diff --git "a/designv11-71.json" "b/designv11-71.json" new file mode 100644--- /dev/null +++ "b/designv11-71.json" @@ -0,0 +1,10997 @@ +[ + { + "image_filename": "designv11_71_0001608_00207721.2020.1799110-Figure1-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0001608_00207721.2020.1799110-Figure1-1.png", + "caption": "Figure 1. 3D-simplified virtual forced model.", + "texts": [ + " Because the 3D artificial potential field method readily falls into the local optimum, the target point is unreachable and the calculation amount is complicated, the 3D simplified virtual force model is established by introducing virtual functions, and the rotation matrix in the space coordinate system is used to determine the model parameters. This obstacle avoidance strategy is combined with UAV motion control and is successfully applied to multiple targets of drone swarms. The 3D simplified virtual force model is shown in Figure 1, and the general idea is as follows: according to the position information of the ith UAV at time t, find the position relationship between the two nearest obstacles (UAVS), namely, p2 and po2, and calculate the expected speed and desired position points pg1 through a 3D particle swarm algorithm based on kinematic constraints. Then, find the relationship between the desired velocity and the rotation matrix angles a and b of the spatial coordinate system to solve the rotation matrix parameters. Finally, based on this force model, only the repulsive forces in the X -axis and Y -axis directions in the XOY Z plane are considered to guide the UAV \u2019s deflection; that is, the combined force FXY . Through vector synthesis, the actual speed requirement of the UAV can be obtained. As shown in Figure 1, the spatial coordinate conversion relationship from xoyz coordinate to XOY Z coordinate is as follows: the rotationmatrixT obtained by rotating the xoyz space coordinate system by the angle a around the x-axis and then rotating the angle b around the y-axis is as follows: T = \u23a1\u23a3\u2212 sin(b) cos(b) 0 0 cos(a)\u2212 sin(a) \u00b7 cos(b) \u2212 sin(a) \u00b7 cos(b) 0 sin(a)+ cos(a) \u00b7 cos(b) sin(b) \u00b7 cos(a) \u23a4\u23a6 (21) In Equation (21), the rotation parameters a and b are determined as follows: according to Equation (17), the expected speed vector of the UAV is VRi(t + 1), and its velocity component in xoyz space coordinates is as follows: VRi(t + 1) = Fg1 = [Fg1x Fg1y Fg1z] (22) the projection vector Fyoz of the gravity (expected velocity) on the yoz plane is as follows: Fyoz = Fg1y + Fg1z (23) According to the vector angle formula, the rotation matrix parameters are as follows: a = arcos ( Fyoz \u00b7 Fg1z \u2016Fyoz\u20162 \u00b7 \u2016Fg1z\u20162 ) (24) or b = arcos ( Fyoz \u00b7 Fg1 \u2016Fyoz\u20162 \u00b7 \u2016Fg1\u20162 ) (25) Then, according to formula (8), the repulsive force values of p2 and po2 can be obtained as Fp2 and Fpo2" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0003259_chicc.2015.7260770-Figure2-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0003259_chicc.2015.7260770-Figure2-1.png", + "caption": "Fig. 2: (a) Movement trajectories of ten agents and the contour map modeled by agent 1 for Case 1 in Region A; (b) RMS errors on the environment model for all agents over 100 iterations for Case 1 in Region A.", + "texts": [], + "surrounding_texts": [ + "A HDRHC approach has been developed for the problem of gradient climbing and consists of two levels. In the first level, a radial basis function network has been used to model the search environment by which a dynamical optimization problem is designed. By solving the dynamical optimization problem, a desired trajectory can be produced, which can guide the agent to trace the peaks of the field of interest in the search environment. In the second level, a cooperative control optimization problem has been proposed. On the basis of the desired trajectories of neighboring agents, a real movement trajectory of the agent can be obtained by solving the cooperative control optimization problem. Finally, the effectiveness of the proposed HDRHC approach has been illustrated through the problem of gradient climbing." + ] + }, + { + "image_filename": "designv11_71_0003144_978-81-322-2301-6_13-Figure1-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0003144_978-81-322-2301-6_13-Figure1-1.png", + "caption": "Fig. 1 Kinect sensor", + "texts": [ + " Section3 clearly explains the proposed algorithm, whereas Sects. 4 and 5 illustrates the experimental results and performance analysis. Section6 concludes with idea about future work. The subsections below explain the Kinect sensor and multi-class support vector machine algorithm briefly. Kinect sensor, consisting ofRGB(red, green, blue) camera, infrared (IR) projector, IR camera andmicrophone, is capable of full human body tracking up to two persons at a time [5, 6]. It looks like a webcam as displayed in Fig. 1. It detects 3D representation of an object using depth sensor [10, 11], which consists of infrared laser [12]. Kinect sensor produces RGB video as the output using 8-bit VGA resolution camera [7]. It tracks the human body using 20 body joint co-ordinates within a finite amount of range 1.23.5m [8, 9].Voice gestures can also be recordedwith the help ofmicrophone array. A light emitting diode (LED) is present in front of the Kinect sensor to ensure that Kinect is running properly. Kinect sensor brings forth 3D information about human" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0001303_s00202-020-01035-1-Figure2-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0001303_s00202-020-01035-1-Figure2-1.png", + "caption": "Fig. 2 The machine subdomains: a the rotor subdomains including regions I and II and the air-gap subdomain including region III and b the stator subdomains including regions IV and V", + "texts": [ + " The coefficients aIII n , bIII n , cIII n and dIII n are determined con- sidering the continuity of magnetic vector potential The continuity of the magnetic vector potential between the internal air-gap subdomain III and the regions I and II leads to The continuity of the magnetic vector potential between the internal air-gap subdomain III and the regions IV and V leads to (18)A V ( t 7 \u22c5 ) = A III ( t 6 \u22c5 ) (19)aV 0 = R 4 0 M + 1 \u222b k+ k A III ( t 6 \u22c5 ) d (20)aV h = 2 \u222b k+ k A III ( t 6 \u22c5 ) cos ( z ( \u2212 j )) d (21) 2A III t2 + 2A III 2 = 0 for { t 6 = ln ( R 3 \u2215R 4 ) \u2264 t \u2264 t 4 = 0 0 \u2264 \u2264 2 (22) A III(t \u22c5 ) = \u221e\ufffd n=1 \u239b \u239c\u239c\u239d 1 n Cosh(n(t\u2212t6)) Sinh(n(t5\u2212t6)) aIII n + 1 n Cosh(n(t\u2212t5)) Sinh(n(t6\u2212t5)) bIII n \u239e \u239f\u239f\u23a0 Cos(n ) + \u221e\ufffd n=1 \u239b\u239c\u239c\u239d 1 n Cosh(n(t\u2212t6)) Sinh(n(t5\u2212t6)) cIII n + 1 n Cosh(n(t\u2212t5)) Sinh(n(t6\u2212t5)) dIII n \u239e\u239f\u239f\u23a0 Sin(n ) (23) A III t \ufffd\ufffd\ufffd\ufffdt=t5 = g( ) = \u23a7 \u23aa\u23a8\u23aa\u23a9 AI t \ufffd\ufffd\ufffdt=t2 for i \u2264 \u2264 i + AII t \ufffd\ufffd\ufffdt=t4 for j \u2264 \u2264 j + 0 elsewhere Interface condition (23) gives Interface condition (24) gives The Poisson equation in the rotor permanent magnet subdomain is given by where t 4 = ( R2 R3 ) and t 3 = 0 , M and M are tangential and radial components of magnetization. The radial and tangential components of parallel magnetization for inset design in Fig.\u00a02a can be expressed as where (24) A III t \ufffd\ufffd\ufffd\ufffdt=t6 = h( ) = \u23a7\u23aa\u23a8\u23aa\u23a9 AV t \ufffd\ufffd\ufffdt=t7 for k \u2264 \u2264 k + AIV t \ufffd\ufffd\ufffdt=t9 for l \u2264 \u2264 l + 0 elsewhere (25) aIII n = { 2 2 \u222b i+ i g( ) cos (n )d for i \u2264 \u2264 i + 2 2 \u222b j+ j g( ) cos (n )d for j \u2264 \u2264 j + (26) cIII n = { 2 2 \u222b i+ i g( ) sin (n )d for i \u2264 \u2264 i + 2 2 \u222b j+ j g( ) sin (n )d for j \u2264 \u2264 j + (27) bIII n = { 2 2 \u222b k+ k h( ) cos (n )d for k \u2264 \u2264 k + 2 2 \u222b l+ l h( ) cos (n )d for l \u2264 \u2264 l + (28) dIII n = { 2 2 \u222b k+ k h( ) sin (n )d for k \u2264 \u2264 k + 2 2 \u222b l+ l h( ) sin (n )d for l \u2264 \u2264 l + (29) 2A II t2 + 2A II 2 = \u2212 0 R 2 e\u2212t ( M \u2212 Mr ) for { t 4 \u2264 t \u2264 t 3 j \u2264 \u2264 j + (30)Mrn = (\u22121)kBr 0 p r \u22c5 [ A 1n ( p" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0002128_j.mechmachtheory.2020.104171-Figure4-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0002128_j.mechmachtheory.2020.104171-Figure4-1.png", + "caption": "Fig. 4. A schematic representation of oil-lubricated housing for measuring the friction torque with sealing (a) one RSSR in each side denoting by M 2 \u00d7RS S R ; (b) one RSSR in the left side (driving side) and a closed lid in the right side denoting by M 1 RS S R + cl osed l id .", + "texts": [ + " The test rig allows one to easily replace the shaft and RSSR with new ones after each test. The rig is instrumented to measure the rotational speed, friction torque, oil temperature, and housing pressure. For measuring the friction torque as a function of the sliding velocity, the test rig operates in room temperature and ambient pressure. The idea for measuring the friction torque for an RSSR is that the total friction torque of the system is measured in the case of sealing by means of two RSSRs on the left and the right side. This friction torque is shown by M 2 \u00d7RSSR (see Fig. 4 a). Once again, the total friction torque of the system is measured in the case of sealing by means of only one RSSR on the left side and a closed lid on the right side instead of sealing by means of an RSSR. In this case, the measured friction torque is shown by M 1 RSSR + cl osed l id (see Fig. 4 b). Thus, the friction torque of an RSSR can be expressed through the difference of the above consecutive measurements. Finally, the friction force of an RSSR can be described through the ratio of the friction torque to the shaft radius given by the following equation. { M RSSR = M 2 \u00d7RSSR \u2212 M 1 RSSR + cl osed l id F f = M RSSR / R 1 (1) Fig. 4 shows the sealing of the oil-lubricated housing for measuring the friction torque given by Eq. (1) schematically. All friction losses, from bearings, groove nuts, etc., are eliminated in this way. Therefore, the only remaining magnitude through subtraction is the friction torque of an individual RSSR. The measurement technique can also be verified by Wennehorst [28] , who also conducted RSSR tests with different oils to measure the friction torque. Moreover, before recording data, the shaft was operated for half an hour at a constant velocity to reach the steady-state condition in order to eliminate the thermal influence and change of lubricating oil film on the friction force" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0003544_cta.2127-Figure15-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0003544_cta.2127-Figure15-1.png", + "caption": "Figure 15. Armature coil locations and flux directions.", + "texts": [ + " It can be considered that the measurement is also close to the calculated value as M31MAX cos45\u00b0 M31MAX cos67:5\u00b0 \u00bc 1:846: Copyright \u00a9 2015 John Wiley & Sons, Ltd. Int. J. Circ. Theor. Appl. 2016; 44:1094\u20131111 DOI: 10.1002/cta So, the absolute error involved here is 1.905 1.846=0.059, and the relative error can then be obtained as 1:9 1:846\u00f0 \u00de 1:846 \u00bc 0:029 This measurement only has a 2.9% error that should be tolerable. In order to clarify the magnetic relationship between the coils on the armature winding, Figure 15 represents in a more clear way in terms of the current and its flux directions. To use the same example already discussed, the mutual inductanceM31 between 8C\u20136C (three coils) and 3C (one coil) can be illustrated in Figure 16. In Figure 16, it can be observed that the flux directions of 8C\u20136C and 3C form a 90\u00b0 angle. This is another way to explain these angles affecting on the measured mutual inductances. So, our calculations mentioned in the preceding text for obtaining the mutual inductances on the armature winding with these inherent geometrical angles can be verified" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0000987_0954406220915499-Figure1-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0000987_0954406220915499-Figure1-1.png", + "caption": "Figure 1. The schematic of dual-axle steering