is challenging to choose the useful ones. Last but not least, performing the fault reconstruction needs a kind of low pass filter. Thus, we explore these uncertainties to construct this approach in the rest of this paper.\nThe proposed approach is constituted of a residual generator, a faulty isolator and a fault signal reconstructor. The residual generator requires an exploration of the combination between the attitude dynamics and NLAR algorithm. Given the residuals, it is challenging to detect and isolate the faulty actuator. Finally, to estimate the efficiency loss of the faulty actuator, we need to reconstruct the fault signals.\nFrom the dynamic model of the hexacopter (1), we realize that the change of the attitude dynamics is a direct reflection of the actuator fault. Meanwhile, if the efficiency loss of the actuator is tiny, the change of the attitude dynamics may be subtle, thus we can make the following approximation, \u03c6\u0307 = \u03c9x, \u03b8\u0307 = \u03c9y, and \u03c8\u0307 = \u03c9z . Then, it follows that,\u23a7\u23aa\u23a8 \u23aa\u23a9 \u03c6\u0308 = Iyy\u2212Izz Ixx \u03b8\u0307\u03c8\u0307 \u2212 Jr Ixx \u03b8\u0307\u03c9\u0393 + \u03c4\u03c6 Ixx , \u03b8\u0308 = Izz\u2212Ixx Iyy \u03c6\u0307\u03c8\u0307 \u2212 Jr Iyy \u03c6\u0307\u03c9\u0393 + \u03c4\u03b8 Iyy ,\n\u03c8\u0308 = Ixx\u2212Iyy Izz \u03c6\u0307\u03b8\u0307 + \u03c4\u03c8 Izz .\n(5)\nFor adopting NLAR [21] algorithm to generate the residuals, we rewrite the approximated attitude dynamics as the state space model, { x\u0307(t) = f(x(t)) + gu(t),\ny(t) = x(t). (6)\nwhere x(t) = [\u03c6\u0307 \u03b8\u0307 \u03c8\u0307]T represents the vector including three angular velocities of the earth frame, g = diag[ 1 Ixx 1 Iyy 1 Izz\n], and y(t) represents the IMU gyroscope output.\nThen we derive the O\u0394 matrix to calculate the corresponding left null space \u03a9\u22a5. The triangular nonlinear observability matrix O\u0394 is expressed by,\nO\u0394 = \u23a1 \u23a2\u23a2\u23a2\u23a2\u23a2\u23a2\u23a2\u23a2\u23a2\u23a3\ny(t) Lfy + Lgyu\nLffy + Lgfyu+ Lfgyu+ Lggyu 2\u23a7\u23a8\n\u23a9 Lfffy + (Lgffy + Lfgfy + Lffgy)u+ (Lggfy + Lgfgy + Lfggy + Lgggy)u\n2+ (2Lgfy + Lfgy)u\u0307+ 3Lggyuu\u0307+ Lgyu\u0308\n\u23ab\u23ac \u23ad\n...\n\u23a4 \u23a5\u23a5\u23a5\u23a5\u23a5\u23a5\u23a5\u23a5\u23a5\u23a6 .\nLky = \u2211n i=1 \u2202y(x) \u2202x ki(x) is the Lie derivative of scalar function y in the direction of vector function k [21]. We utilize the following script to express the Lie derivative, Li(Lj(Lky)) = LiLjLky = Lijky. The parity matrix \u03a9\u22a5" ] }, { "image_filename": "designv11_62_0000458_amr.591-593.1879-Figure4-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0000458_amr.591-593.1879-Figure4-1.png", "caption": "Fig. 4 3D finite element model of an Fig. 5 Boundary conditions of the engine model. industrial turbo engine.", "texts": [ " Conclusion can be drawn that these two models have identical loads since they have similar model and boundary condition. The presented method accurately describes the effects of the unbalance on the solid element model. Unbalance response analysis can be performed conveniently in solid element rotor model with this method. 3D FE Model of an Industrial Turbo Engine. A solid model of an industrial turbo engine is established and a finite element mesh is developed using ANSYS. Part of the whole casing and the low compressor rotor are selected for analysis. Fig. 4 shows the finite element mesh of the selected part of the turbo engine with the top casing half removed. The blades on the rotor disks are represented by solid annuals with the same mass and inertia moment. The stator blades of the engine are modeled as mass elements. There are three support bearings between the casing and rotor, all of which are modeled using COMBIN14 element in ANSYS. The remaining parts are modeled using solid185, one kind of 8 node solid element in ANSYS. The total elements in the whole model are 108,051 and the corresponding nodes 193,544" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0001517_0731684409348345-Figure8-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0001517_0731684409348345-Figure8-1.png", "caption": "Figure 8. Contact line between female and male rotor: (a) Imf\u00bc 4:1, (b) Imf\u00bc 5:1.", "texts": [ " The formation stage of contact line can be regarded as when the imaginary contact line on outside of end section of female and male rotor gradually moves forward to the whole effective length. This process is the feeding stage of high viscosity material. The extinction stage of the contact line can be regarded as when the contact line gradually moves forward to exhaust end from the effective length of rotor. One of the contact lines of female and male rotor on surface of female rotor is shown in Figure 8. The novel kneader cannot be used as a screw pump because of the special teeth ratio Imf. There exists a very large open area in meshing area while Imf\u00bc 4:1 and Imf\u00bc 5:1, at UNIV OF CALIFORNIA SANTA CRUZ on November 25, 2014jrp.sagepub.comDownloaded from as shown in Figure 9. A certain backflow will occur, and then good mixture will be induced by these open areas following axial direction under pressure. The female and male screw rotors of the novel twin-screw kneader machined by the CBN wheel are shown in Figure 10" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0000948_icici-bme.2013.6698502-Figure4-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0000948_icici-bme.2013.6698502-Figure4-1.png", "caption": "Fig. 4 Illustration of the half-plane switching concept [5]", "texts": [ " The pseudo-code design for Orbit Following to calculate the course command and the altitude command is stated as follows [5] Input: Orbit center c = (cn, ce, cd) T, radius \u03c1, direction \u03bb, actual position p=(pn, pe, pd) T, course angle \u03c7, gains korbit, sample rate Ts . 1: hc \u2190\u2212cd 2: d \u2190 3: \u03d5 \u2190 atan2(pe \u2212 ce, pn \u2212 cn) 4: while \u03d5 \u2212 \u03c7 < \u2212\u03c0 do 5: \u03d5 \u2190 \u03d5 + 2\u03c0 6: end while 7: while \u03d5 \u2212 \u03c7 > \u03c0 do 8: \u03d5 \u2190 \u03d5 \u2212 2\u03c0 9: end while 10: \u03c7c(t) = \u03d5 + \u03bb 11: return hc , \u03c7c D. Path Manager To overcome the problems of transition between two waypoints, the Path Manager block is used. The best concept is the half-plane switching [5] as illustrated in Fig. 4. The pseudo-code design of the Path Manager is given as follow [5] Input: Waypoint locations Wp = {w1, w2, w3, \u2026, wn}, Position p = (pn, pe, pd) T , fillet radius R. Require: N \u2265 3 1: if New waypoint path Wp is received then 2: Initialize waypoint index: i \u2190 2, and state machine: state \u2190 1. 3: end if 4: qi\u22121 \u2190 5: qi \u2190 6: \u2190 cos\u22121( ) 7: if state = 1 then 8: flag \u2190 1 9: r \u2190 wi\u22121 10: q \u2190 qi\u22121 11: z \u2190 12: if p H(z, qi\u22121) then 13: state \u2190 2 14: end if 15: else if state = 2 then 16: flag \u2190 2 17: c \u2190 18: \u03c1 \u2190 R 19: \u03bb \u2190 sign(qi\u22121,nqi,e \u2212 qi\u22121,eqi,n) 20: z \u2190 21: if p H(z, qi ) then 22: i \u2190 (i + 1) until i = N \u2212 1" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0000889_elan.201200242-Figure6-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0000889_elan.201200242-Figure6-1.png", "caption": "Fig. 6. (A) Square wave voltammetric signal (generator current topography plot shown, 2 Hz, 20 mV amplitude, 1 mV step potential) for 1 mM indigo carmine in 0.1 M KCl. (B) Collector current topography plot. (C) Generator current contour plot. (D) Collector current contour plot.", "texts": [ " A separate peak response for reduction of indigo carmine and for oxidation of leuco-indigo carmine is observed here probably due to a multistep protonation with a less fast equilibration of intermediates. Both oxidation and reduction mechanism are summarised in Equation 5. Processes IV.: Indigo Carmine Reduction in KCl In the absence of buffer, voltammetric responses become affected by additional pH gradients between generator and collector electrode. In this case new additional phenomena are observed such as a mismatch between the appearance of the generator and collector square wave voltammetry contour features. Figure 6 shows topography and contour plots for the reduction of 1 mM indigo carmine at the gold\u2013gold junction electrode. The square wave voltammetric response at the generator electrode is dominated by the peak at 0.45 V vs. SCE which is consistent with a locally generated proton activity of ca. pH 11 due to the pKa of the leuco-indigo carmine product [32] and the corresponding consumption of ca. 1 mM protons locally at the generator electrode surface. Perhaps surprisingly, when the collector potential is fixed at more negative potentials the collector response (see Figure 6D) shows similarity to that observed previously at pH 7 (see Figure 5D) with a strong peak at ca. 0.28 V vs. SCE. It is possible that the transport of protons (associated with reduced indigo carmine) generates a local pH gradient in KCl solution with conditions at the collector electrode creating a slightly more acidic environment. In future, in order to reveal a more quantitative insight into the resulting pH pattern and into the quantitative implications of the square wave voltammetry data as a function of both, collector and generator potential, a 3D simulation tool capable of treating multistep electron transfer with coupled equilibria at a relatively complex geometry will be desirable" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0000608_j.triboint.2010.11.012-Figure1-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0000608_j.triboint.2010.11.012-Figure1-1.png", "caption": "Fig. 1. Orifice-type restrictors: (a) orifice restrictor and (b) inherently compensated restrictor.", "texts": [ " The simulation results have revealed clearly that the effects of orifice and inherence in series on bearing performance cannot be ignored in the design of aerostatic bearings. & 2010 Elsevier Ltd. All rights reserved. 1. Introduction Stability is very important in both the development of aerostatic bearings and their application to precision machine tools. Many researchers have adopted both analytical and experimental methods to investigate various types of aerostatic bearing that are compensated by orifice-type restrictors. The two typical types of orifice-type restrictors are an orifice restrictor with a pocket, as shown in Fig. 1(a), and an inherently compensated restrictor formed by a feeding hole, as shown in Fig. 1(b). The inherence restriction caused by the inherently compensated restrictor depends on the change of thickness of air film in bearing clearance, while the orifice restriction caused by the orifice restrictor is irrelevant to the thickness of an air film. Lund [1\u20133] analyzed pressurized gas journal bearings and acquired data related to the load capacities as well as the stiffness and damping coefficients that can be used to verify the whirl instability of a rotor supported by aerostatic bearings" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0001467_9783527644117.ch5-Figure5.12-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0001467_9783527644117.ch5-Figure5.12-1.png", "caption": "Figure 5.12 Three - dimensional drawing of the modular stack half - cell showing the central reaction chamber with reference electrode inlet (bottom right) and solution fi lling ports, the working electrode plate with inlet for the", "texts": [ "), usually with little consideration of the stability of such power or of the consequences for powering devices. Comparison of power output from such assemblies under the range of conditions utilized thus becomes, as is evident from previous comparison of bioanode and biocathode performance, problematic. A welcome contribution to methodologies for comparison of bioelectrocatalytic fuel cell electrodes is the use of a standardized fuel cell set - up [25] . The set - up is actually a half - cell, anode, with a platinum mesh as counter electrode, assembled, as presented in Figure 5.12 , to permit fl ow presentation of fuel to a glassy carbon anode. The reproducibility of the glassy carbon anode surface area was verifi ed using voltammetry of a solution - phase ferricyanide redox probe. Reproducibility of voltammetry for deposited polymethylene green and for the voltammetric response of the deposited fi lm to NADH oxidation is presented, with a view for use as a bioanode for NADH - dependent dehydrogenase reactions. Results from the use of such a standardized methodology should prove invaluable for comparison of surfaces, mediators, and enzymes in bioanodes and BFCs, and are eagerly anticipated" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0000249_amm.459.390-Figure1-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0000249_amm.459.390-Figure1-1.png", "caption": "Fig. 1 2D-Axi-symmetric model of tire [8]", "texts": [ " Abaqus solves the problem by computing the current configuration of the acoustic domain periodically creating a new mesh. The new mesh uses the same topology throughout the simulation, but the nodal locations are adjusted periodically so that the deformation of the structural acoustic boundary does not lead to the severe distortion of the acoustic elements. The calculation of updated nodal locations is based on adaptive mesh smoothing. Fig.2 shows the adaptive meshing of the present model. Tire. A 2D model of a passenger car tire (PCR) 175/70R13 considered for the present study is shown in Fig.1. The components of the tire are two steel belts, one capply, one body ply, and one chafer ply. Yeoh material model with constants taken from [8] is used in the creation of this finite element model. The portion of the acoustic cavity is filled by air with bulk modulus of 426 kPa. The meshing of the entire model including air is carried using Hypermesh software. The tire is modeled in Abaqus using CGAX4H (4-noded bilinear element with twist and hybrid formulation) and CGAX3H are the elements for rubber components and SFMGAX1 (2-noded linear axi-symmetric surface element with twist) is the element for belt layers" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0000900_ever.2013.6521527-Figure8-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0000900_ever.2013.6521527-Figure8-1.png", "caption": "Fig. 8. Vector diagram of the PMAREL motor.", "texts": [ " The d- and q-axis stator flux linkages versus d- and q-axis currents of the synchronous REL motor are shown in Fig. 7. Comparing with the results in Fig. 2, it is worth noticing that the behaviour of the main flux linkage (i.e., the d-axis flux Iinkage) remains the same. IV. PM ASSISTED RELUCTANCE MOTOR According to the axis notation given for the REL motor, a permanent magnet (PM) is added so as to produce a PM flux Iinkage Am in the negative q-axis. Therefore, the flux Iinkages become: Ad Ldid Aq Lqiq - Am The corresponding vector diagram is shown in Fig. 8. (17) (18) The advantage of using a PM in the rotor is twofold. At first, a part of the PM flux saturates the iron bridges of the rotor. Secondly, the PM flux linkage tends to reduce the angular distance between the current vector angle and the voltage vector angle, that is, it tends to reduce the power factor angle. As a consequence, there is a slightly higher torque for the same current, with a corresponding increase of the motor efficiency. In addition, a noticeable increase of the power factor is achieved, that implies a reduction of the inverter volt ampere ratings" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0000714_s00521-012-0916-3-Figure2-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0000714_s00521-012-0916-3-Figure2-1.png", "caption": "Fig. 2 Sketch of a traditional equipment for porthole die extrusion", "texts": [ " The Taguchi\u2019s Design of Experiments (TDoE) allows both to highlight the main factors to be considered into the investigation phase and to verify the quality of the investigated NNs starting from the evaluation of some performance indexes (i.e. ANOVA, Normal Probability Plot). The solution results more robust, and the performance costs low [19]. The extrusion by porthole die is a process that is always more used in order to produce hollow components; parts with different sections can be obtained. The process takes place through the complex die structure composed of container, porthole, mandrel, welding chamber and bearing part (Fig. 2); the working sequence consists of three stages that can be identified like dividing, welding and forming stage. The mandrel legs divide the material before that it flows inside the welding chamber, where due to simultaneous action of temperature and pressure, it joins again forming welding lines [31]. Solid joining mechanics occur during their formation similar, for instance, to the friction stir welding [32]; two typologies of welding lines, especially, can be found on the extruded parts, namely longitudinal and transverse lines [33]" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0002642_978-3-319-66866-6_2-Figure4-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0002642_978-3-319-66866-6_2-Figure4-1.png", "caption": "Fig. 4. Definition of design space (a) Relevant load case of the initial model (b) Assembly and application restrictions (c) Result of topology optimization (d) Extended design space", "texts": [ " To validate the design method, a pedal crank is used as a demonstrator. The objective is a maximal weight reduction with constant internal (von Mises) stresses compared to conventional models (respectively material characteristics). Except considering relevant interfaces, the component dimensions are freely selectable. 3.1 Clarification of Requirements The initial model is conventionally manufactured with AlSi10Mg alloy and has a weight of m = 217.45 g. Two interfaces for fixing the bottom bracket and the pedals define the length l = 170 mm, as depicted in Fig. 4-a. According to Sullivan und Chris, different load cases occur during the lifecycle. In assumption of an idealized model, in which torsional forces are neglected (due to the external force introduction), the critical load case is pure bending [16, 29]. In combination with the load vector Fmax = 2.250 N, which represents the maximum occurring forces with an additional safety factor, the bending load is used for further optimization [29]. 3.2 Definition of Design Space As shown in Fig. 4-b, the design space is limited by assembly and application restrictions as well as the size of the process chamber (machine Eosint M280). This rough estimation is used as an input for topology optimization (using material characteristics shown in Fig. 3) in order to narrow down the design space. Figure 4-c shows the optimization results as an iterative elimination of the volume elements with the lowest stresses (Ansys Workbench 17.0). It can be seen that an asymmetric topology occurs. For the following potential analysis and selection of an internal structure, the optimization result is rebuilt as shown in Fig. 4-d. In order to identify a suitable internal structure to be integrated in the design space, the asymmetrical shape is neglected in the first iteration. The results are considered again while dimensioning for stress reduction. Hence, the topology is iteratively adapted to the asymmetric shape. Preliminary investigations at the institute describe the modeling and simulation of digital specimens with internal structures [9]. In addition to the application of computer-aided tools, the results are validated by analyzing physical specimens with a static test bench" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0002311_ls.1385-Figure15-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0002311_ls.1385-Figure15-1.png", "caption": "FIGURE 15 Test bench schematic diagram", "texts": [ "4 \u03bcm, and the value when eg = 0.1 is the minimum, it is 0.3 \u03bcm. A standard ball is placed on the center of worktable of the hydrostatic spindle, and a micrometer is placed on the base of the spindle system, and the probe of the micrometer is touch with the standard ball, the eccentric of the standard ball and the rotation axis is adjust to reach the measurement range of the inductance meter, and the test is start after the spindle system is stable, as shown in Figure 14.The schematic diagram of the test bench is shown in Figure 15. The rotary error result is shown in Figure 16; this shows that the maximum rotary error is 0.26 \u03bcm, and it indicates that the actual orbit of the shaft is less than the value when the eccentricity is 0.1 in Figure 11.The experimental result that verifies the rotation accuracy of the spindle system is not affected by the imbalance basically. The perturbed Reynolds equation is solved by the finite difference method, and the disturbance pressure distribution of oil film is gained. The load increment method and perturbation pressure method are used to the solution of the 4 dynamic stiffness coefficients and damping coefficients" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0000900_ever.2013.6521527-Figure9-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0000900_ever.2013.6521527-Figure9-1.png", "caption": "Fig. 9. Flux lines in the PMAREL motor.", "texts": [], "surrounding_texts": [ "The behaviours of the d-axis flux linkage versus the d-axis current and the q-axis flux linkage versus the q-axis current of the three motors under study are shown in Figs. 2, 7 and 10, for IM, REL and PMAREL motor respectively. The behaviour of the d-axis flux Iinkage is alm ost the same as it is expected since the magnetic circuit is very similar: the motors are characterised by the same stator, the same air gap, and by the same width of the rotor magnetic paths. On the contrary, a different behaviour is observed along the q-axis. The REL motor exhibits a higher slope. This means a higher q-axis inductance and a lower torque-to-current ratio (at least until saturation is limited). This is confirmed by comparing the torque T in Table 11, according to the same stator current (amplitude and angle). The insertion of the PMs tends to reduce the q-axis flux Iinkage for given current, with a corresponding increase of rated torque and, correspondingly, efficiency and power factor, see Table 11." ] }, { "image_filename": "designv11_62_0000254_i2010-10560-0-Figure1-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0000254_i2010-10560-0-Figure1-1.png", "caption": "Fig. 1. Schematic representation of the geometry and the definition of the distortion angles \u03b8 and \u03d5.", "texts": [ "C voltage-induced static domains (without electrohydrodynamic movement) oriented along the initial planar alignment of the nematic director n with an electrically controlled period were observed for the first time by Vistin\u2019 [1,2]. Bobylev and Pikin [3,4] first proved the flexoelectric nature of these domains by developing a theory including three parts in the \u201celectric enthalpy\u201d of the nematic: elastic, flexoelectric and dielectric. They regarded one nematic completely aligned along the X axis (in the plane of the electrodes) excited by a vertical constant electric field applied along the axis Z (see fig. 1). In their theory the nematic is either dielectrically stable (\u0394\u03b5 < 0) or dielectrically unstable (\u0394\u03b5 > 0) while flexoelectrically unstable. The competition between the flexoelectric and dielectric torques leads to threshold development of flexoelectric instabilities appearing in the form of continuous domains along the two angles: the polar (or zenithal one) \u03b8 and the azimuthal one \u03d5 (see fig. 1). The authors considered the case of small deformations (nx \u223c 1, ny \u223c \u03d5, nz \u223c \u03b8, \u03b8 = \u03b8(y, z), \u03d5 = \u03d5(y, z)) and assumed the following simple solution: \u03b8 = \u03b80 cos(qy) cos(\u03c0z/d), \u03d5 = \u03d50 sin(qy) cos(\u03c0z/d) for the case of strong anchoring of the director at both glass plates (the inner surfaces serve as electrode planes). They established two coupled differential equations for the variables \u03b8 and \u03d5 after the minimization of the \u201celectric enthalpy\u201d, accepting isotropic elasticity K11 = K22 = K. Given the requirements for the non-trivial solution they have obtained the dispersion a e-mail: ymarinov@issp", " In vector form it has the following expression: HE = \u222b V { 1 2 (K11(div n)2 + K22(n \u00b7 rotn)2 +K33(n \u00d7 rotn)2) \u2212e1zE \u00b7 ndiv n \u2212 e3xE \u00b7 (rotn \u00d7 n) \u2212\u0394\u03b5 8\u03c0 (n \u00b7 E)2 } dV, (4) where the unit vector n is the nematic director, K11, K22, and K33 are the elastic constants of splay, twist and bend, e1z and e3x are the flexoelectric coefficients of splay and bend, \u0394\u03b5 is the dielectric anisotropy (in our case negative or positive), the vector E is the d.c. electric field. Introducing the following components of n, E: nx = cos \u03b8 cos \u03d5, ny = cos \u03b8 sin \u03d5, nz = sin \u03b8, Ez = E (a planar-oriented nematic layer and vertically applied electric field) (see fig. 1), and accepting that for the case of longitudinal flexoelectric domains nx \u223c 1, finally we have expressed the \u201celectric enthalpy\u201d through the director components ny and nz and their derivatives with respect to y and z as follows: HE = \u222b\u222b { 1 2 ( K11 ( \u2202ny \u2202y + \u2202nz \u2202z )2 +K22 ( \u2202nz \u2202y \u2212 \u2202ny \u2202z )2 ) \u2212e1zEnz ( \u2202ny \u2202y + \u2202nz \u2202z ) \u2212 e3xEny ( \u2202nz \u2202y \u2212 \u2202ny \u2202z ) \u2212\u0394\u03b5 8\u03c0 n2 zE 2 } dydz. (5) The expression (5) shows that only splay and twist deformations enter the elastic energy [6]. The minimization of the \u201celectric enthalpy\u201d with respect to the director components ny and nz and their derivatives with respect to the coordinates y and z yields two equations of Euler-Lagrange" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0000786_aero.2013.6496960-Figure13-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0000786_aero.2013.6496960-Figure13-1.png", "caption": "Figure 13 \u2013 Developed excavator with sensors", "texts": [ " In addition, the expansion plates contained large, circular arc areas to maintain contact with the wall surface, which holds the body position against the rotation action of the EA. These plates cover a large area of the outer surface of the subunit. The remaining uncovered spaces, i.e. the gaps between the propulsion subunits and those between expansion plates, are covered with dustproof material (aluminum evaporation sheets) to prevent soil from entering a subunit. Additional rubber friction sheets are placed on the outer surface of the expansion plates, increasing the friction forces while the robot is moving inside a launcher (see Fig. 13). Table 1 shows the specification for each subunit. The sensor (FlexiForce, Nitta) is attached between the link and the expansion plate (Figure 9). Each link has a sensor, and an expansion plate has two sensors. Each subunit has a total of eight sensors, each of which measure the vertical push force against the wall. The output voltage from a sensor is amplified with an amplified circuit, and it is measured by a computer through an A/D converter. Pushing Force Model of Expansion Plate inside Cylindrical Shape The pushing force of the expansion plates is modeled here", " Finally, we pulled the pull tension scale attached on the bar in the rotational direction. Table 2 shows the coefficient of friction for each experiment. The coefficient of friction was larger than 1.0 in three experimental conditions. We believe that this result is the irregularity of the rubber friction sheets and the support of the four radially situated expansion plates increase the friction. Development of the excavator We developed a subsurface explorer containing excavation and propulsion units with force sensors. Figure 13 shows the developed excavator with four propulsion subunits. The excavation unit connected to the front of the propulsion subunit and the transport part passed through the propulsion parts. A DC motor placed at the rear of the robot rotated the EA. The excavated and transported soil was discharged from discharge ports behind the propulsion unit. Table 3 shows the specifications of the robot. Propulsion Experiments in an Acrylic Pipe We measured the pushing force for four units and the peristaltic crawling movement to determine their relationship in a pipe with a diameter of 130 mm" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0000624_cdc.2012.6427081-Figure1-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0000624_cdc.2012.6427081-Figure1-1.png", "caption": "Fig. 1. Illustration of longitudinal aircraft entities", "texts": [ " The simplified aircraft dynamics can be described by [9] \u03b3\u0307 = L\u0304\u03b1\u03b1\u2212 g VT cos \u03b3 + L\u0304o \u03b1\u0307 = q + g VT cos \u03b3 \u2212 L\u0304o \u2212 L\u0304\u03b1\u03b1 \u03b8\u0307p = q q\u0307 = Mo +M\u03b4\u03b4 (1) with L\u0304o = Lo mVT , L\u0304\u03b1 = L\u03b1 mVT where \u03b3, \u03b1, \u03b8p represent flight path angle, angle of attack and pitch angle, respectively; p is the time rate-of-change of \u03b8p; m and g denote the aircraft\u2019s mass and gravity constant, respectively; VT is the aircraft\u2019s airspeed; L\u03b1 represents the lift curve slope, and Lo represents all of the other contributions to lift, such as Mach number, pitch rate, etc; \u03b4 is control surface deflections; M\u03b4 is the control pitching 978-1-4673-2064-1/12/$31.00 \u00a92012 IEEE 5374978-1-4673-2066-5/12/$31.00 2012 I moment, and Mo is moment contributions from all other sources such as angle of attack, Mach number, etc. Mo is often approximated by Mo = M\u03b1\u03b1 + Mqq. The aircraft longitudinal model is depicted in Fig. 1. At an operating point, Lo, L\u03b1, M\u03b4 , M\u03b1 and Mq are unknown, constant, scalar parameters [10]. We assume that the airspeed is maintained in a small neighborhood of the desired velocity by a simple linear controller such as PI controller, so VT is considered as a constant parameter. Inspection of the dynamics in (1) reveals that the dynamics of the triple (\u03b3, \u03b1, q) is in strict feedback form. Define the state space variables as x1 = \u03b3, x2 = \u03b1, x3 = q and the control input u = \u03b4. Considering external disturbances existing in real systems, the aircraft longitudinal dynamics in state space can easily be described as follows: x\u03071 = f1(x1) + a1x2 +\u22061(x, t) x\u03072 = f2(x1, x2) + x3 +\u22062(x, t) x\u03073 = f3(x2, x3) + a3u+\u22063(x, t) (2) with f1(x1) = \u2212 g VT cosx1 + L\u0304o, f2(x1, x2) = g VT cosx1 \u2212 L\u0304o \u2212 L\u0304\u03b1x2, f3(x2, x3) = M\u03b1x2 +Mqx3, a1 = L\u0304\u03b1 > 0, a3 = M\u03b4 > 0 where x = [x1, x2, x3] T is the state vector; \u2206i(x, t), i = 1, 2, 3 are unknown disturbances and satisfy |\u2206i(x, t)| \u2264 \u03c1i with \u03c1i being a positive constant" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0000983_i2013-13159-0-Figure1-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0000983_i2013-13159-0-Figure1-1.png", "caption": "Fig. 1. The nonlinear spring-pendulum model: (a) the part of the nonlinear spring-pendulum which is similar to the SD oscillator; (b) the rotating part of the nonlinear spring pendulum.", "texts": [ " Moreover, the approximation system can be obtained by simplifying the two-degree-of-freedom system, in which the dynamical behaviors are deteccted using the multi-dimensional method. Numerical simulations are also carried out in sect. 4, which indicate that the dynamical behavior of the approximation system is similar to the original system and its practicality is certified. a e-mail: tianrl@stdu.edu.cn Based on SD oscillator and pendulum model, we constructed a spring-pendulum system under nonlinear irrational restoring force as shown in fig. 1. The system comprises a lumped mass linked with a pair of inclined elastic stiff springs and can vibrate vertically and laterally. The vertical vibration of the system is shown in fig. 1(a) and the lateral vibration in fig. 1(b). Because of geometrical nonlinear characteristics, the system has a strong nonlinear restoring force. The mathematical vibration model of the spring-pendulum system is shown as follows: \u23a7 \u23aa\u23aa\u23a8 \u23aa\u23aa\u23a9 mX \u2032\u2032 + kX ( 1 \u2212 L\u221a X2 + l2 ) \u2212 m(h + X)(\u03d5\u2032)2 + mg(1 \u2212 cos \u03d5) = 0 m\u03d5\u2032\u2032 + 2m (h + X) X \u2032\u03d5\u2032 + mg (h + X) sin \u03d5 = 0 , (1) where X \u2032 = dX dt , \u03c6\u2032 = d\u03c6 dt , m, X, \u03c6, h, g, l, L are the mass of the oscillator, the vertical displacement, the lateral angular displacement, the vertical distance of the oscillator in balance, the acceleration of gravity, the half distance between the fixed points of the two springs and the original length of the springs, respectively" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0000900_ever.2013.6521527-Figure5-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0000900_ever.2013.6521527-Figure5-1.png", "caption": "Fig. 5. Vector diagram of the REL motor.", "texts": [ " The motor torque is computed from Maxwell's stress tensor integrated along the rotor periphery, directly from the field solution. Alternatively, the torque can be computed from the flux linkages as 3 Tim = 2PArdirq (11 ) which yields a satisfactory prediction of the average motor torque [20]. III. SYNCHRONOUS RELUCTANCE MOTOR In the REL machine, the d - q reference frame is determined by the rotor geometry. The d-axis commonly corresponds to the higher permeance path. The vector diagram of the REL motor is show in Fig. 5. The stator resistance and the 3-D effects are not included. According to the d - q axis currents id and iq, the flux Iinkages are expressed as: Ad = Ldid Aq = Lqiq (12) (13) where Ld is the d-axis inductance and Lq is the q-axis inductance. The ratio between the two inductances defines the saliency ratio of the REL motor, that is, \ufffd = Ld/ Lq\u2022 From the vector diagram, it is possible to obtain the follow equations: (14) where 0:; and o:\ufffd are the current vector and flux Iinkage vector angles, respectively, and 'Pi is the inner power factor angle", " In particular, the same stator geometry is considered, together with the stator winding distribution. The flux lines are shown in Fig. 6 In each FE analysis the stator currents id and iq are fixed. They are transformed through Park's transformation in the actual stator currents ia, ib and ic, which are assigned within the slots, according to the winding distribution. From the magnetic field solution, the REL motor performance are computed, by means of relations defined from the vector diagram of Fig. 5. The d- and q-axis stator flux linkages versus d- and q-axis currents of the synchronous REL motor are shown in Fig. 7. Comparing with the results in Fig. 2, it is worth noticing that the behaviour of the main flux linkage (i.e., the d-axis flux Iinkage) remains the same. IV. PM ASSISTED RELUCTANCE MOTOR According to the axis notation given for the REL motor, a permanent magnet (PM) is added so as to produce a PM flux Iinkage Am in the negative q-axis. Therefore, the flux Iinkages become: Ad Ldid Aq Lqiq - Am The corresponding vector diagram is shown in Fig" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0002394_b978-0-08-101022-8.00016-8-Figure17.2-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0002394_b978-0-08-101022-8.00016-8-Figure17.2-1.png", "caption": "Fig. 17.2 Needle-based electrospinning process for growing up nanofibrous fabrics directly onto microchips. On the right, optical microscope image of a classical interdigitated electrode (IDE) coated with aligned fibers [10].", "texts": [ " Due to such an extreme versatility, electrospinning is expected to have several strategic advantages, relative to other nanotechnologies, in the fabrication of advanced tools for diagnostics and health monitoring. Obviously, due to the wide and ever increasing production of scientific papers 388 Electrospun Materials for Tissue Engineering and Biomedical Applications in this topic, this chapter is not surely exhaustive, but it likes to enhance some main strategies that researchers are pursuing to create more and more innovative and high-performance support systems to human health (Fig. 17.2). The use of blood as a diagnostic medium is a routine and prevalent medical practice, not only in providing alerts about general health status but also in furnishing detailed information on individual diseases based on specific constituent biomarkers. A crucial issue arising in diagnosing health status of individuals is that in some pathology, symptoms emerge when the disease is in an advanced stage, and possible treatments can be ineffective or heavily invasive for humans, relative to early diagnosing" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0000431_amr.566.197-Figure5-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0000431_amr.566.197-Figure5-1.png", "caption": "Fig. 5 Polymer bearing race tested at 800N load, 2.1 \u00d710 5 cycles: (a) race overview, (b) flaking failure detail.", "texts": [ " The maximum tensile stress occurs at the edge of contact circle, where cracks initiate. The contact circle radius is similar to the radius curvature of the observed cracks and the cracks grow diagonally in two directions outwards the contact circle under contact surface [7]. The artificially introduced surface holes were observed after testing under the load of 800N, 200N, 100N, and 50N. The observation results were divided into three groups; holes with no cracks, holes with cracks, and pitting failure. Table 2 shows the statistics for each group. Fig. 5 shows image of failures detected on a component race tested under a load of 800N. Fig. 5 (a) is an overview photo of the race and Fig. 5 (b) is a flaking failure detail example. The surface cracks grew in the rolling direction, and in a close distance from one another. Fig. 6 shows a cracking occurred in a specimen tested at 200N. Two cracks were observed at the artificial hole: one at the periphery of the contact track. It starts at the bottom of the artificial hole or surface and grow in the rolling direction. It is hence concluded that the cracking does not follow the \u2018wedge effect\u2019 model.The maximum depth of the crack is 0.54mm" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0000679_20120403-3-de-3010.00065-Figure2-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0000679_20120403-3-de-3010.00065-Figure2-1.png", "caption": "Fig. 2. ALMAV Operational regimes", "texts": [ " Likewise, we study the transition stage providing the modeling and control strategies. The dynamics models of the rotorcrafts are deduced using the Newton-Euler approach. In terms of control algorithms, we are using a hierarchical control law. The ultimate objective of the project is to launch a vehicle about 500 meters with a height of 100 meters, at this point the vehicle collects valuable visual information of the target zone through a pointing-downwards camera. To achieve this goal, let us consider two rotorcrafts MAVs, a GLMAV (Fig. 1) and the ALMAV (Fig. 2). Despite that some mechanical similarities are evident in the vehicles design, the actuators nature and distribution separates the flight control profile. 2 The term \u201dremote\u201d is employed to denote a zone that is beyond of the aerial range of a mini rotorcraft 2.1 GLMAV The GLMAV is a rotary-wing MAV based on a coaxial rotor system, that is to say a dual motor driving two counterrotating blades, of which only one is fully controlled by a swashplate. During the ballistic phase, (BP) the rotors are folded to fit in the projectile shell", " It has a smaller curvature radius because the axial velocity is decreasing, due to the lift generated by the rotors. Fig. 3 depicts the different elements of the GLMAV. 2.2 ALMAV The ALMAV is also a rotary-wing vehicle based on a dual counter-rotating rotors. However, unlike the GLMAV the ALMAV lacks of swash-plate, instead, it incorporates control surfaces (aerodynamic-based torques) to drive its attitude. The design of this vehicle obeys to a specific launching application; the MAV is launched from an air shuttle carrier (Fig. 2) to carry out missions in distant zones from the launching site. Once the vehicle has reached the target zone (hostile location, contaminated area, crowds conflicts, disaster assessment, search & rescue, etc.), the flying robot performs hovering flight to inspect the surrounding environment acquiring and transmitting vital information through an onboard camera to a ground base station (Fig. 4). CESCIT 2012 3-5 April 2012. W\u00fcrzburg, Germany 261 Table 1 summarizes the advantages and drawbacks of both MAVs: The avionics is shared by both MAV configurations" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0001261_robot.2010.5509623-Figure8-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0001261_robot.2010.5509623-Figure8-1.png", "caption": "Fig. 8. Dynamic positioning with environmental disturbances.", "texts": [ " They are functions of the relative angle, \u03b3R, between the wind and platform direction, and are taken from tables. \u03c1w is the density of air in kg/m3, AT and AL are the transverse and lateral projected areas in m2, and L is the overall length of the platform in m. VR is the relative wind speed (Fig. 7), and is given in knots, see [13] and [14]. We also impose a sea current with velocity (in m/s) shown in Fig 7. In order to improve the performance of the controller so as to counterbalance environmental disturbances, we activate the integral part of the controller setting kix = kix = kiy = 0.01. In Fig. 8, the dynamic positioning performance of the controller is illustrated against the environmental disturbances. Fig. 9 (a, b, c, d, e, and f) shows the thrusts of the jets and the corresponding angles. The linear and angular velocities are depicted in Fig. 10 (a, c, and e,) while in Fig. 10 (b, d, and f) we see the position and orientation variables. Again, despite the disturbances and the actuators constraints, the platform is stabilized within the required limits. For the implementation of the above described control system, we utilize three GPS receivers, and two antennas, see Fig" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0000867_physreve.81.022301-Figure1-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0000867_physreve.81.022301-Figure1-1.png", "caption": "FIG. 1. Bottom of a slide line for a two-dimensional granular model defining the Mohr circle and pair in relation to the local tangential FT and normal FN forces between the grains as proposed by Eber 5 .", "texts": [ " 6 where, as in 5 , a local friction angle can be defined by the ratio of the local tangential and normal forces. A simple 2D model can give us insight into the relation between the local force fabric or configuration of forces between grains and the macroscopic behavior of the pile. In Refs. 5,6 , the authors related the normal and tangential forces applied to a plane in the material macroscopic description to the normal and tangential forces between grains composing that plane. Following Eber 5 , we can relate and as depicted in Fig. 1 on a particular plane, to the local grain-grain forces by m = cos + sin cos \u2212 sin , 1 where m is the local microscopic friction, , as defined in Fig. 1, is the angle between the normal force between grains, and is the range of angles over which can vary. The macroscopic friction is given by M = / =tan M, so we can write m = tan m = tan M cos + sin cos \u2212 tan M sin . 2 Solving for tan M, we arrive at tan M = tan m \u2212 tan 1 + tan tan m = tan m \u2212 . 3 As can be seen from the previous equation, the angle is referred to the principal plane that defines and so that such an angle depends on M itself. Referring to Fig. 2, we can derive the relation tan M = tan 2 3 m \u2212 + /4 " ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0002200_978-3-319-51222-8-Figure1.41-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0002200_978-3-319-51222-8-Figure1.41-1.png", "caption": "Fig. 1.41 Application of theReduction principle, a and b moment distributions and classical Mohr, c different paths with the same result, and d Reduction principle", "texts": [ " When we isolate this term on one side, then the minus sign disappears M \u00b7 (tan \u03d5l + tan \u03d5r) = M \u00b7 1 = . . . (1.281) In most books, this is written as M \u00b7 \u0394\u03d5 = M \u00b7 1 = . . . (1.282) because for small angles is tan \u03d5 \u03d5 but some authors go so far as to write \u0394\u03d5 with the dimension of rad or degree which is not correct. M \u00b7 \u0394\u03d5 is an expression 72 1 Basics of exterior work. When you rotate the beam end by 45\u25e6, then the work done by M is not M \u00b7 45\u25e6 but M \u00b7 tan 45\u25e6. The Reduction principle is a particular variant of Mohr\u2019s equation. Mohr would calculate the horizontal displacement ui of the frame in Fig. 1.41 by applying a unit 1.33 Reduction Principle 73 point load X1 = 1 in the direction of ui and then evaluate the integral 1 \u00b7 ui = \u2211 e \u222b le 0 ( MM1 EI + N N1 EA ) dx (1.283) by integrating over all frame elements. But according to theReduction principle, it suffices to apply the point loadX1 = 1 to a substructure of the original structure where the term substructure normally means any statically determinate system contained in the original system, for example, the single post in Fig. 1.41d. This surprising result is better understood when we realize that the node that carries the load can be approached from different starting points, see Fig. 1.41c, and that the sum of the horizontal displacements on each path must be the same, must be ui. This means we can compute ui by integrating, for example, only over the end post in Fig. 1.41d 1 \u00b7 ui = \u222b l 0 ( MM1 EI + N (N1 = 0) EA ) dx . (1.284) The Reduction principle essentially says that the Dirac delta, the X1, need not be applied to the original system but that it can be any substructure that is \u2018contained\u2019 in the original system, see Fig. 1.41d. Contained means that in the transition to the substructure nodes of the original structure can be released but no nodes may be fixed additionally. In the language of mathematics, this means that the Dirichlet boundary may shrink but it may not grow, [2, p. 149]. The admissible substructures are typically the structures which you would eventually choose in the force method. The single post is such a substructure. The Reduction principle is a clever application of Green\u2019s first identity. The sum of all the single identities of the frame elements is zero \u2211 e G (u, u1)e = 0 (1" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0002302_b978-0-12-803581-8.10351-0-Figure7-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0002302_b978-0-12-803581-8.10351-0-Figure7-1.png", "caption": "Fig. 7 Chassis features. Reproduced from Weidner, L.R., Radford, D.W., Fitzhorn, P.A., 2003. A multi-shell assembly approach applied to monocoque chassis design. In: SAE 2002 Transactions, Journal of Passenger Cars \u2013 Mechanical Systems, p. 2486.", "texts": [ " Thus, further validating the potential of the top/bottom, left/right monocoque design approach, which like the pressure vessel design, maximizes the overlap area for enhanced structural bonding. Many features were desired in the monocoque design development, in addition to high torsional rigidity. Some of the noteworthy features to be incorporated included structural sidepods, to improve both stiffness and driver side impact safety, and strakes ahead of the sidepods, which were to aid airflow to the cooling system, the underbody aero-features and yield an added degree of side impact safety, as indicated in Fig. 7. The concepts of spatial complexity are useful in any design for manufacture process. The shape complexity classification for a one piece, complete chassis, as shown in Fig. 7, is U7.7 This is the most complex classification for a manufactured part, meaning that it is undercut-U and irregular-7. For comparison, the idealized chassis (Fig. 1) has half the shape complexity rating, T4 (tube, closed one end). Undercut, or reentrant, geometries are very difficult to create using cost-competitive composites tooling and layup techniques. Features in the top (such as the desired single piece cockpit rim stiffener) and side details of the chassis (strakes) drove the design to the use of a top, bottom, and sides", " The left and right shells incorporate a stepped \u201cjoggle joint\u201d geometry to facilitate preliminary assembly and the integration of the inside facesheet of the cored sandwich structure. Using a joggle joint only in conjunction with the inside facesheet of the side shells and lapping the underbelly of the monocoque with the bottom part creates a joint with high strength and stiffness in an area that is prone to impacts from the road and debris (Fig. 12). Sidepods add side impact protection for the driver. However, there remains a vulnerable area just in front of the sidepod, and behind the front tire. Using strakes and structural sidepods, as shown in Fig. 7, the driver\u2019s knees and hands are better protected, especially from impact with other cars or sharp objects on or near the course. Side impact protection in this region was mandated in Formula 1 in 2000. The incorporation of this strake as an integral portion of the chassis would have been a severe complication in a chassis tooled in only top and bottom halves. However, with the multi-shell design, the strakes can be readily molded, directly into the left/right shells. A cross-section view of the lower strake is seen in Fig" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0000261_iros.2010.5651258-Figure2-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0000261_iros.2010.5651258-Figure2-1.png", "caption": "Fig. 2. Telescopic-legged rimless wheel that can climb up slopes", "texts": [ " To solve this problem, the author proposed a method for generating a gait that guarantees to overcome the potential barrier by asymmetrizing the impact posture as shown in Fig. 1 (b). The primary purpose of this method is to tilt or shift the robot\u2019s center of mass (CoM) forward for overcoming the potential barrier at mid-stance easily. The easiest way to asymmetrize the impact posture is to extend the stance leg during stance phases using the prismatic joints or knee joints. The author numerically investigated the validity of the proposed method using a telescopic-legged rimless wheel model shown in Fig. 2 (a), and performed parameter study in [6]. We showed the F. Asano is with the School of Information Science, Japan Advanced Institute of Science and Technology, 923-1292 Ishikawa, Japan fasano@jaist.ac.jp possibility of stable gait generation on level ground using the model, and extended it to climbing up a steep slope as shown in Fig. 2 (b). The simulation results implied that the generated level gaits are remarkably high-speed and the robot would jump because of the overly rapid motion. Through our investigations, the need of ankle brake has been indicated. The importance of forward-bending impact posture has been also discussed in several related works [2][3][4]. On the other hand, the author has wondered about the meaning of anterior-posterior asymmetry of human foot. As the authors have shown, such as in virtual passive dynamic walking, the stance leg of limit cycle walkers must be driven forward during stance phases to generate an energy-efficient level gait without creating negative input power [5][7]" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0000392_iros.2013.6696761-Figure3-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0000392_iros.2013.6696761-Figure3-1.png", "caption": "Fig. 3. suspension focus", "texts": [ " The suspension consists of a linear spring that is free to move along the y body axis. The rear and front springs have a non-compressed length d0, spring constant k, and compressed lengths d1 and d2, respectively. The force applied by a typical spring is thus: Pz = (d0\u2212d)k; d < d0. (8) Note that the force applied by the spring is state dependent since d is a function of the vehicle orientation \u03b8 . This and the assumption of mass-less wheels allows us to compute an expression that relates the normal force Fn to the traction force Ft . Figure 3 shows the forces acting on a typical wheel: the ground forces consist of the normal and traction forces Fn and Ft ; the suspension applies on the vehicle the forces Pz and Px; consequently, the forces applied on the axle are Fn, Ft , Pz and Px. Since the wheels are assumed mass-less, the axle forces must satisfy equilibrium: Pz +Fnn \u00b7 e2 +Ftt \u00b7 e2 = 0 (9) Px +Fnn \u00b7 e1 +Ftt \u00b7 e1 = 0, (10) where e1 and e2 are the unit vectors in the the suspension coordinates, as shown in Figure 3. We thus obtained two equations (9-10) in 3 unknowns: Px,Fn,Ft (Pz is state dependent and hence known). Since we wish to compute the traction force Ft , we first solve for Fn using (9), then substitute it back into the equations of motion (3). Substituting (8) in (9) and solving for Fn yields: Fn =\u2212aFt +bPz. (11) where a = t \u00b7 e2 n \u00b7 e2 (12) b = 1 n \u00b7 e2 . (13) For a given Pz, equation (11) expresses Fn as a linear function of Ft . This implies that the ground force F \u2208 R 2 must lie on a straight line that crosses the friction cone, as shown schematically in Figure 4" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0000397_00368791111140495-Figure2-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0000397_00368791111140495-Figure2-1.png", "caption": "Figure 2 Rigid rotor model", "texts": [ " With the assumptions of compressible, isothermal laminar flow situated in bearing clearance and perfect gas for air, the non-dimensional Reynolds equation of this air film which is derived from Navier-Stokes and continuity equations can be expressed in the two-dimensional Cartesian coordinates system as shown: \u203a \u203au h3\u203a P2 \u203au \u00fe D L 2 \u203a \u203a z h3\u203a P2 \u203a z \u00bc2L \u203a \u203au \u00f0 P h\u00de\u00fe4L \u203a\u00f0 P h\u00de \u203at \u00f01\u00de where D and L are bearing diameter and length, P and h are non-dimensional pressure and thickness of film, u and z are angular and axial coordinates of bearing, respectively, t \u00bc vt is non-dimensional time, and L is bearing number. If the journal center leave its equilibrium position \u00f0eo;fo\u00de due to small perturbation, that whirls with radial and tangential components represented by Re\u00f0epest\u00de and Re\u00f0eofpe st\u00de, respectively. This small journal whirl can induce small perturbations of the film pressure and film thickness which also can be decomposed into static and dynamic parts. As shown in Figure 2, the non-dimensional film thickness and pressure distributions can be expressed by: h \u00bc ho \u00fe 1pe stcos u\u00fe 1ofpe stsin u \u00f02\u00de and: P \u00bc Po \u00fe 1pe st P1 \u00fe 1ofpe st Pf \u00f03\u00de where ho is the static film thickness and defined by: ho\u00f0u; z\u00de \u00bc 1 \u00fe 1ocos\u00f0u2 fo\u00de; \u00f04\u00de Inherent restriction on stability of rotor-aerostatic bearing system Cheng-Hsien Chen et al. Industrial Lubrication and Tribology Volume 63 \u00b7 Number 4 \u00b7 2011 \u00b7 277\u2013292 1o \u00bc eo=c and fo is eccentric ratio and attitude angle of journal center in static equilibrium, 1p \u00bc ep=c is perturbed eccentric ratio and fp is perturbed angle" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0002610_gt2017-64959-Figure1-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0002610_gt2017-64959-Figure1-1.png", "caption": "Figure 1. BOUNDARY CONDITIONS AND DETAILED SCHEMATICS OF THE PARAMETERS", "texts": [ " The variable parameters are those values which are manipulated in order to obtain the optimum shape. The constants are the values already defined by the limitation of manufacturing and also the constraints of the nature of the problem such as the constraints extracted from fluid dynamic analysis. A section of the disc and blade is selected and due to the symmetry nature of the problem half of this section is modeled. The boundary condition for the model is the symmetry boundary condition for the section lines which prevent displacement in angular direction, see Fig. 1. The centrifugal body loading is applied on the whole model by introducing rotational speed \u03c9 and defining density of the material used. Also, having the weight of the blade, its center of mass radial distance and rotational speed provide us the centrifugal load due to the weight of the blade. As is illustrated in Fig. 1, this force is applied by a uniform pressure proportional to the force and the section area in which blade will attach to the root. The parametric model uses 11 values to be constructed. There are two main guidance lines A and B which provide a canal in which the tooth is located, Fig.1. The curvatures C1 and C2 are two main curvatures on the disc and are tangent to the guidance lines. By changing the radial location of point A and B and also by changing the corresponding angles the location of the curvatures will change. The curvature C2b for the blade is considered as one millimeter less in radius comparing to C2 of the disc so it is not an independent variable. The parameterization of the design means that different sets of parameters provide different shape for the blade and disc joint" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0000035_amr.181-182.361-Figure1-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0000035_amr.181-182.361-Figure1-1.png", "caption": "Fig. 1 Basic arc tooth profile of existing gear pump", "texts": [ " The contact strength, fatigue resistance and pitting resistance of the double arc gear are stronger than those of the traditional involute gear. The flow pulsation and noise can be reduced after the double circular tooth profile is used in the gear pumps. High pressure can be achieved by the sealing of the tooth top and tooth side to solve the leakage problem. The tooth side sealing is completed by a floating lateral plate. This is a mature technology. The tooth top sealing is completed by the chamber scraping of top teeth. Currently, the tooth profile of the double arc gear pump in the market is an arc curve, which is shown in Fig. 1.Theoretically, there is the line contact between the curve and pump body round, but a gap which is the cause of no high pressure exists actually. There are two solutions: (1) the pump body is cast and gears are interference fitted to achieve the chamber scraping. (2) There is no tooth top width of arc gear because the addendum circle is processed by hob. The machining errors of the hob can not guarantee that the tooth tip arcs are in one circle. The largest tooth may achieve the chamber scraping while other teeth can not. There is still a leakage problem. By turning the teeth top along the broken line shown in Fig. 1, an addendum circle as the involute gear pump is formed and the problem is also solved. All rights reserved. No part of contents of this paper may be reproduced or transmitted in any form or by any means without the written permission of Trans Tech Publications, www.ttp.net. (ID: 129.93.16.3, University of Nebraska-Lincoln, Lincoln, USA-03/04/15,06:29:14) The Selection of Asymmetric Double Arc Tooth Profile. In section 1.1, the arc profile has many advantages. However, displacement is reduced because of the smaller gear tooth and excessive removal will affect the continuity of transmission" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0000394_09596518jsce982-Figure3-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0000394_09596518jsce982-Figure3-1.png", "caption": "Fig. 3 Top view of excavator\u2019s work space", "texts": [ " Through such diagrams and given that the excavator\u2019s boom\u2013arm pair may be rotated in the range 0u to 360u, the maximum theoretical total soil volume that could be removed for a certain position of an excavator can be easily calculated. In order to define the optimal initial position of the excavator with respect to the dig geometry, it is important to define a means to represent its performance in a quantitative way. Within this context, a solution to the problem of locating the excavator positioning point(s) that provide removal of the maximum amount of soil will be presented. When seen from a top view (Fig. 3), an excavator\u2019s work space is part of a cycle with its centre E coinciding with the rotation centre of excavator\u2019s carriage, while its radius RE is equal to the maximum digging reach of the bucket end on ground (Fig. 2). Critical points are defined as all points that lie on the intersection of the excavator\u2019s workspace outer limit (bucket leftmost end) and the dig\u2019s upper surface. The excavator\u2019s workspace outer limit represents the extreme case scenario for reaching dig points, as these lie on the edge of its working envelope formed by a totally extended boom\u2013arm pair. An escape and unloading space zone is always necessary and this zone is illustrated in Fig. 3. Angle hE denotes the orientation of excavator\u2019s longitudinal axis with respect to zero-angle line (dash-dot line in Fig. 3). Given that there is no interference of dig contour to excavator\u2019s work space, if Vmax is the maximum volume of soil that could be removed, then this volume is analogous to rotation angle hmax~ hL{hR\u00f0 \u00de and may be easily calculated (or approximated) for any excavator by means of its work space diagram (Fig. 2). Additionally, under the assumption that the bottom surface of the dig \u2013 after its excavation \u2013 does not present any altimetrical irregularities, the problem of volumetric calculations may be reduced to a 2D problem of surface and, eventually, angle calculations" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0003045_eptc.2017.8277474-Figure2-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0003045_eptc.2017.8277474-Figure2-1.png", "caption": "Fig. 2. The structural schematic diagram of", "texts": [ " Firstly, the dynamic model based on Lagrange is established; Then, the full closed-loop control system is adopted; Lastly, Aiming at motion platform, the synchronous planning control method and the integral decoupling gantry control method are used respectively. [4] The synchronization error of the two control modes is compared by experiments. The results show that synchronization error of the latter is smaller, which can improve the control performance and the accuracy of synchronous motion. 2. Dynamic modeling of motion platform 978-1-5386-3042-6/17/$31.00 \u00a92017 IEEE 2017 19th Electronics Packaging Technology Conference2 In this paper, the structure of the gantry is shown in Fig. 2. The platform is composed of three axes (Y1 axis, Y2 axis and X axis). The X axis is a crossbeam. The stator of the linear motor in the X direction is installed on the cross beam. The mover is connected with the worktable to realize the movement in the X direction. The crossbeam is driven by two parallel linear motors to realize the movement in the Y direction. [5] gantry motion platform According to Fig. 2, m1 is the mass of the X directional crossbeam. m2 is the mass of the slide table. The corresponding positions of the motor on the Y1 axis, the Y2 axis and the X axis are y1, y2 and x respectively. \u0398 is the slight rotation angle of the cross beam caused by the non synchronization of the double-linear motor. O is the barycenter of the crossbeam, and o\u2019 is the barycenter of the slide table. L is the length of the crossbeam, o\u2019 is the width of the crossbeam, and e is the vertical distance between the barycenter of the slide table and the barycenter of the cross beam" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0001248_epe.2013.6634388-Figure1-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0001248_epe.2013.6634388-Figure1-1.png", "caption": "Fig. 1: Structure and specifications of magnetic gear.", "texts": [ " However, they have some problems including vibration, acoustic noise, and maintenance concerns. In addition, the mechanical gears require lubricant to reduce the friction. On the other hand, magnetic gears can transmit torque without mechanical contact. Therefore, they have low vibration and acoustic noise, and no lubricant in comparison with conventional mechanical gears. Various types of magnetic gears have been introduced in previous papers [1]\u2013[3]. Among them, a planetary type magnetic gear [4] has attracted interest recently. Fig. 1 shows a basic structure of a planetary type magnetic gear used in the consideration. The magnetic gear consists of an inner and outer rotors with surface\u2013mounted permanent magnet (SPM), and ferromagnetic stationary parts which are called pole pieces. It works as a gear when the pole pieces modulate the magnet fluxes. Transmission torque density of the planetary type magnetic gear is higher than the other types because all the magnets of the inner and outer rotors contribute to generate and transmit torque [5], [6]", " In this paper, first, the torque characteristics, core loss, eddy current loss in permanent magnets, and the efficiency of the magnetic gear are calculated by three\u2013dimensional finite element analysis (3D\u2013FEA). Next, no\u2013load and load test results of a prototype magnetic gear manufactured on the basis of the FEA results are indicated and compared to the ones obtained from 3D\u2013FEA. Finally, based on the FEA and experimental results, an efficiency-improved magnetic gear is manufactured and demonstrated. The maximum efficiency of the improved gear achieves up to over 90%. The magnetic gear shown in Fig. 1 has the SPM rotors. The pole\u2013pairs of the inner and outer rotors are 3 and 31, respectively. Hence, the gear ratio is 1: 10.333 given by the ratio of the inner and outer pole\u2013pairs [5]. The pole pieces are placed between the inner and outer rotors. The number of pole pieces is 34 given by the sum of numbers of the inner and outer pole\u2013pairs. The material of the permanent magnet is sintered Nd\u2013Fe\u2013B of which the remanence Br is 1.25 T and the coercivity Hc is 975 kA/m, respectively. Core material of the pole pieces and the rotor back yokes are non\u2013oriented silicon steel with the thickness of 0", " Table I shows the calculated losses and efficiency of the magnetic gear when the gear operates as a reduction gear and a load torque is 10 N\u2022m which represents 80% of the pull out torque. In this calculation, the mechanical loss is neglected. It is understood that the magnetic gear has high efficiency of 96%. Hence, it is expected that the efficiency of a prototype magnetic gear reaches over 90% even if the mechanical and other losses are actually existed. On the basis of the above results, a prototype magnetic gear was manufactured. The structure of the prototype gear has almost the same specifications shown in Fig. 1, but the stack length of the pole pieces is changed from 10 mm to 16 mm as shown in Fig. 6, in order to be easily fixed them with the gear housing. Fig. 7 shows the general view of the experimental setup. The prototype magnetic gear operates as a reduction gear on this system. The rotational speed of the inner rotor is regulated an arbitrary speed by the servomotor. The load torque is controlled by the hysteresis brake. Fig. 8 shows the input rotational speed versus output rotational speed at no load" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0003666_iemdc.2019.8785290-Figure2-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0003666_iemdc.2019.8785290-Figure2-1.png", "caption": "Fig. 2. Fringing air-gap flux density and virtual conductor in air-gap center used to model the field effect [1]", "texts": [ " This paper proposes one building block on the path towards an alternative, less computationally expensive and yet accurate way of calculating clamping plate losses, with the goal of ultimately allowing their calculation to be incorporated into a more comprehensive tool chain for rapid dimensioning of new generators. Traxler-Samek [1] did some pioneering work in this field, proposing the use of line conductors placed in the air-gap center flush with the stator and rotor core faces to represent the fringing air-gap flux density as seen in fig. 2. The stator and rotor cores are modeled as one infinitely permeable half volume using the method of image currents. Although being computationally inexpensive, this approach has one major drawback: Any reaction of the field to the highly permeable clamping plate is disregarded. This paper proposes to calculate the clamping plate field in open-circuit operation using a Schwarz\u2013Christoffel mapping. The next section is going to present the numerical approach to Schwarz\u2013Christoffel mappings chosen for this work, followed by sections covering the field calculation and flux sink concept" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0002302_b978-0-12-803581-8.10351-0-Figure40-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0002302_b978-0-12-803581-8.10351-0-Figure40-1.png", "caption": "Fig. 40 Top clip interface bracket.", "texts": [], "surrounding_texts": [ "The mono shock wedge is designed to react suspension loads from the spring/damper unit and support the suspension box rear angle wall. The design utilizes the same U-shape cross-section as the intermediate bulkhead as seen in Fig. 29. The wedge will be bonded to the chassis using the two angular bonding flanges and to the main bulkhead with the two vertical bonding flanges shown in Fig. 37(a). The wedge design and bonding flanges allow for transfer of the load from the spring/damper unit to the chassis and then into the main bulkhead and the main structure of the chassis as shown in Fig. 37(b). The layup of the wedge follows that of the intermediate bulkhead, [745, 0, 0\u201390, 90, 745, 0, 0\u201390, 90, 745, 0, 0\u201390, 90], from the outer surface to the inner surface. Ungrouped plies indicate the use of unidirectional prepreg, while grouped plies indicate the woven material." ] }, { "image_filename": "designv11_62_0003514_978-3-030-20131-9_324-Figure65-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0003514_978-3-030-20131-9_324-Figure65-1.png", "caption": "Fig. 65. Schematic diagram of beaded cool pad weaving device.", "texts": [], "surrounding_texts": [ "In chapter 2, three methods of automatic knitting of bead mat are put forward. Among them, Warp and Weft Automatic Weaving Method is the most simple, but the cost is higher and its stability is poor. The lock stitch sewing weaving method is rea lly good, but there exists three difficulties applied in the machine: The cross-sectional size of the hook must be smaller than the size of the bead hole as the hook needle needs to completely pass through the bead, which makes it difficult to hook the string smoothly. It is difficult to guarantee the parallelism between the cross-section of the loop ring and the end of the bead of the first row in steps 4-6. When weaving a larger size beaded cooling pad, threading the string into the longer aligned transverse bead holes becomes hard. So compared with the other two methods, single-line straight-through method, which has good stability and high knitting efficiency and is easy to be realized on the machine, is the focus of the discussion below. Based on this method, an automatic weaving device capable of weaving a beaded cool pad is proposed and designed. 4.1 Feeding Device Design Before weaving the beaded cooling pad, since all the beading arrangements are disordered, it is necessary to design a device that puts the beads into the weaving state in an orderly manner, which is called feeding device. Referring to the feeding mechanism of the firecracker weaving [14-16], Fig.14 shows the schematic diagram of the designed bead feeding mechanism device. Before the device runs, all the beads are placed in the hopper and the two guiding wheels. A small number of longitudinally are placed in a horizontal arrangement. When the two guide wheels rotate in opposite directions, the beads in the hopper will be putted into the guide groove and conveyed to the front of the beading device in an orderly manner. In order to avoid a rigid collision between the feeding device and the ball transported device, the end of the guiding groove is made by a material with better elasticity. 4.2 Design and Working Principle of the Beaded Pad Weaving Device Fig. 15 demonstrated the beaded pad weaving device. Since the figure is only for explaining the movement process of the beaded pad weaving device, the feeding device is not shown in the figure. And there are 7 motors in this device. Control motor A controls the movement of the threading device. Control motor B controls the movement of the movable line of the downlink line. Control motor C controls the movement of the movable stroke switch. Control motor D controls the rotation of the output port and the braided port. Linear motor E S. Ouyang et al.2544 drives the up-line feed ball push block movement. Linear motor F pushes the braided beaded cool pad unit into the braided port. Linear motor G drives the linear motion of the downlink line feed bead block. Before the device is operated, the downlink threading is first performed. After the downlink threading is finished, the string is installed into the beaded pad weaving device. The downlink line with heavy beads at the end is wrapped around the fixed pulley mounted on the frame. Then it passes through the downlink line movable seat, the through hole, the downlink end sleeve, the bead and the braided port successively, to reach the uplink line. And the uplink line is directly connected to the needle of the threading device, the first string of beads are moved to the corresponding position on the weaving port, as shown in Fig.16. After the string installation is completed, the motor A is manually controlled to make the driving roller be located between the two trapezoidal blocks on the movable seat rail of the threading device. The manually controlled linear motor E is to drive the uplink line to send th e beads push block moves, which is external bead conveyed from the feeding device. It causes the holes axis of the bead to coincide with the needle axis of the bead threading device and the up-line bead push block is in the beading state. The specific work ing process of the device is as follows: Method Research and Mechanism Design of Automatic Weaving\u2026 2545 Step 1: Under the drive of the control motor A, the bead threading device moves to the right. When the driving roller moves in the second trapezoidal block on the movable seat rail, the location clamping position of the needle will be changed. As a result, the uplink line smoothly penetrates an external bead provided by the uplink line bead transported device, as shown in Fig.17. Step 2: The bead threading device continues to move to the right. When the movable seat contacts the movable travel switch, the uplink line is just tightened. At this time, some of the motor operation will change as follows: Control motor A reversed means the bead threading device starts to move to the left. Controlling motor B rotated forward means the downlink line movable seat moves to the right for a suitable distance, providing two beads re quired for the next unit downlink line weaving. After that, controlling motor B stops. Linear motor F runs, the weaving beads are pushed into the weaving port and the output port, then moves back to the initial position. Linear motor E reversely drives, the pushing block of uplink line feeding bead returns to the initial position, and is on out feeding condition, as shown in Fig.18. S. Ouyang et al.2546 Step 3: Under the driving of the control motor A, the bead threading device moves to the left. When the threading device contacts the fixed stroke switch, some of the motor operation will change as follows: Linear motor G forward drives, the pushing block of the downlink line feeding bead pushes a shared bead to make the hole axis of the shared bead coincide with the needle axis. Control motor A rotated forward means the bead threading device starts to move to the right. Linear motor E drives forward, the push block of the uplink line feeding bead is in the feeding state. Control motor C rotated forward, the movable travel switch moves to the left for a suitable distance exactly equal to the length of the string required to weave every bead pad unit, as illustrated in Fig. 19. Step 4: When the pushing block of the downlink line in the feeding state, the control motor D is drive to rotate the output port and the braided port counterclockwise by 180\u00b0. Step 5: Under the driving of the control motor A, the bead threading device moves to the right. When the driving roller moves in the first trapezoidal block on the movable seat rail, the location and clamping position of needle will be changed to make the uplink line successfully penetrate into a shared bead provided by the downlink line feeding device. Step 6: The linear motor E is reversely driven to make the pushing block of the downlink feeding bead out of the feeding state, as shown in Fig. 20. Method Research and Mechanism Design of Automatic Weaving\u2026 2547 Step 7: Repeat the actions from steps 1 to 6 until the end of the weaving task. Step 8: When the weaving process is finished, each motor is controlled by software programming to bring the device into an initial state." ] }, { "image_filename": "designv11_62_0000182_carpi.2012.6473352-Figure1-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0000182_carpi.2012.6473352-Figure1-1.png", "caption": "Fig. 1. The safety space model around power tower and the discrete viewing regions", "texts": [ " In these EVAs represented by dotted red circle areas in Fig. 2, the UAQ would clearly observe the corresponding facilities - blue circle areas in Fig. 2. In order to get a better observation perspective, we select the center of each EVA as target viewing regions (TVRs) (the green block in Fig. 3(b)), which are the important regions mentioned in chapter I. UAQ has to visit all these TVRs completely during the inspection. Hereafter, only one side of the power tower is considered as the half hollow cylinder shows in Fig. 1. The UAQ starts inspection from S (the bottom of tower), and visits TVRs one by one. After visited all TVRs, the UAQ moves to G, shown in Fig. 3. It is one complete episode of the inspection flight. Reinforcement learning[9] is defined as a Markov decision process (MDP) with three parameters: the state space S , the action A and the reward R . Mathematically, this theory can be presented as following: an agent experiences a state transition from 1 tt ss by executing the action Aa t in state Ss t , and it will get the reward Rr t 1 at the same time" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0002302_b978-0-12-803581-8.10351-0-Figure39-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0002302_b978-0-12-803581-8.10351-0-Figure39-1.png", "caption": "Fig. 39 Top clip interface bracket.", "texts": [], "surrounding_texts": [ "The mono shock wedge is designed to react suspension loads from the spring/damper unit and support the suspension box rear angle wall. The design utilizes the same U-shape cross-section as the intermediate bulkhead as seen in Fig. 29. The wedge will be bonded to the chassis using the two angular bonding flanges and to the main bulkhead with the two vertical bonding flanges shown in Fig. 37(a). The wedge design and bonding flanges allow for transfer of the load from the spring/damper unit to the chassis and then into the main bulkhead and the main structure of the chassis as shown in Fig. 37(b). The layup of the wedge follows that of the intermediate bulkhead, [745, 0, 0\u201390, 90, 745, 0, 0\u201390, 90, 745, 0, 0\u201390, 90], from the outer surface to the inner surface. Ungrouped plies indicate the use of unidirectional prepreg, while grouped plies indicate the woven material." ] }, { "image_filename": "designv11_62_0003660_iemdc.2019.8785146-Figure2-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0003660_iemdc.2019.8785146-Figure2-1.png", "caption": "Fig. 2 Example of eccentricity", "texts": [ " The concept of a multi-three-phase PMSM is shown in Fig. 1. Since the PMSM has three groups of coils whose powers are provided by three individual inverters, we can individually control the amplitude and phase of a current at each group and reduce losses and vibration. Eccentricity and roundness degradation generate a magnetic unbalance in the magnetic gap between rotors and stators, causing low order harmonics of the electromagnetic force and cogging torque [4]-[5]. Examples of eccentricity are shown in Fig. 2. Static eccentricity means that the stator center and rotational center are different, as in Fig. 2(a), and dynamic eccentricity means that the rotational center and rotor shape center are different, as in Fig. 2(b). Fig. 2 demonstrates that the magnetic gap length differs depending on the spatial position of the stator and that the electromagnetic force between the rotor and stator is unbalanced. Since multi-three-phase PMSMs can individually control the phase current of each group, reducing the magnetic unbalance is possible in the magnetic gap; however, to control the phase current of each group, we must detect the magnetic unbalance. Conventional studies have detected static eccentricity by the current and voltage waveforms at the stator\u2019s coil in salient-pole synchronous motors [6]-[7]" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0000902_eeeic.2013.6549559-Figure1-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0000902_eeeic.2013.6549559-Figure1-1.png", "caption": "Figure 1. Brush making contact with commutator bars", "texts": [ " Finite Element Analysis (FEA)) and the circuit differential equations of the machine (numerical integration techniques). Whereas the differential equations of the motor circuits proposed are generally similar, the main difference in the approaches to model universal motor commutation appears to be the treatment of the brush contact voltage or rather the brush resistance. In this paper the motor circuits are derived using the brush model proposed by the authors in an earlier paper [11]. The brush model proposed was created for universal motors with a brush width (see Fig. 1) to commutator segment width ratio \u03b6 0 [1;2] as the most common ratios in practice are in this range. The resistances used highly depend on the rotor angle \u03b8. Hence, the cases of one coil shorted and two coils shorted by one brush simultaneously as well as an arcing phase can be accounted for with this model. Arcing is the main source of radio frequency electromagnetic noise [16] emitted by the universal motor. The noise is emitted as radiation and gridbound high frequency interference. Due to the high amount of mains operated universal machines (home appliances, power tools) the impact of the interference on other electronic devices and local grid quality is despite measures for suppression not inconsiderable" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0000130_j.physc.2013.04.021-Figure3-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0000130_j.physc.2013.04.021-Figure3-1.png", "caption": "Fig. 3. Prototype of axial flux induction motor.", "texts": [ " And because the directions of torques are arranged, the torques can be treated as internal force. Furthermore, a linear induction motor is adopted to provide propulsive force. In this way, four HTS AFIMs and one linear induction motor compose a minimum system of maglev train. Because HTS AFIMs will be moving along the secondary, the secondary without slots is better than the secondary with slots. In addition, the HTS AFIMs will be immersed directly independently in liquid nitrogen to satisfy the working condition. One prototype of AFIM is manufactured and shown in Fig. 3. 3. Force characteristic analysis In order to achieve a correct and stable result, analytic method and finite element method have been used to model the motor. The primary of the motor is simplified to a linear lamination and the current is assumed to distribute in a sinusoidal form on the top surface of the stator. Then, the secondary can be simplified to a linear lamination too, as shown in Fig. 4. It is assumed that the laminations are composed of a linear material obeying the steady-state Maxwell\u2019s equations in the moving conductor and the linear constitutive laws" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0002302_b978-0-12-803581-8.10351-0-Figure36-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0002302_b978-0-12-803581-8.10351-0-Figure36-1.png", "caption": "Fig. 36 Full assembly of the roll hoop in the main bulkhead.", "texts": [ " The inner flange follows the roll hoop contour with an offset to allow the front half of the bulkhead shell inner flange to fit between the hoop and back half, thus bonding the back and the front together forming a single hollow bulkhead with the steel tube structure inside. The front half is configured in a u-shape to combine the two halves as stated above with the inner flange, and the outer bonding flange is then tucked between the hoop and the chassis. The front vertical wall is the same as the back. This allows for a full 360 degree bonding of the roll hoop, and provides sufficient bonding to the chassis as shown in Fig. 35 and the full assembly of the roll hoop into the bulkhead shown in Fig. 36. The main bulkhead reinforces the chassis cross-section from distorting under load and, in addition to providing rollover protection, supports localized loads imposed by the suspension. The inner bonding flanges of the composite halves closely follow the roll hoop contour to allow the largest opening possible for the driver in the chassis. Following the roll hoop inner perimeter allows for only a 0.05000 loss in the perimeter due to the composite bulkhead thickness and required adhesive. The layup for each half of the composite main bulkhead is [745, 0, 90, 745] through the majority of the bulkhead surface area, including the outer bonding flanges, where 0 degree is considered across the chassis", " The layout was marked and the aluminum plate with blue marking paint and machined on a vertical mil. The angled edges were cut with a 20 degree tapered end mill, and the additional angle variation was accounted for by tramming the head. The rounded ends were done on the grinder. Also, an additional round was added to the top surface edges, to provide an acceptable radius between the bulkhead face and the bonding flange. The main bulkhead requires a pair of tools to create the matching bulkhead sections that sandwich the forward tubular roll hoop as shown in the previous section in Fig. 36. Conceptually, these aluminum tools were produced in a fashion very similar to the front bulkhead tooling of the previous section. The rear shell has a flange in only a single direction from the plate surface, while the front composite laminate has to have inner perimeter flanges to mate with the rear bulkhead shell and an outer perimeter flange to add further bonding area against the inside monocoque surface. The perimeters of the three aluminum plates were waterjet cut. Angled cuts were made at 19" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0001493_ines.2010.5483842-Figure2-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0001493_ines.2010.5483842-Figure2-1.png", "caption": "Figure 2. Stator-field- oriented space-phasor diagram of the synchronous motor with adjustable exciting field and controlled stator flux, operating at unity PF for different values of the load torque.", "texts": [ " DOUBLE-FIELD-ORIENTED VECTOR CONTROL If the exciting current of the Ex-SyM is controlled, there are three degrees of freedom from the control consideration, i.e. three control loops can be formed: two magnetical (for flux and power factor control) and a mechanical one (for speed). The stator flux and the speed are directly controlled, by means of PI controllers, and the PF is indirectly controlled; it is commanded to its maximum value by canceling the stator-field-oriented longitudinal armature reaction ( 0=ssdi \u03bb ), as is shown in Fig. 2. The stator-field orientation ensures the most suitable procedure for the control of the power factor and the resultant stator flux. At unity (maximum) power factor the stator is absorbing a minimum current that may be considered optimal from energetic point of view. The power factor is maximum, if the stator voltage and stator current are in phase. Consequently, the stator-flux vector \u03a8s results perpendicular onto the stator-current space phasor is1,2,3. each phasor corresponding to another load torque value", " In order to achieve the unity power factor, the top of the stator resultant flux phasor \u03a8s must be situated on the semi-circle, which diameter is the space phasor of the exciting flux \u03a8me. given by the exciting winding current in the air gap. This statement results from the stator-voltage equation, written with space phasors, if the derivative of this flux is zero, i.e. it is constant because it is controlled or the machine is in steady state operation mode. In the rotorposition oriented diagram, if the top of \u03a8s is inside the circle, the synchronous motor will operate with leading (capacitive) current [2]. The diagram from Fig. 2 also shows, that not only the stator voltage, but also the exciting current must be adjusted, according to the load torque, in order to keep constant both the resultant stator field (by means of ims the corresponding magnetizing current) and the power factor at unity value [2]. \u2013 238 \u2013 \u2013 239 \u2013 The perpendicularity between the resultant stator-flux and the armature current space phasors can be achieved if the resultant stator-field-oriented longitudinal armature reaction is cancelled. In the proposed control structure, presented in Fig", " It has also two components in the stator-field-oriented reference frame: ( ) ( )\u23aa\u23a9 \u23aa \u23a8 \u23a7 +=\u2212 \u2212+=\u2212 ssqsqseq Ref msssdsdsed ii iii \u03bb\u03bb \u03bb\u03bb \u03c3 \u03c3 1 1 (5) Taking into account the fact that the stator-fieldoriented longitudinal armature reaction is cancelled, Eq. (5) will be simplified as follows: ( )\u23aa\u23a9 \u23aa \u23a8 \u23a7 +=\u2212 \u2212=\u2212 ssqsqseq Ref mssed ii ii \u03bb\u03bb \u03bb \u03c31 (6) The exciting winding is supplied from a DC-to-DC chopper working controlled by a feed-forward carrierwave voltage PWM. With the before presented control structure simulations were performed for validation purpose followed by the experimental implementation. V. SIMULATION RESULTS Based on structure from Fig. 2 simulations were performed in Matlab-Simulink\u00ae environment. The salient pole Ex-SyM rated data are: UsN = 380 V, IsN = 1.52 A, PN = 800 W, fN=50 Hz, nN = 1500 [rpm], cos\u03c6= 0.8 (capacitive). After the starting process the motor will run at the rated speed value corresponding to a frequency of 50 Hz. At t= 1 s a speed reversal is applied, under the full rated load. The mechanical load has reactive character and it is linearly speed-dependent. \u2013 240 \u2013 in stator-flux-oriented coordinate frame. \u2013 241 \u2013 The simulation results shows that the proposed control structure from Fig. 2 is a viable one with improved performances with respect to the conventional VC systems [2]. The results show a good performance of the drive also in transient operation at starting, and also at speed reversal (Fig. 1). The power factor is maximum also during the speed reversal, when the drive is in regenerative running for a short period of time, as is shown in Fig. 5. Unity power factor is realized by canceling the stator-fieldoriented longitudinal armature reaction, as in Fig. 7. VI. CONCLUSIONS The presented control structure uses two types of orientations: resultant stator-field and exciting-field, i" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0002971_s00773-017-0487-1-Figure1-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0002971_s00773-017-0487-1-Figure1-1.png", "caption": "Fig. 1 Coordinate systems", "texts": [ " Moreover, VPMM (Vertical Planer Motion Mechanism) tests are also carried out for obtaining the maneuvering coefficient of the used model. The obtained mathematical model is validated by free running model experiment carried out in the tank, and course keeping stability or controllability are discussed. Furthermore, to clarify the advantage of thrust vectoring system, the authors apply the optimal control theory to the obtained mathematical model, and further improvement of controllability due to the additional fins is discussed. The coordinate systems utilized here are defined in Fig.