mechanism.", + "texts": [ + " On the one hand, it can provide a set of differential equations which come from the engineering practice for studying the phenomenon of multiple limit-cycles in mathematics. On the other hand, the multiple limit-cycles shimmy mechanism of dual-axle steering mechanism is found preliminarily, which can provide a theoretical research method for the multiple limit-cycles vibration of similar mechanism in engineering. Mechanical model and mathematical model A schematic of dual-axle steering mechanism is shown in Figure 1. The steering force generated by the turning of the steering wheel is transmitted to the steering gear through the steering operating mechanism (steering shaft, drive shaft, steering universal joint, and so on). After reducing the speed and increasing the torque through the steering gear, the torque is transmitted to the first rocker. The steering knuckle arm of the first axle is driven to move by the movement of the anterior longitudinal rod. The left steering knuckle and the left wheel mounted on it rotate around the kingpin in the same direction" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0002830_978-94-007-6046-2_42-Figure1-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0002830_978-94-007-6046-2_42-Figure1-1.png", + "caption": "Fig. 1 Pendulum models. (a) Inverted pendulum. (b) Linear inverted pendulum", + "texts": [ + " Unlike a normal inverted pendulum, it has linear dynamics S. Kajita ( ) Humanoid Research Group, Intelligent Systems Research Institute, National Institute of Advanced Industrial Science and Technology (AIST), Tsukuba, Japan e-mail: s.kajita@aist.go.jp \u00a9 Springer Nature B.V. 2019 A. Goswami, P. Vadakkepat (eds.), Humanoid Robotics: A Reference, https://doi.org/10.1007/978-94-007-6046-2_42 905 which are useful to design biped gait and control. To put in the concrete, let us look at the motion equation of a \u201cnormal\u201d inverted pendulum as shown in Fig. 1a. It is given by R D g r sin ; (1) where is the angle of inclination from vertical, r is the length of the pendulum, and g is gravity acceleration. This is a nonlinear differential equation because it has sin on the right side. As a result, high mathematical skill is required to treat it. For example, we must use elliptic integrals to write its analytical solution. Figure 1b shows the linear inverted pendulum. It is assumed that the pendulum has an actuator which keeps the center of mass (CoM) constant height. Assuming that all mass is concentrated on the top of the pendulum, its equation of motion is given by Rx D g z x; (2) where x and z denote the position of the CoM. As we see, this is a linear differential equation with a constant coefficient .g=z/. Note that the same equation holds in the vicinity of D 0 of the normal pendulum dynamics (1). However, the LIP dynamics (2) is held at any value of x and this linearity gives us an ease of mathematical analysis" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0001060_j.ijimpeng.2020.103609-Figure4-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0001060_j.ijimpeng.2020.103609-Figure4-1.png", + "caption": "Fig. 4. Experimental setup (left) and cross-section (right) of pneumatically compacted sandbag specimen.", + "texts": [ + " The inner layer of LDPE storage bag was filled with sand while the outer layer of LDPE storage bag was connected to a tubing to draw air out using a vacuum pump as shown in Fig. 3. The vacuum compacted sandbag specimen was measured to have a negative pressure of \u22120.6 bar. The sandbag specimens were pressurised to selected pressures \u2013 0.5 bar, 1.0 bar, 2.0 bar using an inflatable airbag. Air was delivered to the opening of the airbag (a small incision was made to fit the opening of airbag through), inflating the sandbag as shown in Fig. 4. A pressure gauge was used to ensure that the sandbag was pressured to desired setting. A metal block was used to support the pneumatically compacted sandbag preventing it from toppling. The densities for the vacuum compacted and pneumatically compacted sandbags are expected higher than that of the uncompacted one. However, their values cannot be determined using Eq. (1) because the equation does not include the effect of vacuum and pneumatically compaction on the calculated volume of sandbag. The projectile used for ballistic impact test was a 12 mm diameter chrome bearing steel (AISI E 52,100 steel) sphere weighing 7 g" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0002861_978-3-030-55889-5-Figure5.6-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0002861_978-3-030-55889-5-Figure5.6-1.png", + "caption": "Fig. 5.6 Plate, a principal stresses under a vertical point load, b influence function for \u03c3yy near the corner. If the vector gi points in the direction of the principal stresses, stiffness changes\u2014which generate forces f+ i in the direction of the principal stresses\u2014have the greatest impact on the stress \u03c3yy at the corner", + "texts": [ + "37) and this can be split further into N scalar products of the nodal displacements gi and the equivalent nodal forces f+ i at theN nodes, see (3.230), and therefore the interplay 302 5 Stiffness Changes and Reanalysis of two plots determines the size of J (e): (i) the sensitivity plot of the functional and (ii) the plot of the principal stresses. J (e) = gT f+ = sensitivity plot \u00d7 principal stresses. The sensitivity plot determines the direction and the size of the gi, and from numerical tests we know that the principal stresses determine the direction of the forces f+ i , see Fig. 5.6, while the size of the f + i depends on the size of the element and on the size of the stiffness change in the element. It is the same situation as in Sect. 3.32, only the nodal forces are now the f+ i . 5.8 The Effects Fade Away 303 The further we move away from the source point, the more \u201clinear\u201d an influence function becomes, see Fig. 5.7, since the vector g resembles more and more a vector g = a + x \u00d7 b, and this means, because the vector f+ is orthogonal to such vectors, that the influence of the forces f+ tends to zero with increasing distance from the source point at which we observe the effects which the f+ produce, J (e) = J (uc) \u2212 J (u) = gT f+ (a + x \u00d7 b)T f+ = 0 " + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0001833_ccta41146.2020.9206372-Figure1-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0001833_ccta41146.2020.9206372-Figure1-1.png", + "caption": "Fig. 1. The vehicle-manipulator, which is considered in this paper [4]", + "texts": [], + "surrounding_texts": [ + "I. INTRODUCTION\nRobot systems are used in many different applications. One of them is the field of so-called vehicle-manipulators (VM), which consist of a mobile platform (vehicle) with an attached robot arm (manipulator). Large hydraulic VMs are used in forestry or in road maintenance works [1]. The use of hydraulic cylinders in these applications is widespread due to their good power-to-weight-ratio and robustness against dust and other contaminations in the dirty working area [2]. Currently, such systems are controlled by a dedicated operator and there is a low level of automation available. The operator has to perform a dual-task: Keeping the vehicle on its path and controlling the hydraulic manipulator to fulfil a dedicated working task. This dual-task is psychologically and physically demanding and therefore, the use of automation can increase the efficiency and safety. A simple solution would be to apply automated driving approaches to the vehicle. However, currently available approaches do not consider the additional manipulator e. g. [3]. This leads to poor results in the given setting as the goals of VMs differ from the goals achieved by autonomous vehicles. VMs have to fulfil a specific task with the manipulator. This task often\n1Balint Varga and So\u0308ren Hohmann are with the Institute of Control Systems, Karlsruhe Institute for Technology, 76131 Karlsruhe, Germany balint.varga@kit.edu, soeren.hohmann@kit.edu\n1Arash Shahirpour, Yannick Burkhardt and Stefan Schwab are with Control in Information Technology, FZI, Research Center for Information Technology, 76131 Karlsruhe, Germany schwab@fzi.de\nconsists of following a trajectory. Consequently, a novel shared-control (SC) concept was proposed in a previous work of the authors [4] that shows the following key aspects:\n\u2022 The vehicle is fully automated and thus able to follow its reference trajectory. \u2022 The manipulator is controlled by the human operator. This means that the reference trajectory of the manipulator is determined only by the human operator himself and therefore is not available for the automation. \u2022 The vehicle supports the desired motion of the manipulator based on the inputs of the operator.\nThis concept does neither need to capture the environment of the manipulator nor is a reference trajectory of the manipulator necessary. The supportive movement of the vehicle is only based on the inputs of the operator applied to the joystick. This concept enables a real-life implementation. The effectiveness of the concept is verified using simulations\nin [4]. This contribution provides an additional validating case study of the SC. Furthermore, a personalization of the concept is examined through practical tuning of the most crucial parameters. Finally, the concept is extended with a force feedback (FF) on the haptic user interface. This extension only requires information that is already available for the basic SC. Also it is not necessary to adapt the algorithms of the basic SC concept.\nThe remaining of the paper is structured as follow: In Section II, the state-of-the-art of related applications is\n978-1-7281-7140-1/20/$31.00 \u00a92020 IEEE 542\nAuthorized licensed use limited to: Dalhousie University. Downloaded on November 01,2020 at 19:34:14 UTC from IEEE Xplore. Restrictions apply.", + "presented. This is followed by a detailed discussion of the SC concept and the validation methods in Section III. In Section IV, the simulator is presented with the real-time simulation of the VM and with a graphical user interface. Section V presents the experiment design. The results of the validation are discussed in Section VI. Finally, Section VII concludes the contribution and depicts possible future work.\nThe field of large VMs is related to the modelling of small VMs and to the domain of large hydraulic manipulators with static base. Therefore, the discussion of the state-of-the-art deals with these two major categories.\nNumerous control methods have been developed in the field of small vehicle-manipulators. Such systems have been investigated in simulations and using small demonstrators. While early research in this area [5] and [6] dates back to the \u201990s, it was after 2000 that the research effort in this field gained attention [7], [8], [9] and [10]. The main focus in the control algorithms in these papers is to handle the redundancy of the manipulator. The VMs need to fulfil a dual-task, i.e. following two trajectories: one for the manipulator and one for the mobile platform. In order to achieve this, a control algorithm is presented in [11], where the aforementioned redundancy is solved with the help of two joint dependent variables. This method makes it possible to prioritize the trajectories in situations where the system is incapable of maintaining both trajectories. In [12], the control of a robotic manipulator on a moving base is presented where the controller perceives the movements of the mobile platform as a disturbance to the system. In [9] and [13], the controllers benefit from the flatness characteristics of the system and they are validated by being implemented on custom-made mobile manipulators.\nAnother field is the modelling and shared-control of large manipulators with electric-hydraulic actuators. They are widely used in agriculture and forestry works due to their good power to weight ratio. Their modelling and control are complex tasks because of the highly nonlinear dynamics of the hydraulic cylinder. There are heuristic and data based concepts for the shared-control [14] and [15]. Mathematical models of these highly nonlinear system are presented in [16] and [17]. The downside of using these actuators is that their control is challenging. In [18], an open-loop control concept is discussed for a forestry crane that can handle unreliable sensors which is common in such vehicles. A forestry crane is presented in [16] that has an automated slewing motion controller. The modelling, control and also the implementation of the control concept on real applications are discussed. Simple modelling approaches are presented in [19] and in [20], where the hydraulic cylinder is\nmodelled as a linear transfer function, on which model-based controllers can be designed.\nNone of the above mentioned works deal with the problem how to set up a shared-control concepts if the vehicle with its motion can support the motion of the manipulator.\nThe SC concept discussed in this section was introduced and presented by the authors in detail in [4]. There, it was shown that a supporting motion of the vehicle is necessary in sudden curves, where the hydraulic manipulator cannot maintain its trajectory on its own. Here, only the key properties of the concept are shown.\nThe control model of the vehicle is a kinematic bicycle model in Frene\u0301t coordinates [2], in which the system is modelled by the following state equations:\ns\u0307v = v\n1\u2212 \u03barv \u00b7 dv cos(\u2206\u03b8v)\nd\u0307v = v sin(\u2206\u03b8v) (1)\n\u2206\u03b8\u0307v = v tan(\u03b4)\nL \u2212 s\u0307v\u03barv,\nwhere the path coordinate is sv and v is the vehicle\u2019s speed (see Fig. 2). The state dv is the distance between the reference point PR and the rear axle Pv. \u2206\u03b8\u0307v is the deviation of the vehicle\u2019s yaw angle from the tangent of the reference at the point PR. The variable \u03barv specifies the reference\u2019s curvature and the steering angle of the vehicle is \u03b4. The distance between the front and the rear axle is given by the parameter L. Describing the system as a first order dynamic system the state vector is chosen as\nx = [dv \u2206\u03b8v \u03bav \u03barv]T . (2)\nThe state \u03bav is the vehicle curvature, which is defined as \u03bav = tan \u03b4\nL . General assumptions are that the deviation \u2206\u03b8v is small and the vehicle\u2019s speed is close to the speed of the reference s\u0307v \u2248 v. These assumptions allow the linearisation of the system (1) around the equilibrium x = [0 0 0 0], which leads to the time-variant state space representation\nx\u0307 = A(t)x + Bu + Zz. (3)\n543\nAuthorized licensed use limited to: Dalhousie University. Downloaded on November 01,2020 at 19:34:14 UTC from IEEE Xplore. Restrictions apply." + ] + }, + { + "image_filename": "designv11_71_0000953_0954407020907818-Figure12-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0000953_0954407020907818-Figure12-1.png", + "caption": "Figure 12. Schematic diagram of test bench: (a) driving motor, (b) planetary gearbox, (c) vibration acceleration sensor, (d) normal gear, (e) worn teeth fault and (f) broken teeth fault.", + "texts": [ + " When Nout=100, Ns =100, and Ptrain=10%, most features of the different health conditions are separated and most features with the same health condition are gathered in the corresponding cluster. It is noted that the mapped features learned by the FCSF-based method are clustered better with lower Nout, Ns, and Ptrain values. Compared with SF, FCSF could learn more discriminative features using smaller samples. Case study 2: Fault diagnosis of planetary gear Data description. In this experiment, the gear fault datasets collected from the bearing seat at the drive end of planetary gearbox are employed to demonstrate the FCSF-based diagnosis method. As depicted in Figure 12, the test devices mainly include a driving motor, a planetary gearbox, a piezoelectric vibration acceleration sensor, a tachometer, and a data acquisition system. The rotation speed is maintained at 1500 r/ min. The vibration acquisition system collects timedomain signals under different fault conditions with a sampling frequency of 12.8 kHz. There are four health conditions in this experiment: NC, worn fault (WF), broken tooth (BT), and compound fault between WF and BT (CF). The number of samples in each health condition is 100" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0002861_978-3-030-55889-5-Figure3.20-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0002861_978-3-030-55889-5-Figure3.20-1.png", + "caption": "Fig. 3.20 CST element, a degrees of freedom, b nodal forces which generate the influence function for \u03c3xx at an internal point; the influence function is for each point x the same; A = area of the element", + "texts": [ + "123d) Only the nodes of the element with the source point x carry nodal forces ji , since the shape functions \u03d5i of the nodes lying farther off generate no section forces at the source point x . The calculated influence functions are outside the element with the source point exact, and the computer fixes the error inside the element by adding the local solution (of the GF)\u2014as in the slope deflection method. Example 3.4 Next, we turn to 2-D problems, and we begin with an CST element (constant strain triangle), see Fig. 3.20. In such elements the stresses are constant \u23a1 \u23a3 \u03c3xx \u03c3yy \u03c3xy \u23a4 \u23a6 = E 2 A \u23a1 \u23a3 y23 0 y31 0 y12 0 0 x23 0 x13 0 x21 x32 y23 x13 y31 x21 y12 \u23a4 \u23a6 \u23a1 \u23a2 \u23a2 \u23a2 \u23a2 \u23a2 \u23a2 \u23a3 u1 v1 u2 v2 u3 v3 \u23a4 \u23a5 \u23a5 \u23a5 \u23a5 \u23a5 \u23a5 \u23a6 (3.124) where xi j = xi \u2212 x j and yi j = yi \u2212 y j . 3.14 Weak and Strong Influence Functions 185 To generate the influence function for the stress \u03c3xx at a point x inside an element, we apply at each node the stresses \u03c3xx of the two unit displacement fields \u03d5i (x) of the node as equivalent nodal forces, ji = \u03c3xx (\u03d5i )(x), which means j1 = E 2 A y23 \u00b7 1 j3 = E 2 A y31 \u00b7 1 j5 = E 2 A y12 \u00b7 1 j2 = j4 = j6 = 0 " + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0002964_ijcvr.2019.098008-Figure5-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0002964_ijcvr.2019.098008-Figure5-1.png", + "caption": "Figure 5 Flow chart showing the operating sequence of the stable walking controller", + "texts": [ + " Balancing strategies together with trajectory modifications is also used to enable the robot to walk on uneven and inclined terrains. It is important to note that the trajectory modification changes the walking trajectory based on the angle of inclination of the ground. Stable walking controller basically is a normal walking controller which uses the IMUs to sense any disturbances including ground inclination changes and respond quickly to maintain the balance of the biped robot and prevent fall. The flow chart that represents the operating sequence involved in a stable walking controller is shown in Figure 5. The robot first initialises to the rest position when switched on. Then it shifts its COM slowly laterally towards one side. Then it starts to perform stationary walk. Stationary walking refers to robot shifting COM sideways and lifting its legs up similar to walking motion but without moving forward. Legs only move vertically upward and not forward. This is particularly important for push recovery from large external forces. If the forces are small, it can implement hip-ankle strategy to recover from falling" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0001895_1350650120964295-Figure1-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0001895_1350650120964295-Figure1-1.png", + "caption": "Figure 1. Schematic sketch of the constraints by Koch10 for a sphere-cone contact. The constraints are fulfilled for all indicated points, yet only PSP corresponds to the sphere\u2019s actual contact point.", + "texts": [ + " Furthermore, the intersection points do not correspond to the contact points, even though the discrepancy is low for small penetration values. The comprehensive method introduced by Koch10 (gradient method) allows geometrical approximation of roller and flange as piecewise geometric functions of spheres, cones, tori and polynomials. The contact points are determined by iteratively solving an 8 8 matrix of non-linear equations. The system of equations results out of the following requirements towards contact points P1 and P2 (e.g. for a spherecone contact in Figure 1): the gradients of the implicit surface function rFi must be collinear at the contacts points and the connecting vector between P1 and P2 must also be collinear to the gradients. Additionally, P1 and P2 must be located on the surface Si of their respective bodies. These equations contain eight unknowns: six coordinates of contact points and two scaling factors t1 and t2 resulting from the collinearity requirements. Therefore, the system of equations is determined. However, the system\u2019s nonlinearity requires an iterative calculation of contact points. On the one hand, this approach offers an accurate description of the roller and flange contact and can easily be applied to further geometrical shapes. On the other hand, it does not necessarily converge towards the correct solution (see Figure 1). The runtime for solving the system of linear equations can greatly be reduced by using the solution of the previous time step of the MBS as the initial value. A rather universal approach based on computer graphics for a very detailed geometrical description of roller, flange and contact area is given in Kiekbusch.11 The anticipated contact area is discretized into cells and the penetration of each cell is calculated using a \u2018ray tracing method\u2019 (for further reading, see for example Glassner16)", + " 23, the minimum sparsity of the Jacobian matrix is 25%. Depending on the orientation of the geometries, this value can increase significantly, especially for spheres. This can make the inversion of the Jacobian with numerically fast methods, e.g. \u2018LU-decomposition\u2019,14 difficult to impossible. In such a case, the matrix needs to be inverted using numerically less efficient but more stable algorithms, for example by calculation of the inverse via the Jacobian\u2019s adjugate matrix. The second problem, is shown in Figure 1. All constraints defined by Koch are fulfilled for PSP and ~PSP, even though ~PSP does not correspond to the correct contact point. To which points the algorithm converges highly depends on the given start value. Similar problems can be encountered for cone-torus contact pairings in the semi-analytical approach and by Gupta.7 In this case, however, only a one-dimensional function is minimized, which makes handling of false solutions due to convergence issues less complex. Thirdly, the contact between two tori may occur simultaneously at two points, if their rotational axes are coplanar and the tori overlap each other considerably (e" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0002993_msec2017-2796-Figure8-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0002993_msec2017-2796-Figure8-1.png", + "caption": "FIGURE 8: SUPPORT STRUCTURE GENERATION", + "texts": [ + " A surface roughness evaluation module was developed allowing determining the surface finish of each surface and evaluating the average roughness of the part for different build orientation using Equation 4 and 5. Figure 7 and Table 4 show the results of the developed module to evaluate surface roughness of a part according to different build orientations. 