\u00a01. Now, ov \u2212 xv yv zv is a space\u2011fixed coordinate system, whereas o \u2212 x y z is a ship\u2011fixed coordinate system with its origin fixed at the center of gravity. In these figures, positive direction of moments and thrust vectoring system are also shown. Here Tx , Ty and Tz are forces due to thrust vectoring system in thruster\u2011fixed coordinate system. The definitions of non\u2011dimensionalization in this paper are shown in Table\u00a01. Notice that the non\u2011dimensionalization uses cubic root of the ship volume as a non\u2011dimensionaliza\u2011 tion factor having dimension of the length" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0000389_detc2013-12966-Figure2-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0000389_detc2013-12966-Figure2-1.png", "caption": "Figure 2. HOT AND COLD SPOT REGIONS", "texts": [ " Then solve for the static equilibrium position of the journal and predict the steady state orbit at the journal for an assumed imbalance [14-16]. Next, the hot and cold spot regions for the journal must be determined and the temperature distribution on the journal can then be obtained. Finally, evaluate the rotor thermal bend and calculate the thermal imbalance resulting from the average hot spot temperature gradient using the following equation [ 1, 2, 7, 11, 12 , 13] ddt YMU (2) where Md is the overhung mass. In order to locate hot and cold spot regions, three methods have been used (Figure 2): Method #1: hot spot is defined as the closest point to the bearing sleeve at time instant t=0. In this method, it is also assumed that Oj0, Oj and hot spot H all lie on the same straight line; Method #2: hot spot is defined as the closest point to the bearing sleeve at time instant t=tm. In this configuration, the hot spot H is located on the straight line Ob and Oj; Method #3: hot spot is defined as the point which has the maximum temperature in the orbit. The corresponding equations can be derived to obtain the hot spot and cold spot based on the specific method [7]" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0001680_00207179.2013.813646-Figure1-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0001680_00207179.2013.813646-Figure1-1.png", "caption": "Figure 1. Schematic diagram of the surface vessel showing the control point (CP) and the centre-of-gravity (CG).", "texts": [ " (2010) Nonlinear Sliding mode Outdoor Fahimi and Kleeck (2012) Nonlinear Sliding mode Outdoor Siramdasu and Fahimi (2013) Nonlinear Model predictive Outdoor Notes: LQG = linear quadratic Gaussian; PI = proportional integral; PID = proportional integral derivative. m22v\u0307 + m11ur + d22v = F sin \u03b1, (2) I33r\u0307 + (m22 \u2212 m11)uv + d33r = \u2212(F sin \u03b1)L, (3) where u, v and r are the surge, sway and yaw speeds, respectively, F is the propeller force, \u03b1 is the rudder angle and L is the distance of the centre-of-gravity (CG) from the propeller location on the vessel (Figure 1). The propeller is hinged to the vessel such that its axis can yaw with respect to the vessel. Here, the term rudder angle \u03b1 (as shown in Figure 1) refers to the angle that the axis of a hinged propeller makes with the longitudinal axes of the vessel. The vector F = [F cos \u03b1, F sin \u03b1,\u2212FL sin \u03b1]T contains the longitudinal and lateral components of the propeller force (F), and the propeller\u2019s moment about the vessel\u2019s CG, which has a longitudinal offset L with the propeller location on the vessel (Figure 1). The propeller force is modelled as F = AnC, where A is a constant, C is a dimensionless parameter and n \u2265 0 is the propeller speed. The propeller force can be written as F = [AU1, AU2,\u2212ALU2]T, (4) where U1 = nC cos \u03b1 and U2 = nC sin \u03b1 (5) are selected as the mathematical control inputs. A robotic vessel and a control box are designed and built for experimentation purposes (Figure 2). The specifications of the vessel are: length 80 cm; width 70 cm; mass 8 kg; and diameter of each propeller 4.5 cm" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0002230_ut.2017.7890291-Figure2-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0002230_ut.2017.7890291-Figure2-1.png", "caption": "Fig. 2. Testbed with 3DOF manipulator", "texts": [ " (10), ( ) 2 3 ,2 1 2, ,int1 1 1 ( ) 2 2 i i mid p zmp x p i ii x \u03b8 \u03b8 \u03b8= \u2212 \u039e \u0398 = \u2212 \u0394 + W W (10) The solution in null space can be obtained by conducting a redundancy analysis using Gradient Projection Method(GPM) for the proposed performance index as shown in Eq. (11). 1 2 ; , , , T N \u03b3 \u03ba \u03b8 \u03b8 \u03b8 = \u2212 \u22c5\u2207\u039e \u2202\u039e \u2202\u039e \u2202\u039e\u2207\u039e = \u2202 \u2202 \u2202 (11) where \u03ba and \u2207\u039e represent a positive gradient constant and the gradient vector of performance index respectively. 4. Experiments 4.1 System composition To implement the proposed ZMP algorithm and redundancy resolution method, a testbed composed of a floating vehicle which represent simple underwater vehicle attached with a redundant manipulator was developed. As shown in Fig. 2(b), this manipulator has three DOF (pitch-pitch-pitch) pitching redundant motion. To drive the each joint of the manipulator, 3 motors and its drivers are installed in the control box which can be waterproof. And the driving forces for each joint are transferred by pulley-timing belt mechanism from the motor to each joint. To measure the pitch-angle and pitch-acceleration, attitude and heading reference system(AHRS) is installed in the control box. The floating vehicle is composed of buoyancy materials and equipped with the developed manipulator" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0002495_978-3-319-56802-7_41-Figure2-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0002495_978-3-319-56802-7_41-Figure2-1.png", "caption": "Fig. 2 Kinematic model of a car with two trailers", "texts": [ " Definition 3 The configuration x is a kinematic singularity if the filtration of \u0394 at x is not constant in a neighbourhood of x. If additionally \u0394 is regular, i.e. G (x) has full rank, the configuration x is a non-holonomic kinematic singularity. The difference to Definition 2 is that G may have full rank at a non-holonomic singularity. Only for non-holonomic systems can G (x) be regular but the filtration \u0394 not be regular at x. Corollary 1 The set of non-holonomic singularities, denoted \u03a3nh, is closed in V p. Example 2 Consider the carwith two trailers in Fig. 2. The p = 7 systemcoordinates are (x, y, \u03b81, \u03b82, \u03d5, \u03b1) \u2208 V 7 = R 2 \u00d7 T 5. The kinematic control system is \u239b \u239c \u239c \u239c \u239c \u239c \u239c \u239c \u239c \u239d x\u0307 y\u0307 \u03b8\u0307 \u03b8\u03071 \u03b8\u03072 \u03d5\u0307 \u03b1\u0307 \u239e \u239f \u239f \u239f \u239f \u239f \u239f \u239f \u239f \u23a0 = \u239b \u239c \u239c \u239c \u239c \u239c \u239c \u239c \u239c \u239d R cos\u03d5 cos \u03b8 R cos\u03d5 sin \u03b8 R L sin \u03d5 R L sin \u03d5 \u2212 R L1 cos\u03d5 sin \u03b81 R L1 cos\u03d5 sin \u03b81 \u2212 R L2 cos\u03d5 cos \u03b81 sin \u03b82 0 1 \u239e \u239f \u239f \u239f \u239f \u239f \u239f \u239f \u239f \u23a0 u1 + \u239b \u239c \u239c \u239c \u239c \u239c \u239c \u239d 0 0 0 0 1 0 \u239e \u239f \u239f \u239f \u239f \u239f \u239f \u23a0 u2. (10) The accessibility distribution \u0394 = span (g1, g2) is regular, i.e. has constant dimension for all x \u2208 V 7. Its filtration terminateswith the accessibility algebra\u0394 = R 7, and the system is thus accessible and controllable" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0000263_016918610x551791-Figure5-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0000263_016918610x551791-Figure5-1.png", "caption": "Figure 5. Model 2. Rotate the vector \u2212\u2212\u2192 EM in the direction of movement of the object by the leading angle \u03c6. Translate the vector to the center of the virtual circle by parallel translation. Set the length of the vector to r . Create the target velocity vector \u2212\u2192\u0307 E\u2217 from \u2212\u2192 E to \u2212\u2192 T with a constant velocity v.", "texts": [ " In addition, the centrifugal force and the Coriolis force affect the object when the object is considered to be in a rotating coordinate system. For the case in which the centrifugal force is sufficiently larger than gravity, gravity can be neglected. In this case, a circle is the optimal trajectory for the endeffector to control an object that is rotating with uniform circular motion. Therefore, we used the attractor described in Section 2. We added a compensator to this attractor so that the robot would be able to control the rotational axis of the object to target rotational axis O (Fig. 5). Target position \u2212\u2192 T and target velocity vector \u2212\u2192\u0307 E\u2217 of the end-effector are defined as: (i) Define a virtual circle with radius r around the target rotational axis O . (ii) Define vector \u2212\u2212\u2212\u2192 EM from the position of the end-effector \u2212\u2192 E to the position of object \u2212\u2192 M . (iii) Rotate vector \u2212\u2212\u2212\u2192 EM in the direction of the movement of the object by the leading angle \u03c6. (iv) Define \u2212\u2192 T as the position r \u2212\u2212\u2212\u2192 EM/|\u2212\u2212\u2212\u2192 EM| from the axis of the virtual circle O . (v) Control the end-effector towards position \u2212\u2192 T at a constant velocity v" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0003613_012066-Figure3-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0003613_012066-Figure3-1.png", "caption": "Figure 3. Contour of von-Mises stress on faultless coil spring given the maximum load.", "texts": [], "surrounding_texts": [ "Transport vehicles require a good suspension system to dampen vibration, swings and shocks received as they travel along bumpy, hollow, and uneven roads [1]. These conditions are very uncomfortable and may cause accidents. The suspension is also expected to hold the load during some common vehicle maneuvers such as acceleration, braking or deflection while on the road [2]. The coil spring is one of the main components for dampening vibrations and shocks to the load so as to provide comfort and security while the vehicle is in motion [3]. Depending on the condition of their application, coil springs often sustain fatigue failure. This indicates that the tension received below by the coil spring from the maximum stress of the material while sustaining a dynamic load causes fatigue failure [4-8]. The yield strength of the material is also a criterion of failure. Components of automotive suspension must be changed with a traveling distance of 73,500 km, or every five years [9]. The fault of 13.18 % of 24.2 million vehicle tests was recorded [10]. With the development of computing technology, the numerical analysis method has become particularly suitable for use because it will increase the calculation efficiency, the cost-effectiveness as well as save time. Various numerical analysis methods are widely available, but the finite element analysis (FEA) has proven to be reliable in solving problems in the field of continuum mechanics [11]." ] }, { "image_filename": "designv11_62_0002131_10402004.2017.1285970-Figure14-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0002131_10402004.2017.1285970-Figure14-1.png", "caption": "Figure 14: Power loss (W) in the journal bearing for the Ester sample. Over 70 % of the total losses happen in the region where the viscosity is higher", "texts": [], "surrounding_texts": [ "1. Schirru, M., Mills, R., Dwyer-Joyce, R., Smith, O., and Sutton, M. (2015). Viscosity Measurement in a Lubricant Film Using an Ultrasonically Resonating Matching Layer. Tribology Letters, 60(3) pp. 1-11. 2. Mason, W.P., Baker W.O., McSkimin, H.J. et al. (1948). Measurement of shear elasticity and viscosity of liquids at ultrasonic frequencies. Physical Review. Vol. 75(6) pp. 936- 946 3. Bujard, M.R., (1989) Method of measuring the dynamic viscosity of a viscous fluid utilizing acoustic transducer. US Patent Number 04862384 4. Emmert, S.W., (1986) Apparatus and method for determining the viscosity of a fluid sample. US Patent Number 4721874 5. Greenwood, M.S,. and Bamberger, J.A., (2002) Measurement of viscosity and shear wave velocity of a liquid or slurry for on-line process control. Ultrasonics Vol. 39 pp. 623-630 6. Lamb, J., (1967) Physical properties of fluid lubricants: rheological and viscoelastic behaviour. Proceedings of the Institution of Mechanical Engineers, Conference Proceedings. Vol. 182 pp. 293-310 7. Barlow, A.J., and Lamb, J., (1959) The visco-elastic behaviour of lubricating oils under cyclic shearing stress. Proc. R. Soc. Lond. Vol 253 pp. 52-69 8. Dowson, D. G. R. A. V., Higginson, G. R., and Whitaker, A. V., (1962) \"Elasto- hydrodynamic lubrication: a survey of isothermal solutions.\" Journal of Mechanical Engineering Science 4.2 pp. 121-126. ACCEPTED MANUSCRIPT 27 9. Schirru, M., and Dwyer-Joyce, R.S., (2015) A model for the reflection of shear ultrasonic waves at a thin liquid film and its application to viscometry in journal bearings. Proc IMechE Part J: J Engineering Tribology, [online] accessible from: http://pij.sagepub.com/ 10. Collin, R.E., (1955) Theory and design of wide-band multisection quarter-wave transformers. Proceedings of the IRE Vol.36 pp. 621-629 11. Emmert, S.W., (1986) Apparatus and method for determining the viscosity of a fluid sample. US Patent Number 4721874 12. Bair, S., and Winer, W.O., (1980) Some observations on the relationship between lubricant mechanical and dielectric transitions under pressure, Trans. ASME, Journal of Lubrication Tech., 102, 2, pp. 229-235 13. Hutton, J.F., (1967) Viscoelastic relaxation spectra of lubricating oils and their component fractions. Proc. Roy. Soc. London Mathematical and Physical Sciences. Vol. 304, 1476, pp.65-80 14. Johnson, K. L., & Tevaarwerk, J. L. (1977). Shear behaviour of elastohydrodynamic oil films. Proceedings of the Royal Society of London A: Mathematical, Physical and Engineering Sciences Vol. 356, No. 1685, pp. 215-236 15. Kinsler, E., et al. (2000) Fundamentals of acoustics , 4 th Edition, Wiley, New York 16. Harrison, G., and Barlow, A.J.. (1981) Dynamic Viscosity Measurement. Methods of experimental physics 19 pp. 137-178. 17. Stachowiack, G.W., and Batchelor, A.W., (2001) Engineering Tribology. 2 nd ed. Butterworth Heinemann, pp. 204-210. ACCEPTED MANUSCRIPT 28 18. Raimondi, A.A., and Boyd, J., (1958) A solution for the finite journal bearing and its application to analysis and design. ASLE Transactions. Vol. 1 pp. 159-209 19. Juvinall, R.C., and Kurt, M.M., (2006) Fundamentals of machine component design. 5 th ed., John Wiley & Sons ACCEPTED MANUSCRIPT 32 ACCEPTED MANUSCRIPT 33 ACCEPTED MANUSCRIPT 34 ACCEPTED MANUSCRIPT 35 ACCEPTED MANUSCRIPT 36 ACCEPTED MANUSCRIPT 37 ACCEPTED MANUSCRIPT 38 ACCEPTED MANUSCRIPT 40 ACCEPTED MANUSCRIPT 41 ACCEPTED MANUSCRIPT 42 ACCEPTED MANUSCRIPT 43 ACCEPTED MANUSCRIPT 44 ACCEPTED MANUSCRIPT 45 ACCEPTED MANUSCRIPT 46 ACCEPTED MANUSCRIPT 47" ] }, { "image_filename": "designv11_62_0001423_robio.2011.6181436-Figure1-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0001423_robio.2011.6181436-Figure1-1.png", "caption": "Fig. 1. Two objects grasped by a multifingered hand.", "texts": [ " (A7) The relationship between the finger configuration displacement and reaction force is replaced with a 3D orthogonal spring model. In Assumption (A7), the spring stiffness of the j-th finger is denoted by 33],,diag[ zjyjxjj kkkK . The direction of the stiffness is fixed along the three axes of the initial fingertip coordinate frame. In Assumption (A3), the spring is compressed in the initial configuration to generate the initial fingertip force 0fjjj K xf . The virtual spring model is attached to the fingertip as shown in Fig. 1. When the object configuration is displaced due to an external disturbance, each finger rolls or slides on the object surface, the spring length changes, and the stored potential energy also changes. B. Nomenclature We use the following coordinate frames: b is a base coordinate frame. c is a coordinate frame between two objects. oi is an object coordinate frame fixed in the i-th object. fj is a finger coordinate frame fixed in the j-th finger, where the x- and y-axis are tangent to the normal direction, and the z-axis is along the contact normal direction", " The body velocity of C in L is calculated by 00 0 0][ 0 0 )]([)]([:)(\u02c6 31 1 , C T CC CC CC CC C L C Lb CL M MK MK MT MT tTtTtV , (3) where 22 CM , 22 CK and 21 CT are the metric tensor, curvature and torsion of the body surface, respectively. In the next section, we use the above definition for matrices )( ciCci LciT , )( ojCoj LojT and )( fjCfj LfjT . III. DISPLACEMENT OF FINGER CONFIGURATION A. Configuration Displacement of Two Objects If each object is not detached and penetrated with another object as shown in Fig. 1, the configuration displacement of two objects is represented by the following 11 parameters. 11 21 ],,,,[: TT c T cc T c T cx , (4) where 6],[: TT c T cc x is the configuration of frame c , c is the spin angle between two objects, and 2 ci is the contact location on the i-th object. These parameters are illustrated in Fig. 2. The configuration displacement of each object, oi , is obtained from the following constraint: )]()][(][[ )](][)][(][[ cCci c cc bc bc b ciCci Lci Lci oi oioi boi boi b TTT TTTT " ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0002599_gt2017-64896-Figure3-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0002599_gt2017-64896-Figure3-1.png", "caption": "Figure 3 \u2013 AXIS-SYMMETRIC SMALL PUNCH TEST MODELED IN ANSYS.", "texts": [ " The repetitive calculations involved in iteratively determining the values can be very costly in terms of time and computing power. As such, several studies have made efforts to automate the process. Automation has previously been completed with curve fitting algorithms and neural networks which automatically vary material properties within a certain range [33, 39]. The model created for this study was originally based off of a proven model established by Campitelli et al. [37]. A 2D axissymmetric model was utilized within ANSYS to carry out the simulation, using a static structural analysis, as shown in Figure 3. The upper and lower dies were modeled as rigid, fixed structures. The sample and punch were meshed with quadrilateral elements, with refinements on the edge of the punch and in the center of the sample where the deformation occurs. A friction factor of 0.1 was applied between the specimen and each die, to keep the sample in place during deformation, simulating the clamping contact present in the experiment. Contact and friction of 0.05 was used between the punch and specimen, for contact between the polished punch and polished sample" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0001888_s10483-013-1684-7-Figure1-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0001888_s10483-013-1684-7-Figure1-1.png", "caption": "Fig. 1 Mechanical model of gear-pair system on deformable bearings", "texts": [ " The effect of spall on mesh stiffness is analyzed, and the model also considers the variant stiffness in the presence of spall. The approximate analytical solutions of primary resonance and internal resonance are obtained by the averaging method, then the different bifurcation characteristics caused by the evolvement of spall and transmission error can be revealed by employing the singularity theory for the two-state variable system. 2 Generation of dynamic model with spalling defect The geared rotor-bearing model investigated in the present study is shown in Fig. 1. In this model, friction forces at the mesh point are assumed to be negligible. Thus, the transverse vibrations along the directions of pressure line and the vibrations along the direction perpendicular to the pressure line are uncoupled. Bearings and shafts that support the gears are represented by equivalent damping and stiffness elements as shown in Fig. 1. The damping elements are characterized by linear viscous damp coefficients c1 and c2, and the stiffness elements are defined by k1 and k2. External radial preloads F1 and F2 are also applied to both the rolling element bearings. The model takes into account the so-called static transmission error, and both the stiffness kg(\u03c4) and the static transmission error e(\u03c4) can approximately be considered as time-periodic functions, and the fundamental frequency of both of the quantities equals the gear mesh frequency" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0000636_detc2013-13388-Figure1-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0000636_detc2013-13388-Figure1-1.png", "caption": "FIGURE 1. Schematic of tunnel gear driving system and coordinates.", "texts": [ " Given its structural and force complexity, the scope is limited to the ring gear\u2019s in-plane and in-extensional deflections and pinions\u2019 rigid deflections. Special attention is paid to the time-varying effect of the mesh stiffness on the instability behavior for the equally-spaced configurations. The main results are determined using Floque\u2019t theory and numerical simulations. 2 Copyright \u00a9 2013 by ASME Downloaded From: https://proceedings.asmedigitalcollection.asme.org on 01/06/2019 Terms of Use: http://www.asme.org/about-asme/terms-of-use 2. DYNAMIC EQUATIONS 2.1 Model Description Fig. 1 illustrates a tunnel gear driving system, where identical pinions are equally-spaced such that the angular separation between the adjacent ones is 2\u03c0/N, where N is the pinion number. The ring gear is rotating at a constant angular velocity \u2126. The coordinate o-r\u03b8z is an inertial frame. Without any loss of generality, the first pinion is attached at \u03c81=0, and thus the nth (n=1, 2, 3, \u2026 N) one is at \u03c8n=2\u03c0(n-1)/N. The notation krn represents the mesh springs between the nth pinion and the ring gear", " (2) Potential Energy of the Springs 1) Potential energy of the mesh springs Figure 3 illustrates the nth pinion-ring gear mesh. The relative deflection between the pinion and the ring gear may be written as r p psin cos sinn n n u u v u\u03b8 \u03b8\u03b4 \u03b2 \u03b2 \u03b2 \u03b8 \u2202 = \u2212 + + \u2202 (11) where \u03b2 is the gear pressure angle, and \u03b8 is constrained to the pinions\u2019 position angles ( n\u03c8 ). Using Eq. (11), the elastic mesh potential energy is ( )2 2 m r r0 1 1 2 N n n n n U k d \u03c0 \u03b4 \u03b4 \u03b8 \u03c8 \u03b8 = = \u2212 (12) r \u03d5 \u03a9 rv u\u03b8 pnv O n\u03c8 Base circle of pinion Base circle of ring Mesh direction \u03b2 \u03b2 pnu \u03b2 According to Fig. 1, the overall potential energy of the support springs on the pinions can be written in the inertial frame as ( )2 2 p p p0 1 1 2 N n n n n U k v d \u03c0 \u03b4 \u03b8 \u03c8 \u03b8 = = \u2212 (13) 2.5 Total Energies Equations (6) and (7) yield the total kinetic energy as ( ) 2 22 2 2 20 2 2 p 2 p p20 1 1 2 1 2 N n n n n n u u u u T AR d t t u J m v d r \u03c0 \u03b8 \u03b8 \u03b8 \u03b8 \u03c0 \u03c1 \u03a9 \u03a9 \u03b8 \u03b8 \u03b8\u03b8 \u03b4 \u03b8 \u03c8 \u03b8 = \u2202 \u2202 \u2202 \u2202 = + + + \u2202 \u2202 \u2202 \u2202\u2202 + + \u2212 (14) From Equations (10), (12) and (13), the total potential energy is ( ) ( ) 23 2 2 2 r r3 30 0 1 2 2 p p0 1 1 22 1 2 N n n n n N n n n n u uEI U d k d R k v d \u03c0 \u03c0\u03b8 \u03b8 \u03c0 \u03b8 \u03b4 \u03b4 \u03b8 \u03c8 \u03b8 \u03b8 \u03b8 \u03b4 \u03b8 \u03c8 \u03b8 = = \u2202 \u2202 = + + \u2212 \u2202 \u2202 + \u2212 (15) 4 Copyright \u00a9 2013 by ASME Downloaded From: https://proceedings" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0001223_03091902.2013.785608-Figure18-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0001223_03091902.2013.785608-Figure18-1.png", "caption": "Figure 18. COMSOL model depicting the pressure differences across the spirometer design at Time\u00bc 0 of the measurement.", "texts": [ "04 1 MV Resistors 200 8 0.04 2\u20135/160 1\u201315/160 Proto Board 100 69 0.69 2\u20135/160 1\u201315/160 Proto Board 100 69 0.69 10 10 3/400 Plywood 100 63 0.63 200 200 Pine 30.48 m 100 1 Large-Diameter Flat Washers 200 16 0.08 Wood Screws #12 200 400 52.8 0.132 Steel Machine Screw Hex Nut 400 24 0.06 Load Cells 2 compressive load cell (16375097) strain gauges from technoweighindia.com 200 780 3.9 Plastic Circuit Cover 100 88 0.88 Polyurethane 1150 oz 30 0.3 Total 9.432 Figure 16. Spirometer without mouthpiece. and an outlet of 0 kPa. Figure 18 shows the pressure at Time\u00bc 0 s, while Figure 19 shows the pressure at Time\u00bc 2 s. From the readings, it can be seen that a pressure drop occurs along the narrow cylinder, after the eccentric cone that can be easily detected by the sensor. The user interface for the spirometer was designed to properly engage both the operator and the patient using the device. This was done through a graph and fill bar as shown in Figure 19. The graph of flow rate helped the operator in evaluating spirometer read-out" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0002027_978-3-319-00479-2-Figure18-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0002027_978-3-319-00479-2-Figure18-1.png", "caption": "Fig. 18 FE-Simulation of impact against a block with the animal\u2019s mass, representing the worst case scenario for occupant acceleration load", "texts": [ " The evaluation of occupant safety had been conducted on the basis of further simulations using the finite element simulation model in combination of a multibody occupant model. For this, a validated MADYMO occupant model was used. For this study in a first step it was reviewed, whether the collision with the vehicle front represents a problem for the occupant. The worst case for acceleration load on the occupants due to an animal impact is, when the vehicle impacts the animal\u2019s body with full overlap, corresponding to a configuration in which the animal lies or sits and the whole mass of the animal is effective during the impact (Fig. 18). As a reference test the US-NCAP configuration was chosen because it can be assumed, that for the vehicle design or the design of the restraint systems this configuration is considered. Therefore a FE-simulation with the US-NCAP configuration was made to get the vehicle acceleration, which was used as input for the MADYMO occupant simulation. The setup for the MADYMO simulation was a sled with the mounted seat adopted with respect to the vehicle drawings and a validated HIII dummy (Fig. 19). Typically restraint systems and characteristics were used, but airbag triggering and restraint characteristics were not optimized" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0000484_omae2013-10496-Figure2-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0000484_omae2013-10496-Figure2-1.png", "caption": "Figure 2 - Coordinate systems", "texts": [ " So, robust differentiation techniques must also be applied, also derived by Levant [9]. In a previous paper [19], the HOSM was applied to the DPS with promising results, but only time-domain numerical simulation results were showed. The present paper aims to present small-scale experiments results that confirmed the good results of the controller. DPS is only concerned with the low-frequency horizontal motions of the vessel, that is, surge, sway and yaw. The motions of the vessels are expressed in two separate coordinate systems (Figure 2): one is the inertial system fixed to the Earth, OXYZ; and the other, O\u2019x1 x2 x6, is a vessel-fixed non-inertial reference frame. The origin for this system is the intersection of the midship section with the ship\u2019s longitudinal plane of symmetry. The axes for this system coincide with the principal axes of inertia of the vessel with respect to the origin. The motions along of the axes O\u2019x1, O\u2019x2 e O\u2019x6 are called surge, sway e yaw, respectively. Downloaded From: http://proceedings.asmedigitalcollection" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0000965_icat.2013.6684040-Figure1-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0000965_icat.2013.6684040-Figure1-1.png", "caption": "Figure 1. Key parameters of the model.", "texts": [], "surrounding_texts": [ "In the other words, any trajectory of the system starting in the invariant ellipsoid \u03b5x stays in it for all t >0.\nSimilarly, the output-invariant ellipsoid is\n 1y)CPC(y:Ry 1TTm y ,\nwhere P > 0 is a configuration matrix of the ellipsoid \u03b5x.\nSet of output-invariant ellipsoids allows us to estimate the impact of external disturbances to the system\u2019s output. Thus, selecting from the set of invariant ellipsoids \u03b5y the minimal one for some criterion, we will limit the impact of external influences to the system's output y(t).\nAs a criterion for selecting the minimum invariant ellipsoid it is proposed to use the objective function f(P)=tr(CPCT), that determines the sum of squares of the semi-axes of the output-invariant ellipsoid of the system (1).\nIn order to provide the desired degree of stability it is proposed to use the method based on the construction of an auxiliary polynomial with a prescribed degree of stability.\nIn [2] the theorem that describes an algorithm for constructing a polynomial with a prescribed degree of stability is proved:\nFor any nd \u2013 dimensional vector the degree of stability of the polynomial\nodd, is n if ),,s(~)),(as(\neven; is n if ),,s(~ ),s(\nd * 1d\nd * *\n\n \n \n \n 5\nnot less then a prescribed value > 0, and conversely, if the degree of stability of the polynomial \u0394(s) not less then a value > 0, than nd \u2013 dimensional vector exists and satisfies the identity \u0394(s) \u2261 \u0394*(s,), where\n \nd\n1i\n0 i 1 i 2* ),(as),(as),s(~ , 6\n]2/n[d d , ,2),(a 2 1i 1 i \n2 2i 2 1i 20 i ),(a , ,d,1i \n7\n,),(a 2 0d1d \n},,,...,,,,{ 0d2d1d22211211 . 8\nThus, specifying the arbitrary vector , using (5) \u2013 (8) we can construct the polynomial with the desired degree of stability.\nFor the construction of automatic control system it is proposed to modify the external disturbances compensation method [5] by combining it with the method of providing the desired degree of stability.\nThe developed algorithm is as follows:\n Following the method of compensation of external impacts solving the equation\n.0P,0HHPBBPAAP T1TT with respect to the matrix variable P, expressing it\u2019s components through and , and using obtained P(, ) to determine the coefficients of the controller k1, k2, k3, k4 as functions with parameters , using ratios\n .,PB,K Kxx,PBu 1T\n1T\n\n \n\n\n\n Using u = K(, )x obtaining the closed-loop system.\n Constructing the characteristic polynomial of the closed-loop system depending on ,\n .C,BKAsEdetsd \n Defining the desired degree of stability \u03b7 and construct the auxiliary polynomial \u0394*(s,).\n Equating the coefficients of the corresponding powers in characteristic and auxiliary polynomials obtain the system of equations with a set of parameters , as a solution.\n Solving the minimization problem\n)C),(CP(trmin T 0, \ntaking into account the additional restrictions for parameters , from the previous step.\n Using obtained , according to the formula\nu = YP-1x\nderive vector K of the controller\u2019s coefficients.\nIV. COMPUTER MODEL OF THE CONTROL SYSTEM Let us consider the mathematical model of the sea-going\nship:\nu ,\n),t(dhbaa ),t(dhbaa\n222221\n111211\n\n \n\n \n\n\n\n\n(9)\nHere is an angular velocity relative to the vertical axis, \u03c6 is a course (the turn to port side is considered positive), is a deviation angle of the vertical rudders, \u03b2 is a drift angle (angle between the velocity vector and longitudinal axis of the ship), u is a control, d(t) is a bounded exogenous disturbance:\ndT(t)d(t )\u2264 1, 0 \u2264 t < \u221e. (10)\nThe package MATLAB with the subsystem Simulink is one of the most effective tools to form and use in the researches computer models of dynamic systems.\nLet us consider the mathematical model (1) of the seagoing ship with displacement ton 6000.", "Its coefficients for a fixed velocity possess the following values:\na11 = -0.03408, a12 = 0.56, a21 = 0.015, a22 = -0.306,\nb1 = -0.0099, b2 = -0.00417, h1 = -0.0648, h2 = -0.0046,\nQ = diag( 0.2, 0.2, 0.2, 0.2).\nThe graph of the disturbance is represented on Fig. 2.\nLet\u2019s form Simulink-model of the control system for the marine ship, which scheme is represented on the Fig.3.\nV. THE RESULTS OF SYNTHESIS It is offered to search for the control as a controller with\nmathematical model\nu = k1 k2k3 \u03c6k4 (11)\nwhere k1, k2, k3, k4 are parameters need to be found those provide the desired dynamics of the closed-loop system.\nDeviation of the rudders and the its velocity turn (that is control) are constrained:\n\u2264 30o, u\u2264 3o/sec.\nOne of the effective approaches to the increasing of the digital system\u2019s performance consists in the search of such values of tuned parameters with fixed structure of the\nAs the result of this algorithm for given ship the controller with the coefficients\nk1 = 2.44, k2 = 66.6, k3 = -0.0384, k4 = 0.00133,\nwhich provides the desired dynamics, is obtained. At the same time all technical constraints are taking into account.\nFor the performance testing of the found controller let use it for automatic control of a sea-going ship under bounded exogenous disturbance.\nFig. 4 and Fig. 5 show the graphs of yaw and deviation angle of the vertical rudders.\nVI. CONCLUSION In the work the method of suppression of bounded disturbances based on using of invariant ellipsoids is stated. For the specific model of a marine ship was found the controller, provided the restriction of exogenous disturbances, taking into account all technical requirements.\nREFERENCES [1] Veremey E.I., Korchanov V.M., Pogozhev S.V., Computer modeling\nof control systems for marine movable objects. St.Petersburg: SPbSU, 2002. 370 p.\n[2] Veremey E.I. The synthesis of the laws of multipurpose control for marine movable objects// Gyroscopy and navigation \u2013 2009-\u2116 4. P. 3\u201314.\n[3] Lukomskiy U.A., Korchanov V.M. The control of the marine movable objects - St.Petersburg: Elmor, 1996. \u2013 320 p.\n[4] T.I. Fossen. Guidance and Control of Ocean Vehicles. John Wiley & Sons. New York, 1999.\n[5] Polyak B.T., Scherbakov P.S. Robust stability and control \u2013 \u041c.: Nauka, 2002.-303 p." ] }, { "image_filename": "designv11_62_0000543_0976-8580.93230-Figure1-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0000543_0976-8580.93230-Figure1-1.png", "caption": "Figure 1: The 2-DOF mass conveyor belt system", "texts": [ " The next is organized as follows: In Section 2, the mode coupling system subjected to Coulomb friction to be studied is introduced and modeled. In Section 3, the stability analysis of the system is carried out by using the Routh criterion. And, an analytical calculation method for the limit cycle amplitude based on the extended Harmonic Balance technique is presented in Section 4. The influence of physical parameters on the limit cycles is investigated in Section 5. Conclusion is made at last. 2. SYSTEM DESCRIPTION AND MODELING The mode coupling model usedin this work is a 2-DOF mass-conveyor belt system as shown in Figure 1. The conveyor belt moves with a constant velocity vo. The mass is held by two springs (k1 and k2) and two dampers (C1 and C2) at two directions (x1 and x2). Because of the external force Fe, the mass is always contacts with the belt. And Coulomb friction force exists at the contact surfaces. The origin form of this model is first presented by Hulten [6,7] to study the problem of drum brake squeal, in which only the stiffness is included. The system may be instable because of mode coupling. After that, Website: www", " But as we know, the Coulomb friction force has its own direction which depends on the relative sliding velocity between two contact surfaces. So such an assumption may be not sufficient for the real dynamic behavior of mechanical systems. In this article, therefore, to make up the shortage of above assumption, a complete Coulomb model considering direction is adopted Ffi = \u03bcFni sgn(vo - xi , (i = 1,2) (2) where, xi denotes the absolute velocity of the mass, and sgn(\u2022) is the sign function. Therefore, the motion equation of the nonlinear system shown in Figure 1 is m o o m x x c o o c x x k o o k + + 1 2 1 2 1 2 1 2 = \u2212 \u2212 \u2212 \u2212 + x x k x v x F k x v x o e o 1 2 2 2 1 1 1 2 2 2 2 sgn( ) sgn( ) 2 Fe (3) where, x1 and x2 present the absolute displacements of the mass in the directions of x1 and_x2, respectively. In such a case, the external force Fe will cause a static displacement of the mass. So when there are x x1 2 0= = and x x1 2 0= = the static equilibrium points of the system are: x F k x F k e e 1 0 2 1 2 0 2 2 2 1 2 1 2 1 2 1 = \u2212 + = \u2212 + ( ) ( ) , ( ) ( ) (4) Taking the static equilibrium points as a new origin of coordinate and inserting x x x1 1 1 0' = \u2212 and x x x2 2 2 0' = \u2212 into Equation (3), the following can be obtained as m o o m x x c o o c x x k + + 1 2 1 2 1 2 1 ' ' ' ' k k k x x k x x v xo 1 2 2 1 2 2 2 2 0 1 1 \u2212 = + \u2212 \u2212 \u2212 ' ' ' '( )[sgn( ) ] k x x v xo1 1 1 0 2 1( )[sgn( ) ]' '+ \u2212 \u2212 (5) where, x1 ' and x2 ' represent the displacements of the mass in the new coordinate, respectively" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0002045_iccsce.2012.6487214-Figure5-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0002045_iccsce.2012.6487214-Figure5-1.png", "caption": "Fig. 5: Experimental platform", "texts": [ " Figure 4 depicts the flow chart for the control program to perform the intended functionality. i) Project the state ahead ii) Project the error covariance ahead Predictor 1 i) Compute the Kalman gain ii) Update estimate with measurement zk = iii) Update the error covariance Corrector Initial estimates for and (n = resolution conversion bit) A custom-made experimental platform made from PVC pipes, and acrylic is used to hold the quadrocopter that has been assembled with controller board and other devices. The platform is designed to be smoothly oriented at any angle between 0 to 360\u00ba. Figure 5 shows a photograph of the platform. IV. RESULTS AND DISCUSSION Primarily, to evaluate the performances of the control program with Kalman\u2019s filter 1-D algorithm to filter out noise from accelerometer outputs as well as performing roll and pitch conversion, the value for Xpriori, Ppriori, Q and R must be initialized. The initial value for Xpriori is set 0 and basically, the value is not very essential since it is adapted during the operation. The similar thing also goes for Ppriori. However, it is important to set the Ppriori some value other than zero and possibly high enough to narrow down" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0000418_detc2011-47599-Figure3-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0000418_detc2011-47599-Figure3-1.png", "caption": "FIGURE 3. Screw notations of twists and reciprocal wrenches of the planar two-3R cable-driven closed chain", "texts": [ " 2, the following notations will be used for the twists and wrenches in the examples of this paper: \u2022 $\u0302k: j represents the unit screw of twist of jth joint in kth chain \u2022 $\u0302k: jr represents the unit screw of reciprocal wrench of jth joint in kth chain \u2022 $\u0302i: j\u2212h represents the unit screw of ith cable wrench, acting from jth link to hth link and j 6= h \u2022 $ewk: j represents the screw of external wrench acting on jth link in kth chain \u2022 $ep represents the screw of external wrench acting on end- effector \u2022 ti represents the tension scalar of ith cable Figure 3 presents the planar two-3R cable-driven closed chain that will be used as an example to carry out the proposed force-closure analysis based on reciprocal screw theory. Step 1: Determining the Twist and Reciprocal Wrench Screw Representations From Fig. 3, the unit screws for the pure rotation twists in each open chain, $\u0302k: j, are described with respect to the inertial frame as follows: $\u0302k: j = 0 0 1 ;rk: j\u00d7 0 0 1 ( j = 1,2,3;k = 1,2) (1) where rk: j is the global position vector of the jth joint centre in the kth chain. The unit screws of the various reciprocal wrenches in each open chain, $\u0302k: jr, j = 1,2,3, are described with respect to the 3 Copyright c\u00a9 2011 by ASME Downloaded From: http://proceedings.asmedigitalcollection.asme.org/ on 02/01/2016 Terms of Use: http://www", " As such, the cables will be routed on Chain 1 to provide torque to the active joints and achieve equilibrium (see Fig. 4). Also, all torques on the passive joints in Chain 2 due to external wrenches have to be reflected onto joints in Chain 1 using the relationship obtained in (6). $3:1-0t3 - Cases with active joints selected from different chains For cases where passive joints are chosen from both chains, the reflected torque is found using (6). For example, if the passive joints for the planar two-3R closed chain example in Fig. 3 are chosen to be Joints 1 and 2 on Chain 1, and Joint 3 on Chain 2, the reflected torques on the active joints, i.e., \u03c4r 1:3, \u03c4r 2:1 and \u03c4r 2:2, are determined as follows: \u03c41:1 \u03c41:2 \u03c4r 1:3 = (S1) T (S2) \u2212T \u03c4r 2:1 \u03c4r 2:2 \u03c42:3 (7) Solving the three simultaneous equations in (7) will result in the exact solutions for \u03c4r 1:3, \u03c4r 2:1 and \u03c4r 2:2, provided S2 is of full rank. Step 3: Determining the Joint Torques Required to Sustain the External Wrenches Based on Proposition 1.1, the external wrenches on each link can be uniquely expressed as a linear combination of the reciprocal wrenches of the respective chains" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0000234_icnsc.2013.6548848-FigureI-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0000234_icnsc.2013.6548848-FigureI-1.png", "caption": "Figure I. Lateral vehicle model", "texts": [], "surrounding_texts": [ "For describing the lateral dynamics of the vehicle, the two states lateral-drift and yaw motion model represented in ([) (figure 1) is the one commonly used [14]. jp = _1 (Fe! + FYI )-Ij/ mv e Vi = /z (afFYI - aJyr ) (\\) Where b is the sideslip angle and x is the yaw angle. Other parameters are visible in Figure \\. Most of the works consider the linear version of the model, in this paper, we propose to consider it from a nonlinear point of view. To reflect a more realistic driving behavior and to take into account some dangerous situations (saturated tire-road forces) the consideration of a nonlinear model is required. Since the nonlinearities of the vehicle model are located in the expressions of the lateral forces, considering the nonlinear expression of these last leads to a nonlinear lateral drift model of the vehicle. F; = D; sin (C; tan-I(B;(1- EJa; + tan-I(B; aJ)) (2) Where i = {f, r}. Stands for front and rear wheel respectively, Do Co Bi and Ei are parameters that depend on the tire characteristics. a f and a r are the front and rear sideslip angles of the tire respectively given by: af = -/3 - tan -l( a: l;/ cos( /3) J + Sf ar = -P + tan -I ( a: vi\" cos( fJ)) (3) bf is the steering angle of the front tire. It is well known that tire lateral force saturation occurs even at sideslip angle values, corresponding to rational driving situations included in the normal region and the pseudo-sliding one. The sideslip angles may be then approximated by a simplified expression as follows: a; . ar=-/3--Ij/+or v /3 a, . (4) a, = - +-Ij/ v" ] }, { "image_filename": "designv11_62_0002601_gt2017-64123-Figure9-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0002601_gt2017-64123-Figure9-1.png", "caption": "Figure 9. The parameterized solid air-gas channel sub-model", "texts": [ " The developed parameterized model contains a minimum of simplifications and is equivalent to models used in standard gas-dynamic engineering analysis and in determining strength characteristics of centrifugal compressors impellers. Therefore, the impeller structure obtained as a result of optimization will not require further structural redesign. Parameterized solid sub-models description A parameterized 3D solid model, which includes two sub-models, is developed to automate the optimization process. The solid model is shown in Figure 8. One submodel describes the air and includes an intake part, the airgas channel of the impeller, and the vaneless diffuser outlet (see Figure 9). The intake configuration is not modified during the optimization. The other sub-model describes a centrifugal compressor impeller with its blades and disc (see Figure 10). Each sub-model includes a cyclic-symmetric sector with one short and one full-sized blade. The sub-models are coordinated with each other, so a modification of one submodel leads to corresponding changes in the other submodel. The technological constraints (form separation, foundry equipment dismountability, etc.) and the prohibition of occurrence of singular structures (blade, fillet crossing, etc" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0000224_kem.504-506.1305-Figure6-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0000224_kem.504-506.1305-Figure6-1.png", "caption": "Figure 6: Tool wear VBB (\u00b5m) comparison analysis in the case of En variation and in the case of DOC variation.", "texts": [ " In this case the effect is more evident when En influence is considered. A final remark deductible from this forces analysis is that in machining of Inconel 718 low values of DOC and En should be used in order to obtain low cutting and thrust force components; the same observation can be done for the initial value of cutting force component at the beginning of the cutting operation (Fx,eng): low values of DOC and En lead to reduce initial cutting force components. Tool wear. The experimental results reported in Fig. 6 show some of the tool wear comparison analysis conducted with the aim to better understand the influence of the machining process conditions on the selected tool. In accordance with the international standard normative ISO 3685:1993 the average width of the flank wear land in the regular worn zone B, VBB (Fig. 3), were measured and compared at the different considered levels of En and DOC by optical microscope analysis. It is important to highlight that all the experiments were carried out in conditions of constant removal volume for each run (Vrim=27403mm 3 ). As concern the influence of En parameter, it can be noted in Fig. 6 as highest value for VB is registered for En=45\u00b0. The reason of this behavior is due to the thrust and the cutting force components since for En =45\u00b0 the thrust force component and the cutting force component have comparable values, leading to maximize tool wear in terms of VB. In Fig. 6 is also reported the influence of the DOC influence on tool wear parameters for fixed values of En. Particularly, it can be observed as tool wear increases with the increasing of the DOC. Process feasibility. By analyzing the results reported in Table 2 it was possible to execute an accurate analysis to understand which are the values of the DOC and En (for the fixed values of the cutting speed, feed rate and removal volume chosen in this work) more suitable for performing the cutting process on Inconel 718 such as the catastrophic tool rupture (as the one reported in Fig" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0003613_012066-Figure2-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0003613_012066-Figure2-1.png", "caption": "Figure 2. Shape of the defect formed on the surface of the coil spring: (a) defect up to 0.2 mm; (b) defect up to 3.5 mm.", "texts": [ "1088/1757-899X/523/1/012066 The coil spring was analyzed under different conditions of defects that occurred on the surface of the coil through a FEA. The study was conducted using geometry and similar dimensional values as well as the same material. The coil spring was analyzed under three different conditions, namely, a faultless coil spring, a coil spring with a defect of 0.2 mm, and a coil spring with a defect of 3.5 mm. The shapes of the defect that occurred on the surface of the coil springs are shown in Figure 2. . 3. Results And Discussion Figures 3 and 4 illustrate in detail the occurrence of the von-Mises stress on the faultless coil spring using the FEA, Figures 5 and 6 on the coil spring with a defect 0.2 mm, and Figures 7 and 8 on the coil spring with a defect of 3.5 mm. Based on the results of FEA, the shear stress value of the ASTM A227 coil spring obtained was approaching the allowable shear stress, but its value exceeded the yield strength. Thus, coil springs with such defects undergo plastic deformation and are susceptible to failure" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0003613_012066-Figure5-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0003613_012066-Figure5-1.png", "caption": "Figure 5. Contour of von-Mises stress on a coil spring with a defect of 0.2 mm given the maximum load.", "texts": [], "surrounding_texts": [ "Transport vehicles require a good suspension system to dampen vibration, swings and shocks received as they travel along bumpy, hollow, and uneven roads [1]. These conditions are very uncomfortable and may cause accidents. The suspension is also expected to hold the load during some common vehicle maneuvers such as acceleration, braking or deflection while on the road [2]. The coil spring is one of the main components for dampening vibrations and shocks to the load so as to provide comfort and security while the vehicle is in motion [3]. Depending on the condition of their application, coil springs often sustain fatigue failure. This indicates that the tension received below by the coil spring from the maximum stress of the material while sustaining a dynamic load causes fatigue failure [4-8]. The yield strength of the material is also a criterion of failure. Components of automotive suspension must be changed with a traveling distance of 73,500 km, or every five years [9]. The fault of 13.18 % of 24.2 million vehicle tests was recorded [10]. With the development of computing technology, the numerical analysis method has become particularly suitable for use because it will increase the calculation efficiency, the cost-effectiveness as well as save time. Various numerical analysis methods are widely available, but the finite element analysis (FEA) has proven to be reliable in solving problems in the field of continuum mechanics [11]." ] }, { "image_filename": "designv11_62_0001687_978-90-481-9689-0_68-Figure7-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0001687_978-90-481-9689-0_68-Figure7-1.png", "caption": "Fig. 7 Geometry 3.B of the 3-UPU TPM.", "texts": [], "surrounding_texts": [ "In this section, three manufacturing solutions are presented in order to avoid the leg collision in the Geometries of type B (crossed legs) of the 3-UPU TPM. Geometry 1.B is taken (for clarity) as an example of this type of 3-UPU TPM. The first manufacturing solution S1, is to rebuilt the platform of the manipulator. This is obtained by disconnecting the platform of this geometry from the legs and rotating it by a suitable angle \u03b1 about the z axis of Sb, then connecting again the legs to the platform still keeping the same base axis directions. This means to manufacture a platform with the revolute axis directions rotated of \u03b1 (clockwise in the example shown in Figure 8a) with respect to the Geometry 1.B. This makes it possible to avoid the leg collision. In Figure 8a, the universal joints on the base and on the platform are represented by points for clarity, and the prismatic ones are omitted. After manufacturing the new platform, the coordinates of the center of the universal joint on the platform Ai, i = 1,2,3, are given by: 601 A.H. Chebbi and V. Parenti-Castelli OpAi = cos\u03b1OpA\u2032 i + sin\u03b1OpA\u2032\u2032 i , with OpA\u2032\u2032\u22a5OpA\u2032; and \u2225\u2225OpA\u2032\u2032 \u2225\u2225 = \u2225\u2225OpA\u2032 \u2225\u2225 (2) where A\u2032 i, i = 1,2,3, are the centers of the universal joints on the platform of the Geometry 1.B. The second manufacturing solution S2, schematically shown in Figure 8b, is to rebuild both the base and the platform of the Geometry 1.B in order to have the coordinates of the centers of universal joints at the base and at the platform, respectively Bi and Ai, i = 1,2,3, see Figure 8b, given as follows: ObBi = ObB\u2032 i + eq1i, and OpAi = OpA\u2032 i + eq4i (3) where B\u2032 i and A\u2032 i, i = 1,2,3, are respectively the center of the universal joints in the base and in the platform of the original Geometry 1.B; q1i and q4i, i = 1,2,3, are respectively the unit vectors of the revolute joints on the base and on the platform, which maintain the same directions of the original Geometry 1.B; e is a given distance between the corresponding center of universal joints in the platform of the Geometry 1.B and the platform rebuilt. The third manufacturing solution S3, schematically shown in Figure 8(c), is to rebuilt the second and the third link of each leg of the Geometry 1.B in order to change the physical position of the prismatic pairs on each leg along EiFi, where the coordinate of points Ei and Fi, i = 1,2,3 are given by: ObEi = ObBi + dq2i, and OpFi = OpAi + dq3i (4) where Bi and Ai, i = 1,2,3, are respectively the centers of the universal joints in the base and in the platform of the Geometry 1.B; q2i and q3i, i = 1,2,3, are respectively the unit vectors of the intermediate revolute joints of the i-th leg; d is a given distance between the directions of the prismatic pairs for the Geometry 1.B and the manipulator geometry after rebuilding. 602 Geometric and Manufacturing Issues of the 3-UPU Pure Translational Manipulator" ] }, { "image_filename": "designv11_62_0003356_978-981-13-5799-2_12-Figure12.75-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0003356_978-981-13-5799-2_12-Figure12.75-1.png", "caption": "Fig. 12.75 Coefficient of friction with respect to the tire temperature, sliding velocity and thermal conductivity of a tire [114]", "texts": [ " The heat conduction Qi from the interface to ice per unit length of the track is given by12 Qi \u00bc B Tm T0\u00f0 \u00de=V1=2 s \u00bdJ=length B \u00bc 2kib l=\u00f0pai\u00def g1=2; \u00f012:102\u00de where ki is the thermal conductivity of ice, ai is the thermal diffusivity of ice, l is the contact length, and b is the contact width. Using Eqs. (12.101) and (12.102) and considering the heat required to melt ice, the friction coefficient on ice l is given by l \u00bc Qr \u00feQi \u00feQm\u00f0 \u00de=Fz lr \u00fe li \u00fe lm; \u00f012:103\u00de where Fz is the vertical load and Qm is the heat required to melt ice per unit length of the track. lr is proportional to 1/Vs while li is proportional to 1= ffiffiffiffiffi Vs p . As lm increases, l decreases as shown in Fig. 12.75a, b. This is because the layer of melted water is thick in the region of higher values of lm. l increases with the conductivity of the tread rubber k because the heat conduction from the interface to the tire Qr increases as shown in Fig. 12.75c. 12Note 12.6. Giessler et al. [120] measured the temperature increase on the surface of a winter tire on a hard packed snow track by conducting a longitudinal traction test in the time domain. The maximum friction power for the tire is about 3 kW and the maximum difference in surface temperature over time is about 6 K. Although these data are for the temperature increase of a tire on snow, similar temperature increases may be observed on ice. (2) Hayhoe and Shapley\u2019s model (2-1) Changes in temperature and heat flux at the interface Hayhoe and Shapley [116] developed an analytical model of the friction coefficient of a sliding tire on ice by dividing the contact area into two distinct regions, namely a dry sliding region and viscous flow region" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0002041_isr.2013.6695683-Figure2-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0002041_isr.2013.6695683-Figure2-1.png", "caption": "Fig. 2. The exoskeleton subject to external forces in the single suppOli phase.", "texts": [ " It is the way to estimate the desired GRF using accelerometer attached to human's body and transform for the desired GRF into joint torque. In the following part of this paper, we first explain the proposed control method using GRF estimation. Then, by applying the control method to the exoskeleton model, simulation results demonstrate the effectiveness of it. .. .. - -xexo = xh F:xo = Rexo - g exo (1) (2) In the single support phase, the only one leg is in contact with the ground so that we estimate the desired GRF acing on the leg supporting the exoskeleton. Figure 2 shows forces acting on the exoskeleton in the single support phase, and the sum of them is as follows: I F = mZi mg - R = mZi (3) In the equation (3), m is mass, R is ground reaction force at the center of pressure (CoP), CoP is the contact point on the ground, g is the acceleration of gravity and a is the acceleration at the center of mass (CoM). Since we can determine the acceleration of the exoskeleton the acceleration of the exoskeleton by measuring the acceleration of the wearer, the desired GRF can be estimated like equation (4)" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0001310_(asce)as.1943-5525.0000121-Figure3-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0001310_(asce)as.1943-5525.0000121-Figure3-1.png", "caption": "Fig. 3. Cross-section specification of the wing structure", "texts": [ " (6b), the three terms represent the quasi-steady lift, the lift attributable to apparent mass of the fluid, and the lift attributable to wake velocity, respectively. In the same manner, in Eq. (6c), three terms are the quasi-steady moment, the moment attributable to apparent mass of the fluid, and the moment attributable to wake velocity. The equation of the surface of the wing structure at time t is F\u00f0x; y; z; t\u00de \u00bc za\u00f0x; y; t\u00de w0\u00f0y; t\u00de \u03d5\u00f0y; t\u00de \u00b7 x \u00bc 0 \u00f07\u00de where za denotes the vertical position of any point in the crosssection shown in Fig. 3. Then the no-penetration boundary condition of the flow can be stated as DF\u00f0x; y; t\u00de Dt \u00bc \u2202F \u2202t \u00fe U \u00b7\u2207F \u00fe \u00f0Un \u00fe ub \u00fe uw\u00de \u00b7 \u2202F \u2202x \u00fe v \u00b7 \u2202F \u2202y \u00fe \u00f0wb \u00fe ww\u00de \u00b7 \u2202F \u2202z \u00bc 0 \u00f08\u00de In this paper, the Cartesian velocity components in Eq. (8) are defined as ub \u00bc \u2202\u03a6b=\u2202x, uw \u00bc \u2202\u03a6w=\u2202x, wb \u00bc \u2202\u03a6b=\u2202z, and ww \u00bc \u2202\u03a6w=\u2202z, in which (\u03a6b, \u03a6w) are the velocity potential functions of the bound vortex and the wake, respectively. On the basis of the small perturbation assumption and the thinairfoil theory, practically valid in the case of large aspect ratio wing, the velocity components about the z-axis and the downwash are obtained" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0001713_s11431-011-4589-4-Figure3-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0001713_s11431-011-4589-4-Figure3-1.png", "caption": "Figure 3 Distribution of the stator coils.", "texts": [ " Concentrated coils are applied in the stator poles of the PM spherical motor and divided into three layers. Stator coils of the middle layer are located on the equator of the rotor sphere. Both the upper layer and the lower layer are symmetrical about the equator. Each layer has 18 coils, which are divided into three phases. In order to ensure the same current flowing through each coil and uniform variation of the electromagnetic force, every six coils are connected in series to be one phase. As shown in Figure 3, A1, A1\u2032, A2, A2\u2032, A3 and A3\u2032 constitute one phase, of which A1 and A1\u2032, A2 and A2\u2032, A3 and A3\u2032 are respectively connected in reverse series. The connection schematic diagram of A phase stator coils is shown in Figure 4. The connection mode of A phase, B phase and C phase is the star connection. By controlling the commutation order of the stator coils with a DSP controller, spin motion of the PM spherical motor can be achieved. As shown in Figure 1, the area of the flux lines across the stator coil is much less than the surface area of the permanent magnet" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0000253_iciea.2013.6566377-Figure5-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0000253_iciea.2013.6566377-Figure5-1.png", "caption": "Figure 5. Normalized frequency contents of a usual vibration signal", "texts": [ " From these time-segments spectral contents, the vibration signal time-segment are calculated using (6) = FFT ( ) (6) ! = 0,1,2, \u2026 ,512 Where ! indicates any particular frequency calculated from a time segmented using a window size of . All calculated spectral contents are scaled to 0 to 512 Hz frequency range and amplitudes are normalized by (7): \u0305 = /max( ) (7) Therefore 0\u2264 \u0305 \u2264 1 Normalized frequency contents of a usual vibration signal from a bearing with a particular window size of 1024 for 64 time segments are shown in the Fig. 5. Thus in this section a set of spectral features of the time domain vibration signal has been obtained using multi-size segmentation window which will be used for ANN training and testing. The future of the diagnostic techniques will be self-learning and adaptive. Biologically inspired ANN classifiers with diverse learning capabilities are among the best candidates for such future approaches. A neural network for fault classification can be achieved efficiently due to the fact that neural network is a nonlinear empirical model which can capture the nonlinear system dynamics and do not require knowledge of particular system parameters [7]" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0001344_pedstc.2011.5742411-Figure3-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0001344_pedstc.2011.5742411-Figure3-1.png", "caption": "Figure 3. 2-D cross section of the simulated machine.", "texts": [ " The movement is taken into account by means of the moving band technique, the Maxwell stress tensor and the mechanical oscillation equation. The general field equation describing an electromagnetic system in 2-D domain when the magnetic potential and current density only have axial variation is as, 0 (1) The machine structure in 2-D cases can be divided into 4 different zones with the same equations for elements located in each one. The explanations and formulation of these zones for BDFM are mentioned in subsections A to D. This zone includes the elements of the stator and rotor cores (Fig. 3). There is no independent voltage or current source in this region. The field equation of the elements in zone 1 is stated in (2). 0 (2) By applying the Galerkin and the Euler recurrence methods, the matrix form equation of (2) is obtained as (3). (3) where, is the magnetic potential vector of nodes and and matrices are defined in (4) and (5), respectively [14]. (4) 1 0.5 0.5 0.5 1 0.5 12 0.5 0.5 1 DQ t \u03c3 \u23a1 \u23a4 \u23a2 \u23a5= \u23a2 \u23a5\u0394 \u23a2 \u23a5\u23a3 \u23a6 (5) Where, (6) and are the global co-ordinate of the ith node in a triangular element with three nodes" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0003207_6.2019-1912-Figure1-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0003207_6.2019-1912-Figure1-1.png", "caption": "Fig. 1 Free body diagram of the satellite body, with indicated the manipulator reaction force f0 and torque t0 on the base platform", "texts": [ " Since the base frame of the manipulator {0} is assumed to be rigidly attached to the satellite frame, B, and since there is no relative motion between the two components it follows that \u00db\u03b80 = \u00dc\u03b80 = 0, Eqs. (6) and (7) simplify to \u03c90/F = R0 B \u03c9B/F , (9) \u00db\u03c90/F = R0 B \u00db\u03c9B/F . (10) At every iteration, using the initial conditions \u03c9B/F , \u00db\u03c9B/F , \u00dbvB/F and external forces and torques Fext, the outward and inward routines are run in sequence. These ultimately provide the reaction forces f0 and t0 acting on the manipulator base frame {0} due to the arm\u2019s dynamics. At the satellite body frame B, the free-body diagram (Fig. 1) yields fB =FB + RB 0 f0, (11) tB =TB + RB 0 t0 + P{C,B }/B \u00d7 FB + PB 0/B \u00d7 RB 0 f0, (12) where the vector P{C,B }/B is the offset between the frame with origin at the center of mass of body B and the body frame itself, expressed in the B frame. Ultimately, the equations of motion of a free-flying manipulator can be expressed as follows [14][ M Mmb M\u1d40 mb Mb ] [ \u00dc\u03b8 \u00dcx ] + [ C Cb ] + [ J\u1d40 J\u1d40 b ] Fext = [ \u03c4arm F F B ] , (13) where M, C, J, and \u03c4arm are as in Eq. (5), Mb = diag([m13\u00d73, JB ]) is the inertia matrix of the satellite base, with m = m f + \u2211K k=1 mc,k , where m f is the mass of the satellite frame without VSCMGs and mc,k is the mass of the k-th VSCMG" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0000763_icece.2010.900-Figure1-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0000763_icece.2010.900-Figure1-1.png", "caption": "Figure 1. Vector diagram of synchronous motor at steady state condition", "texts": [ " The traditional stator flux estimation is very sensitive to a variety of error, while the effective realization of the electrically excited synchronous motor DTC system is based on obtaining accurate information of the stator flux. For this, a novel full-order closed-loop stator flux observer is used for eliminating the stator flux track offset caused by conventional integrator in this paper. II. THE ELECTRICALLY EXCITED SYNCHRONOUS MOTOR DTC CONTROL BASE ON NOVEL STATOR FLUX OBSERVER The steady vectorgraph of the electrically excited synchronous motor with unity power factor is shown in Fig.1. By rapidly increasing the angle between stator flux s\u03c8 and air-gap flux m\u03c8 , it can quickly generate a large torque step. Therefore, the synchronous motors of direct torque control have fast dynamic response performance. In other words, by switching the appropriate voltage space vector, the torque angle between the stator flux s\u03c8 and d-axis can be quickly changed, and thus the electromagnetic torque will rapidly increase or decrease. In order to accurately obtain information on the stator flux, a full-order closed-loop stator flux observer is proposed in this paper which is shown in Fig" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0002567_icca.2017.8003033-Figure1-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0002567_icca.2017.8003033-Figure1-1.png", "caption": "Fig. 1. Quadrotor configuration and coordinate systems.", "texts": [ "00 \u00a92017 IEEE 46 The rest of this paper is organized as follows. Section II shows the attitude dynamics of the quadrotor and gives the problem formulation. In Section III, the AFTC method is proposed for the attitude tracking of the quadrotor. Simula- tion in Section IV show the effectiveness of proposed AFTC method. Finally, Section V will give a brief conclusion and future work. II. QUADROTOR ATTITUDE DYNAMICS The structure of the quadrotor used in this research and its coordinate systems are shown in Fig. 1. The inertial frame E, is defined by axes {XE ,YE ,ZE} with ZE pointing upward. The body frame B is defined by {XB,YB,ZB} with origin in the mass center of the quadrotor, XB is in the middle of the rotor 1 and rotor 2, ZB is vertical to the plane of cross structure. Then, as shown in [18], the quadrotor attitude dynamics can be given by d dt (J\u03a9) = \u03c4 (1) where \u03a9 = [ p q r ]T and J = diag [ jx jy jz ] are the angular velocity vector in the body frame and the inertia matrix of quadrotor respectively" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0002865_we.2149-Figure13-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0002865_we.2149-Figure13-1.png", "caption": "FIGURE 13 Conceptual design in exploded view", "texts": [ " Additionally, we demonstrated that the number of required conventional pitch motions can be significantly reduced by using the cage hinge. This also makes it possible to use plain bearings, which offer several advantages. They can withstand the vibrations and high static loads that occur better than rolling element bearings. Furthermore, plain bearings tend to be more cost-effective. Another major advantage of plain bearings is their ability to be segmented. To benefit from this advantage, we propose a design whereby U-shaped plain bearing modules enclose a fixed sliding ring (see Figure 13). Each module contains several slide linings. To reduce friction, the use of porous sintering bronze-based materials is conceivable. These materials usually provide a polymer-based lubricant, eg, PTFE, and are therefore permanently dry lubricated. Nevertheless, the slide linings wear over time and must be replaced. Replacing rolling element bearings requires the costly process of dismounting of the blades, which may reduce the profitability of the turbine. By contrast, the wear parts of the segmented plain bearing can be replaced separately for each module via the inside of the hub", " The cage hinge places emphasis on some requirements: Sufficient power and frequent reversions of the working direction at high response rates are crucial for realizing the small ongoing pitch motions. With this in mind, hydraulic solutions have some advantages over electric servo drives. The hydraulic operating force is available with very short delay times at all hydraulic cylinders, even when starting against full load. In addition, the force is also maintained without difficulty at standstill, which relieves parking brakes or even makes them redundant. Although there is the danger of leakage, modern hydraulic systems offer high reliability and durability. In Figure 13, possible configurations for the actuation mechanism of the cage hinge and the segmented plain bearing are shown. Both use the chosen hydraulic drives. The components described in the previous sections are combined to form an overall concept design. The dimensioned FB is completed by adding a central tube. It passes through the upper plate without touching it and joins the lower plate to the actuation mechanism. This closes the load path between the upper and lower plates, which can be therefore rotated in opposing directions", " It therefore has a negligible effect on the characteristic of the FB and was not included in the design optimization. The FB is directly combined with a plain bearing consisting of 18 segments and the sliding ring. The cage hinge and the plain bearing each have an individual actuation mechanism. The actuation mechanism for realizing the rotation of the cage consists of 3 hydraulic cylinders working together. The plain bearing only requires one cylinder. When considering the assembly as a whole (shown in Figure 13 in exploded view), drawbacks and opportunities can be recognized: \u2022 Weight: Since the FB does not replace the conventional bearing, the weight of the entire bearing unit on the rotor will increase. To a great extent, this depends on the materials used and the structural design. Although it will increase the complexity, the overall weight must be a side criterion for further optimizations. Applying lightweight construction methods will offer several options for effectively reducing the weight. \u2022 Modularity: The cage hinge with the corresponding actuation mechanism on top is functionally completely separate from the plain bearing and its actuation mechanism", " As mentioned previously, the slide linings and other wear parts can be individually checked and replaced in a cost-efficient way, which prevents breakdowns at an early stage. This is particularly beneficial for offshore turbines, whereby long maintenance intervals and quick, simple maintenance measures are of great advantage. \u2022 System control: The FB is (theoretically) free of clearance and friction and the response of the hydraulic cylinders is fast and precise. This facilitates advanced closed-loop control with low latency to obtain the maximum power yield and fatigue load reduction from IPC. As mentioned before, the presented concept design in Figure 13 shows only one possible layout. Several changes can be made, for example, reversed installation of the FB unit to place its actuation mechanism to the side closest to the hub. The design freedom also applies to all parts that are not visible in the figure for reasons of clarity, such as housing, wires, and hoses. For scenarios involving the extensive use of IPC, there is a need for a new blade bearing solution that can handle the unfavorable load conditions of IPC. We therefore propose a new concept approach, which is novel due to the way that its key component is an FB that is based on elastic deformation instead of on rolling and sliding contact" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0000389_detc2013-12966-Figure1-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0000389_detc2013-12966-Figure1-1.png", "caption": "Figure 1. ROTOR WITH OVERHUNG MASS", "texts": [ " The current model for the Morton effect first requires an estimate for the initial mechanical imbalance in the rotor. This imbalance is then put into the dynamic rotor model to obtain the synchronous orbit. A fluid film bearing model is then used to establish the temperature distribution which leads to the thermal imbalance. Finally, an unbalance threshold criterion is used to predict the system stability. The Morton effect induced thermal instability in overhung rotors is focused in the current model. Figure 1 is the schematic of the rotor with overhung mass. H and C in the figure denote hot spot and cold spot, respectively. Yd represents the thermal deflection at the overhung equivalent mass center of gravity. The proposed solution method requires the calculation of the bearing journal response for an initial mechanical imbalance. A nominal initial mechanical imbalance Um, has been defined as the imbalance created from a centrifugal force equal to 10% of the total static rotor weight W. This ratio selection is based on the experimental observation representing the most typical situation in reality" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0001620_978-1-4614-3475-7_2-Figure2.33-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0001620_978-1-4614-3475-7_2-Figure2.33-1.png", "caption": "Fig. 2.33 Example 2.10", "texts": [ "67), the mass of the plate is M = \u221a 2 a \u03c10 \u222b a 0 dx \u222b a 0 \u221a x2 + y2dy = \u221a 2 a \u03c10 \u222b a 0 [ y 2 \u221a x2 + y2 + x2 2 ln ( y+ \u221a x2 + y2 )] dx = \u221a 2 a \u03c10 \u222b a 0 [ a 2 \u221a x2 + a2 + x2 2 ln ( a+ \u221a x2 + a2 x )] dx = \u221a 2 a \u03c10 { a 2 \u222b a 0 \u221a x2 + a2dx+ 1 2 \u222b a 0 x2 ln ( a+ \u221a x2 + a2 x ) dx } = {\u221a 2 a \u03c10 a 2 [ x 2 \u221a x2 + a2 + a2 2 ln ( a+ \u221a x2 + a2 )]}a 0 + { 1 2 \u221a 2 a \u03c10 [ ln ( a+ \u221a x2+a2 x ) + a 3 x 2 \u221a x2 + a2\u2212a 3 a2 2 ln ( a+ \u221a x2+a2 )]}a 0 = 1 2 \u221a 2 a \u03c10 [ ln ( a+ \u221a x2 + a2 x ) + a 3 x 2 \u221a x2 + a2 \u2212 a 3 a2 2 ln ( a+ \u221a x2+a2 )] = \u03c10a2 3 [ 2+ \u221a 2ln ( 1+ \u221a 2 )] . Example 2.10. Revolving the circular area of radius R through 360\u25e6 about the x-axis, a complete torus is generated. The distance between the center of the circle and the x-axis is d, as shown in Fig. 2.33. Find the surface area and the volume of the obtained torus. Solution Using the Guldinus\u2013Pappus formulas A = 2\u03c0yCL, yC C x y O Fig. 2.34 Example 2.11 V = 2\u03c0yCA, and with yC = d the area and the volume are A = (2\u03c0d)(2\u03c0R) = 4\u03c02Rd, V = (2\u03c0d) ( \u03c0R2)= 2\u03c02R2d. Example 2.11. Find the position of the mass center for the semicircular area shown in Fig. 2.34. Solution Rotating the semicircular area with respect to x-axis, a sphere is obtained. The volume of the sphere is given by V = 4\u03c0R3 3 . The area of the semicircular area is A = \u03c0R2 2 " ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0002619_s2075108717030051-Figure2-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0002619_s2075108717030051-Figure2-1.png", "caption": "Fig. 2. Visualized models of \u041c\u0422-2010 (\u0430), \u041c\u0422-2012 (b) in pressure field of inflow velocity: Ry is the vertical component of the resultant hydrodynamic resistance (lifting force), L is the distance between the lift center to AUV center of mass (arm of hydrodynamic moment in vertical plane).", "texts": [ " Analytic and computing operations during AUV dynamics research include the following elements: \u2022 construction of 3D visual model of the vehicle on the basis of structural diagram of the project (Solid Work), \u2022 hydrodynamic calculation of the model by means of virtual f low method (Flow Vision), \u2022 identification of parameters of motion mathematical model (Symbol Toolbox Matlab), \u2022 analysis of stability criteria \u201cin large\u201d and \u201cin small\u201d, \u2022 construction of stability areas in state space and control parameters (Symbol Toolbox Matlab), \u2022 optimization of dynamic processes taking into account requirements to stability and control. Figure 2 shows AUV visualized computer models in elevation profile of pressure field with the angle of attack \u03b1 of 5\u00b0 and inflow velocity of 1 m/s. From the images it is seen that there is pressure concentration near the protruding parts that influences the position of vertical component of resulting hydrodynamic resistance. It is shifted to the stern with respect of the center of mass in MT-2010 and to the bow in MT-2012. It follows from this example that there are differences in hydrodynamic properties of these vehicles due to the difference in size of hull and tail control surfaces" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0000421_detc2012-70904-Figure1-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0000421_detc2012-70904-Figure1-1.png", "caption": "Figure 1 - Example of the Body-Bar graph of a mechanism. (a) The mechanism; (b) The Body-Bar graph with polygons; (c) The Body-Bar graph with revolute joint; (d) Bodies in the graph circuits, body 0 is grey to indicate ground link.", "texts": [ " Definition: The Body-Bar graph G=(B,E) consists of |B| bodies and |E| constraints/bars. When dealing with this type of graph, the main focus is on the bodies and the constraints between them. In mechanisms, each type of kinematic pair corresponds to a set of constraints/bars. Because a higher/lower pair imposes one/two constraint(s) between the bodies, there is(are) one/two bar(s), respectively. The order of the body is defined according to the number of bars incident to it. As an example, the mechanism in Figure 1a has five links (including the ground); thus, in its graph (Figure 1b) there are five bodies. Link 1 is connected to link 2 through a revolute joint: there are two bars between body 1 and body 2, while there is one bar between body 2 and body 4 since the corresponding links are connected through a higher pair. For the sake of simplicity, two bars connecting two bodies can be replaced by a revolute joint resulting in a more compact graph as shown in Figure 1c. In rigidity theory [18] the convention is to draw the bodies as circuits instead of polygons, as shown in Figure 1d. Since the graph is a topological representation, there is, of course, no difference between the variants in Figure 1 since all are topologically equivalent. In the following section we introduce the building blocks of this type of representation and the operations needed to construct these building blocks. In section 4.2 we introduce an efficient algorithm from rigidity theory for decomposing the graphs into building blocks. Since we know how to construct these building blocks, section 6.2 introduces a topological gear train synthesis method in which we construct various building blocks, combine them, and convert them into gear trains" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0000571_1.4023085-Figure1-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0000571_1.4023085-Figure1-1.png", "caption": "Fig. 1 Multilayer film representation", "texts": [ " The following hypotheses are also included in the developments. Hypotheses. \u2013 The convection cooling is assumed to be negligible compared to the conduction cooling across the film thickness [6]; \u2013 A steady state conduction regime is assumed; \u2013 Only viscous heating is considered; \u2013 The heat generation is concentrated in the shearing layer; \u2013 The heat generation is constant only across the shearing layer; \u2013 The shearing layer presents a liquid behavior; \u2013 The shearing layer may be located anywhere in the film thickness. Figure 1 illustrates some of the variables. Both global (x,y,z) and local (x1,y1,z1) coordinate systems are located at the initial point/line of contact. However, the former is situated at the midposition in the film thickness, while the latter is attached to the midposition of the shearing layer. K @2T @z2 \u00fe H \u00bc 0 (8) where H is given by Eq. (9). H \u00bc l _c2 (9) where l is calculated with Eq. (4). The factors f and g define the relative shearing layer thickness and position, respectively. The following expressions establish the relations between the variables: h 2 z h 2 (10a) z1 \u00bc z g h 2 (10b) h 2 1\u00fe g\u00f0 \u00de z1 h 2 1 g\u00f0 \u00de (10c) gj j 1 fj j (10d) Equation (8) is solved accounting for the three lubricant zones illustrated in Fig. 1. Moreover, if b represents the heat proportion flowing to solid 1, Eq. (8) can be solved considering the heat partition. In zones l1 and l2 H\u00bc 0 (or q has a constant value). Consequently, the temperature T presents a linear distribution (@2T=@z2 \u00bc 0). Considering the Fourier\u2019s law (Eq. (11)), and integrating, Eq. (12) gives the temperature distribution q \u00bc K @T @z1 (11) T \u00bc qz1 K \u00fe c (12) where c is the integration constant. Since H and q are related by the first thermodynamic law (Eq. (13)), replacing q and considering that (T z1 \u00bc fh=2\u00f0 \u00de \u00bc Tf 1) leads to Eq", " Consequently, a reliable evaluation of the inlet temperature rise caused by rolling/sliding conditions may be derived from the Wilson and Sheu thermal correction factor. The Gupta expression (Eq. (6)) can also be considered. However, since these factors imply semi-empirical formulations, adding the sliding contribution to the Ref. [32] analysis appears to be preferable. Therefore, for rolling/sliding conditions, successive integrations of the heat equation (Eqs. (8) and (9)), considering the boundary conditions as T\u00bc TS at z \u00bc 6h=2 and _c given by Eq. (38) for the global coordinate system of Fig. 1 leads to the inlet temperature distribution (Eq. (39)). _c \u00bc @u @z \u00bc z l @p @x \u00fe u1 u2 h \u00bc C1z\u00fe C2 (38) T TS \u00bcDT \u00bc l K C1 12 z4 h 2 4 ! \u00fe C1C2 6 z3 h 2 2 z !( \u00feC2 2 2 z2 h 2 2 !) (39) where h in the inlet may be assumed to vary from h1 (\u00bc1) far from the contact zone to hc in the contact zone (x\u00bc \u2013a). Moreover, since secondary inlet flows mix the lubricant, it may be assumed that the averaged value of DT offers a reasonable evaluation of the inlet lubricant temperature. If Eq. (39) is integrated with respect to z, the average temperature increase across the film thickness is given by 1 h \u00f0h 2 h 2 DTdz \u00bc lh2 48K h2 5 C2 1 \u00fe 5C2 2 (40) The average inlet temperature increase is finally obtained by the integration of Eq" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0002329_s40997-017-0082-4-Figure2-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0002329_s40997-017-0082-4-Figure2-1.png", "caption": "Fig. 2 Shaft and bush dimensions and assembled specimens", "texts": [ " Then, these values were compared to investigate the effects of the roughness on the friction coefficient of the interference joints. In this research, an interference joint of a shaft and bush was selected for experimental study. Internal ring of a standard roller bearing type of NA6906 was chosen as bush. Dimension of standard roller bearing with designation tag NA6906 is shown in Fig. 1. Shafts were made by VCN150 (AISI4340) steel with 30.030 mm diameter and effective length of 30 mm as depicted in Fig. 2. VCN150 is a heat-treatable and low alloy steel containing chromium, nickel and molybdenum. It has high toughness and strength in the heat-treated condition. VCN150 is used in most industry sectors. Typical applications of this alloy are heavy duty shafts, gears, axles, spindles, couplings and pins. By considering the diameters of the shaft and bush, there is 30-lm diametric interference in the joint. For determination of the mechanical properties, materials of the shaft and bush were analyzed by spectrometry tests; the ingredients of the materials agreed with AISI4140 and DIN100 Cr6 for shaft and bush, respectively" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0002029_978-3-319-02609-1-Figure4.4-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0002029_978-3-319-02609-1-Figure4.4-1.png", "caption": "Fig. 4.4 Schematic illustrations of the 1-DOF bond model of Case Study III. (A) Binding and non-binding surfaces of a building block. (B) Cross-sectional view of two building blocks of radius R assembled with an offset b. (C) Three instances of self-assembled structures with different alignment. The 0-DOF variant of the case study leads only to self-assembled structures with perfect alignment, but it is less computationally expensive.", "texts": [ " (iv) Finally, the robot that has just aggregated receives a message equal to the chain size and (v) does not aggregate because this value is too large. 4.3 Case Study III: Self-assembly of Microscale Components 43 Components Case Study III This case study investigates the SA of cylindrical building blocks with one hemispherical end (Section 3.1.3). Their geometry prevents the formation of unstructured aggregates or chains; rather, they tend to bind to each other in a pairwise fashion (see also Figure 4.4). Because of the experimental difficulties discussed in Section 3.1.3, we rely solely on simulations to investigate this type of system; yet, these numerical tools allow us to gain useful insights into the applicability of our methodological framework. We shall stress that these simulations serve as baseline for higher level models, this is why we describe them as an integral part of the case study. For studying relatively small systems (N0 \u2264 100), we rely on dynamic rigid body simulations implemented in Webots (Figure 4", " Upon collision of two building blocks, if one contact point belongs to the binding surface of both building blocks and their relative bearing is smaller than a given threshold, then both building blocks align along 4.3 Case Study III: Self-assembly of Microscale Components 45 the average direction common to both their axes, and a fixed joint is attached to them. The resulting bond force is computed as follows: Fbond = Farea \u00b7A(b, R) (4.2) where Farea is the surface force per unit area, and A(b, R) is the overlap area3 between two disks of radius R and inter-center distance b (see Proposition 1 of [144] and Figure 4.4B). The condition Fext > Fbond is checked at each time step, and if it is not verified, then the bond is destroyed. This 1-DOF bond model is studied using realistic simulations. Case Study III 2-DOFIn the third variant, we consider a 2-DOF bond model that is characterized by an energy \u0394E given by a Gaussianlike function of the relative alignment \u03be = (\u03b81 \u03b82) T of the colliding building blocks: 3 Since b is the only variable parameter of A(b, r), the bond has indeed a single DOF, even though building blocks may move with respect to each along two distinct dimensions" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0003679_j.mechmachtheory.2019.05.026-Figure1-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0003679_j.mechmachtheory.2019.05.026-Figure1-1.png", "caption": "Fig. 1. Illustration of complete hob profile cross section perpendicular to rotation axis C 0 .", "texts": [ " In the homogenous coordinate notation system used in the present study, a position vector a x i + a y j + a z k is written in the form of a column matrix a = [ a x a y a z 1 ] T . Similarly, the unit directional vector is written as n = [ n x n y n z 0 ] T . The drawing approach provides a feasible means of obtaining the critical points of a hob. However, it is not easily implemented in computer code. Consequently, the present study employs an algebraic method to derive the hob profile instead. Fig. 1 shows the complete hob profile cross section perpendicular to the rotation axis, C 0 . Fig. 2 shows the section profile of a single cutting tooth on the hob with a negative rake angle \u03c8 and outer radius R . The origin of the hob coordinate system ( xyz ) 0 is located at the rotation axis C 0 . For a brand new hob, the following basic design parameters are known in advance: (1) the relief parameter s cs ; (2) the outer radius R = R 0 ; and (3) the number of hob gashes n gash . Given the value of n gash , the interval angle \u03c6 between neighboring gashes is obtained simply as \u03c6 = 2 \u03c0 n gash " ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0002601_gt2017-64123-Figure17-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0002601_gt2017-64123-Figure17-1.png", "caption": "Figure 17. Equivalent von Mises stress distribution in the optimized structure", "texts": [ " Thus, during the optimization, 33 parameters vary: 4 parameters of the air-gas channel; 9 parameters of short blade; 11 parameters of the full-size blade; 9 disk parameters. The time of the optimization was 4 weeks. During this period, 300 configurations were analyzed. The original impeller models and the structure obtained as a result of the optimization are shown in Figure 15. Gas-dynamic and strength characteristics of the optimized impeller Figure 16 shows the velocity field distribution for the middle diameter of an optimized impeller. Figure 17 shows the equivalent von Mises stress distribution in the optimized structure, normalized relative to the maximum stress in the original design. The distribution of the long-term strength factors in the optimized full-length and short blades is shown in Figure 18. The fatigue life distribution of the optimized structure is shown in Figure 19. A Campbell diagram of the optimized structure is shown in Figure 20. The natural frequencies and harmonics values are normalized relative to the first natural frequency" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0000568_s11431-010-4273-0-Figure7-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0000568_s11431-010-4273-0-Figure7-1.png", "caption": "Figure 7 3D sketch of the air cannon test rig.", "texts": [ " Based on the above analysis and many model computations, the optimum model is designed with a 50 mm piston sleeve inlet length and 65 mm nozzle diameter, as compared with the original model`s 29 and 51 mm, respectively, especially using the Laval nozzle with a 70 mm outlet diameter. Computation result shows that the impulse force increases 73% for 0.8 MPa vessel pressure. Unsteady simulation demonstrates that the valid work time is about 100 milliseconds before the impulse force decreases by 30% of the peak value which is reached about 3\u20136 milliseconds after the air cannon works. For measuring the impulse force well, completely rigid connection was used in the test rig as shown in Figure 7. There were the force measurement system and exhausting system which mainly includes air source, pipeline, air cannon and the plate with force sensor. The right-hand part was the force measure system consisting of two flanges which connected exhausting nozzle and plate with the force sensor. The structure made sure the concentricity between nozzle and sensor expediently. It also could adjust the distance between the nozzle and the plate. The sensor measured the force when the air flow impacted the plate" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0001385_ssd.2010.5585589-Figure4-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0001385_ssd.2010.5585589-Figure4-1.png", "caption": "Figure 4. mesh of the magnetic circuit", "texts": [], "surrounding_texts": [ "The study of the behavior of electromagnetic systems depends on the conditions of their operation. Indeed, several formulations are possible such as electrostatics, electrodynamics, the magnetostatic, the magnetodynamic, etc .. Since the induction motor has a dynamic that we can not neglect in the study, our choice fell on a study magneto-evolving, and this allows to follow the temporal and transient behavior of the machine The software used in this study is Flux2D [9]. It allows achieving of a scheme of magnetic circuit in two dimensions. Our choice is oriented toward the plane perpendicular to the axis of rotation of the machine and which develops the electromagnetic field. The magneto-evolutionary model is represented by the following equation: Rot: rotationnel A : magnetic vector potential (Wb/m) J current density uniform (Aim) f1 : magnetic permeability (Him) H : magnetic field (Aim) (j e : electrical conductivity (1/ Q.m ) T: Time (seconds) The magnetic circuit shown below (Figure. 1 ) is the induction motor whose geometry is made of data provided by the manufacturer of induction machine of 5 k W. This is mounted in a test-bed at the laboratory AMPERE Lyonl. The use of the rotating air gap, a function Flux2D, allows us to consider the rotation of rotor in magneto evolutionary study, without making a new mesh of the machine at each position of the rotor. 20 I 0 7th International Multi-Conference on Systems, Signals and Devices Figurel. Circuit magnetic of induction motor the distribution of the winding machine is shown in Figure I according to a Polar periodicity. The rotating air gap is represented (fig.l). It is possible to model the half or a quarter of the machine considering its symmetry, but with the introduction of fault, this consideration wilJ no more bevalid, we therefore used the full representation of the machine. In this modelling, the magnetic circuit saturation is taken into account. The stator and rotor are affected by the material whose magnetization curve is shown below (Fig.2). We must also consider the magnetization curve of the rotor having an influence in the distribution of flux lines Figure 3 shows the mesh made on the magnetic circuit of the machine. We note that the mesh is denser in the vicinity of the air gap, in this place where is realized the electromagnetic exchanges between stator and rotor. The mesh is coarser towards the tree and the outside of the breech to reduce the computing time without loss of information (figA). The external carcass of the machine is conductive; it is generally linked to land. For this reason, we have applied in the simulation the Dirichlet condition on the whole contour of the machine with zero vector potential. 3. ELECTRICAL CIRCUIT OF STATOR Electrical circuit of stator (fig.S) contains coils which wilJ be linked to magnetic scheme. They represent the conductors who wilJ be housed in the stator slots. Between two coils connected, there is a resistance representing the resistance of the head coil (2) and an inductance which represents the leakage inductance slot and head coils (3). 20 I 0 7th International Multi-Conference on Systems, Signals and Devices (2) (3) The calculation of leakage inductance takes into account the permeance of stator slots because it is considered that the surface of the slot in the magnetic circuit is completely filled with copper. Against the permeance of the isthmus and the slot head are considered in the magnetic representation. The assignment of elements of stator electrical circuits to magnetic circuit is performed in an interactive manner through the software flux 2d. 4. SHORT-CIRCUIT FAULTS SIMULATION The fault of short circuit is represented schematically by the connection between two points in the winding. A short circuit manifested usually in of coils of phases different and in the head coil because in these that conductors of different phases intersect. In the stator windings, there are five cases of faults: - short-circuit tum to tum - Short circuit coil to coil - Open circuit - Short circuit phase to phase - Short-circuit coil to ground. They are clarified in Figure 9 In this study, short-circuits are simulated in the electrical circuit model by connecting the coils with a conductor. 4.1 Short-circuit turn to turn Among these five faults, short circuit turn to turn represents the origin and cause of other faults. The persistence of the latter will promotes the emergence of other cases of short circuit. A short-circuit tum to tum in the same coil may occur either between the coil and neutral or between turns in the middle of winding 4.1.1. Short-circuit turns-neutral This short circuit is simulated by the connecting of turns with a conductor linked to neutral, as shown in Figure 6 Figure 7 shows the stator currents in the case of short circuit turns neutral. We note that the current amplitude in the phase affected increases of 35% relative to the other two phases. However the currents of two phases unaffected have increased of 52% compared to the healthy case. More important is the number of turns shorted, stronger is the increase of currants. 4.1.2. Short-circuit in middle of the coil If the short circuit occurs in the middle of the coil (20% of turns is short circuited) (Figure 8), we record, in Figure 9, an increase of currents in the phase affected. A slight variation of the amplitude is noticed for the other phases. The difference from the short-circuit turns-neutral is the asymmetry between the currents of phases uninfected by the fault. coil 2010 7th International Multi-Conference on Systems, Signals and Devices 4.2 Short circuit coil to coil A short circuit between different coils close to the power supply (Fig. 1 0) cause very strong currents that would lead to the merger of power conductors. Therefore the immediate halt of the machine through the protection relay. However, a short circuit close to neutral between two coils causes an imbalance (fig. 11) and don't cause shutdown of the machine. 30 , , , , - - -1- - - 1-- ---,- - --,-- - ,- - \" I \" , \" 20 -j , I , ,--. 10 \u03b42n, 0 s2n < \u03b42n. (67) It is shown that the responses byModified and New type sliding mode control do not bring the chattering in the neighbourhood of the sector boundaries. Simulation results of three types of the sliding sector control are shown in Figures 6\u20139. By the previous sliding sector control, the trajectory of the state moves near the sliding sector boundary" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0000568_s11431-010-4273-0-Figure9-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0000568_s11431-010-4273-0-Figure9-1.png", "caption": "Figure 9 Axisymmetric free jet flow sketch.", "texts": [ " The right-hand part was the force measure system consisting of two flanges which connected exhausting nozzle and plate with the force sensor. The structure made sure the concentricity between nozzle and sensor expediently. It also could adjust the distance between the nozzle and the plate. The sensor measured the force when the air flow impacted the plate. The test rig photo is shown in Figure 8. In the test, compressed air flowed out from the nozzle and generated a free jet flow which had the expansion angle as shown in Figure 9. The free jet flow satisfied the momentum conservation in every section. The plate diameter influenced the force measurement due to its position. Ref. [5] gave the experimental value of the expansion angle tangent Bx as follows: Bx=db/dx=0.086. Three L values of the distance between the plate and nozzle were chosen, including 60 mm which is about one nozzle diameter, 120 mm and 420 mm. The corresponding plate diameter D was calculated as follows: D= D0+2*Bx*L=144.76 mm. So, two plates with 150 and 220 mm diameter, respectively, were used in the test rig to compare the measure impulse forces" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0001344_pedstc.2011.5742411-Figure2-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0001344_pedstc.2011.5742411-Figure2-1.png", "caption": "Figure 2. Structure of the nested loop rotor with 3 loops per nest", "texts": [], "surrounding_texts": [ "Keywords-Brushless Doubly-Fed Machine; 2-D finite element model; magnetodynamic; stator fault.\nI. INTRODUCTION The idea of having two stator windings for variable speed applications is proposed in [1-3]. Dual stator winding machines have been categorized as split-wound and selfcascaded [4]. The self-cascade machine or Brushless DoublyFed Machine (BDFM) was introduced by Hunt. In a BDFM stator windings configuration is similar to stator winding of an induction machine. One of these windings is connected directly to the grid, which is called power winding. The other is fed by a partially VA-rated (e.g. 30% of machine\u2019s rating) bidirectional converter [5], which is called control winding. These windings have different pole number to avoid having coupling between them and usually are excited at different frequencies [6].\nBDFM has three different modes of operation depend on the connection of the control winding. First, simple induction mode when the control winding is open-circuited. Second, cascade mode when the control winding is short-circuited. Third, synchronous mode when the control winding is supplied with a certain frequency regarding the rotor speed. The BDFM is expected to run more efficiently in synchronous mode. In this mode, the rotor speed is independent of the torque, in contrast to the two other modes.\nIt requires a special rotor structure that has some nested loops on the circumference of the rotor to incorporate the effects of cascade connection [7-9]. The BDFM and its rotor structure for 3 loops per nest are shown in Figs. 1 and 2, respectively.\nDuring stator fault, the air gap field may contain components having all possible pole pair numbers. In this situation the magnetic coupling between power and control windings is not equal to zero anymore. So the coils apply forces to each other. As a result, the effective performance of the machine will decreases and sever destruction may occur. Therefore detection of these faults is important.\nIn this paper, to analyze stator faults a finite element model of the machine is introduced then simulation results of the finite element model is compared to the dynamic model of the machine [11], which used generalized harmonic analysis through the coupled circuit [12], to verify the accuracy of the finite element model. The following assumptions are made [13]: \u2022 The stator and the rotor are modeled by two smooth\nconcentric cylinders of infinitely permeable iron. \u2022 The stator windings and the rotor bars are replaced by\nequivalent point conductors lying on the stator and rotor surfaces, respectively.\n\u2022 Flux crosses the air gap in radial lines.\n978-1-61284-421-3/11/$26.00 \u00a92011 IEEE 169", "II. FINITE ELEMENT MODEL A 2-D magnetodynamic finite element model is represented to verify the validation of the proposed dynamic model. This model allowing the simulation of the healthy and faulted machine with its special rotor bar configuration. The machine is modeled in a 2-D domain using the Maxwell equations to formulate the field behavior and the finite element method (FEM) to discrete the domain. The formulation uses the magnetic vector potential as unknown, the Galerkin method to obtain the set of equations to be solved numerically, the Euler recurrence method to discrete the temporal derivatives and the successive approximation and Newton-Raphson method to consider the nonlinear characteristic of magnetic material. The movement is taken into account by means of the moving band technique, the Maxwell stress tensor and the mechanical oscillation equation. The general field equation describing an electromagnetic system in 2-D domain when the magnetic potential and current density only have axial variation is as,\n0 (1)\nThe machine structure in 2-D cases can be divided into 4 different zones with the same equations for elements located in each one. The explanations and formulation of these zones for BDFM are mentioned in subsections A to D.\nThis zone includes the elements of the stator and rotor cores (Fig. 3). There is no independent voltage or current source in this region. The field equation of the elements in zone 1 is stated in (2).\n0 (2)\nBy applying the Galerkin and the Euler recurrence methods, the matrix form equation of (2) is obtained as (3).\n(3) where, is the magnetic potential vector of nodes and and\nmatrices are defined in (4) and (5), respectively [14].\n(4)\n1 0.5 0.5 0.5 1 0.5\n12 0.5 0.5 1\nDQ t \u03c3 \u23a1 \u23a4 \u23a2 \u23a5= \u23a2 \u23a5\u0394 \u23a2 \u23a5\u23a3 \u23a6\n(5)\nWhere,\n(6)\nand are the global co-ordinate of the ith node in a triangular element with three nodes.\nTo consider the nonlinear characteristic of the magnetic material of the stator and rotor core (e.g. dependency of to\n), the set of equations must be solved recursively. A desired convergence rate can be obtained by using the result of the successive approximation method after several iterations as the initial value in the Newton-Raphson method [14]. The Newton-Raphson recursive equation is stated in (7) for the elements of zone 1. \u2206 \u2206 , \u2206 (7)\nis the Jacobian matrix and its procedure to extract is explained in [14].\nThe elements located in the air gap and interior regions of the stator and rotor slots except the conductor regions, are in this zone. There are no independent source and induced currents and the magnetic permeability is constant. The matrix form field equation of the elements in this zone can be written as (8).\n\u2206 0 (8)\nThe special issue about this zone is the moving band locating in the air gap. Fig. 4 provides some details of the mesh, especially in the air gap region. Introducing the movingband technique involves splitting the air gap into three layers, where the middle one is the moving band. In order to minimize the element distortion during the rotation, only the elements on the moving band were remeshed at every step time [15]. The remeshing scheme during the rotor movement is explained in [14-16].\nThe elements in this zone discrete the rotor conductor's regions. The field equation and its matrix form are as (9) and (10), respectively.\n0 (9)\n\u2206 \u2206 (10) where," ] }, { "image_filename": "designv11_62_0000761_2013-01-2221-Figure5-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0000761_2013-01-2221-Figure5-1.png", "caption": "Figure 5. The mounting device used to mount the AE system on the wheel", "texts": [], "surrounding_texts": [ "AEWMS VALIDATION The system validation comprises three tests for healthy, in-service, and faulty bearings that are working in real applications such as automotive wheel bearing. It is designed to assess the operation of the Easy AE wireless monitoring system in detecting bearing failures at an early stage, and to investigate its ability to provide information to users about its status. The test was achieved by first measuring a faulty wheel bearing already run for more than 50\u00d7103 km, using several runs for validation, and then this bearing was replaced with a new one and the same test repeated for several runs.\nThe new wheel bearing was left to operate in rough service till it reached about 20\u00d7103 km of its service life, then retested again after that. All the three tests were carried out at the same wheel speed (190 rpm), and the AEASL system output was selected to be the comparison feature.\nDuring the test of the new bearing, the system computed the AEASL feature to less than 25dB, whereupon the AEASL reached 55.5dB for the faulty bearing. Also, the AEASL signature of the new bearing was compared to the in-service bearing at the same running conditions, and it was found that the AEASL increased from about 25dB to 33.5dB as shown in Figures 7 and 8.\nBy inspecting the faulty bearing raceways, it is evident that this increase of the AEASL feature of about 20dB is attributed to the severe wear found in the outer raceway, occurred in two locations: the first is macro-pitting in the middle of the outer raceway, and the second is uneven wear at both sides of the bearing outer raceway, as shown in Figure 9.", "The study has presented a new cost-effective acoustic emission wireless monitoring system that was able to detect different conditions of wheel bearings and clearly identify its condition using only one AE sensor placed on the drive axle shaft. The AE average signal level was selected to indicate the condition of the bearings, and successfully discriminated between the different bearing conditions. The type of failure identified can be considered as a common failure in most bearings, and its monitoring if of considerable interest for many applications. The system can solve a major problem in applications where the sensing points are far from the acquisition and analysis point. The \u2018EASY AE\u2019 monitoring system has been experimentally validated on an automotive application. The information provided by the system was considered sufficiently accurate and demonstrated the robustness of the information provided by the system. The system is being developed for use on rotating machinery including those in aircraft and wind turbines, and can be adapted for other applications such as structural monitoring. Further publications will follow to provide results of testing the system on other applications, including results during long term testing.\n1. SKF (1983) Bearing Maintenance and Replacement Guide Catalogue\n3600E. 2. ONSY, A., BICKER, R., SHAW, B. A., ROWLAND, C. W. & KENT,\nT. (2008b) Monitoring Bending Fatigue Failure in Helical Gears Using Acoustic Emission, Vibration, and On-Line Oil Debris Analysis: A Comparative Study. Proceedings of the Fifth International Conference on Condition Monitoring & Machinery Failure Prevention Technologies, UK, Edinburgh, 1108-1118. 3. ONSY, A., BICKER, R., SHAW, B. A., ROWLAND, C. W. & KENT, T. (2009) Intelligent Health Monitoring Of Power Transmission Systems; An Experimental Validation. Proceedings of the 2009 Conference of the Society for Machinery Failure Prevention Technology, USA, Dayton, Ohio, 499-518. 4. ONSY, A., BICKER, R., SHAW, B. A., (2010) Intelligent Diagnostic Health Management of Power Transmission Systems: An Experimental Validation. International Journal of COMADEM, Vol.13 (2) pp. 46-58, ISSN 1363-7681. 5. VALLEN-SYSTEM GmbH (2004) Acoustic Emission Signal Conditioner ASCOP-P User Manual, man402E. 6. TAN, C.K., IRVING, P. & MBA, D. (2005) Diagnostics and Prognostics with Acoustic Emission, Vibration and Spectrometric Oil Analysis for Spur Gears: A Comparative Study. Insight: NonDestructive Testing and Condition Monitoring, 47, 478-480.\n7. MURRAY, R., SVOBODA, V. & SVEC, P. (2004) Condition Monitoring Using AE Method. Physical Acoustics Ltd, DGZfPProceedings BB 90-CD Lecture 28. 8. TOUTOUNTZAKIS, T., TAN, C.K., & MBA, D. (2005) Application of Acoustic Emission to Seeded Gear Fault Detection. NDT and E International, 38, 27-36. 9. TAN, C.K. & MBA, D. (2005a) Identification of the Acoustic Emission Source During a Comparative Study on Diagnosis of a Spur Gearbox. Tribology International, 38, 469-480. 10. COLE, P. & BRADSHAW, T. (2004) AE Signals From Process Monitoring. Physical Acoustics Ltd, DGZfP-Proceedings BB 90-CD, Lecture 36. 11. JOHNSON, J.E. (2005) Identifying Common Ultrasonic Predictive Failure Signatures in Bearing Elements for the Development of an Automated Condition Based Ultrasonic Monitoring Controller, M.Sc. Thesis, East Tennessee State University, USA. 12. AL-GHAMD, A.M. & MBA, D. (2006) A Comparative Experimental Study on the Use of Acoustic Emission and Vibration Analysis for Bearing Defect Identification and Estimation of Defect Size. Mechanical Systems and Signal Processing, 20, 1537-1571. 13. LI, C. J. & LI, S.Y (1995) Acoustic emission analysis for bearing condition monitoring. Wear 185, 67-74. 14. JAMALUDIN, N., MBA, D., & BANNISTER, R. H. (2001) Condition monitoring of slow-speed rolling element bearings using stress waves. Proc Instn Mech Engrs Vol 215 Part E. 15. ELFORJANI, M. & MBA, D. Acoustic Emissions Observed from a Naturally Degrading Slow Speed Bearing and Shaft COMADEM 2010. 16. AL-BALUSHI, K.R. & SAMANTA, B. (2002) Gear Fault Diagnosis Using Energy-Based Features of Acoustic Emission Signals. Proceedings of the Institution of Mechanical Engineers. Part I: Journal of Systems and Control Engineering, 216, 249-263. 17. THOTA, S., VANKAMAMIDI, S., MUTHIREDDY, R., WAYNE, S. & QI, G. (2008) Detection and Analysis of Horizontal Conveyor Bearing Failures Using Acoustic Emission Technique. AEWG-51 & International Symposium on AE, Memphis, USA. 18. Onsy, A., Shaw, B., and Zhang, J., \u201cMonitoring the Progression of Micro-Pitting in Spur Geared Transmission Systems Using Online Health Monitoring Techniques,\u201d SAE Int. J. Aerosp. 4(2):1301-1315, 2011, doi:10.4271/2011-01-2700. 19. Onsy, A., Bicker, R., and Shaw, B., \u201cPredictive Health Monitoring of Gear Surface Fatigue Failure Using Model-Based Parametric Method Algorithms; An Experimental Validation,\u201d SAE Int. J. Aerosp. 6(1): 2013, doi:10.4271/2013-01-0624. 20. SAMUEL, P.D. & PINES, D. J. (2005) A Review of Vibration-Based Techniques for Helicopter Transmission Diagnostics. Journal of Sound and Vibration, 282, 475-508. 21. JARDINE, A.K.S., LIN, B., & BAN, J.D. (2006) A Review On Machinery Diagnostics And Prognostics Implementing Condition-Based Maintenance, Mechanical Systems and Signal Processing, 20, 1483-1510. 22. BRAUN, S. (1986) Mechanical Signature Analysis - Theory and applications. London: Academic Press Inc." ] }, { "image_filename": "designv11_62_0001959_cjme.2012.01.179-Figure7-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0001959_cjme.2012.01.179-Figure7-1.png", "caption": "Fig. 7. Sketch map of influence of grip length on grip rigidity", "texts": [ " However, when the basic interference reaches 12 m, the radial displacement increase instead than that of the interference is 10 m. It is because that the maximum equivalent stress exceed the yield strength of LSFH when the basic interference reaches 12 m (refer to Table 2). Therefore, to guarantee the match of LSFH and cutter not only has a good radial grip rigidity but also the strength of LSFH do not exceed the yield strength, the most reasonable interference should be between 8 m and 10 m. Sketch map of influence of grip length on grip rigidity is shown in Fig. 7 when the basic interference is 10 m. Where L is the length of cutter grip; A is the point for radial displacement analysis; F is the applied loading. The simulation results of static radial grip rigidity of the matching of LSFH and cutter under different grip length are shown in Fig. 8. Fig. 8 shows that radial grip rigidity will decrease with the decrease of grip length. When the grip length is less than the maximum matching length (18 mm), the static radial grip rigidity will decline sharply. For example, when the grip length is 8 mm and the applied loading is 250 N, the radial displacement is 24", " In high speed milling, it is necessary to consider the influence of the centrifugal force on the radial grip rigidity in an actual machining condition. The radial grip rigidity curves under different spindle speed are shown in Fig. 9 when the basic interference is 10 m and the grip length is 23 mm. Fig. 9 shows that the influence of the rotational speed on radial grip rigidity can be ignored when the spindle speed less than 20 kr/min since the diameter of LSFH and cutter are usually small. The radial grip rigidity decline sharply when the spindle speed reaches 40 kr/min. For example, the radial deformation of point A (refer to Fig. 7) at spindle speed 40 kr/min is 90.612 m which is almost two times of that when spindle speed is 0 r/min. Fig. 10 shows the relationship between spindle speed and radial deformation when the applied is 200 N. It can be found that the deformation curve fit with the 4 polynomial very well (degree of fitting R2=1), that is to say, the higher the speed, the greater impact of the centrifugal force on the radial grip rigidity. Therefore, in the actual high speed case, the larger interference fit should be selected to offset the effect of centrifugal force on the radial grip rigidity on condition that the strength of LSFH meets requirements", " Test system for static radial grip rigidity is shown in Fig. 11. Dynamometer type Kistler9265B is mounted on the machining center bench of DMU 60T. Radial force on the cutter is achieved by controlling the micro movement of spindle box and recorded and displayed by data processing system. The non-contact displacement sensor, which has a resolution of 0.02 \u00b5m, is fixed on the sensor supporter. The clamping position of the sensor supporter is align with the restrained end of FEA model and the measure position is the same with the point A (refer to Fig. 7). Thus, measured deformations will eliminate the affect of connection stiffness of LSFH and shark, Shark and spindle. In this study the fit of LSFH and cutter is 6H3/n2 which has the maximal interference, least interference and basic interference is 10 m, 6 m, and 8 m, respectively [10]. The measured and the simulated result when the grip length is 23 mm are shown in Fig. 12. It can be found that the measured result agrees with the simulated result very well and the relative error only 4.2% when the applied load is 250 N, which shows the above finite element model is reasonable" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0000249_amm.459.390-Figure3-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0000249_amm.459.390-Figure3-1.png", "caption": "Fig. 3 Full tire-air model", "texts": [ " Since the meshes are different for air and the tire, the in-built option TIE in Abaqus is used. In the 2D model an air-liner is created which represents air surface and tire-liner as tire surface. The air-liner is defined as slave surface and tire-liner is defined as master surface. Since the air has to adjust according to the tire motion these surfaces is tied using surface based TIE option. The elements in the 2D model are rotated 360\u02da to generate the full tire model used for simulation. The full tire model with air and road contact is shown in Fig.3. At first analysis is done at inflation and footprint loads. After this, the undulations from the road are given as a sinusoidal input and are modeled with an excitation force of 10 N to create pressure variation in the tire acoustic cavity. Effect of inflation pressure on air cavity Tire is a pneumatic system which supports vehicle load. Compressed gas or air is filled inside the tire to create tension in the carcass plies enabling it to carry the vehicle load. In the space of just one month a passenger car tire can lose 6-12 kPa (1-2 PSI) of air pressure" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0003711_itec.2019.8790540-Figure12-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0003711_itec.2019.8790540-Figure12-1.png", "caption": "Fig. 12: 2D Magnet Eddy Current Losses at 3000 RPM", "texts": [ " 11 shows the core loss density data comparison between the published measurement data from the steel manufacturer and the fit data from the modified Berttoti\u2019s core loss model. Fig. 11 Magnet Eddy Current Losses \u2013 3D FEA Fig. 11 shows the magnet eddy current losses in 3D. For 2D, the eddy current losses can be computed by using the contraint of zero total current over the magnet cross section. The resulting 2D eddy current losses are more pessimistic than the 3D eddy current losses due to the lack of end magnet eddy current return path (Fig 12). Matched Sine Delta Matched Sine Delta Speed rpm 5000.00 5000.00 7500.00 7500.00 Electrical Freq. Hz 333.33 333.33 500.00 500.00 Shaft Torque N.m 130.09 131.30 -0.92% 75.44 75.17 0.35% EM Torque N.m 132.09 133.10 -0.76% 78.04 78.06 -0.02% Phase Current A rms 94.86 94.13 0.78% 70.63 70.05 0.84% Torque Ripples % 1.66% 19.93% -91.69% 1.87% 25.93% -92.78% Elec. Power W 75442.71 75823.44 -0.50% 67813.71 67435.61 0.56% Mec. Power W 68117.20 68750.84 -0.92% 59250.63 59041.61 0.35% Efficiency % 90.29% 90" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0003210_joe.2018.8514-Figure4-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0003210_joe.2018.8514-Figure4-1.png", "caption": "Fig. 4 Layout of the scuppers", "texts": [ "0/) 2860 Therefore, it can be seen that the installation of box of bird guard at the cross-arm of the tower has little influence on the electric field and potential distribution of the insulator, the installation of box of bird guard will not affect the operation of the transmission line. As the water is easy to accumulate inside the box of bird guard, in practical application, four scuppers are designed at the four corners of the box of bird guard to facilitate the discharge of water. The layout of the bottom scupper of the box is shown in Fig. 4. 3.1.1 Equivalence of the gap between the water column and insulator: In order to analyse the breakdown characteristic of the air gap around the insulator when the water falls from the box of bird guard, a three-dimensional electrostatic field model is established to analyse the electric field and potential distribution around the insulator. The conductivity of rainwater in nature is about 100 \u03bcs/cm, and it can exceed 1 ms/cm in some areas [12, 13], besides, when the water falls from the bird box, it dissolves part of the filth in the box, thus it can have an even higher conductivity" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0000253_iciea.2013.6566377-Figure11-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0000253_iciea.2013.6566377-Figure11-1.png", "caption": "Figure 11. Outer-race fault with window size 1024 for 64 segments", "texts": [], "surrounding_texts": [ "The future of the diagnostic techniques will be self-learning and adaptive. Biologically inspired ANN classifiers with diverse learning capabilities are among the best candidates for such future approaches. A neural network for fault classification can be achieved efficiently due to the fact that neural network is a nonlinear empirical model which can capture the nonlinear system dynamics and do not require knowledge of particular system parameters [7]. Neural network design includes input layer, hidden layers and output layer as shown in Fig. 6. Number of input layer neurons is equal to the input features (7) and the number of output neurons is equal to number of classes. There can be one or more than one hidden layers with different number of neurons. Most of the time increasing the number of hidden layers neurons guarantees good learning but it requires more training time and computational cost. Therefore minimum number of the hidden layers and neurons meeting certain classification accuracy is preferred. Neural Network pattern classifier learns system dynamics in the form of weighted links between the neurons during the training phase. Depending upon problem-nature training can be of the supervised or unsupervised type [4, 9]. In unsupervised learning the targets for input features are unknown while for supervised learning targets are known. We are using supervised learning because for each spectral features training input we have associated targets. For output classification we are using one versus all classifier. Before starting the training the training data is randomly shuffled and divided into training, validation and test sets with different percentages. MSE (Mean Square Error) is used to measure the accuracy of the training, 2013 IEEE 8th Conference on Industrial Electronics and Applications (ICIEA) 263 validation and test sets. Validation set is used as check to stop learning and testing data set is used to measure over fitting of the classifier. Neurons in different layers are connected through weighted links. These weights are tuned during the training of the ANN as shown in Fig. 7. Weights are randomly initialized before starting the training. During training, with random weights, output of the network is used to calculate the MSE using target values. If the calculated MSE is of acceptable value then training can be stopped, otherwise based on the MSE, weights are tuned using back propagation algorithm. Thus, after training, ANN can be used for classification. IV. RESULTS AND DISCUSSION Methodology developed in the previous sections will now be tested practically. Fig. 8 illustrates the experimental setup for recording the actual vibration data sets [10]. Experiments were conducted with four types of bearings including one normal and three faulty bearings with faults in inner-race, ball and outer-race. Faults in the bearing were created by electro discharge machining. Faulty bearings are supporting the shaft of the motor and the load is 2HP with a speed of 1750 r/min. The data have been collected through accelerometers using a 16-channel digital-audio-tape recorder and sampled at the rate of 12000 samples per second. Time vibrations recorded and converted to spectral features for these four signals are shown in the Figs. 5, 9, 10 and 11 respectively with a window size of 1024 for 64 segments. 264 2013 IEEE 8th Conference on Industrial Electronics and Applications (ICIEA) In this paper we have used only one hidden layer with different number of neurons to experimentally check the best training that can be achieved with minimum number of neurons. Number of input features, 513, is equal to spectral features of the time segment. Output layer of the ANN contains four neurons because of four classes as shown in Fig. 12. For training ANN, the spectral contents from different signals with different window size are grouped into a set with respective class (8): ($ , % ) = {($ , % ), ($ , % ), ($ , % ), \u2026 , ($&, %&)} (8) Window size used in this experiment are 256, 512, 1024 and 2048 samples. In (8) p' represents input spectral pattern for any of the four classes with any specific window size. t' represents target class or output for this particular input. Four of the target classes used in this experiment are: *- 1; % = [1 0 0 0] Normal signal *- 2; % = [0 1 0 0] Inner-race fault *- 3; % = [0 0 1 0] Bearing fault *- 4; % = [0 0 0 1] Outer-race fault This grouped data was randomly shuffled and then divided into training, validation and test sets with a respective proportion of 60%, 20% and 20% of the total grouped contents given by (8). With random weight initialization, the network was trained using feed forward back propagation algorithm with different number of hidden layers neurons. Training the network, more than hundred times, for different number of hidden layer neurons, resulted in almost similar minimum MSE. Minimum number of hidden layer neurons that gave acceptable accuracy was 2. Accuracy achieved with two neurons in the hidden layer and more than two neurons are comparable. But using higher number of hidden layer neurons increases the computational time. Fig. 13 shows, as the number of hidden layer neuron increases the learning time increases almost exponentially. Thus 2 hidden layer neurons are best selection for this particular scenario. The appropriate ANN architecture selected is shown in the Fig. 12. As the vibration signals from bearing are quasi-stationary therefore the problem of optimum window selection has been addressed by using multi-sized time-domain segmentation-window for augmented feature selection. Classifier trained with these augmented spectral features has shown 100% accuracy complying with the supposition that smaller windows will capture the signatures appearing for short duration or higher frequencies and larger windows will address the low frequencies contents. The trained network is tested with all the window sizes, 256, 512, 1024 and 2048 2013 IEEE 8th Conference on Industrial Electronics and Applications (ICIEA) 265 with classification accuracy of 100%. Thus multi-window approach offers a range of windows to fit a variety of transients appearing in the vibration signal providing a comprehensive data set for training and testing of the real time rotary machines vibration. V. CONCLUSIONS In this paper, multi-size-window time segmentation based spectral features augmentation for neural network bearing fault classification has been presented. Augmented spectral features of vibration signal, in rotary machines, calculated using multi-size time-segmentation-window have been used to train and test the neural network classifier. Classification results have shown that classifier, with multisize-window time-segmented spectral features, has learned the dynamics of quasi-stationary vibration signals efficiently for real time scenarios with 100% accuracy." ] }, { "image_filename": "designv11_62_0000518_iros.2013.6696697-Figure1-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0000518_iros.2013.6696697-Figure1-1.png", "caption": "Fig. 1. Quadrotor platform", "texts": [ " As marker generators, laser diodes are located on the quadrotor hardware, and vision sensors can detect those projected laser markers on the ground all the time without missing them when the quadrotor is in stable status. Tilting motion of a quadrotor gives distorted laser marker patterns and those patterns can be analyzed to estimate the attitude of the quadrotor. Also, the altitude can be calculated by using laser markers, without adding other sensor. We will demonstrate the detailed method and results with a real robot. Fig. 1 shows the quadrotor hardware that we used. This is propelled by four brushless DC motors (BLDC motor) and those speeds are controlled by ESC (Electric Speed Control) module. These ESC modules use I2C interfaces to communicate and those are connected to ATmega2560 microcontroller. ATmega2560 controls the attitude of quadrotor by adjusting the speed of the rotors with control algorithm. For the image processing, OV9655 image sensor and STM32F407IGT6 micro controller are used. As its internal memory is not enough to handle the image data, external high speed SRAM are used as a buffer" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0001594_eej.22414-Figure15-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0001594_eej.22414-Figure15-1.png", "caption": "Fig. 15. Master\u2013slave robot. [Color figure can be viewed in the online issue, which is available at wileyonlinelibrary.com.]", "texts": [ " As regards the disturbance torque occurring when the arm is in contact with the environment, the disturbance T\u0302dis estimated by the disturbance observer coincides with the reaction force Treac provided that the reaction is produced only by the environment. However, the disturbance torque includes torsional torque Ttors, gravity torque Tg, and other components other than the reaction Treac, and thus we estimated the reaction force by subtracting the mentioned disturbance components from the estimated disturbance T\u0302dis (Fig. 14). The reaction force was estimated as follows: The master\u2013slave robot used in the experiments is shown in Fig. 15. The master\u2013slave robot was driven by a magnetic gear reducer with a speed reduction ratio of 16:1. A force sensor was attached to the master\u2019s arm for verification of the experimental results, but the sensor values were not used in feedback control. On the slave\u2019s arm, the motor output was reduced by a conventional gear reducer (reduction ratio 16:1), and a force sensor (uniaxial load cell) was attached to the end of the arm. In the experiments, the reaction force estimated by Eq. (12) was taken as the reaction on the master side, thus providing force sensorless bilateral control" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0003985_icamechs.2019.8861668-Figure10-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0003985_icamechs.2019.8861668-Figure10-1.png", "caption": "Fig. 10. Mechanism of the trunk", "texts": [ " However,, when a vertical force is applied to support the weight of the robot the rubbers of the joint engages, and the joint becomes locked. Therefore, the stiffness of the leg becomes high and the leg can transmit the force from the motor to the ground as a driving force. The advantage of the proposed mechanism in comparison to the conventional mechanism is that the joint can be locked at any angle as shown Fig. 9. Thus, it can produce a driving force when moving through a narrow space. The mechanism of the leg when flexed is given in Fig. 9. The trunk is illustrated in Fig. 10. It is flexible and can adapt to the rough ground during locomotion. The trunk was made of duplex bellows as shown in Fig. 10 i.e. smaller bellows inside larger bellows for the realization of different stiffness i.e. for small flexure and large flexure, respectively. The trunk is made up of series of connected rectangular prisms, and has high stiffness in the vertical direction for upliftment of its weight and low stiffness in the horizontal direction to allow for the robot to turn. Two wires were installed on both sides of the trunk and were connected to an active pulley with a servomotor as shown in Fig. 11. This allowed for the robot to turn by pulling the wires using the servomotor" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0002791_b978-0-12-812138-2.00003-9-Figure3.33-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0002791_b978-0-12-812138-2.00003-9-Figure3.33-1.png", "caption": "FIGURE 3.33", "texts": [ " The skin effect causes the effective area of the cross section of the conductor to reduce at higher frequencies, thus increasing the effective resistance of the conductor. At standstill, the rotor frequency is high since it equals to the stator frequency, whereas at a normal operating speed, it is low at about 1 3 Hz in a 60-Hz motor. As a result, the effective resistance of the rotor circuit decreases as the operating speed increases. Besides the skin effect, the rotor conductor can be properly designed to have a higher resistance at the starting. In this case, the rotor conductor is made in the form of a deep-bar as shown in Fig. 3.33B. In the current-carrying deep-bar, the leakage flux surrounding the bottom part of the deep-bar conductor is greater (A) (B) Comparison of rotor bars. (A) Normal rotor and (B) deep-bar rotor. Torque speed characteristics for different values of the rotor resistance. than that of the upper part. The leakage flux lines are shown by the dotted lines on the right of Fig. 3.33B. The leakage inductance of the conductor elements at the bottom part is larger than that of the conductor elements at the upper part. Therefore the AC current in the upper part of the conductor will flow more easily than in the bottom part with a large leakage reactance. This nonuniform current distribution makes the effective resistance of the bar to increase. Since the leakage reactance is proportional to the frequency, this nonuniform current distribution is more pronounced at the starting of a high rotor frequency" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0000285_iccme.2013.6548271-Figure2-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0000285_iccme.2013.6548271-Figure2-1.png", "caption": "Fig. 2 Distribution of Robot Coordinate Systems", "texts": [ " So the pose parameters of such parallel robot could be measured successfully using optical position method. 369978-1-4673-2971-2/13/$31.00 \u00a92013 IEEE Proceedings of 2013 ICME International Conference on Complex Medical Engineering May 25 - 28, Beijing, China The 6-UPS parallel robot is composed of six identical linear actuators which are connected between two rigid platforms: the base platform and the mobile platform. Its base platform is mounted on a support and a certain drill is attached to the mobile platform. It belongs to the rod-length-changing parallel robot. As shown in Fig. 2, mobile coordinate system and base coordinate system are respectively fixed on the mobile platform and the base platform. The pose relationship between these two coordinate systems can be expressed as [ , , , , , ]T x y zP P P . At the initial process of the surgical parallel robot, each linear actuator moves along the slide way, and stops at the original position determined by the photoelectric switches. When the parallel robot is in its original position, the mobile surface coordinate system is regarded as a reference to define the base coordinate system and the mobile coordinate system", " When the end-effector is at a random pose, using the Polaris Spectra System to track the passive markers, the homogeneous transformation matrix relating the mobile platform coordinate system to the base coordinate system can be expressed as: 1 ' '( )R R M T O O M T O OT T T T T (3) Finally, the pose parameters 0 0 0( , , , , , )x y z of the endeffector can be calculated by the following equation: 11 12 13 14 21 22 23 24 31 32 33 34 41 42 43 44 ( , , ) ( , ) ( , ) ( , ) 0 0 0 1 R O x y z x y z T T p p p R z R y R x t t t tc c c s s s c c s c s s p t t t ts c s s s c c s s c c s p t t t ts c s c c p t t t t (4) That is: 0 0 0, ,x y zx p y p z p 32 33 31 21 11atan2( , ), arcsin( ), atan2( , )t t t t t where ( ~ ) ( ~ ) 2 2 ( ~ ) . In order to evaluate the effectiveness of the kinematic calibration method, the paper defines some relevant concepts, including position deviation and orientation deviation, as follows. As seen in Fig. 2, the pose matrix of the tool end D is expressed by T TX P in the optical coordinate system { }M . In this expression, T[ ]P x y z is its position matrix and T[ ] is its orientation matrix. In the same way, define the expression T T d d dX P , T[ ]d d d dP x y z and T[ ]d d d d , to represent the given pose matrix of that in the optical coordinate system { }M . The measured pose matrix of that through optical tracking system in the optical coordinate system { }M is T T r r rX P , T[ ]r r r rP x y z and T[ ]r r r r " ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0002302_b978-0-12-803581-8.10351-0-Figure43-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0002302_b978-0-12-803581-8.10351-0-Figure43-1.png", "caption": "Fig. 43 Clip interface assembly.", "texts": [ " This design was used for both right and left sides; however, since the stud and collar attachment to the rear clip had to be slightly above the bottom of the chassis, the bottom brackets had to be designed separately. They were made to carry loads from the rear clip to the front monocoque sidewalls, bottom surface of the monocoque shell and the back wall of the monocoque. These lower interface brackets are shown in Figs. 41 and 42. The four clip interface brackets were placed at the far corners of the chassis in order to gain as much torsional stiffness as possible within the given geometry. This configuration can be seen in the chassis assembly image shown in Fig. 43. By placing the clip interface brackets close to the corners of the chassis, reasonable load transfer from the side walls to the top and bottom shells is attained. The design of the brackets was such that the main flat bonding area of the bracket was the same thickness as the Nomex honeycomb core that the brackets replaced. This allows a smooth transition of the carbon fiber/epoxy facesheets from the core to the brackets to decrease stress concentrations at the core\u2013bracket interface. In addition to the flat bonding area that relies on the shear strength of the adhesive, mechanically locking pockets were designed into the brackets to increase pull-out strength" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0003066_hnicem.2017.8269562-Figure7-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0003066_hnicem.2017.8269562-Figure7-1.png", "caption": "Fig. 7. Designed fuzzy logic surface plot", "texts": [ " The fuzzy rules are chosen by the authors and are summarized as follows: If \u03b5 is VeryLow & \u03b4 is Low then Abort If \u03b5 is Low & \u03b4 is Low then Abort If \u03b5 is Average & \u03b4 is Low then Abort If \u03b5 is High & \u03b4 is Low then Abort If \u03b5 is VeryHigh & \u03b4 is Low then Abort If \u03b5 is VeryLow & \u03b4 is Medium then Abort If \u03b5 is Low & \u03b4 is Medium then Abort If \u03b5 is Average & \u03b4 is Medium then Continue If \u03b5 is High & \u03b4 is Medium then Continue If \u03b5 is VeryHigh & \u03b4 is Medium then Continue If \u03b5 is VeryLow & \u03b4 is High then Abort If \u03b5 is Low & \u03b4 is High then Abort If \u03b5 is Average & \u03b4 is High then Continue If \u03b5 is High & \u03b4 is High then Continue If \u03b5 is VeryHigh & \u03b4 is High then Continue The fuzzy membership functions for Battery Percentage and Degree of Ability to Hover are shown in Figure 4 and Figure 5 respectively. Figure 6 shows the fuzzy membership functions of the output, that is whether to abort or to continue the mission. MAV is small yet it is a good aerial vehicle to control because of its manueverability when compared to larger aerial vehicles. With this, we can create a robust controller that determines if a certain MAV has the capacity to finish a certain mission depending on the different constraints. Whereas in this case, these constraints are the \u03b5 and \u03b4. Figure 7 shows the surface plot of the Battery Percentage and Degree of Ability to Hover againts the output. Using the Matlab\u2019s Fuzzy Logic Toolbox, the highest value of S is 0.769. As example, if we set \u03b5 to 50 and \u03b4 to 0.5 we can get S equal to 0.769 as shown in Figure 8. This is the same value when set \u03b5 to 100 and \u03b4 to 1 which means that the MAV will still continue the desired mission. Setting \u03b5 to 30.7 and \u03b4 to 0.414, the value of S will be 0.27 which means that the FTC controller tells the MAV to abort the mission" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0003033_coase.2017.8256301-Figure5-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0003033_coase.2017.8256301-Figure5-1.png", "caption": "Fig. 5: Schematic of Pgoal, Pobstacles calculation. (a) Add Gaussian noise at the state change point (needle rotation or needle lateral motion point). (b) Check if the needle tip reaches the target position and if the path collides with obstacles. (c) Repeate (a) and (b) in enough times, then calculate Pgoal and Pobstacles.", "texts": [ " \u03b11 is negative and \u03b12 is positive, beacause cost needs to be minimized. Both measurement error and external disturbance in the operation will cause needle to offset from the desired trajectory. Therefore, Pgoal and Pobstacles are used as the performance criteria of the first two attributes, rather than the absolute value of position error and the distance between the needle trajectory and obstacles. In this paper, Pgoal and Pobstacles are calculated by adding Gaussian noise at the point where the needle state changes take place, shown in Fig. 5. Needle insertion, needle rotation, and needle lateral motion are the three mainly available inputs to control needle[23]. Needle rotation and needle lateral motion could cause more tissue trauma to change the needle tip direction in surgery. In addition, needle rotation and needle lateral motion make it more difficult to realize precise control and needle deformation prediction. Thus, path realization difficulty coefficient H is defined. H increases with the increase of needle rotation number and needle lateral motion displacement" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0001056_icelmach.2010.5608084-Figure3-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0001056_icelmach.2010.5608084-Figure3-1.png", "caption": "Fig. 3. Hysteresis loop L1 before overexcitation and loop L\u2032 1 after overexcitation.", "texts": [ "2 shows a terminal voltage pattern with the shortduration overexcitation and the changes of the input current, 978-1-4244-4175-4/10/$25.00 \u00a92010 IEEE efficiency and power factor. The short-duration overexcitation means that the terminal voltage Vi applied to the hysteresis motor running at synchronous speed is continuously increased up to nVi (n > 1) and then continuously decreased to Vi. The ratio n is called an overexcitation factor. To simplify discussion, we consider a case when the load is always synchronous pull-out torque and the motor runs at synchronous speed. Fig.3 shows the hysteresis loop in the hysteresis ring before and after overexitation [3]. Loops L1 and L\u2032 1 show the B-H loop before and after it, respectively, where a value of n is 2 and Bm is the maximum flux density in the hysteresis ring. It is clear from the figure that loop L\u2032 1 has a larger area than loop L1. The torque of the hysteresis motor is represented by [3] T = K \u222e BdH (1) where K is a constant, B and H are flux density and field intensity at an arbitrary point inside the rotor ring, \u222e BdH is an area of the hysteresis loop" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0002163_s10483-017-2182-6-Figure1-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0002163_s10483-017-2182-6-Figure1-1.png", "caption": "Fig. 1 Principal coordinates (x(s), y(s), z(s)) and Frenet frame (T (s), N(s), B(s)) along curve R", "texts": [ " Using the simplified equation from the Euler-Lagrange equations, in Section 4, we discuss the effect of the intrinsic twisting on the twisted ribbons. Finally, some conclusions are given. In our earlier works[37\u201340], the polymer chain was modelled as a thin elastic rod. Assume that the central axis curve R of the rod is a smooth inextensible curve in the three-dimensional Euclidean space[32\u201333,35]. Then, the geometrical configuration of the rod is determined by the curvature \u03ba(s), the torsion \u03c4(s) of R, and the twisting angle \u03c7 between the Frenet frame and the reference frame on the cross section (see Fig. 1). The orthonormal Frenet frame on R is denoted by (T (s), N(s), B(s)), where T , N , and B are, respectively, the unit tangent, normal, and binormal vectors, respectively. Let the position vector of a point on R be represented by r(s) = (x(s), y(s), z(s)), where s \u2208 (a, b) is the arc-length parameter. With the differential geometry theory[41], we can write the curvature and torsion of a curve as follows: \u03ba = |r\u2032(s) \u00d7 r\u2032\u2032(s)| |r\u2032(s)| 3 , \u03c4 = (r\u2032(s), r\u2032\u2032(s), r\u2032\u2032\u2032(s)) |r\u2032(s) \u00d7 r\u2032\u2032(s)| 2 , (1) where r\u2032(s), r\u2032\u2032(s), and r\u2032\u2032\u2032(s) are the first-, second-, and third-order derivatives of r(s) with respect to the parameter s, respectively, and the symbol (,) denotes the triple product" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0003275_b978-0-12-814062-8.00014-5-Figure12.1-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0003275_b978-0-12-814062-8.00014-5-Figure12.1-1.png", "caption": "Figure 12.1 Scheme of the MMC fabrication via the SLM process. SLM, Selective laser melting; MMC, metal matrix composites.", "texts": [ " In the experiments, nano titanium carbide powders\u2014TiC (991%, 40 60 nm, cubic), titanium diboride-TiB2 (981%, 2 12 \u03bcm) were used, which were both produced (US Research Nanomaterials Inc., Houston, United States) and nano tungsten carbide 2 WC (991%, 50 80 nm, hex) produced by Plasmotherm (Moscow, Russia). The distribution of all the aforementioned powders by size was analyzed by means of an optical granulometer ALPAGA 500NANO (OCCHIO s.a., Belgium). A unique experimental-technological setup for the PBF process of powder FG MMCs based on the nickel alloy NiCrBSi matrix was constructed in the Samara branch of the LPI (Fig. 12.1) and included [32,41]: two lasers on YAG: Nd13 (Kwant-60, radiation wavelength 1.064 \u03bcm, power up to 25 W), and ytterbium fiber laser LK-100-B (radiation wavelength 1.07 \u03bcm, power up to 100 W), operating in a continuous mode independently from each other; own deflectors for each wavelength for realizing the laser influence (LI) scanning over the powder surface; soft process management via a personal computer; interchangeable focusing lenses; mechanism for delivering and leveling several types of powder mixtures simultaneously; cylindrical platform moving in the vertical direction, on which a 3D parts are layerwise fabricated; and in situ diagnostics of the SLM process. The procedure for the FGS creation and gradient 3D parts, based on the MMC with the NiCrBSi alloy matrix, is shown in Fig. 12.1 and was developed before [29,32,37]. The following analysis techniques were used in the framework of this study. The composition and crystal structure of the MMC submicron and nanoscale structures were determined from X-ray diffraction (XRD) on a DRON-3M diffractometer (Cu K\u03b1 radiation). The phase composition of the samples was determined using the Xray database PDF2, 1999 release, and the computer program\u2014SearchMatch ver. 3.010. The quantitative analysis of the diffractograms was carried out using the Rietveld program PowderCell 2", "2A C shows the appearance of the 3D cubes from the NiCrBSi alloy after the SLM in argon: (A) the power of the LI changed from 95 to 50 W from the substrate to the top, but without heating the platform; (B) the regime did not include heating of the platform, the power of LI was constant; (C) the regime included heating of the platform Tp5 300 C. Finally, Fig. 12.3 shows the appearance of gradient 3D cubes from the MMC based on the nickel alloy NiCrBSi in which the content of the nanoceramic doping additive from 5% by volume at the base, 10% by volume, increases from platform to top in the middle and further to 15% by volume, according to the scheme shown in Fig. 12.1. As noted earlier, in this study we worked with: (1) TiC; (2) TiB2; and (3) WC nanoceramics. It is clearly visible that structure of all 3D parts received is far from perfect. There is a significant porosity and roughness of the fused structure without additional HIP (result will be presented in a separate contribution), the shape of 3D parts has significant deviations from the claimed (cube, the base area is 53 5 mm2). In the sample of Fig. 12.2B, there is a significant shrinkage (the regime without heating of the platform)" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0003711_itec.2019.8790540-Figure1-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0003711_itec.2019.8790540-Figure1-1.png", "caption": "Fig. 1b: 2D FEA Model of 48 Slots 8 Pole 2010 Prius IPM Motor \u2013 Reconstructed Full Motor Model", "texts": [ " Afterward, a customized user scripting program is run on top of the FEA tool to iteratively search at each rotor position over 1 electric cycle the corresponding motor current to yield the desired torque with minimum ripples. The current harmonics are then derived for all the data points of the maximum torque-speed curve and can be used as look-up tables for the motor drive controller to reduce the IPM torque ripples at different torque-speed operating points. The impacts of reducing the motor torque ripples with current harmonic injection are assessed through FEA simulations on the motor efficiency by comparing the motor losses. A. Toyota Prius 2010 IPM Motor Characteristics The motor model as shown in Fig. 1 is based on the 2010 Prius IPM motor [1] with 48 slots, 8 poles and V-shaped interior NdFeB magnets. From [1], the 2010 Prius IPM motor and drive have the main following characteristics: 650 V DC bus voltage, 177 max A rms, 60 kW peak power, 207 N.m peak torque, 13500 rpm max speed. B. FEA Model A 2D FEA model with rotating motion and external circuit connection to represent the motor windings and current power supply is used. First, we did a no load simulation at 3000 rpm to compute the motor cogging torque and back emf" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0000649_1.4001727-Figure2-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0000649_1.4001727-Figure2-1.png", "caption": "Fig. 2 Three-dimensional Reuleaux\u2019s method", "texts": [ " The angle of rotation will be the angle between the ine connecting the pole to any point of the body at the first posiion and the line connecting the pole to the same point after the otation. Reuleaux\u2019s method can be used for the kinematic regisration problem 14 . Three-Dimensional Generalization of Reuleaux\u2019s ethod Let La= la , l\u0304a and Lb= lb , l\u0304b be a pair of lines given in their l\u00fccker coordinates that belong to the rigid body before a dis- lacement. The homologous lines La = la , l\u0304a and Lb = lb , l\u0304b corespond to the same lines after the rigid body is displaced to a new onfiguration. A line can be defined by two points or by the interection of two planes. Figure 2 shows the three-dimensional genralization of Reuleaux\u2019s method. For each pair of homologous ines, we find the common perpendicular of the pair, and then use q. A2 derived in the Appendix to find the Pl\u00fccker coordinates f the common perpendicular between the two lines. The common erpendicular between La and La is Sa, and between Lb and Lb is b. They are written as see Fig. 2 41011-2 / Vol. 2, NOVEMBER 2010 om: http://mechanismsrobotics.asmedigitalcollection.asme.org/ on 03/06/2 Sa = sa, s\u0304a = La La Sb = sb, s\u0304b = Lb Lb Now, we find the midline of each pair, which is the line that is at the half distance between the two homologous lines and is half the angle between them. Midlines can be represented as Ma = ma,m\u0304a = 1 2 La + La Mb = mb,m\u0304b = 1 2 Lb + Lb Next, we find the lines perpendicular to the midlines Ma and Mb, as well as the common perpendiculars Sa and Sb see Fig. 2 . Again, Eq. A2 is used to find the Pl\u00fccker coordinates of these lines Ca = ca, c\u0304a = Ma Sa Cb = cb, c\u0304b = Mb Sb The lines Ca and Cb correspond to the bisecting lines in the two-dimensional version of Reuleaux\u2019s method. The line coordinates of the screw axis that represents the motion from the first to the second position of the given lines will be the normalized common perpendicular between Ca and Cb, namely S = Ca Cb = s, s\u0304 1 The screw parameters of a helical motion consist of a line and a pitch" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0001347_is3c.2012.79-Figure1-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0001347_is3c.2012.79-Figure1-1.png", "caption": "Fig. 1. A three dimensional planar redundant manipulator configuration", "texts": [ " The projection matrix onto the null space of J can be presented as: TT NNNNP 1)( \u2212= (8) Thus, equation (7) becomes: HkNN TT \u2207=\u03b8 (9) Then adopt the pseudoinverse of NT to equation (9), and it becomes: vNNIHNkN TT )( ++ \u2212+\u2207=\u03b8 (10) Where 1)( \u2212+ = NNNN T and v is an arbitrary vector that controls the direction and speed of sampling the configurations of the constraint locus. II. KINEMATICS OF THREE DIMENSIONAL PLANAR Consider the six degrees of freedom three dimensional planar manipulator shown in Fig. 1, where li denotes the i-th link, i denotes the i-th joint angle, and (x,y,z) is the target point. To find the position coordinates (x,y,z) the following equations can be used: (11) As long as the manipulability measure M of manipulator is based on the Jacobian matrix J, the Jacobian matrix of the manipulator is calculated as: \u2202 \u2202 \u2202 \u2202 \u2202 \u2202 \u2202 \u2202 \u2202 \u2202 \u2202 \u2202 \u2202 \u2202 \u2202 \u2202 \u2202 \u2202 \u2202 \u2202 \u2202 \u2202 \u2202 \u2202 \u2202 \u2202 \u2202 \u2202 \u2202 \u2202 \u2202 \u2202 \u2202 \u2202 \u2202 \u2202 = 654321 654321 654321 \u03b8\u03b8\u03b8\u03b8\u03b8\u03b8 \u03b8\u03b8\u03b8\u03b8\u03b8\u03b8 \u03b8\u03b8\u03b8\u03b8\u03b8\u03b8 tptptptptptp tptptptptptp tptptptptptp zzzzzz yyyyyy xxxxxx J (14) The manipulability measure of the manipulator can be calculated using equations (4) and (14)" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0000836_detc2013-13594-Figure2-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0000836_detc2013-13594-Figure2-1.png", "caption": "Figure 2. 14-DOF HYPOID GEARED ROTOR SYSTEM MODEL.", "texts": [ " Finally, the mesh stiffness is defined by )./( 0eeFk Lm (5) Above formulation is applied to each contact interface, thus the multi-point mesh parameters in one mesh cycle can be obtain and applied to the dynamic model. The number of contact tooth pairs at any given time is a function of load, pinion roll angle and tooth geometry. The coupled multi-body dynamic and vibration model, proposed by Tao [4], is employed in this paper. The pinion, gear, engine and load component are considered as rigid rotors, as shown in Figure 2. The engine and load component each has only one rotational coordinate. The shaft-bearing assembly stiffness is used to represent the flexibility at both pinion and Downloaded From: http://proceedings.asmedigitalcollection.asme.org/ on 01/31/2016 Terms of Use: http://www.asme.org/about-asme/terms-of-use 3 Copyright \u00a9 2013 by ASME gear lumped support. Multi-point mesh coupling is applied between the engaging gear teeth. The equations of motion in matrix form can be written as }{}]{[}]{[}]{[}]{[}]{[ FxGxGxKxCxM a (6) where T Lgzgygxggg pzpypxpppD zyx zyxx },,,,,, ," ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0003762_1.5111940-Figure1-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0003762_1.5111940-Figure1-1.png", "caption": "Fig. 1. A representation of the interactions between one particle on the rim of the wheel and the other particles of the wheel. The particle in the center is i; the one to the left is i 1 and the one to the right i\u00fe 1. D~rc;i is the vector from the central particle to the ith particle on the rim. The unit vector n\u0302i bisects the exterior angle between D~r i 1 and D~r i.", "texts": [ " However, this is sufficient to investigate the interactions of a wheel decelerating along its direction of motion. The wheel is modeled as a set of point masses (particles) connected to each other with damped springs. A single particle mc at the center represents the axle of the wheel. In the case of a bicycle wheel, for example, this would stand in for the axle, spokes, and rim. The outside of the wheel (the tire) is composed of n particles each of mass m, initially arrayed in a circle centered on the central mass. Figure 1 shows the connections between one of the particles on the tire and the other components of the wheel. This discussion will use i to index the n particles around the rim of the wheel, with i running from 0 to n 1. Each particle i connects to particles i 1 and i\u00fe 1 via a spring with spring constant k and equilibrium length l0 and a linear damper with damping constant b; the connections between all of the rim particles are identical. Particle n 1 connects in this manner to particles n 2 and 0. D~ri is the twodimensional displacement vector from particle i to particle i\u00fe 1 (or to particle 0 in the case of i\u00bc n 1); Dr\u0302 i is the unit vector in the same direction. Each rim particle is also connected to the central particle mc by a spring with spring constant kc (c for \u201ccentral\u201d) and equilibrium length lc0 and a linear damper with constant bc. D~rc;i is the two-dimensional displacement vector from the central particle to particle i. We define n\u0302i as the unit vector normal to the wheel. This direction bisects the angle exterior to the wheel between D~ri 1 and D~ri. /i is the angle between D~ri 1 and D~ri, with a positive angle shown in Fig. 1. Figure 2 is a free-body diagram showing all of the internal forces (i.e., gravity and excluding ground interactions) acting on one of the rim particles. ~Fi is the force of mass i\u00fe 1 pulling or pushing on mass i and is 721 Am. J. Phys., Vol. 87, No. 9, September 2019 Robert Knop 721 ~Fi \u00bc k \u00f0Dri l0\u00de \u00fe b \u00f0D _~r i Dr\u0302 i\u00de h i Dr\u0302 i: (1) Notice that the damping only considers the component of the relative velocity of the two particles along the direction of their separation. That is, there will only be damping if the distance between the two point masses is changing", " A small but nontrivial fraction is dissipated in the ground damping springs (5%) and in work that the pressure force does on the wheel (8%). The work from pressure would go into the temperature of the gas in the wheel, but this simulation does not include any thermodynamics. It is interesting that the force doing most of the work is different from the external force providing the most impulse. The overall picture of rolling friction in this work is consistent with modern descriptions such as those that appear in Fig. 2 of Ref. 7, Fig. 1 of Ref. 8, and Fig. 1(b) of Ref. 9. That is, there are three external forces on the wheel. Gravity acts downwards from the center of mass. The normal force acts upwards, but at an effective contact point that is forward (in the direction of motion) of the point on the ground directly below the center of mass. Static friction acts backwards; it is the force that slows down the linear speed the wheel, but its torque would tend to increase the rotation rate. The torque from the offset normal force is large enough to counteract the torque from static friction and to slow down the rotation at the rate necessary to maintain rolling without slipping" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0001600_tasc.2010.2100801-Figure3-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0001600_tasc.2010.2100801-Figure3-1.png", "caption": "Fig. 3. Coil tilt about the horizontal plane (rotation around the Y axis).", "texts": [ " The technique is based on numerical reconstruction of possible geometrical distortions and misalignments of the coil using measured data on a spatial coil field. The technique is quite universal and can be applied to any coils. Standard distortions of the PF1 coil related to the accepted manufacture/assembly tolerances include [4]: 1) Linear shift along the X and Y axes, , (Fig. 2). 2) Tilt about the horizontal plane, ; at low tilt angles , the coil centerline is tilted with , where R is the coil radius (see Fig. 3). 1051-8223/$26.00 \u00a9 2011 IEEE 3) In-plane ellipticity, , , that is a deformation of the coil shape in the plane XY defined in terms of two elliptic half-axes , and an angle of rotation around Z-axis . Fig. 4 illustrates ellipticity at , , . 4) Vertical displacement of the coil centerline, or so called warping, occurred at winding, characterized by an amplitude and angle , where corresponds to the maximum deviation of the coil centerline relative to the horizontal plane XY; is the toroidal angle corresponding to this maximum deviation, ", " This value is artificially introduced with the use of an additional ferromagnetic element in order to make allowance for effect of steel structures of the building. The dummy cross-section was 32 mm 7.5 mm. The coordinate system X, Y, Z was related to the measuring bench so that the vertical axis Z coincided with the rotation axis of the measuring unit. The coordinates of the coil center in the plane XY were measured with a ruler graduated in millimeters. The maximum tilt of the coil edges relative to the horizontal plane was as low as 1 mm. The measured horizontal components Nx and Ny (see Fig. 3) of the unit tilt vector were found to be close zero, the maximum measurement error was 0.0025. The main vertical component Nz was evaluated as . The measured parameters are summarized in Table I. In this table Rc\u2014average coil radius; Xc, Yc, Zc\u2013initial coordinates of the coil center; Nx, Ny\u2013components of the unit tilt vector. When the measured data were processed, the terrestrial field value was subtracted from the readouts. The radar-type diagram in Fig. 8 illustrates distribution of the coil field for the reference position when the coil center coincides with the rotation axis of the measuring unit" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0000269_s12206-012-0712-0-Figure1-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0000269_s12206-012-0712-0-Figure1-1.png", "caption": "Fig. 1. Schematic view of undeformed rough bodies.", "texts": [ " Then, the cutoff frequency effects on the rough point contact and on the rough flat surface contact will be investigated by using the CGM and its improved form. Comparison will also be made between the contact modes. Finally, conclusions will be drawn. 2. Problem formulation and solution method *Corresponding author. Tel.: +86 23 6511482, Fax.: +86 23 65106195 E-mail address: fmmeng@xjtu.edu.cn \u2020 Recommended by Editor Sung-Lim Ko \u00a9 KSME & Springer 2012 Consider a contact where a upper body (body 1) is pushed down on a upper body (body 1) by a normal force F , as shown in Fig. 1. The Cartesian coordinate system is attached with the x- and y-axes aligned with the contacting surface of body 2 and the z-axis normal to the surface. In Fig. 1, the symbol h denotes the average separation between the contacting surfaces of bodies 1 and 2 which have the Poisson\u2019s ratio iv , Young's modulus iE , and three-dimensional surface roughness ( , )iz x y measured from the profile mean line of body i, where the subscript i (i = 1, 2) refers to body 1 or 2. This contact problem (Fig. 1) can be solved by using several technologies, as stated in the introduction. In the present study, the contact is solved by the CGM and its improved form, since the computational results of the two methods can be used for analyzing interaction between asperities on rough surfaces, which is often excluded in the foresaid statistical contact models or fractal contact theories. Here, the rough point contact problem is considered first. In this case, body i is assumed to be sphere i (i = 1, 2). Moreover, for the sake of simplification, spheres 1 and 2 are assumed to have an equivalent Poisson\u2019s ratio v , equivalent Young's modulus E , composite surface roughness ( , )Z x y , and equivalent radius R superimposed with ( , )", " Therefore, to obtain the elastic half-space solutions of the point contact, the following assumptions are made: (1) Each contacting surface is assumed to be very large, compared with actual contact area occurring only at the tips of asperities on the surface. (2) The real contact area is a small fraction of the contacting surface, which ensures that the analysis results for point contact will not be affected by the actual boundaries of the contacting bodies. (3) Only the effect of the normal force F along the zdirection is considered on the contacting body 1 shown in Fig. 1, since the surface slopes are small. (4) Only a simple elastic-perfectly plastic deformation model is adopted so that the contact pressure is limited to plasticity limit value. (5) No relative sliding takes place between the contacting bodies along the x or y -direction to simplify the contact analysis. Note that for flat surface contact analyses, only assumptions 3-5 are used. According to the above assumptions, in the rough point contact case, the Boussinesq\u2019s elastic theory can be employed to solve the composite deformation of bodies (i" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0000165_jjap.49.04dk16-Figure4-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0000165_jjap.49.04dk16-Figure4-1.png", "caption": "Fig. 4. (Color online) Layers constituting a membrane switch: (a) fixed layer, (b) flexible layer, (c) spacer layer, and (d) entire switch array. The size of one pixel is 7 mm2.", "texts": [ "jp Japanese Journal of Applied Physics 49 (2010) 04DK16 REGULAR PAPER 04DK16-1 # 2010 The Japan Society of Applied Physics layer was restored by elastic motion to the off state, thereby opening signal contacts. The EPD sheet was driven by connecting the fixed contact of each signal electrode to the pixel electrode of EPD via a blind through-hole (BTH). The detailed composition of the membrane switch is shown in Fig. 3. The switch has a relatively simple structure consisting of four layers; a fixed layer (layer A), a wall spacer (layer B), a flexible layer (layer C) and an insulating layer (SP1200). Layers A and C were prepared by using flexible copper clad lamination films (FCCL). Figure 4 shows an actual image of all four layers: the fixed layer covered with the isolation layer (a), the flexible layer (b), the spacer layer observed as a frame region around the switch (c), and the entire array of 16 16 membrane switches (d). The isolation layer over the fixed layer, which has an opening at the contact position of the signal electrodes, is a transparent 1-mm-thick foil, and is barely observable in Fig. 4(a). It is employed to avoid the shortcircuiting of the electrostatic electrodes. The isolation layer is also effective for lowering driving voltage by increasing the effective dielectric constant between the electrodes. The spacer layer is about 25 mm thick, which defines the air gap between the electrodes. It also defines the perimeter of one pixel. The electrostatic electrodes on the flexible and fixed layers are connected in rows and columns, receptivity, to a driving circuit for addressing individual pixels" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0003807_sahcn.2019.8824932-Figure1-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0003807_sahcn.2019.8824932-Figure1-1.png", "caption": "Fig. 1. Fleet of three UAVs in a triangular formation", "texts": [ " A state corresponds to the situation of the environment as perceived by the UAV. In our context, we assume that only the following information is available at the followers: the current view of the RSSI value between neighboring UAVs, a partial history of this information and the previous chosen actions. Then, we define a state as a tuple composed by (i) the RSSI between the UAVs at the previous time interval, (ii) the last action taken by the UAVs and (iii) the RSSIs at the current time interval. In the case of a triangular formation (see Figure 1) , the RSSI values needed for the algorithm are those between the leader and the first follower, the leader and the second follower, and between the two followers. Each transition from a state to another provides a reward which can be positive or negative. The algorithm tries to maximize the sum of the rewards the agent is receiving over the training period. The reward value is determined by comparing the position of the follower UAVs, obtained after the action, and the initial formation. For each iteration, the smaller the distance between the new positions and the correct positions, the higher the reward" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0001838_ma302559w-Figure2-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0001838_ma302559w-Figure2-1.png", "caption": "Figure 2. Schematic diagram of the shear deformation device for film specimens used in this work. The direction of the incident X-ray beam is set parallel to either the OZ-axis (a) and the OY-axis (b). Part (a) shows the cross section normal to the OZ-axis of the shear cell. The OX-, OY-, and OZ-axis are set parallel to shear direction, velocity gradient direction, and neutral (or vorticity) direction.", "texts": [ "1021/ma302559w | Macromolecules 2013, 46, 1549\u22121562 al.22 Okamoto et al. and Shin et al. carried out the strain-phaseresolved DSAXS experiments simultaneously with stress measurements. Almdal et al.19 studied the influence of LAOS on the orientation of the BCC-sphere in the neat poly(ethylenepropylene)-block-poly(ethylethylene) (PEP\u2212PEE) using SANS where the PEP block was deuterated and had the volume fraction of 0.83. They obtained results that the (110) plane of BCC lattice was aligned parallel to the shear plane (the OXZ-plane defined later in Figure 2) and the \u27e81 \u030511\u27e9 axis is oriented parallel to the shear direction (the OX-axis in Figure 2). The scattering patterns showed that the twinned BCC structure (B/B\u2032-twin to be defined later in section IV-1) is formed with its twinning plane (112) vertical to the shear plane. Ackerson et al.23 reported a similar orientation for colloidal suspensions under shear deformation. Koppi et al.20 investigated the shear effects as a function of shear rate and T for the PEP\u2212PEE dibcp having different volume fraction of PEP of 0.25 and molecular weight by using SANS and rheological measurements under LAOS", " Figure 1a schematically shows the temperature dependence of the microdomain structure of the sample. The order\u2212order transition temperature (TOOT) between hex\u2212cyl and BCC-sphere was determined to be 183 \u00b0C, while the lattice disordering-ordering transition temperature (TLDOT) was determined to be 215 \u00b0C and the dimicellization/micellization transition temperature was estimated as high as 280 \u00b0C, using SAXS measurements.24,25 Figure 1b will be described later in section II-5. II-2. Shear Apparatus. Figure 2 shows the shear deformation device of film samples under LAOS together with the definition of the Cartesian coordinate OXYZ, where the OX-, OY-, and OZ-axis are dx.doi.org/10.1021/ma302559w | Macromolecules 2013, 46, 1549\u221215621550 taken along the shear direction, the velocity gradient direction, and the neutral (or vorticity) direction, respectively. The 2D-SAXS patterns were taken with the incident X-ray beam along either the OZ-axis (part a) or the OY-axis (part b). The device in Figure 2a shows its cross section normal to the OZ-axis. As-cast films of thickness 0.5 mm, length of ca. 7 mm, and width 5 mm along the Z-axis were stacked to make the total thickness of the sample equal to be 3 mm along the Yaxis as shown in Figure 2. The stacked samples were sandwiched between two metal plates, and LAOS was imposed on the specimen by moving one of the metal plates parallel to the other. We used two beryllium plates fixed in the brass frames for both side of the incident and scattered X-ray beam. The thickness of the beryllium plates used was 0.3 mm. 2D DSAXS patterns with the incident X-ray beam parallel to the OY-axis and the OZ-axis were taken by rotating the sample sandwiched by the shear device by 90\u00b0 around the OX-axis" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0001060_12.908589-Figure9-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0001060_12.908589-Figure9-1.png", "caption": "Fig. 9 Sketch of scanner optics \u00bbLASSY\u00ab and path of laser beam through optics", "texts": [ " Specifically at laser processes the difference of process start and the running process can be dramatically. Taking this into consideration, some specific parameters are available. For catching a glimpse of it by the operator a graphic display shows all relevant process data during the process. All settings and process data inclusive the temperature images are stored in customized folders. Profibus is standard interface to machines and devices. Hard wiring is possible as well. The scanning optics \u00bbLASSY\u00ab (Fig. 9 and 10) has been developed for industrial use. Compact size and low weight make \u00bbLASSY\u00ab a component to all kind of laser treatment machines. The scanning optics is suited for lasers in the wavelength range from 808 to 1070 nm (diode, Nd:YAG; Yb:YAG) and has no focusing component. A laser optics with a rectangular or round focal spot with minimum focal length of 300 mm minimum is necessary affront the input of the scanner optics. The spot size of that optics depends on the required hardening depth" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0000024_amr.102-104.605-Figure5-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0000024_amr.102-104.605-Figure5-1.png", "caption": "Fig. 5 Assembly drawing of three kinds of layer face milling cutter", "texts": [ " Under the conditions of all the constraints are found, the subsequent parameters change with the previous parameters, and thereby a similar new model is generated. Following this modeling ideas, the main angle of 45 \u00b0 layers face milling cutter is got, as shown in Fig. 4. According to modeling parametric design process, the face milling cutters with main angle 60\u00b0 and 75\u00b0are got, which are based on layer face milling cutter with 45\u00b0 main angle. Three face milling cutter assembly models are got after combining tool blade, tool folder, tool cushion and cutter body, as shown in Fig. 5. Layer Face Milling Cutter Modal Analysis. Modal is natural vibration a characteristic of mechanical structure, and each modal has a specific natural frequency, damping ratio and modal shape. Modal parameters can be obtained by calculation or test analysis, and the analysis process is modal analysis. The ultimate goal of modal analysis is to identify the modal parameters of system, which can provide basis for tool geometry optimization and vibration restrain analysis [6]. Based on this, in order to analyze the impact of the main angle of layer face milling cutter on natural frequencies, this paper carries out modal analysis based on parametric modeling" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0001059_1.4001772-Figure8-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0001059_1.4001772-Figure8-1.png", "caption": "Fig. 8 Mobile platform traversing an S shaped trajectory", "texts": [ " This example is just to familiarize the eader with the formulation. A thorough development of motion ynthesis for the mobile platform including the higher order mo- ion synthesis will be dealt with in an ensuing paper. 31015-8 / Vol. 2, AUGUST 2010 om: http://mechanismsrobotics.asmedigitalcollection.asme.org/pdfaccess. 7 Numerical Example of Motion Planning Consider a mobile platform traversing an \u201cS\u201d shaped trajectory composed of two circular arcs such that its curvature changes from concave to convex at point C, as shown in Fig. 8. In this case, the platform is always aligned with the direction of travel such that all the ICs for the velocity, acceleration, jerk, etc., are located at the center of the curvature. When the mobile platform crosses point C, the normal acceleration, jerk, etc., instantaneously switch to the opposite direction resulting in shock and motion uncertainty. Using IC based motion programming; we can remove this crossover shock and uncertainty with a dexterous platform as follows. To prevent the shock, we put a restriction on the motion, whereby, point C becomes a stationary inflection point" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0001318_med.2013.6608714-Figure1-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0001318_med.2013.6608714-Figure1-1.png", "caption": "Fig. 1. Projectile side and front views.", "texts": [ " Fin-stabilized projectiles use aerodynamic surfaces located at their aft section whereas spin-stabilized maintain a high airframe spin rate which guarantees gyroscopic dynamic stability. Even though the latter are simpler to implement, since they use the initial spin rate of the ballistic weapon, the canard actuators need to be of extremely high bandwidth since they must modulate the control signal to this high frequency. The solution to this inconvenience is the so-called dual-spin course correction fuse concept shown in Fig. 1 which uses a co-axial motor to decouple the projectile body from the guidance fuse, also containing the software and hardware needed for guidance. *This work was in part financially supported by the DGA (Delegation Ge\u0301ne\u0301rale de l\u2019Armement). 1S. Theodoulis, V. Gassmann, T. Brunner and P. Wernert are with the Guidance, Navigation & Control Department of the French-German Research Institute of Saint-Louis (ISL), 5 rue du Ge\u0301ne\u0301ral Cassagnou, 68301, Saint-Louis, France, Spilios.Thedoulis at isl.eu The feedback controller needed in order to manage the roll-channel motion of the nose works in two main modes: during the first (corresponding to the initial projectile flight phase, spanning from the launch up to its trajectory apogee), progressively decreases the nose roll rate to zero whereas during the second one (corresponding to the guidance phase and up to the impact time) it orientates the angular position of the nose to desired positions which are computed by an external guidance loop" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0000969_012001-Figure1-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0000969_012001-Figure1-1.png", "caption": "Figure 1. Perspective view of a vehicle car model.", "texts": [ " In order to solve the interaction model two main branches had been developed [21]: to solve bridge and vehicle motions separately (iterative schemes for stablishing compatibility are needed) [3, 6, 7, 9, 11, 13, 21]; and to solve a single coupled equations set [2, 5, 8, 19, 22, 23]. A three-dimensional high speed railway vehicle has been modeled using rigid bodies that are associated to masses and inertias. Each car is independent of other cars, due to that, different cars do not share any degree of freedom. Five different rigid bodies are used in a single vehicle definition: a car body, two bogies and four wheelsets (Figure 1 and 2). Each body has six degrees of freedom (three displacements and three rotations) but for all of them, longitudinal displacements are impossed. Furthermore, rotations of wheelsets around his main axle are constrained. Thus, each car has 31 degrees of freedom. The bodies are linked together by linear springs and dampers, which simulate the behavior of the two suspension levels of passengers railway vehicles. Vertical and lateral springs and dampers are used for modelling primary and secondary suspensions" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0002046_nuicone.2013.6780154-Figure1-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0002046_nuicone.2013.6780154-Figure1-1.png", "caption": "Fig: 1 Six- Phase PMSM Motor (Stator winding)", "texts": [], "surrounding_texts": [ "1 Abstract-- this paper presents a complete modeling and performance analysis of six phase Permanent Magnet Synchronous Motor (PMSM) drive system supplied by two voltage source inverters. The Six phase PMSM drive system is reliable, reduces the stator current per phase, and posses high degree of freedom as compared to conventional three phase drives. Vector control technique is used with two Pulse width modulated (PWM) inverters. The proposed model is simulated under various load conditions such as no load, steady load and dynamic load. Computer simulations obtained under various operating conditions have been presented and found satisfactory. Keywords\u2014 Multiphase, Permanent magnet synchronous motor (PMSM), PWM, Inverter.\nI. INTRODUCTION\nMultiphase variable speed drive has received growing interest since the second half of the 1990s because of several advantages of multiphase and superiority of PMSM drive system over other drive system. This growing interest is cause by the fact that this machine can provide notable improvements in various aspects of performance as compared to conventional three phase motor and six phase induction Motors. Historical technical reasons that required adopting the multiphase drive solution instead of three phase are listed below [1,2,3,4,5,6]:-\n1) For a given motor output power, Multiphase variable speed drive reduces the stator current per phase. 2) The use of more than three phases offers an improved reliability. 3) Pulsating torques produced by time harmonic components in the excitation waveform is reduced by Multiphase machines. 4) Fault tolerant drives 5) Higher degree of freedom The main application areas of multiphase machines especially in motor drives are ship propulsion, traction (including Electric and HEV) and the concept of More-electric aircraft. Other suitable applications are locomotive traction, aerospace and high power Applications.\nII. THREE PHASE DRIVES vs. SIX PHASE DRIVES\nAs stated above at point no.1, the use of multiphase drive in place of three phase drive, helped to overcome the problems related to high power applications by using current limited devices. For a given motor power, an increase in phase number determines a reduction in per phase power, enabling the use of smaller power electronic devices for every inverter leg, without increasing the per phase voltage. This is still a solution adopted for high power applications, such as transport and ship propulsion drives. In fact, commercial evaluation of large multiphase machines for ship propulsion is in progress. The improved reliability features of multiphase drives, listed at point no.2, enable their use also in faulty conditions; in fact if one phase of a multiphase machine becomes open circuit, the machine is able to self start and to run with only a different rating, which depends on post fault strategy of control and in the number of the phases. In the three phase case, the loss of one phase determines an important de rating of the machine while it is running. Furthermore the machine is not self starting and, for this purpose, it requires external means. Finally, the advantages derived from statement at point no.3 were very important in the 60s, when three phase inverter fed ac drives operate with six step mode. Time harmonic of voltages and currents introduced by this operation mode produced low frequency torque ripple, leading to difficulties on speed control and noise production. Since in a n phase machine torque pulsations are caused by supply time harmonics of the order 2n\u00b11, which result in torque ripple harmonic 2n times higher than the supply frequency, an increase in the number of phases seem the best solution to the problem. This aspect of multiphase drives has lost importance since the discover of PWM of VSI, which allows the control of inverter output voltage harmonic content. As stated at point no.4.The increased number of phases in drives offers considerable benefits because of the capability to continue operation when a single or multiple phase loss occurs [1,2,3]. Three phase drive is sensitive to different kinds of faults, both in motor phase and in inverter leg. When one of these faults does occur in one phase, the drive operation has to be stopped for a non programmed maintenance schedule. The motor in faulty conditions is able to run but it is not still self starting. The cost of this can be high, which justifies the development of fault tolerant motor drive systems. On the contrary, in post fault condition, multiphase machines can continue to be operated with an asymmetrical winding structure and unbalanced excitation, producing a higher fraction of their rated torques with little pulsations when compared to the three phase machines.\nPerformance Analysis of Two Inverter Fed Six Phase PMSM Drive\nAnurag Singh Tomer Satya Prakash Dubey Associate Prof. ( Deptt. Of Electrical Engineering) Prof.( Deptt. Of Electrical Engineering)\nRCET, BHILAI (C.G.) RCET, BHILAI (C.G.)\n2013 Nirma University International Conference on Engineering (NUiCONE)\n978-1-4799-0727-4/13/$31.00 \u00a92013 IEEE", "III. SUPERIORITY OF PMSM MOTOR\nThere are variety of ac servo drives on the market contending with both the DC brush Machines and AC servo drives. Two types of permanent-magnet ac motor drives are available in the industry. These are Permanent Magnet Synchronous Motor (PMSM) drives with sinusoidal flux distribution and the brushless DC Motor (BLDC) drive with trapezoidal flux distribution [1,2]. Recent availability of high energy-density permanent magnet (PM) materials at affordable prices, continuing breakthroughs and reduction in cost of powerful fast, digital signal processors and micro-controllers combined with the noteworthy advances in semiconductor switches and modern control technologies have opened up new possibilities for permanent magnet /brushless motor drives in order to meet competitive worldwide market demands. In addition to this PMSM have several advantages over the DC brush motor and induction motor for low power application. (Listed below) PMSM have the following advantages over DC Motors [3,7,8,9,10,11]:-\n1) Less audible noise, longer life, Spark less (no fire hazard), 2) Higher speed, higher power density and smaller size, Better heat transfer. PMSM have the following advantages over Induction Motors:1) Higher efficiency, Higher power factor, Higher power density for lower than 10 KW applications, resulting in smaller size 2) Better heat transfer. 3) The above comparison shows that the PMSM are\nsuperior to the induction motor /DC motor.\nIV. MODELLING OF SIX PHASE PMSM The six phase permanent magnet synchronous machine has\ntwo identical, balanced, star connected assumed stator\nwindings. Commonly, these sets of winding can have a phase shift of 0, 30 and 60 degrees. Zero degree phase shift is\nsimilar to three phase system. Sixty degree phase shift forms\nsymmetrical arrangement and can be reduced to three phase system because two phases of different stars are always\ncollinear. Thirty degree phase shift forms unsymmetrical\narrangement which cannot be further simplified. The thirty degree phase shift arrangement is most favorable with respect\nto voltage harmonic distortion and torque pulsation.\nTherefore, thirty degree phase shift between star connections has been preferable arrangement and also in this paper. But,\nthe actual machine used for implementation in the laboratory\nhas 33.2725 degree separation between the two three phase groups [11].\nIn developing the mathematical model the following assumptions are made [11]:-\ni. The set of three-phase stator winding is symmetrical.\nii. The capacitance of all the windings can be neglected.\niii. Each of the distributed windings may be represented\nby a concentrated winding. iv. The change in the inductance of the stator windings\ndue to rotor position is sinusoidal and does not\ncontain higher harmonics. v. Hysteresis loss and eddy current losses are ignored.\nvi. The magnetic circuits are linear (not saturated) and\nthe inductance values do not depend on the current. In this study, a six-phase PMSM with two three-phase Windings are adopted where ABC winding is spatially 30 electrical degrees phase led to XYZ winding [12,13].