5 Copyright \u00a9 2017 ASME Downloaded From: http://proceedings.asmedigitalcollection.asme.org/ on 08/25/2017 Terms of Use: http://www.asme.org/about-asme/terms-of-use A module was developed to detect the overhanging surfaces and generate the needed support with estimating the total volume and the total contact area. Figure 8 and Table 5 shows the result of the developed module Table 6 shows the result of the developed module to evaluate the build time and cost of a part according to different build orientations as shown in Figure 8: TABLE 6: BUILD TIME AND COST Orientation 1 2 3 Material used (Kg) 0.113 0.5947 0.1284 Material for support (Kg) 0.012 0.493 0.0274 Material cost ($) 34.266 180.183 38.9 Time for support (h) 15.055 30.754 15.776 Total exposure time (h) 8.992 29.224 9.634 Total recoating time (h) 14.544 10.011 14.622 Consumed energy (KWh) 17.35 56.39 18.58 Energy cost ($) 3.32 17.354 3.346 Indirect cost ($) 78.915 199.801 158.389 Build time (h) 23.536 39.235 24.256 Build cost ($) 116.481 397.339 201.035 6. GRAPHICAL USER INTERFACE A graphical user interface was developed using MATLAB allowing the user to load the CAD model, define the load direction and set the different part requirement as presented in Figure 9 and Figure 10 The user has also the possibility to control the importance weights values and thus customize the optimization process" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0001211_012001-Figure1-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0001211_012001-Figure1-1.png", + "caption": "Figure 1. Static distribution of vertical loads along axes of fully loaded road train.", + "texts": [ + " However, the use of road trains is often limited due to inadequate haulage capacity, when it comes to the transportation of heavy machinery, equipment and weapons in tough country road conditions [5]\u2013[9]. Development of road trains with a sufficiently high level of traction dynamics is thus a pressing scientific and practical issue, which may be effectively solved by introducing an active drive of link element wheels [10]\u2013[14]. The paper presents a mathematical model of planar motion of a two-unit road train. The subject of research is a road train with 120,000 kg gross vehicle weight consisting of a four-axle tractor with electromechanical transmission (EMT) and a three-axle semitrailer. Figure 1 shows the static distribution of vertical loads. IASF-2019 IOP Conf. Series: Materials Science and Engineering 819 (2020) 012001 IOP Publishing doi:10.1088/1757-899X/819/1/012001 In the base case (road train with a passive semitrailer), each wheel drive is implemented using 60 kW traction electric motor (TEM). So, the total power of all TEMs of the tractor is 480 kW. This is the maximum power transmitted from the powertrain to TEMs. The power of active road train\u2019s TEMs has been selected so as to have equal power-to-weight ratio in all cases" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0000830_0954410020911832-Figure2-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0000830_0954410020911832-Figure2-1.png", + "caption": "Figure 2. Uncontrollable force schematic diagram.", + "texts": [ + " It is possible, by substituting these vectors into the definition of aerodynamic forces and moments, to derive the forces and moments considering the effects of atmospheric wind. Moreover, because the projectile investigated in this paper is spin stabilized, whose and are both small (usually less than 15 ) when the projectile flies stably. Therefore, it is reasonable to approximate sin , cos 1, sin and cos 1 for simplicity. Uncontrollable forces in VRF. The uncontrolled forces of projectile mainly include the drag force, lift force, Magnus force, and gravity force, as shown in Figure 2. The relative velocity vector to the wind ~vr in VRF can be calculated by equation (6), and according to equation (4) the unit vector of projectile\u2019s axis-direction in VRF can be expressed as ~x \u00bc cos cos cos sin sin T 1 T . The relative velocity vector ~vr is substituted into the vector statements of forces illustrated in McCoy.24 It is then simplified with sin , cos 1, sin , and cos 1, under the assumption that and are small. ~x and ~vr (calculated as equation (6)) are substituted into the definitions of drag and lift forces, and the approximation of drag, lift force acting on the projectile in VRF can be computed as: Drag force : FDV \u00bc 0:5 SCDvr~vr \u00bc 0:5 vrSCD v wx2 wy2 wz2 T ; Lift force : FLV \u00bc 0:5 SCL ~vr \u00f0~x ~vr\u00de \u00bc SCL 2 wy2\u00f0v wx2\u00de wz2\u00f0v wx2\u00de \u00f0v wx2\u00de 2 \u00few2 z2 wy2wz2 wy2wz2 \u00f0v wx2\u00de 2 \u00few2 y2 2 64 3 75 0 B@ \u00fe w2 y2\u00few2 z2 \u00f0v wx2\u00dewy2 \u00f0v wx2\u00dewz2 2 64 3 75 1 CA Considering the pretty high rolling speed of a guided spin-stabilized projectile, which is a great challenge for the design of control system and actuator, the actuator is usually decoupled with projectile whose rolling speed is much smaller than the main part of projectile bod" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0002878_ijmic.2020.108911-Figure2-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0002878_ijmic.2020.108911-Figure2-1.png", + "caption": "Figure 2 Unconstrained two pendulums model (see online version for colours)", + "texts": [ + " Accordingly, the current paper is interested in achieving the desired position, speed and time with respect to the motion. This goal is satisfied by designing H\u221e loop shaping robust controller, and the SLS is modelled using dual inverted pendulum. Dynamic behaviour of SLS is sophisticated compared to the movement of other humanoid gestures (Hosseini and Bazargan Lari, 2019). To simplify this system, we propose a model with two pendulums and two joints to represent the lower limb system as shown in Figure 2 with the following assumptions: \u2022 first pendulum represents the thigh while the second is the shank \u2022 first joint represents the hip, which connects pendulum 1 to basin while second joint represents the knee, which is connects pendulum 1 and 2 \u2022 walking occurs by moving hip and knee joints \u2022 ankle impact on overall motion is negligible. Accordingly, the mathematical model of the SLS will have the following equations: ( ) ( ) ( ) ( ) 1 2 2 1 2 2 1 22 1 1 2 2 1 2 2 1 2 1 2 1 1 1 3 cos 3 2 sin 2 sin 2 2 m m m l l l m l l m m gl \u03b8 \u03b8 \u03b8 \u03b8 \u03b8 \u03b8 \u03b8 \u03b8 \u03c4 + \u2212 + \u2212 + + + = (1) ( ) ( ) 2 2 1 2 1 1 22 2 2 2 2 1 2 1 1 2 2 2 2 2 cos 3 2 sin sin 2 2 m l lm l m l l m gl \u03b8 \u03b8 \u03b8 \u03b8 \u03b8 \u03b8 \u03b8 \u03b8 \u03c4 \u2212 + \u2212 \u2212 + = (2) where m1 and m2 are, respectively, the thigh and shank masses and they are 0" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0000761_s00202-020-00962-3-Figure8-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0000761_s00202-020-00962-3-Figure8-1.png", + "caption": "Fig. 8 No-load flux densities for 4 winding pole models", + "texts": [ + " All the pole ratio models are analyzed for optimal tooth thickness, and the optimal value of outer/inner stator tooth thickness is obtained for optimal torque characteristics at each pole ratio as presented in Fig. 7. As shown in Fig. 7, the optimal value of outer/inner stator tooth thickness attains the highest value at pole ratio of 23 and then starts reducing after that point. Stator winding and rotor pole numbers affect the flux linkage and flux density in the PMVM as presented in (2), and the useful flux decreases by an increasing the number of winding and rotor pole. The no-load flux densities of all the models for pole ratios ranging from 2 to 26 in 4 stator winding poles are shown in Fig. 8. According to Fig. 8, the no-load flux density in the models with lower pole ratios is higher and reduces gradually by the increase in pole ratios from 5 to 26. Cogging torque and torque ripple are the two important factors to analyze the torque performance of any machine. Cogging torque is the result of the interaction between stator teeth and permanent magnets. As indicated by (3) in Reference [23], the models with higher values ofCT should have a higher cogging torque. Table 3 shows the DDSWmodel with different pole ratios and other parameters for the model with Table 3 Cogging torque of DDSW with different pole ratios at 4 winding poles Pole ratio 2 5 8 11 14 17 20 23 26 Ss 6 12 18 24 30 36 42 48 54 Pr 8 20 32 44 56 68 80 92 104 Nc 24 60 288 264 840 612 1680 1104 2088 CT 2 4 2 4 2 4 2 4 2 Cogging torque (Nm) 5" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0000468_ro-man46459.2019.8956434-Figure1-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0000468_ro-man46459.2019.8956434-Figure1-1.png", + "caption": "Fig. 1: Kinematic Model of passive flexible needle.", + "texts": [ + " Next, simulation results and their discussions are elaborated in section V. Finally; the conclusions are made in section VI. The passive flexible needle has attained the curvature due to asymmetric bevel-tip. When we insert the needle inside the tissue, it will automatically bend since force is applied by the tissue on the sloppy tip part of the needle. If simultaneously insertion with rotation at the base of the needle is applied, then the needle attained the curve path in a different direction as shown in Fig. 1. The insertion and rotation velocity are the two inputs of the kinematic model for the passive flexible bevel-tip needle. The model is a generalized nonholonomic unicycle model. Furthermore, the torsional compliance of the needle shaft is neglected in our study. Frame P is the inertial world reference frame, and frame Q is attached to the tip of the needle. In the kinematic model, l1 is the distance from reference frame P to the needle tip which is shown in Fig. 1. Relative to the universal frame P, parameter \u03b8 is the front wheel angle which is depending upon the insertion length. The passive flexible needle has two DOFs: insertion and rotation (relative to i1, and i2 respectively see Fig. 1). The position and orientation of the connection joint relative to frame P can be described compactly by a 4\u00d74 homogeneous transformation matrix. gPQ = [ RPQ pPQ 0 1 ] \u2208 SE(3) (1) v = p\u03071i1 + p\u03072i2 (2) where i1 is the insertion velocity, i2 is the rotation velocity of the needle v1 = [0,0,1,0,k,0]T (which relates to insertion velocity) v2 = [0,0,0,0,0,1]T (which related to rotation velocity) Here, after insertion, the curvature attained by the needle is given as k = tan\u03b8 l1 . Due to the insertion input i1, the needle moves rectilinearly along the z-axis direction of its reference frame. Furthermore, the bevel-tip passive needle curves its path around the y-axis of the body reference frame, (during this insertion) which represents the orientation of the needle tip. The combined insertion and the passive curvature of the needle is depicted in Fig. 1. The spinning of the needle is the primary input to the needle which causes to follow the straight path. Insertion is the secondary input and assumes to be constant. We have not considered an issue during the needle is pulling back from the tissue region. We assume that it follows the same path during its forward insertion. In order to define the orientation between the frames P and Q, we have utilized the Z-X-Y fixed angles as generalized coordinates for parameterizing the rotation matrix RPQ. Let \u03b3 , \u03b1 and \u03b2 be the roll, pitch, and yaw of the needle in the plane, respectively" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0002099_s12206-020-1029-z-Figure5-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0002099_s12206-020-1029-z-Figure5-1.png", + "caption": "Fig. 5. 3 1T R type CDPM.", + "texts": [ + " Among the structures shown in Table 5, the 3 1T R type CDPM constrained by two moment constraint wrenches is one of promising new structures. Its motion has Sch\u00f6nflies motion which is characterized by the 3-DOF translational motion and the 1-DOF rotational motion about the fixed axis. In particular, it should be noted that numerous robots having the Sch\u00f6nflies motion have been developed and employed in various operations in the industry. Thus, in this work, the structure is implemented as a haptic device to verify its potential applications. Fig. 5 shows the conceptual schematic of the CDPM and Fig. 6 shows its practical design as the cable-driven haptic device (CDHD). In Fig. 6(a), a handgrip is connected to the serial linkage which consists of both a RR type planar limb and a ball spline. Note that two-dimensional moment constraint ( c =M \u02c6 \u02c6cx cyM x M y+ ) in Fig. 5 is realized by the serial linkage. The similar serial linkage is attached additionally to the upper part of the handgrip to reinforce the rigidity of the CDHD. And the lower end of the handgrip is connected by three cables via a passive revolute joint. The role of this revolute joint is to make the wires not rotate such that the wires are not entangled when the handgrip rotates. The pulley is attached around the lower ball spline and driven by two cables winded around. The ball spline allows the passive translational motion along its axial direction while it allows the active rotational motion about its axis to the handgrip from the pulley" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0000254_ecce.2019.8911883-Figure3-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0000254_ecce.2019.8911883-Figure3-1.png", + "caption": "Fig. 3. Flux contour plots of conventional and proposed FSPM machines: (a1) Conventional FSPM, pole ratio=3.5 (a2) Proposed FSPM, pole ratio=3.5 (b1) Conventional FSPM, pole ratio=8 (b2) Proposed FSPM, pole ratio=8 (c1) Conventional FSPM, pole ratio=17 (c2) Proposed FSPM, pole ratio=17", + "texts": [ + " In this part, the influence of pole ratio on the flux barrier effect between the proposed and conventional FSPMs will be comparatively analyzed via finite element analysis (FEA) method. 