\nThe phase voltage and flux linkage equations in the stationary reference frame for ABC winding and XYZ winding of sixphase PMSM are shown as:\ndt d IRV ABC ABCSABC \u03c6+= (1)\nMABCXYZABCABC ILIL '1211 \u03c6\u03c6 ++= (2)\ndt d IRV XYZ XYZSXYZ \u03c6+= (3)\nMXYZABCXYZXYZ ILIL '2122 \u03c6\u03c6 ++= (4)\nWhere Rs = diag [Rs, Rs, Rs]T is the stator resistance vector; VABC= [VA VB VC]T is the phase voltage vector of ABC winding; IABC= [iA iB iC]T is the current vector of ABC winding; Vxyz= [Vx Vy Vz]T is the phase voltage vector of winding; iXYZ= [ix iy iz]T is the current vector of XYZ winding; \u00d8ABC=[\u00d8A \u00d8B \u00d8C]T is the stator flux linkage vector of win ABC ding; \u00d8XYZ=[\u00d8X \u00d8Y \u00d8Z]T is the stator flux linkage vector of xyz winding; L11 is the stator inductance vector of win ABC winding;L22 is the stator inductance vector of XYZ winding; L11 and L22 are the mutual", "3 inductance vectors; \u00d8\u2019MABC\u2019 is the permanent-magnet flux linkage vector of ABC winding; \u00d8\u2019MXYZ is the permanentmagnet flux linkage vector of XYZ winding [14,15,16]. In order to control the six-phase PMSM, the following Transformation matrixes have been used to transfer the above Equations into the synchronous rotating reference frame:\n+\u2212 +\u2212 =\n2 1 2 1 2 1\n)120sin()120sin(sin )120cos()120cos(cos\n3 21 00\n00\neee\neee Tqd \u03b8\u03b8\u03b8 \u03b8\u03b8\u03b8\n(5)\n+\u2212\u2212 +\u2212\u2212 =\n2 1 2 1 2 1\n)90sin()150sin()30sin( )90cos()150cos()30cos(\n3 22 000\n000\neee\neee Tqd \u03b8\u03b8\u03b8 \u03b8\u03b8\u03b8\n(6)\nwhere Tqd1 is the transformation matrix for ABC winding; Tqd 2 is the transformation matrix for XYZ winding; e is the rotor flux angle [15]. Moreover, the machine model of a\nsix-phase PMSM can be described in synchronous rotating\nreference frame as follows\n)( 111 1 1111 PMdde q\nqqsq IL dt\ndI LIRv \u03c6\u03c9 +++= (7)\n111 1 1111 qqe d ddsd IL dt\ndI LIRv \u03c9\u2212+= (8)\n)( 222 2 2222 PMdde q\nqqsq IL dt\ndI LIRv \u03c6\u03c9 +++= (9)\n222 2 2222 qqe d ddsd IL dt\ndI LIRv \u03c9\u2212+= (10)\nre P \u03c9\u03c9 2 = (11)\nWhere v d1 and v q1 are the d-q axis voltages of ABC winding; v d 2 and v q2 are the d-q axis voltages of XYZ winding; id1 and iq1 are the d-q axis currents of ABC winding; id2 and iq2 are the d-q axis currents of XYZ winding; L d11 and L q11 are the d-q axis inductances of ABC winding; L d 22 and L q22 are the d-q axis inductances of XYZ winding; r is the rotor angular velocity; e is the electrical angular velocity; \u00d8PM is the permanent magnet flux linkage; P is the no. of pole pairs of six phase PMSM. As assumed that winding sets are identical (L\nq11 =L q22 = L q and L d11 =L d22 = L d ). Furthermore, the developed\nelectric torque e Te can be represented by the following equation:\n( ) ( )( )[ ]22112122 3 qdqdqdqqPMe IIIILLII P T +\u2212++= \u03c6 (12) However, the electromagnetic torque cannot be estimated accurately in a general case without knowledge of the currents of both winding sets and the inductance parameters that describe the magnetic coupling between them. In addition, the mechanical dynamic equation of the sixphase PMSM is:\nLr r e TB dt\nd JT ++= \u03c9\u03c9\n(13)\nWhere J is the inertia of six-phase PMSM; B is the damping\nCoefficient; TL is the load torque [14,15,16].\nV. MACHINE PARAMETER AND SIMULINK MODEL\nThe Six phase PMSM parameter is given in following table: 1 for the simulation purpose [14,15].\nTABLE: I MACHINE PARAMETER S.NO. NAME RATING 1. Nominal power Pn 25Kw 2. Nominal voltage Vn 380 volts 3. Nominal speed nn 350RPM(36.5rad/s) 4. No.of Poles 8 5. Stator Resistance Rs 0.64 ohm 6. PM flux Linkage \u00d8PM 2.04 wb\n7. Ld,Lq 24mH,31.4mH 8. Inertia J .014Nm/(rad/sec2) 9. Damping coefficient B .0124Nm/(rad/sec) The following simulation model of complete six phase PMSM drive system is simulated under different load conditions. The reference speed is set at 36.5rad/sec. PI controller is used for torque generating current component.\nFig:2 Simulink model of Six phase PMSM drive\nVI. SIMULATION RESULTS\nThe presented model of six phase PMSM drive system has been\nsimulated under various load conditions for 0.6 Seconds.\nCase I: No-Load operation:\nThe model is simulated at no-load (TL = 0) and at rated reference speed ( r = 36.5 rad/sec.). Fig.3 shows the six phase stator current, rotor speed, and Torque response of the proposed scheme. The simulation results show that at 0.075sec. Speed and Torque reaches, it\u2019s rated value." ] }, { "image_filename": "designv11_62_0002200_978-3-319-51222-8-Figure2.10-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0002200_978-3-319-51222-8-Figure2.10-1.png", "caption": "Fig. 2.10 Influence functions for M(x) and V (x)", "texts": [ "2 Influence Function for M(x) To set the stage for the influence function of M(x), we install a hinge at the source point x, spread it by one unit G \u2032 2(x\u2212) \u2212 G \u2032 2(x+) = 1 , (2.35) and formulate with both parts, GL 2 and GR 2 , Betti\u2019s theorem B (G2, w) = B (GL 2 , w)(0,x) +B (GL 2 , w)(x,l) = 0 + 0 . (2.36) The work done at the beam ends is zero and of the work done at the source point x, the interface between the two halves, only the term G \u2032 2(x\u2212) M(x) \u2212 G \u2032 2(x+) M(x) = 1 \u00b7 M(x) (2.37) is left over and so it follows that 1 \u00b7 M(x) = \u222b l 0 G2(y, x) p(y) dy . (2.38) Influence Function for V(x) In this case, we install a shear hinge (see Fig. 2.10e), which we spread by one unit, (\u22121), G3(x\u2212) \u2212 G3(x+) = 1 , (2.39) and we formulate with both parts GL 3 and GR 3 Betti\u2019s theorem B (G3, w) = B (GL 3 , w)(0,x) +B (GR 3 , w)(x,l) = 0 + 0 . (2.40) At the interface, at the source point x, all terms cancel up to G3(x\u2212) V (x) \u2212 G3(x+) V (x) = 1 \u00b7 V (x) (2.41) and so the equation W1,2 = 0 provides the result 1 \u00b7 V (x) = \u222b l 0 G3(y, x) p(y) dy . (2.42) 2.3 Influence Functions for Force Terms 89 Theoretically, you only need to know the influence function for the zero-order displacement and the influence functions for the higher-order derivatives are simply obtained by differentiating this influence function with regard to the source point x (see Fig" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0001662_vppc.2011.6043205-Figure2-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0001662_vppc.2011.6043205-Figure2-1.png", "caption": "Fig. 2. Finite Element Mesh of an Interior PM Machine", "texts": [ "(1), the PDE associated with 1D heat transfer is given as follows: c \u2202T \u2202t \u2212 k \u22022T \u2202x2 = qwind (14) The boundary temperatures representing Dirichlet boundary conditions on both ends of the winding end-turns are determined by the hot-spot temperature of the winding that is calculated by the 2D model. The order of the finite element model discussed in the prequel is then reduced in three ways, which are introduced as follows. An interior permanent magnetic (IPM) machine that is suitable for use in a HEV is used in this section for illustration purposes. Its geometries and finite element mesh are shown in Fig. 2. The method proposed is based upon the simulation of the eigenmodes of the dynamic thermal system which is in the form of state variable description. Although the full\u2013order finite-element dynamic model can have thousands of states (corresponding to the number of finite element nodes), we are mainly interested in the most heavily excited eigenmodes, as these tend to dominate the system dynamic response. We therefore consider the Jordan normal form of the system; i.e., we perform a change of basis x = E\u22121 ti, where E is a matrix whose columns consist of eigenvectors of the A matrix", " The extent of excitation of each eigenmode is evaluated by the following equation: Condi = \u2212\u03bb\u22121 i bcond,i (18) Corei = \u2212\u03bb\u22121 i bcore,i (19) where Condi and Corei is the extent of excitation of conduction and core losses of the ith eigenmode, \u03bbi = 1 \u03c4i is the eigenvalue of the ith eigenmode, bcond,i is the ith element of the column of B corresponding to the conduction losses, and bcore,i is the ith element of the column of B corresponding to the core losses. The \u2212\u03bb\u22121 i term in Eq.(18, 19) can be seen as a filter to increase the weights of slow eigenmodes, as the system response is mainly governed by the slowest eigenmodes. By setting an appropriate threshold value, the most excited eigenmodes can be chosen by calculating Eq.(18, 19). Fig. 3 shows the effect of conduction and core losses on the system of the slowest 100 eigenmodes in the stator of the IPM machine shown in Fig. 2. As can be seen from Fig. 3, a relatively small number of eigenmodes are significantly excited. Most eigenmodes are not excited at all. As the heat generation in each slot is identical under the assumptions of balanced phase currents, the symmetry of the excitation requires a tooth-slot symmetry in the temperature profile. Thus, all the eigenmodes that are not geometrically symmetric are not excited. Fig. 4 shows the first eight modes of the stator mesh, inversely ordered by time constant (i.e., Mode \u20181\u2019 corresponds to the largest time constant, Mode \u20182\u2019 the second largest, etc", " However, the minimum \u201cdynamic\u201d eigenvalue chosen in the reduced\u2013order model is relatively large, which can be simulated with a relatively large step length in an explicit method (1sec is chosen). This shows another advantage of the reduced\u2013order model over the full-order model \u2013 the reduced model can be solved by an explicit method instead of an implicit method, thus being more computationally efficient, as it avoids the need for a matrix\u2013vector solve. The full\u2013order thermal FEA model has 2458 states variables for the stator, as shown in Fig. 2. By using the model order reduction techniques presented here, the reduced-order stator model can have less than ten states while maintaining satisfying accuracy. Thus, the number of calculations is dramatically reduced. The reduced\u2013order thermal model, built in Simulink R\u00a9, is shown in Fig. 6. The stator slot hot-spot temperatures that are calculated from the full and reduced\u2013order thermal models due to a step in conduction and core losses are shown in Fig. 7. The \u201cjump\u201d at t = 0 in the reduced\u2013order solutions is due to the earlier assumption that the \u201cstatic\u201d eigenmodes in reduced\u2013 order models instantaneously converge to their quasi-steadystate values. As can be seen from the figure, the accuracy of the reduced-order model is improved by increasing the number of excited eigenmodes. The maximum relative error between full\u2013 and reduced\u2013order models is shown in Table I. As can be seen from the table, the error of the 5th-order reduced model remains less than 2% for the entire time span. The internal temperatures of the IPM machine in Fig. 2 are evaluated by the full\u2013 and reduced\u2013 order models described in this paper. The full-order FEA model has 2458 and 1466 states in the stator and rotor, respectively, while the reduced\u2013order model has only 10 and 5 states in the stator and rotor. The comparison of the two models is shown in Table II. As can be seen from the table, the computation time is reduced by over three orders of magnitude. It takes only 1.9sec to simulate a driving cycle from 0 to 1400sec for the reduced\u2013order model. Thus, the reduced model is feasible for real-time simulation" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0003085_intellisys.2017.8324247-Figure3-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0003085_intellisys.2017.8324247-Figure3-1.png", "caption": "Fig. 3. The parameters for multiple steps.", "texts": [ " Note that the work expressions shown in (4) and (5) are useful to analyze the work done by the joint\u2019s driving motors of the robot during walking. IV. TRAJECTORY GENERATION BY USING ZMP METHOD In many cases of bipedal robots, the walking locomotion is performed with a fixed hip height and the ZMP placed inside the foot boundary in order to sustain maximum stability [18]. The desired position and velocity can be set as follows [18]: (6) (7) (8) Where: , is the height of the hip, is the position of center of mass, the gravity , is the ZMP location, is the desired traveling distance and is the desired time to take the desired distance (Fig.3). However, the foot trajectory is set as follow: (9) V. SIMULATION AND DISCUSSION This section presents the simulations realized with AMESim and the results obtained for different conditions and cases. Simulations have been performed separately for each leg, with a 2D model, to allow easy analysis of the results by decoupling the behaviour of each subsystem instead of analyzing only the global system. The motion simulated correspond to walking bipedal robot in straight line at a constant velocity of 2 m/s with the hip at a constant height of 0,49 m (no vertical acceleration)" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0003356_978-981-13-5799-2_12-Figure12.40-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0003356_978-981-13-5799-2_12-Figure12.40-1.png", "caption": "Fig. 12.40 Hydroplaning simulation of a rolling tire [42]", "texts": [ ", the patent of which was granted for Bridgestone in 2002 [46], studies were conducted on the critical hydroplaning speed [47, 48], the effect of the tire pattern on hydroplaning [49] and the braking distance on a wet road [50, 51]. Osawa and Nakajima of Bridgestone [52] developed the riblet wall technology that suppresses the turbulence inside the main groove by CFD and validated to improve the hydroplaning of a tire. (2) Validation of hydroplaning simulation The hydroplaning simulation of a rolling tire with a practical tread pattern is shown in Fig. 12.40, where water drains into the two circumferential grooves and lateral grooves. The velocity of the tire is 60 km/h, the tire size is 205/55R16, the inflation pressure is 220 kPa, and the load is 4.5 kN. Prescribed velocities are applied to the tire model in the horizontal and rotational directions. The coefficient of friction between the tire and road is zero. The depth of water is 10 mm, and flow boundaries around the water pool are wall boundaries through which no water can flow. An explicit FEA/FVM scheme is applied in this simulation, and water flow is evaluated after the hydrodynamic force becomes stable" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0002050_peam.2012.6612506-Figure3-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0002050_peam.2012.6612506-Figure3-1.png", "caption": "Figure 3.The simulation model of skeleton defects of wire ropes", "texts": [ "The signals get amplified and filtered through the signal preprocessor,the analog signals generated by which are then converted into digital signals easy to be treat analysed.After the treatment by DSP or other processing units,digital signals are input for further analysis ,calculation and post treatment which are aimed at extracting target variables of faults like the types,degree and the location.The ultimate analysis results would be sent to the display.In terms of severe defects that has been set by the system ,it will alarm and automatically control the treatmen. Define The ANSYS simulation model of belt testing is established as figure 3 ,according to the principle of the magnetic flux leakage. In the simulation model,6 slender cylindrical steel bars act as the wire ropes,the length of which equals that of the wire ropes being studied.The reason for the processing feasibility is that practically the wire ropes are made by twining steel wire axisymmetrical in diameter.Whether the steel bars or the wire ropes are stimulated ,it is similar that the magnetic induction line would form a closed circuit.Consequently replacing wire ropes with steel bars has such little effect on the leakage magnetic field that the structure difference can certainly be ignored " ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0002584_iemdc.2017.8002303-Figure1-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0002584_iemdc.2017.8002303-Figure1-1.png", "caption": "Fig. 1: Cross-section and dq axis reference frame of the reluctance machine.", "texts": [ " For machines like induction and reluctance machines, selfexcitation can be achieved by the connection of capacitors across the supply terminals, allowing them to be used as a stand-alone generator [2]\u2013[5]. The self-excited induction generator (SEIG) offers certain advantages over a conventional synchronous generator as a source of isolated power supply, such as low cost, robustness, brushless rotor, reduced size, absence of DC source for excitation and low maintenance requirements [2], [5], [6]. However, the main drawback is that both magnitude and frequency of the generated voltage are severely affected by the connected loads. The self-excited reluctance generator (SERG, shown in Fig. 1) provides an alternative solution [7]. It has almost all the advantages of the induction generator, and in addition, the frequency is directly proportional to the rotor speed and with no copper losses in the rotor [8]. In literatures, some analytical models were built to analyze the performance of the self-excited reluctance generator. Abdel-Kader attempted to develop an equivalent circuit for the reluctance generator in the same manner as the induction generator, but ignored the effect of the saliency ratio which is an essential parameter in reluctance machines [4]", " After that, attention is paid on the required minimum residual magnetism of the rotor for self-excitation in reluctance generator. At last, the capability of self-excitation in reluctance generators by connecting charged capacitors is discussed. To simplify the analytical model, some assumptions are fixed: 1) Space harmonics and time harmonics are ignored. 2) Core loss in the machine is neglected. 3) Only the d-axis magnetizing inductance is assumed to be affected by magnetic saturation, while q-axis inductance is considered to be constant. According to the d-q axis reference frame in Fig. 1, equations of the reluctance machine in the transient state are given,\u23a7\u23aa\u23a8 \u23aa\u23a9 vd = Rsid + d\u03bbd dt \u2212 \u03c9e\u03bbq vq = Rsiq + d\u03bbq dt + \u03c9e\u03bbd (1) { \u03bbd = Ldid + \u039bres \u03bbq = Lqiq (2) where vd, vq , \u03bbd, \u03bbq and id, iq are the d- and q-axis voltages, flux linkages and currents, respectively. Rs is the phase stator resistance, \u03c9e is the electrical angular velocity, \u039bres is the residual flux linkage. Ld and Lq are the d- and q-axis inductances. The analytical model is represented by the equivalent circuit shown in Fig" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0003952_s12206-019-0936-3-Figure12-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0003952_s12206-019-0936-3-Figure12-1.png", "caption": "Fig. 12. Schematic of the flight phase.", "texts": [ " Whether the robot has entered the flight process should be determined to distinguish this process from the takeoff process. The robot leaves the ground and enters the flight process when the support force of the ground to the robot is zero and the velocity of the center of mass of the trunk of the robot along vertical direction is greater than zero. That is, 4 0 0 N y F v \u00a3\u00ec\u00ef \u00ed >\u00ef\u00ee (43) where FN is the support force of the ground to the robot. When the robot is in the flight phase, the jumping leg remains fully extended, and the robot can be equivalent to a rigid body (Fig. 12). The mass of the equivalent rigid body is M, and the moment of inertia is Jo. The robot is only affected by gravity after leaving the ground without the effect of air resistance on the robot, and the action point of gravity is in the equivalent center of mass of the equivalent rigid body. Therefore, the motion form of the robot during flight can be regarded as the coupled motion of oblique projection motion of the equivalent center of mass and rotation motion of the equivalent rigid body around the center of mass. Fixed coordinate system O-XY is established by taking the initial position of the equivalent center of mass of the robot during flight as the coordinate origin. The positive direction of X axis is horizontal to the right, and the positive direction of Y axis is vertical upward (Fig. 12). The position and posture of the equivalent rigid body PQ during flight in coordinate system O-XY can be expressed as ( ) ( ) ( ) ( ) ( ) ( ) 0 4 0 4 02 4 0 0 0 21 2 x y y t t X t v t t v t Y t v t t g t t t t g t dtw \u00ec = \u00d7\u00ef \u00ef \u00e6 \u00f6\u00d7\u00ef = \u00d7 - \u00d7 \u00a3 \u00a3 +\u00e7 \u00f7\u00ed \u00e7 \u00f7\u00ef \u00e8 \u00f8 \u00efQ =\u00ef\u00ee \u00f2 (44) where t0 is the time when the robot completes the takeoff process and leaves the ground. (v4x(t0), v4y(t0)) is the velocity of the robot along x2 axis and y2 axis when the robot completes the takeoff process and leaves the ground" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0000392_iros.2013.6696761-Figure6-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0000392_iros.2013.6696761-Figure6-1.png", "caption": "Fig. 6. The bump", "texts": [ " Even though the \u201drigid suspension\u201d has a constant length, we assume that the longitudinal force Pz is continuous, similarly to the real suspension model. This allows us to compute the traction forces Ft1,Ft2 at a given time step, using the forces Pz1,Pz2 from a previous time step. Once Ft1,Ft2 are computed, Pz1,Pz2 are updated for the next time step. The following examples demonstrate our approach to control allocation for a longitudinal vehicle model moving over bumpy terrain. The vehicle parameters, shown in Table I, are of a small dune buggy [13]. Figure 6 shows the bump produced by the exponential function: Z = 0.2e\u22122(x\u22123)2 (18) for x = [0,6]. Figure 7 shows the velocity limit curve and the time optimal velocity profile over the bump. The limit curve exhibits two drops, each occurring when one of the wheels passes over the bump. The drop in the velocity profile is caused by the convex nature of the bump, which forces the vehicle to slow down in order to maintain contact with the ground [11]. Also shown in Figure 7 is the time optimal speed profile from rest to rest" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0003613_012066-Figure7-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0003613_012066-Figure7-1.png", "caption": "Figure 7. Contour of von-Mises stress on a coil spring with a defect of 3.5 mm given the maximum load.", "texts": [], "surrounding_texts": [ "Transport vehicles require a good suspension system to dampen vibration, swings and shocks received as they travel along bumpy, hollow, and uneven roads [1]. These conditions are very uncomfortable and may cause accidents. The suspension is also expected to hold the load during some common vehicle maneuvers such as acceleration, braking or deflection while on the road [2]. The coil spring is one of the main components for dampening vibrations and shocks to the load so as to provide comfort and security while the vehicle is in motion [3]. Depending on the condition of their application, coil springs often sustain fatigue failure. This indicates that the tension received below by the coil spring from the maximum stress of the material while sustaining a dynamic load causes fatigue failure [4-8]. The yield strength of the material is also a criterion of failure. Components of automotive suspension must be changed with a traveling distance of 73,500 km, or every five years [9]. The fault of 13.18 % of 24.2 million vehicle tests was recorded [10]. With the development of computing technology, the numerical analysis method has become particularly suitable for use because it will increase the calculation efficiency, the cost-effectiveness as well as save time. Various numerical analysis methods are widely available, but the finite element analysis (FEA) has proven to be reliable in solving problems in the field of continuum mechanics [11]." ] }, { "image_filename": "designv11_62_0001021_icsens.2013.6688443-Figure3-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0001021_icsens.2013.6688443-Figure3-1.png", "caption": "Figure 3. Prototype drawing from SolidWorks.", "texts": [], "surrounding_texts": [ "nearly every metal ion carrying more than unit positive charge [7]. The stoichiometry for EDTA is 1:1, which means it binds to metal ions like Ca2+ and Mg2+ in a 1:1 ratio. The stability of a metal-EDTA complex is dependent on the pH of the solution. The complexes of divalent metals are stable in ammoniacal solution [7]. By varying pH of a solution, it can be decided which metals will be titrated with EDTA and which will not. For Ca2+ and Mg2+, the complexation reaction with Y4- and corresponding formation constants Kf are [8]:\nThe formation constants with higher values mean that the above reactions will go to completion if EDTA exists in its completely deprotonated form. Moreover, the greater the stability of complexes the sharper the end points of the titration would be. To ensure that the titrant is in its completely deprotonated form, pH values greater than 12 are necessary [8]. However, at pH 12 both calcium and magnesium precipitates in the solution. Therefore, titration is generally carried out at pH 10 at which portion of EDTA is in HY3- form and additional complexation reactions and conditional formation constants are given as [8]:\nBoth EDTA and Ca2+ and Mg2+ ions are colorless. Therefore, a visual indication is needed to detect the end-point in EDTA titration. The most common technique is to use a metal ion indicator. Metal ion indicators are compounds whose color changes when they bind to a metal ion [2]. There are several metal ion indicators available such as Eriochrome black T, Calmagite, Murexide, Xylenol orange and Pyrocatechol violet. Useful indicators must bind metal less strongly than EDTA does [2]. Both Eriochrome black T and Calmagite have been used to detect water hardness. However, Calmagite provides a sharper end point and has a longer shelf life than Eriochrome black T [8]. 1-(2-Hydroxy-5-methyl-1-phenylazo)-2napthanol-4-sulphonic acid or calmagite was proposed as a general purpose metallochromic indicator, though principally for calcium and magnesium, by Diehl and Lindstrom [7]. When calmagite is added to a water sample at pH 10, it will\nreact with the Ca2+ and Mg2+ and the color of the solution will turn into wine red. As EDTA is added, it will first react with any free Ca2+ and Mg2+ and then start to react with the colored complex. EDTA will effectively remove the metal ions from the indicator and tightly bind with them. As the metal ions unbound from the indicator, the color of the solution starts changing to blue which also indicates the end point of the titration. It is worth mentioning that when magnesium content in the solution is low relative to the calcium content,\nmagnesium/EDTA is added to aid in getting a sharper end point [7][8]. Moreover, the indicator binds too weakly to the Ca2+ ions (K\u2019f = 4.4 x 103) and therefore to get a sharp end point Mg/EDTA is added in the solution [8]. The complete reaction process can be shown as follows [9]:\nSpectrophotometry is any procedure that uses light to measure chemical concentrations [2]. Spectrophotometric analyses that use visible radiation are called colorimetric analyses. The intensity of light decreases as it passes through a solution. The amount of light passed through and absorbed by a solution is given by Transmittance (T) and Absorbance (A) respectively.\n0P PT =\nT P P A log)log( 0 \u2212==\nwhere, P0=Initial intensity and P= Final intensity of light.\nAbsorbance is related to the concentration of the sample, the path length that light has to cover and molar absorptivity of a particular substance. The relationship is expressed as Beer\u2019s law given as follows:\nA = \u03b5bc\nwhere \u03b5 is molar absorptivity expressed in M-1cm-1, b is path length expressed in cm and c is the concentration of the solution expressed in M (moles/liter).\nIII. EXPERIMENTAL SETUP The prototype of the sensor was designed initially in Solidworks. Then it was printed using a 3D printer, Dimension Elite, which uses fused deposition modeling technology to print prototypes. The material used to print the prototype was Acrylonitrile Butadiene Styrene (ABS), a common thermoplastic. The overall dimension of the prototype was 30 mm \u00d7 40 mm \u00d7 15 mm. A 10 mm \u00d7 30 mm \u00d7 10 mm channel was embedded to hold the water sample. The prototype had two housings for two LEDs \u2013 one blue and one red, and two housings for two photodiodes. The path length that light had to cover in this prototype was 10 mm. As the prototype was non-transparent, a square opening had been created for the photodiode housing and an iris was created for\nCa2+ + Y4- \u2192 CaY2- Kf = 5.01 x 1010\nMg2+ + Y4- \u2192 MgY2- Kf = 4.9 x 108\nCa2+ + HY3- \u2192 CaY2- + H+ K\u2019f = 1.8 x 1010\nMg2+ + HY3- \u2192 MgY2- + H+ K\u2019f = 1.7 x 108\nMg2+ + HIn2- (blue) \u2192 MgIn- (wine red) + H+\n{EDTA forms a weaker complex with Mg2+ than Ca2+. Ca2+ reacts with EDTA first}\nCa2+ + MgIn- (wine red) + Y4- \u2192 CaY2- + MgIn- (wine red)\n{After all Ca2+ is titrated, EDTA reacts with MgIn-}\nMgIn- (wine red) + Y4- \u2192 MgY2- + In3- (colorless)\nIn3- (colorless) + H2O \u2192 Hin2- (blue) + OH-", "the LED housing. All the openings were closed off with plastic slides appropriately cut to the size of the openings. Then epoxy was applied between the slides and the prototype body to prevent any leakage of the sample. A circular inlet and an outlet were created to inject sample in and out of the channel.\nBefore the experiment, 20 drops of distill water were taken into a 30 cm3 clean vial and a small scoop of calmagite powder was mixed at pH 10. The color of the solution turned blue indicating absence of Ca2+ and Mg2+ ions. The experiment was performed in a semi-dark room to prevent the photodiodes detecting external lights. Two LEDs were inserted into the LED housings and connected to two separate 9V batteries. Then two photodiodes were inserted into the housings and the leads were connected to one Fluke brand and one Agilent brand digital multimeters. Water sample was collected from a nearby tap and preserved in a clean plastic cup. Then using 5 ml syringe water was poured into the channel. The LEDs were turned on and the top portion of the channel, completely open, was covered with a plastic slide colored in black. Voltage readings were taken from the multimeters and recorded, 0.50 V and 0.46 V for blue and red light respectively. Then 2 drops of pH 10 buffer and 2 drops of unknown concentration of calmagite solution, prepared by mixing distilled water with calmagite powder, were poured into the channel using separate clean plastic pipets. A stick was used to mix the solution with the water sample which then turned wine red. The top of the channel was covered with the black slide and voltage readings were recorded. The voltage across the photodiode detecting the blue light dropped by 0.11 V compared to the previous voltage reading whereas for the red light it only dropped by 0.03 V. After the measurement, the cover had been taken off and 2 drops of 0.2 M EDTA disodium salt (EDTA-Na2) solution, purchased from SigmaAldrich, were added to the sample. As soon as the solution was mixed with the same stick, the sample turned blue. The opening of the channel was covered with the black plastic and voltage readings were taken from the multimeters. The voltage reading for the photodiode detecting the red light dropped significantly to 0.21 V. On the other hand, the voltage reading for the blue LED jumped to 0.44 V. After the readings were recorded, the lights were turned off and the photodiodes and LEDs were carefully taken out of their housings. The sample solution was then discarded into the sink and the prototype was rinsed in cold water for few minutes under the tap. The\nprototype was then rinsed with distilled water and finally wiped with a TechniCloth\u00ae II nonwoven wiper. The same process had been repeated 9 more times and finally the voltage readings were plotted with standard deviation error bars. It was observed that the voltage readings for some experiments such as number 4, 7 & 8 did not follow the trend of the other experiments. This was attributed to the disturbance of the test setup especially dislocation of photodiodes due to unintentional touch.\nTen new tests were repeated at a later time with variation in one test parameter. Previously, an unknown concentration of calmagite was used whereas during these ten new tests a known concentration, 0.005 M, of calmagite had been used. During these tests, 4 drops of calmagite were needed instead of 2 drops during the previous 10 sets of test.\nIV. RESULTS Fig. 5 and Fig. 6 show the voltage readings with standard deviation error bars of the two photodiodes for the two different sets of 10 experiments respectively, plotted using MS Excel. The x-axis of the graphs shows the experiment number and y-axis shows the voltage readings recorded by the multimeters across the two photodiodes. Fig. 7 shows the color change when calmagite is added to the water sample at pH 10 and finally when EDTA is added to the sample. It should be noted that the black plastic cover has been placed on top of the structure where the LEDs are located in order to prevent interference while taking picture.", "V. CONCLUSION In this paper, we have presented a sensor that detects hardness in water sample based on concepts of complexometric titration and colorimetric analysis. The sensor prototype has been designed using 3D printing technology which houses two LEDs, blue and red, and two photodiodes. A channel embedded in the prototype holds water sample. When calmagite is added to the sample it turns red prompting voltage drop at the photodiode detecting blue light. Then when EDTA is added to the solution it turns blue prompting voltage drop across the other photodiode. Therefore, we have shown through the experiments the presence of calcium and magnesium ions in the water sample, an indicator of water hardness. The 3D prototype is very shiny and reflects light internally and externally. In our future work, we plan to modify the prototype design to minimize the light reflection. We also plan to investigate a mixture mechanism when calmagite and EDTA are added to the sample and another mechanism to determine the total hardness amount in terms of CaCO3.\nThe authors would like to thank Dr. Bader Aldalali and Dr. Xuezhen Huang of Micro-Nano Sensors and Actuators Laboratory at UW Madison for their help in procuring chemicals and electronic components required for the experiments.\n[1] Fairfax Water. (2013, July 29). Explanation of Water Hardness\n[Online]. Available: http://www.fcwa.org/water/hardness.htm [2] D. C. Harris, Quantitative Chemical Analysis, 6th ed. New York: W.H.\nFreeman and Company, 2003. [3] World Health Organization. (2013, July 1). Calcium and magnesium in\ndrinking-water: public health significance [Online]. Available: http://whqlibdoc.who.int/publications/2009/9789241563550_eng.pdf.\n[4] R. A. C Lima et al., \u201cHardness screening of water using a flow-batch photometric system,\u201d A. Chim. Acta, vol. 518, iss. 1-2, pp. 25-30, August 2004.\n[5] J. Saurina, E. L\u00f3pez-Aviles,A. Le Moal and S. Hern\u00e1ndez-Cassou, \u201cDetermination of calcium and total hardness in natural waters using a potentiometric sensor array,\u201d A. Chim. Acta, vol. 464, iss. 1, pp. 89- 98, July 2002.\n[6] H. Perlman. (2013, July 29). Water Properties and Measurements [Online]. Available: http://ga.water.usgs.gov/edu/characteristics.html [7] T. S. West, Complexometry with EDTA and related reagents, 3rd ed. Poole, UK: BDH Chemicals Ltd, 1969.\n[8] M. C. Yappert and D.B. DuPr\u00e9, \u201cComplexometric Titrations: Competition of Complexing Agents in the Determination of Water Hardness with EDTA,\u201d J. Chem. Educ., vol. 74, no. 12, pp. 1422-1423, December 1997.\n[9] S. E. Bialkowski. (2013, August 1). Titrimetric Water Hardness Determination [Online]. Available: http://ion.chem.usu.edu/~sbialkow/ Classes/361/Hardness.html" ] }, { "image_filename": "designv11_62_0002166_ipack2011-52061-Figure2-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0002166_ipack2011-52061-Figure2-1.png", "caption": "Figure 2. Axisymmetric mesh with the initial shape of the droplet (colored in black).", "texts": [ " A hemispherical cap of water of the required droplet size is initialized in the domain. Any shape changes due to the presence of the electrode needle at the top of the droplet are neglected. Acceleration due to gravity is applied throughout the domain in the negative y-direction. A no-slip boundary condition is specified at the bottom wall. The remaining domain boundaries are specified-pressure boundaries, set at a gauge pressure of zero [29]. The computational domain, mesh and boundary conditions are shown in Figure 2 along with the initial droplet shape shaded black. The contact angle model based on the contact line force balance discussed above is implemented as a contact angle boundary condition on the contact line. It is implemented in the VOF-CSF model through user-defined functions (UDFs) in FLUENT. In this section we present results from a simulation of the transient motion of a 5 \u00b5l water droplet on a smooth Teflon surface. The droplet is actuated using different step input voltages. The predictions from the simulations are compared and validated against experiments" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0003653_j.promfg.2019.06.199-Figure1-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0003653_j.promfg.2019.06.199-Figure1-1.png", "caption": "Fig. 1 L-PBF Build (www.3dspectratech.com)", "texts": [ " This is an open access article under the CC BY-NC-ND license (http://creativecommons.org/licenses/by-nc-nd/3.0/) Peer-review under responsibility of the Scientific Committee of NAMRI/SME. Keywords: Machining; AM Part Machining; AM Part Fixturing; Laser Powder Bed Fusion Post Processing; Inconel 718 Machining Laser powder bed fusion (L-PBF) prints metallic, near net shape parts by using a laser beam to locally melt powder and weld it to an adjacent solid region. The parts are fabricated layer-by-layer up from a build plate (see Fig. 1). An underlying network of solid supports and lattice supports connects the part to the build plate. Lattice supports also connect intra-part regions where overhangs exist. After the build is complete, the plate and attached parts are thermally stress relieved. Subsequently a wire electric discharge machining (WEDM) Available online at www.sciencedirect.com ScienceDirect Procedia Manufacturing 00 (2019) 000\u2013000 www.elsevier.com/locate/procedia 2351-9789 \u00a9 2019 The Authors. Published by Elsevier B", " This is an open access article under the CC BY-NC-ND license (http://creativecommons.org/licenses/by-nc-nd/3.0/) Peer-review under responsibility of the Scientific Committee of NAMRI/SME. Keywords: Machining; AM Part Machining; AM Part Fixturing; Laser Powder Bed Fusion Post Processing; Inconel 718 Machining Laser powder bed fu ion (L-PBF) prints metallic, near net shape parts by using a laser beam to locally melt powder and weld it to an adjacent solid region. The parts are fabricated layer-by-layer up from a build plate (see Fig. 1). An underlying network of solid supports and lattice supports connects the part to the build plate. Lattice supports also connect intr -part r gions where ove hangs exist. After the build is co plete, the plate and attached parts are thermally stress relieved. Subsequently a wire electric discharge machining (WEDM) 464 Edward De Meter et al. / Procedia Manufacturing 34 (2019) 463\u2013474 2 Author name / Procedia Manufacturing 00 (2019) 000\u2013000 process is typically used to separate the parts from the build plate", " In doing so, it will reveal two new technical issues related to the machining of L-PBF parts, stress relaxation distortion and powder sludge generation. In the initial stage of this research, the authors attempted to identify common attributes of L-PBF parts that affect machining process design. To achieve this, the authors worked with AM process engineers at the Center for Innovative Materials Processing Through Direct Digital Deposition (CIMP-3D). Recent builds and archived images of builds were examined as were images of builds taken from the internet (see Fig. 1). Best practices related to solid support design and lattice support design were also discussed. It was determined that parts fabricated using the L-PBF process have three things in common. All have an underlying support network that connects the main solid geometry to the build plate. This network typically consists of both solid metal and lattice. The purpose of this network is to support loose powder for overhangs, extract heat from the weld zone, and restrain part distortion from the large shrinkage stresses that develop during the build" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0002226_ut.2017.7890311-Figure2-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0002226_ut.2017.7890311-Figure2-1.png", "caption": "Figure 2 Equipment layout of general purpose type AUV", "texts": [ " SBP is an acoustic instrument to investigate precisely geological layer structure under several tens of meters below the seabed mainly for exploration of hydrothermal deposits. The second AUV (called general purpose type AUV) can change the payload according to the purpose of investigation. Multi beam echo sounder (MBES) is equipped for survey of seabed topography in the second AUV. And high altitude image mapping device will be equipped for survey of seabed resource. Table 1 shows the main specification of the two AUVs. Lengths of both AUVs are less than 4 meters. Figure 1 and Figure 2 show the equipment layout of each AUV. Two AUVs are equipped with the same acoustic positioning device and underwater acoustic communication device for the simultaneous operation of multiple AUVs. 2.2 Patterned behavior The AUVs can behave underwater with the combination of predefined behavior pattern [2]. The patterned behavior control helps AUV\u2019s operators to set up mission plan easier. Table 2 shows predefined behavior patterns of our AUVs. 3. Sea Trial After finishing the assembling, the land test and the water tank test, we conducted the sea trial for each AUV in Suruga Bay" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0003031_icciautom.2017.8258701-Figure2-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0003031_icciautom.2017.8258701-Figure2-1.png", "caption": "Figure 2. workspace area enhancement of redundant mechanism relative to", "texts": [ " (11), a given point belongs to the workspace of the mechanism; if the stroke limitation of the distal actuators is satisfy, for at least one position of redundant actuators. Also, for the reachable workspace, the method is same as the constant orientation except in step one, the end effector orientation is limitless. In this research, the stroke limitation of the distal prismatic actuators are selected as 200( ) 400( )\u2264 \u2264imm mm (29) In what follows, the workspace of the proposed kinematically redundant mechanism is obtained and the results are compared with the ones in similar size non-redundant mechanism depicted in Fig. 2. The complete kinematics analysis of 3 - RPR mechanism was presented in [15]. Also, two cases for the redundant actuators range are considered. In one case, wide range of redundant actuators are considered, while in the second case, the range of redundant actuators are limited and considered as 1 2 3 / 6 5 / 6, 5 / 6 3 / 2, / 2 / 6 \u03c0 < < \u03c0 \u03c0 < < \u03c0 \u2212\u03c0 < < \u03c0 \u03b8 \u03b8 \u03b8 (30) The constant orientation workspace area of the mechanism is reported in Tables 1, 2 and 3 and the reachable workspace area is reported in table 4 and 5" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0000583_icelmach.2010.5608226-Figure12-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0000583_icelmach.2010.5608226-Figure12-1.png", "caption": "Fig. 12: WFSM with a serial inductance.", "texts": [], "surrounding_texts": [ "A 1.5 kW wound field synchronous machine. The exciter field coils are on the rotor and its armature is on the stator. the machine parameters are given as follows: . Rexc = 0.5 \u2126 Lexc = 28 mH Mse = 75.3 mH . Ld = Lq = 0.37 H Rs = 2.8 \u2126 The used load is: Rf = 530 \u2126 and Lf = 1.4 H . Fig 10 shows the armature voltages obtained by simulation and experimentation. From this figure: - The commutation angle is 27\u25e6 by simulation and 26\u25e6 by experimentation; - During the commutation between two phases, the voltage in the third phase is influenced. This result shows that the classical circuit presented in Fig 4 is not sufficient to describe the machine behavior. The armature currents in the Park frame and the excitation current are given in Fig 11. Simulation results are only presented. As shown in this figure, the excitation current (iexc) oscillates at the same frequency of id and iq . This oscillation reduces the duration of the commutation. From this figure, the excitation current and the d-axis current have a very similar shape. This result validates the previous theory given in subsection B. The following equality can be deduced: (iexc)max \u2212 (iexc)min \u2248 kexc( (id)max \u2212 (id)min) Where kexc is the reduction coefficient between the excitation and the d-axis armature windings. This coefficient is used to refer the variables and the machine parameters to the armature windings reference frame. As shown in Fig 9, if (L\u0303\u03c3exc << Lad), most of the id harmonics go through L\u0303\u03c3exc and the above approximation will be more justified. To verify the influence of the excitation current oscillation, an inductance (L = 2 H) is putted in serial with the excitation winding. This inductance is used to filter the excitation current and eliminate its oscillation. The commutation angle was 27\u25e6 without the using of the inductance L. By using L = 2 H , the simulation and experimental tests give 35\u25e6 for the commutation angle. The machine behavior is similar to a PMSM\u2019s, the coupling between phases is reduced. This result shows the influence of the excitation winding." ] }, { "image_filename": "designv11_62_0002605_s10015-017-0385-y-Figure3-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0002605_s10015-017-0385-y-Figure3-1.png", "caption": "Fig. 3 Schematic diagram of experimental environment", "texts": [ " The error between the actual constrained force fn obtained from the experimental result and the desired force fnd is considered to be proportional to (ft \u2212 K\u0302tfnd) given the condition that the changing of q is limited in small extent. When defining \u0394f = fnd \u2212 fn, Eq. (15) can be changed into Considering that the change of the manipulator\u2019s shape during grinding dose not large, B(q)jT R seems to be generally constant. In this report, the tangential grinding force that reduces the error of the constrained force is obtained from the experiment, by determining the correct coefficient K\u0302t to have the \u0394f to minimized. The control performance is confirmed in the next section. Figure\u00a03 shows experimental environment. Our robot is two link SCARA manipulator with a control period of 0.6\u00a0ms. The work-piece is iron. Figure\u00a04 shows the appearance of placed work-piece. The desired grinding force is set as constant, and the grinding resistance coefficient K\u0302t, which has been explained in Sect.\u00a04, is changed as a experimental parameter. The appropriate K\u0302t can be determined from the following experiment. With the condition that the change of (15)fn = fnd + B(q)JT R (ft \u2212 K\u0302tfnd). (16)\u0394f = B(q)jT R (K\u0302tfnd \u2212 ft)" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0002696_chicc.2017.8028324-Figure8-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0002696_chicc.2017.8028324-Figure8-1.png", "caption": "Fig. 8: The static scene and the ToF cameras. The ceiling of the room in the picture is not visible.", "texts": [ " In the proposed control strategy, two sliding mode controllers are used to minimize the distance error ed(t) and the orientation error ea(t) respectively. Then, the control input u(t) is the orthogonal projection of the resultant vector of the outputs of the two sliding mode controllers. To confirm the performance of the proposed navigation algorithm in static environment, we build a static scene, which is a closed room with some walls. There are four ToF cameras deployed in this scene to detect the obstacles (see Fig. 8). The size of the room is 21m\u00d713.5m\u00d75m. The resolution of the ToF cameras used in the simulation is 128\u00d7 128. The measurement range of the ToF cameras is 15m and the maximum angle of the view is 135\u25e6. The depth images obtained by the four ToF cameras are shown in Fig. 9, which are the real-time measurements of the environment. The model of the flying robot is a quadcopter. The Fig. 10 shows the trajectory of the flying robot in the closed room travelling from the initial point to the target without any collision" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0000974_amr.712-715.709-Figure1-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0000974_amr.712-715.709-Figure1-1.png", "caption": "Fig. 1 Laser cladding device positioned on internal fillets of a crankshaft", "texts": [ " The aim of the research was to design a device and method for renovating and repairing crankshaft journal surfaces in-situ by means of laser build-up. The essence of this technical research is a functional platform for repairing and renovating crankshaft journal surfaces. Principal design of the in-situ laser cladding machine The aforementioned goals are achieved by placing the laser cladding nozzle positioning and guidance device directly on the crankshaft journal fillets. These fillets as a rule are not damaged or worn out and thus preserve the original manufacturer\u2019s crankshaft dimensions. Therefore the internal fillets (see Fig. 1) can be used as a reference surface when positioning the laser build-up nozzle guidance platform. Alternatively and especially for relatively smaller crankshafts, the external filets or radii can be successfully used as a reference and base surface \u2013 cf. Fig. 2. The device is composed of two guideways for positioning it onto crankshaft fillets and two frame parts, each part of which is fitted to its respective guideway. Additionally the device comprises two upper rods by means of which both frame parts are in fixed connection to each other, whereby the upper rods are positioned in the upper part of the frame, and two lower rods by means of which both frame parts are in fixed connection to each other" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0002791_b978-0-12-812138-2.00003-9-Figure3.5-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0002791_b978-0-12-812138-2.00003-9-Figure3.5-1.png", "caption": "FIGURE 3.5", "texts": [ " Similar to the stator, a rotor has a cylindrical iron core that consists of silicon steel laminations. It should be noted that the electric circuit (windings or conductors) in the iron core of the rotor is not fed by an external power source but its current flows by the induced EMF. There are two types of a rotor used in induction motors: squirrel-cage rotor and wound rotor. A squirrel-cage rotor has a laminated iron core with slots for placing skewed conductors, which may be a copper, aluminum, or alloy bar. These rotor bars are short-circuited at both ends through end rings as shown in Fig. 3.5. Stator phase winding. (A) Stator core and (B) sinusoidally distributed winding and mmf. Squirrel-cage rotor. This rotor gets its name from its structure\u2019s resemblance to a squirrel cage. The rotor conductors are not placed exactly parallel to the shaft but are skewed by one slot-pitch to reduce cogging torque, and this allows the motor to run quietly. Because of its simple and rugged construction, about 95% of the induction motors use the squirrel-cage rotor. Similar to the windings on a stator, a wound-rotor as shown in Fig" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0003674_rpj-01-2019-0025-Figure1-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0003674_rpj-01-2019-0025-Figure1-1.png", "caption": "Figure 1 CAD rendering of the hybrid manufacturing equipment", "texts": [ " One route to overcome these limitations is a combinatorial manufacturing approach where multiple digitally driven processes are interleaved to create a single, integrated system (Lorenz et al., 2015). To date, these manufacturing approaches have been prevalent in the metal AM systems, but as yet have not been widely implanted for the production of engineering ceramics (Soares et al., 2018). This paper presents a hybrid AM approach where both additive and subtractive computer-controlled processes are integrated together resulting in a process that has the flexibility of AMwith the surface finish and tolerances achieved using subtractive processes. Figure 1 shows a CAD rendering of the hybrid manufacturing equipment. The process was developed using 96Wt.% per cent alumina with an average primary particle size of 2-3 mm. The raw feedstock has been formulated into a high viscosity, aqueous paste with a measured moisture content of between 18-22 per cent, using a proprietary binding medium developed byMorgan Advanced Materials. Sacrificial support was produced using a commercially available hot end (E3D V6, UK) and extruder (Bulldog Lite, UK) with 1.75 mm polylactic acid (PLA) filament" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0002980_978-94-007-7194-9_71-1-Figure6-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0002980_978-94-007-7194-9_71-1-Figure6-1.png", "caption": "Fig. 6 The system is instantaneously balanced if there exist admissible contact forces fci that can support its", "texts": [ " The ZMP and FRI indicators however lack genericity as being specifically designed for biped locomotion scenarios. Although generalizations of these indicators can be found in the literature to extend them to more complex multi-contact situations, reverting to the essential definition of balance provides practical solutions to the definition problem of balanced states and metrics. One the one hand, an instantaneous approach to balance considers the system to be dynamically balanced if there exist admissible contact forces that can support its motion, as illustrated in Fig. 6. Regarding the system as a whole, it essentially states that the Newton-Euler equations of motion (2) and contact mechanics are satisfied. A balance stability margin can henceforth be defined as the quantification of either the admissible motions around the current state or the disturbance wrenches that can be supported [2]. motion PLx ;M Rx On the other hand, a long-term approach to balance requires the consideration of viability. Pratt and Tedrake propose to this aim to approach viability through the notion of capturability in [47]" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0000253_iciea.2013.6566377-Figure9-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0000253_iciea.2013.6566377-Figure9-1.png", "caption": "Figure 9. Ball fault with window size 1024 for 64 segments", "texts": [], "surrounding_texts": [ "The future of the diagnostic techniques will be self-learning and adaptive. Biologically inspired ANN classifiers with diverse learning capabilities are among the best candidates for such future approaches. A neural network for fault classification can be achieved efficiently due to the fact that neural network is a nonlinear empirical model which can capture the nonlinear system dynamics and do not require knowledge of particular system parameters [7]. Neural network design includes input layer, hidden layers and output layer as shown in Fig. 6. Number of input layer neurons is equal to the input features (7) and the number of output neurons is equal to number of classes. There can be one or more than one hidden layers with different number of neurons. Most of the time increasing the number of hidden layers neurons guarantees good learning but it requires more training time and computational cost. Therefore minimum number of the hidden layers and neurons meeting certain classification accuracy is preferred. Neural Network pattern classifier learns system dynamics in the form of weighted links between the neurons during the training phase. Depending upon problem-nature training can be of the supervised or unsupervised type [4, 9]. In unsupervised learning the targets for input features are unknown while for supervised learning targets are known. We are using supervised learning because for each spectral features training input we have associated targets. For output classification we are using one versus all classifier. Before starting the training the training data is randomly shuffled and divided into training, validation and test sets with different percentages. MSE (Mean Square Error) is used to measure the accuracy of the training, 2013 IEEE 8th Conference on Industrial Electronics and Applications (ICIEA) 263 validation and test sets. Validation set is used as check to stop learning and testing data set is used to measure over fitting of the classifier. Neurons in different layers are connected through weighted links. These weights are tuned during the training of the ANN as shown in Fig. 7. Weights are randomly initialized before starting the training. During training, with random weights, output of the network is used to calculate the MSE using target values. If the calculated MSE is of acceptable value then training can be stopped, otherwise based on the MSE, weights are tuned using back propagation algorithm. Thus, after training, ANN can be used for classification. IV. RESULTS AND DISCUSSION Methodology developed in the previous sections will now be tested practically. Fig. 8 illustrates the experimental setup for recording the actual vibration data sets [10]. Experiments were conducted with four types of bearings including one normal and three faulty bearings with faults in inner-race, ball and outer-race. Faults in the bearing were created by electro discharge machining. Faulty bearings are supporting the shaft of the motor and the load is 2HP with a speed of 1750 r/min. The data have been collected through accelerometers using a 16-channel digital-audio-tape recorder and sampled at the rate of 12000 samples per second. Time vibrations recorded and converted to spectral features for these four signals are shown in the Figs. 5, 9, 10 and 11 respectively with a window size of 1024 for 64 segments. 264 2013 IEEE 8th Conference on Industrial Electronics and Applications (ICIEA) In this paper we have used only one hidden layer with different number of neurons to experimentally check the best training that can be achieved with minimum number of neurons. Number of input features, 513, is equal to spectral features of the time segment. Output layer of the ANN contains four neurons because of four classes as shown in Fig. 12. For training ANN, the spectral contents from different signals with different window size are grouped into a set with respective class (8): ($ , % ) = {($ , % ), ($ , % ), ($ , % ), \u2026 , ($&, %&)} (8) Window size used in this experiment are 256, 512, 1024 and 2048 samples. In (8) p' represents input spectral pattern for any of the four classes with any specific window size. t' represents target class or output for this particular input. Four of the target classes used in this experiment are: *- 1; % = [1 0 0 0] Normal signal *- 2; % = [0 1 0 0] Inner-race fault *- 3; % = [0 0 1 0] Bearing fault *- 4; % = [0 0 0 1] Outer-race fault This grouped data was randomly shuffled and then divided into training, validation and test sets with a respective proportion of 60%, 20% and 20% of the total grouped contents given by (8). With random weight initialization, the network was trained using feed forward back propagation algorithm with different number of hidden layers neurons. Training the network, more than hundred times, for different number of hidden layer neurons, resulted in almost similar minimum MSE. Minimum number of hidden layer neurons that gave acceptable accuracy was 2. Accuracy achieved with two neurons in the hidden layer and more than two neurons are comparable. But using higher number of hidden layer neurons increases the computational time. Fig. 13 shows, as the number of hidden layer neuron increases the learning time increases almost exponentially. Thus 2 hidden layer neurons are best selection for this particular scenario. The appropriate ANN architecture selected is shown in the Fig. 12. As the vibration signals from bearing are quasi-stationary therefore the problem of optimum window selection has been addressed by using multi-sized time-domain segmentation-window for augmented feature selection. Classifier trained with these augmented spectral features has shown 100% accuracy complying with the supposition that smaller windows will capture the signatures appearing for short duration or higher frequencies and larger windows will address the low frequencies contents. The trained network is tested with all the window sizes, 256, 512, 1024 and 2048 2013 IEEE 8th Conference on Industrial Electronics and Applications (ICIEA) 265 with classification accuracy of 100%. Thus multi-window approach offers a range of windows to fit a variety of transients appearing in the vibration signal providing a comprehensive data set for training and testing of the real time rotary machines vibration. V. CONCLUSIONS In this paper, multi-size-window time segmentation based spectral features augmentation for neural network bearing fault classification has been presented. Augmented spectral features of vibration signal, in rotary machines, calculated using multi-size time-segmentation-window have been used to train and test the neural network classifier. Classification results have shown that classifier, with multisize-window time-segmented spectral features, has learned the dynamics of quasi-stationary vibration signals efficiently for real time scenarios with 100% accuracy." ] }, { "image_filename": "designv11_62_0001363_rast.2011.5966982-Figure6-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0001363_rast.2011.5966982-Figure6-1.png", "caption": "Figure 6. The circle centered at O2 with radius r2.", "texts": [ " 2) The Determination of Position of Vertex P2: In this phase, the lengths of L3 and L4 are varied discretely with respect to the related constraints. The coordinates (p2x, p2y, p2z) of vertex P2 are determined by considering L3, L4, and the coordinates (p1x, p1y, p1z) of vertex P1 determined in previous phase. In order to determine P2 (p2x, p2y, p2z) the geometrical relation between P1 and P2 is taken into account. Vertex P2 may be located on the sphere centered at P1 with radius d1, as shown in Fig 5. Let t be the axis on x-y plane which is perpendicular to the line B3B4a and passes through O2 as shown in Fig. 6. Varying lengths of L3 and L4 and keeping P1 fixed, vertex P2 moves in the circle centered at O2 with radius r2, which lies on t-z plane. In order to determine the coordinates (p2x, p2y, p2z) of vertex P2, it is necessary to figure out whether or not the sphere centered at P1 and the circle centered at O2 intersect. This intersection may exist, providing that the intersection on x-y plane between the projection of the sphere and the axis t exists. The projection on x-y plane of the sphere is the circle centered P\u01311 with radius d1" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0003036_imece2017-70301-Figure1-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0003036_imece2017-70301-Figure1-1.png", "caption": "Figure 1\uff0eSketch of non-contact loading device", "texts": [ " Next, an experimental spindle test rig with constant pressure preload is established, and the measurements of stiffness and natural frequency at different rotating states are performed. At last, the effects of rotation speed and temperature on the dynamic performance of measured spindle are discussed based on the measurement results. The non-contact electromagnetic loading device is the main component of the whole measurement system, and it can apply non-contact load to the dummy tool when the measured spindle rotates at high speed state. Fig.1 shows the principle of a noncontact electromagnetic loading device. The device contains iron core, coil wound, and dummy tool. The coil wound and iron core form a pair of adjacent magnetic poles together, and another pair of magnetic poles is at the symmetrical position. When current is applied into the coil wound, a closed magnetic circuit emerges inside the pair of magnetic poles, dummy tool, and the air gap. The dummy tool is held by the tool holder, and the electromagnetic attractive force produced by the electromagnetic field can be given as [9]: 2 0 2 0 cos 4 2 sA NI F C (1) where 0 is vacuum permeability( 7 0 4 10 /H m ), sA denotes the effective cross-sectional area of the magnetic pole, N is the coil turns, I is the current which is applied into the coil wound, 0C represents the air gap between the magnetic poles and dummy tool, is the angle between the two adjacent magnetic poles. In Eq. (1), the relationship between the current I and electromagnetic force F is nonlinear. In order to control the load applied to the spindle, the currents in the two opposite pairs of magnetic poles are set as 0 ci i and 0 ci i respectively (Fig.1), suppose that the air gap is invariant, then the load applied to the dummy tool can be given as: 2 0 0 1 2 2 0 cos 2 s cA N i i F F F C (2) where 0i is called bias current, and ci represents the control current. When the bias current 0i is DC, and remains constant, it can be found that the load applied F to the dummy tool has a liner relationship with the control current ci . Then the electromagnetic force can be controlled by adjusting the control current ci . During the measurement of stiffness of rotating spindle, ci is set as DC, and the amplitude of ci can be adjusted" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0002146_indiancc.2017.7846459-Figure1-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0002146_indiancc.2017.7846459-Figure1-1.png", "caption": "Fig. 1: 3-DOF laboratory helicopter model [3]", "texts": [ " An estimation model for (21) can be built as H(t)W\u0302 \u00af (t) = k\u0302(t) (22) Similarly, Equations (21) and (22) gives the error dynamics and its derivative, as H(t)W\u0303 \u00af (t) = k\u0303(t) H(t) \u02d9\u0303W \u00af (t) = \u02d9\u0303k(t) (23) Let us take a Lyapunov function candidate as V(t) = W\u0303 \u00af T W\u0303 \u00af 2 + k\u0303T k\u0303 2 (24) and using adaptive law for weights at second level as \u02d9\u0302W \u00af (t) = \u2212H(t)k\u0303(t) (25) = \u2212HT(t)H(t)W\u0302 \u00af (t) + HT(t)k(t) the time derivative of V(t) becomes V\u0307(t) =\u2212k\u0303(t)T k\u0303(t)\u2212 (H(t)Tk\u0303(t))T(H(t)T k\u0303(t)) (26) which is negative semidefinite, confirms V(t) in (24) as a valid Lyapunov function. A 3-DOF helicopter model as shown in Fig. 1 is simulated to validate the theoretical developments discussed in previous sections. The 3-DOF helicopter model used here has two rotors, which are front rotor and rear rotor. The helicopter body has 3-DOF: the elevation \u03b5 , pitch \u03b8 and travel \u03c6 angles. The aim here is to control elevation and travel using the voltages generated by two DC motors attached to the propellers. The dynamical equation of helicopter model in state space form is given as [2]: x\u03071 = x2 x\u03072 = \u03d11cosx1 + \u03d12sinx1 + \u03d13x2 + \u03d14cosx3u1 x\u03073 = x4 x\u03074 = \u03d15cosx3 + \u03d16sinx3 + \u03d17x4 + \u03d18u2 x\u03075 = x6 x\u03076 = \u03d19x6 + \u03d110sinx3u1 (27) where the state variables x1 is the elevation angle, x2 is the elevation angular velocity, x3 is the pitch angle, x4 is the pitch angular velocity, x5 is the travel and x6 is travel angular velocity" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0003476_s00604-019-3461-2-Figure2-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0003476_s00604-019-3461-2-Figure2-1.png", "caption": "Fig. 2 A closer look at individual electrode deposition, focusing on the cross section and top view of the electrode. The formation of Au-MRE and Au-PS-MRE are obtained by the following steps; (a) second plasmaetching, (b) PS ring removal (for Au-MRE) and (c) final electrodeposition of gold. Au-PS-MRE is achieved by direct deposition of gold in the presence of PS ring", "texts": [ " The details of the simulation parameters are given in ESM (Part 2). Understanding the formation of microring array structure; Au-MRE and Au-PS-MRE We propose a novel two-step patterning to create the gold ring structures with a large d/r ratio based on the concept of utilizing second etching masks. After the first step of plasma etching, the areas underneath PS spheres were adhered to the substrate surface. A mild sonication in DI water later led to the removal of PS spheres while maintaining the ring-shaped residues (Figs. 1(2) and Fig. 2 (PS ring on ITO after the first etching) and Fig. S2 in ESM). The composition of PS residues changed to graphitic materials as the result of the highly reactive components of plasma causing the chain scission of the PS molecules [22, 23]. From the previous study [20], ITO conductivities decreased in the plasma-exposed area resulting in the selective deposition of gold within the area underneath the PS sphere mask. The gold disc size therefore depends on the PS sphere mask size. In the process, we expect to increase the d/r ratio for less overlapping diffusion domains by adding the second step of plasma etching using PS rings as the secondary mask", " The reduced mask size with constant interspacing can increase the d/r ratio. Furthermore, the whole surface was secondly etched to reduce the conductivity in the region uncovered by the PS ring mask. The Au ring structure was obtained after electrodeposition of gold on relatively high conductive area (Au-MRE in Figs. 1, 2 and 3a). However, the PS ring thickness was decreased from 8.45 \u00b1 0.23\u03bcm to 1.36 \u00b1 0.54\u03bcm after the second etching as a result from direct etching [24\u201326]. The second plasma-etching process can be understood referring to Fig. 2a. The second etched PS ring was damaged due to direct etching, which resulted in a decrease of ring height (area in dashed line disappeared) and ring thickness (rectangular area or striped area disappeared). PS ring thickness is the length difference value between outer radius (rout) and inner radius (rin) as shown in Fig. 2. The results suggested that it was not possible to make more than two etching steps because of the decreasing material thickness. In order to reduce the gold ring size, we devised another synthesis route based on the idea of keeping the PS mask on while performing gold deposition. Unlike the process described for making Au-MRE, here the metal was directly deposited in the presence of PS film mask (Fig. 2c (Au-PSMRE)). Site selective growth of metal was accomplished due to different exposure to the plating solution. Since PS rings were formed as a result of PS sphere etching, they are porous in nature. Therefore, the rings allowed for subsequent gold deposition underneath the PS ring mask (Inset B of Fig.3b). This technique offers a favorable access of metal ion towards ITO surfaces nearing the ring edges causing the arrangement of gold nanoparticles around the edges (Inset A of Fig.3b). As a result, the Au ring covered with PS structure; Au-PS-MRE can be obtained as shown in Figs. 1, 2 and 3b. It can also be seen that the Au ring is specifically formed on the ITO surface where the PS ring was damaged by the second plasma-etching (striped area of the second etched PS ring in Fig.2). This can be explained by the different conductivities on the ITO surface after second etching as shown in Fig. S3 in ESM. The area underneath the damaged PS ring during the second etching has higher conductivity than that of the gap area (low-conductive region, C in Fig. S3) since it was protected by the mask in the initial step. We refer to this area as a medium-conductive region (B in Fig. S3). AuNPs deposition is initiated within the PS protecting area (high-conductive region, A in Fig. S3)" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0001159_jjap.51.030206-Figure1-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0001159_jjap.51.030206-Figure1-1.png", "caption": "Fig. 1. (a) Schematic illustration of cell pairing through a microslit in a microchannel. ES: ES cells. So: Somatic cells. Black arrows indicate schematic flow streamlines. (b) A close-up view of paired cells. Reproduced from Proc. Micro Total Analysis Systems 2010, p. 687.", "texts": [ " We arranged a hydrodynamic weir to trap the cells within the vicinity of a microslit in a microchannel. The microslit acts as a space to keep the cells in contact with each other. Single-step pairing was achieved without flow switching, by introducing each cell suspension from different inlets into the microchannel at the same time. As an application of the pairing, we report cell fusion through the microslit which regulated cell fusion efficiency and subsequent movement of contents of the cells (cytoplasm, nuclei, and other organelles). The principle of the cell pairing is shown in Fig. 1. Two different types of cells are introduced from a separate inlet into the microchannel [Fig. 1(a)]. In this study, embryonic stem (ES) cell and somatic cell were used. A hydrodynamic weir arranged in the microchannel consists of spaces for each cell and a microslit connecting the two spaces. A schematic illustration of the flow profile around the weir is also shown in Fig. 1(a). A part of cells are trapped and paired through the microslit as shown in Fig. 1(b), while other cells pass without being trapped. In contrast to previous report wherein different types of cells were simply trapped into a weir after sequential injection of the suspensions,6) the cells are trapped into each side of the weir, and single-step pairing is achieved. Afterward, we can study natural communications between the paired cells by the observation of dye transfer,5) or we can study \u2018\u2018forced\u2019\u2019 communications by intentional fusion of the paired cells using appropriate reagents" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0000961_3ca.2010.5533559-Figure1-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0000961_3ca.2010.5533559-Figure1-1.png", "caption": "Fig . 1 : Inver ted pendulum system.", "texts": [ " In this paper, we proposed an adaptive controller for the posture control problem of inverted pendulum control to achieve the property of the asymptotical stability even though the mismatched perturbations existed in the plant. This paper is organized as follows. In Section 2, we briefly introduce the inverted pendulum to be controlled. Section 3 gives the design of the robust controllers. The analysis of stability is presented in Section 4. In Section 5, a numerical simulation is used to show the applicability of the proposed design technique. II. Preliminaries and Problem Formulations The inverted pendulum system (Figure 1) is composed of a rigid pole and a cart. The rigid pole is hinged on the cart. The cart can move left or right along the rail. The objective of controlling the inverted pendulum is to drive the cart so that the rigid pole can be balanced, that is, 0 . The dynamic equations of the inverted pendulum are governed by [8] 2 2 2 2 2 ( ) cos( ) sin( ) 0 cos( ) sin( ) sin( ), (1) m g m g in m m k k k k p m M p m r r m m p m p m mg where M is the mass of the cart, m is the mass of the rigid pole, L is the half the actual rigid pole length, F is the input force to the cart, p is the position of the cart, is the angular position of the rigid pole, in is the voltage input, m is the armature resistance, gk is the internal gear ratio, mk is the motor torque constant, and r is the motor gear radius" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0002932_10_2017_44-Figure8-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0002932_10_2017_44-Figure8-1.png", "caption": "Fig. 8 Different micro- and macroscale light harvesting (photo)electrochemical modules developed for current production and characterization: (a) Cost effective MFC [68], (b) Single chamber photoelectrochemical system [69], (c) Flow-through system [70], (d) Single chamber BPV [51], (e) BPV for suspended cells [18], (f) Microfluidic BPV [3], (g) Microsized bio-solar cell [71], and (h) Scalable micro-liter sized bio-solar cell [71]. Images were reproduced with kind permission from individual journals", "texts": [ " Devices were not optimized in most of the reported studies, which instead focused on either photocurrent generation or fundamental analysis, as diagnostic toolboxes for study of electron transport. These toolboxes are not yet fully optimized due to their complexity, especially of electronic interfacing, and these instruments are generally treated as black boxes by most researchers. This limits understanding of the experimental setups and can result in overinterpreted conclusions. Photoelectroactivity of cyanobacteria can be characterized by measuring open circuit potential or by applying potentiostatic methods. Luimstra and coworkers developed a simple and cost-effective device (Fig. 8a) that was reported to obtain reproducible photoelectroactivity (i.e., power and current outputs) using algae and cyanobacteria. The device was also recommended for benthic varieties, especially those that attach to surfaces [68]. Cereda and coworkers developed an electrochemical characterization method for rapid and quantitative characterization of the electroactivity of Synechocystis sp. PCC 6803 [24]. The module allowed potentiostatically controlled experiments; rapid read-outs were recorded without any processing time for biofilm growth [24]. However, these studies also lacked characterization of the biofilm and, especially, of the link between biofilm physiology and electroactivity. An ideal module should combine online biofilm monitoring using microscopy techniques and correlate it with electron production. Inglesby and coworkers developed a singlechamber photoelectrochemical system for in situ microscopy investigation of the photosynthetic filamentous cyanobacterium Arhtrospira maxima (Fig. 8b). The system allowed study of the effects of temperature, light intensity, photosynthetic activity, and current density [69]. Attachment of cyanobacteria to the electrode surface and subsequent biofilm formation is still a major challenge, especially for continuous, real-time, and online microscopy monitoring of growth. In a recent study, St\u20acockl and coworkers demonstrated a flow-through system that allowed simultaneous imaging and monitoring of current outputs of Shewanella oneidensis MR-1 [70] (Fig. 8c). The flow-through setup was complemented with electrochemical impedance spectroscopy and confocal microscopy [70]. It would be promising to adapt such microscopically coupled electrochemical modules for investigation of photosynthetic organisms. In another characterization study, McCormick and coworkers incorporated an oxygen sensor directly into the photoelectrochemical module and correlated the peak power densities of Chlorella vulgaris and Synechococcus sp. with photosynthetic oxygen evolution (Fig. 8d) [51]. Ryu and coworkers were able to demonstrate photosynthetic electroactivity at the single cell level. A nanoelectrode was inserted into the immediate vicinity of the thylakoids of the green alga Chlamydomonas reinhardtii and electrons from the photosynthetic activity were directly harvested [72]. Such a technique can be very useful for estimating the available electron pool size in a cell. However, it was not possible to position the nanoelectrode accurately in the thylakoid membrane and, therefore, precise conclusions about the source of electrons from the photosystems could not be derived", " Scaleup of BESs by increasing their size often increases the internal resistances and electrochemical losses. Therefore, connecting smaller units is preferred. Additive voltage of 2 V was reported when four small-scale BESs were connected in series; the system was able to power a digital clock [51]. Similarly, Madiraju and coworkers stacked three two-chamber electrochemical modules to obtain high power densities [53]. Lee and Choi demonstrated another stacking method by connecting multiple small-scale units, as shown in Fig. 8g and h. Accordingly, Wei and coworkers obtained 1.28 V operating voltage by connecting nine microscale modules in series [74]. Microscale photobioelectrochemical reactor modules have shown tremendous process performance and typically high current densities. Lee and coworkers used a 300-\u03bcL module and obtained high current densities of up to 3.2 mA m 2 [71] (Fig. 8g and h). Bombelli and coworkers constructed a 400-nL module that enabled a current density of 1,050 mA m 2, which is the highest value reported to date for nano- and microscale devices [3] (Fig. 8f). These current ranges were achieved as a result of low internal resistances and high surface-to-volume ratios. Most of the microsystems were operated in batch mode. Under these circumstances, the stability of the biofilm should be carefully studied. Excess biofilm growth and clogging have to be taken into account in order to use these systems continuously in long runs. For photoelectrochemical modules to be competitive with existing PV technologies, several issues need to be addressed. The overall light conversion efficiency of photoelectrochemical modules depends on various biological, technical, and process parameters" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0000092_pedg.2010.5545826-Figure3-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0000092_pedg.2010.5545826-Figure3-1.png", "caption": "Fig. 3. Relationship between the phase current and the phase inductance I.", "texts": [ " According to the literature[5], the analytic formula of the generation voltage can be derived as follows, 2 1 max 2 ( ) [ ( )]C G U i L K U \u03b8 \u03b8 \u03c9 \u03b8 \u03b8 \u03b8 \u03b8 \u2212 \u2212 \u2212 \u2212 = \u2212 2( )d\u03b8 \u03b8 \u03b8\u2264 \u2264 (3) 2 1 max 2 ( ) G U i L U \u03b8 \u03b8 \u03c9 \u03b8 \u03b8 \u2212 \u2212 = \u2212 3( )d\u03b8 \u03b8 \u03b8\u2264 \u2264 (4) Where, a\u03b8 \u3001 b\u03b8 \u3001 c\u03b8 \u3001 d\u03b8 are related to the pole arc coefficient of the generator stator and rotor, 1\u03b8 is the turn-on angle of the power converter, 2\u03b8 is the turn-off angle of the power converter, Lmax is the maximum value of the phase inductance. The generation voltage is related to the turn-on angle of the power converter, the turn-off angle of the power converter and the angular velocity of the generator in the certain generator with the certain excitation voltage. While the rotor speed of the generator is high, the rotational electromotive force, Li\u03c9 \u03b8 \u2202 \u2202 , is greater than the power generation voltage, GU , the phase current will be raised sequentially after the main switches are turned off, as shown in Fig. 3. In the high rotor speed operation state, the phase current will be raised sequentially a long time after the main switches are turned off. The longer the freewheeling time of phase currrent is, the easier the establishing generation voltage is. While the rotor speed of the generator is low, the rotational electromotive force, Li\u03c9 \u03b8 \u2202 \u2202 , is smaller than the power generation voltage, GU , the phase current will be dropped immediately after the main switches are turned off, as shown in Fig. 4. In the low rotor speed operation state, the phase current will be dropped immediately after the main switches are turned off" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0001287_el.2012.3326-Figure1-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0001287_el.2012.3326-Figure1-1.png", "caption": "Fig. 1 Positioning angle of shaft a 3-DOF motor b Positioning angle of shaft", "texts": [ " The motion of the 3-DOF motor can be classified as rotating motion and positioning motion. The rotating motion is identical to the motion of the rotary machine, and the rotating angle is represented as ur . The shaft can be tilted as well as rotated. Therefore, the motion is named as the positioning motion in this Letter. The positioning angles are represented as ua and ub in the spherical co-ordinate system and the angles could be measured by encoders which are placed on the motor frame shown in Fig. 1a. The anglesua, ub and ur are described in Fig. 1b. ua is the inclination angle measured from the z-axis of the stator reference frame and ub is the azimuth angle of its orthogonal projection on the xy-plane of the stator reference frame. Fig. 1b shows the shaft direction on a 3D plane when the shaft is tilted to ua and ub. In Fig. 1b, the superscript \u2018s\u2019 on the notation of axes indicates the stator reference frame. Using this co-ordinate system we could calculate the shaft positioning angle from the stator reference frame. Also, the angles ua and ub cannot be obtained from the encoder signal. To calculate these angles, encoder signal ux and uy should be transformed into ua and ub through the following equation: ua = tan\u22121 1/ tan2 (ux)+ tan2 (uy) \u221a( ) ub = tan\u22121 (tan (uy)/\u2212 tan (ux)) (1) Using this co-ordinating system or co-ordinate transformation, we can not only obtain the shaft positioning angle but also the position of the magnet pole and coils are measured" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0003190_978-3-030-05321-5_8-Figure8.45-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0003190_978-3-030-05321-5_8-Figure8.45-1.png", "caption": "Fig. 8.45 Cross-section of tough hand", "texts": [ " In the bucket mode, in order to form a concave shape, all the root joints are wide open, and the two tip joints of the fixed fingers are closed. In this case, the opening angle of the fixed fingers is designed to be larger than that of the rotating fingers. Furthermore, the rotating fingers are designed to be shorter than the fixed fingers so that all the four fingertips can be arranged in line to ensure satisfactory bucket performance. However, this difference in the finger lengths, in turn, may lead to degraded performance of the device as a hand. The finger lengths were thus considered an optimal compromise. Figure8.45 shows the cross-sectional view of the hand. A fixed finger has three joints, but two of these joints are interlocked by a link bar and bend simultaneously. Thus, a fixed finger has two degrees of freedom. The rotating finger has one joint that is located near the spin axis. In all fingers, the bending motion of one degree of freedom is generated by one linear hydraulic actuator (cylinder). All the linear actuators employed this hand have the same specification. Hereafter, in some cases, this linear actuator is called a bending (control) actuator" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0000860_icma.2012.6282812-Figure3-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0000860_icma.2012.6282812-Figure3-1.png", "caption": "Fig. 3 A point in the tip and its projection on image plane of up-vision", "texts": [ " The HTM Te is calculated as (2) in which the character in the bracket is omitted. The HTM 0T1 can be derived base on the rigid body kinematics as (3). cos sin 0 0 sin cos 0 0 0 0 1 0 0 0 0 1 Ti (1) ( , , ) ( , ) ( , ) cos sin sin sin cos 0 cos sin 0 sin cos cos cos 0 0 0 1 Te x y z y x y x y y x x x x y x y x y z Trans Rot y Rot x (2) 0 1T T Te i (3) Supposing that there is a point p\u2019 on the tip surface plane XaOaYa and point p is its projection on the image plane XiOiYi of up-vision shown as Fig. 3, the X and Y coordinates of point p on XiOiYi plane are just the coordinates of point p\u2019 with respect to frame OiXiYiZi . This can be written as (4). ' '1 0 1i i i i i p p p p px y z x y (4) The coordinates of point p\u2019 with respect to frame OiXiYiZi can be calculated using HTM iTa which represents the transformation of frame OaXaYaZa with respect to frame OiXiYiZi. The HTM iTa is the same with 0T1 .Supposing that equations axp\u2019=a and ayp\u2019=b are satisfied, the following (5) can be derived. ' ' 0' ' 1 '' ' 11 1 T T i a p p i a ip p a ai a pp p ax x by y zz z (5) By substituting the results of ixp\u2019 and iyp\u2019 calculated by (5) into (4), the X and Y coordinates of point p can be derived as (6) and (7) by eliminating the quadratic terms" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0001517_0731684409348345-Figure1-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0001517_0731684409348345-Figure1-1.png", "caption": "Figure 1. Assembly drawing of the novel kneader reactor.", "texts": [ " But these mechanisms can be improved by changing the profiles of twin-shaft kneaders based on gear engagement theory. For screw pump, the fill level of the plug-flow reactor depends on the intrinsic self-cleaning properties of the reactor itself. This phenomenon is well known as the \u2018compulsory transportation effect\u2019, or \u2018pumping effect\u2019. Based on these observations and working principles of traditional twinscrew extruders, a novel continuous twin-screw kneader reactor, composed of self-cleaning, compulsory transportation, and mixture, as well as plasticizing, is proposed in this article (Figure 1). The core component of the novel twin-screw kneader reactor is one pair of mutally engaged screw rotors. The key to improving the overall performance of the kneader lies in the design, analysis and optimization of the tooth profile of screw rotors. Therefore, geometric analysis, optimization, and tooth profiling of screw rotors of the kneader reactor are discussed in following sections. DESIGN THEORIES OF ROTOR PROFILES The profiles of the novel twin-screw kneader are designed using the design method of twin-screw pump and spatial engaging theory", " Meanwhile, the teeth number ratio imf and lead length Hm,Hf can directly affect the overall performance of the kneader, therefore, imf and Hm,Hf should be used as the design variables when analyzing the geometric properties of the twin-screw kneader. In order to analyze the geometric properties of the twin-screw kneader, the volumetric efficiency Cn should be calculated first. The area of end section of the screw rotor can be expressed as: Sf \u00bc Pi 1 1 2 R ui\u00fe1 ui y0fxf x0fyf du Sm \u00bc Pi 1 1 2 R ui\u00fe1 ui y0mxm x0mym du 8>>< >>: , \u00f09\u00de where Sf is the area of end section of the female rotor, and Sm is the area of end section of the male rotor, and a is the meshing angle of the two screw rotors and can be expressed as follows (Figure 1): \u00bc 2 arctan ffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi 4RaH H2 p =\u00f02Ra H\u00de h i , \u00f010\u00de where Ra is the tip circle radius of female and male rotor, and H is the groove depth of the rotor. The area At of the end section of the barrel is: At \u00bc 2 R2 a 2 2 \u00fe Ra H 2 Ra sin 2 : \u00bc \u00f02 \u00deR2 a \u00fe \u00f02Ra H\u00deRa sin 2 \u00f011\u00de Supposing the length of rotor is l, the flow area f at the end section and the volume V between teeth of the kneader are: f \u00bc At Sf Sm V \u00bc f l : \u00f012\u00de at UNIV OF CALIFORNIA SANTA CRUZ on November 25, 2014jrp" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0001027_cdc.2013.6760635-Figure1-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0001027_cdc.2013.6760635-Figure1-1.png", "caption": "Fig. 1. Geometric view of Doppler tracking", "texts": [ " Notice that, in the linear case, the correction and prediction steps can be carried out by a Kalman Filter (possibly in information form to avoid the transformation from the covariance to the information form needed for consensus). Conversely, in the nonlinear case proper modifications are needed [14]; the prediction and correction steps can be carried out by any nonlinear filter, e.g. Extended Kalman Filter (EKF), Unscented Kalman Filter (UKF) or Particle Filter (PF). A Doppler sensor measures the frequency shift fd of a received target echo. By the Doppler effect, such a shift is proportional to the radial speed \u03c1\u0307 (see fig. 1) according to the relationship fd 4 = fT \u2212 fR = 2 \u03bb \u03c1\u0307 = 2 \u03bb \u03bd cos\u03b1 (7) where: fT and fR denote the transmitted and, respectively, received signal frequencies; \u03bb = c/fT is the transmitted signal wavelength, c being the light speed; \u03c1\u0307 = \u03bd cos\u03b1, \u03bd being the target speed and \u03b1 the angle between the target direction and the sensor-target line, referred to in the sequel as the relative heading. Let us consider a target, moving in the (\u03be, \u03b7) plane, with kinematic state xt = [\u03bet \u03b7t \u03be\u0307t \u03b7\u0307t] \u2032 evolving in time according to the discrete-time motion model (1)", " In order to decompose the kinematic state x with respect to its observability properties from the Doppler measurements yt, t = 0 1, . . . one can resort to geometrical considerations and define a suitable transformed set of coordinates. To this end let \u03c1t = |pt| be the range; \u03bdt = |p\u0307t| be the speed; \u03b8t = atan2 (\u03b7t, \u03bet) be the azimuth; and \u03b1t 4 = atan2 ( \u03b7\u0307t, \u03be\u0307t ) \u2212 atan2 (\u03b7t, \u03bet) be the difference of the heading angle minus the azimuth, i.e. the relative heading. The geometrical meaning of such transformed coordinates is clearly depicted in fig. 1. Next, the aim is to re-express the constant velocity motion model in the transformed state vector z 4 = [\u03c1, \u03bd, \u03b1, \u03b8] \u2032 instead of the original state vector x. From the definitions of the transformed coordinates, the following result can be readily stated. Proposition 2: Consider the transformed set of coordinates z 4 = [\u03c1, \u03bd, \u03b1, \u03b8] \u2032, then the state equation turns out to be \u03c1t+1 = [ \u03c12 t + T 2 s \u03bd 2 t + 2Ts\u03c1t\u03bdt cos\u03b1t ] 1 2 \u03bdt+1 = \u03bdt \u03b1t+1 = \u03b1t \u2212 atan2 (Ts\u03bdt sin\u03b1t, \u03c1t + Ts\u03bdt cos\u03b1t) \u03b8t+1 = \u03b8t + atan2 (Ts\u03bdt sin\u03b1t, \u03c1t + Ts\u03bdt cos\u03b1t) (9) and the Doppler observation equation takes the form yt = 2 \u03bb \u03bdt cos\u03b1t " ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0003015_2017-36-0413-Figure7-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0003015_2017-36-0413-Figure7-1.png", "caption": "Figure 7. Composite rear bumper at 90\u00b0.", "texts": [], "surrounding_texts": [ "The analysis of the bumper, both for composite orthotropic material and for isotropic steel material, were performed under the same contour conditions and force application. The results demonstrate the equivalent stresses for the steel and composite bumpers." ] }, { "image_filename": "designv11_62_0003923_b978-0-12-814245-5.00018-9-Figure18.11-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0003923_b978-0-12-814245-5.00018-9-Figure18.11-1.png", "caption": "FIGURE 18.11 When \u03d55 0, the arc lies on the X Z plane, the position of the tip is shown in the figure.", "texts": [ " A p p licatio n s o f Fle xib le R o b o ts By choosing i5 2, we have, \u03ba5 s2 l2 2 dscos\u03d5 Substituting s5 \u00f0l1 1 l2 1 l3 1 l4\u00de=4, \u03ba5 l1 1 l2 1 l3 1 l4\u00f0 \u00de=4 2 l2 2 d l1 1 l2 1 l3 1 l4\u00f0 \u00de=4 cos\u03d5 5 l1 1 l2 1 l3 1 l4\u00f0 \u00de2 4l2 2 d l1 1 l2 1 l3 1 l4\u00f0 \u00decos\u03d5 Since tan\u03d55 \u00f0l3 2 l1\u00de=\u00f0l4 2 l2\u00de, we get, cos\u03d55 l4 2 l2ffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi \u00f0l42l2\u00de2 1 \u00f0l32l1\u00de2 q Substituting back to the above equation, \u03ba5 l1 1 l2 1 l3 1 l4\u00f0 \u00de2 4l2\u00f0 \u00de= 2 d l1 1 l2 1 l3 1 l4\u00f0 \u00de\u00f0 \u00de l4 2 l2\u00f0 \u00de ffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi l42l2\u00f0 \u00de2 1 l32l1\u00f0 \u00de2 q 5 \u00f0l1 2 3l2 1 l3 1 l4\u00de ffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi l42l2\u00f0 \u00de2 1 l32l1\u00f0 \u00de2 q 2 d l1 1 l2 1 l3 1 l4\u00f0 \u00de\u00f0l4 2 l2\u00de In summary, we successfully define the manipulator-specific mapping g, converting actuator lengths l1; l2; l3; l4\u00f0 \u00de to arc parameters (s;\u03ba;\u03d5\u00de. By using the arc geometry method to define the continuum manipulator-independent mapping h, the manipulator is modeled as a piecewise arc, as shown in Fig. 18.11. Considering when \u03d55 0, the arc lies on the X Z plane. The position of the tip is calculated as: r 12 cos\u03b8\u00f0 \u00de; 0; rsin\u03b8. 