12-stator-slot FSPM machines with 14-, 16-, 17- rotor-tooth, whose pole ratios are 3.5, 8 and 17, respectively, are chosen as example, and the main parameters are listed in Table I. According to the above analysis, when the combination of stator and rotor slot is the same, the winding structure of the two machines is also the same. Fig. 3 shows the flux contour plots of the conventional and proposed FSPM machines. When at a low pole ratio, i.e. 3.5, as presented in Fig. 3 (a1) and (a2), the conventional FSPM is with a little higher flux density in the stator yoke than that of the proposed FSPM, because the \u201cflux barrier effect\u201d is poor and the usage of PMs is larger. When at a middle pole ratio, as shown in Fig. 3 (b1) and (b2), the stator yoke flux densities of the conventional and proposed FSPM machines are similar. However, when pole ratio grows to 17, as illustrated in Fig. 3 (c1) and (c2), the flux barrier effect in the conventional FSPM machine is obviously strengthened, which leads to lower stator yoke flux density than that of the proposed FSPM machine. Fig.4 compares the influences of pole ratio on the phase back-EMF of the two FSPM machines (As the pole ratio of the conventional and proposed FSPM machines is the same, the combination of slot and pole is also the same). It can be seen that the back-EMF of the proposed FSPM machine is lower than that of the conventional FSPM machine, when the pole ratio is smaller than 8" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0001975_icsse50014.2020.9219290-Figure1-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0001975_icsse50014.2020.9219290-Figure1-1.png", + "caption": "Fig. 1. Geometric diagram of tracking control of USV.", + "texts": [ + " iii) Simulation studies indicate the presented DRPIDC strategy can guarantee the superiority and robustness the tracking control of the unknown USV system. The rest of this paper is organized as: Section II gives the necessary description for the USV system. Section III focus on designing the data-driven robust PID controller. Section IV gives simulation studies of closed-loop USV system under DRPIDC scheme. Finally, Section V draws some conclusions of this paper. II. PROBLEM FORMULATION As shown in Fig.1, to facilitate describing USVs systems, the earth-fixed frame and the body-fixed frame, i.e., E-frame {E (OXY )} and B-frame {B (ox\u2032y\u2032)}, are respectively defined in the paper. Define position \u03b7 = [x, y, \u03c8] T and velocity vectors \u03c5 = [u, v, r] T with respect to the E-frame and B-frame, respectively. Authorized licensed use limited to: Auckland University of Technology. Downloaded on November 14,2020 at 22:48:06 UTC from IEEE Xplore. Restrictions apply. According to [19], the dynamics equation of a USV can be modeled by{ \u03b7\u0307 = Rm(\u03c8)\u03c5 M(t)\u03c5\u0307 = \u03c4 \u2212 C(\u03c5)\u03c5 \u2212D(\u03c5)\u03c5 \u2212 g (\u03b7, \u03c5) + d(t) (1) where \u03c4 = [\u03c41, \u03c42, \u03c43] T denotes control input, M(t) denotes the unknown time-varying inertia matrix, C(\u03c5) and D(\u03c5) are the coriolis matrix and damping matrix, respectively, both of which are considered unknown" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0001375_i2mtc43012.2020.9129496-Figure1-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0001375_i2mtc43012.2020.9129496-Figure1-1.png", + "caption": "Fig. 1. Ball-bearing. (a) Parts of the bearing. (b) Dimensions.", + "texts": [ + " Finally, Section VII contains the conclusion of this This full text paper was peer-reviewed at the direction of IEEE Instrumentation and Measurement Society prior to the acceptance and publication. 978-1-7281-4460-3/20/$31.00 \u00a92020 IEEE 1 Authorized licensed use limited to: University College London. Downloaded on July 05,2020 at 20:18:31 UTC from IEEE Xplore. Restrictions apply. paper. According to their construction and application, the shape and the elements of the bearing can be diverse. In this paper, it refers to ball-bearing as bearing, which is constituted by balls (rolling elements), inner raceway (IRW), outer raceway (ORW), and cage (see Fig. 1.a). The dimensions and the number of balls in a bearing are subject to ball diameter (Db), and pitch diameter (Dc), as depicted in Fig. 1.b. The faults can be classified as localized or distributed defects [31], [32]. These defects result in vibrations in which their magnitude is related to the grade of damage. In turn, these vibrations generate alterations in the current signal. A. Vibration Analysis An ORW fault generates small impulses when the ball is in contact with the defect, and this can be found as a spurious frequency in the vibration spectrum [32], localized where (1) suggest. Forw = Nb 2 Fr ( 1\u2212 Db Dc cos\u03b2 ) (1) where Nb is the number of the balls in the bearing, Fr is the frequency of the rotor, Db is the ball diameter, Dc is the pitch diameter of the bearing, and \u03b2 is the contact angle between the ball and the defect" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0003658_amm.592-594.1381-Figure2-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0003658_amm.592-594.1381-Figure2-1.png", + "caption": "Fig 2: Cross section of a micropolar fluid film journal bearing", + "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: 141.211.4.224, University of Michigan Library, Media Union Library, Ann Arbor, USA-11/07/15,19:01:07) of the bearing, is the geometric center of the rotor, is the geometric center of the journal is shown in Fig.1. being the offset distance between the journal center and rotor center and u being the attitude angle of the X-coordinate shown in Fig.2. By applying force of equilibrium to the rotor bearing system, the forces applied to the journal center and the equations of motion of in the Cartesian coordinates and the equations of motion of the bearing center could be written as: = + = \u2212 /2 (1) = \u2212 = \u2212 /2 (2) + + \u2212 = (3) + + \u2212 = \u2212 g (4) + + + = (5) + + + = \u2212 g + (6) Where and are the resulting damping forces in the radial and tangential directions. g is the acceleration of gravity, and are the components of the micropolar fluid film forces" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0002902_978-94-007-6046-2_43-Figure3-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0002902_978-94-007-6046-2_43-Figure3-1.png", + "caption": "Fig. 3 Early adoption of the SLIP model in robotics. (a) Planar hopper model with spring legs used to test running control algorithms for the Raibert hopper [45]. (b) Bipedal runner model of McGeer [36] for studying running stability", + "texts": [ + " In the late 1970s, Raibert had learned from Bizzi about the importance of the springy characteristics of muscles and tendons in the control of animal limb movements and embarked with the development of a computer-controlled pogo stick robot on a long and prolific carrier studying legged locomotion and robots [45]. Although the pogo stick robot was a more complicated machine, the model studied by Raibert and his colleagues to test gait control in simulation included main features of the planar SLIP (Fig. 3a), and they later analyzed the planar SLIP system more explicitly to better predict foot placements for symmetric stance phases of the machine [45]. The first formal study that analyzed running stability with a version of the SLIP model was presented by McGeer [36]. Interested, like Raibert, in building legged machines, McGeer incorporated spring leg behavior into a more realistic biped model whose legs had mass and inertia and were connected at the hip by a torsional spring (Fig. 3b). He analyzed the stride function of this model and predicted mechanical design parameters that lead to periodic running motions. Some of these motions were passively stable. For others, he showed that they could be stabilized with an LQR-derived, active feedback control on leg thrust. Since these early adoptions, the SLIP model has evolved into the main template for the design and control of running robots. Some legged robots are direct instantiations of the model. Many others actively embed spring \u2013 mass behavior using control", + " The neutral placement arose from the intuition that legged hopping and running are symmetric gaits and the symmetric leg placement approximately equals half the distance that the CoM advances in stance. While body attitude has no equivalent in the point mass SLIP, the thrust and forward speed controls of the Raibert hopper can be interpreted in this model. They provide intuitive examples for the control of system energy by changing the spring rest length in stance and of gait stability by leg placement in flight. McGeer [36] combined intuition with more formal analysis of the return map of his running model to study its control (Fig. 3b). The initial version of his model had only passive spring actuation of the legs and hip. The resulting hip oscillation created a passive swing-leg placement control. The return map of this model at liftoff (denoted with a superscript) zlo iC1 D R.zlo i ; p/; (7) had five elements in the state vector z. Parameter sets p can be found in simulation for which the model indicates steady running with zlo iC1 D zlo i D zlo . Some of these solutions have eigenvalues j of the Jacobian DzRj that lie within the unit circle (j j j < 1), which showed that the running model can be passively stable" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0001895_1350650120964295-Figure9-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0001895_1350650120964295-Figure9-1.png", + "caption": "Figure 9. Occurrence of multiple contact points for cone-torus contact.", + "texts": [ + " Second, the torus of the roller and the tori of the flange are often parallel at the initial time step of MBS. This may lead, depending on the overlap, to multiple contact points. When the roller skews in the following time steps, the contact now only occurs in the proximity of one of the previous contact points. Since the direction of skew cannot be predicted in situ at the beginning of the MBS, a good start value for the gradient method cannot be given. Therefore, correct convergence cannot be guaranteed. If the rotational axis erot;TO is normal to the cone\u2019s lateral surface (e.g. in Figure 9), two pairs of contact points also exist for torus-cone contacts. The semianalytical approach can only detect one contact, just as the approaches by Gupta7 or Koch.10 However, it is highly unlikely that two contacts occur for a torus-cone combination in a REB, since in most cases conical frustums are used, whose heights are significantly smaller than the tori\u2019s major radii. Numerical efficiency is tested by runtime comparison of the semi-analytical approach with the gradient method.10 Contact detection for a full roller-flange contact at the inner ring (Figure 2) is performed with geometrical parameters from Table 1", + " In addition to the parameters of the base figures, the inner race radius RIR, and the initial axial distance of the geometrical centres of TORO and TOFL,2 (DTO;ax) as shown in Figure 2 are given to enable performance comparison within further studies. The algorithms are implemented in the programming language C, which is commonly used for MBS. Solutions of the numerical schemes are considered sufficiently exact, if the Euclidean norm of the difference between the (exact) analytical solution and the numerical solution is less than 0.1 mm. For torus-cone and torus-torus contacts, analytical solutions can be found for special geometric arrangements. An example for such an arrangement is given in Figure 9. The inner ring is fixed in space and the outer ring rotates at 500 rpm. The roller\u2019s location and orientation is defined by superposition of kinematic rolling and a randomized sequence of skewing, tilting and translational displacement in all spatial directions. A typical minimal step size in MBS of REB (1e-6 s) is used for modelling of a time span of 100ms. Small step sizes lead to small changes of the contact location within a time step and are beneficial for the gradient method,10 since the initial value of the numerical schema is close to the solution" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0003805_02678292.2015.1006153-Figure3-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0003805_02678292.2015.1006153-Figure3-1.png", + "caption": "Figure 3. (a) Schematic of a +1/2 disclination loop attached to a +1 line and director field in the cross-sectional plane in a conical capillary. Disclination bending stiffness arises when \u2202 ?n=\u2202sj1 0 and/or \u2202 ?n=\u2202sj2 0 (adapted from ref. 10); (b) schematic of the coordinate system and geometry of two m = +1/2 lines emanating from the branch point at x = xb, y = 0. The x-axis is along the capillary axis. At the centre of the loop, the capillary radius is R0, which decreases by moving to the left and increases by moving to the right. The angle between the x-axis and the tangent vector t is \u03c6, and N is the unit normal. Regarding loop log radiuses, XB1 and XB2, the branch point angles, \u03c61 and \u03c62, are not equal but both are around 60\u00b0. y0 is the loop short radius which is the half-separation distance between the two +1/2 lines. \u03c9b1 and \u03c9b2 are the branch point velocities. At the top and the bottom of the loop, (x0 \u00b1 y0), \u03ba = 0 and at the BP, (XB1,0) and (XB2,0), \u03ba = \u03ba* (adapted from Ref. [10]).", + "texts": [ + "[1] For NLCs, one can use a traceless tensor order parameter Q = \u039b(nn-I/3) + P(mm-ll), where (\u039b, P) are the uniaxial and biaxial order parameters and (n, m, l) are the orthonormal directors.[1] The singular defect-associated order parameter heterogeneities result in the singular core energy per unit volume. When the disclination line is curved, additional distortions arise because the normal planes to the line are splayed and this splay contributes to the line energy in the form of a bending energy that is of the same order of magnitude as tension energy when the director curvature and the line curvature are of the same magnitude. Figure 3a shows the orientation distortions associated with a curved +1/2 disclination line. Orientation gradients on a plane perpendicular to the line ?n\u00f0 \u00de contain splay-bend distortion and generate line tension. The additional orientation gradient distortions along the arc length d ?n=ds\u00f0 \u00de generate bending elastic storage. Along the arc length, the bending stiffness is not constant, resulting in asymmetric loops. Figure 3b shows the schematic of a +1/2 disclination loop attached to +1 line segments in a cone, relevant to this paper. The +1 straight and two +1/2 curved disclinations meet at the BP. For a disclination branching process in a cylinder,[7] the aperture angle between the two 1/2 disclinations at the BP is \u03c61 = \u03c62 \u2248 60\u00b0, resulting in an elliptic lens loop. For a cone, we show (Appendix) that these angles can be different, \u03c61 \u03c62 and more complex ovoidal loop shapes emerge. The disclination branching process that mediates the PP and PR textures is characterised by (i) the branch aperture angle, (ii) the shape and curvature of the m = +1/2 lines and (iii) the characteristic branching region thicknesses XB1 and XB2" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0003802_icolim.2014.6934343-Figure1-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0003802_icolim.2014.6934343-Figure1-1.png", + "caption": "Fig 1. Geometry of the basket of the IA V. The lengths displayed are in meters.", + "texts": [ + " In this paper we first introduce the three geometric configurations we have used and then we analize the results of each configuration. Finally we critically review the results and derive some best practice procedures. III. TEST SETUP We have considered and built a representative geometry of a basket of an IA V: this is made of a box that is I meter by 2 meters with a thickness of 10 cm. The IA V has a banister with a height of 1 meter and has a square section of 4 cm. A sketch of the basket is depicted in Fig 1. L. Barbieri et al. \u2022 Insulated aerial vehicles and high voltage live working in Italy: tests and studies to assess the safety aspects We have considered both a single bundle and a two bundle configuration. The first one is depicted in Fig 2. The bundle is made of three cilindric conductors each one placed on a vertex of an equilateral triangle with edges of 40 cm. Each conductor has a diameter of 4 cm. The bundle is 12 meter long. Each end is connected to a large sphere with a diameter of 2 meters to avoid any boundary effect" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0002830_978-94-007-6046-2_42-Figure2-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0002830_978-94-007-6046-2_42-Figure2-1.png", + "caption": "Fig. 2 A walking machine of minimum configuration by D.C. Witt [19]. (a) A prototype mechanism for powered lower-limb prostheses. (b) Model for the lateral motion", + "texts": [ + " By this feature, the LIP Model has been used to analyze the dynamics of biped walking, to create biped gaits, and to design their controllers as described later. In this section we overview some of the important studies that inspired the linear inverted pendulum. Since an inverted pendulum is a natural choice to model a biped walking robot, it has been used from the very beginning of the robotics research. One of the pioneering works was done by D.C. Witt in 1968 [19]. He conceived a prototype mechanism for a powered lower-limb prostheses as Fig. 2a. To analyze its dynamics, he assumed that all mass is concentrated on the body and the leg mass is negligible. Moreover, the equation of motion was derived based on the small-angle approximations, i.e., on the assumption that all the motion does not largely deviate from the vertical axis of gravity. Figure 2b is the model to analyze the lateral motion, whose dynamics was expressed as R D \u02db2. F 0/; (3) where is the lateral inclination of the robot, \u02db2 g=h, and is a parameter determined from the lateral distance between the left and the right feet. The variable F 0 takes C1 or 1 depending on the leg support. Then, Witt introduced the following \u201cnormal coordinates\u201d to get further simplification: y1 C \u02db 1 P (4) y2 \u02db 1 P (5) By substituting (3) into above equations, he derived Py1 D \u02db.y1 C F 0/; (6) Py2 D \u02db", + " Although the magnitude of u1 is strictly limited by the foot size, it is useful to compensate model errors and disturbance which must be concerned in the real walking robot control. Figure 7 shows walking simulation of a birdlike biped with leg mass based on a real hardware. The leg mass was treated as modeling error, and the ideal LIPM trajectory was followed by compensation control implemented by the ankle torque [7]. As previously mentioned, inverted pendulum models relies on an assumption of negligible leg mass with respect to the body. As Witt illustrated in Fig. 2a, if a robot carries its power supply, sensors, and controllers on its body, the leg mass becomes less and less relevant with respect to the center of mass dynamics. However, the technologies were premature to realize such walking machine in the 1970s and 1980s at which actuators were too weak and other instruments were too heavy. This problem was solved around 1990s, and that triggered later prolific research activities on biped walking robots. The 3D version of linear inverted pendulum mode (3D-LIPM) was derived by Hara, Yokogawa, and Sadao [4] for point contact foot" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0001701_0954405420949757-Figure4-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0001701_0954405420949757-Figure4-1.png", + "caption": "Figure 4. Unit tangent vectors of discrete point.", + "texts": [ + " , n0i and n00i be unit normal vectors of convex side and concave side at P0i and P00i . Unit normal vectors of face cone at P0i and P00i are denoted by c0i and c00i , which are calculated as c0i = f0i 3 (r03 f0i)= f 0 i 3 (r03 f0i)j j c00= f00i 3 (r03 f00i)= f 00 i 3 (r03 f00i)j j \u00f015\u00de Let t0i and t00i be unit tangent vectors of tooth crest line at P0i and P00i , which are calculated as follow t0i = c0i 3 n0i= c 0 i 3 n0ij j t00i = c00i 3 n00i= c 00 i 3 n00ij j \u00f016\u00de c0i, c 00 i , n 0 i, n 00 i , t 0 i and t00i are shown in Figure 4. Upper and lower edges calculation for chamfering Auxiliary circles for edges calculating Let the plane through P0i perpendicular to t0i be planeP0 i, the plane through P00i perpendicular to t00i be planeP00 i . Auxiliary circle R0i is drawn on P0 i with P0i as its center and r0i as its radius, and auxiliary circle R00i is drawn on P00 i with P00i as its center and r00i as its radius, as shown in Figure 5. Let the intersection of R0i and convex side of tooth surface be S0i, R 0 i and face cone be A0i,R 00 i and concave side of tooth surface be S00i , R 00 i and face cone be A00i " + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0000715_morse48060.2019.8998749-Figure1-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0000715_morse48060.2019.8998749-Figure1-1.png", + "caption": "Fig. 1. Mechanical model and vector diagram of the robot.", + "texts": [ + " After that, the simulation results and the numerous comparisons are illustrated to verify the accuracy and effectiveness of the proposed controller and its superiority to the traditional DSC controller. In the last section, the conclusion is given. II. MATHEMATICAL MODEL In this section, the kinematic and dynamic model of the robot would be constructed. The pistons move vertically simultaneously at the same time with the rotation of the rotating shaft attached below the lower panel could perform all types of robot movements. Fig.1 illustrates the car driving simulator model with the movements that are translation along the OZ axis , rotation along the vertical OZ, rotation along OX axis, and rotation along the OY axis and they are defined by zP , \u03b3 , \u03b1 , and \u03b2 . Assuming that [ ], , , T zp p \u03b1 \u03b2 \u03b3= is known and the forward kinematic equation of the robot is to compute [ ]1 2 3 T q l l l \u03b3= with il is the length of the leg i ( 1,2,3)i = and \u03b3 is the rotation angle around OZ axis. The global coordinate and kinematic parameters are demonstrated in the vector diagram in Fig. 1 with a is the radius of an upper panel, b is the radius of a lower panel, and c is a distance between O and AO . As show Fig. 1, a loop-closure equation can be written for each limb of 4- DOF car motion simulator: i i i A i AA B OP PB O A OO= + \u2212 \u2212 (1) with 1, 2,3i = . The coordinates of A1, A2, A3: 1 *sin( ) *cos( ) 0 6 6 T A a a \u03c0 \u03c0 = , [ ]2 0 0 T A a= \u2212 , and [ ]3 0 0 T A a= \u2212 . And the coordinates of B1, B2, B3: 1 *sin( ) *cos( ) 0 6 6 T B a a \u03c0 \u03c0 = , [ ]2 0 0 T B a= \u2212 , 3 *sin( ) *cos( ) 0 6 6 T B a a \u03c0 \u03c0 = \u2212 . Besides, a center of the lower panel and a center of the upper panel are [ ]0 0 T A zO a= and [ ]0 0 T zP p= , respectively" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0003915_amm.607.405-Figure1-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0003915_amm.607.405-Figure1-1.png", + "caption": "Fig. 1 Dynamic finite element analysis models", + "texts": [ + " Then we simulate its natural frequency and modal shape, compare them with the experimental results, two shows good agreement. Finite element analysis model The marine gearbox works along two lines: forward of two levels of transmission, reverse of three levels of transmission. The parameters of gear system are shown in table 1. Using ANSYS software, we respectively establish the dynamic finite element analysis models of the marine gearbox's housing and the gear-shaft-bearing-housing coupled system. The finite element model are shown in Fig 1. The housing model has 86 418 nodes, 317 154 elements. The gear system model has 130 666 nodes, 476 902 elements, including 76 truss elements and 56 spring elements. 