314 Handbook of Robotic and Image-Guided Surgery The transformation of the coordinate frame follows this sequence: (1) rotating about the Y-axis by \u03b8 to align with the tip frame; (2) translating to tip by r 12 cos\u03b8\u00f0 \u00de; 0; rsin\u03b8; (3) rotating about Z-axis by \u03d5. The homogeneous transformation matrix is given by, T 5 cos\u03d5 2sin\u03d5 0 0 sin\u03d5 cos\u03d5 0 0 0 0 1 0 0 0 0 1 0 BBB@ 1 CCCA cos\u03b8 0 sin\u03b8 r 12 cos\u03b8\u00f0 \u00de 0 1 0 0 2sin\u03b8 0 cos\u03b8 rsin\u03b8 0 0 0 1 0 BBB@ 1 CCCA 5 cos\u03d5cos\u03b8 2sin\u03d5 cos\u03d5sin\u03b8 rcos\u03d5\u00f012 cos\u03b8\u00de sin\u03d5cos\u03b8 cos\u03d5 sin\u03d5sin\u03b8 rsin\u03d5\u00f012 cos\u03b8\u00de 2sin\u03b8 0 cos\u03b8 rsin\u03b8 0 0 0 1 0 BBB@ 1 CCCA By substituting \u03ba5 1=r and \u03b85\u03bas, we get the matrix represented by configuration space variables, T 5 cos\u03d5cos\u03bas 2sin\u03d5 cos\u03d5sin\u03bas cos\u03d5\u00f012 cos\u03bas\u00de \u03ba sin\u03d5cos\u03bas cos\u03d5 sin\u03d5sin\u03bas sin\u03d5\u00f012 cos\u03bas\u00de \u03ba 2sin\u03bas 0 cos\u03bas sin\u03bas \u03ba 0 0 0 1 0 BBBBBBBBBB@ 1 CCCCCCCCCCA Note that the orientation of the frame after the above transformation does not rotate about the new axis, which does not represent the true orientation in the real world" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0002302_b978-0-12-803581-8.10351-0-Figure16-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0002302_b978-0-12-803581-8.10351-0-Figure16-1.png", "caption": "Fig. 16 Monocoque \u2013 paired side shells.", "texts": [ " While this chapter is meant as an overview of the product development process, from design, through tooling development, layup and assembly, and not a detailed design study, the discussion in this section does suggest materials and ply stacking sequences consistent with a lightweight monocoque chassis. These specifics help demonstrate the details required to fully develop such a product. In much of the following discussion, portions of the monocoque shell are excluded in an attempt to show more conservative performance. The carbon fiber type, layers and orientation for different portions of the chassis are described. Fig. 16 shows the two side shells, combined, without the top or bottom shells in place to illustrate the basic laminates used in the side shells. The inside of the pair of side shells defines the driver envelope. When the top and bottom shells are added, they result in additional plies of composite structure in the overlap regions, generating the full required structure. The principal structural design requirement is torsion. The basic structure of the monocoque consists of five layers of plain weave surrounding \u00bc00 Nomex honeycomb core", " These strakes play several different roles, including guiding external airflow. However, from a structural standpoint, the lower strake is of primary interest. This lower strake provides added side impact protection in that region between the front wheel/suspension and the sidepod. The layup is the basic stacking sequence given above, but the strake is filled with a preshaped structural foam shape and capped with additional layers of prepreg on the inside surface during side shell manufacture, creating a stiff beam structure in the region shaded in red in Fig. 16. The monocoque has two added unidirectional plies in for the suspension box region (shaded in blue in Fig. 16) with fibers running fore and aft, for enhanced front crush protection. This region of the monocoque consists of five layers of plain weave, and two additional layers of unidirectional prepreg surrounding 6 mm (\u00bc00) thick Nomex honeycomb core. 745;745;0; 0290; core; 0290;0;745\u00bd Fig. 17 shows the top shell of the multi-shell monocoque design, which completes the structure of the cockpit rim stiffener, includes the structural sidepod and head restraint structure, and provides the contour for the nonstructural suspension box cover" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0000744_isie.2013.6563755-Figure5-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0000744_isie.2013.6563755-Figure5-1.png", "caption": "Fig. 5. Proposed solution", "texts": [ " Therefore a compensation for the backlash is necessary. Even if the backlash is presented on a thrust wire, it does not affect much if the wire is used to a continuous action of either pulling or pushing of an object. Considering this into account, this method is developed to lay two thrust wires in parallel and keep one wire tensed and the other pushed while the operation. This is can also be considered as a modified case of tension control for the tendon driven systems [14]. This idea is illustrated in Fig. 5. For the modeling purpose of this concept, the position measurement direction was used positive for the pushing direction of the thrust wire. The advantage of this method is the non-necessity of additional sensors to measure the backlash and its effects. The backlash is simply compensated with the use of two parallel thrust wires with thrust control. In this research, bilateral control is used with thrust control to achieve the required task. In this experiment, four linear actuators are used. One is taken as the master, two actuators for slaves and the rest was used to measure the end effector position. Indices used for this system modeling is shown in table II. To model the setup and the control following transformation matrices were used. From the actuator space to the modal space the position tracking and the law of action and reaction is also to be realized. By considering the directions of forces and position of Fig. 5, when both position and force are in the same direction it will be the joint force. When the force and position are in opposite directions, the force will be thrust force. Based on that, the transformation matrix can be written as (3). Also the transformation between the modal space and the joint space has the thrust control of the two thrust wires. Therefore the transformation matrices can be written as (4). 1 11 1 (3) 1 11 1 (4) Here denotes the transformation matrix between actuator space and the modal space" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0002166_ipack2011-52061-Figure1-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0002166_ipack2011-52061-Figure1-1.png", "caption": "Figure 1. Droplet system for transient droplet spreading under electrical actuation. Experimental image of a conducting droplet (water) with a conducting needle is shown in the inset.", "texts": [ " The time history of the droplet shape, contact line radius and contact line speeds are computed and compared with measurements from experiments conducted as part of this work for validating the model. The effect of contact line friction on the transient droplet motion is analyzed. An approximate mathematical model is also developed based on the contact line force balance to help interpret the detailed VOF computations. The model is used to predict the transient contact line motion and the maximum contact line velocity. It is found to be in good agreement with experiments and with the detailed VOF model. The experimental setup is shown in Figure 1. A conducting droplet is placed on a thin layer of dielectric coated with a Teflon surface and surrounded by air. The experimental droplet response to the applied actuation is recorded at 1000-2000 fps using a high-speed camera (Photron 1024 PCI). All the images are processed using MATLAB [28] to determine the dynamic contact angle and the interfacial contact radius. The experimental uncertainties in contact angle and contact radius measurements are \u00b12\u00b0 and \u00b10.03 mm, respectively. A highly-conducting silicon wafer with a 1 \u00b5m thick thermally grown oxide layer is utilized as the substrate. The substrate is spin-coated with 1% Teflon solution (DuPont, Wilmington, DE) forming a 0.5 \u00b5m thick later to impart hydrophobicity. An aluminum needle of 125 \u03bcm diameter is inserted into the droplet from the top, as shown in Figure 1. Copyright \u00a9 2011 by ASME Downloaded From: http://proceedings.asmedigitalcollection.asme.org/pdfaccess.ashx?url=/data/conferences/interpack2011/70206/ on 05/26/2017 Terms of Use: http://www.asme.org/about-asme/terms-of-use 3 In the current work, the volume of fluid-continuum surface force (VOF-CSF) model in the commercial fluid dynamics software package, FLUENT [29], is used. A custom contactangle model based on the force balance at the contact line is implemented using user defined functions to capture the effects of surface tension, electrowetting and dynamic contact line forces", " The implementation of the contact line friction is similar to that of Keshavarz-Motamed et al. [26]. An image of the experimental droplet and the aluminum needle is shown as an inset. A voltage difference is provided between the silicon wafer and the needle to actuate the droplet. A 5 \u00b1 0.1 \u00b5l de-ionized water droplet is used in all the experiments. The initial contact angle and the contact radius of the droplet in the absence of electrical actuation are 120\u00b0 \u00b1 2\u00b0 and 0.98 \u00b1 0.03 mm, respectively. The simulation mimics the experimental setup, which is shown in Figure 1. Because the droplet motion is axisymmetric in nature, a 2D axisymmetric VOF computation is performed in FLUENT. A square grid in the r-y plane is used. A hemispherical cap of water of the required droplet size is initialized in the domain. Any shape changes due to the presence of the electrode needle at the top of the droplet are neglected. Acceleration due to gravity is applied throughout the domain in the negative y-direction. A no-slip boundary condition is specified at the bottom wall. The remaining domain boundaries are specified-pressure boundaries, set at a gauge pressure of zero [29]", " The agreement is reasonable, and predictions of the contact line radius fall within experimental error bands. The maximum spreading, as well as the time to attain it, are accurately predicted by the model. The predicted droplet shapes also match reasonably well (Figure 3). The experimental droplet images lie in between the s = 0.1 and s = 0.9 contour lines obtained from the numerical solution, validating the numerical methodology. However, due to the asymmetry of the needle, as seen in the inset of Figure 1 and Figure 5, some of the droplets are seen to have non-axisymmetric shapes. Copyright \u00a9 2011 by ASME Downloaded From: http://proceedings.asmedigitalcollection.asme.org/pdfaccess.ashx?url=/data/conferences/interpack2011/70206/ on 05/26/2017 Terms of Use: http://www.asme.org/about-asme/terms-of-use 6 The predicted time history of the normalized height of the droplet and the contact radius are shown in Figure 7. The contact radius and height are normalized based on the initial and final states given by 0 50 t ms t ms r r r r and 50 0 t ms t ms h h h h respectively" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0000913_j.proeng.2012.01.1070-Figure3-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0000913_j.proeng.2012.01.1070-Figure3-1.png", "caption": "Fig. 3. Structure response with no initial gap: (a) temperature field; (b) von Mises stress of the heat pipe", "texts": [ " After the imposition of the boundary conditions procedure, the assembled equations can be solved by traditional techniques In mechanical analysis, the weak form of governing equation on elemental level can be expressed by: N 1 d d d d e b e eb e e N e e e e eb e e e e e e u u eb g w w u u n w f w n (4) Where Ng denotes the gap function between element e and its neighbour eb, u is the penalty parameter. In the present paper the temperature and stress fields of each contact body are considered alternatively. In our study, since material properties are temperature-dependent, and thermal contact resistance depends on both of the temperature and contact pressure, the iterative method is employed to solve the nonlinear equilibrium equations. 4. Results and discussions It can be seen from Fig. 3 that the stagnation point temperature is 1050 K, that is just 50 K higher than the temperature of the working fluid of the heat pipe, and is totally under the capacity of the C/C material. But the maximum von Mises stress of the heat pipe can up to 1150 MPa, which is far beyond the allowable temperature of the heat pipe material, and is very dangerous. In order to reduce the thermal mismatch stress and smooth the stress distributions, an initial radial gap between the heat pipe and the C/C material is widely used", " As the exact analysis of thermal structures with initial gaps always refers to contact problems, it is really a hard work, so in the present paper the effective thermal expansion coefficient of the heat pipe is applied to simulate the effect of initial gaps: eff gap work 0= T T r (5) Where is the real thermal expansion coefficient of the heat pipe material, and gap is the width of the initial gap, workT is the temperature of the working fluid of the heat pipe, 0T is the reference temperature, and r is the outer radius of the heat pipe. (a) The effect of thermal contact resistance with different inital gaps are given in Fig. 4. Numerical results show that, using heat-pipe-cooled leading edge can greatly reduce the maximum temperature to make sure that the stagnation temperature is in the tolerance zone (as shown in Figure 3) , which is a very effective thermal protection concept. At the same time, using initial gap can greatly reduce the stress level of the structure, while increase the interface thermal contact resistance and stagnation temperature(as shown in Figure 4a). Structure strength and temperature should be considered simultaneously with the thermal contact resistance in structure safety evaluation of structures with initial gaps. If the inital gap is too small, the von Mises stress of the structure may cause structural failure because of the huge difference of thermal expansion coefficient of heat pipe and C/C material" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0001493_ines.2010.5483842-Figure9-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0001493_ines.2010.5483842-Figure9-1.png", "caption": "Figure 9. The trajectory of the armature-current in natural stator-fixed coordinate frame.", "texts": [], "surrounding_texts": [ "For the self-commutation of the Ex-SyM it is important an accurate information about the rotor position. This was realized using an incremental encoder, mounted on the motor shaft. The mounting is realized in a manner, that the encoder index signal is synchronized with respect to the rotor position (i.e. the direction of the field-winding). The encoder generates a number of pulses proportional to the angular position of the shaft. It gives information also about the sense of the rotating motion: positive values for direct and negative ones for reverse running. The counter resets to zero at every full rotation. The amplitude of this saw-tooth signal will be equal to the encoder resolution. In order to obtain a continuous signal, useful in the present application, where the Coordinate Transformation blocks use the sine and cosine of the angular positions, this enc\u03b8 encoder position signal has to be processed accordingly, in order to obtain a position signal \u03b8 between ]2 ,0[ \u03c0 for direct, and ]2 ,0[ \u03c0\u2212 for reverse rotation respectively, based on the following expression: \u03c0 \u03b8 \u03b8 2 r enc N = , (4) where Nr is the number of increments/revolution (the encoder resolution). Applying the sine and cosine trigonometrical functions to the above position signal, continuous sine wave signals are obtained that the period is equal to 2\u03c0 mechanical angle. On the other hand, the computation of the angular speed, as a derivative of the position \u03b8 is not accurate, because of the zero crossing, where the obtained speed signal is infinite, and it cannot be processed by the controller board. In the following, a simple procedure is presented that avoids the zero crossings, and gives an accurate result. The derivatives of functions \u03b8sin and \u03b8cos are also shifted with 2/\u03c0 . The magnitude of the obtained signals will be equal with the angular speed, \u03c9 . The sign of the position signal \u03b8 will give the direction of the rotation. The computation of the angular speed is based on expression: ( ) ( ) ( ) ( )\u03b8\u03b8\u03b8\u03b8\u03c9\u03c9 sign dt d dt dsign 22 cossin \u239f \u23a0 \u239e \u239c \u239d \u239b+\u239f \u23a0 \u239e \u239c \u239d \u239b== . (5) and its simulation may be processed with the structure presented in Fig. 1 [4]. IV. DOUBLE-FIELD-ORIENTED VECTOR CONTROL" ] }, { "image_filename": "designv11_62_0002302_b978-0-12-803581-8.10351-0-Figure68-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0002302_b978-0-12-803581-8.10351-0-Figure68-1.png", "caption": "Fig. 68 Pro/manufacture model of lower strake piece and work piece.", "texts": [ " It was decided that this would be beneficial given the potential difficulty in removing a mold from this geometry. The manufacture of the sidepod entry tooling followed the procedure of the ergonomic backboard, with the exception that the geometry was thin enough that each of the two left and right side tools could be produced from a single block of tooling foam. In this case, these tools were manufactured without integral vacuum flanges. Each inlet was tooled, and molded, in two pieces due to the contour which did not lend itself to a simple removal direction/parting plane. Fig. 68 shows the small lower strake tool simulated within the block of foam and the machined tool is shown, immediately after machining in Fig. 69. Machining was accomplished by rough cutting with a 0.7500 diameter end mill, followed by finish machining with a 0.500 ballnose end mill. The foam tooling was sanded to 1500 grit, coated with polyurethane primer paint and sanded down to a fine finish. The final four tools for the sidepod entry are shown in Fig. 70. The bulkhead designs described in detail in the previous chapter are crossframes spanning the width of the monocoque at several locations along the length" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0003269_tmag.2019.2899725-Figure3-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0003269_tmag.2019.2899725-Figure3-1.png", "caption": "Fig. 3. Model of a problem consisting of a copper shell with a cylindrical coil settled in the center of the annulus. (a) Top view with dimensions. (b) 3-D view with symmetry plane at the bottom.", "texts": [ " This will be shown in Section IV to suppress any additional oscillations in the solution. Numerical investigations are carried out on an example involving a cylindrical shell made of copper included in the stationary annulus stat, and a cylindrical coil placed in the center of a cylinder constituting the moving domain mov. There is a symmetry plane perpendicular to the axis of the shell, and the axes of the coil and of the shell are at the right angle to each other. Table III summarizes the dimensions of the problem model, as shown in Fig. 3. Note that, the volume of the moving domain has an invariant cross section at the right angles to the direction of motion. This enables the verification of the time transient simulation at steady state by the single stationary solution obtained by the A, V \u2013V potential formulation including the motion term u \u00d7 B(Minkowski transformation [5]). The time transient simulation is carried out by the T, \u2013 formulation, where the curl of the current vector potential T describes the eddy current density J in the conducting domain with the excitation current density written as J0 = curlT0 [6]" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0001983_s1061934813090049-Figure2-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0001983_s1061934813090049-Figure2-1.png", "caption": "Fig. 2. Cyclic voltammograms obtained with 2.7 \u00d7 10\u20133 to 10.8 mM NADPH solutions prepared in 0.1 M, pH 7.4, PBS buffer solutions and the calibration line obtained from I currents at +0.225 V.", "texts": [ " JOURNAL OF ANALYTICAL CHEMISTRY Vol. 68 No. 9 2013 GLUTAMATE SOL\u2013GEL AMPEROMETRIC BIOSENSOR 797 The ability of the chemically modified electrode to reduce the oxidation potential of NADPH was after wards evaluated at different pH conditions with cofac tor concentrations varying from 2.70 \u00d7 10\u20133 mM to 10.74 mM. A progressive shift of the oxidation poten tial in negative direction was observed with pH in crease but best results concerning sensitivity were ob tained with NADPH dissolved with 0.1 M, pH 7.4 PBS buffer solution (Fig. 2). An evident increment propor tional to cofactor concentration up to the highest as 798 JOURNAL OF ANALYTICAL CHEMISTRY Vol. 68 No. 9 2013 GOMES et al. sayed concentration was noticed as a consequence of polymer reduction by the cofactor together with the decrease of the cathodic current when sweeping the potential in opposite direction. For the higher concen trations of NADPH even a mediated spontaneous ox idation occurred at null potential. This behaviour is not observed at the bare electrode if MG is simply add ed to the solution of the cofactor" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0001189_jtam-2013-0001-Figure5-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0001189_jtam-2013-0001-Figure5-1.png", "caption": "Fig. 5. Belt pulley \u2013 position 2", "texts": [ " This force has the following relation with respect to the force P \u03c4 b : (5) Rn b = mP \u03c4 b , Unauthenticated Download Date | 5/15/17 2:19 PM where m is a coefficient, dependent on the state of the band-saw blade (m = 0.5 for a sharp band-saw blade and m = 1 for a very blunt band-saw blade). R\u03a3 is the total resistance force on the woodworking detail. This force is calculated for each individual case. Figure 4 shows the woodworking detail and the forces and the velocities described above. The weight of the feeding wheel 3 is: (6) Ge 3 = m3g, where m3 is its mass and g is the acceleration of gravity. The belt pulley position 2 is shown in Fig. 5, below: This belt pulley transmits the driving moment from the electric motor to the basic shaft. The force P e 2 creates the driving moment Me 2z with respect to the axis of rotation. It is calculated from the known expression, cited below: (7) Me 2z = P e 2 r2. Unauthenticated Download Date | 5/15/17 2:19 PM This force is equal to the force P e 1 in the belt gear. The electric motor creates a moment which can be calculated from the following expression: (8) Me 1 = Ne/\u03c9e, where Ne is the power of the electric motor for cutting and feeding of the woodworking detail, \u03c9e is the angular velocity of the electric motor", " The moments of this force to the coordinate axes are calculated by the following relations: Me R3x = Re 3ze sin \u03d5 = (Rn b + R\u03a3) 2 e cos \u03b1 sin \u03c9t, Me R3y = Re 3z(r3 + e cos \u03d5) = (Rn b + R\u03a3) 2 cos \u03b1(r3 + e cos \u03c9t),(17) Me R3z = 0. Unauthenticated Download Date | 5/15/17 2:19 PM The moments of the weight Ge 3 with respect to the coordinate axes are calculated by the following dependencies: Me G3x = Ge 3e sin \u03b1 = m3ge sin \u03b1, Me G3y = 0,(18) Me G3z = Ge 3e cos \u03b1 cos \u03d5 = m3ge cos \u03b1 cos \u03c9t. The force P e 3y is calculated from the dependence below: (19) P e 3y = 2F e 3 + P e 3 = 2F e 3 + K\u2206(\u03bb)bHu V . The components P e 2x and P e 2y of the force P e 2 shown in Fig. 5, are determined by the dependencies below: (20) P e 2x = P e 2 cos \u03b2 = Ne cos \u03b2 \u03c9r2 , P e 2y = P e 2 sin \u03b2 = Ne sin\u03b2 \u03c9r2 . The components P e 5x and P e 5y of the force P e 5 shown in Fig. 6, are determined by the following relations: (21) P e 5x = P e 5 cos \u03c7 = Puu cos \u03c7 \u03c9r5 , P e 5y = P e 5 sin \u03c7 = Puu sin \u03c7 \u03c9r5 . The support reactions are determined by the equilibrium conditions. (22) \u2211 zi = 0, \u2211 MAxi = 0, \u2211 MBxi = 0, \u2211 MAyi = 0, \u2211 MByi = 0. We write equations (22) as follows: Re 3z \u2212 Ae z = 0, Me R3x + Me G3x + ( Ge 3 \u2212 P e 3y ) a1 + P e 5yb1 \u2212 Be y(b1 + c1) \u2212P e 2y(b1 + c1 + d1) = 0, Me R3x + Me G3x + ( Ge 3 \u2212 P e 3y ) (a1 + b1 + c1) + Ae y(b1 + c1) \u2212P e 5yc1 \u2212 P e 2yd1 = 0,(23) \u2212Me R3y \u2212 P e 5xb1 + Be x(b1 + c1) \u2212 P e 2x(b1 + c1 + d1) = 0, \u2212Me R3y \u2212 Ae x(b1 + c1) + P e 5xc1 \u2212 P e 2xd1 = 0" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0001482_iros.2011.6095088-Figure2-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0001482_iros.2011.6095088-Figure2-1.png", "caption": "Fig. 2. The backbone curve of a flexible strip is the curve running along the center line of the strip from one end to the other. The endto-end transformation along the curve can be represented with a homogeneous transformation matrix, .", "texts": [ " The Kinematics of Sheet Bending A sheet of material that is too stiff to deform in the plane but capable of bending is often referred to as a ruled or developable surface. The internal constraints imposed by such shapes force the bending curvature of the surface to be simple \u2013 that is, it is possible to draw straight lines of zero curvature through any point on the surface that lies completely along the surface, resulting in the local area being equivalent to a section of a cone. A side effect of this constraint is that the deformation of a whole sheet can be described by the bending occurring on a single backbone drawn across the sheet, as shown in Fig. 2. Such a curve, running end-to-end between two links of a robot, can be parameterized in terms of its arc length, 0, . The local coordinate frame of this backbone curve forms a rotation matrix, , whose local axes are depicted in Fig. 2, with tangent to the backbone and perpendicular to the backbone but along the sheet. Again, there is assumed to be no extension or shearing in the plane, only bending of magnitude about a principal axis of rotation at some angle, , as shown in Fig. 3. The change in as a function of is given by the well-known relationship: (1) The matrix is the skew-symmetric form of the rotation rate vector, 0 cos 0cos 0 sin 0 sin 0 (2) The evolution of the backbone position, , is also defined in terms of the local coordinate axes as described by , 100" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0002791_b978-0-12-812138-2.00003-9-Figure3.6-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0002791_b978-0-12-812138-2.00003-9-Figure3.6-1.png", "caption": "FIGURE 3.6", "texts": [ " (A) Stator core and (B) sinusoidally distributed winding and mmf. Squirrel-cage rotor. This rotor gets its name from its structure\u2019s resemblance to a squirrel cage. The rotor conductors are not placed exactly parallel to the shaft but are skewed by one slot-pitch to reduce cogging torque, and this allows the motor to run quietly. Because of its simple and rugged construction, about 95% of the induction motors use the squirrel-cage rotor. Similar to the windings on a stator, a wound-rotor as shown in Fig. 3.6 has a set of three-phase windings, which are usually Y-connected. The rotor windings are tied to the slip rings on the rotor\u2019s shaft and thus can be accessible through the brushes. Due to this configuration, in a wound-rotor type induction motor, the rotor resistance can be varied by connecting external resistors to the rotor windings via the brushes. This allows the torque speed characteristics of the induction motor to be varied as needed. Now, we will discuss in detail the squirrel-cage induction motor, which is the most common type of an induction motor", " This results in a lower slip and thereby a higher efficiency. Therefore we can see that there are conflicting requirements of rotor resistance for a satisfactory performance, i.e., a high resistance at starting and low resistance at the normal operating speed. For this reason, different types of rotors have been developed to achieve a variation in rotor resistance with operating speed. In a wound-rotor induction motor, external resistance can be connected to the rotor windings through slip rings as shown in Fig. 3.6. Thus a satisfactory performance in both the starting and running conditions can be obtained by varying the value of the external resistance with the rotor speed. For squirrel-cage induction motors, the rotor resistance depends on the operating speed. This is because the effective resistance of the rotor conductor may be changed with the rotor frequency due to the skin effect. The skin effect is the tendency of AC current to flow near the surface of the conductor. The skin effect causes the effective area of the cross section of the conductor to reduce at higher frequencies, thus increasing the effective resistance of the conductor" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0001661_tasc.2011.2109050-Figure4-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0001661_tasc.2011.2109050-Figure4-1.png", "caption": "Fig. 4. Cross section of the levitated conveyer.", "texts": [ " By installing bigger electromagnets over the levitated conveyor, 1051-8223/$26.00 \u00a9 2011 IEEE pulse-field magnetization is easily applied to superconductors in the levitated conveyor and larger pulse-field magnetization is easily obtained. Since the magnetic flux density distributions in superconductors are uniform along the magnetic rails, it is necessary for the electromagnets to produce magnetic flux densities similar to those from the magnetic rails. The specifications of electromagnets are shown in Table III. D. Levitated Conveyor The levitated conveyor is shown in Fig. 4. The conveyor is a rectangular container made of stainless steel for saving liquid nitrogen. Four superconductors (SmBaCuOx, ) are installed at the bottom of the container for levitation. In the experiment, the superconductors are field-cooled by using liquid nitrogen. The magnetic flux densities were measured at a distance 1.0 mm from the magnetic rail surface. The measured area consists of a magnetic rail and a half of the LIM. Fig. 5 shows the magnetic flux density distribution at a distance 1" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0001738_1.5062560-Figure1-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0001738_1.5062560-Figure1-1.png", "caption": "Figure 1: Schematic diagram of a laser cladding setup and model-part domains", "texts": [ " Complete modelling of the Laser Direct Metal Deposition process requires the application of two different computational-mathematical methods \u2013 computational fluid dynamics and the finite element method. This is due to the nature of the modelled processes: LDMD consists in parts of fluid behaviour (e.g. powder stream, melt pool and clad creation) and in other parts of solid behaviour (solidified clad, metallurgical evolution, stress and distortion). The model presented here is based on four modelparts; three of which are based on CFD and one of which is FEM-based. A schematic diagram of the modelling domains is shown in Figure 1. Model-part 1 encompasses the powder-feeder and nozzle and marks the beginning of a chain of partial models along the laser cladding process. Since Part 1 of the model does not encounter any interfering influence from outside of this system it may be assumed a stable system, which allows it to be simulated by a steady-state method, where the state of gas pressure, velocity and powder mass flow rate do not change during the process. The model tracks individual parcels and assumes one-way coupling between the gas and particles, as the volume fraction is very small" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0002740_2017.11.45-Figure1-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0002740_2017.11.45-Figure1-1.png", "caption": "Figure 1. Schematic illustration of the stepwise fabrication process and testing principle of the novel H2O2 electrochemical biosensor", "texts": [ " The suspension of PCN-333(Al)-GO composite (2 mg mL -1 ) was dropped on the surface of polished glassy carbon electrode (GCE). Finally, the as-prepared MP-11/GO-PCN-333 (Al) /GCE was obtained by immersing PCN-333(Al)-GO/GCE in a MP-11 solution (1 mg mL -1 ) for 12 h and then dried in air. To further improve the performance of the electrode, the resulted electrode was rinsed with fresh buffer solution for several times to remove the weakly absorbed molecules and then dried in air. The detailed process of the MP-11/PCN-333(Al)-GO/GCE was shown in Fig. 1. Int. J. Electrochem. Sci., Vol. 12, 2017 10393 A three-electrode system of a platinum wire as the counter electrode, a saturated calomel electrode (SCE) as the reference electrode and a bare or modified GCE as the working electrode was used. Cyclic voltammograms (CVs) were operated in 0.1 M KCl solution containing 5.0 mM Fe(CN)6 3\u2212/4- at room temperature or N2-saturated 0.2 M PBS (pH 7.0). Electrochemical impedance spectroscopy (EIS) was performed in 0.1 M KCl solution containing 5.0 mM Fe(CN)6 3\u2212/4- at open circuit potential in the frequency range from 0" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0003950_s11029-019-09824-x-Figure1-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0003950_s11029-019-09824-x-Figure1-1.png", "caption": "Fig. 1. Schematic of the experimental setup.", "texts": [ " This can be substantiated by the results of experimental studies on slow and fast cyclic deformation of plexiglass performed by the Authors, where the coefficient of energy absorption in the slow process turned out to be by an order of magnitude smaller than that obtained in the fast deformation process. For studying the hereditary properties of rubber in short-term shear creep, it is advisable to use vertically fixed test specimens (in order to exclude the static component of deflection caused by the gravity force of test specimens). For this purpose, an experimental setup developed previously was modernized [25], and its schematic is shown in Fig. 1. The installation consists a base (1) and a force rack (2), rigidly connected together. A cantilever (3) with a grip (4) at its end is fixed on the force rack. The test specimen (5) is clamped with the help of separated rigid plates connected to the cantilever by bolted joints, preventing the test specimen from rotation in the cross section of its fixation. For mounting a laser sensor of displacement (7), a mobile platform (6) is installed on the force rack, whose position along the force rack can be changed for measuring the deflection w of the free end of test specimen when its arm of crane changes" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0003753_s40964-019-00095-5-Figure1-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0003753_s40964-019-00095-5-Figure1-1.png", "caption": "Fig. 1 Section of a powder bed in a PBF-LB system during exposure with the relevant process parameters", "texts": [ " All samples examined in this study were generated with an EOSINT M270 PBF-LB system (EOS GmbH, Germany). The powder material used was a commercially available EN 1.4404 (ASTM 316L) stainless steel (d50 = 26.28\u00a0\u00b5m; d90 = 42.56\u00a0 \u00b5m). The energy is supplied to the powder bed with a 200\u00a0W Yb fiber laser with a wavelength of \u03bb = 1064\u00a0nm, which generates a Gaussian beam profile in continuous wave operation. Recommended settings of the system-integrated beam expander for this material lead to a laser spot diameter (Fig.\u00a01) of \u03bb = 132.8\u00a0\u00b5m at the level of the layer to be scanned. The exposure of the individual powder layers was carried out according to a stripe pattern with parallel scanning vectors (Fig.\u00a02). From powder layer (layer n\u00a0\u2212\u00a01) to powder 1 3 layer (layer n), the orientation of the scanning vectors rotated by 67\u00b0 along the z-axis (building direction). For the generation of fully dense samples, an area energy density EA (formula\u00a01) of 2.00\u00a0J/mm2 was applied. The designated layer thickness (Fig.\u00a01) is 20\u00a0\u00b5m and was kept constant for the sample production of this study. Design of experiments (DoE) was used to determine an appropriate scope of experiments and to verify and interpret the results. This section describes the creation of the test design to investigate the influence of selected process parameters on pore characteristics in thin-walled porous structures of 316L. The energy input per area into the powder bed by the laser beam, referred to as area energy density EA, is defined as In Fig.\u00a01, the process parameters contained in Formula\u00a01 are shown graphically. P describes the power of the laser beam, which exposes the powder surface with the scan speed vs in the scanning tracks of a defined distance to each other, designated as \u201chatch distance h\u201d. In the course of preliminary investigations, a process window was determined which is spanned by the parameters P, vs and h for staking out the design space (Fig.\u00a03). To enable a well-arranged two-dimensional visualization, the line energy density EL is found as a quotient of laser power P and scan speed vs on the abscissa" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0003015_2017-36-0413-Figure3-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0003015_2017-36-0413-Figure3-1.png", "caption": "Figure 3. Off-axis rupture strength. [9]", "texts": [ " [7] Description of the elements at the rear bumper The materials used in the simulation of the rear bumper were the 1020 steel and epoxy composite with glass fibers. Tables 2 and 3 describe the mechanical properties of the composite material and the metallic material. The orientation of the fibers causes influence on some of the properties of the composite material. Unidirectional composites have an important elastic deformation, which is a good advantage in some cases. Unidirectional materials oriented at 0\u00b0 have greater resistance to rupture, this resistance tends to decrease according to the increase of the orientation angle of the fiber, as shown in Figure 3. [9] The study used as a reference the dimensions of the Mercedes Benz Atego 1719 4x2-truck model with total gross weight of 16000kg. [10] The bumper is made of AISI 1020 carbon steel with dimensions of 2400mm in length, 200mm in height and 2mm in thickness. The vertical support bracket also uses the same material, with ratios 60x600x150mm and 2mm thick. Figure 4 shows the final component of the rear bumper assembly of the truck. For the glass fiber reinforced epoxy composite the dimensions of the main beam and flaps were the same, except for the thickness of the laminate, which in this case will be 3mm (1mm each layer)" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0002054_amer.math.monthly.124.3.265-Figure8-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0002054_amer.math.monthly.124.3.265-Figure8-1.png", "caption": "Figure 8. An infinite Kepler\u2013Poinsot polyhedron \u201cof Archimedean type\u201d: the \u201cregular {6,8}\u201d with open cube tunnels, a {4, 6, 4, 4, 6, 4}.", "texts": [ " Although this may help to explain the 3D image, it doesn\u2019t seem to help in, for instance, getting the dual {8, 6} of the {6, 8} (see below). Another way to get a better understanding for the shape of this regular, infinite, Kepler\u2013Poinsot polyhedron is to introduce prismatic tunnels (open cubes) between the open cubohemioctahedra. This turns the {6, 8} into an \u201cArchimedean type\u201d infinite Kepler\u2013Poinsot polyhedron {4, 6, 4, 4, 6, 4}, that is, a polyhedron composed from two types of regular polygons, in this case, hexagons and squares (see Figure 8). One might think about obtaining yet another new regular polyhedron by forming the dual of the {6, 8} as that should be an {8, 6}. \u201cDual\u201d is used here in the sense that the March 2017] NOTES 267 This content downloaded from 134.148.10.13 on Tue, 07 Mar 2017 06:17:19 UTC All use subject to http://about.jstor.org/terms vertices of one polyhedron correspond to the faces of the other. Thus, the tetrahedron is self-dual, while the octahedron and the cube, the icosahedron and the dodecahedron, the small stellated dodecahedron and the great dodecahedron, and, finally, the great stellated dodecahedron and the great icosahedron are duals of each other" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0002776_cadiag.2017.8075685-Figure1-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0002776_cadiag.2017.8075685-Figure1-1.png", "caption": "Figure 1. Stator and rotor references frames.", "texts": [ " 1;2;3;4 vib mis rf n f n = = (1) When a p pair-poles IM is supposed to be supplied by a perfect balanced power supply with main frequency sf , the distributed stator coils will generate a set of harmonics ( 6 1k\u03bd = \u00b1 ) of rotating magneto-motive forces (MMF) at each angular position, referred into the stator reference frame by s x\u03b1 [16,17] : ( , ) K . .cos( . )s s x s s xM t I t p\u03bd \u03bd\u03b1 \u03c9 \u03bd \u03b1= (2) with K\u03bd constant depending on the machine\u2019s winding. Regarding the stiff relation linking the stator reference to the rotor reference (Fig.1): s r x x r\u03b1 \u03b1 \u03b8= + , (3) and the expression of the IM\u2019 slip (s), the MMF could be reformulated by: [ ] 0( , ) K .cos( . ) 1 .(1 ) . r q r x s x m n s r M t I t p m s n \u03bd \u03bd \u03bd\u03b1 \u03bd \u03b1 \u03c9 \u03c9 = \u03a9 \u03a9 = \u2212 \u2212 \u2212 . (4) These MMF harmonics are superposed to the fundamental one, and are rotating into the forward sense ( 6 1k\u03bd = + ) as well as into the backward sense ( 6 1k\u03bd = \u2212 ). Given a supposed constant air-gap length (g), the MMFs interact with the each kth rotor\u2019s mesh ( 1, 2,... rk N= ) to produce circulating currents into the rotor\u2019s bars and the rotor\u2019s end-rings of the form: ( )0( ) ( , )" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0002274_s12206-017-0308-9-Figure10-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0002274_s12206-017-0308-9-Figure10-1.png", "caption": "Fig. 10. Combined path pattern for solid filling cylindrical part: (a) CAD drawing of the combined path pattern by spiral strategy; (b) CAD drawing of the combined path pattern by zig-zag/raster strategy; (c) the first deposited layer using the zig-zag/raster pattern.", "texts": [ " 8(a)-(c) show the SMDM path procedure used for produce the hollow filling area of square part. Figs. 9(a) and (b) show another path pattern used to produce this part. The real deposited layer of this part is shown in Fig. 9(c). Furthermore, the CAMDM system ready to create the Solid filling area features. In order to show the two types of the path patters the deposition of the solid circular feature part was carried out by both the spiral pattern and the zigzag pattern as shown in Figs. 10(a) and (b), respectively. Fig. 10(c) presents the first few layers of the real deposited part manufactured using zig-zag path pattern. The control system used in the SMD machine is a feed forward (open-loop) control type. Post-processing covers 3D object modeling based on CAD drawing, planning and slicing of the pattern in uniform thickness layers. The CAD drawing is imported using an on-line programming method. So the desired geometry according to the CAD-specification is sliced into layers of uniform height. For each layer the paths along which the beads should be deposited to fill the layer [41]" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0003788_msf.969.756-Figure1-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0003788_msf.969.756-Figure1-1.png", "caption": "Fig. 1. Specimen Work material and Tool holder", "texts": [ " [8] used Taguchi method with GRA in the optimization of turning operations with multiple performance characteristics. From the above studies, it has been found that the GRA is one of the important optimization techniques and it could be successfully applied in the machining process, such as turning. In this context, the present study has been an attempt to optimize the process parameters in the turning of Inconel 625. The work material selected for this experiment was Inconel 625. The work specimen is a circular bar in the shape of 60 mm long and diameter 27 mm, as shown in Fig. 1 (a). The cutting tool used was PVD coated carbide insert having ISO coding CNMG 120408 HRM AH 8015, manufactured by Tungaloy, Japan. Tool holder and insert geometry are depicted Fig. 1(b). Independent variables conceptualised for the experiment were cutting speed, feed and depth of cut. The response variables were material removal rate and surface roughness. The levels of machining parameters were selected based on literature review, recommendation of tool manufacturer and pilot studies conducted before performing actual experiments. The independent variables and their respective designations are represented in Table 1. C depth of cut mm 0.2 0.4 0.6 0.8 The process of dry turning was done on Smarturn CNC lathe with Fanuc CNC control system" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0000802_kem.450.27-Figure3-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0000802_kem.450.27-Figure3-1.png", "caption": "Fig. 3 Product design- in back view", "texts": [ " Strengthen the folding points It was suggested to use high-density rubber pads to strengthen folding points, make tenons tense and fixed, and make the frames more firm and closed together. Then, the stability and safety during riding might be increased. 5. Use big and small gear wheels combination If using the same size of gear wheels, it would be inefficient and slow when riding. Therefore, imbalanced gear wheels combination was set to adjust and bring greater speed as well as increasing the efficiency of riding. The new designs of the improved folding bicycle in this study are shown in Fig. 3 and Fig. 4. The Contradiction Matrix of TRIZ is a powerful tool, especially when the 40 Inventive Principles can assist designers create innovative ideas. In this research, principles in TRIZ were applied to improve the capability and degrees of comfort of the folding bicycles. The main points in the designs of this study are double shock absorbers, removable hidden stand and asymmetric wheels.. The main goal of double shock absorbers is to balance the supports from both sides, to concern the structure system of the frame, and to achieve the vibration-absorbing effect" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0003803_s0025654419010023-Figure11-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0003803_s0025654419010023-Figure11-1.png", "caption": "Fig. 11.", "texts": [ " Similar to the case of blocking or slipping of the front wheels and by virtue of the considered model for skidding of both axes of the apparatus under any initial conditions given via the variables vy, \u03c9z, \u03c91 at moments t \u223c 1 (T \u223c T2), the device enters the non-slip mode of unlocked (non-slip) wheels (in the longitudinal and transverse directions). It should be noted that the mentioned wheels are the front ones. Further, the movement proceeds according to the scenarios of skidding with locked or spinning rear wheels with non-slip front wheels from sections 3.3, 3.4. Let us discuss how the rotation of the front wheels of the vehicle affects its skid. Fig. 11 shows straight line (4.19) for \u03b4 = 0 and straight lines (4.19) symmetrically located relative to it, which correspond to the same value and different signs of the angle \u03b4 of rotation of the front wheel of the bicycle model of the apparatus. When choosing any of the listed straight lines, the movement along the phase trajectories (4.18) occurs towards it. The area of applicability of the considered model of motion of the apparatus, determined by the constraints (4.9) (it is inside and a small external neighborhood of the circle in Fig. 11) is almost completely contained in the strip i2z a \u03c9z \u2212 vx|\u03b4| < vy < i2z a \u03c9z + vx|\u03b4| (4.20) MECHANICS OF SOLIDS Vol. 54 No. 1 2019 where the condition (3.14), which in dimensionless variables has the fom sign\u03c9z = \u2212sign\u03b4 (4.21) cannot be satisfied at all points during the movement to the straight line (4.19). Thus, the skidding of the apparatus as a whole will be minimal when \u03b4 = 0. If the initial values of the variables vy, \u03c9z meet the condition C2 = 0, then the skidding of the apparatus will be completely stopped at moments of time t \u223c 1" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0000055_s12346-012-0085-x-Figure2-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0000055_s12346-012-0085-x-Figure2-1.png", "caption": "Fig. 2 System (2) in Example 2", "texts": [ " (2) We remark that M(x, y) = x2 + y2 +1, N (x, y) = 2M(x, y)(1\u2212 x4)\u2212 x3, F(x) = x2, g(x) = x3 and h(y) = y3. Taking the supplement function \u03d5(x) = x6, we see that the condition (C2) is satisfied for |x | \u2264 \u03b4, where \u03b4 is the constant in the interval (0, 1) satisfying the equation 2\u03b44 + \u03b43 \u2212 2 = 0. In facts, we have N (x, y) > 2(1 \u2212 x4) \u2212 x3 > 0, F(x) \u2212 \u03d5(x) > 0 and \u03d5 \u2032 (x)[F(x) \u2212 \u03d5(x)] dh dy [h\u22121(\u03d5(x))] \u2212 N (x, y)g(x) M(x, y) = x6 M(x, y) > 0 for 0 < |x | \u2264 \u03b4. Thus, we conclude from Theorem 1 that the system has local homoclinic orbits. See Fig. 2 for the phase portrait of the system. Example 3 Consider the Li\u00e9nard perturbation system \u03b5 x\u0307 = y \u2212 \u03bb ( x3 3 \u2212 x2 2 ) , y\u0307 = \u2212k(x3 \u2212 a), (3) where \u03b5 > 0, \u03bb > 0 and k > 0. The case of a = 0 and \u03bb = 1 has been treated in [3] or [6]. By using some transformation, system (3) is transformed to the following system x\u0307 = y \u2212 \u03bb 6 x{2x2 + 3(2\u03b1 + 1)x + 6\u03b1(\u03b1 + 1)} y\u0307 = \u2212\u03b5kx(x2 + 3\u03b1x + 3\u03b12), (4) where \u03b1 is a real constant satisfying the equation \u03b13 + a = 0. Then we note that the system has a unique equilibrium point (0, 0)" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0002053_caidcd.2010.5681281-Figure4-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0002053_caidcd.2010.5681281-Figure4-1.png", "caption": "Figure 4. multi-body dynamic model of off-road vehicles", "texts": [], "surrounding_texts": [ "random input under the soft terrain The deformation road has the impact to the ride comfort of the off-road vehicle, call the TIRSUB.DLL dynamic database of the tire subroutine in the ADAMSNIEW. In the c-class road, with speed of 18km/h, the simulation is performed on different roads that defined according to the parameters of Table 1. Fig.8 shows that, with the surfuce hardness decrease, the tire center subsidence increase, the curve became smoother, and there is a lag in the phase. Fig.9 shows that the front tire load have a certain amplitude at the beginning, then trend to stable, and they have a center amplitude stability in the range of body weight due to the incentive of the ground is not flat to the vehicle. This is consistent with the reference's conclusion [I]; the results of dynamic simulation are shown in Table 3. From the Table 3, the vehIcle under the same speed, when the road hardness lower the soil subsidence has been increased, which increase almost eight times from the hard road to the soft road. The displacement Rms of the vehicle mass center has also increased three times form the hard road to the soft road, this show that the vehicle travelled on the soft terrain has not only the tire deformation but also has the surface deformation. The increase of the soft soil subsidence cause the vehicle body vertical amplitude increase, and the soil is more softer the soil subsidence is more greater, leading to greater vertical amplitude of body. Besides, tire load and suspension deflections were changed because that road roughness excitation makes the vehicle and tire to vibrate and has influence on the vehicle's ride comfort. With the same suspension stiffuess and damping, the vehicle body acceleration and suspension deflection change with the soil soften. So in this trend it can be seen the suspension stiffuess and damping suit to the hard road but do not suit to the soft road, the suspension stiffuess and damping values should be design by the road which the off vehicle used drive. VI. CONCLUSION Based on interaction between the tire and soft pavement, the subroutine that calculates the force between the tire and soft terrain is established to realize the calling in ADAMS, and then the subroutine is used in the simulation of whole vehicle model. The ride performance characteristics in all states were recognized and analyzed by means of simulation when the vehicle runs the same speed and different deformation degree of the road. The simulation result shows that the model was validity and feasibility, and could predict the nde comfort charactenstIcs of the vehIcle that runs on the soft terrain. According to the characteristics got by simulation, it showed the nonlinear properties of the soft terrain has great influence in the ride comfort. The research offers the experience of construction design and analysis on research of the ride performance of the vehicle runs on soft pavement. The model was only considered the pressure model of the soft terrain, and was not considered the shear deformation of the soft soil which has influence on the vehicle's ride comfort. Moreover when the off-road vehicle was driving on the road the rear tire usually rolls in the track that the front tire left, so the front and rear tire have the different soft soil conditions when traveling on the road. But in this model the front and rear tire have the same soft soil conditions. Establishing the model that take into account the shear deformation of the soft soil and the pressure model which under the repeat loading can be more comprehensive and accurate to assessment the ride comfort of the vehicle when it driving on the soft terrain .It is need further research in the future. ACKNOWLEDGEMENTS This research was supported by the Specialized Research Fund Doctoral Program of Higher Education of China under grant no.20070532008 and the Science Fund of State Key Laboratory of Advanced Design and Manufacturing for Vehicle Body of Hunan University of China under grant nO.60870002." ] }, { "image_filename": "designv11_62_0002053_caidcd.2010.5681281-Figure3-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0002053_caidcd.2010.5681281-Figure3-1.png", "caption": "Figure 3. The model of the equivalent tire interaction with the soft terrain", "texts": [ " Schmid can be used to describe the tire interaction with the road [8]. I.C. Schmid thought that when the flexible tire drive on the soft terrain, the part of the tread on the ground have the deformation under the pressure of the ground which is shown in Fig.2, from the side of the tire ,the contour of tire on the ground is a curve, this curve can be replaced by using the equivalent rigid tire which the Diameter is D' and have the same curvature with the tire on the ground contour, shown in the Fig.2. In Fig.2 and Fig.3, D is the tire diameter, L is the length that the arc of the tire contact with the ground project in the horizontal plane. Zo is the amount of sinkage, fa is the radial elastic deformation of the tire. The relationship of the tire load with the amount of sinkage: b(kc / b + krp)z(n+o5)..[if F \"., -------'----- Z n -+1 2 (2) The relationship of the equivalent diameter D' with the tire diameter: (3) According to the sinkage amount hypothesis proposed by the Soviet scholar Age Yi Jin[9] Zo = [K/\ufffd + K, /; ,where , the k K and n are equal to the (k, +bk" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0001878_detc2011-48146-Figure2-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0001878_detc2011-48146-Figure2-1.png", "caption": "Figure 2 \u2013 An example of finding the Grubler\u2019s correction using the force method.", "texts": [ "1 The force method for finding Grubler\u2019s Correction In this method we search for the maximum number of independent forces that can act in the links, such that the force equilibrium around each joint is satisfied. Since each inner force defines a set of links where there are inner forces, termed self-stress set, this method searches the maximum independent self-stress sets. This number of independent inner forces is the Grubler\u2019s correction. An example of applying the force method to calculate the exact dof is given in Fig.2. For the system in Fig.2.a, Grubler\u2019s equation yields zero dof. In Fig.2b we apply inner force in link 7 that defines the self-stress set \u2013 {7,8}. Applying an inner force within another link, not included in the existent self-stress sets (in this case link 5), yields a self-stress in links {5,6,1,2,7,8} defining the second set as shown in Fig.2 c. The only links that do not belong to any set are {3,4}. However, applying an inner force in one of them, for example in link 4, does not define any self-stress set since around joint C there cannot be equilibrium of forces, as shown in Fig. 2d. In summary, there are two self-stresses, as shown in Fig. 2e. Thus the correction number is two, meaning, there are 2 dof for this specific geometry of the system. In Fig.2f we can see the two independent motions, which in this case are both infinitesimal motions. Joint B is mobile and links {5,4,3,1} can rotate with one dof, thus the exact number of dof for this configuration is two. A CB D 1 2 5 6 43 8 7 A CB D 1 2 5 6 43 8 7 A CB D 1 2 5 6 43 8 7 A CB D 1 2 5 6 43 8 7 (a) (b) (c) (d) 2 Copyright \u00a9 2011 by ASME Downloaded From: http://proceedings.asmedigitalcollection.asme.org/ on 09/01/2017 Terms of Use: http://www.asme.org/about-asme/terms-of-use The pairs of the following links (1,2), (5,6),(7,8) and (3,4) are collinear Another method for finding Grubler\u2019s correction, i", "2 The Assembly method for finding the Grubler\u2019s correction In this method we start to assemble the system and during the assembly we find the number of self-stresses. The case where a self-stress occurs is as follows: The assembly self-stress rule : If you have to add a link during the construction of the system in such a way that it is inserted between joints in which at least two of them are immobile, we will have an additional self-stress For the sake of clarity, we apply the assembly method to the same problem that we had applied to the force method, as appears in Fig. 2. It is easy to verify that there is no problem assembling the first 5 links, (Figs. 3a \u20133d), since at most one end joint of the links is immobile. The first problem arises in step \u2013 \u2018e\u2019 where link 5 has to be added while joint C is restricted to be collinear with link 5, thus it cannot rotate and cannot move since joint D is immobile. The other end joint of link 5 is a pinned joint and thus is of course immobile. Therefore, according to the assembly self-stress rule, a new self-stress has been created" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0002570_icuas.2017.7991425-Figure5-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0002570_icuas.2017.7991425-Figure5-1.png", "caption": "Fig. 5: Scenario S-I. The introduction of two rendez-vous Points placed in a proper position allows the recomposition of the flight formation when possible. Xk is the domain to which RVWs connecting the k-th CFSO pair belongs", "texts": [ " For this reason, the higher level optimization problem is solved by adopting a general purpose Genetic Algorithm (GA), whose objective function involves the finding of optimal trajectories at the lower level. Parameters used in the optimization procedure are shown in Table I. As for the choice of the domain Xk where to search RVWs between pairs of CFSO n and m, the following possibly nonconvex region is chosen: CoH (VCFSOm \u222a VCFSOn ) + \u2212 (CoH (VCFSOm ) \u222a CoH (VCFSOn )) where VCFSOk is the set of vertices in the k-th CFSO, and CoH(V ) is the convex hull of vertices in V (see Fig. 5 and Fig.7). In order to show the effectiveness of proposed strategy, two different simulation scenarios have been considered. In both simulations, the fleet is composed of five UAVs. Table I, shows the parameters characterizing the described simulations. In the first simulation scenario S-I, twelve obstacles (see Fig. 4), have been considered. By virtue of (3), we can distinguish three CFSOs, two of these are touched by UAV trajectories (gray filled obstacles). Therefore, there is the need to optimize two RVWs to rejoin the flight formation as shown in Fig. 5. For simulation scenario S-II, we have considered twentyseven obstacles, that compose seven CFSOs. Fig. 6 shows flight trajectories optimized by any UAV independently. Only three of the seven CFSOs are activated by flight paths optimized independently by UAVs, and therefore four RVWs must be optimized, to reestablish the flight formation in between relevant CFSOs. The introduction of intermediate RVWs to re-compose the flight formation allows to reduce the spread of UAVs. Fig. 7 shows optimized RVWs and recomputed flight path for each UAV" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0001613_mec.2011.6025789-Figure2-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0001613_mec.2011.6025789-Figure2-1.png", "caption": "Figure 2. Heave compensation system", "texts": [], "surrounding_texts": [ "According to the Newton's laws of motion, the vertical direction motion equation can be obtained. The wire rope length under deep water is much longer than which out of water, to calculate the mass of wire rope conveniently, all the wire rope is supposed under water and converted to the payload mass, also suppose vessel static, then analysis payload and force on wire rope, the equation (1) can be obtained. 1( ) ( ) 2 EAMg Vg ml Al g l l \u03c1 \u03c1\u2212 + \u2212 = \u0394 1 \u03c1 and g represent density of water 1024 kg/m3 and constant of gravity 9.81 m/s2, respectively. Due to vessel heave motion, A linear mass-spring-damper system consisted of payload and wire rope will be oscillated, see its kinetics equation (2) 1( ) ( ) 2 m m eq p v p v p d p p p EAm z z R z C z R z C A z z l i i \u03b8 \u03b8 \u03c1= \u2212 \u2212 + \u2212 \u2212 \u2212 (2) The equivalent mass eqm may be written as 33 1 3eqm M A ml= + + (3) The term 33A denotes added mass, which represents a part of ambient water followed payload motion [7]." ] }, { "image_filename": "designv11_62_0003009_cobep.2017.8257273-Figure9-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0003009_cobep.2017.8257273-Figure9-1.png", "caption": "Fig. 9. : Mounting of the CET-System", "texts": [ " Figure 8 shows the circuit board of the secondary side power electronics with the FBR and the resonance capacitorC2p. The circuit board is mounted directly onto the rotor of the CETsystem with the secondary side coil L2 and connected to the rotor of the EESM. The rotating CET-system was dimensioned with analytical calculation [5] and the magnetic path was simulated in a 2D and 3D FEM-simulation [10]. To verify the calculation and the simulation a prototype was built and mounted on a test bench. Figure 9 shows the mounting of the different parts of the rotating contactless energy system to the test bench. Beginning with the backside of the EESM without the slip rings the bearing cover is opened. The first part of the stator of the CET-system is mounted onto the drilled holes of the bearing cover. Subsequently the rotational part is mounted onto the rotor of the EESM instead of the slip rings and the brackets for the middle ferrites are inserted and mounted to the first part of the stator of the CET-system" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0002356_jmech.2017.38-Figure1-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0002356_jmech.2017.38-Figure1-1.png", "caption": "Fig. 1 Sketch of cam-follower oblique-impact model.", "texts": [ " The two basic functions of the LCP are Upr function and Sgn function: {0} 0 {1} 0 Upr ( ,0] 0 and Sgn [ 1,1] 0 0 { 1} 0 x x x x x x x (1) In mechanics, Upr function is used to simulate geometry and dynamics of unilateral constraint, and Sgn function is used to simulate various kinds of dry friction. In the following part, LCP method will be used to solve the equations of cam-follower oblique-impact system. 3. OBLIQUE-IMPACT MODEL A cam-follower with oblique-impact on the contact point is studied in Fig. 1. The summary of the general notations used in Fig. 1 is shown in nomenclature. By extending the transient impact hypothesis and considering the tangential slip, the oblique collision model is established. In present work, the normal contact between rigid bodies is characterized by a set-valued force law called Signorini\u2019s condition [17]. Figure 2 shows two convex rigid bodies apart from each other by a relative normal gap or distance denoted by gN. The relative normal gap is nonnegative due to bodies\u2019 impenetrability condition, being the two bodies in contact with each other when gN = 0", " These two conditions can be summarized by a set-valued force law as [17] Sgn( )T N T (5) Tangential Coulomb\u2019s contact states can be expressed by Sgn function. It also represents a complementarity behavior. https:/www.cambridge.org/core/terms. https://doi.org/10.1017/jmech.2017.38 Downloaded from https:/www.cambridge.org/core. Columbia University Libraries, on 30 May 2017 at 09:45:34, subject to the Cambridge Core terms of use, available at Journal of Mechanics 3 4. DYNAMIC EQUATIONS OF THE FOLLOWER For the model shown in Fig. 1, the dynamics of the follower under different speed conditions are considered. There are three states of the follower and cam, separation, contact and impact. The dynamic equations of the follower can be expressed as [18] 0Mu h separation (6) - - 0N N T TMu h w w contact (7) -( - ) - - 0N N T TM u u w w impact (8) where u is a generalized velocity and here is the follower\u2019s angular velocity f ; M = Jo = ml2/3 + mr2/4 is the follower\u2019s moment of inertia; h = mgl cos(\u03b8f)/2 is the follower\u2019s external force; w is the generalized coefficient in contact point; \u039b is the impact force of the follower, u+ and u- is the velocity before and after the impact point, respectively. Notice that, \u03bbN is a continuous boundary parameter when there is no impact. \u03bbT is a non-continuous parameter when there is a transition from viscous to slip or change in the direction of relative speed. N denotes the normal direction and T denotes the tangential direction. The relationship between wN, wT and contact point\u2019s relative speeds are: ,N T N N T T w w w u w u t t (9) In order to get the unknown parameters wN and wT, we need to calculate the distance function in Fig. 1 0[ cos( )]sin( ) sin( )cos( )N c f c fg x e e r R and 0[ cos( )]sin( ) sin( ) cos( )T c f c fg x e e . According to the results of reference [19]: w g wq t (10) Therefore, 0( cos( )) cos( ) sin( )sin( )N c f c fw x e e and 0( cos( ))sin( ) sin( )cos( )T c f c fw x e e . There are three equations for the cam-follower impact system. In order to quickly and accurately analyze the complex behaviors of cam-follower oblique-impact system, it is better to establish one unified mathematical equation which can describe separation, impact and contact motion" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0001222_sii.2010.5708342-Figure1-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0001222_sii.2010.5708342-Figure1-1.png", "caption": "Fig. 1 A sample of scanned data", "texts": [ " An experiment is conducted to validate the proposed method, as described in Section 4. Finally, the conclusion is given in Section 5. The objective is to grasp an unknown object in a minimal amount of time by using partial shape information of an unknown object. The following assumptions are made: Robot: \u30fb The mobile robot has two driving wheels and a caster. \u30fb It is equipped with a 2-degrees-of-freedom gripper. Laser range finder: \u30fb A laser range finder is installed on the top of the robot at an inclined angle (Fig. 1) in such a way that the robot can detect and measure an object on the ground. \u30fb The parameters of the laser range finder are detection angle \u03b1, and detection range D. Object: \u30fb The object is irregularly shaped and cannot be deformed. \u30fb The object can be grasped by the mobile robot. \u30fb The grasping part of the object is lower than the installation height of the laser finder. \u30fb The object can be lifted by a single mobile robot with a gripper. \u30fb There are no other obstacles in the environment. For a robot to grasp the object through 3D model construction or a learning mechanism takes a great deal of time", " Step 4: Path planning The robot grasps the object using the information calculated in Step 3 or moves to the next scan position to detect a possible grasping point. By extracting the feature and calculating the grasping position from the scanned data directly, the robot can identify the part of the object that can be grasped quickly. The problem is which feature should be extracted from the scanned data and how the correct grasping position can be calculated utilizing the available partial information. The laser range finder installed on the top of the robot at an inclined angle acquires data for each scanning cycle, as shown in Fig. 1. The object is scanned N times, and the feature is extracted from these scanned data in order to calculate the grasping position. - 300 - SI International 2010 In this study, a mobile robot with a gripper is used to grasp the object. The object can be grasped and lifted with the gripper if the following two conditions are satisfied: \u25cf There are two parallel flat surfaces on the object (Fig. 2). \u25cf The distance between the two flat surfaces is not larger than the maximum opening width of the gripper" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0002027_978-3-319-00479-2-Figure15-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0002027_978-3-319-00479-2-Figure15-1.png", "caption": "Fig. 15 FE-model of camel (left side) and FE-model of the moose (right side)", "texts": [ " To recognize the vulnerability and to identify the mechanisms finite element simulations were performed. Therefore for this analysis a vehicle was used, where a suitable model and FE mesh (Fig. 14) were available. Due to the lack of information regarding statistical data a simulation matrix was created, which was derived from the PC-Crash simulation for the assessment of the accident type, with velocity of vehicle and animal and impact angle of the animal as variation parameter. For the animal the FE models of a camel and of a moose were used. The camel model has a mass of 450 kg (Fig. 15), left side), the moose model has a mass of 350 kg (Fig. 15), right side). A simulation matrix was established with a set of 48 different simulation configurations, with different configurations: \u2022 Fully/partially covered \u2022 Different obstacle (camel, moose) \u2022 Angle of the obstacle (0\u00b0, 45\u00b0, 90\u00b0) \u2022 Velocity of the obstacle (0 kph, 10 kph) \u2022 Velocity of the car (50 kph, 100 kph). In Fig. 16 one of the impact configurations can be seen. There were differences of the mass of the animals, the moose model had a mass of 350 kg, the mass of the camel was 450 kg, and also geometric differences, but the simulation results were similar, there were slight deformations at the front bumper but relatively 73Safety Measures for Avoiding or Mitigating the Occupant heavy deformations at the A-pillar and the roof" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0002971_s00773-017-0487-1-Figure3-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0002971_s00773-017-0487-1-Figure3-1.png", "caption": "Fig. 3 VPMM instruments in NSRC", "texts": [ " In this paper, the authors carry out the PMM (planer motion mechanism) test and free running model test in the large towing tank of the NSRC (Naval Systems Research Center). The NSRC\u2019s large towing tank has length of 248\u00a0m, breadth of 12.5\u00a0m and depth of 7\u00a0m, so it has enough width and depth for the execution of those tests. Our tank test facility allows us to conduct not only PMM tests for the surface ships but also that for under\u2011 water model. The PMM tests in later case are sometimes called VPMM (Vertical PMM), and the authors show the schematic view of test setting and PMM facility in Figs.\u00a03 and 4, respectively. As shown in Fig.\u00a03, there exists two 3\u2011component load cells (3\u2011component balance), and every component of forces and moments except surge force can be measured. Only the surge force is measured using sim\u2011 ple cantilever system inside of the model. In this paper, the moment is measured around the center of buoyancy. Table 2 Principal particulars of the used model Items Value Model length: L ( m) 2.500 Diameter: D ( m) 0.3400 Volume: \u2207 ( m3) 0.1798 Wetted surface area: S ( m2) 2.271 Fig. 2 Photographs of the model and thrust vectoring system 1 3 In this chapter, the authors show the test results and their analyses" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0001433_pi-a.1962.0130-Figure10-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0001433_pi-a.1962.0130-Figure10-1.png", "caption": "Fig. 10.\u2014General arrangement of 30kVA air-cooled brushless generator.", "texts": [ " The main exciter has ample capacity to sustain a steady short-circuit current from the generator of 3 p.u. The ability of the rotating rectifiers connected in the field circuit to withstand the transient currents and voltages induced by sudden short-circuits has been proved by numerous applications of sudden shortcircuits of the three principal kinds, i.e. symmetrical 3-phase, line to line and single line to neutral, with no rectifier failure or damage. The general arrangement of another generator of similar construction is shown in Fig. 10. This is a 6-pole generator rated at 30kVA O-75p.f. 3-phase 200 V 400 c/s. The layout of the principal parts and the excitation system are the same as for the 50kVA generator described above, except in the following respects: (a) The bearings are pressure lubricated by oil. (b) Solid steel is used for the revolving field system, in which the poles are tapered to give an increased cross-sectional area towards their roots and the pole-shoes are bolted on. This is a 6-pole generator weighing 951b. Cooling air is provided from an engine-driven blower and both bearings are lubricated by oil under pressure from the engine" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0003623_1.4044296-Figure2-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0003623_1.4044296-Figure2-1.png", "caption": "Fig. 2 Images of compressor insert from the design and assembly process: (a) CAD model, (b) gages attached, (c) FEA strain plot, and (d ) insert installed in backplate", "texts": [ " In this ball bearing TC design, there are both axial and radial clearances around the outer race to allow for the proper functioning of the squeeze film damper. This allows the outer race to make contact with either the turbine side or the backplate on the compressor side of the housing (indicated in Fig. 1(a)) depending on the direction of the axial load. Thus, the measurement of axial load required a sensor on each side of the bearing. Compressor Side Sensor. In order to measure the thrust loads on the compressor side, a pocket was machined into the backplate to accept an instrumented beam insert. Figure 2(a) illustrates a CAD model of the modified backplate and insert. The insert was composed of three triangular cantilever beams oriented at 120 deg angles from each other. Figure 2(b) depicts how each beam on the insert was instrumented with a strain gage to measure load. The strain gages were prewired polyimide carrier gages with a nominal 120 \u03a9 resistance and a 0.6 mm constantan grid. The gages were bonded to the beams using a cyanoacrylate strain gage glue. The strain gage lead wires were routed through the oil drain of the TC to the data acquisition system. Each beam was designed with a triangular cross section to maintain a constant strain across the surface of the sensor which reduced variability between beams due to differences in gage orientation. This was confirmed using finite element analysis (FEA) as shown in Fig. 2(c). Figure 2(d ) depicts the completed sensor with the strain gage lead wires strain relieved and routed through the oil drain. Prior to assembly of the TC, the instrumented beams were calibrated. First, the heights of the beams were verified to be level with the surface of the backplate. Small differences were corrected with shim stock. The rod holding the load was carefully located along the axis of the TC, then up to 30 kg of load (100 N total per beam) was applied gradually to develop the calibration curve illustrated in Fig" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0001036_s10015-013-0132-y-Figure1-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0001036_s10015-013-0132-y-Figure1-1.png", "caption": "Fig. 1 Principle of walking motion with 3 linked robot", "texts": [ " The This work was presented in part at the 18th International Symposium on Artificial Life and Robotics, Daejeon, Korea, January 30\u2013February 1, 2013. J. H. Lee (&) S. Okamoto Department of Mechanical Engineering, Graduate School of Science and Engineering, Ehime University, 3 Bunkyocho, Matsuyama 790-8577, Japan e-mail: jhlee@ehime-u.ac.jp S. Okamoto e-mail: okamoto.shingo.mh@ehime-u.ac.jp H. Koike Mitsubishi Electric Engineering Company Ltd., Himeji, Japan e-mail: x840013b@mails.cc.ehime-u.ac.jp K. Tani GS Yuasa International Ltd., Kyoto, Japan e-mail: keiya.tani@jp.gs-yuasa.com principle of walking motion of the robot is depicted in Fig. 1, which consists of three links; the torso, the swing and the stance legs. In an initial state, the robot rotates its swing leg in forward direction as in \u2018 ,\u2019 while the stance leg also rotates on the foot end at the same time. After moving forward of the swing leg as in \u2018`\u2019, both legs alternate their modes between \u2018swing\u2019 and \u2018stance\u2019 with each other as in \u2018\u00b4.\u2019 Then, the new swing leg that was the stance leg in the previous cycle starts its rotating motion. Alternating the swing and the stance modes between legs, the robot move forward resultantly" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0002311_ls.1385-Figure12-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0002311_ls.1385-Figure12-1.png", "caption": "FIGURE 12 Rotation error and the orbit of the rotor", "texts": [ "4 for simulation, respectively. It can be seen from the figure, with the increase of eccentric mass distance of the rotor, that driving trajectory of the rotor continues to expand. When the eccentric mass distance is 0.4, axis orbit of the rotor and bearing clearance circle come into contact, and the rotor contacts the bearing. This will seriously affect the stability of the rotor, and the situation is not allowed to happen. The maximum X and Y values in the centerline trajectory are listed in the table below. Figure 12 shows the relationship between the orbit of the shaft and the corresponding rotation error of the spindle; in ideal condition, point C(a,b) is the center of the circle with radius R, however, point O is the center of the actual path with the point Pi, and OPi is ri, |sin(\u03b8i\u2212\u03d5)|\u2264 1 then we can get the \u03b5i \u00bc ri\u2212R\u2212e cos \u03b8i\u2212\u03d5\u00f0 \u00de: (43) Equation 43 is \u03b5i=\u0394ri\u2212\u0394R\u2212 a cos \u03b8i\u2212 b sin \u03b8i, which \u0394R \u00bc 1 n \u2211 n i\u00bc1 \u0394ri; a \u00bc 2 n \u2211 n i\u00bc1 \u0394ri cos \u03b8i; b \u00bc 2 n \u2211 n i\u00bc1 \u0394ri sin \u03b8i: Figure 13 shows the rotation error with different eccentric mass distance" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0002302_b978-0-12-803581-8.10351-0-Figure22-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0002302_b978-0-12-803581-8.10351-0-Figure22-1.png", "caption": "Fig. 22 Front bulkhead design with donut bonded lips.", "texts": [ " Under the rules governing the competition that this monocoque was originally developed for, the front bulkhead had to be shown to be equivalent in structural performance to a welded steel tube frame in terms of bending buckling and tension. In addition to the rules-required structural equivalency, this monocoque design includes loading of the front bulkhead from twisting forces that originate from the lower front suspension inputs. The loadings from the suspension will tend to twist the front monocoque tube structure from its original rectangular shape into a more parallelogramlike shape. Three designs were considered. The first was with an open hole in the middle and a composite lip would extend backward as in Fig. 22. It was thought that the inner flange would sufficiently increase the stiffness for the suspension twisting loading, and a hole would provide access to the footwell portion of the driver compartment. The next design had a cutout, again for access, but a phenolic insert was incorporated around the rim of the cutout, as seen in Fig. 23, in hopes of further increasing the stiffness. The third design was a flat plate as shown in Fig. 24. All three bulkheads were analyzed with a similar layup of (0\u201390, 90, 90, 0, 0, 0\u201390, \u00bc00 honeycomb core, 0\u201390, 0, 0, 90, 90, 0\u201390)" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0000720_amm.307.304-Figure2-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0000720_amm.307.304-Figure2-1.png", "caption": "Figure 2 Replacing the layers of bevel gear face width with spur gears", "texts": [ "159, Pennsylvania State University, University Park, USA-28/05/15,11:41:40) The number of teeth of virtual spur gears and the equal diameter of them can be found from the relations [4]: \u03b3cos bevel spur z z = (1) \u03b3cos bevel spur d d = (2) where: Zspur Number of teeth in virtual spur gear Zbevel Number of teeth in bevel gear dbevel Pitch diameter of bevel gear dspur Pitch diameter of virtual spur gear \u03b3 Pitch cone angle In this study, for higher accuracy face width of each tooth of bevel gear replaced with multiple pairs of spur gears such as shown in Fig. 2. In other words, each layer along the face width is replaced with a pair of spur gear. The load distribution is not constant along the face width as mentioned in reference [2] and shown in Fig. 2. The load changes linearly so that the maximum load is applied on the heel and the minimum at toe. Moreover, whether there is one pair of gear or two pair of gear in contact will affect the load distribution for bevel gear. After replacing the bevel gear with multiple spur gears, the load acting on each layer of bevel gear will apply on its equivalent spur gear. In order to determine the Hertzian stress, it is important to determine the equal radii of curvatures of bevel gear at each point. Considering the equivalent geometry as shown in Fig" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0002495_978-3-319-56802-7_41-Figure1-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0002495_978-3-319-56802-7_41-Figure1-1.png", "caption": "Fig. 1 Kinematic model of a 7 DOF KUKA LWR", "texts": [ " Definition 2 (Singularity\u2014filtration criterion) A configuration q \u2208 V n of an SM is a forward kinematic singularity if and only if the length \u03ba(q) > \u03ba0. The condition \u03ba > \u03ba0 is equivalent to say that the filtration of D is not constant in a neighbourhoodofq. Since thefiltration reveals the effect of higher-order infinitesimal motions it allows for identification of the joint motions that lead the SM out of a singularity [7, 12]. Example 1 Consider the redundant 7 DOF (anthropomorphic) SM in the reference configuration q = 0 in Fig. 1. The SM has 7 revolute joints, and its c-space is V n = T 7. The joint screw coordinate vectors in the reference configuration w.r.t. to the global frame are Y1,3,5,7 = \u239b \u239c \u239c \u239c \u239c \u239c \u239c \u239d 0 0 1 0 0 0 \u239e \u239f \u239f \u239f \u239f \u239f \u239f \u23a0 ,Y2 = \u239b \u239c \u239c \u239c \u239c \u239c \u239c \u239d 0 \u22121 0 \u2212L3 \u2212 L5 \u2212 L E 0 0 \u239e \u239f \u239f \u239f \u239f \u239f \u239f \u23a0 ,Y4 = \u239b \u239c \u239c \u239c \u239c \u239c \u239c \u239d 0 1 0 L5 + L E 0 0 \u239e \u239f \u239f \u239f \u239f \u239f \u239f \u23a0 ,Y6 = \u239b \u239c \u239c \u239c \u239c \u239c \u239c \u239d 0 \u22121 0 \u2212L E 0 0 \u239e \u239f \u239f \u239f \u239f \u239f \u239f \u23a0 . The Jacobian J (0) = (Y1,Y2,Y3,Y4,Y5,Y6,Y7) has rank J (0) = dim D0 = 3. In regular configurations q it is rank J (q) = 6" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0003970_s11370-019-00294-7-Figure7-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0003970_s11370-019-00294-7-Figure7-1.png", "caption": "Fig. 7 Design comparison of parallel RRR derived from a serial manipulator", "texts": [ " That is because the freedom to locate (S\u03021)i is diverse as shown in Fig.\u00a06, while the line denoted as s\u0302T 23 is fixed along the link. It is also remarked that the static relation, the dual relation of (34), is denoted as (34)?\u0307? = \u23a1\u23a2\u23a2\u23a2\u23a3 (\ud835\udf141)1 (\ud835\udf141)2 (\ud835\udf141)3 \u23a4\u23a5\u23a5\u23a5\u23a6 = \u23a1 \u23a2\u23a2\u23a2\u23a2\u23a2\u23a2\u23a2\u23a2\u23a2\u23a3 s\u0302T 23 s\u0302T 23 (S\u03021)1 s\u0302T 31 s\u0302T 31 (S\u03021)2 s\u0302T 12 s\u0302T 12 (S\u03021)3 \u23a4 \u23a5\u23a5\u23a5\u23a5\u23a5\u23a5\u23a5\u23a5\u23a5\u23a6 T\u0302 . 1 3 Three columns of (35) represent actuation wrench, which are actually common in both serial and parallel manipulators. In serial manipulator, they are directed along each link as shown in Fig.\u00a07. On the other hand, they are directed along each second link in the parallel manipulator. However, their directions are the same. In order to make comparison between serial manipulator and its derived parallel manipulator, the following conditions should be met in (28) and (34). What to declare in this section is summarized as follows: (i) Geometry of parallel manipulator is expressed by ray coordinate (ii) Geometry of its inversion is found as the geometry of serial manipulator expressed in ray coordinate A common 3-RPR parallel manipulator has been shown (with active prismatic joints only) in Fig", " The kinematic model of parallel manipulator is interconnected to that of serial manipulator and vice versa. Contrary to general statement saying that serial has better isotropy and parallel has better force and power transmission characteristics in general, the so-called inter-kinematic relation allows us to compare serial and parallel manipulators in a more rigorous way. This is the main contribution of this paper. The main principle lies on common actuation wrench applied to the output space for both serial and parallel manipulators as explained in Fig.\u00a07. Without loss of generality, the serial manipulator derived from parallel has better isotropy, while the parallel manipulator derived from serial can be designed to have better force and power transmission. In case of force transmission ratio, the parallel manipulator can be optimized to outperform serial manipulator. Acknowledgements This work was supported by the Technology Innovation Program (or Industrial Strategic Technology Development Program-Artificial intelligence bio-robot medical convergence project) (20001257, Artificial intelligence algorithm-based vascular intervention robot system for reducing radiation exposure and achieving 0" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0002693_chicc.2017.8028408-Figure1-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0002693_chicc.2017.8028408-Figure1-1.png", "caption": "Fig. 1: A geometrical illustration of LOS guidance", "texts": [], "surrounding_texts": [ "Definition 1 [25]. The system x\u0307 = f(t, x, u), (1) where f : R \u00d7 Rn \u00d7 Rm \u2192 Rn is continuously differentiable, is said to be input-to-state stable (ISS), if there exist class KL function \u03c3 and class K function \u03ba\u0304, such that for any bounded input u and any initial condition x(0), it holds that \u2016x(t)\u2016 \u2264 \u03c3(\u2016x(0)\u2016, t) + \u03ba\u0304(\u2016u\u2016), (2) for all t \u2265 0. Lemma 3 [25, 26]. Suppose that for the system (1), there exists a smooth function V : Rn \u2192 R+ such that for all x \u2208 Rn and u \u2208 Rm , \u03ba1(\u2016x\u2016) \u2264 V (t, x) \u2264 \u03ba2(\u2016x\u2016), (3) and \u2202V \u2202t + \u2202V \u2202x f(t, x, u) \u2264 \u2212W (\u2016x\u2016), \u2200\u2016x\u2016 \u2265 (\u2016u\u2016), (4) where \u03ba1, \u03ba2 are class K\u221e functions, and W , are class K functions. Then, the system (1) is ISS with \u03ba\u0304 = \u03ba\u22121 1 \u25e6\u03ba2 \u25e6 . Lemma 4 (Cascade Stability [27]). Consider the cascade system x\u03071 = f1(t, x1, x2), (5) x\u03072 = f2(t, x2), (6) if the system (5), with x2 as input, is ISS and the origin of (6) is GUAS, then the origin of the cascade system (5) and (6) is GUAS." ] }, { "image_filename": "designv11_62_0000969_012001-Figure6-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0000969_012001-Figure6-1.png", "caption": "Figure 6. Wheel-rail contact pair.", "texts": [], "surrounding_texts": [ "Once contact ellipse is obtained, tangential problem is solved using FastSim. FastSim is a compromise between efficiency and an adequate solution accuracy. Fundamental assumptions of this contact model are: \u2022 normal contact is independent of the tangential, \u2022 Hertz contact theory and their assumptions are valid, thus it can be applied, \u2022 friction coefficient is constant throughout contact ellipse and independent of local slip, \u2022 contact between a wheel and a rail occurs in a single area, which is not always true. The main feature of FastSim is that it relates, by the Simplified Theory of Elasticity, displacements and shear stresses in a linear way. Developing this relationship the following set of differential equations is obtained: \u2202\u03c4xc(xc.yc) \u2202xc = 3 c11 G 8 a \u03be \u2212 4 \u221a b c23 G \u03c0 a3/2 \u03c6 yc (2a) \u2202\u03c4yc(xc.yc) \u2202xc = 3 c22 G 8 a \u03b7 + 4 \u221a b c23 G \u03c0 a3/2 \u03c6 xc (2b) (2c) where xc and yc are the local axes of the contact ellipse, \u03c4ic(xc, yc) shear stress in ic local contact direction in point (xc, yc) of the ellipse; \u03be, \u03b7 and \u03c6 are creepages; c11, c22 and c23 are Kalker coeficients that depend on a/b ratio and Poisson\u2019s ratio \u03bd [28]; and G is the shear stress modulus. Creepages can be defined as: \u03be = Vxc,B1 \u2212 Vxc,B2 V , (3a) \u03b7 = Vyc,B1 \u2212 Vyc,B2 V , (3b) \u03c6 = \u03b8zc,B1 \u2212 \u03b8zc,B2 V . (3c) (3d) taking into account that Vic,Bj is the solid rigid velocity of body j in contact direction ic in the contact point (ellipse center); \u03b8zc,Bj is the solid rigid rotation around contact axis zc of body j in the contact point; and V is the impose translation velocity of train. For solving (2) in contact area it is necessary to make a discretization of that area, like the one shown in Figure 9(a). Integration over this ellipse is made over stripes with constant contact coordinate yc = cnt. The stress value in ellipse boundary of the stripe (xc > 0) is known (\u03c4xc = 0 and \u03c4yc = 0). With this value a new stress value can be calculated using an Euler Explicit integration schema (figure 9(b)). Each step, shear stress in a new point is calculated. This stress \u03c4 (xc, yc) must fulfill the Coulomb\u2019s Law. Taking into account that normal stresses have an ellipsoidal distribution, stress in each point must fulfill: |\u03c4 (xc, yc)| \u2264 \u03bc 3 N 2 \u03c0 a \u221a 1\u2212 (xc a )2 \u2212 (yc b )2 (4) where \u03bc is the friction coefficient. If this condition is fulfilled, contact at this point is stick, else, it is slip. If contact is slip, the absolute value of shear stress will be normal stress in this point times friction coeficient. The direction and sense of the stress vector will be those obtained before applying the Coulomb\u2019s Law. Thus, applying this algorithm, shear stress distributions could be like those shown in Figure 10. When the shear stress distribution is known, resultants forces in local contact directions can be obtained. These forces that will be introduce in the vehicle and structure models." ] }, { "image_filename": "designv11_62_0000715_amm.130-134.610-Figure2-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0000715_amm.130-134.610-Figure2-1.png", "caption": "Fig. 2 Engaging position changed of gears of external gear pumps in trapped-oil cycle", "texts": [ " Formulas (3) can be transformed as the following simultaneous differential equation expressed with X: X''+Cb1YK X'+ Cb1K X= Cb2 M2 (4) In which, X '' is the vibration acceleration, and X''= rb(\u03b81''\uff0d\u03b82'') ; Cb1= rb 2 / I1+ rb 2 / I2 ; Cb2= rb/ I1/\u03b72+ rb/ I2/\u03b72 (5) The solution of Equation (4) mainly depends on the calculations of K and M2, detailed calculation method of M2 is described in Reference [9] and it will be used directly here. Hereafter is the simple description only when the K calculation is given. Figure 2(a)~(h) describe the whole process of tooth profile of gear O2 from entering into engagement in Figure 2(a) to withdrawing from engagement in Figure 2(h). For specific implication of each figure, see Reference [8]. In these figures, Figures 2(a)~(b) and Figures (g)~(h) are pair tooth engagement zone and Figures (b)~(g) are single tooth meshing zone. In the figures, the marked zones are the situations of trapped oil zones 1 and 2 in different engagement areas, and the trapped oil pressures are expressed with p1 and p2 respectively, \u201c\u25cb\u201d and \u201c\u25a1\u201d express the meshing positions N1, N2 and backlash positions H1, H2, FN1, FN2 and FH1, FH2 express the forces acted on gear O2 in N1, N2 and H1, H2, their acting positions and directions are shown as Figure 2. Assuming s as radius of curvature on gear O1 in engaging position N1 in Figure 2, then the s when the eight corresponding special points in Figures2 (a) ~(h) enter into meshing respectively is assumed as sa, sb, sc, sd, se, sf, sg and sh, they are known functions [8] of tooth shape parameters, and s ab expresses the interval [sa,sb], the definitions of other intervals are similar to this definition. According to comparison of oil-film extrusion stiffness, contact stiffness and strapped oil stiffness [7-8] , the oil-film extrusion stiffness can be neglected. Therefore, by combining the calculations of contact stiffness and trapped oil stiffness, defined with the linear stiffness, the following equations can be obtained: KN1+ KN2\uff0dKH1+ KT1\uff0dKT2 (s\u2208s ab ) KN1\uff0dKH1+ KT1 (s\u2208s bd ) K= KN1\uff0dKH1\uff0dKH2+ KT1\uff0dKT2 (s\u2208s de ) (6) KN1\uff0dKH2\uff0dKT2 (s\u2208s eg ) KN1+ KN2\uff0dKH2+ KT1\uff0dKT2 (s\u2208s gh ) In which, KN1 and KN2 are time varying contact stiffness in N1, N2; KH1 and KH2 are time varying contact stiffness in H1, H2; KT1 and KT2 are time varying trapped oil stiffness of trapped oil 1, 2" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0003356_978-981-13-5799-2_12-Figure12.71-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0003356_978-981-13-5799-2_12-Figure12.71-1.png", "caption": "Fig. 12.71 Pattern of a studless tire [103]", "texts": [ " After the use of studded tires was prohibited in Japan in 1990, winter tires without studs but with many sipes, called studless tires, were developed in Japan. The sedimentation of dust decreased but the winter-specific accidents increased as studless Temperature (\u00b0C) D ep th o f q ua si -li qu id la ye r ( \u212b) 20 40 60 0 0 \u22121 \u22123 \u22124 \u22125 \u22126 \u22127 \u22122 Fig. 12.67 Depth of the quasi-liquid layer versus temperature (reproduced from Ref. [94] with the permission of Hokkaido University Press) tires were increasingly adopted as shown in Fig. 12.70. An example of a studless tire is shown in Fig. 12.71. Because studded tires have been prohibited in many countries, studies on friction on ice have been conducted in northeast Asia and Europe. Japanese research has concentrated on increasing the friction coefficient on ice rather than the tire performance on snow because it is difficult to drive a vehicle on ice. Additionally, the interaction between ice and the tire surface can be observed through ice using an indoor apparatus. Such observation is not possible for the interaction of a tire and snow" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0002088_6.2017-1068-Figure1-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0002088_6.2017-1068-Figure1-1.png", "caption": "Figure 1: Schematic view of the quadrotor configuration, left a top view and right an intersection through the y-z plane seen from behind.", "texts": [ " Then, in Section V the simulation setup is described and in Section VI the results are presented and discussed. Finally, Section VII gives the conclusion and Section VIII some recommendations for future work. A quadrotor has four engines symmetrically placed around its center of gravity, the location where the main flight control board is situated. The basic working principle is that all control is done by changing the thrust of two sets of counter-rotating propellers. A schematic view of a quadrotor system with body reference frame B and the forces and moments can be found in Figure 1. As can be seen the system is 2 of 15 American Institute of Aeronautics and Astronautics D ow nl oa de d by U N IV E R SI T Y O F C O L O R A D O o n Ja nu ar y 13 , 2 01 7 | h ttp :// ar c. ai aa .o rg | D O I: 1 0. 