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: 128.210.126.199, Purdue University Libraries, West Lafayette, USA-08/07/15,09:33:40) From the modal analysis results, we can optimize the gearbox's structure and avoid its inherent frequency" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0001543_isie45063.2020.9152461-Figure2-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0001543_isie45063.2020.9152461-Figure2-1.png", + "caption": "Fig. 2. A BDFRG phasor diagram with key vectors and reference frames.", + "texts": [ + " Consequently, the proposed method can provide much higher accuracy and fewer parameter dependence than the MRAS observers reported in [18], [19]. The fundamental angular position and velocity relationships for the BDFRG are as follows [10], [11], [17]: \u03b8rm = \u03b8p + \u03b8s pp + ps \u21d2 \u03c9rm = \u03c9p + \u03c9s pp + ps = \u03c9p pr \ufe38\ufe37\ufe37\ufe38 \u03c9syn (1 + \u03c9s \u03c9p ) (1) where pp and ps are primary and secondary pole-pairs, respectively, pr = pp + ps is the number of the rotor poles, whereas the definitions of \u03b8p and \u03b8s angles are self-evident from Fig. 2. The rotor velocity (\u03c9rm) is locked with the applied frequencies to the windings (i.e. \u03c9p and \u03c9s) meaning that the BDFRG is essentially an electrically excited synchronous machine. Indeed, the synchronous speed (\u03c9syn) is achieved when the secondary winding is DC (i.e \u03c9s = 0). However, unlike a classical synchronous or induction pr-pole machine, \u03c9syn of the BDFRG with pr rotor poles is halved according to (1) making it a medium-speed in nature. Therefore, a more compact, two-stage gear-set option would be feasible for wind turbines in place of a vulnerable three-stage companion used for DFIGs [2]", + " The variable speed range of a BDFRG wind turbine around \u03c9syn i.e. [\u03c9min, \u03c9max] can be defined by the following ratio: r = \u03c9max \u03c9min = \u03c9p + \u03c9s \u03c9p \u2212 \u03c9s =\u21d2 \u03c9s \u03c9p = r \u2212 1 r + 1 (3) For a usual r = 2, the secondary frequency is limited to \u03c9s = \u03c9p/3. As a result, Ps \u2248 0.25Pm from (2), which implies a partially rated power converter by analogy to DFIG. Further details about the BDFRG operation can be found in [4], [5]. III. VOLTAGE ORIENTED POWER CONTROL The space-vector dynamic model of the BDFRM in rotating reference frames (Fig. 2) can be formulated as [10], [11], [17]: up = Rpip + d\u03bbp dt + j\u03c9p\u03bbp (4) us = Rsis + d\u03bbs dt + j\u03c9s\u03bbs (5) \u03bbp = Lpip + Lmi\u2217sm (6) = Lpipd + Lmismd \ufe38 \ufe37\ufe37 \ufe38 \u03bbpd + j (Lpipq \u2212 Lmismq ) \ufe38 \ufe37\ufe37 \ufe38 \u03bbpq (7) \u03bbs = Lsis + Lmi\u2217pm = \u03c3Lsis + Lm Lp \u03bb\u2217 p \ufe38 \ufe37\ufe37 \ufe38 \u03bbm (8) = \u03c3Lsisd + \u03bbmd \ufe38 \ufe37\ufe37 \ufe38 \u03bbsd + j (\u03c3Lsisq + \u03bbmq) \ufe38 \ufe37\ufe37 \ufe38 \u03bbsq (9) 1518 Authorized licensed use limited to: Carleton University. Downloaded on August 04,2020 at 10:42:49 UTC from IEEE Xplore. Restrictions apply. where Lm is the magnetizing inductance, Lp and Ls are the 3- phase self-inductances of the primary and secondary windings, respectively, \u03c3 = 1 \u2212 L2 m/(LpLs) is the leakage coefficient, and \u03bbm is the mutual flux linkage magnitude [4]. By aligning the qp-axis with the line voltage vector as shown in Fig. 2, upq = up, upd = 0 and the primary apparent power expression can be simplified as: Sp = Pp + jQp = 1.5upi \u2217 p = 1.5jup(ipd \u2212 jipq), so that the machine side converter (MSC) controlled real and reactive powers injected to the grid (Fig. 1) become: Pp = 1.5upipq and Qp = 1.5upipd. Given that \u03bbpd \u226b \u03bbpq and applying (4)-(9), these power relationships can be further expanded into more insightful exact (VOC) and approximate (FOC) forms: Pp = 3 2 \u03bbpq + Lmisq Lp up \u2248 3 2 Lm Lp upisq (10) Qp = 3 2 \u03bbpd \u2212 Lmisd Lp up \u2248 3 2 ( u2 p \u03c9pLp \u2212 Lm Lp upisd) (11) The nearly decoupled Pp and Qp control in the \u03c9s rotating ds\u2212qs frame (Fig. 2) can be achieved by varying the respective secondary current components, isq and isd, in a closed-loop fashion as illustrated in Fig. 3. This can be effectively done by using appropriately tuned PI regulators without having to know the machine inductances considering that Pp \u221d isq and Qp \u221d isd with both \u03c9p and up being constant by the primary winding direct grid connection. The main idea of a MRAS observer is to quantify a desired variable (e.g. rotor speed) by comparing the values of a suitably selected quantity (e", + " secondary current vector) obtained from reference and adaptive models. The former should be independent of the estimated variable in question, and the latter dependent on it. To get accurate estimates, an adaptive strategy of some kind (e.g. PI controller) must be applied to eliminate the differences between the two models\u2019 outputs through a stable fast converging iterative process [20]. The secondary-current based MRAS observer is depicted in Fig. 3. The reference model outputs are the stationary \u03b1\u2212\u03b2 frame secondary currents (Fig. 2) calculated from measurements using the following expression for a Y-connected winding with an isolated neutral and a positive phase sequence: iss = is\u03b1 + jis\u03b2 = isa + j(isa + 2isb)/ \u221a 3 (12) The \u03b1 \u2212 \u03b2 frame counterparts of the primary voltage and flux equations, (4) and (6), can be written as follows: ups = up\u03b1 + jup\u03b2 = Rpips + d\u03bbps dt (13) \u03bbps = \u03bbp\u03b1 + j\u03bbp\u03b2 = Lpips + Lmi\u2217sse j\u03b8r (14) The frequency modulated secondary current vector in (14), i\u2217sm = i\u2217sse j\u03b8r , rotating at \u03c9p (Fig. 2) is a function of the rotor \u2018electrical\u2019 position (\u03b8r = pr\u03b8rm). Therefore, this term is estimated by the adaptive model using the measured line voltages and currents in the following way: i\u0302 \u2217 ss = i\u0302s\u03b1 \u2212 ji\u0302s\u03b2 = \u03bbps \u2212 Lpips Lm e\u2212j\u03b8r (15) i\u0302s\u03b1 = \u03bbp\u03b1 \u2212 Lpip\u03b1 Lm cos \u03b8r + \u03bbp\u03b2 \u2212 Lpip\u03b2 Lm sin \u03b8r (16) i\u0302s\u03b2 = \u03bbp\u03b1 \u2212 Lpip\u03b1 Lm sin \u03b8r \u2212 \u03bbp\u03b2 \u2212 Lpip\u03b2 Lm cos \u03b8r (17) \u03bbps = \u03bbp\u03b1 + j\u03bbp\u03b2 = \u222b (ups \u2212Rpips )dt (18) where ip\u03b1 and ip\u03b2 come from (12) but for the primary phase current measurements, and the corresponding up\u03b1 and up\u03b2 in terms of the line-to-line voltages are given as: ups = up\u03b1 + jup\u03b2 = uab + uac 3 + j ubc\u221a 3 (19) 1519 Authorized licensed use limited to: Carleton University" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0002062_j.cma.2020.113498-Figure2-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0002062_j.cma.2020.113498-Figure2-1.png", + "caption": "Fig. 2. Schematic representation of the physical model. A sphere of diameter D is confined by a prismatic enclosure of dimensions D \u00d7 4D \u00d7 6D. The sphere oscillates in the z direction with amplitude A.", + "texts": [ + " Note that after completing the pre-computing stage, the time integration of the current methodology is based on a double implementation of the original generic driver (of either the first or the second type), namely, for the first time when calculating H\u22121RHSn\u22121,n and for the second time when calculating H\u22121[RH Sn\u22121,n \u2212 RF] products exactly the same procedure as that for any fully explicit formulation of the direct forcing IBM. As a result, the developed semi-implicit methodology not only provides a more accurate imposition of the kinematic constraints of no-slip on the surfaces of the periodically oscillating immersed bodies, but it is also as time efficient as its fully explicit counterpart. 4. Results and discussion 4.1. Flow around a transversely oscillating sphere An oscillating sphere of diameter D in an otherwise quiescent fluid confined by a rectangular prism of dimensions 4D \u00d7 4D \u00d7 6D is considered (see Fig. 2). The sphere oscillates in the z direction with periodic velocity Uz given by: Uz = Umax sin(\u03c9T ), (15) where \u03c9 is the angular oscillating frequency, and T is dimensional time. By setting the location of the center of the coordinates in the middle of the prism, the vertical coordinate of the center of the sphere can be obtained directly by integrating Eq. (15) over time to yield: Z = \u2212 Umax \u03c9 cos(\u03c9T ). (16) o-slip boundary conditions are applied to the surface of the sphere and all the walls of the computational domain" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0002613_ieeeconf38699.2020.9389062-Figure2-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0002613_ieeeconf38699.2020.9389062-Figure2-1.png", + "caption": "Fig. 2. The improved artificial potential field method.", + "texts": [], + "surrounding_texts": [ + "Keywords\u2014multi-UUVs, obstacle avoidance, multi-beam sonars, artificial potential field method\nI. INTRODUCTION\nArtificial potential field method is the main method of obstacle avoiding for multiple UUV formations [1-3]. The traditional method always builds artificial potential field by making a circle, the avoidance field, centered on the center of obstacle. But according to the motion characteristics of the UUV, the longitudinal velocity is much greater than the lateral velocity. The traditional artificial potential field method may cause a waste of obstacle avoidance space. In this paper, an improved artificial potential field method, based on optimizing obstacle avoidance space, is proposed to solve the control problem of collaborative obstacle avoidance for multiple UUV formations.\nII. THE IMPROVED ARTIFICIAL POTENIAL FIELD OF METHOD\nA. The Smallest Circular Area Restricted by the limitation of sight angle and detection distance, the multi-beam sonar carried by the single UUV cannot obtain all of obstacle contour information. This section proposes a method to identify static obstacle by combining the information getting from forward-looking sonars of multi-UUVs in the formation.\nThe obstacle contour information, gained by the forwardlooking sonar from the UUV, is prepared to facilitate the application of the improved artificial potential field method below. Here we need to get the coordinate information of the three contour extreme points from the obstacle, as\n, and . Then we can get a smallest circular area which can cover the obstacle with three contour extreme points , and . Set the radius of the circular area as , and the center point of\nas , then\n(1)\nwhere is the Euclidean distance. When is the largest distance, then\n(2)\nThe refers to the smallest circular area which includes the full projections of obstacle and UUV, here the two projections are tangent. Set as an arbitrary point in the projection of obstacle, then for any point , we have . Based on the improved artificial potential field method, we proposed the rule of obstacle avoidance with the conception of the smallest circular area .\nFig. 1. The obstacles are disscoeried by multi-UUVs .\nB. The Improved Artificial Potential Field Method Aiming at the shortcomings of the traditional artificial potential field method, this paper proposes an improved artificial potential field method through optimizing the obstacle avoidance space. This method is proposed based on that the UUV can only detect the extreme feature points, the contour information, of the side facing the UUV formation.\nBy constructing an improved artificial potential field function, the obstacle avoidance potential field space is optimized. The distinguishing features of the method are:\n(1) Use the smallest circular area to include the obstacle area.\n(2) Optimized the obstacle potential field action space.\n(3) The reference obstacle avoidance point dynamically moves on the boundary of the smallest circle of coverage.\n1 obtp 2 obtp 3 obtp obtW\n1 1 1( , )obt obt obtp x y 2 2 2( , )obt obt obtp x y 3 3 3( , )obt obt obtp x y\nobtW obtR\nobtW ( , )obt obt obtO x y\n0 1 2 2 3 3 1max{ ( , ) / 2, ( , ) / 2, ( , ) / 2}obt obt obt obt obt obt btR p p p p p pr r r=\n( )r \u00d7 ( , )obt obt i jp pr\n( ) ( )1 1, 2 2 obt obt obt obt obt i j obt i jx x x y y y= + = +\nobtW\nbp { , }b b bp x y=\nb obtp \u00ceW\nobtW\ntxq\n1 obtp\n2 obtp\n3 obtp\nobtO\n{ }E\nobtR\nobtW\nx\ny\n2lx\n1lx\n978-1-7281-5446-6/20/$31.00 \u00a92020 IEEE\nGl ob\nal O\nce an\ns 2 02\n0: S\nin ga\npo re\n- U\n.S . G\nul f C\noa st\n| 9\n78 -1\n-7 28\n1- 54\n46 -6\n/2 0/\n$3 1.\n00 \u00a9\n20 20\nIE EE\n| D\nO I:\n10 .1\n10 9/\nIE EE\nCO N\nF3 86\n99 .2\n02 0.