25 14 /6 .2 01 7- 10 68 under-actuated with input moments around each axis, but input forces only in the negative z-direction. Another reference frame E is defined as North-East-Down Earth-Fixed with the xE and yE-axis pointing North and East and the zE-axis pointing down. The rotation of the body frame B in the Earth-Fixed frame E is defined using the unit quaternion q = [q0 q1 q2 q3]T " ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0000711_detc2011-48564-Figure2-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0000711_detc2011-48564-Figure2-1.png", "caption": "Figure 2. a) HOUSING STRUCTURE CONSIDERED IN THIS STUDY; b) EXCITATION AND RESPONSE POINTS.", "texts": [ " However, the resultant size of the mass and stiffness matrices may be prohibitively large. In fact, in most cases, after these mass and stiffness matrices are coupled with the nonlinear time-varying hypoid geared rotor assembly, the computational requirement is too large to be practical. Alternatively, the modal model of the housing extracted from an appropriate set of frequency response functions (FRFs) is coupled with the internal geared rotor assembly. The housing structure considered in this study is shown in Fig 2(a). There are 14 response points shown in Fig 2(b) on the housing surface and bearing locations; each possesses 3 orthogonal translation coordinates. Also, 3 of these 14 points are excited externally. The first order formulation, as described in Eq. (13-15), of the Unified Matrix Polynomial Approach [13-14] is then applied to extract the modal parameters from the FRFs )]([ \u03c9H as shown in Fig 3. From the equation, ]][[)](][[)](][[)( 001 1 IBHAHAj =+ \u03c9\u03c9\u03c9 . (13) either ][ 1A or ][ oA is assumed to be the identity matrix. This results in two equations whose solutions bound the \u201ctrue\u201d solution given by 3 Copyright \u00a9 2011 by ASME Downloaded From: http://proceedings", " Coupling with the first 10 modes of housing is believed to be representative when the mesh frequency is less than 500 Hz, since the natural frequency of the tenth mode is around 1000 Hz, as shown in Fig 4. 8 Copyright \u00a9 2011 by ASME Downloaded From: http://proceedings.asmedigitalcollection.asme.org/ on 02/04/2016 Terms of Use: http://www.asme.org/about-asme/terms-of-use 9 Copyright \u00a9 ASME 2011 Figures 11 and 12 show the comparison of dynamic mesh force and housing surface acceleration, respectively, using 5- mode, 10-modes and no coupling cases. No housing surface acceleration data is available for the \u201cno coupling\u201d case, and the Y-direction acceleration of Point 1, defined previously in Fig 2(b), is plotted in Fig 12. The results indicate that the 5- modes coupling assumption is accurate enough for this case to predict dynamic mesh force and housing acceleration when the mesh frequency is less than 500 Hz. Based on the curve shown, the dynamic mesh force appears to be sensitive to the housing elastic characteristics. Comparisons of housing surface accelerations at other locations and directions also support this conclusion. For simplification purposes, these results are not included as an additional proof, and the data will not be shown in this study. External excitation on the housing surface then is applied simultaneously with the internal TE excitation to study the effect of external excitation on the dynamic mesh force and housing surface acceleration. Two cases with the external excitation exerted on different locations, Points 1 and 2 as shown in Fig 2(b), on the housing are studied. The external excitation has a magnitude of 2000 N and a frequency of 360 Hz. As indicated by Fig 13, the dynamic mesh force is more sensitive to the external excitation exerted on Point 2. This is similar to the linear cases, where the sensitivity of response to excitation is determined by the mode shape. A FFT of dynamic mesh force for a single mesh frequency with a value of 300 Hz for the two cases is plotted in Figs 14 and 15. As shown, the magnitudes of dynamic mesh force at 360 Hz for the two cases are different, while the magnitudes of the dynamic mesh forces at mesh frequency and harmonics show very little difference" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0003741_012077-Figure4-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0003741_012077-Figure4-1.png", "caption": "Figure 4. Road barrier in: (a) compact configuration; (b) partially deployed configuration; (c) totally deployed configuration, [9].", "texts": [ "1088/1757-899X/591/1/012077 If we consider 1 the rotational angle of the driver link, the area A1 of a wall formed by this mechanism in the deployed configuration will be equation (3): 1 2 1 sin2 AB lA (3) considering that ADAB ll . Another view of the mechanism with overlapped compact, partially deployed and totally deployed configurations is presented in figure 2. Two examples of deployable structures that may use this mechanism will be given here. The first example is a mobile house (see figure 3). Another application of the mechanism described here road barrier, represented in folded, partially deployed and totally deployed configurations (figure 4). Modern Technologies in Industrial Engineering VII, (ModTech2019) IOP Conf. Series: Materials Science and Engineering 591 (2019) 012077 IOP Publishing doi:10.1088/1757-899X/591/1/012077 The first simulations have been done, considering the coordinates of the mechanism nodes in the fully extended configuration as illustrated in figure 5. We have to note that no dimensional synthesis of this mechanism has been realized till now. As consequence, the mechanism could not fold in the optimum compact configuration" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0003952_s12206-019-0936-3-Figure6-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0003952_s12206-019-0936-3-Figure6-1.png", "caption": "Fig. 6. Structure of the buffering leg.", "texts": [ " This condition is helpful for the robot to take off rapidly. The bionic design principle is used for the buffering leg. The buffering leg includes the femur link and tibia link, and the length of each link is proportional to the leg of locust. Each part is connected by hip, knee, and ankle joint, and each joint can be decreased to an equivalent rotation pair. Two springs are used to limit movements of the hip and knee joints for storing energy to achieve buffering effect, and the connection points are shown in Fig. 6. Table 1. Constraint conditions of the optimization parameters. l1-l5 (mm) \u03b8 (\u00ba) \u03b21 (\u00ba) a (mm) b (mm) [0, 20] [\u22125, 5] [75, 125] [\u221215, 15] [-5, 5] Table 2. Mechanism parameters of the four-bar jumping leg. The spring is used to achieve energy storage and rapid release. One end of the spring is connected to link A2A3 at point H1, and the other end is connected to link A1A4 at point H2 (Fig. 5). The contact point B1 between jumping leg and ground can be simplified as a rotating pair. In Fig. 5, the expression of the center of mass in coordinate system o2-x2y2z2 can be shown in Eq", " 13 shows the relationship between the position of the center of mass in coordinate system Ot-XtYt and the rotation angle of the robot during flight. When the center of mass of the robot moves along the positive direction of Xt axis and negative direction of Yt axis, the rotation angle of the trunk during flight decreases. The maximum rotation angle of the trunk is 547.3\u00b0, and the robot rotates one and a half circle in the air. By contrast, the minimum rotation angle of the trunk is 154.5\u00b0. Therefore, the position of the center of mass of the robot can be adjusted in accordance with Fig. 13 to obtain good motion stability during flight. Fig. 6 shows that the end of the buffering leg cannot slide on the ground during buffering when the ground is very rough. In accordance with Ref. [28], the kinematics of buffering leg can be obtained. 2 2 2 2 2 1 1 2 2 2 arccos arctan 2 b b b b l s h l h sl s h b + + - = + + (49) 2 2 2 2 1 2 2 1 2 arccos 2 b b b b b l l s h l l b p + - - = - (50) 3 2 1b b bb b b= - (51) where lb1 and lb2 are the lengths of tibia link (AB) and femur link (BC), respectively. h is the distance from point C to the ground, and s is the projection distance of the buffering leg on the ground" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0002755_icems.2017.8056505-Figure1-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0002755_icems.2017.8056505-Figure1-1.png", "caption": "Fig. 1 Magnetic flux distribution and demagnetization at revers magnetization field current supply", "texts": [ " Since the magnetic flux is relaxed in the rotor core, IPMSM is more resistant to demagnetization than a surface permanent magnet synchronous motor (SPMSM). If the magnet is deeply embedded in the rotor core in a V shape, the demagnetization durability is improved. One conceivable countermeasure for lower efficiency due to copper loss increase under high-temperature environments is adopting a concentrated winding motor with low winding resistance. However, in concentrated winding motors, the number of stator teeth per rotor pole is small, and local irreversible demagnetization of the permanent magnet tends to occur. Figure 1 shows the magnetic flux lines and the demagnetization ratio when the reverse magnetic field current is applied by finite element analysis (FEA). The distributed winding motor in Fig. 1(a) and the concentrated winding motor in Fig. 1(b) provide the same current with a number of turns at which the back-EMF is the same. The concentrated winding motor occurs in the strong local demagnetization at the corner of the permanent magnets (Fig. 1(b)). Some slight local demagnetization is seen from the corner portion of the magnet even from the center of the magnetic pole. As mentioned above, changing the magnetic field orientation of the magnet-embedded motor might improve the demagnetization characteristics under a high-temperature environment. Some studies have argued that the magnetic field orientation influences radial stress [1], the measurement of permanent magnet magnetization tilted in the magnetizing direction [2], and the relationship between the magnetizing direction and irreversible demagnetization [3]" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0000392_iros.2013.6696761-Figure5-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0000392_iros.2013.6696761-Figure5-1.png", "caption": "Fig. 5. A planar vehicle model with suspension", "texts": [ " The system is therefore indeterminate since it is driven by two control inputs (the front and rear traction forces), of which only one can be determined from the equations of motion. To resolve this inherent indeterminacy, we add a virtual suspension system that in effect adds two degrees-of-freedom to the system. Adding degrees-of-freedom to the system relaxes the coupling between the three equations of motion, which allows us to solve for the two traction forces required to move the system at the given speed and acceleration along the path. The vehicle with the suspension is shown in Figure 5. The suspension consists of a linear spring that is free to move along the y body axis. The rear and front springs have a non-compressed length d0, spring constant k, and compressed lengths d1 and d2, respectively. The force applied by a typical spring is thus: Pz = (d0\u2212d)k; d < d0. (8) Note that the force applied by the spring is state dependent since d is a function of the vehicle orientation \u03b8 . This and the assumption of mass-less wheels allows us to compute an expression that relates the normal force Fn to the traction force Ft " ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0001259_sav-2010-0604-Figure4-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0001259_sav-2010-0604-Figure4-1.png", "caption": "Fig. 4. Connector realized.", "texts": [ " Among them, the connected body can be a part, a surface, a cell or a node. There are three main steps as stated as follows. Step 1: Create connector In this step, the connector\u2019s location, connect when, connect what, layers and connect rule are defined, shown in Fig. 2. Step 2: Realize connector The process of realizing connector is shown in Fig. 3. Because the connection cell is a rigid cell, its material property is not needed. With a projection tolerance of 10, the obtained connectors and new connect cells are shown in Fig. 4. Step 3: Check the quality of the connector The state of a connector can be classified as unconsummated connector, succeed realized connector and failed connector. They can also be displayed in green, yellow and red colors in Hypermesh. The successfully finished connections of the car roof and its crossbeam are shown in Fig. 5. Step 4: Final meshing of BIW After the welding points are all setup, the whole model of the BIW based on FEM is achieved, shown in Fig. 6 and Table 3. With the above finite element model of the BIW, the modal analyses of it are carried out based on Nastran where Block Lanczos method is adopted" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0001259_sav-2010-0604-Figure7-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0001259_sav-2010-0604-Figure7-1.png", "caption": "Fig. 7. The 1st \u2013 7th order vibration shapes.", "texts": [ " The successfully finished connections of the car roof and its crossbeam are shown in Fig. 5. Step 4: Final meshing of BIW After the welding points are all setup, the whole model of the BIW based on FEM is achieved, shown in Fig. 6 and Table 3. With the above finite element model of the BIW, the modal analyses of it are carried out based on Nastran where Block Lanczos method is adopted. The calculated natural frequencies and vibration shapes are obtained, listed in Table 4. Some typical vibration shapes are shown in Fig. 7. The material properties and the boundary conditions of the BIW will change the modes effectively, especially the welding connections. Take the roof panel as an example to illustrate the influence, in which the roof panel is modeled with single or double side constraints, as shown in Fig. 8. For the first case of the model with single side constraints, the calculated modes from 1st to 4th order are listed in Table 5. The calculated modes from 1st to 4th order are listed in Table 6 for the roof panel modeled with double side constraints" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0001590_6.2013-1503-Figure8-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0001590_6.2013-1503-Figure8-1.png", "caption": "Figure 8. Detailed Design View of the Wing Structure", "texts": [ " Nevertheless, much can be gleaned from their efforts. Phase III proceeded with the detailed design of the wing alone with two different tip configurations (the 214 ft raked and 214 ft blended) for comparison. The detailed design consisted of seven distinct areas of examination detailed below: - A CAD model in NX containing detailed parts for the wing\u2019s spars, ribs, stringers and skin. This in turn facilitated higher fidelity FEA and CFD analysis along with a modicum of optimization. The current model can be seen in Figure 8. - A CFD model in Star CCM to determine critical flight loads and accurately model compressibility effects not captured in Phases I and II. - An FEA analysis of the CAD model with CFD generated loads for critical load cases. In addition, modal analysis from the FEA analysis permitted basic aeroelasticity analysis. - Aeroelasticity analysis using the NASA proprietary tool to ensure that the wing possesses adequate flutter margins. - Optimization of selected wing structures using Optistruct to eliminate as much unnecessary weight from the CAD model as possible via topology optimization" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0001553_appeec.2011.5749147-Figure1-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0001553_appeec.2011.5749147-Figure1-1.png", "caption": "Figure 1. Structure diagram of MFC 1", "texts": [ " The sediment was pumped through the two solutions at same flow-rate, and the electricity signal would be measured and appeared in the form of polarization and power density curve. II. EXPERIMENTAL APPARATUS AND PROCEDURES In this work, a double-chamber MFC and two solutions would be used for simplifying the experimental operation. The experiment for the two solutions will be carried out at the same time and condition. The difference for these two MFC would mainly indicated that cases of the flow channel with / without at the anode. Here, MFC 1 set will be with plate and shown in Figure 1, and contrast to MFC2. Here, the plate whose dimension is 5mm high and 1.4 mm thickness will be embedded and vertical to the slanted wall of convergent flow channel. 978-1-4244-6255-1/11/$26.00 \u00a92011 IEEE The anode and cathode reaction chambers are separated by embedding a proton exchange membrane (PEM) in the middle of the SMFC and connected by an external circuit. The MFC is made of PMMA whose thickness is 5mm, and each volume chamber is 4.515mL. The material of electrode for anode and cathode is carbon cloth which total area is 9" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0002200_978-3-319-51222-8-Figure3.3-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0002200_978-3-319-51222-8-Figure3.3-1.png", "caption": "Fig. 3.3 Floor plate, a system, b deflection under gravity load, and c influence function for the deflection w at a node x", "texts": [ "5 m is not the exact value w(x) = 2.75 m. At this point, the reader may object and say that the FE-program computes the nodal values by solving the system K w = f and any value in between by a linear interpolation. This is correct, but these values are just as largeas if the FE-programhad calculated these displacements with the approximate kernels Gh(y, x). This is the important observation. 144 3 Finite Elements And this holds true for all linear, self-adjoint differential equations. The deflection surface of the slab in Fig. 3.3 was computed (theoretically) as follows: The FE-program placed successively at each node xi a force P = 1, determined the corresponding deflection surface Gh( y, xi ), and evaluated the integral wh(xi ) = \u222b \u03a9 Gh( y, xi ) g( y) d\u03a9 y = Volume of Gh \u00d7 g , (3.16) 3.2 Why the Nodal Values Are Exact 145 where g [kN/m2] is the self-weight of the plate. Of course, the FE-program did it differently, and it solved the system K w = f , but the nodal values wi are just as large as if the FE-program had computed these values with the influence function (3" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0003420_978-3-030-16943-5_55-Figure1-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0003420_978-3-030-16943-5_55-Figure1-1.png", "caption": "Fig. 1. Schematic representation of the composites cross-sections.", "texts": [ " Next the vacuum contactor was mounted to the outer part of the vacuum foil and the whole system was connected to vacuum pump. The vacuum bag method was carried out with the constant pressure 0,8 bar during 1 h. The composites contained layers from basalt fibers, flax fibers and both flax and basalt fibers in different configuration. The samples were described as BBB; FFF; BFB; BF; FBF adequately to the type of the incorporated fibrous reinforcement. The arrangement of fibers in the composites is shown in Fig. 1. First type of sample described as BBB contain only basalt fiber fabric, while FFF sample contain only flax fiber fabric. Hybrid epoxy composites reinforced with both flax fiber and basalt fiber were de-signed as BFB and FBF were external laminate layers were made of 2 layers of basalt fabrics and 2 layers of flax fabrics, respectively. Sample BF contain equal amount of layers from flax and basalt. The dynamic-mechanical properties of the composites, were evaluated using DMTA methods (Anton Paar MCR 301, Austria) in a torsion mode, operating at frequency 1 Hz in the temperature range between 25 \u00b0C and 180 \u00b0C, and at the heating rate 2 \u00b0C/ min" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0001433_pi-a.1962.0130-Figure4-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0001433_pi-a.1962.0130-Figure4-1.png", "caption": "Fig. 4.\u2014Schematic of connection for cascade generator.", "texts": [ " This generator is usually coupled, together with a hydraulic pump, to a small high-speed turbine and it is convenient and economical to use the generator as a motor for checking and testing purposes on the ground. A polyphase induction machine is practically ideal for this dual role. (6.3) Cascade Induction Generator Fig. 3 shows a typical general arrangement of this type of machine. Essentially it comprises two ordinary induction machines arranged in a common housing with their rotor cores mounted on a common shaft. The rotor windings are connected together so that the m.m.f.s due to the rotor currents rotate, with respect to the shaft, either in the same direction or in opposite directions (see Fig. 4). A scheme has been described3 whereby both the stator windings of a machine of this kind, when connected to common busbars, would operate as a variable-speed constant-frequency generator without the inherent limitations of a single induction machine. It has been clearly shown,4 however, that this scheme was based upon an erroneous or incomplete understanding of the physical principles involved and it is no longer the subject of serious pursuit. When the frequencies of the currents in the two stator windings are the same the cascade induction generator functions in its asynchronous mode, i" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0000431_amr.566.197-Figure3-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0000431_amr.566.197-Figure3-1.png", "caption": "Fig. 3 Shearing stress model scheme. Fig. 4 Hertzian crack model scheme.", "texts": [ " In this model, in a crack growth mechanism called the \u2018wedge effect\u2019, cracks develop in mode I. When applied to the present tests, this model suggests that the inner surface of holes is evenly pressurized by the fluid (Fig. 2 (c)), and this pressure is the initial cause of cracking (Fig. 2 (c)). (a) rolling contact fatigue machine, (b) detail of the water-lubricated tank, (c) top race and retainer, (d) bottom race, (e) cross section of the microscopic hole. The second considered option [6] was the effects of shearing stress. The scheme for this model is shown in Fig. 3: cracks initiate caused by the shear stress and grow into the material in mode II. The third considered was the Hertzian crack theory and the scheme for it is shown in Fig. 4. Hertzian crack theory is used in cases of contact between a ball and a flat surface. The maximum tensile stress occurs at the edge of contact circle, where cracks initiate. The contact circle radius is similar to the radius curvature of the observed cracks and the cracks grow diagonally in two directions outwards the contact circle under contact surface [7]" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0003970_s11370-019-00294-7-Figure8-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0003970_s11370-019-00294-7-Figure8-1.png", "caption": "Fig. 8 Parallel RPR-to-serial manipulator a a 3-RPR parallel manipulator and b its derived serial manipulator", "texts": [ " However, their directions are the same. In order to make comparison between serial manipulator and its derived parallel manipulator, the following conditions should be met in (28) and (34). What to declare in this section is summarized as follows: (i) Geometry of parallel manipulator is expressed by ray coordinate (ii) Geometry of its inversion is found as the geometry of serial manipulator expressed in ray coordinate A common 3-RPR parallel manipulator has been shown (with active prismatic joints only) in Fig.\u00a08a. The coordinates of instantaneous motion of moving platform is given as The instantaneous first-order kinematics of the platform, which is common in three chains, is given by (35)T u = [ s\u0302 23 s\u0302T 23 (S\u03021)1 s\u0302 31 s\u0302T 31 (S\u03021)2 s\u0302 12 s\u0302T 12 (S\u03021)3 ] T \ud835\udf19 . s\u0302T 23 S\u03021 = s\u0302T 23 (S\u03021)1, s\u0302T 31 S\u03022 = s\u0302T 31 (S\u03021)2, s\u0302T 12 S\u03023 = s\u0302T 12 (S\u03021)3. (36)T\u0302 = [ v0x v0y \ud835\udf14x ]T . where 1i and 3i are the angular joint speeds of the first and third revolute joint in ith chain, respectively.(S\u03021)i and (S\u03023)i are the axis coordinates of the first and third revolute joint in ith chain", " Then, we can design a serial manipulator with three revolute joints by setting as (37)T\u0302 = \ud835\udf141i(S\u03021)i + vi(s\u03022)i + \ud835\udf143i(S\u03023)i, (38)(s\u03022) T i T\u0302 = (s\u03022) T i ( \ud835\udf141i(S\u03021)i + vi(s\u03022)i + \ud835\udf143i(S\u03023)i ) = vi. (39)v = jT T\u0302 (40)T\u0302 = (jT )\u22121v. (41)T\u0302 = \ufffd s\u0302 23 s\u0302T 23 (S\u03022)1 s\u0302 31 s\u0302T 31 (S\u03022)2 s\u0302 12 s\u0302T 12 (S\u03022)3 \ufffd\u23a1\u23a2\u23a2\u23a3 v1 v2 v3 \u23a4\u23a5\u23a5\u23a6 , 1 3 and inserting (42) into (41) yields where 1, 2 and 3 can be regarded as the joint angular speeds of the three revolute joints of 3-RRR planar serial manipulator generating the same end-effector twist as shown in Fig.\u00a08b. It is noted that the three columns of forward Jacobian of serial manipulator in (43) are lines expressed in axis coordinates. Therefore, it is remarked that the Jacobian of parallel manipulator in (39) composed of three lines expressed by ray coordinates is converted into lines in axis coordinates. Another common 3-RRR parallel manipulator has been shown (with active revolute joints only) in Fig.\u00a09a. The instantaneous first-order kinematics of the platform, which is common in three chains, can be represented as Here, we assume that the first joint of each chain is active joint" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0000971_1.4818823-Figure2-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0000971_1.4818823-Figure2-1.png", "caption": "FIG. 2. The schematic of computational domain and boundary conditions.", "texts": [ " r and z are the radial and axial distance of an arbitrary point of laser, and x\u00f0z\u00de is the radius of beam profile at height z. x\u00f0z\u00de \u00bc x0 ffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi 1\u00fe z R 2 s ; (13) where R is the Rayleigh length. Two models, bead on-plate and inside-groove, were set up for comparison. Joint type was the only difference between the two models. Here, the case with groove was chosen to be depicted in details since bead-on-plate models have been studied in many research reports. Figure 2 shows the coordinate system and computational domain used in the study. In order to avoid the complexity of treating the keyhole, the keyhole is excluded from the computational domain in this paper. The radius of the simplified keyhole on the plate surface is 0.5 mm. ANSYS ICEM is used to set the geometry, the grid and the boundary types of the three-dimensional computational domain, where 10 mm x 10 mm, 10 mm y 10 mm, 0 z 20 mm. The laser focus locates on (0, 0, 0). The diameter, height, and the inclination angle of the side nozzle are 4 mm, 10 mm, and 45 , respectively" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0002791_b978-0-12-812138-2.00003-9-Figure3.2-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0002791_b978-0-12-812138-2.00003-9-Figure3.2-1.png", "caption": "FIGURE 3.2", "texts": [ " Induction motors are most widely used as general-purpose motors in many industrial applications because of their low cost and rugged construction. Traditionally, induction motors had been connected directly to the line voltage of 60 or 50 Hz and operated at a nearly constant speed. However, recently, variable speed drives have been made possible by power electronic converters such as inverter, so induction motors are widely used in many applications requiring speed control. The structure of a typical induction motor is shown in Fig. 3.2. An induction motor has cylindrical stator and rotor configurations, which are separated by a uniform radial air gap. The stator and the rotor are made up of an iron core, which has windings inserted inside. The iron core is not a single solid lump but consists of a stack of insulated laminations of silicon steel, usually with a thickness of about 0.3 0.5 mm to reduce eddy current losses. The iron core is made from a ferromagnetic material such as steel, soft iron, or various nickel alloys to produce magnetic flux efficiently and reduce hysteresis losses" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0001460_j.sna.2010.05.007-Figure1-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0001460_j.sna.2010.05.007-Figure1-1.png", "caption": "Fig. 1. Schematics of Micro USM I.", "texts": [ " From the above background, reducing the total number of irings may put ultrasonic motors beyond the mere possible canidate for multi-DOF system actuators. Hence, this study proposes driving method for multiple ultrasonic motors using common line ignal input. . Basics of traveling wave ultrasonic motor There are two kinds of ultrasonic motors according to the differnce of basic driving principles. One is a traveling wave ultrasonic otor, and the other is a standing wave ultrasonic motor [9]. The ormer generally has higher performance than the later. Therefore, his study adopts the traveling wave ultrasonic motor as a research ubject. Fig. 1 shows a bar-shaped traveling wave ultrasonic motor [10]. t mainly consists of metal rings, a shat, and a piezoelectric ceramic ing (stacked PZT). The metal/piezoelectric ceramic rings are unied with the center shaft, which is screwed in the top ring, making he stator to be, so called, a bolt clamped Langevin vibrator. The rst natural bending vibration mode in the x-/y-direction could be electively excited by inputting an appropriate sinusoidal signal, hose frequency is corresponding to the natural frequency of the ibration mode, to the stacked PZT ring" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0000649_1.4001727-Figure10-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0000649_1.4001727-Figure10-1.png", "caption": "Fig. 10 Instantaneous screw axis ruled surface", "texts": [ " If the motion or the trajectory of each of the wo trajectory lines is given as parametric ruled surfaces Ri ui , u\u0304i , i=1,2 , then the central normals Ni i=1,2 and the entral tangents Ti i=1,2 are found from Eqs. 12 and 13 as Ni = ni ni , n\u0304i ni i = 1,2 Ti = ui ni ni ,ui n\u0304i ni + u\u0304i ni ni i = 1,2 19 he central normals Ni i=1,2 correspond to the normal to the elocity vector in the two-dimensional approach. In the threeimensional case, the instantaneous screw axis of the motion will e the common perpendicular between the two normals N1 and 2, as shown in Fig. 10. The instantaneous screw axis is written as S = s, s\u0304 = s, \u2212 1 s \u00b7 s s \u00b7 n\u03041 n2 s + s \u00b7 n\u03042 s n1 20 here s=n1 n2. If the given data are the point features of the body and their elocity vectors rather than two line trajectories, we form two ines Ri= ui , u\u0304i , i=1,2, with the Pl\u00fccker coordinates that belong o the body. Now, we have to write the velocity screw or the entral tangent Ti and the normal Ni to each line in terms of point eatures that coincide with the lines and their velocity vectors. For each line, we know the coordinates of two points and their elocity vectors" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0000775_gt2012-68986-Figure10-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0000775_gt2012-68986-Figure10-1.png", "caption": "FIGURE 10. Simplified shroud contact", "texts": [ " 9 shows a zoom into the response of x1. The comparison shows the numerical and analytical calculation of the displacement of x1. The results of the two different computation methods match very well. Nonlinear system To validate the computation of the nonlinear system the results of the analytical and the numerical approach are compared. The system investigated here consists of two 2-DOF systems that are coupled by the Masing model as described above. The model can be interpreted as a strongly simplified shroud contact, see Fig. 10. The system parameters are given in Table 2. Due to the study of parameters given in the next section viscous damping was set to zero. The system is excited at mass 3 and 4. The phase angular difference between F3 6 Copyright c\u00a9 2012 by ASME Downloaded From: http://proceedings.asmedigitalcollection.asme.org/ on 02/02/2018 Terms of Use: http://www.asme.org/about-asme/terms-of-use and F4 is \u03c6 = \u03c0 2 . The parameters are also selected due to the fact that on the interval between 0s and 100s sliding and thus friction damping occurs" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0001284_0954405411424039-Figure2-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0001284_0954405411424039-Figure2-1.png", "caption": "Fig. 2 Structure of the assembly machine for assembled camshafts", "texts": [], "surrounding_texts": [ "Keywords: assembled camshaft, S-type curve acceleration/deceleration, numerical control, driving equipment, phase angle\n1 INTRODUCTION\nThe processing technology of knurling-joining for the assembly of camshafts (Fig. 1) is an advanced processing technique [1\u20135]. Compared with the traditional one-body camshaft manufacturing process [6] and other processing techniques for camshaft assembly such as welding, sintering diffusion, and tube expander, the processing technology of knurling-joining has unparalleled advantages [7]. The assembly machine for assembled camshafts is specific equipment developed on the basis of knurlingjoining technology. It assembles the flange, cam, and core shaft that are made separately according to a certain order, axial dimension, and phase angle. In the process of assembling the camshaft, a high speed of assembly must be ensured (assembly time of 90\u2013120 s per piece). At the same time, assembling precision should be ensured (error of axial dimension of the cam should be 60.2 mm and error of\nphase angle of the cam should be 630#). The axial dimension of a cam is shown in Fig. 8.\nAs for the assembly machine, under the condition of ensuring the mechanical manufacturing accuracy, the control precision of its numerical control (NC) system is very important. At present, acceleration and deceleration control in the NC system applies an S-type curve acceleration and deceleration algorithm that makes the change of the speed stable in the process of feeding and processing; i.e. the acceleration and deceleration of the system has\n*Corresponding author: Roll Forging Institute, Jilin University,\nChangchun 130025, People\u2019s Republic of China.\nemail: linbj@jlu.edu.cn\nProc. IMechE Vol. 226 Part B: J. Engineering Manufacture\nby guest on March 5, 2015pib.sagepub.comDownloaded from", "flexibility. However, since the S-type curve acceleration and deceleration algorithm involves seven phases and there are many parameters involved, the length of the program segment varies a lot and it is very complicated to realize. This affects the working efficiency. For the assembly machine, seeking high processing speed and orientation precision has very important practical meaning. In the current paper, the S-type curve acceleration and deceleration algorithm is researched, the acceleration and deceleration phase of the algorithm is revised, and the lowspeed orientation phase is cancelled. Therefore the servo system can orientate at high speed and high precision. The practicability of this new algorithm has been verified by experiment.\n2 WORKING PRINCIPLE OF ASSEMBLY\nMACHINE FOR ASSEMBLED CAMSHAFT\nAs shown in Figs 2 and 3, in assembling the camshaft, the machine does the up and down movement along the Z axis and the rotating movement along axes A and B. Assisted by the movement of the cylinder, the machine finishes the assembly of the camshaft. The flowchart of the work procedure is shown in Fig. 4.\nIn this process, it is very important to choose a suitable acceleration and deceleration algorithm to control the orientation precision and response time.\n3 THE ACCELERATION AND DECELERATION\nCONTROL ALGORITHM OFTEN USED BY NC SYSTEMS\nAcceleration and deceleration control of the driving equipment is the core technique of the NC system.\nAcceleration and deceleration control algorithms suitable for different kinds of machines are proposed in many documents, such as straight line, exponent, and S-type curve acceleration and deceleration [8]. The S-type curve acceleration and deceleration algorithm is the one often used nowadays in NC systems. Attenuating the acceleration in the starting phase, this algorithm can ensure that the function of the motor is brought into full play and thus reduce the impact of the start-up [9\u201311].\nThe performance process of S-type curve acceleration and deceleration can be divided into seven phases: plus-acceleration phase, uniform acceleration phase, minus-acceleration phase, uniform speed phase, plus-deceleration phase, uniform deceleration phase, and minus-deceleration phase. The S-type curve acceleration and deceleration process is not uniform acceleration and uniform deceleration (Fig. 5); the acceleration on any point is continuously changing [12, 13].\nIn Fig. 5, J is the plus-acceleration, i.e. the change rate of acceleration given by J = da=dt; t is the coordinate of time, where ti (i = 0, 1, . . . , 7) represents the transition point between phases; ti (i = 0, 1, . . . , 7) is the coordinate of local time, i.e. the time indicated under the condition that the starting point of each phase is used as the zero point, ti?t?ti 1 (i = 0, 1, . . . , 7); Ti (i = 0, 1, . . . , 7) is the continuous performance time of each phase; Vmax is the maximum stable speed; and Amax is the maximum acceleration. In this process\nT1 \u00bc T3 \u00bc T5 \u00bc T7 \u00bc Amax J (1)\nT2 \u00bc T6 \u00bc V Amax \u00bc Amax J (2)\nThe acceleration a, feeding speed v, and feeding displacement s in the S-type curve acceleration and\ndeceleration can be formulated as\nProc. IMechE Vol. 226 Part B: J. Engineering Manufacture\nby guest on March 5, 2015pib.sagepub.comDownloaded from", "a(t) =\nJt, 0 < t \\ T1 A, T1 < t \\ T2 A Jt, T2 < t \\ T3 0, T3 < t \\ T4\nJt, T4 < t \\ T5 A, T5 < t \\ T6 A + Jt, T6 < t \\ T7\n8>>>>>< >>>>:\n(3)\nv(t) =\nv0 + 1 2 Jt2, 0 < t \\ T1 v1 + At, T1 < t \\ T2 v2 + At 1 2 Jt2, T2 < t \\ T3 v3, T3 < t \\ T4 v4 1 2 Jt2, T4 < t \\ T5 v5 Jt, T5 < t \\ T6 v6 Jt + 1 2 Jt2, T6 < t \\ T7 8>>>>< >>>>>:\n(4)\nand\ns(t) =\nv0t + 1 6 Jt3, 0 < t \\ T1 s1 + v1t + 1 2 At2, T1 < t \\ T2 s2 + v2t + 1 2 At2 1 6 Jt3, T2 < t \\ T3 s3 + v3t, T3 < t \\ T4 s4 + v4t 1 6 Jt3, T4 < t \\ T5 s5 + v5t 1 2 Jt2, T5 < t \\ T6 s6 + v6t Jt2 + 1 6 Jt3, T6 < t \\ T7 8>>>>>< >>>>>:\n(5)\nIn equations (3), A is the acceleration. In equation (4)\nv1 = v0 + 1\n2 JT 2 1\nv2 = v1 + AT2\nv3 = v4 = v2 + AT3 1\n2 JT 2 3\nProc. IMechE Vol. 226 Part B: J. Engineering Manufacture\nby guest on March 5, 2015pib.sagepub.comDownloaded from" ] }, { "image_filename": "designv11_62_0001594_eej.22414-Figure1-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0001594_eej.22414-Figure1-1.png", "caption": "Fig. 1. Two types of magnetic gears. [Color figure can be viewed in the online issue, which is available at wileyonlinelibrary.com.]", "texts": [ " In Section 3 we confirm the effects of cogging torque on speed control via preliminary experiments and propose improvement of response via feedforward compensation. Suppression of oscillations of a two-inertia system is discussed in Section 4 using preliminary experiments. In Section 5 we develop a reaction force estimation mechanism with regard to torsion torque, and introduce force sensorless bilateral control. Conclusions are formulated in Section 6. 2.1 Structure of magnetic gears The two types of magnetic gears considered in this study are shown in Fig. 1. The gears in Fig. 1(a) make possible changing the direction of force transmission between two axes on the same plane, thus producing the same effect as bevel gears. Below we call this type magnetic coupling. These magnetic gears are used in experiments with speed control and position control; six permanent magnets are arranged on both the gear and pinion. Figure 1(b) shows magnetic gears providing speed reduction. This gearbox is composed of a two-stage spur gear train. Each stage has a reduction ratio of 4:1 and the total ratio is 16:1. This gearbox is used in experiments with force sensorless bilateral control of master\u2013slave robots. Below we call this type a magnetic gear reducer. There is an axial gap between the magnetic gear and the pinion, thus providing a contactless arrangement. The gear and pinion are equipped with rare earth magnets offering a high maximum energy product (SmCo5, Nd2Fe14B, etc" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0001223_03091902.2013.785608-Figure17-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0001223_03091902.2013.785608-Figure17-1.png", "caption": "Figure 17. Mouthpiece in SolidWorks.", "texts": [ "4 inches) for males and 17 cm (6.8 inches) for females. In terms of hand size across the palm, the average for males was 8.4 cm (3.3 inches) and 7.5 cm (2.9 inches) for females [19]. Both the length and outer diameter of the spirometer correlated with these dimensions seen in Figure 16. The basic geometry of the spirometer consisted of two cylinders with a cone connector, providing a sleek transition from the inlet to outlet tube as shown in Figure 16. The mouthpiece was shaped in an elliptical design shown in Figure 17. This shape was chosen due to ease of handling for the patient; however, the fundamentals of how the pressure sensor works also made this geometry most suitable. Based on average mouth sizes, the diameter of the mouthpiece was designed to be 1 inch. The sensor chosen was a differential pressure sensor. In order for it to function, one port of the sensor measures the pressure near the base of the large outlet cylinder, while the other port measures near the base of the inlet smaller cylinder (model of small cylinder can be seen in Figure 12)" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0002865_we.2149-Figure1-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0002865_we.2149-Figure1-1.png", "caption": "FIGURE 1 Flexure bearings: A, cartwheel hinge; B, cruciform hinge", "texts": [ " In contrast to this, flexure bearings (FBs) are based on a completely different functional principle. They realize the necessary motion by way of an elastic deformation of the mate- Wind Energy. 2017;1\u201310. wileyonlinelibrary.com/journal/we Copyright \u00a9 2017 John Wiley & Sons, Ltd. 1 rial instead of movements between surfaces of solids.7 These displacements may be translational, rotary, or a combination of both. Examples of well-known rotational hinges are the cartwheel hinge and the cruciform hinge depicted in Figure 1. Our research work was driven by the hypothesis that FBs can handle the unfavorable load conditions associated with IPC, especially the small swivel angles, better than conventional rolling bearings. On the basis of this, we propose a novel pitch bearing concept. An FB is combined with a conventional bearing. The FB allows pitching of a few degrees by elastic deformation alone, which is used to realize all small, ongoing motions required for IPC. Meanwhile, the conventional bearing is only used for pitch angle changes exceeding the range of the FB" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0000786_aero.2013.6496960-Figure2-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0000786_aero.2013.6496960-Figure2-1.png", "caption": "Figure 2 \u2013 Concept of the underground explorer with the propulsion and excavation units", "texts": [ " Thus, when compared with other locomotion strategies in terms of space required, the large contracted area with this type of mechanism is useful for moving a narrow object such as a pipe and perforating soil. This mechanism is therefore suitable for a subsurface explorer robot. As a subsurface explorer excavates deeper, greater earth pressure is exerted from the surrounding walls because of regolith. The density of the regolith in front of the excavator also makes downward excavation difficult. Thus, we propose a novel subsurface explorer, as illustrated in Figure 2. Our robot consists of two units: one for propulsion and the other for excavation. We adopted the peristaltic crawling of an earthworm as the mechanism for the propulsion unit, with the excavation tool modeled after an EA. The propulsion unit of the robot consists of four parts, corresponding to the segments of an earthworm. Each part can expand in a radial direction as it contracts in an axial direction and generates a large amount of friction between the surrounding regolith and its body. The propulsion unit controls the contact and non-contact of an expansion plate with the surrounding wall, maintains the rotation reaction of the excavation unit and firmly pushes the excavation unit into the front regolith" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0000530_embc.2012.6347434-Figure1-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0000530_embc.2012.6347434-Figure1-1.png", "caption": "Fig. 1. Position constraints for verification with ADINA. The position of the top nodes are fixed, the x position of the middle nodes are moved 50% of the width( free for y, and z directions), and the position of the lowest nodes are moved 30% of the height in the z direction ( free for x and y directions).", "texts": [ " Unfortunately, this does not satisfy the inf-sup criteria of LadyszhenskayaBabuska-Brezzi-KikuchiiLBBK) [3]. For a surgical simulator, however, it is reasonable to adopt that shape function because (1) the criteria indirectly show that a situation in which the equations are greater than the unknown has never happened in a surgical simulator, and (2) the process is highly timecritical. We used TABLE II for material properties. To adjust the runtime conditions, we used 4\u03b1, as explained later, in the simulator as ADINA\u2019s bulk modulus \u03ba. We performed the large deformation indicated in Fig. 1 using the above runtime conditions. As seen in the figure, the positions of the top nodes are fixed, the x position of the middle nodes are moved 50% of the width (free for y and z directions), and the position of the bottom nodes are moved 30% of the height in the z direction (free for x and y directions). B. Verification with ADINA TABLE III compares the deformation between ADINA and the implemented simulator for a nearly-incompressible hyperelastic model. Subtable (a) compares the displacement, 1 \u03a6\u0303prj = \u222b WMd\u2126 + \u03ba 2 \u222b Prj(WV )2\u03bbd\u2126 + \u03a6ext, Prj(\u03baWV ) = \u03bb and subtable (b) indicates the relative error on each node and the average relative error for all nodes" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0003738_ls.1476-Figure1-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0003738_ls.1476-Figure1-1.png", "caption": "FIGURE 1 A, Mesh adopted for the simulations; B, back\u2010to\u2010back test rig", "texts": [ " In this case, also the source term _m has to be removed from the continuity equation and the sub\u2010 index \u2223v refers to the gaseous phase (air). The rotation of the boundaries of the gears during the simulation implies a significant topological modification of the computational domain and grid, which determines, after few time steps, an unacceptable distortion of the initial mesh, which needs to be updated in order to grant the stability of the numerical solution. For this purpose, the authors have developed a methodology13 consisting in a globlal\u2010remeshing approach (Figure 1A). The grid is completely substituted when it loses a quality compatible with the solution process. The adopted meshing strategy is efficient, and the update of the grid is possible in a very low time. If compared with the local\u2010remeshing technique, typically implemented by many general\u2010purpose commercial software, the approach adopted achieves a much more regular element size, because with local\u2010 remeshing, the volume to be re\u2010meshed is not known in advance and therefore it is not controllable. The larger time step that can be kept for the same Co number leads to a strong reduction of the simulation times, with a reduction of the computational effort estimated at \u221293.5%, as shown in Concli and Gorla.14 The abovementioned approach was applied to a standard back\u2010to\u2010back test rig configuration (Figure 1B). The gearbox under analysis has a very simple parallelepiped case (test gearbox), the dimensions of which are 145 \u00d7 290 \u00d7 h185 mm. The considered gears are standard C\u2010PT gears: The macro\u2010geometrical parameters are listed in Table 1. The lubricant modeled in the simulations is an FVA2 gear\u2010oil, the properties of which are listed in Table 2. The simulations were performed for two different conditions: complete\u2010 (100% filling) and dip\u2010lubrication (up to the axis\u201450% filling). Two rotational speeds, 420 and 840 rpm corresponding to 19" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0001166_gt2012-69356-Figure1-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0001166_gt2012-69356-Figure1-1.png", "caption": "Figure 1: Schematic of a Compliant Plate Seal with Intermediate Plate", "texts": [ "asmedigitalcollection.asme.org on 10/23/2018 Terms of Use: http://www.asme.org/about-asme/terms-of-use 2 Copyright \u00a9 2012 by ASME DP Differential pressure e Effective clearance Q Flow parameter ?\u0307? Mass flow rate (leakage) T Temperature Specific heat capacity ratio R Gas constant IP Intermediate plate IP\u2013FG Intermediate Plate \u2013 Front Gap IP\u2013BG Intermediate Plate \u2013 Back Gap 1/rev Once per revolution RPM Rotations per minute SC Effectiveness of self-correcting behavior The Compliant Plate Seal as shown in schematic Figure 1 consists of compliant plates (around 0.125-0.275 mm thick) attached to the stator at the seal outer diameter, in a circumferential fashion around a rotor. The compliant plates have a slot extending radially inwards from the seal outer diameter, and an intermediate plate extends inwards into this slot from stator. The compliant plates are inclined and the angle with respect to the circumferential direction is called the cant angle, which is typically around 30-60 degrees. The compliant plate stack and the intermediate plate are separated by small gaps on the high pressure side and low pressure side, known as the intermediate plate front gap (IP\u2013FG) and the intermediate plate back gap (IP\u2013BG) respectively" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0000001_detc2011-48237-Figure1-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0000001_detc2011-48237-Figure1-1.png", "caption": "FIGURE 1. SCHEMATIC VIEW OF A REDUNDANT MANIPULATOR.", "texts": [ " A FT B F 0 0 C1 0 0 q\u0308 \u03bb U = D1 +Q D2 y\u0308 (6) Equation (6) can be solved for the second derivatives of the generalized coordinates. In order to obtain the motion of the system generalized coordinates can be double integrated. The required control inputs vector U and the Lagrange multipliers \u03bb will be automatically calculated during the process. The DYSIM [19] software will be used to construct the equations of motion automatically, and to build both forwards and inverse dynamic models for the system. In this study DYSIM will operate within the MatLab/Simulink environment [20]. Consider the n-links manipulator shown in Fig. 1, where absolute joint angles are denoted by \u03b8i and joint lengths are denoted by li for the ith link. The manipulator task consists of transporting a load mass ml from an initial point Pi to a destination Pf in Cartesian space. This system can be modeled by using the Lagrange\u2019s equation of motion as described in the previous section. Dysim is utilized to automatically developed a dynamic model of the system. Cartesian coordinates of the centre of gravity and absolute angles of each link plus the Cartesian coordinates of the load are selected as generalized coordinates, i", " Since fx and fy are generated from the optimization parameters, this constraints can be checked before calling the inverse dynamics. Redundant Links Reach. Another requirement for the motion defined by the optimization parameters to be realizable is that the distance between the end of the last redundant link (point at xr, yr) and the end effector (point at fx, fy) must be less than or equal to (ln\u22121 + ln). This is equivalent to replacing the redundant links by a single virtual link Rrm as shown in Fig. 1. This constraint can be formulated as follows: max \u221a ( fx \u2212 xr)2 +( fy \u2212 yr)2 \u2264 ln\u22121 + ln (14) where xr and yr can be calculated from the optimization parameters as follows: xr = n\u22122 \u2211 i=1 li cos\u03b8i and yr = n\u22122 \u2211 i=1 li sin\u03b8i (15) To demonstrate the proposed method, a 5-link redundant (with three DOF redundancy) manipulator is introduced. The manipulator has five identical links, and each link length is selected as 0.4 m. Each link has an inertia of 0.0001 kgm2. All the links have identical mass of 1 kg each" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0003514_978-3-030-20131-9_324-Figure20-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0003514_978-3-030-20131-9_324-Figure20-1.png", "caption": "Fig. 20. Beaded string process state diagram of the uplink line penetrating into the common beads.", "texts": [ " Step 4: When the pushing block of the downlink line in the feeding state, the control motor D is drive to rotate the output port and the braided port counterclockwise by 180\u00b0. Step 5: Under the driving of the control motor A, the bead threading device moves to the right. When the driving roller moves in the first trapezoidal block on the movable seat rail, the location and clamping position of needle will be changed to make the uplink line successfully penetrate into a shared bead provided by the downlink line feeding device. Step 6: The linear motor E is reversely driven to make the pushing block of the downlink feeding bead out of the feeding state, as shown in Fig. 20. Method Research and Mechanism Design of Automatic Weaving\u2026 2547 Step 7: Repeat the actions from steps 1 to 6 until the end of the weaving task. Step 8: When the weaving process is finished, each motor is controlled by software programming to bring the device into an initial state. Based on the problem existed for the hand-weaving method, the automatic weaving method and mechanism device are studied and discussed. The weaving route was extracted by discussing the hand-weaving method of the beaded cool pad" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0003356_978-981-13-5799-2_12-Figure12.111-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0003356_978-981-13-5799-2_12-Figure12.111-1.png", "caption": "Fig. 12.111 Flow distribution in soil at a slip ratio of 30% (reproduced from Ref. [153] with the permission of Tire Sci. Technol.)", "texts": [ " The tire size is 540/ 65R30, the inflation pressure is 240 kPa, and the load is 32.86 kN. The tire is modeled using Lagrangian elements, while a soil bin having a depth of 0.326 m is modeled using Eulerian elements. The coefficient of friction between the tire and the soil surface is ignored. Figure 12.110 shows the calculated rut shape of a tire with a slip ratio of nearly 100% and the calculated cone penetration resistance of soft soil. The tire sinkage is nearly 110 mm, and the result is in good qualitative agreement with common phenomena. Figure 12.111 is a flow diagram of a tire driving at a slip ratio of 30% on hard soil. Most of the soil beneath the tire flows backward, and the motion of the soil is limited to the shallow part of the soil region aligned to the top surface of the tire. The gross traction is dominant, and the motion resistance is low. The gross traction at a slip ratio of 30% is produced by pushing the soil between the lugs backward. The right figure in Fig. 12.112 shows the motion resistance distribution per unit area in this state" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0002350_0954406217712280-Figure2-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0002350_0954406217712280-Figure2-1.png", "caption": "Figure 2. KISSsoft visualized 3D model of the FZG test equipment.", "texts": [ " Note that the standard geometry is referred to as non-tip relieved (NTR) and the modified as tip relieved (TR). Material data used are linear elastic steel of full density 7.85 kg dm\u20133, elasticity of 210GPa and Poisson\u2019s ratio of 0.3 and PM reduced density of 7.25 kg dm\u20133, elasticity of 150GPa and Poisson\u2019s ratio of 0.28. A commercial transmission software such as KISSsoft handles and presents the input and output parameters for gear manufacturing, geometry and dimension criteria in a structured and systematic manner. With this tool, the gear geometry Figure 2 has been modelled and analysed. The model is based on five shafts modelled as beam elements, supported by eight cylindrical roller bearings SKF NJ406 and one ball bearing SKF6306. The shafts are connected by rigid couplings. At the bottom connection close to the ball bearing, the pretension torque for the specific load stage is set. The driving torque on the input shaft is then found by iterative search for the torque balancing the frictional losses at a given speed. The gear rating calculations are done according to ISO 6336:2006 Method B" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0002375_iea.2017.7939175-Figure4-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0002375_iea.2017.7939175-Figure4-1.png", "caption": "Figure 4. (a) Parts designed via SolidWork 2014x64 Edition software and (b) Parameter setting via Cura 14.01 software", "texts": [ " The variable factor and fixed level factor are assigned, as shown in Table I. The experimental campaign involved specimen made of stainless steel PLA, filament developed by Proto-Pasta Inc., a widely used composite material for FDM processed parts. The dimension of specimen is 10mm \u00d7 17.5mm \u00d7 60mm (width \u00d7 height \u00d7 length), as shown in Fig. 3. The specimen were designed via SolidWork 2014x64 Edition software and exported as STL file. The STL file is imported to Cura 14.01 software for parameter setting and model printing, as shown in Fig 4. Before proceeding with experimentation, it was necessary to control the mass of the stainless steel PLA filament used for experimentation. In order to perform this experiment, the mass of the stainless steel PLA was calculated using Archimedes principle, are as follows: SSPPLA = m / V (1) m = SSPPLA \u00d7 V (2) where SSPPLA is the nominal density of the stainless steel PLA is given as 2.38 g/cm3, m is the mass of material (g) and V is the volume of material (cm3). The experiment was conducted by using standard nozzle size 0" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0000177_roman.2013.6628430-Figure1-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0000177_roman.2013.6628430-Figure1-1.png", "caption": "Figure 1. Exterior of FRC.", "texts": [ " INTRODUCTION The purpose of this research is to achieve visual information augmentation technology with multiple projectors system, which is called FRC[1] in this project. Current computers are mainly for a single user with restricted input and output devices. They can be used for cooperation of multi-user. However in such a case, the computers are connected through network and users use the computers no more than as communication tools for the cooperation. One of the main purposes of FRC is to overcome the limitation of current computer. Fig.1 shows the prototype of FRC. Since it have two sets of camera sensor and image projector module with separately drivable pan-tilt actuators, the FRC is able to sense users\u2019 gesture and project CG images directly on the wall or the floor simultaneously. By putting the FRC on a table, the table will be changed into a smart table that helps cooperation between the users. The users do not need to be aware of input or output device. The table where the FRC is put on becomes an area of input and output" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0001189_jtam-2013-0001-Figure2-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0001189_jtam-2013-0001-Figure2-1.png", "caption": "Fig. 2. Dynamic model", "texts": [ " These reactions are caused by the external loading and the kinematics and mass characteristics of the rotating disk. The expressions for the full dynamic reactions are obtained. These expressions allow the parameters of the machines to be chosen in such a way that the loading in the shaft and the bearings to be minimal. Figure 1 shows scheme of the band saws [1, 2, 3, 4, 5]. We define the following symbols: 1, 2, 5, 6 \u2013 belt pulleys, E \u2013 electric motor, 3 and 4 \u2013 feeding wheels, A \u2013 band-saw blade, 7 and 8 \u2013 chain-wheels. The dynamic model, shown in Fig. 2 is used for solving the problems. The feeding wheel 3 and the belt pulleys 2 and 5 perform rotation with a constant angular velocity \u03c9 about the axis of rotation AB. In this case, the mechanical system (the feeding wheel, the belt pulleys and the basic shaft) describes an angle \u03d5 = \u03c9t. Unauthenticated Download Date | 5/15/17 2:19 PM We choose the following coordinate systems: Fixed coordinate system O3xyz, moving coordinate system O3x1y1z1, which moves along with the feeding wheel. In the initial moment (\u03d5 = 0), the axes of the two coordinate systems coincide" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0003257_1.5090680-Figure2-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0003257_1.5090680-Figure2-1.png", "caption": "FIGURE 2. Five-parameter MDH model", "texts": [ " 1 1 1 1 1( , , , , ) ( , ) ( , ) ( , ) ( , ) ( , )i i i i i i i i i i i iT a d Trans x a Rot x Trans z d Rot z Rot y 1 1 1 1 1 1 11 1 1 1 1 1 1 0 0 0 1 i i i i i i i i i i i i i i i i i i i ii i i i i i i i i i i i i i i i c c s c s a s c c s s c c s c s s c d s T s s c c s c s s s s c c d c (1) Here, 1i iT is the forward kinematic expression of the connecting rod i to the connecting rod 1i , c means cos , s means sin , ia is the length of the connecting rod i , id is the offset of the connecting rod i , i is the torsion angle of the connecting rod i , i is the rotation angle between the adjacent two joint axes 1iZ and iZ on the plane i i iX O Z , and i is the joint angle of the connecting rod i ,as shown in Fig. 2: 020026-2 The MDH parameters of each joint of the manipulator are shown in Table 1. The homogeneous transformation matrix of each joint can be obtained by substituting the parameters of each joint into the transformation matrix. The transformation matrix of the position relationship between the end-effector of the manipulator and the base can be obtained by multiplying the homogeneous transformation matrix to the left, such as formula (2): 0 0 1 2 3 4 5 1 2 3 4 5T T T T T T (2) ERROR MODEL OF PSO-RBF NEURAL NETWORK Particle Swarm Optimization Particle swarm optimization (PSO) algorithm is a random global optimization algorithm, and it is based on the principle of simulated bird predation behavior [7]" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0002302_b978-0-12-803581-8.10351-0-Figure6-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0002302_b978-0-12-803581-8.10351-0-Figure6-1.png", "caption": "Fig. 6 U-tube pressure vessel components. Reproduced from Radford, D.W., Fuqua, P.C., Weidner, L.R., 2004. Tooling development for a multishell monocoque chassis design. In: 36th International SAMPE Technical Conference, San Diego, CA, November 15\u201318, 2004, pp.1063\u20131077.", "texts": [ "6 Previous research on multi-shell pressure vessels has shown that a multi-shell approach is viable in high-stress applications.6 Pressure vessels were developed and evaluated for two reasons: to investigate the viability of packaging internal components and to allow more complex-shaped pressure vessels (e.g., U-shapes) to be created. In addition, the design was cost constrained, thus placing an emphasis on manufacturability. The pressure vessel design used four mating parts, two inner and two outer, each fully overlapping (Fig. 6). This approach generates a large lap-bond area, enabling the high performance under internal pressure, distributed loading. This design and manufacturing method allowed the inner and outer shells to be tooled separately. This gave control over the bond surfaces, ensuring a high efficiency joint, and made for simplified tooling and assembly. It was shown that very good structural properties could be generated from this approach and that the weight could be held in check. FEAs of various specific cases of shell design were undertaken, and the models accurately represented the pressures and forms of failure realized during hydroburst tests" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0003126_infocomtech.2017.8340599-Figure1-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0003126_infocomtech.2017.8340599-Figure1-1.png", "caption": "Fig. 1: Rotary inverted pendulum", "texts": [ " Finally the LQR is implemented on to the rotary inverted pendulum; is one of the benchmark problems in the world of control system as it is a nonlinear and highly unstable system [6]. The paper starts with a brief introduction in section I, section II speaks of the rotary inverted pendulum system, section III discusses the LQR. In section IV a brief idea of the optimization techniques used is discussed, section V describes the simulation work and implementation. Finally in section VI results are drawn followed by references. II. ROTARY INVERTED PENDULUM The rotary pendulum used for the paper is given in fig. 1 which has two degree of freedom with a single actuator. It is also known as Furuta pendulum as it was developed by K.Futura at Tokyo Institute of Technology [7]. The total setup has four main parts which is the pendulum arm connected to a rigid link at one end and to the pendulum on the other side. The servo motor provides the necessary actuating signal or the control signal. is the pendulum arm angle and is the pendulum angle. The assumptions considered for modeling of the system are [9]: Zero initial conditions are inferred when system initiates from equilibrium state" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0000149_15421406.2011.571975-Figure1-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0000149_15421406.2011.571975-Figure1-1.png", "caption": "Figure 1. Schematic representation of the geometry of the liquid crystal cell and the flexoelectric bending of initially homeotroic nematic layer with minute amount of SWCNTs in a transversal electric field E.", "texts": [ " Additionally we have measured the value of the surface energy for the pure nematic E7 and the E7=SWCNTs mixture. Keywords Flexoelectric bending; nematic E7; nematic-carbon nanotube interactions The discovery of the piezoelectricity (flexoelectricity) in the nematic liquid crystals began in 1969 with the famous theoretical prediction by Meyer [1]. Among the many flexoelectric effects discovered in nematics (see for instance the book by Pikin [2] and the review by Petrov [3]), a typical flexoelectric effect is the bending (splaying) of weakly anchored nematic in a transversal electric field (Fig. 1). The early observation of the bending of homeotropic MBBA (p-methoxy benzylidene p\u2019-butyl aniline) in a transversal d.c. electric field had been made by Haas, Adams and Flannery [4]. Helfrich [5] first proved the flexoelectric nature of this bending by developing a simple theory, based on minimization of the electric enthalpy including the elastic, flexoelectric and dielectric energies. Schmidt, Schadt and Helfrich [6] reconsidered experimentally the bending of the nematic MBBA and estimated the value of the bend Address correspondence to Y" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0000329_s1068366612030105-Figure1-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0000329_s1068366612030105-Figure1-1.png", "caption": "Fig. 1. Model for mechano roling fatigue tests.", "texts": [ " From the viewpoint of tribo fatigue, the friction coefficient in force systems is determined as a function of the contact pressure p and the cyclic stresses \u03c3 resulting from the off contact load: (2) According to the data of [4\u20136], depending on the conditions of friction or testing it can be (3) i.e., the cyclic stresses can either increase or reduce the friction coefficient. In this work, we carried out the systematic compar ative analysis of relations (1)\u2013(3) based on the experi mental results; all tests were performed under rolling friction conditions so that is the coefficient of resistance to rolling in the friction pair and is the coefficient of resistance to rolling in the force system. A modification of the roller/shaft force system (steel 18KhGT/steel 18KhGT) was tested (Fig. 1); the test method was published earlier [7]. NF p S W N ( ) , F f p F p \u03c4 = = S N ( , ) ( , ) . F p f p F \u03c3 \u03c3 = ( , )f p \u03c3 ( ),f p r( )f p f= ( , )f p f \u03c3 \u03c3 = Keywords: force (tribo fatigue) system, tribopair, friction, law of friction, coefficient of resistance to rolling, contact load, contact area. DOI: 10.3103/S1068366612030105 204 JOURNAL OF FRICTION AND WEAR Vol. 33 No. 3 2012 SOSNOVSKII et al. RESULTS OF TESTING THE STEEL 18KhGT/STEEL 18KhGT FORCE SYSTEM This force system was tested using the developed method at three values of the contact stresses (p0 = 2000 MPa, p0 = 3200 MPa, and p0 = 5600 MPa), which correspond to the range of elastic and elasto plastic deformations (see the dependence of the max imum contact pressure on the approach of the axes \u2014insert in Fig. 2). The bending load Q was assigned stepwise 5 min after the beginning of the tests. The ini tial bending load was Q = 160 N (the bending stresses were \u03c3a = 160 MPa). The increment of the bending stresses at each loading step was \u0394\u03c3i = 40 MPa = const; the duration of the step was ni = 30000 cycles. The bending load was directed downwards to produce ten sile stresses in the contact zone and upwards to pro duce compressive stresses (see Fig. 1). The degree of slippage was constant during the tests and amounted to \u03bb = 3%. The test results are presented as the graphs of the dependence of the average coefficient of resistance to rolling (under various values of the contact stresses ) on the cycle stress amplitude \u03c3a at the given load ing step (Fig. 2). Sixty values of f \u03c3 obtained during the test at one loading step (10 min) correspond to each point on the graph at \u03c3a = const; this provides a suffi cient accuracy of the results. The data presented in Fig" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0003514_978-3-030-20131-9_324-Figure76-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0003514_978-3-030-20131-9_324-Figure76-1.png", "caption": "Fig. 76. Schematic diagram of string installation.", "texts": [], "surrounding_texts": [ "In chapter 2, three methods of automatic knitting of bead mat are put forward. Among them, Warp and Weft Automatic Weaving Method is the most simple, but the cost is higher and its stability is poor. The lock stitch sewing weaving method is rea lly good, but there exists three difficulties applied in the machine: The cross-sectional size of the hook must be smaller than the size of the bead hole as the hook needle needs to completely pass through the bead, which makes it difficult to hook the string smoothly. It is difficult to guarantee the parallelism between the cross-section of the loop ring and the end of the bead of the first row in steps 4-6. When weaving a larger size beaded cooling pad, threading the string into the longer aligned transverse bead holes becomes hard. So compared with the other two methods, single-line straight-through method, which has good stability and high knitting efficiency and is easy to be realized on the machine, is the focus of the discussion below. Based on this method, an automatic weaving device capable of weaving a beaded cool pad is proposed and designed. 4.1 Feeding Device Design Before weaving the beaded cooling pad, since all the beading arrangements are disordered, it is necessary to design a device that puts the beads into the weaving state in an orderly manner, which is called feeding device. Referring to the feeding mechanism of the firecracker weaving [14-16], Fig.14 shows the schematic diagram of the designed bead feeding mechanism device. Before the device runs, all the beads are placed in the hopper and the two guiding wheels. A small number of longitudinally are placed in a horizontal arrangement. When the two guide wheels rotate in opposite directions, the beads in the hopper will be putted into the guide groove and conveyed to the front of the beading device in an orderly manner. In order to avoid a rigid collision between the feeding device and the ball transported device, the end of the guiding groove is made by a material with better elasticity. 4.2 Design and Working Principle of the Beaded Pad Weaving Device Fig. 15 demonstrated the beaded pad weaving device. Since the figure is only for explaining the movement process of the beaded pad weaving device, the feeding device is not shown in the figure. And there are 7 motors in this device. Control motor A controls the movement of the threading device. Control motor B controls the movement of the movable line of the downlink line. Control motor C controls the movement of the movable stroke switch. Control motor D controls the rotation of the output port and the braided port. Linear motor E S. Ouyang et al.2544 drives the up-line feed ball push block movement. Linear motor F pushes the braided beaded cool pad unit into the braided port. Linear motor G drives the linear motion of the downlink line feed bead block. Before the device is operated, the downlink threading is first performed. After the downlink threading is finished, the string is installed into the beaded pad weaving device. The downlink line with heavy beads at the end is wrapped around the fixed pulley mounted on the frame. Then it passes through the downlink line movable seat, the through hole, the downlink end sleeve, the bead and the braided port successively, to reach the uplink line. And the uplink line is directly connected to the needle of the threading device, the first string of beads are moved to the corresponding position on the weaving port, as shown in Fig.16. After the string installation is completed, the motor A is manually controlled to make the driving roller be located between the two trapezoidal blocks on the movable seat rail of the threading device. The manually controlled linear motor E is to drive the uplink line to send th e beads push block moves, which is external bead conveyed from the feeding device. It causes the holes axis of the bead to coincide with the needle axis of the bead threading device and the up-line bead push block is in the beading state. The specific work ing process of the device is as follows: Method Research and Mechanism Design of Automatic Weaving\u2026 2545 Step 1: Under the drive of the control motor A, the bead threading device moves to the right. When the driving roller moves in the second trapezoidal block on the movable seat rail, the location clamping position of the needle will be changed. As a result, the uplink line smoothly penetrates an external bead provided by the uplink line bead transported device, as shown in Fig.17. Step 2: The bead threading device continues to move to the right. When the movable seat contacts the movable travel switch, the uplink line is just tightened. At this time, some of the motor operation will change as follows: Control motor A reversed means the bead threading device starts to move to the left. Controlling motor B rotated forward means the downlink line movable seat moves to the right for a suitable distance, providing two beads re quired for the next unit downlink line weaving. After that, controlling motor B stops. Linear motor F runs, the weaving beads are pushed into the weaving port and the output port, then moves back to the initial position. Linear motor E reversely drives, the pushing block of uplink line feeding bead returns to the initial position, and is on out feeding condition, as shown in Fig.18. S. Ouyang et al.2546 Step 3: Under the driving of the control motor A, the bead threading device moves to the left. When the threading device contacts the fixed stroke switch, some of the motor operation will change as follows: Linear motor G forward drives, the pushing block of the downlink line feeding bead pushes a shared bead to make the hole axis of the shared bead coincide with the needle axis. Control motor A rotated forward means the bead threading device starts to move to the right. Linear motor E drives forward, the push block of the uplink line feeding bead is in the feeding state. Control motor C rotated forward, the movable travel switch moves to the left for a suitable distance exactly equal to the length of the string required to weave every bead pad unit, as illustrated in Fig. 19. Step 4: When the pushing block of the downlink line in the feeding state, the control motor D is drive to rotate the output port and the braided port counterclockwise by 180\u00b0. Step 5: Under the driving of the control motor A, the bead threading device moves to the right. When the driving roller moves in the first trapezoidal block on the movable seat rail, the location and clamping position of needle will be changed to make the uplink line successfully penetrate into a shared bead provided by the downlink line feeding device. Step 6: The linear motor E is reversely driven to make the pushing block of the downlink feeding bead out of the feeding state, as shown in Fig. 20. Method Research and Mechanism Design of Automatic Weaving\u2026 2547 Step 7: Repeat the actions from steps 1 to 6 until the end of the weaving task. Step 8: When the weaving process is finished, each motor is controlled by software programming to bring the device into an initial state." ] }, { "image_filename": "designv11_62_0001778_ijtc2010-41048-Figure1-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0001778_ijtc2010-41048-Figure1-1.png", "caption": "Fig. 1 Relative position of the pinion and the gear in mesh", "texts": [ " The systems of linear equations, obtained by using the finite difference approximation of the Reynolds, elasticity, energy, and Laplace's equations, are solved by the successive-over-relaxation method. The details of the presented theoretical background are described in Refs. [5,6]. The computer program BevEHD, based on the theoretical background presented, has been applied for the investigation of the influence of pinion\u2019s running offset ( a\u2206 ) and axial adjustment ( b\u2206 ) errors, and of angular misalignments of the pinion axis ( h\u03b5 and v\u03b5 , Fig. 1) on maximum oil film pressure ( maxp ) and temperature ( maxT ), EHD load carrying capacity (W), and on power losses in the oil film ( Tf ). The investigation was carried out for the spiral bevel gear pair of the following main design data: numbers of teeth 11 and 41, module 4.414 mm, face width 31.911 mm, mean spiral angle on pinion 35 deg. The values of lubrication characteristics and operating parameters are: ambient lubricant viscosity Pas19361.00 =\u03b7 , minimum oil film thickness m27.1h0 \u00b5= , pinion\u2019s revolution per minute 2000 rpm" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0001321_icma.2012.6283381-Figure2-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0001321_icma.2012.6283381-Figure2-1.png", "caption": "Fig. 2 2D model of digit II, manus of G. gecko. Phalanx III was omitted as it\u2019s a hook-like bone, and wasn\u2019t so important for G. gecko\u2019s climbing. Joint I and II are revolute joints. l1, and l2 are the length of Phalanx I and Phalanx II. d is the thickness of the adhesive pad. The shaded portion is the adhesive pad.", "texts": [ "1a)), numbered I to V from the preaxial to the postaxial (Fig. 1b)). The phalangeal fomula of the digits is 2-3-4-5-3 from digit I to V [16], including the claw as the distalmost phalanx except digit I which contains no claw. Metacarpals are extension of digits, and they all resemble each other in structure from digit I to V. The adhesive pad covers the distalmost phalanges of the digit. The pes differs little from the manus from the functional viewpoint. Despite the complexity, a 2-Dimension (2D) model of digit II of G. gecko\u2019s manus was built (Fig. 2). We focused on the key dimensions between interphalangeal joints and the thicknesses of the adhesive pad. Rectangular link-bars as phalanges were connected by revolute joints. A tape of elastic material as tissues of the digit was hollowed to hold the linkbars inside, part of which was adhesive on the plantar surface from the distalmost point to the middle of Phalanx I. The bars and the flexible material were stuck together as the bones and muscles in biological tissues, meaning there\u2019s no longitudinal sliding between them" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0000948_icici-bme.2013.6698502-Figure8-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0000948_icici-bme.2013.6698502-Figure8-1.png", "caption": "Fig. 8 HILS implementation", "texts": [ " The first sample time is chosen for 10 milliseconds and the second is chosen for 0.1 second. The former is used for the state estimation process and is determined by considering the aggregated time consumptions in HILS implementation consisting of three UART communication delays, state estimation process and Flight Control System. This time consumption is ranged between 6 until 9 milliseconds. The second is used for the trajectory planning block and is determined from the GPS sensor frequency (10Hz). The overview of HILS implementation is illustrated in Fig. 8. IV. TESTING AND ANALYSIS A. Trajectory Planning The validation of Trajectory Planning block is done in two separated planes, the lateral plane and longitudinal plane. For the testing in the lateral plane, the waypoint locations are assumed in Table 1. The rocket will pass the waypoint in sequence (i.e. from smallest to the biggest order). The radius of the desired orbit is set to 2 meters. Hence we get the trajectory plan that links between waypoints as in Fig. 9. As shown in Fig. 9, the rocket should pass the waypoints in counterclockwise direction" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0000802_kem.450.27-Figure4-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0000802_kem.450.27-Figure4-1.png", "caption": "Fig. 4 Product design- in front view", "texts": [ " Strengthen the folding points It was suggested to use high-density rubber pads to strengthen folding points, make tenons tense and fixed, and make the frames more firm and closed together. Then, the stability and safety during riding might be increased. 5. Use big and small gear wheels combination If using the same size of gear wheels, it would be inefficient and slow when riding. Therefore, imbalanced gear wheels combination was set to adjust and bring greater speed as well as increasing the efficiency of riding. The new designs of the improved folding bicycle in this study are shown in Fig. 3 and Fig. 4. The Contradiction Matrix of TRIZ is a powerful tool, especially when the 40 Inventive Principles can assist designers create innovative ideas. In this research, principles in TRIZ were applied to improve the capability and degrees of comfort of the folding bicycles. The main points in the designs of this study are double shock absorbers, removable hidden stand and asymmetric wheels.. The main goal of double shock absorbers is to balance the supports from both sides, to concern the structure system of the frame, and to achieve the vibration-absorbing effect" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0000185_cecnet.2011.5768595-Figure1-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0000185_cecnet.2011.5768595-Figure1-1.png", "caption": "Figure 1. Anti shovel device", "texts": [ "eywords-excavating force; dipper handle; Analysis; calculation I. INTRODUCTION Excavator anti shovel device consists of swing arm, kangaroo bar, connecting rod, movable arm hydro-cylinder, bucket cylinder and so on. The basic form is shown in figure 1. As is shown in figure 1, according to the practical work requirements, the variables of hydraulic excavator working device performance against shovel are confirmed. And the represent sizes of each parameter are as follows: 1 l CF= , 2 l FQ= , 3 l QV= . The weight of 1 G , geocentric coordinates of swing arm are 1 XG , 1 YG ,the weight of bucket is 2 G , geocentric coordinates of bucket are 2 XG , 2 YG , the weight of basket is 3 G , geocentric coordinates of basket are 3 XG , 3 YG .The weight of complete machine is G , the centroid position parameter is rs, coordinates of swing arm and frame hinge point are XC , YC " ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0003085_intellisys.2017.8324247-Figure2-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0003085_intellisys.2017.8324247-Figure2-1.png", "caption": "Fig. 2. Representative body segment used in the inverse dynamics model.", "texts": [], "surrounding_texts": [ "In order to have an effective analysis on the effect of mono and biarticular muscles on the total work of each leg of robot during walking, we consider the bipedal robot model with springs like mono and biarticular muscles as shown in Fig. 1. The bipedal model shown in Fig. 1 consists of four leg limbs( with upper body, three joints (hip, knee and ankle joints) and four linear springs which are represented by the red dashed lines. The springs S1, S2, S4 correspond to the biarticular muscles: rectus femoris (RF), biceps femoris (BF) and gastrocnemius (GAS) in human legs respectively. Additionally, the spring S3 that correspond to a monoarticular muscle: tibialis anterior (TA) in human legs is also used in this model. The force generated in these springs is calculated as follows: (1) where and denote the stiffness coefficient and the displacement (lengthening or compression) of spring respectively, and in our case the intrinsic damping factor of springs is neglected. The point , represent the spring attachment. The center of mass of the system is located at the hip joint ( ). The limb mass are defined at the center of each limb. The two limbs and are considered in our model to remain parallel during the motion. ( ) represents the upper body attached to the hip joint. ) refers to the foot in contact with the horizontal ground or moving parallel to it. III. DYNAMIC MODELING OF A BIPEDAL ROBOT WITH" ] }, { "image_filename": "designv11_62_0002200_978-3-319-51222-8-Figure3.42-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0002200_978-3-319-51222-8-Figure3.42-1.png", "caption": "Fig. 3.42 Principal stresses in a cellular beam with fixed vertical edges", "texts": [ " In the plate, the situation obviously is different. Doubling the modulus E does not split the slope of the influence function for \u03c3yy (now in vertical direction) in half, but the decrease must be less, perhaps because the stiffness in the neighboring elements does not change. How rigid zones in a slab attract loads can be seen in Fig. 3.41 where the uniform load is mostly carried by the more rigid zones of the slab, h = 40 cm versus h = 20 cm for the inlets. In the extreme case that the inlets have zero stiffness, E = 0, we see results as in Fig. 3.42 where the stresses are forced to flow around the holes. Drop panels and column capitals cause the bending moments to migrate to the column capitals. The discontinuity in the slab thickness makes that the bending moments myy in a horizontal cross section perpendicular to the discontinuity jump, while the moments mxx are not affected (see Fig. 3.43). In a vertical cross section through the center, it would be the other way around. 200 3 Finite Elements In FE-analysis, any linear functional J (uh) = gT f is the scalar product of the vector g, the nodal values of the influence function, and the vector f of the equivalent nodal forces of the load" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0000394_09596518jsce982-Figure5-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0000394_09596518jsce982-Figure5-1.png", "caption": "Fig. 5 Case 2: excavator\u2019s work space intersects one dig side", "texts": [ " It can be easily proven that Sc also increases and becomes maximum as the point E approaches the leftmost allowable point Ea. This point lies on a horizontal line with y 5 yE, while its ordinate f depends upon: (a) the dimensions of the excavator\u2019s carriage, (b) the position of the pivot point on it, and (c) the orientation angle and the dimensions of EUZ. Given that the zone should not interfere, in any case with dig surface, it can be easily proved that the frontal position of the excavator (Fig. 5) provides the optimum solution and the corresponding point will be E2. Its exact position upon line segment expressing y 5 yE depends upon the current dimensions of the excavator\u2019s carriage and EUZ. The lateral position of the excavator is inferior since the useful circular section of the arm\u2013boom is restricted owing to the interference with EUZ. On line e\u2013e9, which passes from E2, there is more than one point that satisfies the above constraint. These points form a segment PQ along this line and possess the property of being pivot- and positioning points for the excavator that provide maximum surface for the circular segment (equation (5))" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0001039_detc2011-48755-Figure4-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0001039_detc2011-48755-Figure4-1.png", "caption": "Figure 4. Demonstration to roll/pitch axis model", "texts": [ " Actuator Dynamics The thrust generated by each propeller is modeled using the following first-order system model F k u s (15) where u, is the PWM input to the actuator, \u03c9 is the actuator bandwidth and k is a positive gain. These parameters were calculated and verified through experimental studies. A state variable, v, will be used to represent the actuator dynamics, which is defined as follows: v u s (16) 4.2. Roll and Pitch Model Assuming that rotations about the x and y axes are decoupled, the motion in roll/pitch axis can be modeled as shown in Figure 4. As illustrated in this figure, two propellers contribute to the motion in each axis. The thrust generated by each motor can be calculated from Eq. (15) and used as corresponding input. The rotation around the center of gravity is produced by the difference in the generated thrusts. The roll/pitch angle can be formulated using the following dynamics: (17) where roll PitchJ J J (18) are the rotational inertia of the device in roll and pitch axes. L is the distance between the propeller and the center of gravity, and \u0394F = F1\u2212 F2 (19) represents the difference between the forces generated by the propellers" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0002602_b978-0-12-811820-7.00005-7-Figure3.7-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0002602_b978-0-12-811820-7.00005-7-Figure3.7-1.png", "caption": "FIGURE 3.7 Hot-film anemometry convection diagram: natural convection alone is applied to the substrate side and bottom surfaces. Smooth surface convection is applied to the top substrate surface. Rough surface convection is applied to the activated deposition elements shown on the back half of the plate. Originally published by ASME in The International Journal of Advanced Manufacturing Technology 79.1-4 (2015): 307\u2013320.", "texts": [ " As the exponential decay functions cannot capture the asymmetry of the anemometry data (due to the asymmetry in the gas flows), an integrated convection comparison is used to ensure an equivalent rate of cooling that is applied in the simulations to the measured convection. The point-by-point error gives the average percent of the fit function deviation from the convection measurements. The integrated values and percent errors are presented in Table 3.5. For the three forced convection models, natural convection is applied to those regions unaffected by the gas flows \u2013 the bottom and sides of the substrate. The three convection regions for the hot-film method, natural, smooth surface, and rough surface, are depicted in Figure 3.7. On the sides and the bottom, hfree = 9 W/m2 K is applied. For regions on the surface of the mesh, either in the areas with no deposition or before the activation of deposition elements, the smooth surface convection is applied. Activated deposition element free surfaces (shown in Figure 3.5) are given the rough surface convection values. This methodology attempts to capture the evolving nature of the clad surface and its impact upon the rate of cooling on the part during cladding. The six convective bounds considered are shown in Figure 3" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0002827_cadiag.2017.8075643-Figure8-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0002827_cadiag.2017.8075643-Figure8-1.png", "caption": "Fig. 8. And example of the Autopilot animation appearance", "texts": [ " Compared to other simulator for small UAV, we demonstrate the possibility to apply advanced control techniques to reduce the trajectory tracking error with wind gust. The validity of the autopilot design concepts and control strategy is proven through different simulation results, including Penguin Be model. This autopilot is a useful tool for preparing UAV mission and path planning. It ensures mission achievement with success by overcoming wind gusts, to avoid trajectory trucking failure. This autopilot simulator includes also a visual module to better understanding of the output results and prevent experimental accident (figure 8). This autopilot project will include in the near future more developed dynamic equation and other modules allowing smooth waypoint switching, trajectory replanning and obstacle avoidance techniques [17], useful in environment including none fly zones. A module for target air or ground trucking allows using it in military, rescue or many other particular events. [1] Ross, P.E., 2014. Open-source drones for fun and profit. IEEE Spectrum,51(3), pp.54-59. [2] Stojcsics, D., 2014. Autonomous waypoint-based guidance methods for small size unmanned aerial vehicles" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0002289_j.robot.2017.04.002-Figure7-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0002289_j.robot.2017.04.002-Figure7-1.png", "caption": "Fig. 7. Tangential and radial components of the wheel force.", "texts": [ " 6 explains the notations used in dynamic modeling: h is the height of CoG from the ground plane, rA1, rA2, rA3 are vectors in the robot\u2019s coordinate system, those pointing to the wheel\u2019s gripping points from CoG, F 1, F 2, F 3 are wheel forces, FCoG and \u03c4 CoG are the force and torque vectors related to CoG. The axial component (free rolling direction, see vfree on Fig. 3) of the wheel force vector is zero. Therefore, the wheel forces can be split to load and drive components using radial and tangential unit vectors respectively, see Fig. 7. F i = Fi,drive \u00b7 ei,tangential + Fi,load \u00b7 ei,radial (8) where ei,radial and ei,tangential are the unit vectors in the wheel coordinate system. The sum of load forces is equal to the vertical force caused by gravity, therefore the force and torque equations of the robot body: FCoG = [FCoG,x FCoG,y mg ] = F 1 + F 2 + F 3 (9) \u03c4 CoG = [ 0 0 \u03c4CoG,z ] = rA1 \u00d7 F 1 + rA2 \u00d7 F 2 + rA3 \u00d7 F 3 (10) where m is the mass of the rigid body. In case of kiwi robot, drive and load directional force components are expressed as Eqs" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0002793_detc2017-68262-Figure6-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0002793_detc2017-68262-Figure6-1.png", "caption": "FIGURE 6: PRINT ORIENTATION AND GEOMETRY WITHOUT ADDED OFFSET. ARROWS INDICATE BUILD DIRECTION.", "texts": [ " Part (or print) orientation during the AM process plays a large part in the part functionality and the ability to use material efficiently. For simplicity, this analysis assumes print orientation will not affect part functionality, and that the part will be generated in the orientation that is most material efficient. The part orientation in the DED wire feed process will be in the direction that requires the least amount of support material. Using the reference orientation in Fig. 5, the most efficient orientation would be along the z-axis. In Fig. 6, the geometry of the part can be seen with the arrows representing the build direction. 3 Copyright \u00a9 2017 ASME Downloaded From: http://proceedings.asmedigitalcollection.asme.org/ on 11/09/2017 Terms of Use: http://www.asme.org/about-asme/terms-of-use The holes in the side have been filled and support material for the side protrusions and the top feature have been extended to meet the base. In addition to the support material, the DED process adds material to all features because of low deposition (or printing) accuracy", " The printing for the SLM process is based on the shortest dimension of the bounding box since there must be enough powder to fill a printing chamber. In polymer powder bed printing processes, a part would not require support material because the part is completely submerged and always surrounded by powder. This is not true for metal based powder bed methods. Support material is required for SLM. The print geometry without added offsets would be the same as DED because both require support material. The print geometry and build direction can be seen in Fig. 6. The SLM powder bed process can print to +/- 0.05 mm [11], so the additional material added is less than that of the wire feed method. The surface area of the print geometry is multiplied by 0.05 mm, which is added to the part volume to determine the print volume of the part. Layer thickness for DED wire feed process is equal to 50-200 \u00b5m, whereas for SLS powder bed it is equal to 20-100 \u00b5m [11]. Given these orientations and offsets, the print mass for the DED wire feed method would be 44.32 kg (10,005 cm3)" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0000850_icma.2013.6617998-Figure5-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0000850_icma.2013.6617998-Figure5-1.png", "caption": "Fig 5. (a) Top foil die (b) Bump foil die", "texts": [ "4 The wave profile of the foil journal bearing The manufacturing process consists of three primary steps. The first is to cut annealed oils to size. This is followed by forming either by rolling for the top foils or pressing against a corrugated tool steel die. The final step is to heat treat the formed foils to develop favorable spring properties and strength [9]. In ref 9, the top foil was formed by rolling, but the top foil was formed like the bump foil with pressing against an arc tool steel die. The rolling method mentioned in ref 9 is not suitable for forming multi-leaf bearings. Figure 5 shows the foils geometry as a starting point. For top foils dies, the semicircle has a diameter of 60 mm, a slot has a depth of 5 mm and a width of 0.5 mm as shown in figure 5(a). Each bump in the bump die has a width of 2.6 mm and a depth or height of 0.4 mm and separated by a flat segment with a length of 0.8 mm as shown in figure 5(b). (a) Figure 6 is photos of top foil dies and bump foil dies. The top foils were made from sheet approximately 0.18 mm thick. Thicker foils will result in higher stiffness but at the expense of lower compliance [11]. The bearing stiffness is largely dictated by the foil thickness and bump design. The test bearings\u2019 inside diameters were sized for 50 mm shafts. The forming process begins by cutting the foil to length (and width) using metal shears and/or wire (EDM). After being cut to the desired dimensions, Strip was placed between the upper and lower dies and one end of the strip was put into the slot, the upper and lower dies were fixed using pins" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0000674_j.proeng.2011.04.497-Figure2-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0000674_j.proeng.2011.04.497-Figure2-1.png", "caption": "Fig. 2. Definition of direction of critical plane for 2D model", "texts": [ " and p denote the shear stress and engineering plastic strain, b and m are material constants, Y is the plastic strain energy density on the critical plane, and D is the fatigue damage. For 1070 steel: 37.0b , 9.0m , MPa1500f , -3 0 mMJ5189D and MPa3000 . The material constants used in the plasticity model of 1070 steel were given in the literature [6]. With the detailed stress-strain response outputted from the FE analysis, the damage of points induced by a loading cycle can be obtained by equations (5). The fatigue initial crack life is D D N f 0 (5) where D denotes the damage increment of every cyclic loading, fN denotes the initiation fatigue crack life. As Fig. 2 showed, to simplify the explanation, the orientation of a plane is represented by an angle, , the angle made by the normal of the material plane and the x direction. For 2D model, after the angle of initiation point was known, the crack growth direction was got. Several cases were simulated to find the influence of kp0 ratio to RCF. Normalized quantities are used in the results. The length is normalized in terms of the half contact width, a , the stresses are normalized by the yield stress in shear, k , the shear strain is normalized by Gk , where G is the shear elasticity modulus" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0003710_icuas.2019.8797766-Figure1-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0003710_icuas.2019.8797766-Figure1-1.png", "caption": "Fig. 1. Body and Inertial Frames of a Quadcopter [10]", "texts": [ " (1) The state vector s can be decomposed into four vector functions: s = [ s1 s2 s3 s4 ] , (2) where each of them are defined as s1 = [ x y z ] s2 = [ vx vy vz ] s3 = [ q0 q1 q2 ] s4 = [ \u03c9x \u03c9y \u03c9z ] . (3) The inertial frame is set on the ground where gravity is pointing in the negative z-direction. The body frame is defined on the quadcopter where the x and y axes point in the direction of the arms, assuming symmetrical frame, and U.S. Government work not protected by U.S. copyright 513 the z-axis points up in the same direction as the motor axes. Figure 1 shows the inertial frame axes denoted by a1, a2, a3 and the body frame axes denoted by b1, b2, b3 [10]. A rotation matrix following the (3-2-1) sequence is applied to rotate the quadcopter from the inertial frame to the body frame [9]. The motor thrust vector in the body frame is T b = 4\u2211 i=1 Ti = k \u23a1\u23a3 0 0\u22114 i=1 \u03c9 2 i \u23a4\u23a6 , (4) where \u03c92 i is the squared angular rate of ith motor and k is the motor thrust coefficient. A simple drag vector for quadcopters is FD = \u23a1\u23a3\u2212kdvx \u2212kdvy \u2212kdvz \u23a4\u23a6 , (5) where kd is the constant friction coefficient (may or may not be unique)" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0001897_jnn.2012.5743-Figure1-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0001897_jnn.2012.5743-Figure1-1.png", "caption": "Fig. 1. The yellow, blue and small green balls represent paramagnetic beads, nonmagnetic beads and MR nanoparticles respectively. When an external magnetic field is applied, at the \u201cequator\u201d of a paramagnetic particle, the total magnetic field equals the dipolar field of the paramagnetic particle subtracted from the external field, which is a minimum according to Eq. (2), so the magnetic force drives nonmagnetic particles to this area.", "texts": [ " In contrast to non-magnetic particles, for paramagnetic particles, with PM > f , the potential energy is strictly negative, which causes the particle to move towards regions of maximum magnetic field, H r = Hmax . Hence, in mixed suspensions of two types of particles, the magnetic force drives non-magnetic particle towards the \u2018equator\u2019 of the paramagnetic particle, where the total magnetic field is equal to the dipolar field of the paramagnetic particle subtracted from the external field. In a similar way, the paramagnetic particles move towards the equator of the non-magnetic particles, where the total magnetic field equals the dipolar field of the non-magnetic particle added to the external field. Figure 1 shows a schematic diagram for this principle. 2082 J. Nanosci. Nanotechnol. 12, 2081\u20132088, 2012 Delivered by Ingenta to: Nanyang Technological University IP: 188.72.126.164 On: Sun, 05 Jun 2016 11:54:43 Copyright: American Scientific Publishers R E S E A R C H A R T IC L E In fact, the size ratio between the paramagnetic particles and the non-magnetic particles is a very important parameter to determine the configuration of the ring-structure. When the system consists of large paramagnetic particles and small non-magnetic particles, the larger paramagnetic particles will act as a core and will be surrounded by pairs of smaller non-magnetic particles around their equatorial areas" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0002527_aset.2017.7983670-Figure1-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0002527_aset.2017.7983670-Figure1-1.png", "caption": "Fig. 1. Modeling of the parallel inverted pendulum", "texts": [], "surrounding_texts": [ "\u2022 Ai \u2208 0. If the curve has endpoints, then the list mentioned above should also include the singularity that corresponds to the graph of the function g(x) = |x|. For a generic two-parameter family of smooth curves, the boundary \u03a3 of the set of local transitivity can have only one more singularity distinct from the singularities of the convex hull of an individual curve in general position [25], namely, the indicated singularity of the graph of the function y = |x| at the origin. It turns out that this list coincides exactly with the list of generic local singularities (up to diffeomorphisms of the plane) of the boundary of the zone of local transitivity for control systems on the plane whose indicatrices are smooth curves with endpoints. Moreover, it is the same as the list of singularities of convex hulls and boundaries of the zones of local transitivity in the case where the indicatrices are the images Controllability of non-linear systems 261 of a manifold U of arbitrary dimension under smooth maps which form families in general position. This being the case, it can be assumed that the manifold U has boundaries and corners. The techniques of the proofs of these (and some other) facts are discussed below. Let us turn to the main case of a three-dimensional manifold M . First we shall assume that the indicatrix is a C\u221e-smooth closed space curve. The list of generic local singularities of the convex hull of a curve in R3 up to diffeomorphisms of the ambient space was obtained in [29] and [6]. It includes the first six of the normal forms presented below in Theorem 1. In order to describe the singularities of the boundaries of the convex hulls of space curves with endpoints we introduce the following notation. A closed convex simple piecewise smooth curve \u03b3 that lies in the plane z = 1 of the space R3 = {(x, y, z)} is called a simple pivot curve if it consists of alternating segments of straight lines and strongly convex arcs and is C1-smooth at their common points. A curve \u03b3 is called a pivot curve with k corners if it contains the sides of k angles less than \u03c0 connected by convex pieces consisting of straight segments and arcs and is C1-smooth at all common endpoints except the vertices of the angles. Some sides may be common for adjacent angles. If we replace a straight-line side of an angle by a convex curve arc in this definition, then we obtain the definition of a pivot curve with k corners and a curved side. The conic surface formed by the segments that join the origin with the points of a pivot curve which is either simple (k = 0) or has k = 1, 2, . . . corners (possibly, with curved sides) will be called a k-cone. We note that the germs of these cones at the origin have functional invariants with respect to the action of diffeomorphisms. Indeed, the tangent vectors at the apex form the tangent cone and sweep out a convex curve on the sphere of directions. The tangent cone is subject only to linear transformations under the action of diffeomorphisms. Theorem 1 [27]. The germ at an arbitrary point of the convex hull boundary of a generic connected space curve with endpoints can be reduced by an appropriate diffeomorphism of R3 to the germ at the origin of the graph of one of the following functions z = fi(x, y), where i = 1, . . . , 7 (Figs. 2\u20135): 262 A.A. Davydov and V.M. Zakalyukin 1) f1 = 0 (the germ of a smooth surface); 2) f2(x, y) = |x| (edge); 3) f3(x, y) = { 0 for x 6 0, x2 for x > 0 (adjacency); 4) f4(x, y) = x2 for y 6 x, x > 0, y2 for y > 0, y > x, 0 for y 6 0, x 6 0 (bow); 5) f5(x, y) = 0 for y 6 0, x 6 0, x2 for y 6 \u2212x, x > 0, y2 for y > 0, y 6 \u2212x, 1 2 (x2 + y2)\u2212 y \u2212 x for x + y > 0 (stern); 6) f6(x, y) = minz\u2208R{z4 + xz2 + yz} (truncated swallowtail); 7) f7(x, y) = y2 + x for x > 0, y2 for y 6 0, x 6 0, (1\u2212 x)y2 for y > 0, x 6 0 (bend); or to the germ at the origin of a k-cone with k = 0, 1, 2. Remark 2. Generic singularities of the convex hull of a closed space curve (without endpoints) belong to the first six classes mentioned in this theorem. Remark 3. The smooth surface 1) may be strictly convex, ruled (Fig. 2, point B), or flat (Fig. 2, point A). The \u2018edge\u2019 2) arises at a generic point of the initial curve itself (Fig. 2, point C). The singularity 3) appears, in particular, at the points of adjacency of a ruled surface and a flat one (Fig. 2, point D). The germs 4) and 5) correspond to vertices E and F of the flat triangles which inevitably appear on the boundary of the convex hull of the curve. The cone apex coincides with the endpoint (Fig. 5, point K). The truncated swallowtail appears at a point of the curve such that the tangent line at this point also intersects the curve at some other point (Fig. 3, point J). Controllability of non-linear systems 263 The list of generic local singularities of the boundary of the transitivity zone for a control system on a three-dimensional manifold whose indicatrix is a closed space curve is given by the following theorem. Theorem 2. For a generic family of smooth curves rm : S1 \u2192 R3 depending on a three-dimensional parameter m = (x, y, z) \u2208 R3 the list of all local singularities on the boundary \u03a3 of the transitivity zone (considered up to diffeomorphisms of R3) consists of the germs 1)\u20137) mentioned in Theorem 1 and the germs 9), 12), 14) described in Theorem 3 below. For a space curve with endpoints the list of generic singularities of the transitivity zone is as follows. Theorem 3. For a generic three-parameter family of connected space curves with endpoints, the germ at any point of the boundary \u03a3 of the transitivity zone can be reduced by an appropriate diffeomorphism either to one of the germs 1)\u20137) of the boundaries of the convex hulls of generic curves in R3 listed in Theorem 1, or to the germ at the origin of the graph of one of the following functions : 8) f8(x, y) = 0 for x 6 0, y 6 0, y2 for y2 > x, y > 0, x for y 6 \u221a x , x > 0 (cut); 9) f9(x, y) = 0 for y 6 0, x + \u03b1y 6 0, (x + \u03b1y)2 1 + \u03b12 for y 6 \u03b1x, x + \u03b1y > 0, \u03b1 \u0338= 0, y2 for y > 0, x 6 0, x2 + y2 for x > 0, y > \u03b1x (adjacency of four surfaces, a cup); 10) f10(x, y) = \u2212y for y > 0, 0 for y 6 0, x > y, y(x\u2212 y)2 for y 6 0, x 6 y (helmet); 11) f11(x, y) = { |x| for y > 0, |x|+ y2 for y < 0 (butterfly); 264 A.A. Davydov and V.M. Zakalyukin or to the germ at the origin of the union of three surfaces with boundaries determined by the conditions 12) z = 0, y 6 0, y = x2, z 6 \u22124x2, z = \u2212t2, y = 1 4 z + tx, 0 6 t 2x 6 1 (adjacency with Whitney umbrella); or to the germ at the origin of the union of two surfaces with common boundary given by the conditions (book) 13) z = 0 for y, x > 0 or for x 6 0, y > 1 4 x2, z = 2t3 + xt2 for 3t2 + 2tx + y = 0, t > max { 0,\u2212x 2 } ; or to the germ at the apex of the lateral surface of a pyramid with n ridges for n = 3, 4, 5, that is, to the germ of the boundary of the domain given by the inequalities 14) x, y, z > 0 for n = 3, x, y, z > 0, z \u2212 x + y > 0 for n = 4, x, y, z > 0, z \u2212 x + y > 0, z \u2212 \u03b1x\u2212 \u03b2y > 0 for n = 5 (for n = 5 the normal form has two scalar invariants \u03b1 and \u03b2); or, finally, to the germ of a k-cone with k = 0, 1, 2. Controllability of non-linear systems 265 Remark 4. The boundaries a and b of the surfaces I and II of the germ 12) are given by the equations z = 0, y = 0 and y = x2, z = \u22124x2, respectively. These curves are tangent to each other at the origin, which is their unique common point. The surface III is a piece of a Whitney umbrella bounded by these curves and tangent to the surfaces I and II along a and b (Fig. 9). The boundary H(Ix) of the convex hull of a generic C\u221e-smooth surface Ix may have singularities which are only C1-smooth. Generic local singularities of H(Ix) [5] are either germs of a smooth surface or germs of a C1-smooth surface of \u2018adjacency\u2019, \u2018cup\u2019, or \u2018truncated swallowtail\u2019 type. For surfaces with boundary and corners, the following assertions hold. Theorem 4. The list of generic local singularities of the convex hull boundary of a smooth compact surface in R3 with a smooth boundary consists, up to diffeomorphisms, of the singularities 1)\u20137), 9), and 11). Theorem 5. Generic local singularities of convex hull boundaries for smooth compact surfaces in R3 with boundaries and corners are either the singularities 1)\u20137), 9), 10), or k-cones with arbitrary k. Theorem 6. The list of local singularities (up to diffeomorphisms of the parameter space) of the boundary of the set of local transitivity for a generic three-parameter family of smooth compact surfaces in R3 consists of the germs 1)\u20136), 9), 12), 14) of normal forms. Theorem 7. The list of local singularities (up to diffeomorphisms of the parameter space) of the boundary of the set of local transitivity for a three-parameter generic family of smooth compact surfaces with a smooth boundary consists of the singularities 1)\u201314). Theorem 8. The list of local singularities (up to diffeomorphisms of the parameter space) of the boundary of the zone of local transitivity for a generic three-parameter family of smooth compact surfaces with boundary and corners consists of all the singularities 1)\u201314) and k-cones with arbitrary k = 1, 2, . . . . Finally, let us consider the case where the indicatrices in three-dimensional space are the images of smooth maps of a given compact n-dimensional manifold Un of controls. 266 A.A. Davydov and V.M. Zakalyukin Theorem 9. The list of local singularities (up to diffeomorphisms of the target space) of the convex hull boundary of the image of a generic smooth map V : Un \u2192 R3 of a smooth n-dimensional compact manifold Un coincides with the list of local singularities of the convex hull boundary of a compact space curve for n = 1, or with the list of local singularities of the convex hull boundary of a smooth compact surface for n > 2, respectively. For n = 2 the list of local singularities of the convex hull boundary of a smooth closed surface also contains the \u2018adjacency with Whitney umbrella\u2019 singularity 12) described in Theorem 3. Theorem 10. The list of local singularities (up to diffeomorphisms of the parameter space) of the boundary of the zone of local transitivity for a generic threeparameter family of smooth maps V : Un \u2192 R3 of a smooth compact manifold Un of dimension n > 3 coincides with the list of germs 1)\u20136), 9), 12), and 14) of normal forms in Theorem 6. Theorem 11. The list of local singularities (up to diffeomorphisms of the target space) of the convex hull boundary of the image of a generic map V : Un \u2192 R3 of a smooth n-dimensional compact manifold Un with boundary and corners coincides with the list of local singularities of the convex hull boundary of a compact space curve with endpoints for n = 1, or with the list of local singularities of the convex hull boundary of a smooth compact surface with boundary and corners for n > 2, respectively. For n = 2 the list also contains the \u2018adjacency with Whitney umbrella\u2019 singularity 12) described in Theorem 3. Theorem 12. The list of local singularities (up to diffeomorphisms of the parameter space) of the boundary of the zone of local transitivity for a generic threeparameter family of smooth maps V : Un \u2192 R3 of a smooth n-dimensional compact manifold Un with boundary and corners coincides with the list of generic local singularities of the convex hull boundary of a smooth compact space curve with endpoints for n = 1, or with the list of generic local singularities of the convex hull boundary of a smooth compact surface with boundary and corners for n > 2, respectively. Remark 5. The lists described above show that generically the boundary of the convex hull and the boundary of the zone of local transitivity are Lipschitz. Apparently, this is true in any dimension. Remark 6. Moreover, in all cases we have the following statement, which provides an example of the principle [19] that \u2018good cases\u2019 dominate in many control-theory constructions, in contrast to Arnold\u2019s principle of \u2018fragility of the good\u2019, which is typical in singularity theory [30]. Consider a germ K of the generic transitivity zone whose base point corresponds to the origin being in a C1-smooth germ of the convex hull. Then either the boundary \u03a3 is smooth, or K is on the \u2018larger\u2019 side of the germ of the complement of \u03a3, which cannot be embedded in a half-space in any smooth local coordinates. Remark 7. The boundaries of the convex hulls of curves and surfaces contain ruled and flat pieces. These features are lost under a diffeomorphism: different arrangements of flat domains and straight lines on ruled pieces may correspond to diffeomorphic singularities. The statements of Theorems 1\u201312 contain only this rough Controllability of non-linear systems 267 classification up to diffeomorphisms. However, all possible cases of the affine structure of singularities can be found by the methods described below. 3.2. Legendre transformations and support hyperplanes. In this section we present the basic constructions and propositions ([25], [27]) which are involved in the proofs of the theorems formulated in the previous section and which may prove to be useful in the study of singularities in many other problems. First, let us note that if the indicatrix is given as the image of a map Rn \u2192 Rm with n > m, m = 2, 3, then, according to Thom\u2019s and Mather\u2019s classical results on the classification of singularities of generic maps (see, for instance, [17]), in the generic case only points of smoothness of the visible contour can appear on the boundary of the convex hull of the indicatrix. The only exception is the case of the \u2018Whitney umbrella\u2019 singularity R2 \u2192 R3. In all other cases it suffices to consider the indicatrices which are embedded curves and surfaces (possibly, with boundaries and corners), and their families. Surfaces with boundaries and corners, as well as curves with endpoints, can be treated uniformly as particular cases of stratified submanifolds embedded in R3. A collection J = {I1, . . . , Is} of closed embedded submanifolds I kj j (strata) of dimensions kj will be called a stratified submanifold if it contains a unique stratum of highest dimension, and any other stratum of lower dimension belongs entirely to some other stratum of higher dimension. For indicatrices in three-dimensional space we only deal with one-dimensional manifolds I1 with endpoints or with a smooth surface I2 in R3 on which there is a smooth curve I1 (a boundary) or two mutually transversal curves I1 1 and I1 2 (the sides of a corner) whose intersection point is treated as a separate stratum. However, the main constructions described below remain valid in the general multidimensional case. Thus, in the general case to a point q of a k-dimensional submanifold I we assign the set (diffeomorphic to the sphere S3\u2212k\u22121) of germs at q of all co-oriented planes tangent to the submanifold at this point. All such germs form a smooth Legendrian submanifold LI \u2248 I\u00d7S3\u2212k\u22121 in the space ST \u2217R3 of co-oriented contact elements. Forgetting the base point of the germ, we obtain the projection \u03c0\u2217 on the space of all co-oriented planes of this submanifold, which is a Legendre map of the submanifold LI to the dual space R\u03023. The image (wavefront) I\u0302 = \u03c0\u2217(LI) of this projection is called the Legendre transform of the initial submanifold I, or the dual surface to I. Denote by I\u0302A the germ of the dual surface to I at the set of all planes tangent to I at the point A. The Legendre transform for a stratified submanifold J is defined as the collection J\u0302 of the Legendre transforms \u03c0\u2217(LIj ) of all the strata Ij . A manifold Jc with boundary or corners is regarded as a subset of the stratified submanifold determined by the corresponding inequalities. Therefore, the corresponding Legendrian submanifold LJc is taken to be the subset of tangent elements at the points of Jc, and the Legendre transform J\u0302c is formed by the corresponding subsets of the dual surfaces of the strata of J . The space of all co-oriented planes in R\u03023 is fibred over the sphere of unit normals (with one-dimensional fibre of parallel planes). The boundary H(\u0393) of the convex hull of the compact subset \u0393 is determined by the set P (\u0393) of support planes of \u0393. Namely, for any unit normal we choose, among all the parallel planes intersecting 268 A.A. Davydov and V.M. Zakalyukin this hull, a plane P that corresponds to the maximum value of the coordinate whose gradient is directed along this normal. We note that a plane P \u2208 R\u03023 is a support plane of J if and only if it is tangent to some of the strata Ij at one or more points, or in other words, belongs to the image \u03c0\u2217LJ . For any support plane P we denote by SP = \u03b3\u22121(P ) the set of points at which this plane is tangent to \u0393. We call it the base of P and its points the base points. The number \u00b5P of distinct points of the base SP will be called its multiplicity. Thom\u2019s transversality theorem [17] for the space of multi-jets of maps or families of maps imposes some restrictions on the possible values of the multiplicity of support planes, types of Legendre maps \u03c0\u2217, and so on. Moreover, the condition that the origin O belongs to the support plane under consideration also imposes some constraints on the possible types of generic singularities of the Legendre map. For instance, for a generic surface \u0393 the multi-germs \u03c0\u2217 |SP are Legendre stable. Each germ is Legendre equivalent either to a germ with a singularity of A1 type (in a neighbourhood of such a point the surface is the graph of a Morse function and its quadratic form is negative definite) or to a germ with a singularity of A3 type (which corresponds to the Legendre map of the graph of the function h = x4 + (y\u2212 x2)2). The number of points in each base SP is at most three for singularities of A1 type and at most one for a singularity of A3 type. Thus, it is feasible to list all possible singularities of dual surfaces and to classify singularities of the convex hulls and the transitivity zones using the following advantageous properties of the Legendre transform. 1. The Legendre transformation of a generic hypersurface repeated twice is the identity map. Indeed, a Legendrian submanifold has two Legendre projections: \u03c0\u2217 and the standard projection \u03c0 : ST \u2217R3 \u2192 R of the fibre bundle. A smooth Legendrian submanifold LI is uniquely determined by its wavefront I\u0302 provided that the regular points of \u03c0\u2217 are dense in LI [31]. Hence the repeated Legendre transformation of the hypersurface I\u0302 yields the same Legendrian submanifold LI and the projection \u03c0, whose image \u03c0(LI) coincides with I. 2. The Legendrian submanifold LI2 of a smooth surface I2 has a regular projection on the surface itself: the tangent plane at each point is unique. In a neighbourhood of the convex hull of the base SP of a support plane P tangent to I2, the convex hull boundary H(I) is determined by the support planes which are close to P . The germ of the dual surface (I\u03022, P ), where P is the plane tangent to the generic surface I2 at a point q, is smooth provided that q is not a parabolic point. At a generic point of a parabolic line \u03b4, this germ is diffeomorphic to a semicubic cylinder (that is, to the bifurcation diagram of an A2 singularity). At isolated points of the line \u03b4, the Legendre transform has a singularity of A3 type, and the germ of (I\u03022, P ) is diffeomorphic to a swallowtail. We note that the interior points of A2 type cannot belong to the convex hull, since the tangent plane divides the surface in a neighbourhood of the point q. 3. The Legendrian submanifold LI1 for a smooth curve I1 is swept out by the circles Sq containing the tangent elements to I1 at its points q. The dual surface I\u03021 is ruled: in the affine chart the circles Sm correspond to straight lines. The set of support planes P tangent to the curve at a fixed point q forms a connected closed arc Eq on Sq. A support plane P passing through q \u2208 I1 can be rotated about the Controllability of non-linear systems 269 tangent to the curve until it touches the curve at some other point or becomes an osculating plane. The germ of the convex hull near the point q of a one-dimensional stratum I1 is determined by the germ Su(I1) on the arc Eq of support planes passing through q. On a generic curve some isolated points may have simple flattening (in a neighbourhood of such a point in the canonical Frenet frame the curve has the parametric form q1 = t, q2 = t2 + \u00b7 \u00b7 \u00b7 , q3 = t4 + \u00b7 \u00b7 \u00b7 ), while at all other regular points the germ of the curve has the canonical form q1 = t, q2 = t2 + \u00b7 \u00b7 \u00b7 , q3 = t3 + \u00b7 \u00b7 \u00b7 . In a neighbourhood of the osculating plane, which is given in both cases by the equation q3 = 0, the dual surface has a singularity of A3 type (swallowtail) in the flattening case and a singularity of A2 type (semicubic cylinder) in the regular case. The osculating plane at a regular interior point (one which is not an endpoint) cannot be a support plane: the curve is located on both sides of the plane (however, at an endpoint the osculating plane can be a support plane). 4. Any point q \u2208 R3 corresponds in the dual space to the plane q\u0302 consisting of all planes in R3 passing through q. The germ of the convex hull in a neighbourhood of a point q which is a zero-dimensional stratum of J (that is, a corner vertex or a curve endpoint) is determined by the germ of Su(J\u0302) on the convex domain Uq \u2282 q\u0302 consisting of all support planes passing through q. 5. The Legendre transforms I\u03021 and I\u03022 of two strata I1 \u2282 I2 are tangent along their intersection I\u03022 \u2229 I\u03021, which consists of the tangent planes to the stratum I\u03021 at the points of I\u03022. Indeed, the surfaces I\u0302i = \u03c0\u2217(LIi ), i = 1, 2, intersect along the set of common tangent planes at the points of the smaller stratum I1. Due to the involution property of the Legendre transformation, the planes tangent to \u03c0\u2217(LI) are points of the submanifold I itself. Therefore, the planes tangent to I\u03021 and I\u03022 at their intersection points correspond to the common points in I1 \u2229 I2 and hence coincide, as required. If the quadratic form of the surface I\u03022 is non-degenerate in the directions tangent to I2 and transversal to I1, then the tangency of I\u03021 and I\u03022 is of the first order. 6. A point Q \u2208 R3 corresponds in the dual space to the plane Q\u0302 consisting of all planes in R3 passing through Q. A point Q belongs to the convex hull boundary of a submanifold B if and only if the plane Q\u0302 is negatively supporting for the subset Su(B) of support planes of B: the open negative half-space of Q\u0302 does not contain points of Su(B), but the plane itself contains such points. In particular, the point O belongs to a convex surface X \u2282 R3 if and only if the dual plane O\u0302 is tangent to the dual surface X\u0302. 7. If O belongs to the convex hull of the base of the support plane for some generic three-parameter family of submanifolds B of dimension k, then the singularities of the Legendre transform at each base point are generic singularities of the Legendre transforms of k-dimensional submanifolds. Indeed, the dimension of the convex hull of the base consisting of l points is less than or equal to l\u2212 1. In order for the origin O to belong to the convex hull of the base, at least 3\u2212 l + 1 conditions must be satisfied. For l points to fall on the osculating plane, l conditions must hold. A degenerate point appears on the base if at least one condition holds. Thus, in this case more than three independent conditions must be satisfied, which is generically impossible for three parameters. 270 A.A. Davydov and V.M. Zakalyukin Now a classification of generic singularities of the boundary of the transitivity zone is obtained by listing all possible positions of the origin in the base of a support plane, describing the corresponding families of dual surfaces and their supporting subsets, and, finally, applying the following results in singularity theory. A diffeomorphism \u03a6: R\u03023 \u2192 R\u03023 mapping the plane O\u0302 onto itself and preserving its positive half-space is called admissible. If \u03a6 maps the support part Su(J1) of the Legendre transform of a stratified manifold J1 to the Legendre transform Su(J2) of another stratified manifold, then, clearly, the origin O belongs to the convex hull boundary H(J1) if and only if it belongs to H(J2). Thus, a family of admissible diffeomorphisms \u03a6m, fibred over a diffeomorphism of the parameter space, acts on the parameter-dependent families of surfaces I\u0302j and maps the respective transitivity zones to each other. Reducing the equations of the dual strata I\u0302j to normal forms via admissible diffeomorphisms, we obtain normal forms of the transitivity zone. For this purpose, consider the action of the diffeomorphism \u03a6m as a contact transformation on the product of the equations of all the components [32]. Denote by Pi(x), i = 0, . . . , n, the polynomials of degree ki in x \u2208 R of the form P0 = xk0 + \u2211k0\u22122 j=0 xja0j , Pi = xki + \u2211ki\u22121 j=0 xjaij , i = 1, . . . , n. Denote by a \u2208 RN the vector of coefficients of all these polynomials. We note that P0 is the standard miniversal deformation of the singularity Ak0\u22121, and Pi for i = 1, . . . , n are versal deformations of the singularities Aki\u22121. Lemma. Let gi(x, b), i = 0, . . . , n, be functions of x with parameters b \u2208 RN such that the values of gi(x, 0) and the values of their derivatives up to order ki \u2212 1 vanish at the origin (the function gi(x, 0) has singularity Aki\u22121 at the origin), and let g(x, b) = \u220fn i=0 gi(x, b). Then there is a contact equivalence consisting of a diffeomorphism (x, b) 7\u2192 ( X(x, b), B(b) ) and a non-zero function \u03d5(x, b) which reduces the function g to the form g(x, b) = \u03d5(x, b) \u220fn i=0 Pi(X, B). The assertion of the lemma is equivalent to the versality of the map described below with respect to a special group of equivalences. Consider the map G : Rk \u2192 Rn+1, G : x 7\u2192 ( g0(x), . . . , gn(x) ) , and in the target space consider the collection Y = \u22c3 {yi = 0} of coordinate hyperplanes. Let D be the group of diffeomor- phisms of Rn+1 which preserve Y ; that is, they have the form \u03b8 : (y0, . . . , yn) 7\u2192( h0(y)y0, . . . , hn(y)yn ) . Two maps G1 and G2 are called Y -contact equivalent if for some family of diffeomorphisms \u03b8x in the group D that depend on a parameter x we have G2(x) = \u03b8x \u25e6G1(X(x)) for some change of variables x 7\u2192 X(x). For a family Gm : Rk \u00d7 RN \u2192 Rn+1 of maps G depending on parameters a in RN , a parameter-dependent contact equivalence is defined in a natural way: the parameters a are replaced via a diffeomorphism by parameters b; a diffeomorphism of the form x 7\u2192 X(x, a) acts on the variables x; a family of diffeomorphisms \u03b8x,a : Rn+1 \u2192 Rn+1 acts on the target space Gm. The notions of versal and infinitesimally versal deformations of a map G are straightforward. Obviously, for this group D of transformations an analogue of the versality theorem holds, since the group is geometrical in the sense of J. Damon. We note that the map P\u0303 : (x, a) 7\u2192 (P0, . . . , Pn) is an infinitesimally versal deformation of the map x 7\u2192 (xk0 0 , . . . , xkn n ). The versality of the map P\u0303 implies the assertion of the lemma: under an equivalence in the group D each component Controllability of non-linear systems 271 is multiplied by a non-zero factor; hence the product is multiplied by a non-zero factor." ] }, { "image_filename": "designv11_62_0003814_codit.2019.8820321-Figure1-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0003814_codit.2019.8820321-Figure1-1.png", "caption": "Fig. 1. Qualitative schematic view of flow vectors in ground effect.", "texts": [ " Detecting disturbances and reacting to them is essential for completing tasks safely and effectively. When dealing with certain classical control system configurations, very large disturbances cannot be rejected in short times. A large part of the solutions is based on robustifying the control algorithms. The caveat of these approaches is that the computational burden is high and sometimes extremely complicated for real-time implementation. The goal of this paper is to present a control scheme for rejecting disturbances that are induced on the quadrotor by the presence of a ground plane, see Figure 1 for a qualitative flow schematic. The general control structure followed by this work is a cascade control approach, where a reliable lowlevel controller is present as the inner loop, and the proposed control scheme is running as an outer loop. The proposal for The authors are with the Department of Computer Science at the Instituto Nacional de Astrof\u0131\u0301sica, O\u0301ptica y Electro\u0301nica (INAOE), 72840 Puebla, Mexico. The third author is also an Honorary Senior Research Fellow at the University of Bristol in the UK", " Figure 7 shows the real drone and its evolution over a test. Results from a test of the multi-controller are shown in Figure 8. The transitions between u1 and u2 can be appreciated in this plot. There is a tendency of the positive pulses to be short in time, whereas the negative pulses tend to be wide in time. This behavior can be explained by considering the ground effect. The presence of the ground below the rotor-craft produces a cushion of air that makes easier to move away from the ground (see Figure 1), so positive command pulses are likely to generate more movement than negative command pulses in ground effect. However, this tendency is blurred sometimes because the inner stabilization of the drone is also affecting the overall behavior. This result demonstrates the ability of the multi-controller structure to adapt to the situation. The performance of the multi-controller structure compared the classical controller is shown in Figure 9. We have plotted the response of the vertical position of five runs" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0002207_j.compfluid.2017.02.022-Figure2-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0002207_j.compfluid.2017.02.022-Figure2-1.png", "caption": "Fig. 2. Sketch of experimental setup.", "texts": [ "65 m/s in the horizontal orientation by adjusting the longating length of rubber bands in an underwater ejection deice, which converted potential of rubber bands into kinetic enrgy. A high speed camera (JAC GC-PX100) was placed on one side f the flume to record the entire gliding process at 200 frames per econd. A breakwater was arranged near the top of the hands of wimmer to reduce the impact of disturbance caused by the ejecion device on the swimmer\u2019s glide motion. Details of the experi- ental setup are shown in Fig. 2 . The CAD model of the mannequin was generated using hinoceros, and then a physical scaled model was built by 3D rinter. The mass center of the 3D scaled model was positioned t 0.52 L from the finger tip, as shown in Fig. 3 . In the initial poition, h indicates the initial distance from the gravity center to he water surface. After gliding for some time ( t ), x indicates the isplacement of the model\u2019s gravity center in the horizontal direcion, z indicates the displacement of gravity center in the vertical irection, and \u03b1 is the pitching angle of the model" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0001099_amr.853.625-Figure4-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0001099_amr.853.625-Figure4-1.png", "caption": "Fig. 4 Scraping tool", "texts": [ " It generates far less voice than oil pressure system on standby and consumes less power. So the cost of production is effectively decreased. According to air hydraulic principle it meets the demand of high speed and high output power. Besides, it is easy to adjust output power. Get the desired operating pressure as long as you adjust air pressure boosting device. The unit with solid structure and simple operation can work effectively. Design of scraping tool. Self-designed scraping tool is shown in Fig. 4. By imitating the scraping movements that workers do there is a bent angle of 120 degrees at two thirds of tool holder in length to make sure the tool bit and the workpiece are at an angle. So that when the workpiece is moved horizontally on the table a scraping action is composed on the workpiece as long as the scraping blade moves up and down. We can obtain scraping effects with different levels of precision through adjusting tool bit width and feed speed of operating floor. Since scraping tool bit gets in touch with the workpiece, there are some friction and extrusion pressure in scraping process" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0002302_b978-0-12-803581-8.10351-0-Figure25-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0002302_b978-0-12-803581-8.10351-0-Figure25-1.png", "caption": "Fig. 25 Intermediate bulkhead.", "texts": [ " The results suggested that the solid version would perform the best and it was determined that forward footwell access was not critical, as access from the suspension box area on top of the footwell was available. The intermediate bulkhead design is constructed in the U-shaped configuration that matches the chassis contour to add some torsional rigidity; however, the principal function is to carry the steering rack loads, the suspension rocker loads and a portion of the cross chassis loads from the upper suspension arms. Fig. 25 shows the design which incorporates a 7.5 degree indent, an arc for the steering shaft to pass through and an enlarged bonding flange in the front to support the fasteners for the rocker pivot posts. The 7.5 degree indent is for the angle of the steering shaft sits at to have the steering wheel at its maximum height. The arc across the bottom of the bulkhead allows for the steering shaft to pass below the bulkhead and connect to the rack. The arc allows for the steering shaft and the rack to be removed from the bulkhead as a single component rather than disassembling the two parts" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0002791_b978-0-12-812138-2.00003-9-Figure3.14-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0002791_b978-0-12-812138-2.00003-9-Figure3.14-1.png", "caption": "FIGURE 3.14", "texts": [ " For example, for a two-pole induction motor fed by currents of 60 Hz, the synchronous speed is 3600 r/min. To analyze the operating characteristic and performance of an induction motor, we usually use an equivalent circuit based on voltage equations that describe the behavior of the motor. In this section we will start with the steady-state equivalent circuit, and in Chapter 4, we will introduce the equivalent circuit that is useful for characteristics in the transient state as well as in the steady state. Consider a two-pole, three-phase, wye-connected symmetrical induction motor as shown in Fig. 3.14. It is assumed that the rotor windings, which may be a wound or a squirrelcage type, are approximated as three-phase windings. Both the stator and rotor windings are distributed windings, but Fig. 3.14 shows only the centered coils on the axis of each phase. We will initially explore the voltage equations for the stator windings. mmf Distribution by phase as current in the four-pole motor. The voltage equation for the stator windings consists of the voltage drop of the winding resistance Rs and the induced voltage proportional to the rate of change over time of the stator flux linkage \u03bbs. Thus the voltage equation for the stator windings is written by vs 5Rsis 1 d\u03bbs dt (3.6) where Rs is the resistance of the stator winding and \u03bbs is the flux linkage of the stator winding" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0003541_978-3-030-20216-3_7-Figure2-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0003541_978-3-030-20216-3_7-Figure2-1.png", "caption": "Fig. 2. The schematic diagram shows the ZVS circuit connected to induction coil. The solder was kept inside the aluminium tube. CNC machine allows the granite bed to move in 3D. Support beams holds the workpiece and vibration motor attached to it.", "texts": [ " As the solder temperature exceed the setpoint, LabVIEW send the signal to digital module to send the 0 (Low) volt to MOSFET based control circuit, lead to turn off the IH circuit (see Fig. 1). The control circuit receives the high/low signal and boost it using IR2110 chip and send signals to the gate of MOSFET (IRFP250 N) to turn the IH on/off. The solid solder material is loaded inside the aluminium tube upto a height of 55 mm. The aluminum tube has ID of 10 mm and OD of 11 mm. A brass nozzle of 400-micron diameter is attached at the bottom of the aluminium tube and kept inside the induction coil suspended by a beam support as shown in Fig. 2. A bar-type DC vibration motor is tightly attached to aluminium tube through a beam support. At a supply of 2.5 V, the vibration motor have a fixed 129.9 rotations, per second (RPS) measured by analysing a high-speed video of vibration motor using MATLAB. The IH and nozzle was kept stationary, and the granite bed was attached with Control numeric control (CNC) machine to move it in 3D. We have used a lead-screw based CNC machine with three stepper motor for x, y and z-axis, controlled with the stepper motor driver at 200 microstepping" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0002476_yac.2017.7967544-Figure2-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0002476_yac.2017.7967544-Figure2-1.png", "caption": "Fig. 2. Illustration example of square formation.", "texts": [ " The assumption of the existence of the virtual leader is reasonable, because the trajectory of the virtual leader we need to track can be regarded as that of a specified real UAV with no external control inputs whose dynamics are described by 0 0 0 ( )= ( ) ( )=g( ), x t v t v t t (3) where g( )t is a continuously differentiable function. Therefore, the time-varying formation control problem is transformed into time-varying formation tracking problem. To make the definition more comprehensible, we take a time-invariant formation of square as an example, which is shown in Fig.2. Consider a multi-UAV system containing four UAVs, which means {1,2,3,4}I = . Since the formation is invariant, ( )ih t is equal to ih . Choose the formation center as the virtual leader. If equation (2) in Definition 2.1 is satisfied, and = , ,i jh h i j I\u2200 \u2208 as showed in Fig.2, the square formation is achieved. Let ( ) [ ( ), ( )]T T T i i it x t v t\u03be = , and 0 0 0( )=[ ( ), ( )]T T Tt x t v t\u03be . It is worth noting that ( )ih t represents the relative offset vector of [ ( ), ( )]T i ix t v t with respect to [ (0), (0)]T i ix v if Definition 2.1 is satisfied. Therefore, we can achieve any desired formation by choosing the appropriate ( )ih t . Next, the paper will focus on the time-varying formation tracking control law design based on NFTSM. In this section, we will firstly present the expression of the NFTSM, after which the control law based on NFTSM is proposed" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0003668_978-3-030-24314-2_70-Figure5-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0003668_978-3-030-24314-2_70-Figure5-1.png", "caption": "Fig. 5. Proposed Air jet cooling system [13]", "texts": [ " The study concluded that machinability of AM components is reduced and hence author recommended to use different cutting parameters for AM manufactured components. Heat accumulation is the main concern/issue on WAAM process and this significantly elevates the work piece temperature and this in turn affects the product. Montevecchi et al. [13], adopted a jet impingement procedure which avoids the excessive accumulation of heat in the work piece and enhance the convective heat transfer from work piece to the surroundings. A proposed air jet cooling system is shown in the below Fig. 5 and the parts of proposed system are 1- Torch, 2- Work piece Deposited material, 3- Air hose 4- Work piece substrate. Simulation studies were conducted in this study with Finite Element analysis to evaluate the effectiveness and the results of study indicated the proposed jet impingement significantly reduced the heat accumulation. Heat accumulation is a core issue in the wire arc additive manufacturing and influences the mechanical properties of the fabricated component. Binato et al. [14] conducted experiments Fig" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0000554_amr.175-176.490-Figure1-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0000554_amr.175-176.490-Figure1-1.png", "caption": "Fig. 1 Warp\u2019s Elongation and Frictional Moving Deformation at the Front and the Back of the Shed while Beating Up", "texts": [ " Therefore, the tension of warp at the front of the shed while beating up should be: 1 210 101101 cos\u03b1 CSTP lCTTTTT b \u2212+ +\u2206+=\u2033+\u2032+= (3) where T1 is the tension of warp at the front of the shed(cN), T0 is the warp static tension of warp(cN), P is the beating force on each warp(cN), Tb0 is the fabric tension at the instant of starting beating (cN, calculated according to each warp), \u03b11 is the half angle of front shed(\u00ba) and C2 is the stiffness coefficient of fabric(cN /mm). Let the frictional movement volume of warp in the heald eye to be S, and the compensation rate of warp elongation at the back of the shed that caused by the downward swing of back rest to be S2, as in Fig. 1, the absolute elongation \u25b3l1\u00b4and \u25b3l2\u00b4of warp at the front and the back of the shed while beating up should be: Sll \u2212\u2206=\u2032\u2206 11 (4) 222 SSll \u2212+\u2206=\u2032\u2206 (5) If the friction of warp stop stand on the warp is not considered, the warp tension at the front and the back of the shed while beating up and the tension difference between them should be: ( ) 1 210 10 1 210 101 coscos \u03b1\u03b1 CSTP CSlT CSTP ClTT bb \u2212+ +\u00d7\u2212\u2206+= \u2212+ +\u00d7\u2032\u2206+= (6) ( ) CSSlTClTT \u00d7\u2212+\u2206+=\u00d7\u2032\u2206+= 220202 (7) ( ) 1 210 22121 cos 2 \u03b1 CSTP CSSllTTT b \u2212+ +\u00d7+\u2212\u2206\u2212\u2206=\u2212=\u2206 (8) Where T2 is the tension of warp at the back of the shed(cN), \u25b3T is the tension difference between warp at the front and the back of the shed(cN), C is the stiffness coefficient of warp(cN/mm)" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0002887_acs.2838-Figure7-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0002887_acs.2838-Figure7-1.png", "caption": "FIGURE 7 A, Diagram of elastic beam; B, Equivalent model of elastic beam40", "texts": [ " It can be seen that the saturation nonlinearity is undertaken using the adaptive method. In addition, since the amplitude of the control is reasonable, it can be concluded that the proposed control signal is implementable using real physical actuators. Figure 6 illustrates the time response of the update parameter \u03a6\u0302. Clearly, this parameter is bounded and it approaches to a fixed value. This means that the update parameter does not make the system to be internally destabilized. In this example, a smooth and discontinuous oscillator, which is composed of an elastic beam (see Figure 7A) and has found useful applications in engineering disciplines,40 is taken into account. As depicted in Figure 7B, this oscillator is equivalent to a snap-through truss system including a mass, a pair of inclined elastic springs, and a rigid support. The nonlinear vibration equation of this nonautonomous system with a viscous damping and an external harmonic excitation is stated as follows40: \u23a7\u23aa\u23a8\u23aa\u23a9 x\u0307 = y y\u0307 = \u2212x ( 1 \u2212 1\u221a x2+a2 ) \u2212 2\u03b6\ud835\udc66 + f0 cos \u03c9\ud835\udc61, (39) where a= 0.5, \u03b6 = 0.01 \u221a 2, f0 = 0.8, and \u03c9 = 0.75 \u221a 2 are system parameters. Figures 8 and 9 show the chaotic attractor and oscillatory vibrations of the nonautonomous elastic beam model (39)" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0003657_iciaict.2019.8784844-Figure1-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0003657_iciaict.2019.8784844-Figure1-1.png", "caption": "Fig. 1. Hexacopter model with the frames", "texts": [ " INTRODUCTION Unmanned Aerial Vehicles (UAVs) are becoming more prominent nowadays due to their potential applications in high-quality photography [1], agriculture [2], monitoring [3], rescue [4], defense [5] and transportation [6]. UAVs have different configurations and sizes based on their applications. The most common type of UAVs is the rotating wing. Rotating wing UAVs or multi-rotors are usually named based on their rotors number. For instance, the quadcopter has four rotors while hexacopter has six rotors as shown in Fig.1. Quadcopter modeling and control have often been studied in the literature [7]\u2013[10] while hexacopter modeling and control are still not well-established. Controlling hexacopter is a nontrivial task since it has non-linear dynamics, over-actuated, prone to disturbances and more affected by the aerodynamic forces than the quadcopter. On the other hand, it offers promising features such as its capability to handle higher payloads than the quadcopter and its ability to tolerate actuator faults. Hence, it might be utilized in more applications than a quadcopter" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0003983_00423114.2019.1662925-Figure3-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0003983_00423114.2019.1662925-Figure3-1.png", "caption": "Figure 3. Conventional railway vehicle.", "texts": [ " The high speed rolling stock TSI (2008) was superseded by the Loc & Pas TSI (2014) [15], which has not defined this criterion any more. This section deals with the experimental set up. First we describe the car architecture. After that an overview of the failure scenarios will be given. The analysed vehicle is a commuter train with four cars (two single-deck motor cars and two double-deck trailer cars) and a maximum speed of 160 km/h. It is a conventional railway vehicle, which typically consists of four wheel-sets, two bogies (leading and trailing bogie) and a car body, as shown in Figure 3. The standard set up for the bogie is one yaw damper at each side. Before a railway vehicle can be used in passenger mode, it has to pass on-track test runs for vehicle acceptance for running behaviour. These tests usually include rides with demounted anti-yaw dampers. At the on-track tests for Rhein Ruhr Express (RRX) additional rides with a reduced yaw damper and modified bushing elements have been Figure 4. Bogie with demounted yaw damper. executed. Test rides took place in the Deutsche Bahn (DB) net in Germany" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0003542_icps.2019.8733379-Figure4-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0003542_icps.2019.8733379-Figure4-1.png", "caption": "Figure 4: implementation diagram for the Three phase short-circuit test of the SG", "texts": [], "surrounding_texts": [ "A. SG parameter identification based on sudden short-circuit\ntest\nThis test is done by applying a sudden three phase short-circuit on the terminals of the SG while running in a well-known permanent state. The resulting current in one phase is expressed as [12] :\n \n \n0 0\n0 0\n1 1 1 1 1 ( ) cos\n1 1 cos\n2\nd d\na\nt t T T\na c n\nd d d d d\nt\nTc\nd d\ni t V e e t x x x x x\nV e\nx x\n \n\n \n \n\n \n \n \n(10)\nThis current is composed of different terms identified here as:\n A damped pseudo-periodic term of angular frequency\n\ud835\udf14, with an initial amplitude of \ud835\udc49\ud835\udc500\n\ud835\udc65\ud835\udc51\u2032\u2032 during the sub-\ntransient phase, and with a final amplitude of \ud835\udc49\ud835\udc500\n\ud835\udc65\ud835\udc51\n A damped aperiodic term whose initial amplitude depends on the initial position, with respect to the inducing field, and the phase considered\n It may also contain another damped pseudo-periodic term pulsed at 2\ud835\udf14, which have been neglected here due to its usually low amplitude.\nEquation (10) can then be rewritten as:\n \n \n' '' ' ''\n0\n0\ncos\ncos\nd d a\nt t t\nT T T\ncc d d n ar\nAC n DC\ni t I I e I e t I e\nI t t i t\n \n \n \n \n(11)\nWhere some components are defined as in IEEE Std 115 (2009) for the purpose of this procedure:\n Upper and lower envelopes of the short-circuit current which are obtained by framing the term cos(\u03c9\ud835\udc5b\ud835\udc61 + \ud835\udf03) between -1 and +1.\n The constant alternative term, with is obtained at steady state \ud835\udc61 \u2192 \u221e\n A direct term DCi t to be extracted from the short-\ncircuit current.\nIn a practical analysis, after the upper and lower envelope have been obtained, the fundamental and direct components of the short-circuit current are calculated as follows:\n \n0\n_ ( ) _ ( )\n2\n1 1 1 1 1 d d\nAC\nt t\nT T\nc\nd d d d d\nu env t l env t I t\nV e e x x x x x\n \n \n \n \n(12)\n \n 0 0\n_ ( ) _ ( )\n2\n1 1 cos\n2 a a\nDC\ntt\nT Tc ar\nd q\nu env t l env t i t\nV e I e\nx x \n\n \n \n(13)\nd-axis parameters can then be extracted from the obtained set of data using least-square fitting algorithm. This process is summarized the flowchart in Figure 2.\nB. SG parameter identification based on load rejection test\nThe load rejection test consists in suddenly suppressing a specific load initially connected to the generator, with field voltage kept constant. Typically, the disconnected load may be purely inductive or capacitive. The SG parameters can then be extracted from the measured terminal voltage and field current data [2, 10]. The simplified approach illustrated in this paper relies on the hypothesis that the adjustment required due to the generator saturation level can be made after its parameter have been extracted, during the validation process.\nFrom the measured terminal voltages, the d-q voltages are\nderived using de well known Park transformation:\n0\n2 2 cos( ) cos( ) cos( )\n3 3\n2 2 2 sin( ) sin( ) sin( ) , 3 3 3\n1 1 1 2 2 2\nd a\nq b\nc\nV V\nP V P V\nV V\n \n \n \n \n \n(14)\nThe q-axis terminal voltage after a load rejection event is\nexpress as [14] :\n' '' 0 0' '' '( ) ( ) ( )d d\nd d d\nt t T T\nq d d d d dv t E x I I x x e I x x e (15)\nConsidering the specific case of a capacitive load rejection test, the branch resistance can be neglected leading to a nearzero internal angle and a load angle \ud835\udf11 = \u2212 \ud835\udf0b\n2 , the steady state", "d-q axis values for terminal voltages and currents are then obtained:\n0d\nq\nV\nV E\n , and 0 d q\nI I\nI\n \n(16)\nWhere \ud835\udc38 and \ud835\udc3c are their respective measured RMS values. Equation (15) then becomes:\n' '' 0 0' '' '( ) ( ) ( )d d\nd d d\nt t T T\nq d dV t E x I I x x e I x x e (17)\nThe least-square optimization algorithm can then be implemented using equation (17) to identify the d-axis parameters of the SG. The flowchart in Figure 3 summarizes the simplified approach to identify SG parameters from a load rejection test.\nIV. EXPERIMENTAL APPLICATION\nThe load rejection and short-circuit tests have been conducted on a 1.5kVA Synchronous Generator whose parameters are given in Table below.\nA. Sudden short-circuit experiment and analysis\nAnalysis of the phase a current has been done using the developed method and the result are shown in Figure 5.\nB. Load rejection test experiment\nThe experimental setup for the load rejection test on this same SG is given in below.", "The procedure described in section III was applied on the recorded data, the results are shown in Figure 7, where the small window presents the measured values of \ud835\udc49\ud835\udc5e.\nC. Estimated results and discussion\nTable II resumes the estimated dynamic parameters for the\nstudied SG.\nFrom these estimated values, we can see a slight difference between the results from the SCT and LRT, this is explained by the fact that the LRT has been done with the generator under saturated condition. An adjustment should thus be made from LRT results based on the saturation analysis of the SG [10].\nV. CONCLUSION\nThe work presented a review of dynamic equations of a synchronous generator in order to predict its behavior under short-circuit and load rejection events. The link between classical equivalent circuit and dynamic parameters have been discussed and a simplified method to estimate transient, subtransient and steady state parameters from short-circuit and load rejection tests have been presented and applied. Comparison between the results from these two tests records shows that despite the slight difference, the method can be an effective way to quickly identify generator parameters using readily available computer resources.\nVI. REFERENCES\n[1] R. Wamkeue, I. Kamwa, and X. Dai-Do, \"Short-circuit test based\nmaximum likelihood estimation of stability model of large generators,\" IEEE Transactions on Energy Conversion, vol. 14, no. 2, pp. 167-174, 1999.\n[2] IEEE, \"Guide for Test Procedures for Synchronous Machines,\"\nIEEE Power & Energy Society, 2009.\n[3] I. Kamwa, M. Pilote, P. Viarouge, B. Mpanda-Mabwe, M. Crappe,\nand R. Mahfoudi, \"Experience with computer-aided graphical analysis of sudden-short-circuit oscillograms of large synchronous machines,\" IEEE Transactions on Energy Conversion, vol. 10, no. 3, pp. 407-414, 1995.\n[4] I. Kamwa, P. Viarouge, and J. Dickinson, \"Direct estimation of the\ngeneralised equivalent circuits of synchronous machines from short-circuit oscillographs,\" IEE Proceedings C - Generation, Transmission and Distribution, vol. 137, no. 6, pp. 445-452, 1990.\n[5] F. P. d. Mello and J. R. Ribeiro, \"Derivation of synchronous\nmachine parameters from tests,\" IEEE Transactions on Power Apparatus and Systems, vol. 96, no. 4, pp. 1211-1218, 1977.\n[6] R. Wamkeue, C. Jolette, and I. Kamwa, \"Alternative approaches\nfor linear analysis and prediction of a synchronous generator under partial-and full-load rejection tests,\" IET Electric Power Applications, vol. 1, no. 4, pp. 581-590, 2007.\n[7] C. Han, X. Wu, and P. Ma, \"Identification of synchronous\ngenerator parameters based on 3-phase sudden short-circuit current,\" in 2011 4th International Conference on Electric Utility Deregulation and Restructuring and Power Technologies (DRPT), 2011, pp. 959-962.\n[8] S. Tahan and I. Kamwa, \"A two-factor saturation model for\nsynchronous machines with multiple rotor circuits,\" IEEE Transactions on Energy Conversion, vol. 10, no. 4, pp. 609-616, 1995.\n[9] A. E. Fitzgerald, C. Kingsley, S. D. Umans, and B. James, Electric\nmachinery. McGraw-Hill New York, 2003.\n[10] R. Wamkeue, F. Baetscher, and I. Kamwa, \"Hybrid-state-model-\nbased time-domain identification of synchronous machine parameters from saturated load rejection test records,\" IEEE Transactions on Energy Conversion, vol. 23, no. 1, pp. 68-77, 2008.\n[11] D. Hiramatsu et al., \"Analytical study on generator load rejection\ncharacteristic using advanced equivalent circuit,\" in 2006 IEEE Power Engineering Society General Meeting, 2006, p. 8 pp.\n[12] I. Kamwa, M. Pilote, H. Carle, P. Viarouge, B. Mpanda-Mabwe,\nand M. Crappe, \"Computer software to automate the graphical analysis of sudden-short-circuit oscillograms of large synchronous machines,\" IEEE Transactions on Energy Conversion, vol. 10, no. 3, pp. 399-406, 1995.\n[13] A. Keyhani, S. Hao, and G. Dayal, \"Maximum likelihood\nestimation of solid-rotor synchronous machine parameters from SSFR test data,\" IEEE Transactions on Energy Conversion, vol. 4, no. 3, pp. 551-558, 1989.\n[14] E. d. C. Bortoni and J. A. Jardini, \"Identification of synchronous\nmachine parameters using load rejection test data,\" IEEE Transactions on Energy Conversion, vol. 17, no. 2, pp. 242-247, 2002." ] }, { "image_filename": "designv11_62_0002803_sled.2017.8078423-Figure8-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0002803_sled.2017.8078423-Figure8-1.png", "caption": "Fig. 8. The sensorless controller and test installation with test motor (left, ECX 19L), and coupled motor (right, EC 32) with position sensors.", "texts": [ " The gains are calculated by pole placement and optimized for different speed ranges, specifically to accommodate a low signal-to-noise ratio at low speed and allow fast tracking at high speed, and are then selected during operation by a gain scheduler. V. EXPERIMENTAL RESULTS The maxon sensorless controller consists of a low-cost 32- bit RISC ARM processor, MOSFET power electronics and shunt current measurement on the negative potential branch of the DC link. The current controller loop runs at 25 kHz. The PWM frequency is 50 kHz. During signal injection the current transient is sampled at 400 kHz. The low-cost on-chip ADC of the MCU is used for collecting current measurements. The test setup is illustrated in Fig. 8. A maxon ECX 19 L motor is coupled to a load motor (EC 32) with Hall sensors and an optical incremental encoder with 500 counts per turn serving as ground truth. Fig. 9 shows the accuracy of the position estimate from signal injection at standstill. Notice that there is a periodic position-dependent component in the estimation error. Acceptable accuracy of the position estimate from signal injection requires correct compensation of the harmonic offset. Accuracy can then be optimized by tuning a range of parameters such as the excitation amplitude and the number of excitation directions" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0000831_amm.86.352-Figure1-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0000831_amm.86.352-Figure1-1.png", "caption": "Fig. 1 Contact zone and contact lines of the worm pair.", "texts": [ " In general, the limit line of curvature interference appears only in the sub-conjugate area 2A \u03a3 . Corresponding to 2A \u03a3 , the rotation angle of the worm can be figured out from Eq. (1), and is expressed as follows: ( ) 1 10 2 2 arccos A A C B \u03d5 \u03d5 \u03d5= \u2212 + \u2212 + , (2) where auxiliary angle 10 \u03d5 should be determined by its sin and cosine values. Herein ( )2 2 10sin B A C B\u03d5 = \u2212 + and ( ) ( )2 2 10cos A C A C B\u03d5 = \u2212 \u2212 + . Hereinafter, the determination of the curvature interference limit point is explained by taking point I shown in Fig. 1 as an example. The major parameters of the worm pair in question are noted as follows: center distance 125 a mm= , transmission ratio 12 32 2i = , the number of worm threads 1 2Z = (right-hand rotation), reference diameter of the hourglass worm at its throat 1 46 d mm= , radius of the grinding wheel 150 dr mm= , radius of the working arc profile of the grinding wheel 51 mm\u03c1 = , setting angle of the axis of the grinding wheel 19\u03b2 = , modification quantity of the center distance during machining the worm 0" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0000916_s1068371210100020-Figure1-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0000916_s1068371210100020-Figure1-1.png", "caption": "Fig. 1. Arrangement of the active part in inertial drive: (1) screw; (2) carrier plate; (3) piezoelectric plates; (4) inertial element.", "texts": [ " Another trend in the development of an inertial piezodrive is the application of kinematic pairs, which allows one to reduce the relation between the value of an object\u2019s finite movement to the value of piezoele ment deformation used to acquire this movement. Figures 1 and 2 present an inertial rotation linear drive that contains a backlash free screw\u2013nut pair (produced by Haydon Switch & Instruments). A car rier plate 2 in which two piezoelectric plates 3 are can tilever fitted in one end is joined to the screw 1 (Fig. 1) in the given pair. An inertial element 4, the weight of which considerably exceeds the weight of the element 2, is mounted at the other end of the plates 3. The drive operates as follows. The bottom of the nut is fixed on the stationary platform and the movable object is attached to the screw 1. A dissymmetrical saw tooth control signal (Fig. 3) is applied to the elec trodes of the piezoelectric plates 3. The electrodes are switched so that, if one plate shrinks, the other stretches. When forming a gently sloping leading edge of the control signal, piezoelectric plates 3 bend and the inertial element 4 slowly shifts, thus allowing the frictional force to restrict the rotation of the screw in the nut" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0001609_0954406211427833-Figure13-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0001609_0954406211427833-Figure13-1.png", "caption": "Figure 13. Fixture and gear blank mounted on a five-axis CNC machining centre.", "texts": [ " Therefore, a conical end-mill cutter is specially developed. This shape can be defined as the revolved form of the rack-cutter profile. The end-mill cutter with a semi-cone angle of 20 was manufactured on a five-axis CNC tool and cutter grinder, and a photograph of the conical cutter is shown in Figure 12. Each tooth is machined individually. A machining fixture was designed and manufactured. Using this fixture, both convex and concave gear teeth are machined. The assembly of fixture and gear blank was mounted on the rotary table, as shown in Figure 13. The fixture axis was aligned with the rotary table axis and it was ensured that the centre of the spherical gear blank coincides with the tilting axis of the machine. For machining each tooth in a row, the gear blank has to be tilted in the interval of 12 each time inside the fixture, as shown in Figure 14. at University of Ulster Library on May 14, 2015pic.sagepub.comDownloaded from CAD model of the concave gear is imported into the manufacturing module of Pro/E. The work piece is created to define the stock" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0000458_amr.591-593.1879-Figure2-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0000458_amr.591-593.1879-Figure2-1.png", "caption": "Fig. 2 3D and beam model of a disk on a flexible shaft: (a) 3D element model (b) beam element model(with element shape on).", "texts": [ " In this way the equivalent unbalance force acts as it is applied on the center node and rotates around the spinning axis. Other vector synthesis effects, such as moment caused by component tilt, can also be modeled with the method presented here. Numerical Example. A simple example is considered to illustrate the validity and accuracy of the unbalance equivalent modeling method in 3D solid rotor model. Harmonic response analyses are performed on both beam element model and solid element model as depicted in Fig. 2. The shaft of sample model has a length 0.12 m, inner radius Ri = 0.003 m, outer radius Ro = 0.004 m. A disk, with a radius Rd = 0.02 m and a thickness Hd = 0.002 m, is fixed at half of the shaft length. The material properties are given as : modulus E = 210 Gpa, density \u03c1 = 7800 kg/m 3 , Poisson' ratio \u00b5 = 0.3. An unbalance of 0.005 Kg\u00b7m is assumed at the disk and a constant damping 3% is applied during the calculation procedure to limit the vibration when going through the critical speed. Rigid constraints are applied to both ends of the rotor model" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0000566_1.4000814-Figure1-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0000566_1.4000814-Figure1-1.png", "caption": "Fig. 1 Inertia-frame and body-frame coordinate systems", "texts": [ " The proof of convergence of the algorithm is given in Ref. 4 . 3 A Case Study on VMUV In this section, the proposed identification method is used to identify the VMUV. The following equation of motion has been simulated completely in MATLAB, and input-output data for system identification is derived from this simulator. 3.1 General Equations of Motion. The kinematics and dynamic equations of vehicle motion can be developed using an inertial coordinate frame I and a body-fixed coordinate frame B, as showed in Fig. 1. Thus, the following notation is required. MARCH 2010, Vol. 132 / 024501-3 Terms of Use: http://asme.org/terms c t t g s b 0 Downloaded Fr The dynamic equations in the VMUV are written around the enter of buoyancy CB because the position of its center of mass CM is time varying unlike the AUV. Therefore, VMUV equaions are more complicated than AUV equations. The full equaions of motion that describe the dynamics of the VMUV are iven in Ref. 18 . These equations consist of translational surge, way, and heave and rotational roll, pitch and yaw equations" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0003190_978-3-030-05321-5_8-Figure8.10-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0003190_978-3-030-05321-5_8-Figure8.10-1.png", "caption": "Fig. 8.10 Prototypes of hydraulic cylinders using lightweight alloy", "texts": [ "15MPa with almost no leakage of the working fluid. As shown in Fig. 8.9, the developed cylinder successfully achieved twice the output density of that of cylinders regulated by Japanese Industrial Standards (JIS) and the International Organization for Standardization (ISO). Note that the output density is the ratio of the maximum thrust force to the mass of the cylinder here. Based on the discussion above, we developed cylinders for hydraulic robots operated by 35MPa whose diameter is 20\u201360mm. Figure8.10a, b are examples of the developed cylinders. These cylinders are applied to a tough robotic hand (the details of which will be mentioned in Sect. 8.6). Figure8.11a shows the result of measuring the sliding pressure of the developed cylinder (Fig. 8.10a). In the measurement, the cylinder drove with no load for its full stroke of 100mm. Figure8.11a, b show the time response of the operating pressures and the position of the pistons on the pushing side and the pulling side, respectively. Although the pressure increases near the stroke end point, the operating pressure achieved is <0.01MPa. Oscillating torque actuators that can be operated with 35MPa were also developed, and the output density reached higher values, as shown in Fig. 8.12. The developed motors were designed to drive with no load operated by<0" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0002503_med.2017.7984292-Figure1-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0002503_med.2017.7984292-Figure1-1.png", "caption": "Fig. 1. Quadrotor Airframe And Reference Frames.", "texts": [ " The NDO-SMC trajectory tracking control for the quadrotor UAV system devised in this paper is an extension of the work presented in [7] in which the control objective is stabilization only. The main challenge in this work was the design of a nonlinear disturbance observer combined with sliding mode controller to achieve trajectory tracking rather than stabilization. A condition was imposed on the observer gain whose satisfaction warrants complete attenuation of mismatched disturbances. The dynamical model of the considered quadrotor UAV, shown in Fig.1, is originally described in details in [17] and again employed in [6]. The position of the center of the quadrotor\u2019s mass is denoted 978-1-5090-4533-4/17/$31.00 \u00a92017 IEEE 1269 by the vector \u039e = [x, y, z]T . This position vector is expressed relatively with respect to an inertial frame (I) associated with the unit vector basis (e1, e2, e3). The attitude is denoted by \u0398 = [\u03c6, \u03b8, \u03c8]. These three angles are the Euler angles yaw (\u2212\u03c0 < \u03c8 < \u03c0), pitch (\u2212\u03c0 2 < \u03b8 < \u03c0 2 ), and roll (\u2212\u03c0 2 < \u03c6 < \u03c0 2 ) that define the orientation vector of the quadrotor in space expressed in body frame (B) fixed to the body of the quadrotor UAV and associated with vector basis (eb1, e b 2, e b 3)" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0002156_j.triboint.2017.02.022-Figure7-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0002156_j.triboint.2017.02.022-Figure7-1.png", "caption": "Fig. 7. Toroidal traction variator with power transmission from an input to an output disc through an intermediate roller of adjustable angle. (Inlaid photograph courtesy of Torotrak (UK)).", "texts": [ " (A4)) c1, c2 constants in the fluid mass-density function d sum of roughness local heights of the contact countersurfaces dz,x, dz,y auxiliary variables (Eq. (A5)) dzz,x, dzz,y auxiliary variables (Eq. (A6)) D R R R x R y+ \u2212 \u2212 \u2212 \u2212x y x y 2 2 2 2 Dx, Dy contact-ellipse semi-axial lengths E elastic modulus f, fx, fy functions (Eqs. (7)\u2013(9)) F contact normal load (Fig. 1) G E/(2+2\u03bd); shear modulus h film thickness (Fig. 1; Eq. (1)) hc, hmin central and minimum film thickness H height of the variator in Fig. 7 i, j; is, js discrete coordinates of square m \u03b1 \u03b710 /[5. 1\u2219ln( )+49. 317]9 0 n contact efficiency (Eq. (A15)) Nx, Ny numbers of partitions in Ox and Oy p pressure P tangent plane of the contact (Fig. 1) r radius of the variator in Fig. 7 rx, ry radii of curvature of the roller Rx, Ry radii of curvature s side semi-length of square S contact area (Fig. 1) sgn(\u2022) sign function ur, ud tangential velocities of the roller and the output disc U (ur+ud)/2; rolling velocity V ur\u2212ud; nominal sliding velocity w normal elastic displacement (Eq. (2)) wp pressure-induced normal elastic displacement (Eq. (3)) ws normal elastic displacement (Eqs. (6) and (11)) wx, wy traction-induced normal elastic displacements (Eqs. (4) and (5)) x, X, y, Y, z coordinates Greek symbols \u03b1 pressure-viscosity coefficient \u03b2 x X y Y( \u2212 ) + ( \u2212 )2 2 \u03b3, \u03b4 coefficients in the \u03c4L-function \u03b5 normalized correction (Eq", " (14) and (16)) \u03bbp \u03bd p x y \u03c0E(1\u2212 ) ( , )/( )2 \u03bbx \u03bd \u03c4 \u03c0G(2 \u22121) /(4 )zx \u03bby \u03bd \u03c4 \u03c0G(2 \u22121) /(4 )zy \u039b parameter of the rheological model [8,21] \u00b5 coefficient of friction \u00b5r, \u00b5d traction coefficients of the roller and disc (Eq. (A14)) \u03bd Poisson's ratio \u03bek i i\u2212 \u2212 (\u22121) /2k s (k=1, 2) \u03c1 \u239b \u239d\u239c \u239e \u23a0\u239f\u03c1 1+ p p0 0.6 \u2219 10 1 + 1.7 \u2219 10 \u22129 \u22129 ; fluid mass density \u03c10 fluid mass density at temperature \u03b8 and atmospheric pressure \u03c4L \u03c40 + \u03b3p \u2212 \u03b4\u03b8; limiting shear stress \u03c4zx, \u03c4zy tangential tractions \u03c4zx (r), \u03c4zy (r) tangential tractions on the surface of the roller \u03c40 constant in the \u03c4L-function \u03c6 roller angle (Fig. 7) \u03c9 angular velocity of the output disc (Fig. 7) dry or lubricated by a film of thickness h as in the figure. An orthogonal coordinate system Oxyz is defined, with its origin located at the nominal contact point on one of the solids. Also depicted in the figure is the contact area (S) in the tangent plane (P) of the contact. The local distance of the counter-surfaces is generally given by [8]. h x y h D x y d x y d w x y w( , ) = (0, 0) + ( , ) + ( , ) \u2212 (0, 0) + ( , ) \u2212 (0, 0)s s (1) In Eq. (1), D R R R x R y\u2245 + \u2212 \u2212 \u2212 \u2212x y x y 2 2 2 2 is the geometrical separation of the approximately ellipsoidal, unloaded counter-surfaces under dry conditions with Rx and Ry being their radii of curvature in Ox and Oy; d(x, y) stands for the sum of the local roughness heights of the counter-surfaces, again for unloaded and dry conditions; ws(x, y) is the sum of the normal elastic displacements of the counter-surfaces under operating conditions; finally, h(0,0), d(0,0) and w(0,0) correspond to h, d and ws, respectively, at the datum of displacements, which is taken to coincide with the centre of the contact (0,0)", "3 N or m al iz ed co rr ec tio n, \u03b5 0 5000 10000 15000 20000 25000 30000 35000 40000 To ta ln um be r of pa rti tio ns ,N = (2 N x+ 1) (2 N y+ 1) \u03b5max; <\u03c4zx, \u03c4zy> \u03b5max; <\u03c4zx> \u03b5mean; <\u03c4zx, \u03c4zy> \u03b5mean; <\u03c4zx> N Fig. 5. Normalized correction as a function of the number of contact partitions for a square contact under uniform pressure p and tractions \u03c4zx=\u00b5p or \u03c4zx=\u03c4zy=\u00b5p. Traction drives as in continuously or infinitely variable transmissions (CVTs or IVTs) present some of the most challenging kinematical and tribological problems met in mechanical engineering applications. An example is that of the variator in a toroidal traction drive (Fig. 7) where power is transmitted from an input disc to a coaxial output disc, through an intermediate roller of adjustable angle. The EHD contact between a disc and the roller is elliptical and the kinematics include rolling, sliding and spinning motion. A model of this type of nonNewtonian EHL problem has been validated by the author in Ref. [8] and is based on a generalized Reynolds equation as detailed in Appendix A. The same model and proprietary computer software TORO [8,15,16] are used here to explore the role of tangential tractions on lubricant film thickness, traction coefficients and contact efficiency for the contact between a roller and an output disc (Fig. 7). The film thickness calculation in Nikas [8,15,16] involves the same analytical Eqs. (3)\u2013(9) presented earlier in this article. Published studies on EHD contacts with combined rolling, sliding and spinning motion are very rare in the literature; for example, the theoretical analysis by Li et al. [17], which contains a brief literature review, and the even rarer experimental study by Jiang et al. [18]. Even so and to the best of the author's knowledge, published studies treat the problem through a classic Reynolds equation", " [24], which deals with Newtonian lubrication of rolling-sliding-spinning elliptical contacts. The aforementioned studies [21\u201324], in accordance with the vast majority of published studies, ignore the tangential-traction effect on film thickness. Table 1 provides the input data for the analysis, similarly to data used in Nikas [8,16] for an IVT developed by Torotrak in England. Let us begin with the rolling velocity, U, defined as the mean value of the tangential velocities of the roller and output disc in the rolling direction (Ox) in Fig. 7. This is varied between 2.15 and 9.85 m/s (Fig. 8) whilst keeping the nominal sliding velocity (the absolute difference of tangential velocities of the roller and output disc in Ox) constant at 0.3 m/s. Fig. 8(a) shows that the effect on hc is very similar between the three models, with relative differences small enough to be ignored: the maximum difference between Nikas [8] and Hamrock and Dowson [21] is 40 nm and that between Nikas [8] and Zou et al. [24] is 12 nm. Using the author's model [8], no effect of tangential tractions on hc is found", " As far as the EHD traction coefficients of the roller and disc are concerned, the results show that they are in error by about \u22124.6% to +14.2% (1.6% on average for the absolute value) when the traction factor is ignored. Contact efficiency in this case is also in error by up to 3.0% (0.3% on average) by ignoring the traction factor, which can be significant for high power input; for example, at 300 kW (number suggested by the IVT manufacturer [15,16]), an efficiency error of 1.0% for the variator in Fig. 7 with two discs and an intermediate roller equates to output power error of 300 \u00d7 2 \u00d7 1.0/100 = 6 kW. Finally, the effect of the ellipticity ratio is shown in Fig. 10. The ellipticity ratio is defined equal to Dy/Dx, where Dy and Dx stand for semi-axis length of the contact ellipse in the transverse (Oy) and the rolling (Ox) axis, respectively. The ratio is varied by varying the crown radius of the roller (ry) between 14 and 38 mm. Comparing the results of the three models in terms of hc in Fig. 10(a) shows that there is very close agreement between the author's results and those by Zou et al", " A generalized Reynolds equation was developed in Nikas [8,15,16] to allow for easy adaptation of various rheological laws and incorporate the variation of dynamic viscosity along the thickness of an EHL film. This is the principally correct way to treat an EHL problem where fluid viscosity varies with strain rate [19] as in traction fluids used in IVTs. Adapting the analysis of Nikas [8] to the kinematical requirements of Section 3 of this article, the generalized Reynolds Eq. (A1) and accompanying equations for power transmitted from the roller to the output disc (Fig. 7) are as follows. \u23aa \u23aa \u23aa \u23aa\u23aa \u23aa \u23aa \u23aa\u23a7 \u23a8 \u23a9 \u23a1 \u23a3 \u23a2\u23a2 \u23a4 \u23a6 \u23a5\u23a5 \u23ab \u23ac \u23ad \u23a7 \u23a8 \u23a9 \u23a1 \u23a3 \u23a2\u23a2 \u23a4 \u23a6 \u23a5\u23a5 \u23ab \u23ac \u23adx bu d c h c h d p x V \u03c9y c h d y c h c h d d p y \u03c9xd c h \u2202 \u2202 + \u2212 ( ) ( ) \u2202 \u2202 \u2212 + ( ) = \u2202 \u2202 ( ) ( ) \u2212 \u2202 \u2202 \u2212 ( )zz x zz x z x z x z x z x zz y z y z y zz y z y z y r , , , , , , , , , , , , (A1) where V u u= \u2212r d is the nominal sliding velocity with ur and u \u03c9 H r \u03c6= [ + sin( )]d standing for the tangential velocities of the roller and the disc, respectively, at the centre of the contact, and \u222bb \u03c1dz\u2254 h 0 (A2) \u222bc z dz \u03b7 q x y( )\u2254 \u2032 , ( = , )z q z q , 0 (A3) \u222bc z z dz \u03b7 q x y( )\u2254 \u2032 \u2032 , ( = , )zz q z q , 0 (A4) \u222bd \u03c1c z dz q x y\u2254 ( ) , ( = , )z q h z q, 0 , (A5) \u222bd \u03c1c z dz q x y\u2254 ( ) , ( = , )zz q h zz q, 0 , (A6) In Eqs", " In the numerical analysis of Section 3 and in accordance with Nikas [8], when the shear stress limit is exceeded (\u03c4 \u03c4 \u03c4+ >zx zy 2 2 L 2), the following treatment is performed: \u23a7 \u23a8\u23aa \u23a9\u23aa \u23ab \u23ac\u23aa \u23ad\u23aa u v \u03c4 \u03c4 \u03c4 \u03c4 \u03c4 \u03c4 \u03c4 \u03c4 \u03c4 \u03c4 \u03c4 \u03c4 > : \u2265 \u21d2 { \u2254 sgn( ) , \u2254 0} < \u21d2 \u2254 sgn( ) \u2212 zx zx zx zy zx zy zy zx L (new) L (new) L (new) L 2 2 (A10) \u23a7 \u23a8 \u23aa\u23aa \u23a9 \u23aa\u23aa \u23a7\u23a8\u23a9 \u23ab\u23ac\u23ad \u23ab \u23ac \u23aa\u23aa \u23ad \u23aa\u23aa u v \u03c4 \u03c4 \u03c4 \u03c4 \u03c4 \u03c4 \u03c4 \u03c4 \u03c4 \u03c4 \u03c4 \u03c4 < : \u2265 \u21d2 \u2254 sgn( ) , \u2254 0 < \u21d2 \u2254 sgn( ) \u2212 zy zy zy zx zy zx zx zy L (new) L (new) L (new) L 2 2 (A11) \u23a7\u23a8\u23a9 \u23ab\u23ac\u23adu v \u03c4 \u03c4 \u03c4 \u03c4 \u03c4 \u03c4 = : \u2254 sgn( ) 2 , \u2254 sgn( ) 2zx zx zy zy (new) L (new) L (A12) where xsgn( )\u2254 1 for x > 0 and xsgn( )\u2254 \u2212 1 for x < 0. Subsequently, Eq. (A9) are integrated with respect to z after substitution of the shear stresses \u03c4 \u03c4 z q x y= + ( = , )zq zq p q (r) \u2202 \u2202 with the following result: \u23a7 \u23a8 \u23aa\u23aa\u23aa\u23aa \u23a9 \u23aa\u23aa\u23aa\u23aa \u23ab \u23ac \u23aa\u23aa\u23aa\u23aa \u23ad \u23aa\u23aa\u23aa\u23aa \u23aa \u23aa \u23aa \u23aa \u23aa \u23aa \u23aa \u23aa \u23a7 \u23a8 \u23a9 \u23a1 \u23a3\u23a2 \u23a4 \u23a6\u23a5 \u23ab \u23ac \u23ad \u23a7 \u23a8 \u23a9 \u23a1 \u23a3\u23a2 \u23a4 \u23a6\u23a5 \u23ab \u23ac \u23ad \u222b \u222b z V \u03c9y z \u03c9x d = + d = h \u03c4 z \u03b7 \u03c4 z h \u03c4 z \u03b7 \u03c4 z 0 \u2212 \u2212 1 \u2212 + 0 + 1 \u2212 + zx p x \u03c4 zx p x \u039b \u039b zy p y \u03c4 zy p y \u039b \u039b (r) \u2202 \u2202 1 L (r) \u2202 \u2202 1/ (r) \u2202 \u2202 1 L (r) \u2202 \u2202 1/ (A13) where index \u201c(r)\u201d denotes the roller of the IVT (Fig. 7). Eq. (A13) are solved numerically for the unknown tractions \u03c4zx (r) and \u03c4zy (r) at each point (x, y). In the preceding computational scheme, the generalized Reynolds Eq. (A1) is solved via a Successive Over-relaxation (SOR) method (section 7 in Nikas [8]) by account of the no-slip boundary conditions listed under Eq. (A8), the cavitation constraint p\u22650 and the load balance condition \u222c p x y Fd d = . The three-dimensional computational mesh to cover the fluid film in the contact has 400,000 grid-points in the examples of the present article. In case of film collapse, solid contacts are treated as detailed in section 6 of Nikas [8]; it is noted that the results presented in Section 3 of this article do not involve local film collapse, except in one case shown in Fig. 8(b), which does not affect the general discussion. Finally, for the evaluation of the results in this article, the traction coefficients (\u03bcr and \u03bcd) and the contact efficiency (n) for the transmission of power from the roller to the output disc in Fig. 7 are calculated as follows: \u222c \u03bc \u03c4 x y F d= d d , (q: r, )zx q (q) (A14) n \u03bc u \u03bcu = d d r r (A15) where dummy index q in Eq. (A14) indicates either the roller (r) or the disc (d). [1] Rajagopal KR, Szeri AZ. On an inconsistency in the derivation of the equations of elastohydrodynamic lubrication. Proc R Soc Lond A 2003;459:2771\u201386. [2] Dowson D, Higginson GR. Elasto-hydrodynamic lubrication, the fundamentals of roller and gear lubrication. Oxford: Pergamon Press; 1966, [ISBN: 9780080114729]. [3] Johnson KL" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0002109_1464419317689946-Figure5-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0002109_1464419317689946-Figure5-1.png", "caption": "Figure 5. Definition of points P and Q at the convex geometries of moving bodies. (b) Describes the area between moving bodies in (a).", "texts": [ " The radius for body i (see Figure 4) with a wavy surface can be described using a Fourier series approach as rwave,i si\u00f0 \u00de \u00bc ri \u00fe Xnh j\u00bc1 Ajsin\u00f0 jsi \u00fe j \u00de \u00f07\u00de where ri is mean radius, si is the angular surface parameter of body i, Aj is amplitude, j is the phase angle of waviness associated with harmonic j, and nh is the number of harmonics used to represent the waviness of the surface. For a two-dimensional ball bearing system, the definition of the local position vector ui,P si,P of a point P on the surface of a body i is ui,P si,P \u00bc rwave,i\u00f0si,P\u00de cos\u00f0si,P\u00de sin\u00f0si,P\u00de \u00f08\u00de When the geometry of the bodies is approximately circular, as shown in Figure 5, a good estimate of the position of the contact points can be easily obtained. Contact points lie approximately in the line that joins the geometric centre of the bodies, and the following result holds i \u00fe si,P \u00fe j \u00fe sj,Q \u00f09\u00de Contact forces. The normal contact forces act on bodies in contact in the normal direction to the surfaces and opposite sign due to the principle of action and reaction. The normal contact force vector applied on body i is Fn, i \u00bc f , _ ni \u00f010\u00de where f , _ is the modulus of the force and is calculated as a function of indentation and its time derivative _ , and ni is a unit vector normal to the geometry at the point of contact in the inwards direction" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0000826_kem.516.203-Figure1-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0000826_kem.516.203-Figure1-1.png", "caption": "Fig. 1 Schematic illustration of laser sintering Fig. 1 process", "texts": [ " In-process monitoring of strain change and temperature at the base plate was proposed in order to investigate the thermal and strain behaviour during the sintering process. Additionally, different conditions of laser scanning direction and energy density were also applied to understand their temperature and strain relationships. These results were compared to their deformation results which were obtained by using laser displacement sensor [4]. Laser Sintering Process. The sintering process of metal powder is illustrated in Fig. 1. The system consists of a powder table, sintering table, recoater blade and a Yb:fibre laser (IPG Photonic Corp.: YLR-SM). By using CAD, a 3D model was divided into sliced layers where every layer thickness was 50 \u00b5m. Before the sintering process started, a sandblasted steel base plate was placed on the forming table. In order to produce a 50 \u00b5m thick sintered layer, the powder table was lifted up while the forming table was moved down respectively. Then, the powder from the powder table was deposited on a base plate by the recoater blade" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0003837_iet-csr.2019.0001-Figure13-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0003837_iet-csr.2019.0001-Figure13-1.png", "caption": "Fig. 13 Khepera IV robot [8]", "texts": [ " Finally, a comparison experiment has been conducted by emulating Khepera IV robot using Webots simulator, where the changes in the environment make Focused D* path not functional anymore, and FreeD* algorithm is used instead to overcome this issue. Fig. 12a shows the initially obtained path (red dot line) using Theta* algorithm where no information about the new change is available. In addition, Fig. 12a demonstrates how safe and smooth the final FreeD* path is in the presence of unexpected obstacle/damage, whereas Fig. 12b shows a closer look on how the robot moves to avoid the collision. To gauge performance in practical deployment scenarios, experiments are conducted using Khepera IV robots [8]. Khepera IV robot is depicted in Fig. 13 and its specifications are enlisted in Table 1. The diameter of Khepera IV is 140 mm and the height is 58 mm. It is a two-wheeled mobile base platform, and is equipped with several sensors such as infrared and ultrasonic for both long range and short range object detection and differential drive odometry. This setup has been used to explore the navigation problems Khepera IV robot will face during its movement towards the target as discussed in Section 3.1. Path planning is one of the most important issues in the robotics field and finding the shortest path by a simple and effective way is what the proposed approach is trying to achieve" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0000842_icecc.2011.6067774-Figure3-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0000842_icecc.2011.6067774-Figure3-1.png", "caption": "Figure 3. Pull-pressure sensors installation", "texts": [ " The rollers uphold the weight of the conveyor belt. With the continual scrolling of rollers, we couldn\u2019t install pressure sensor in the place of touching with the conveyor belt to measure pressure on it. So, we use pull-pressure sensors to measure the size of the conveyor belt bearing pressure in this system. In order to effectively measure fault signal, the pull-pressure sensors are installed both sides of rollers in the tail of loading point. The roller is three segments roller. The installation of pull-pressure sensors is shown in Fig. 3. In order to accurate measure, to expand the possible fault position of the starting point, we select rollers near by the tail of loading point to install pull-pressure sensors. We all installed six sensors in the whole system. During this range, if a foreign body make a lot of pressure on conveyor belt, one or several pull-pressure sensors will output a big value of voltage(or a big current, related with the chosen sensor types).And the combined output of six sensors is a big value. So we could measure the whole pressure of six sensors in real-time to conclude whether there is a very big pressure on the surface of conveyor belt" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0002570_icuas.2017.7991425-Figure3-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0002570_icuas.2017.7991425-Figure3-1.png", "caption": "Fig. 3: Example of flyable paths connecting point A to point B in the presence of one obstacle and fixed initial and final flight directions.", "texts": [ " If the altitude difference between the starting and target point is too large to satisfy the flight path climb/descent angle constraints, the path length is artificially augmented by adding a helix at the beginning or at the end of the path (see Fig.2). In order to generate 2D flyable paths in the presence of polygonal obstacles or no-fly zones, the overall path can be computed as a sequence of Dubins paths, whose circumferences are centered in the vertices of obstacles, assuming that starting and ending directions dA and dB can be locally achieved by using Dubins circles with a negligible increase of the path length (see Fig.3). It can be demonstrated that the shortest piecewise linear path in the presence of polygonal obstacles is a sequence of connections among waypoints including A,B and the vertices of the so called effective obstacles (see Definition 2) [16]. Hence the optimal flight path can be found solving a shortest path problem over a suitable graph called Visibility Graph. Definition 6: The Visibility Graph (VG) GV = {WV , EV } is an undirected weighted graph whose node set WV includes points A and B, and all the obstacles vertices", " Then the flyable path can be computed using Dubins circles, as follows: \u2022 Step1 Find an optimal piecewise linear path over the RVG described by a sequence of Ns+2 nodes (A, v1, . . . , vNs , B) including A and B. \u2022 Step 2 For each intermediate node h= 1, . . ., Ns of the optimal sequence, define a circle Ch, with radius Rmin centered in the corresponding polygon vertex. \u2022 Step 3 Define Dubins circles CARight and CALeft, CBRight and CBLeft, centered at the starting and target points A and B. \u2022 Step 4 Build Dubins trajectories (see Fig. 3) based on four possible sequences of circles, namely \u2013 (CARight, C1, . . . ,CNs , CBRight ), \u2013 (CARight, C1, . . . ,CNs , CBLeft ), \u2013 (CALeft, C1, . . . ,CNs , CBLeft ), \u2013 (CALeft, C1, . . . ,CNs , CBRight ), and select the shortest. Although the above procedure has been proved to be quite efficient for a single UAV, if the UAV is inserted in a flight formation, its application, in the presence of obstacles, can easily break the formation if a CFSO in encountered (see Fig.4). For this reason, the introduction of intermediate RVWs to recompose the flight formation is proposed in the following section" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0000806_6.2010-1206-Figure4-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0000806_6.2010-1206-Figure4-1.png", "caption": "Figure 4. CAD representation of concept with solar panels and batteries", "texts": [], "surrounding_texts": [ " Fuel cells\n Piston engine + batteries\n Fuel cells + solar cells\nA. Hybrid Propulsion The main challenge with an hybrid propulsion is to find a suitable way to connect the two\ndifferent engines and to study the best way to combine the both engine. The hybrid propulsion is directly inspired from the current development in the automotive industry.\nAmerican Institute of Aeronautics and Astronautics\n5", "B. Solar Panels and Batteries Propulsion Solar powered aircraft have been fascinating man kind for some decades[6-8], and several design have been presented from the first solar powered manned aircraft in 1979 by Solar Riser8, to the recent Solar Impulse project1, design for a round the world trip. The common issues for all those configurations have been the small available amount of power and the need for extreme low weight.\nAmerican Institute of Aeronautics and Astronautics\n6", "C. Fuel Cells Propulsion Fuel cell propulsion has been demonstrated by Boeing2 on a Dimona aircraft that has been\nmodified in order to keep the weight down. Other studies on the usage of fuel cells have been presented1-3. The usage of fuel cells have been proposed in many case for mide size UAV4, and Georgia Tech have been flying a UAV demonstrator powered by fuel cells.\nHere 3 different configurations are presented and compare to the Lancair Legacy. The main differences between them being a compromise between range or higer cruise speed.\nAmerican Institute of Aeronautics and Astronautics\n7" ] }, { "image_filename": "designv11_62_0000986_j.proeng.2011.03.137-Figure15-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0000986_j.proeng.2011.03.137-Figure15-1.png", "caption": "Figure 15 The mechanism with AB different from EF.", "texts": [ " M=6m-5p-b1= 134546 (8) M=3m-2p-b1= 134243 (9) The mobility of the mechanism is one. A point E, situated in the middle of the element BC, describes a circular trajectory, with the dimension of the radius equal to the dimension of the element AB. We join an element EF on the point E (Figure 14). We present two situations: The length of the element EF is equal to the length of the element AB and the element EF is parallel to the segment AB (Figure 14). The length of the element EF is different from the length of the AB element (Figure15). For the situation shown in figure 14 we calculate the mobility of the mechanism. For that purpose, we cut twice the frame, so that the number of the joints becomes equal to the number of the elements, i.e. 6 (Figure 16). The mobility number of the element 6 is b1=3 (Ty, Tz, Rx), like in the previous example. Let\u2019s calculate the mobility number of the element 5. With this aim in view, we check the possible independent movements of the extreme element 5 relative to the frame 0 (Figure 16). The point E is situated on the elements 2 and 4, at the same time", " Tybeing very small, the point E3 must be very close to the point E1 (because we displace the point F from F1 to F2), but at the same time it must be very close to the point E, which is the initial position (Figure 17). But this is not possible and the independent movement Ty of the extreme element 5 is not possible, too. The same procedure is applied to the independent movement Tz, and the result is identical. But the movements Ty(Tz) and Rx are possible. That means the spatiality of the element 5 is: b=2. 1=0-2-3-6=-bb-=M 6 1k kp21 6 1i i f-f (10) For the situation shown in figure 15 we calculate the mobility of the mechanism. We cut twice the frame, so that the number of the joints becomes equal to the number of the elements, 6 (Figure 18). The mobility number of the element 6 is b1=3 (Ty, Tz, Rx), similar to the previous example, but the mobility number of the element 5 is different because the length of the element EF is different from the length of the element AB. For that, we check the possible independent movements of the extreme element 5 relative to the frame 0. Mentally, we realize an infinitesimal translation Ty of element 5 from the point F1 to the point F2 (Figure 18)" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0000511_14786435.2011.580285-Figure4-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0000511_14786435.2011.580285-Figure4-1.png", "caption": "Figure 4. (a) Shear test-piece viewed along the axis, showing the central rigid region and the plastic annulus. (b) Circuit, c, within the zone of plastic deformation of the disc.", "texts": [ " For velocities or normal strain-rates prescribed on C it follows by Equation (14) since, if the above pair of strain-rate tensor fields are solutions so also is their difference, that Z D \u00f0_e2 _e1\u00de:\u00f0_e2 _e1\u00dedV \u00bc 0: \u00f018\u00de Since each term in the integrand of the left-hand member of this equation is positive, irrespective of the signs of the strain-rate components or the relative magnitudes of the individual strain-rates, it follows that _e1 \u00bc _e2 in D: \u00f019\u00de From the Stokes\u2019 integral theorem and Equation (3) we have for a closed circuit c within the plastic region Z c v dr \u00bc 0, \u00f020\u00de where r is the position vector of any point on c. We now apply this theorem to the disc. Figure 3 shows a shear test piece, the subject of the following analysis. On one end of a section of round bar a thin disc has been machined, while the remainder has been reduced in diameter, leaving a rod on which a flat has been milled allowing this portion to be gripped and rotated or, equivalently, to allow it to be held fixed whereas the edge of the disc is clamped and rotated. Figure 4a shows the inner portion of the disc (the rigid central region and plastic annulus) viewed along the axis of symmetry. The central rigid portion of the test-piece has radius a and the outer boundary to the plastic region has an instantaneous radius . A point within the plastic annulus has plane polar coordinates \u00f0r, \u00de, a r , 0 2 and increases as the rigid disc centre E \u00f0r a\u00de (Figure 4a) is rotated. Also shown, in Figure 4b, is a closed circuit c with sides 1 and 2 at different radial distances from the disc centre but located entirely within the plastic zone in the disc in a meridian plane. From Equation (20) it immediately follows that, if the rotation rate vector field is to be directed everywhere parallel to the axis of symmetry, v1 \u00bc v2 \u00f021\u00de and therefore, if N is the unit vector directed along the axis (Figure 4b), v N \u00bc constant: \u00f022\u00de Choosing the unit vectors T and B, as shown in Figure 4a, we have v \u00bc 1 2hThNhB hTT hNN hBB @ @T @ @N @ @B hTv 0 0 2 6664 3 7775, \u00f023\u00de where v is the velocity vector within the plastic annulus. Writing (cylindrical polar coordinates) r for B, for T and z for N this leads to v T \u00bc 0, v B \u00bc 0, if @ @z \u00f0h v\u00de \u00bc 0, v N \u00bc 1 2h hr @ \u00f0h v\u00de @r : \u00f024\u00de Inserting the values h \u00bc r, hr \u00bc 1 for the arc-length parameters we obtain v N \u00bc 1 2r @ @r \u00f0rv\u00de \u00bc A, \u00f025\u00de where A equals a constant. Integrating, this leads to v \u00bc Ar\u00fe C r , \u00f026\u00de where C equals a second constant; a result which may also have been obtained from the Euler\u2013Lagrange equations associated with Equation (2)", " Measurements of incremental displacements within the existing plastic zone were then made as above. These, reduced to normalised increments of polar angle deflection, are shown by the circles in Figure 5, where they are compared to the values calculated from (Equation (30)) (solid curve). The distortions of an initially radial line on the same disc rotated through a total angle of 14 from the initially undeformed state, with an associated movement of the outer boundary from the edge of the central rigid region, E (Figure 4a), in the disc, are shown by circles in Figure 6; the x, y coordinates being referred to the axes shown in the inset. The results in Figure 5 test the velocity field, which is described by Equation (28). The solid curve in Figure 6 gives the theoretical total displacements, which are described by Equation (47). In the case of an incompressible solid flowing plastically by means of the mechanism of extended slip it has been shown, on the basis of Equation (3), that: (1) the rotation-rate vector field obeys a circuital theorem (Equation (20)), (2) the strain-rate tensor is solenoidal (Equation (12))", " Whether the yield curve is bi-linear or the plastic modulus is merely positive the solution is unique in all three types of problem (Equation (19)). The Hencky\u2013Prandtl slip-line field theory, completed by the addition of Equation (2), as it follows from Equation (3), has been applied in this paper to the Dirichlet problem consisting of a thin disc of material. If the disc is rotated slowly within a rigid circular domain E (the rotation-rate vector field remaining everywhere normal to the plane of the disc) it is found that an expanding plastic annulus forms in the disc round the domain E (Figure 4a) and the following have been proven theoretically: (1) The outer radius of the plastic annulus expands according to Equation (38), which relates increments in the radius, , of the outer plastic-rigid boundary in the disc to increments in the angle of rotation, , of the rigid central region, r a, of the disc. (2) The velocity is entirely circumferential and is described by Equation (28). (3) The total circumferential displacements in the disc are described by Equations (46) or (47), in the case where the solid exhibits a bi-linear yield curve" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0001959_cjme.2012.01.179-Figure2-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0001959_cjme.2012.01.179-Figure2-1.png", "caption": "Fig. 2. Structure of LSFH", "texts": [ " The fourth chapter describes the experimental measurement of the radial grip rigidity. The final chapter is a conclusion of the full text LSFH is a clamping component between the shank and cutter (Fig. 1), and there are no accessory parts when the LSFH grips the cutting tool. The connection between the shank and lengthened shrink-fit holder is a taper interference fit when lengthened shrink-fit holder is pulled to shank by a thread drawbar which fixed in shank. The structure of LSFH in this paper is shown in Fig. 2. Experiment modal analysis (EMA) is used in this study to obtain an equivalent FEA model which agrees with the practicable machining state. The combined amendment procedure and the solid model of the LSFH and cutter are presented in Fig. 3, where n and n present the frequency and modal respectively; v, , and E present Poisson\u2019s ratio, density and Young\u2019s modulus, respectively. The details of the method of EMA can be found in Ref. [6]. The material of LSFH is an alloy with a large thermal expansion coefficient and the material of cutter is sintered-carbide" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0001390_ukricis.2010.5898098-Figure1-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0001390_ukricis.2010.5898098-Figure1-1.png", "caption": "Figure 1. Twin rotor multi input multi output system", "texts": [ " Therefore, this work is done using a parametric approach utilizing the input and output data of the system in order to develop the inverse model of a TRMS. II. TWIN ROTOR MIMO SYSTEM The twin-rotor multiple-input multiple-output (MIMO) system (TRMS) is a laboratory set-up developed by Feedback Instruments Limited [7] for control experiments. Its behaviour in certain aspects resembles that of a helicopter. For example, it possesses a strong cross-coupling between the collective (main rotor) and the tail rotor, like a helicopter. A schematic diagram of the TRMS used in this work is shown in Figure 1. It is driven by two DC motors. Its two propellers are perpendicular to each other and joined by a beam pivoted on its base that can rotate freely in the horizontal and vertical planes. The beam can thus be moved by changing the input voltage in order to control the rotational speed of the propellers. The system is equipped with a pendulum counterweight hanging from the beam, which is used for balancing the angular momentum in steady-state or with load. The system is balanced in such a way that when the motors are switched off, the main rotor end of the beam is lowered" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0002336_17452759.2017.1325132-Figure1-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0002336_17452759.2017.1325132-Figure1-1.png", "caption": "Figure 1. The point \u2018A\u2019 location of Cartesian and spiral coordinates systems for 2D and 3D space accordingly (a) 1D space, (b) 2D space and (c) 3D space.", "texts": [ " Currently, there are a number of laser system same \u2018Blu-ray\u2019 technology and industrial lasers allowRPS (Shulunov 2016b) achieving the micron and submicron perforation precision with a layer thickness of about 1 \u03bcm or less, pulse repetition rates in the Megahertz range (nanosecond fibre), while DPSS lasers perforation frequency increases up to 8 MHz (Lin et al. 2001). RPS AMT has important advantages over the dominant AMT currently available on the market, such as hundredfold accelerated performance, more predictable mechanical properties and smoother surfaces of the parts made because whole powder layers are simultaneously sintered into final parts. The research compares different conformal coordinate transformation algorithms to convert symmetrical and non-symmetrical objects from raster image data to spiral coordinates. Spiral coordinate system (Figure 1), derived from the spiral of Archimedes, is used for 2D and 3D spaces. It is different from the Cartesian coordinate systems requiring two coordinates (x, y) to determine the point\u2019s location in flatness (Figure 1(b)) and three coordinates (x, y, z) in 3D case (Figure 1(c)) because (l ) and (l, z) coordinates accordingly are enough for the spiral one. The point\u2019s position is defined by the spiral length within the accuracy, which depends on constant \u2018h\u2019 (the distance between successive turnings), in RPS, with \u2018h\u2019 being the height of a compressed ribbon. Three algorithms for conversion of 3D objects into ribbon This part includes a short explanation of algorithms for spiral coordinates, while the formulas and the software of conformal conversion of three-dimensional objects into ribbons are described in detail in Shulunov (2016a), Shulunov (2016c), Shulunov and Esheeva (2017a) and Shulunov and Esheeva 2017b)" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0000912_icelmach.2012.6350192-Figure9-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0000912_icelmach.2012.6350192-Figure9-1.png", "caption": "FIGURE 9: Experimental setup complete drive and optical encoder.", "texts": [ " The estimation error during the dynamic phase, when the speed reference changes from 1\u2032000 to \u22121\u2032000 rpm and viceversa, decreases from the initial \u223c200 rpm to a few rpm only. The electrical position estimation error is shown in Fig. 7. The position estimation error starts at \u223c10\u25e6, and after a transition, it vanishes as 6 soon as the electrical parameter estimations converge to the real values. The technique is applied on a 5-pole pairs BLDC motor used in automotive industry, coupled with its reduction stage, Fig. 9. For debug and illustration purpose, an optical encoder is added on the back of the drive. The drive electronics is shown in Fig. 10. The DSP used is a TMS320F28335, 150 MHz clock, from Texas Instruments. The drive datasheet values are listed in Table III. Note that the mechanical parameter values are given for the motor only. Friction constant is evaluated on the motor bearings datasheet and the inertia is that of the rotor only. Friction constant and inertia of the reduction stage are unknown" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0001196_ccdc.2011.5968649-Figure7-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0001196_ccdc.2011.5968649-Figure7-1.png", "caption": "Fig. 7. semi-physic experiment platform)", "texts": [ " It can be seen that the ship course is controlled as desired in Fig. 5. When altering the course demand, the overshoot is very small. Generally speaking, the overshoot is zero. This 2598 2011 Chinese Control and Decision Conference (CCDC) is very ideal. Fig. 6 is magnified zone of Fig. 5. The course drifts from 44.95\u2218 to 45.05\u2218 which make coincidence with the fact that the yaw accuracy is 0.1\u2218. In order to verify the performance and real-time of the proposed autopilot, the experiment is taken on semi-physic platform as Fig. 7 shown. Autopilot is based on PC104 (128MB RAM, 300MHz clock) embedded VxWorks R\u20dd OS. The rudder is driven by hydraulic power whose maximum rudder deflection is about 25\u2218 to both sides and maximum rudder rate is 5\u2218/s. The ship motion simulator simulates the ship motion with none-linear equation. The entire adaptive autopilot algorithm is described with C language. At 6kn speed, the parameters of vessel converge to \ud835\udc4e1=- 1.84, \ud835\udc4e2=-0.84, \ud835\udc4f0=0.0016, \ud835\udc4f1=0.0155, while at 10kn speed, to \ud835\udc4e1=-1.82, \ud835\udc4e2=-0" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0001198_2011-01-2157-Figure1-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0001198_2011-01-2157-Figure1-1.png", "caption": "Figure 1. Free Body Diagram of the subject vehicle", "texts": [], "surrounding_texts": [ "The mathematical model of both the tires and the vehicle are used as the basis of the developed control algorithm. First a velocity profile is generated (Figure 2) to input the desired program of motion. For the purposes of this control algorithm, the vehicle accelerates in a straight path. The second function input into the model is terrain input forces to simulate driving along rough terrain. These forces are applied by varying the normal load to each wheel. Two important assumptions are made. The first assumption is that when the weight is removed from one wheel, it is added to the other, meaning the weight is not leaving the vehicle, but rather shifting from one side to the other. The second assumption made is to limit the weight shift to 20% of the total vehicle weight. Such assumptions simplify the justification of the proposed algorithms. More precise modeling of the dynamic normal reactions will be done in future research. The random terrain inputs are created using a sinusoidal signal generator as well as a Gaussian white noise signal for random inputs (Figure 3). The numerical values of the signal wave at any given time (t) are multiplied by 20% of the user defined static wheel normal loads (Ww\u2032 and Ww\u2033). This value is then added back to the dynamic normal load (equations 13 and 14). (13) (14) The defined velocity and wheel normal loads are inputs to the simulation model. The model includes the developed control algorithm (Figure 4) to control the angular velocity of two independent driven wheels by equating the slip of each tire. First the input velocity profile is differentiated to obtain the acceleration which is then used in conjunction with the input terrain forces to calculate the generalized slip of the vehicle, S\u03b4a (equation 10). The slip of each tire is set equal to this vehicle slip. The tire slips are used to compute the desired angular velocities for each wheel (equation 11). Then the applied torque for each wheel is set by multiplying a proportional gain by the difference between the desired wheel angular velocity and the actual measured angular velocity. The torque for each wheel is sent forward to the tire model from [8] which uses the torque and wheel inertia to determine the angular acceleration of the wheel, which is then integrated to obtain the angular velocities. These angular velocities are used to compute the linear velocities of the wheels, and (equation 12). Additionally, there are two sets of feedback loops used in the control algorithm in order to improve the tracking performance. The first set of feedback loops uses the torque at each wheel and (equations 12 and 13) to calculate the slip at each wheel. This is then compared with the input slip and sent back through the control algorithm. The second set of feedback loops sends the angular velocities of each wheel, obtained by integrating the angular acceleration of each wheel, back to be compared with the angular velocities obtained from the slip of each tire." ] }, { "image_filename": "designv11_62_0003036_imece2017-70301-Figure4-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0003036_imece2017-70301-Figure4-1.png", "caption": "Figure 4\uff0eSketch of the measurement system", "texts": [ "3 shows the power amplifier designed for the loading device. The input of the power amplifier is control voltage signal cv , the outputs are 0 ci i and 0 ci i respectively. The 2 Copyright \u00a9 2017 ASME Downloaded From: http://proceedings.asmedigitalcollection.asme.org/ on 01/15/2018 Terms of Use: http://www.asme.org/about-asme/terms-of-use relationship between the control voltage and control current is linear. Then, different types of load can be applied to the dummy tool by altering the type and amplitude of the control voltage signal cv . Fig.4 shows the overview of the whole measurement system, two tension-compression-type force sensors and two eddy-current-type displacement sensors are used to measure the load applied to the dummy tool and the response displacements respectively. A data acquisition card (NI-USB6363) is exploited for the data acquisition and the output of analog control voltage signal cv . The specifications of measurement system are shown in Tab.1. Table 1. Specifications of the measurement system Parameter value Loading force Fmax 200 N Coil current Imax 10 A Swept range 0~1000 Hz Diameter of dummy tool 40 mm Air gap 1 mm As shown in Fig" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0003716_6.2019-4392-Figure16-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0003716_6.2019-4392-Figure16-1.png", "caption": "Fig. 16 Printed CAD Model of Center Annulus", "texts": [ " 15 is the opening of the regenerative channels inside the constant velocity manifold. The reason they open at the top of the manifold respective to viewing the nozzle in its print direction is so theoretically, the fluid can fill the cavity first and then equally travel through the regenerative channels. D ow nl oa de d by U N IV E R SI T Y O F G L A SG O W o n Se pt em be r 2, 2 01 9 | h ttp :// ar c. ai aa .o rg | D O I: 1 0. 25 14 /6 .2 01 9- 43 92 The center annulus was a straightforward design for printing and Fig. 16 showcases the print direction in which it was printed. The major printing consideration to make was the support material needed to support the flange. The support design was like the chamber support, where a solid 45\u00b0 structure was designed underneath the flange. This, like the nozzle and chamber, suffered from dross formation but did not matter because it was to be removed via precision machining. The sensor ports were also designed in a similar fashion to the chamber and nozzle with pilot holes" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0001212_tia.2010.2070052-Figure2-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0001212_tia.2010.2070052-Figure2-1.png", "caption": "Fig. 2. Arrangement of search coils.", "texts": [ " Thus, wide gap length makes a long axial length of the target ring. For example, the diameter of 20 mm is required in the displacement sensor head to measure rotor displacement at 7-mm magnetic gap. Considering fringing flux of the sensor head coil and to avoid interference of thrust shaft movement, the axial length of 30 mm or large is necessary in the sensor target ring. Therefore, to reduce shaft length, the integration of the bearingless motor and the displacement sensor is an important subject. Fig. 2 shows an arrangement of search coils for rotor displacement estimation. Search coils A and B are set up in the stator teeth located in the x-axis negative direction. Search coils C and D are set up in the stator teeth in the x-axis positive direction. The number of turns of search coils is all the same. The search coil A and B are connected and referred as search coil AB. The search coils C and D are also connected and referred as search coil CD. Fig. 3 shows connection methods. Fig. 3(a) shows a method of connecting search coils in opposite magnetomotive force (MMF) direction" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0003981_embc.2019.8857739-Figure1-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0003981_embc.2019.8857739-Figure1-1.png", "caption": "Fig. 1. Tool illustration. (a) The cannulation tool is inserted into the eyeball to perform cannulation. The needle is used to puncture the retinal vein, and the injection tube is used to deliver the medicine. The scleral force, denoted as Fs, is the contact force between the tool shaft and sclerotomy, the tip force, denoted as Ft , is the force applied at the tool apex, and the insertion depth is the distance between the tool tip and sclerotomy port. (b) The dimensions of the cannulation tool. The tool is constructed with two materials, an 8 mm long nitinol tube at the tip and a stainless steel tube for the remainder of the tool shaft. The FBG sensors are located in three segments along with tool shaft. (c). The section view of the tool shaft. Three fibers are attached evenly around the tool shaft.", "texts": [ " He is with School of Mechanical Engineering and Automation at Beihang University, Beijing, 100191 China, and also with LCSR at the Johns Hopkins University, Baltimore, MD 21218 USA (email: changyanhe@jhu.edu) E. Yang, I. Iordachita are with LCSR at the Johns Hopkins University, Baltimore, MD 21218 USA (e-mail: eyang31@jhu.edu, iordachita@jhu.edu) practice. One of the greatest challenges of the RVC procedure is that the required force applied by the needle to puncture the vein (puncture force) as shown in Fig. 1 (a) is well below human sensing capability [2], which potentially could lead to unintentional, excessive manipulation forces on retinal vessel and cause further iatrogenic injury. Additionally, at the sclerotomy incision, sustained contact force between the tool shaft and the sclera (scleral force) can also potentially lead to scleral damage. Therefore, successful implementation of RVC requires not only detection of vessel puncture, but also monitoring of the scleral force in order to inform the surgeon of unsafe tool-to-tissue interaction forces and improve surgical outcome", " This work presents the design and finite element analysis (FEA) of a new sensorized tool with dual stiffness that uses FBG sensors to detect both tip force and scleral force, as well as measure insertion depth. Tool calibration and validation were performed using a previously developed Steady Hand Eye Robot (SHER) research platform [10], and the linear correlations between the FBG wavelength readings and the measurements, i.e., the tip force, the scleral force and the insertion depth as described in Fig. 1 (a), were calculated. As a proof of concept for RVC use, the sensorized tool was assembled with a 3D printed handle, and a motorized cannula needle was fabricated and inserted into the tool shaft. The resulting cannulation tool can be utilized to enhance RVC procedure safety by providing surgeons with essential tissue manipulation information. The mechanical design of the sensorized tool improves upon previous dual force sensing tools [8] in the material and structural properties of the tool shaft", " The tip end of the tool consists of a 8 mm length of nitinol tubing (Young\u2019s modulus = 83 GPa, under austenite status), and the remaining shaft is a 70 mm (45 mm of which is the sensing part) length of stainless steel tubing (SS304, Young\u2019s modulus = 203 GPa). In addition to the segment interfacing between the 23 Ga outer tubes, two more segments of 26 Ga tubing are utilized at the tool tip and the distal end of the tool shaft to keep the inner injection tube centered. The injection tube itself is a 31 Ga (\u03c6 = 0.26mm) tube, and a 36 Ga (\u03c6 = 110\u00b5 m) beveled microneedle is glued at the tip with adhesive to serve as the cannula needle, which is angled at 45\u25e6, as shown in Fig. 1 (b). In order to measure transverse forces at the tool tip and sclera, three FBG fibers are arranged at 120\u25e6 intervals around the circumference of the tool shaft to sense the wavelength shift due to tool deflection as shown in Fig. 1 (c). Each fiber contains three separate 3mm-long FBG sensors located at 3 mm, 28 mm, and 34 mm from the tool tip as shown in Fig. 1 (b), for a total of 9 FBG sensors. The first segment of FBG sensors (FBG-I), which is closest to the tool tip, is capable of measuring the puncture force. The second and third segments of FBG sensors (FBG-II and FBG-III) are located outside of the eye and together are capable of measuring the scleral force as well as the insertion depth [8]. The FBG fibers themselves are secured to the outside of the tool shaft using the medical device adhesive. To maintain equal spacing among FBG fibers along the length of the round tool shaft, a custom jig was created out of polymer clay to hold both the tool shaft and the fibers in place during fabrication as shown in Fig" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0002865_we.2149-Figure5-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0002865_we.2149-Figure5-1.png", "caption": "FIGURE 5 Reduced forces acting at the rotor (A) and resulting reaction forces in the bearing unit (B)", "texts": [ " This section describes the design of the FB. Firstly, we will analyze the loads that need to be carried by the FB. We will then present the topological concept of the flexure and the corresponding method for dimensioning its shape so that the flexure will withstand the loads analyzed previously and offer the required elastic angular range. Within this article, the dimensioning is based on static assumptions. Further, it is a conservative approach since all loads taken into account are presumed to be extreme values. Figure 5 shows a simplified model of the reduced external forces acting at the rotor (A) and the reaction forces in the bearing unit (B). Again, the previous simulation data are used to determine the following values. The axial force Faxial undergoes a swelling, sinusoidal oscillation in phase with the rotor rotation. The mean value of the axial force corresponds to the centrifugal force. Unless the blades are in feathering position, the axial force is always positive because with the rotor nominal speed being a median wind speed of 14 m/s, the centrifugal force is greater than the gravitational forces" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0001158_074683410x480230-Figure4-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0001158_074683410x480230-Figure4-1.png", "caption": "Figure 4.", "texts": [ " 2, MARCH 2010 THE COLLEGE MATHEMATICS JOURNAL 131 Since we now have explicit expressions for \u03b1, d and s as functions of t we can compute the coordinates (x, y) of F \u2032 by substituting these expressions into (1): x(t) = 1 4 ln ( 2t + 2 \u221a t2 + 1 4 ) (2) y(t) = 1 2 \u221a t2 + 1 4 . (3) Now that we have expressed x and y in terms of the parameter t , let\u2019s try to eliminate t . Solving for t in (2) gives t = e4x \u2212 e\u22124x 4 . Substituting this last expression in (3), we obtain y = 1 2 \u221a( e4x \u2212 e\u22124x 4 )2 + 1 4 . Further simplification yields y = 1 4 ( e4x + e\u22124x 2 ) = 1 4 cosh 4x . Hence the locus of the focus is the catenary of Figure 4. x y We have demonstrated a geometric connection between the parabola and the catenary\u2014specifically, that the locus of the focus of a parabola which rolls on the x-axis without slipping is a catenary. 132 \u00a9 THE MATHEMATICAL ASSOCIATION OF AMERICA We close with some related questions for the reader: 1. What is the locus of the focus of the parabola y = x2 as it rolls along some other curve (such as another parabola or perhaps an ellipse or hyperbola) which is tangent to the parabola? It can be shown, for example, that the locus of the focus of the parabola y = x2 as it rolls along the parabola y = \u2212x2 is simply the directrix of the second parabola" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0000954_s1001-6058(10)60054-6-Figure4-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0000954_s1001-6058(10)60054-6-Figure4-1.png", "caption": "Fig. 4 Comparisons on surface streamlines (Left: Exp; Right: URANS)", "texts": [ " Figure 3 presents the comparisons on the timeaveraged surface pressure coefficients from two angles of view, where the contours have used the same legend as for the experimental measurement. The left figures are the measurements, while the right figures are our numerical results. The overall URANS results match the measurements very well, and local difference can be observed, especially on the surface of the streaimwise strut between the four wheels. Fig. 3 Comparisons on Cp (Left: Exp; Right: URANS) Figure 4 presents the time-averaged surface friction patterns in comparison with the experimental oil visualization. It is shown that the URANS computations have produced very similar mean flow features over the outward wheel surfaces, where the flow is attached and being visualized in the experiment. URANS results also present a small scope of second separation over the upstream wheel. The flow separation on the backsides of both the front and the rear wheels by computations differ from those of experimental visualization" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0002791_b978-0-12-812138-2.00003-9-Figure3.56-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0002791_b978-0-12-812138-2.00003-9-Figure3.56-1.png", "caption": "FIGURE 3.56", "texts": [ "74) Considering the total induced voltage, the voltage drop of the stator resistance Rs, and leakage reactance Xal, the stator voltage equation of Eq. (3.69) can be expressed as the phasor representation by Vs 5 IsRs 1 Is jXal 1Es 5 IsRs 1 Is jXal 1 Is jXar 1Ef 5 IsRs 1 Is jXs 1Ef 5 IsZs 1Ef (3.75) where, Xs\u00f05Xar 1Xal\u00de is the synchronous reactance and Zs\u00f05Rs 1 jXs\u00de is the synchronous impedance. By using Eq. (3.75), the complete per phase equivalent circuit of a cylindrical rotor synchronous motor under the steady-state condition is shown in Fig. 3.56. Excitation voltage induced by the field flux. Per phase equivalent circuit of a cylindrical rotor synchronous motor. The torque of a cylindrical rotor synchronous motor is obtained by dividing the output power by the synchronous speed \u03c9s. The input power of a synchronous motor is given by P5 3VsIscos \u03b8 (3.76) where \u03b8 is the phase angle between the stator voltage Vs and the stator current Is. Eq. (3.75) of a synchronous motor can be represented by the phasor diagram as shown in Fig. 3.57. The phase angle \u03b4 between the stator voltage Vs and the excitation voltage Ef is an important factor of power transfer and stability in synchronous machines and is usually termed the power angle (or load angle)" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0000417_edpc.2013.6689738-Figure8-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0000417_edpc.2013.6689738-Figure8-1.png", "caption": "Figure 8: Angle error influence on torque (BS vs. FW)", "texts": [ " e d , meas q , meas m d q d , ideal a , b, c el q , ideal a , b, c el T ( I ,I ) 1,5 p ( ( L L ) ( I D( I I I , ))) ( I Q( I I I , )) \u03a6 \u0394 \u0394 \u0394 \u0398 \u0394 \u0394 \u0394 \u0398 = \u22c5 \u22c5 + \u2212 \u22c5 + \u22c5 + (14) The derived equations are valid for every operating point. The impact on the different operating points is highlighted in the next subsection. B. Impact of Sensor Errors in Base Speed and Field Weakening In the BS area, the electric drive is operated according to the MTPA strategy. The current vectors are roughly perpendicular to the torque curves and the torque lines density is relatively low (compared to FW) as seen in Fig. 8. In the FW area this changes to the opposite. This results in increased torque sensitivity in respect to the angle measurement error. The rotor position sensor requirements are therefore higher in the FW area than in the BS area. This is illustrated in Fig. 8. One example calculation is performed in order to illustrate the effect of an angle error on the torque. 20 Nm is set as a reference value and the Id and Iq current reference components are calculated with the MTPA strategy and substituted in the torque equation (9) (Iq =203A and Id = -68A). For the same torque in the FW area, the current reference values are Id =- 311A and Iq =143A at a speed of 2489 rpm. Fig. 9 shows the generated torque in respect to the electrical angle error for the two operating points" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0002283_978-3-319-57078-5_15-Figure5-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0002283_978-3-319-57078-5_15-Figure5-1.png", "caption": "Fig. 5. Support construction in data preparation (a), additively manufactured brake mount on the processing platform (b) and re-processed component (c).", "texts": [ " The complete CADmodel is usually imported in STL format to the software for the SLM device (RDESIGNER) after the smoothing and sliced in horizontal layers with defined layer thickness here. If need be, a support structure is required dependent on the geometry of the component. This serves to stabilize the component and also protects against the deformation of the component which could occur during the cooling process due to its own proper weight or warpage. Additionally, increased internal tensions can be reduced with better heat dissipation in using such a support structure [18]. The support structure (see Fig. 5a and b) is automatically generated during data processing. Lastly, the SLM device for the model is transferred to a manufacturer-specific print data file. The manufacturing equipment for the newly generated brake mount is the SLM device ReaLizer SLM250. The device consists of a sieving station, processing area and main tank as well as a system control. In the processing area of the device, parts with dimensions with a maximum of 248 248 240 mm can be manufactured. Using a microwelding process, the component emerges, layer by layer, almost with its final contours on a bed of powder in the processing area with the aid of the laser at an output of 400 W", " Once the components have been uncovered, the base plate on which the components were produced can be removed from the processing area and the parts can be freed from the support constructions. Pliers are used for the rough removal of the support construction. Then the components are filed, and, in the last step, the surface is treated with a sandblaster. The surface of the brake mount, in particular, on which the support construction was bonded to the component must be reworked. The surface here is distinctly rougher where it was bonded to the support construction than the parts that had no contact with the support construction (see Fig. 5c). Already through the use of additive manufacturing the material used can be significantly reduced. While a raw volume of the bulk material of 40.8 cm3 is necessary for the machining operations, only material volume of 12.7 cm3 is required for the additive manufacturing process. This represents a reduction of material consumption of approximately 68%. But also the generation of a component plays a significant role for sustainability and thereby also the CO2 emissions [18]. Components with a high ratio of blank to pre-manufactured capacities exhibit a markedly better result for additive manufacturing than massive components" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0000003_1.4005570-Figure1-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0000003_1.4005570-Figure1-1.png", "caption": "Fig. 1 Link n with simple joints in equilibrium", "texts": [ " Consequently, the focus of this paper will be a procedure for calculating the angular acceleration \u20achn and linear acceleration \u20acsn comprising the dual acceleration of link n with simple joints relative to the previous link n-1 expressed as d dt nV\u0302n n ;n 1 \u00bc 0 0 \u20achn \u00fe e\u20acsn 8< : 9= ; (5) Calculating accelerations is an important aspect of evaluating a mechanism as demonstrated in many studies of particular mechanisms, some recent examples of which include Gong and Zhang [7], Huang et al. [8], Li and Wen [9], Rui et al. [10], and Tian and Wu [11], and it is considered that the automatic-computation formulation which can be applied on a general basis to a large variety of mechanisms both planar and spatial will be a useful contribution. In this study, we will consider links with simple joints (revolute, prismatic, or cylindrical) on proximal end n and distal end n\u00fe1 as is shown in Fig. 1. In this case, coordinate-transformation matrix B AT\u0302 takes the form of the joint-link modeling matrix n n\u00fe1 M\u0302 which can be expressed in terms of dual angles h\u0302n \u00bc hn \u00fe e sn and a\u0302n \u00bc an \u00fe e an where angle hn represents the rotation of joint n, length sn the translation of joint n, angle an the twist of link n, and length an the length of link n, and letters s and c represent the sine and cosine functions, as n n\u00fe1 M\u0302 \u00bc ch\u0302n ca\u0302nsh\u0302n sa\u0302nsh\u0302n sh\u0302n ca\u0302nch\u0302n sa\u0302nch\u0302n 0 sa\u0302n ca\u0302n 2 64 3 75 (6) The relative velocities of an open-loop mechanism of N links are related by the equation 1V\u03021 1N \u00bc Xn\u00bcN n\u00bc2 1 nM\u0302nV\u0302n n;n 1 (7) Contributed by the Mechanisms and Robotics Committee of ASME for publication in the JOURNAL OF MECHANISMS AND ROBOTICS" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0002302_b978-0-12-803581-8.10351-0-Figure15-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0002302_b978-0-12-803581-8.10351-0-Figure15-1.png", "caption": "Fig. 15 Monocoque chassis assembly. Reproduced from Radford, D.W., Fuqua, P.C., Weidner, L.R., 2004. Tooling development for a multi-shell monocoque chassis design. In: 36th International SAMPE Technical Conference, San Diego, CA, November 15\u201318, 2004, pp.1063\u20131077.", "texts": [ " Where localized external loads are introduced, added shear and compressive strengths are necessary, leading to the replacement of the honeycomb core with a phenolic of equivalent thickness to serve as a hardpoint. Material proprieties were obtained from manufactures data, and from published experimental values. The monocoque is assembled from multiple shells, which is meant to allow eased fabrication, but also to allow the incorporation of foam-filled side strakes and structural sidepods, to enhance the side impact safety. Helmet side shields, near the rear of the cockpit opening are also integral to the design. The multi-shell manufacturing and assembly concept is shown in Fig. 15. While this chapter is meant as an overview of the product development process, from design, through tooling development, layup and assembly, and not a detailed design study, the discussion in this section does suggest materials and ply stacking sequences consistent with a lightweight monocoque chassis. These specifics help demonstrate the details required to fully develop such a product. In much of the following discussion, portions of the monocoque shell are excluded in an attempt to show more conservative performance" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0003038_cdc.2017.8264622-Figure2-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0003038_cdc.2017.8264622-Figure2-1.png", "caption": "Fig. 2. Original and Expanded Networks For Example 1", "texts": [ " Then, the resulting \u201dexpanded\u201d network can be represented with associated graph G\u0304 = (V\u0304 , E\u0304) where V\u0304 = Vnf \u222aV \u2032 and E\u0304 = Enf \u222a E\u2032. The neighborhood definitions for the expanded network are exactly the same as in the original network. In-neighbor, out-neighbor and inclusive neighbor of the node i in the expanded network G\u0304 are denoted by N\u0304+ i , N\u0304\u2212 i N\u0304i, respectively. The StrBYZ behaviour is illustrated by the expanded network representation in the following example. Example 1: Consider the network G = (V,E) with n = 3 given in Fig. 2(a) where V = {1, 2, 3} and E = {(2, 1), (3, 1), (1, 2), (3, 2), (1, 3), (2, 3)}. Suppose that node 1 is faulty. Since N\u2212 1 = {2, 3} where |N\u2212 1 | = 2, there should be 2 copies of node 1 whose in-neighbors are exactly the same with node 1. Hence, we have V \u2032 = {1\u2032, 1\u2032\u2032} where N+ 1 = N\u0304+ 1\u2032 = N\u0304+ 1\u2032\u2032 = \u2205. The resulting expanded network G\u0304 = (V\u0304 , E\u0304) is given in Fig. 2(b). The edge set of the expanded graph is E\u0304 = Enf \u222aE\u2032 where Enf = {(3, 2), (2, 3)} and E\u2032 = {(2, 1\u2032), (3, 1\u2032\u2032), (1\u2032, 2), (1\u2032\u2032, 2), (1\u2032, 3), (1\u2032\u2032, 3)}. Since E\u2032 = E\u20321 \u222a E\u20322, we also have E\u20321 = {(2, 1\u2032), (3, 1\u2032\u2032)} and E\u20322 = {(1\u2032, 2), (1\u2032\u2032, 2), (1\u2032, 3), (1\u2032\u2032, 3)}. By using Definition 7, we can model the misbehaviour of StrBYZ agents. To this end, let the state value for each j \u2208 V \u2032i be represented in (4) by xij [k]. Let x[k] =[ x1[k], \u00b7 \u00b7 \u00b7 , xnnf [k] ]T and x\u0303[k] = [ x\u03031[k], \u00b7 \u00b7 \u00b7 , x\u0303f [k] ]T be the state vectors of the non-faulty and faulty nodes where x\u0303j [k] = [ xij [k] ] i\u2208N\u2212i for each j \u2208 F " ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0000285_iccme.2013.6548271-Figure3-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0000285_iccme.2013.6548271-Figure3-1.png", "caption": "Fig. 3 The parallel robot model", "texts": [ " The pose relationship between these two coordinate systems can be expressed as [ , , , , , ]T x y zP P P . At the initial process of the surgical parallel robot, each linear actuator moves along the slide way, and stops at the original position determined by the photoelectric switches. When the parallel robot is in its original position, the mobile surface coordinate system is regarded as a reference to define the base coordinate system and the mobile coordinate system. The parallel robot with a medical drill attached to its mobile platform can be abstracted as a model shown in Fig. 3. In this model, six legs are mounted on the base platform at spherical joints 1B to 6B , and arranged symmetrically on the base platform, on a circle. For the six universal joints 1P to 6P on the mobile platform, the geometric configuration is similar to that of base platform. The position of joints ( 1,2,...,6)iP i on the mobile coordinate system is described as ],,[ 000 iziyix PPP . In the same way, the position of joints ( 1,2,...,6)iB i on the base coordinate system is described as ],,[ iziyix BBB " ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0003190_978-3-030-05321-5_8-Figure8.20-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0003190_978-3-030-05321-5_8-Figure8.20-1.png", "caption": "Fig. 8.20 Schematic of McKibben artificial muscle", "texts": [ " (The left and right photographs in Fig. 8.19 are for three axes and multiple axes, respectively.) Appropriate usage of these components is expected to have great potential for composing small, high-pressure hydraulic systems. In this project, we studied a hydraulically-operated McKibben artificial muscle. McKibben pneumatic artificial muscles were studied in artificial limb research in the 1960s [30] and this artificial muscle was recently commercialized by Bridgestone Corporation for robotic applications. Figure8.20 shows a schematic representation of a McKibben artificial muscle. The structure is very simple, and is composed of a rubber tube and braded cords. When the tube is pressurized, it shrinks in the axial direction by means of the braded cords, which has a geometric structure that is essentially a network of pantographs. It is called an artificial muscle because the behavior is akin to that of a muscle in that when it expands in the radial direction, it simultaneously shrinks in the axial direction" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0000426_sled-precede.2013.6684505-Figure8-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0000426_sled-precede.2013.6684505-Figure8-1.png", "caption": "Fig. 8. Flux distribution of concentrated winding 14-pole/18-slot IPMSM", "texts": [ " The maximum speed possible by using sensorless DTC was only 1450rpm which was far less than that was achieved with certain modification to the conventional FOC control method as described in [11]. The calculated maximum speed based on conventional dq model is 1881 rpm which is closer to the maximum speed achieved with the un-modified control of the FSCW. The DTC is based on the mathematical model of the machine which assumes sinusoidal MMF. However, the fractional-slot, concentrated, non-overlapping winding produces MMF which is rich in harmonics and sub-harmonics. It is useful to compare the flux linkage density in the air-gap by using the finite element analysis method. The Fig. 8 shows magnetic flux distribution of the FSCW IPMSM. Fig. 9 shows the flux density distribution due to the stator current only for the concentrated and distributed windings. As can be seen from the figure, the concentrated winding produces a fundamental component together with several significant harmonics and sub-harmonics. Since the concentrated winding produces MMF which is rich in harmonics and sub-harmonics, the conventional mathematical model of the IPM machine with DW is not accurate for the FSCW IPMSM" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0002770_celc.201700781-Figure1-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0002770_celc.201700781-Figure1-1.png", "caption": "Figure 1. Experimental procedure scheme. The picture is showing the porous ITO in the wet state.", "texts": [ " This communication describes in detail that original protocol and provides structural (SEM and XRD) and physicochemical (electrochemistry and transparency) characterizations of the resulting material. The porous ITO electrode has then been modified electrochemically with mesostructured methylated silica in order to increase the available surface area and applied to combined electrochemical and spectral detection of industrial dyes (methylene blue, Meldola\u2019s blue and methylene green). Right after the electrospinning procedure (see Figure 1), the obtained material is a composite of indium tin hydroxide and polyvinylpyrrolidone (PVP) that is quite flexible, but it is neither conductive nor transparent. The layer is also readily solubilized by any polar solvent (mainly due to the solubilization of PVP, the stabilizing polymer). A calcination is needed to convert this composite material into indium tin oxide. The goal of this study is to immobilize the electrospun ITO nanofibers in one step on a fused silica substrate. For that reason a polystyrene (PS) coating is used to bypass the problem of shrinking by providing a layer that liquefies at the adequate moment to avoid breaking of ITO during calcination", " Afterwards, they were treated at 1000 \u02daC for 1h under nitrogen atmosphere (with temperature ramp of 100 \u02daC h-1) unless stated otherwise in results section. Nitrogen gas was treated with help of two adsorbent cartridges (oxygen adsorbent and water adsorbent) provided from \u201cSpectron Gas Control Systems GmbH, Germany\u201d. Then, the samples have been left to cool down in the oven. Finally, after heat treatment samples were rinsed with water and then with ethanol. The whole procedure of preparing the samples for further experiments is presented in graphic form on Figure 1. Mesostructured methylated silica deposition The functionalization of ITO plates or ITO nanofilament layers on fused silica with methylated mesostructured silica was achieved by applying -1.3 V for 20 s or longer if specified in deposition solution according to previously published procedure.[37] Preparation of the sol was as follows: 0.47 g of cethyltrimethylammonium bromide was dissolved in mixture of 20 ml of ethanol and 20 ml of 0.1 M solution of NaNO3 in water. After the complete dissolution 0" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0002758_icems.2017.8055934-Figure2-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0002758_icems.2017.8055934-Figure2-1.png", "caption": "Fig. 2. Open circuit flux density distributions of machines with different magnetizations", "texts": [ " The cross section of a prototype machine is shown in Fig. 1, and the detailed parameters are listed in Table I. It should be noticed that only the 12-slot/10-pole machine is shown, while the machine with other pole number will also be investigated which shares the same stator and rotor specifications. In addition, machines having different magnetizations, i.e. parallel, radial and Halbach magnetizations, are considered and the corresponding open circuit flux density distributions are shown in Fig. 2. The magnet thickness is chosen as 9mm to show the characteristics of different magnetizations clearly. But the influence of magnets will be investigated in later sections. 978-1-5386-3246-8/17/$31.00 \u00a92017 IEEE The radial and tangential flux densities of machines with different magnetizations are shown in Fig. 3 and Fig. 4, respectively. As shown, the machines with Halbach array have the highest fundamental airgap flux density in both radial and tangential fields, and the machines with radial magnetization have the lowest value" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0000065_s1068371210080092-Figure2-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0000065_s1068371210080092-Figure2-1.png", "caption": "Fig. 2. Electric circuit of two step vibroexciter.", "texts": [ " The processes that occur in electromechanical systems, which are a collection of electrical and mechanical devices, can be investigated using electromechanical analogy. The investigation of the processes that take place in the mechanical system can be replaced by investigations of the processes that occur in an electric circuit [3]. Let us investigate the electromechanical system of a two step vibroexciter for generating a mathematical model. A block diagram of the two step vibroexiter is presented in Fig. 1 and its electric circuit is given in Fig. 2. A two step vibroexiter consists of bed 1; cores 2, 3; winding systems 4, 5; springs 6, 7, 8; stiff element 10; tips 11, 12; ropes 13, 14, and vessels 15. If the core windings are connected to the ac source, diode D1 operates for one semi period and diode D2 operates for the other semi period. If the current passes through winding 4, the stiff element 9 is attracted to the core 2 and, after that, the tip 12 is lowed by ropes 13, 14 and the weight 11 is raised. As a result, the distance between weights 11 and 12 increases" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0000665_amr.139-141.394-Figure1-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0000665_amr.139-141.394-Figure1-1.png", "caption": "Fig. 1 Schematic chart of the ultrasonic aided hot-dip aluminizing system", "texts": [], "surrounding_texts": [ "Preparation of Hot-dip Aluminum Coating on Stainless Steel. In the experiments, the fundus material was austenitic stainless steel plate whose thickness is 2mm and of model ZG1Cr18Ni9Ti. Its compositions are shown as Table 1. The material of hot-dip was commercially pure aluminum whose purity is more than 99.9%. Its compositions are shown as Table 2. The stainless steel plate was cut into rectangular gobbets of the dimensions 30mm\u00d720mm\u00d72mm. The surface was washed using deionized water and dried. The hot-dip was carried out using electric resistance furnace crucible. Micro-arc Oxidation of Hot-dip Aluminum Coating on Stainless Steel. The stainless steel samples after hot-dip aluminum was polished using sand paper of granularity 200, 500 and 1000 respectively, washed in acetone solution using ultrasonic, washed using ethanol, dried and carried out micro-arc oxidation. The schematic chart of the micro-arc oxidation system was shown as Fig. 2. The voltage of the pulse power supply was adjustable from 0V to 600V. The current strength was adjustable from 0A to 20A. The pulse frequency was 1000Hz. The duty ratio of the pulse was 30%. tester of model TT260. The hardness of coating was measured using micro hardness tester of model HVS 1000. The surface roughness of coating was measured using roughness tester of model TR110. The surface appearance of the coating was observed using scanning electronic microscope of model JSM5600LV." ] }, { "image_filename": "designv11_62_0002302_b978-0-12-803581-8.10351-0-Figure48-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0002302_b978-0-12-803581-8.10351-0-Figure48-1.png", "caption": "Fig. 48 Pedal attachment system configuration.", "texts": [ " The control pedal package was designed to be inserted through the top of the monocoque, hidden under the suspension cowling, above the driver\u2019s feet. Rather than simply making a cutout in the top of the monocoque and attaching a metal pedal plate and assembly, it was determined that a more rigid composite component should be created to replace and reinforce the upper monocoque footwell region, much in the style of the cockpit rim stiffener. The control pedal area reinforcement was designed from the size and geometry of the base plate of the pedal assembly shown in Fig. 48. The sides were extended 0.10 in. for the rounded corners that would be need for the layup. The reinforced pedal mounting attachment system is dropped in through the top of the suspension box and bonded permanently in place using the integral flange. The pedal attachment reinforcement system design has a layup of an alternating 745, 0\u201390 degrees orientation of prepreg weave for a total of 15 layers. Pedal loading can be exceptionally high during braking and a rigid mounting area in the monocoque is critical" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0001063_lindi.2012.6319503-Figure2-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0001063_lindi.2012.6319503-Figure2-1.png", "caption": "Figure 2. New grinding machine construction for grinding nonconventional helical surfaces [7]", "texts": [ "1, the solution for the rotor 1 is well known and is similar to the grinding of worms. The grinding of the rotary house 2 is more complicated because of the inner surface, but also possible [4]. The determination of grinding wheel profile is possible using modern software, like HeliCAD [4] or Surface Constructor [5,6], and the profile of the determined surface of revolution grinding wheel can be dressed by CNC dressers. The grinding of helical surfaces having non-constant lead needs the special technology and machine introduced in [4]. The grinding machine shown in Fig. 2 is also needed for the finishing process of spiroid and globoid worms. The theoretically exact grinding of such type of worms was an unresolved task in the past, except for some special types of spiroid and globoid worms, for example involute spiroid worms. The problem is that the surface of revolution shaped classical grinding wheel cannot change its shape during grinding. However, the grinding of helical surfaces characterized by changing diameter along the threads or having non-constant lead needs this ability, because the contact line between the machined surface and grinding wheel changes its shape during grinding" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0001504_amr.154-155.1468-Figure1-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0001504_amr.154-155.1468-Figure1-1.png", "caption": "Fig. 1. Gaussian Fig. 2. Double ellipsoidal Fig. 3. Geometric model and meshing", "texts": [ " The thermal physical nature of material properties changed rapidly with the change of temperature. Its controlling equation of heat conduction is: QT T T Tc t x x y y z z \u03c1 \u03bb \u03bb \u03bb \u2202 \u2202 \u2202 \u2202 \u2202 \u2202 \u2202 = + + + \u2202 \u2202 \u2202 \u2202 \u2202 \u2202 \u2202 . (1) In the equation, \u03c1 , c and \u03bb are the density of material, the specific thermal capacity of material and the heat conductance of material. These are the functions of temperature. Q is the intensity of inner heat source. Heat source model. Though Gaussian (surface mode as shown in Fig. 1) type of heat source may be used for the low penetration arc welding processes, it does not reflect the action of arc pressure on the molten pool surface. Goldak et al. [6] proposed a double ellipsoidal heat source model as shown in Fig. 2, which has the capability of analyzing the thermal field of deep penetration welds. The heat flux distributions inside the front and rear quadrant of the double ellipsoidal heat source can be expressed as [7]. Geometry model and mesh. Because the welded Hastelloy C-276 alloy sheets are symmetrical about the weld center-line, only half of the weldment is modeled" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0002302_b978-0-12-803581-8.10351-0-Figure33-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0002302_b978-0-12-803581-8.10351-0-Figure33-1.png", "caption": "Fig. 33 Roll hoop design for outer perimeter.", "texts": [ " To increase the efficiency of this structure required by rules, it was also used to react loads from the suspension. Under the rules of the competition that this monocoque was originally designed for, the main forward roll hoop was constrained to be a metal tube structure. The design of this tube structure will not be covered in this discussion. However, how it is incorporated into the chassis structure is an important component of the composite chassis design for manufacture. The resulting metal tube roll hoop is shown in Fig. 33. The main bulkhead is constructed to constrain the tubular metal roll hoop movement, support the chassis in torsion, and allow for fast access in and out of the chassis. The bulkhead is designed as two carbon fiber halves that will be bonded together to form a single component shown in Fig. 34. The roll hoop is constrained from movement by the bonding of the two halves. The back half (drivers side) of the bulkhead is configured in an s-shape cross-section. The s-shape allows the outer flange to bond to the chassis to restrict movement of the bulkhead, and the main wall bonds directly to the roll hoop" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0001686_100702-Figure2-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0001686_100702-Figure2-1.png", "caption": "Fig. 2. The schematic drawing of a robotic manipulator with kinematics parameters.", "texts": [ " (38) For simplicity, here we use the notation \u039b instead of \u039b\u22121T Sp\u22121, then equation (38) can be expressed as \u2206VT (k+1) = (\u03b5(k)\u2212\u039b\u03b5(k)p)T(\u03b5(k)\u2212\u039b\u03b5(k)p)\u2212 \u03b5(k)T \u03b5(k) = \u2212\u03b5(k)T \u039b\u03b5(k)p\u2212 (\u03b5(k)p)T \u039b T \u03b5(k)+(\u03b5(k)p)T \u039b T \u039b\u03b5(k) = \u2212(\u03b5(k)p)T [(\u03b5(k)1\u2212p)T \u039b +\u039b\u03b5(k)1\u2212p\u2212\u039b T \u039b ] \u00d7 \u03b5(k)p. (39) When \u039b is chosen as shown in Eq. (15), we have \u2206VT (k+1) = ( \u03bb [\u039b ]\u2212\u03bbmin[2(S\u03b5(k))1\u2212p/T ] ) \u2016\u03b5(k)p\u20162 < 0. (40) The inequality (40) indicates that the proposed controller guarantees the stability of the flexible-joint mechanical system (1). The results obtained from the simulation of the proposed control scheme on a two-link RLFJ manipulator are shown in this section. In the simulation, the aim is to make the RLFJ manipulator whose kinematics parameters are shown in Fig. 2 track desired trajectories q1d = (25\u2212 28e\u2212t + 7e\u22124t)/20 and q2d = (5+4e\u2212t\u2212 e\u22124t)/4 from the initial position [q1i,q2i] = [1,1.5]. Robotic dynamic parameters are given by M(q) = [ (m1 +m2)l2 1 +m2l2 2 +2m2l1l2 cos(q2) m2l2 2 +m2l1l2 cos(q2) m2l2 2 +m2l1l2 cos(q2) m2l2 2 ] , g(q) = [ \u2212(m1 +m2)gl1 sin(q1)\u2212m2gl2 sin(q1 +q2) \u2212m2gl2 sin(q1 +q2)) ] , C = [ \u2212m2l1l2 sin(q2)(2q\u03071q\u03072 + q\u03072 2) \u2212m2l1l2 sin(q2)q\u03071q\u03072, ] , 100702-5 where mi and li are given in Table 1. For i = 1,2, mi, and li denote mass and length of the i-th link" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0002599_gt2017-64896-Figure4-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0002599_gt2017-64896-Figure4-1.png", "caption": "Figure 4 \u2013 STRAIN DISTRIBUTION WITHIN THE SPT SAMPLE AT A DISPLACEMENT OF 0.1MM.", "texts": [ "org/about-asme/terms-of-use various material types. A parametric study was also implemented to test the effects of various material properties on SPT curves with varying conditions and their interactions with one another, thus gauging the effectiveness of SPT to evolving material properties such as those encountered in SLM materials. Simulations comparing the model shown to that originally referenced by Campitelli et al. matched well in terms of the physical model deformation and strain distribution, which is shown in Figure 4. As the punch displaces the sample, strain begins to build up directly below the punch on the bottom of the sample, as the rest of the sample begins to bend at the radius of the bottom die. As the displacement increases the portion of the sample below the punch will begin to compress and stretch, while the portion of the sample between the punch and the lower die begins to stretch, as seen in Figure 4, which shows the plastic strain distribution in a 0.5mm sample of 316L at a displacement of 0.1mm. Necking typically occurs most in the free area between the punch and the lower die, concentrated at the area near the punch. The area directly below the punch will continue to stretch and eventually fracture, the position of the fracture will depend on the material in question and its microstructure, which will be especially relevant for the layered structures present in SLM materials [40]. The model was then tested for flexibility with material choice against another set of published experimental SPT results by Cuesta et al" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0001752_detc2011-48226-Figure2-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0001752_detc2011-48226-Figure2-1.png", "caption": "FIGURE 2. WRENCH GRAPH OF THE 4-RUU PM IN P3.", "texts": [ " Accordingly, the four actuation forces can be expressed as: F1 = ab ; F2 = cd ; F3 = ef ; F4 = gh (4) Now let x = (x; 0) and y = (y; 0). Hence, line xy collects all points at infinity corresponding to directions orthogonal to z. Let j = (z; 0), i = (m1; 0), k = (m2; 0), l = (m3; 0) and m = (m4; 0). Accordingly, the four constraint moments are expressed as: M1 = ij ; M2 = kj ; M3 = lj ; M4 = mj (5) where i, k, l and m belong to xy. A wrench graph, representing the projective lines associated with the wrenches of the 4- RUU PM in P3, is given in Fig. 2. Due to the redundancy of constraints, a superbracket of the 4-RUU PM can be composed of the four actuation forces Fi (i = 1, . . . ,4) in addition to two among the four constraint mo- ments expressed in Eqn. (5). Thus, one can write ( 4 2 ) = C2 4 = 6 superbrackets S j ( j = 1, . . . ,6). However, a parallel singularity occurs when the six possible superbrackets vanish simultaneously. For example, the superbracket S1 involving the two constraint moments ij and kj takes the form: S1 = [ab, ef, cd, gh, ij, kj] (6) 4 Copyright c\u00a9 2011 by ASME Downloaded From: http://proceedings", " Points a, c and j belong to the same pro- jective line T12. Thus, j= c\u2212a and [abej] = [(c\u2212j)bej] = [cbej] = [ecbj]. Accordingly, B = [adfh][ecbj]\u2212 [abfh][ecdj] = [a \u2022 dfh][ec \u2022 bj] = (afh)\u2227 (ecj)\u2227 (db) (14) where the dotted letters stand for the permuted elements as explained in [13, 26]. From Eqn. (14), term B is the meet of three geometric entities, namely, 1. (afh) is a finite plane having f3 \u00d7 f4 as normal vector; 2. (ecj) is the finite plane containing the finite points e and c and the direction z. Since plane (ecj) contains lines T12 and T34 (Fig. 2), the line at infinity of plane (ecj) = span(T12, T34) can be expressed by (uj) where u = (u; 0) and u is the unit vector of a finite line crossing T12 and T34. Accordingly, plane (ecj) has u\u00d7 z as normal vector; 3. (db) is the line at infinity of all parallel finite planes containing directions f1 and f2, i.e., having f1 \u00d7 f2 as normal vector. An actuation singularity occurs iff term B of Eqn. (14) vanishes. It amounts to the following vector form: ( (f3 \u00d7 f4)\u00d7 (u\u00d7 z) ) \u2022 (f2 \u00d7 f1) = 0 (15) From Eqs" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0002883_s10527-017-9732-5-Figure3-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0002883_s10527-017-9732-5-Figure3-1.png", "caption": "Fig. 3. Exoskeleton: a) structural diagram of exoskeleton; b) general view of exoskeleton for rehabilitation; 1) thigh; 2) knee; 3) foot; 4) cas ing; 5) surface.", "texts": [ " Simulation of the trajectory of the center of mass during verticalization with parameters kp = 300 and ki = 0 and at optimum parameters shows that movement of the center of mass along the optimum trajectory has essentially zero oscilla tion, such that the deviation from the specified trajectory (a vertical straight line) is no greater than 10%. Mathematical modeling showed that the patient could rise stably from the \u201csitting\u201d position in any start ing conditions. The exoskeleton is a bipedal mechanism driven by linear actuators (Fig. 3). The exoskeleton, consisting of four elements, 1 4, is established on support surface 5. The elements are sequentially connected via electric drives and hinges. Contact with the surface along which movement occurs is mediated by the contact surfaces of the feet. The device is fitted with a total of six linear drives, two of which are attached via hinges to the case, with a further four drives on each hip and knee joint. The drive shaft is attached via a hinge to one of the elements of the exoskeleton" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0002302_b978-0-12-803581-8.10351-0-Figure64-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0002302_b978-0-12-803581-8.10351-0-Figure64-1.png", "caption": "Fig. 64 Computer-aided design (CAD) model of backboard tool.", "texts": [ " The ergonomically contoured backboard tool described in the previous chapter was constructed from high-density polyurethane foam tooling board. A five-axis machining system was not available and thus the limitations of a three-axis machine had to be considered. Since z-axis travel of CNC mill that was available was limited to 4 in., it was not possible to machine the tool from a single block. In order to overcome this obstacle the tool was made from three 4 in. thick sheets of foam, two of which are bonded together and then bolted to the third, shown in Fig. 64 in green, after machining is complete. The top of the side bolsters had to be formed by hand in advance of bonding due to the difficulties of machining such small parts. The lower section of the ergonomic backboard tool is shown during the machining operation in Fig. 65. The tool was then bolted together to form the complete geometry as shown in Fig. 66. After bonding and bolting the tool together, any voids were filled with auto body filler and sanded smooth. Once the filler was sanded to contour the entire tool was progressively sanded from 100 to 220 girt sand paper" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0002302_b978-0-12-803581-8.10351-0-Figure5-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0002302_b978-0-12-803581-8.10351-0-Figure5-1.png", "caption": "Fig. 5 Multi-shell monocoque concept.", "texts": [ " In fact, the current trend in the design of the elite composite monocoques used in Formula 1 is to generate molds that allow the complete shell to be laid up in a single part.4 Thus this suggests a composite monocoque design approach that utilizes both a top/bottom and left/right pair of shells to overcome the structural, safety, and manufacturability issues. A solution to the cost and complexity of conventional monocoque design that will be the focus of this chapter makes use of multiple overlapping shells to generate complex features, including the cockpit rim stiffener.5 This approach shown in Fig. 5 utilizes both a top/bottom and left/right set of shells to overcome the structural, safety, and manufacturability issues. The multishell approach is based on earlier work on multi-shell composite pressure vessels.6 Previous research on multi-shell pressure vessels has shown that a multi-shell approach is viable in high-stress applications.6 Pressure vessels were developed and evaluated for two reasons: to investigate the viability of packaging internal components and to allow more complex-shaped pressure vessels (e" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0002029_978-3-319-02609-1-Figure3.2-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0002029_978-3-319-02609-1-Figure3.2-1.png", "caption": "Fig. 3.2 (a) Details of a Lily module, including the latching mechanism composed of four permanent magnets with different pole orientation, north-south (NS) and south-north (SN), respectively. Each module is endowed with a 2-color visual marker for tracking purposes. (b) CAD model of the Lily module used as blueprint for 3D printing. The four permanent magnets attached to each side wall as well as the aluminum block at the bottom of the module are also shown.", "texts": [ " The group sizes vary between 4 and 20 Alice robots, depending on the case study. Lily is a 3-cm-sized water-floating block endowed with four SmCo permanent magnets (one on each side\u2019s center) for mutual attraction and latching, as 1 The number and the location of these sensors can be on at the center of the board, or two on each side, depending on the case study. More details are provided in Chapter 4. 3.1 Experimental Platforms 31 well as with a 2-color visual marker for tracking purposes (see Figure 3.2a). Each module has a cuboidal, centro-symmetric shape specifically engineered for improving the relative alignment of the assembled blocks (Figure 3.2b); the main body and the cap of each module is manufactured using 3D printing. An extra aluminum block is added at the bottom of each module in order to reach an overall weight of about 17.3 g (the buoyancy limit is 21.9 g), and to improve flotation stability. Importantly, Lily is not self-locomoted; instead, it is stirred by the fluid flow produced by pumps located along the tank perimeter (Figure 3.3a). The experimental setup consists of a circular water tank of approximately 30 cm in diameter, with six inlets connected to four diaphragm pumps" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0003753_s40964-019-00095-5-Figure2-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0003753_s40964-019-00095-5-Figure2-1.png", "caption": "Fig. 2 Scanning strategy applied for the samples produced within the scope of this study illustrated by means of a section from the powder bed in a PBF-LB system. From the lower to the upper sketch, the scan vectors are rotated by 67\u00b0", "texts": [ "4404 (ASTM 316L) stainless steel (d50 = 26.28\u00a0\u00b5m; d90 = 42.56\u00a0 \u00b5m). The energy is supplied to the powder bed with a 200\u00a0W Yb fiber laser with a wavelength of \u03bb = 1064\u00a0nm, which generates a Gaussian beam profile in continuous wave operation. Recommended settings of the system-integrated beam expander for this material lead to a laser spot diameter (Fig.\u00a01) of \u03bb = 132.8\u00a0\u00b5m at the level of the layer to be scanned. The exposure of the individual powder layers was carried out according to a stripe pattern with parallel scanning vectors (Fig.\u00a02). From powder layer (layer n\u00a0\u2212\u00a01) to powder 1 3 layer (layer n), the orientation of the scanning vectors rotated by 67\u00b0 along the z-axis (building direction). For the generation of fully dense samples, an area energy density EA (formula\u00a01) of 2.00\u00a0J/mm2 was applied. The designated layer thickness (Fig.\u00a01) is 20\u00a0\u00b5m and was kept constant for the sample production of this study. Design of experiments (DoE) was used to determine an appropriate scope of experiments and to verify and interpret the results" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0002169_el.2016.4171-Figure2-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0002169_el.2016.4171-Figure2-1.png", "caption": "Fig. 2 Concept of integrated control of feedback and feedforward control using shift technique of central position", "texts": [ " In other words, the central position of the feedback control shifted to the opposite of the user position, in proportion with the estimated speed. This led to robust results and a consistent control pattern to aid with unreliable predictions. This was especially true for low-speed walking, because the position-based smoothing technique was naturally incorporated. In addition, this method maintained the direction of the position-based feedback control while increasing the control magnitude. This can play an important role in the omnidirectional treadmill, as it overcomes uncertainty of estimated walking direction. Fig. 2 shows the relationships of user position (PU), original central position (PC), and shifted central position (PSC). The integrated speed control algorithm is described in the following equations, where VT and VE represent the final treadmill speed and estimated user speed, respectively. Feedback proportional gain (Kp) and feedforward gain (Ke) were applied in order to fine tune the target system. The smoothing function (S) represents the dead and dampened zone handling, which was implemented using the boundary radius of each zone, on the basis of previous work [7]" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0001637_0954410012442633-Figure1-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0001637_0954410012442633-Figure1-1.png", "caption": "Figure 1. Schematic of an ALV.", "texts": [ " Since the chief guidance commands are given in pitch channel, this work is concerned with study of controlling the longitudinal motion (pitch channel). The article is arranged as follows: the Problem formulation section covers the problem formulation accompanied by concise dynamic characteristics of the ALV. In the sections Controllers setup and Controls Implementation and simulations, the control system arrangement and results are developed and evaluated, respectively, along with a brief derivation of sliding-mode and feedback linearization control. The last section concludes the paper. Assuming rigid airframe for an ALV (Figure 1), Newton\u2019s second law of motion leads to well-known 6-DOF equations of motion of an ALV as follows2,18 Fx \u00bc m _u\u00fe qw rv\u00f0 \u00de Fy \u00bc m _v\u00fe ru pw\u00f0 \u00de Fz \u00bc m _w\u00fe pv qu\u00f0 \u00de Mx \u00bc Ixx _p My \u00bc Iyy _q\u00fe rp Ixx Izz\u00f0 \u00de Mz \u00bc Izz _r Izx _p\u00fe pq Iyy Ixx 8>>>>>>< >>>>>>: \u00f01\u00de where m and Iij correspond to mass and moments of inertia and the vectors (u, v, w) and (p, q, r) are linear and angular velocities, respectively, all stated in body frame. The objective is to make the ALV track the commanded pitch angle. Although the strength of backstepping controller is laid in its ability to handle the dynamic non-linearities, the control design being considered in the literature, usually is derived by a linearized ALV model; this is mostly due to the fact that the ALV attitude control systems which are prevalently designed through linearized equations of motion, are perfectly acceptable" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0001687_978-90-481-9689-0_68-Figure2-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0001687_978-90-481-9689-0_68-Figure2-1.png", "caption": "Fig. 2 Singularity loci for the 3-UPU TPM Geometry 1.A.", "texts": [ " By geometric inspection, it can be seen that this condition occurs when two axes of the revolute pairs of the platform (q4i, q4 j, i = 1,2, j = 2,3, i = j) projects on the two corresponding axes of the base (q1i, q1 j), providing the projection direction is along the shortest distance of the two axes. Condition (ii) is a concise and geometric definition of singularity occurrence and it represents a powerful geometric tool for detecting this type of singularity. In the following, two known geometries of the manipulator are recalled to show the main properties of the 3-UPU TPM [1,4]. For the first geometry, defined as Geometry 1.A, the axes of the three revolute pairs in the base/platform are coplanar and intersect at three points Ci, i = 1,2,3 as shown in Figure 2; here, only the revolute pairs in the base/platform are represented for clarity, all the other ones are omitted. The same simplification has been adopted in the next figures. A reference system Sb fixed to the base with origin Ob (the centre of the circle with radius b defined by the points Bi, i = 1,2,3) is chosen. Axes x and y are on the plane \u03c0 defined by the three points Ci, i = 1,2,3, with x axis through point B1, z axis is normal to the plane \u03c0 , while y axis is taken according to the right hand rule" ], "surrounding_texts": [] }, { "image_filename": "designv11_62_0001140_s1672-6529(13)60226-7-Figure2-1.png", "original_path": "designv11-62/openalex_figure/designv11_62_0001140_s1672-6529(13)60226-7-Figure2-1.png", "caption": "Fig. 2 High-speed 3D video recording system.", "texts": [ " This signal was also used as the starting point to analyze the video images in the motion analysis system. The effective area was calibrated using a 0.6 m \u00d7 0.4 m \u00d7 0.2 m calibration frame with 16 non-colinear points that approximately filled the overlapping region of the four cameras before the experiments. The calibration frame was recorded by the four cameras and was then removed from the effective area. The global coordinate system X-Y-Z was defined based on the result of the calibration analyzed through 3D motion analysis (see Fig. 2). Fig. 3 shows the body-centered coordinate system x-y-z located at the MC, the x-axis along the body symmetrical axis toward the heads of the crab, the y-axis laterally to the left, and the z-axis perpendicular to the xy-plane. The location of the MC on the dorsal carapace for each crab was determined by suspending the crab from strings attached to two chelas[2,18]. We recorded the crab while suspended from each location toward the dorsal carapace and defined the extended line of the string in each image using the Adobe Photoshop SC5 Extended software" ], "surrounding_texts": [] } ]