\n93 89\n06 2\nAuthorized licensed use limited to: Carleton University. Downloaded on June 19,2021 at 15:57:47 UTC from IEEE Xplore. Restrictions apply.", + "(4) The volume and direction of the repulsion force from obstacle to UUV are changing with the movement of UUV and the reference obstacle avoidance point.\nAs shown in Fig.1, refers to the smallest circular area, means the reference obstacle point which is moving on the boundary of over time. ,\n, are three contour extreme points from the obstacle. And is the center point coordinate of\n, is the coordinate of the target point. represents the coordinate of UUVs, and there must be for any . Here, we also have as the force of the target point, as the force of reference obstacle point and the force from other UUVs. Set\n, then we have the coordinate of obstacle is in the coordinate system, where is\nthe distance between UUV and . So we can get the coordinate of each contour extreme point in the coordinate system as below,\n(3)\nTake , into consideration, the coordinate of reference obstacle point at time is obtained,\n(4)\nTo optimize the potential space, we put forward the potential function of the dynamic reference obstacle point based on the artificial potential field method. Set the distance vector between UUV and dynamic boundary point as , for any time , we can get the potential function of the position as .\nDesign the potential function of obstacle and target point respectively as:\n(1) The potential function of the reference obstacle point is\n(5)\nwhere are coefficients of the potential function, and is the radius of action of potential field.\n(2) The potential function of the target point is\n(6)\nwhere are the positive coefficients of potential function.\nSo, for the ith UUV, the total potential field is\n(7)\nThen we can get the gradient function of the obstacle.\n(8)\nwhere the gradient functions along the axis are:\n(9)\n(10)\nAs the potential functions are based on distance, then the gradient function is\n(11)\nwhere is the unit pointing vector from UUV to the target point.\n(12)\n(13)\nThrough the analysis of Eq. (8) and (9), it can be seen that an elliptical potential field area is built based on the smallest covering circle domain, which optimizes the potential field space. The area increases the axial distance between the obstacle and UUV, it also reduces the vertical distance in the UUV action direction. Besides, an advantage of is that the potential field function gained by the UUV is bounded when the target point is far, and the gradient function applied on the UUV changes slightly, which can be regarded as constant.\nIII. COORDINATION CONTROL FOR OBSTACLE AVOIDANCE Considering about the virtual external force of UUV in the potential field, the control input of UUV can be divided into repulsion control input , gravitation control input and formation coordination control input . Then the control input of the ith UUV is:\n(14)\nobtW\n( )( )= ( ), ( )tp t x t y t\nobtW 1 obtp\n2 obtp 3 obtp\n( , )obt obt obtO x y\nobtW ( , )tar tar tarO x y ( )lp t\n( )l obtp t \u00cfW 0t \u00b3 ( )tF t\n( )oF t\n( )inF t\n=[cos ,sin ]Ti iq qS { } { }[ , ]B B T s i i ix y = \u00d7 Sr { }B s\nir obt ip\n{ }E\n{ }\n{ } cos sin sin cos obt B lji i\nobt B lji i\nxx x yy y q q q q \u00e9 \u00f9 \u00e9 \u00f9 \u00e9 \u00f9\u00e9 \u00f9 = +\u00ea \u00fa \u00ea \u00fa \u00ea \u00fa\u00ea \u00fa-\u00eb \u00fb \u00eb \u00fb\u00eb \u00fb \u00eb \u00fb\nobtO obtR\n( )tp t t\n( )( ) ( ) l obt\nt obt obt l obt\np t Op t O R p t O - = + -\n( ) ( )l tt t= -!l p p t [ ]( ) ( ), ( ) Tl l lt x t y t=p\n( ) ( ) ( )2 2 2 2 2 ( ) ( ) ( ) ( ) exp ( )\n0\nl t l t\nrep t\nx t x t y t y t l d\nU t\nl d\ng a b \u00ec \u00ec \u00fc- -\u00ef \u00ef\u00ef - - \u00a3\u00ef \u00ed \u00fd= \u00ed \u00ef \u00ef\u00ee \u00fe \u00ef\n>\u00ef\u00ee\n!\n!\np\n, ,g a b d\n( ) 2( ) ( ( ), ) ( ( ), )att l l tar\nl tar\nU t t t k kr l r l = + + p p O p O\n,l k\n( ( )) ( ( ))att l rep lU U t U t= +p p\n( ) ( ) ( )( ) ( ) ( )\n0\nrep l rep l\nrep l\nU t U t l d\nU t x y\nl d\n\u00ec \u00b6 \u00b6\u00e9 \u00f9 \u00a3\u00ef\u00ea \u00fa\u00d1 = \u00b6 \u00b6\u00ed\u00eb \u00fb\n\u00ef >\u00ee\n! ! !\n!\np p p\n,x y\n( ) ( )\n2\n2\n2 2\n2 2\n( ( )) 2 ( ( ) ( ))\n( ) ( ) ( ) ( ) exp\nrep l l t\nl t l t\nU t x t x t x\nx t x t y t y t\ng a\na b\n\u00b6 - = -\n\u00b6 \u00ec \u00fc- -\u00ef \u00ef\u00d7 - -\u00ed \u00fd \u00ef \u00ef\u00ee \u00fe !\np\n( ) ( )\n2\n2\n2 2\n2 2\n( ( )) 2 ( ( ) ( ))\n( ) ( ) ( ) ( ) exp\nrep l l t\nl t l t\nU t y t y t y\nx t x t y t y t\ng b\na b\n\u00b6 - = -\n\u00b6\n\u00ec \u00fc- -\u00ef \u00ef\u00d7 - -\u00ed \u00fd \u00ef \u00ef\u00ee \u00fe\n!\np\n( ( ))( ( )) att l att l att U tU t r \u00b6 \u00d1 = \u00b6 !pp n\natt !n\n( )22 ( ( )) ( ( ), ) att l\nl tar\nU t t k k r l r l \u00b6 = - \u00b6 + p p O\n( )22 ( ( )) ( ( ), ) att l\nl tar\nU t t k k r l r l \u00b6 = - \u00b6 + p p O\ntarU\n( )o tu ( )t tu ( )in tu\n( ) ( ) ( ) ( )i i i i t o int t t t= + +u u u u\nAuthorized licensed use limited to: Carleton University. Downloaded on June 19,2021 at 15:57:47 UTC from IEEE Xplore. Restrictions apply.", + "Here we design the negative gradient of the potential function of the target point as , and use the negative gradient of the potential function of UUV as , that is\n(15)\n(16)\nThe coordination control input that ensure the formation stable convergence between UUVs is\n(17)\nwhere and are the control gain of the formation coordination control input.\nTake the feedback linearized Second-order integration model,\n(18)\nWhen UUVs are getting into the potential field area, the UUV system would shift as:\n(19)\nwhere , \uff0cand is the coefficient of influence of the current position on the state of UUV.\nAssume: When the formation is facing an obstacle, after building the obstacle covering area through the outline information of the obstacle. The UUV formation would get diverse obstacle avoidance routes due to different directions of the virtual force, based on the potential field effect of the dynamic obstacle reference point on the boundary of , and the space position of the potential field where the UUV is located. That is, the UUV formation will divide into smallscale subsystems according to different potential fields.\nWe define the generalized energy functions of the two subsystems as below:\n(20)\n(21)\nwhere and represent the number of UUVs in each of two subsystems, and .\nThe total potential energy functions of the two subsystems are as:\n(22)\n(23)\nTheorem: For the UUV formation with the targeted linearized model, the improved obstacle potential field function and the target point potential field function are effective in the potential field environment constructed by the coordinated controller and improved artificial potential field method. The initial generalized energy function of the system is bounded; are sufficient. Then, the formation can be transformed autonomously and the following conclusion holds: the subsystems remain convergent with a transformation of the formation, as\nProof: Take the derivative of the generalized energy function of the subsystem as follow:\n(24)\nExpand the function we can get:\n(25)\nDerivative the total potential energy function of the subsystem based on its definition:\n(26)\nAnd we have the following function after substituting:\n(27)\nThen the following inequality holds:\n(28)\nWe can write it as:\n(29)\nwhere .\nAccording to the characteristics of Laplacian matrix, must be a semi-definite. Then we derive , and the formation would be stable convergence.\nIV. SIMULATION AND RESULT In this simulation, assume that there are 4 UUVs in a formation, and set the initial positions as , , where , the initial velocity are , then set the potential coefficients as , set the coordinate of obstacle is , . Set the coordinate of target point as , , . And the interval between\n( )i t tu\n( )i o tu\n( ) ( ( ))i t att lt U t= -\u00d1u p\n( ) ( ( ))i o rep lt U t= -\u00d1u p\n( )i in tu\n( ) ( )( ) ( ) ( ) ( ) x v i i i in i in p ij j i v ij j i j N j N t t a t b tw \u00ce \u00ce\n\u00e9 \u00f9 = - + - + -\u00ea \u00fa\n\u00ea \u00fa\u00eb \u00fb \u00e5 \u00e5u x K K x x K v v\ninK w\n( ) ( ) ( ) ( ) i i i i t t t t = = ! ! x v v u\n( )\n( )\n( ) ( ) ( ) ( ( )) ( ( )) ( )\n( ) ( ) ( )\n( ) ( ) ( )\nx i\nv i\ni i\ni att i rep i i\np in ij j i j N\nv in ij j i j N\nt t t U t U t t\na t t t\nb t t t\nw\n\u00ce\n\u00ce\n= = -\u00d1 -\u00d1 -\n+ -\n+ -\n\u00e5\n\u00e5\n!\n! x v v x x x\nK x x\nK v v\np in in p=K K K v in in v=K K K w\n0W\n0W\n1 1\n1 1 1 1\n( ( ), ( )) ( ) ( ) 2 ( ( )) ( ) ( ) m m\nT T i i i i i i Q t t t t t t tw = = = + Y +\u00e5 \u00e5v v x xz V z\n2 2\n2 2 1 1\n( ( ), ( )) ( ) ( ) 2 ( ( )) ( ) ( ) m m\nT T j j j j j j Q t t t t t t tw = = = + Y +\u00e5 \u00e5v v x xz V z\n1m 2m\n1 2m m m+ =\n( ) 1\n1 1\n( ( )) ( ( )) ( ( )) m\ni i rep att i t U t U t = Y = +\u00e5 x xz\n( ) 2\n2 1\n( ( )) ( ( )) ( ( )) m\ni i rep att i t U t U t = Y = +\u00e5 x xz\n0 ( ( ), ( ))Q t tz V\n1 2( ( ), ( )) 0, ( ( ), ( )) 0Q t t Q t t< 2) can be obtained from the following equation: move too fast. In this paper, we set Vmax = 5 mm/s [4]. The adaptive A-causality control uses the deadline of each position information, and the deadline is set to the generation time of the position information + fj. (> 0) seconds. If the information is received by the time , it is The two systems explained in Section 2.1 are used to conduct cooperative work of carrying an object together (see Figure 2 (a\u00bb. Since the robot arms of the systems need to carry the object at the same timing, the position information of the robot arm in each system is transmitted to the other system . Then , the robot arm of which operation timing is delayed is controlled by using the information ; that is, we can advance the operation timing of the robot arm which is delayed. When the network delay between the systems is negligibly small , even if the position of the robot arm in one system changes, the position information is immediately transmitted to the other system", + " In the experiment, in order to move the robot arms in almost the same way always, we carry out work of pushing the top block piled up front and behind the initial position of the wooden stick held by the two toggle clamp hands of the robot arms for about 15 seconds (it takes about 5 seconds to drop the front block, and around 10 seconds 96 Authorized licensed use limited to: University of Gothenburg. Downloaded on December 20,2020 at 05:40:18 UTC from IEEE Xplore. Restrictions apply. to drop the behind block.) [7] . Figure 2 shows the positional relationships between the wooden stick and building blocks; the position difference between the front and behind brocks is 80 mm . The height of the building blocks in the front is the same as that of the building blocks in the back . To ensure more stable operation, the robot arm is prevented from moving in the left/right (y-axis) and up/down (z-axis) directions, and it moves only forward/backward (x-axis). To investigate the effect of the robot position control using force information, we added 0 ms and 400 ms as the additional delay between the two robot systems and performed each task 10 times with and without the robot position control using force information " + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0002276_s12541-020-00431-8-Figure9-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0002276_s12541-020-00431-8-Figure9-1.png", + "caption": "Fig. 9 Virtual testing lab model", + "texts": [ + "\u00a07, the test rig is created by connecting actuators that can apply the vertical displacement to the wheel center of the CTBA model. The input profile used in this study is shown in Fig.\u00a08. A virtual testing lab (VTL) is an analysis method that uses a drive signal created based on the signal measured in vehicle durability tests. In this study, the wheel center force and moment signal measured by a wheel force transducer (WFT) during accelerated durability road (Belgian) tests are used as 1 3 the VTL drive signals. As shown in Fig.\u00a09, a test rig is created by connecting an actuator that can apply 6 DOF forces and moments to the wheel center of the CTBA model. The input profile used in this study is shown in Appendix\u00a01. In this section, the results of the two test modes on the beam element based CTBA model are compared with the results of the flexible body CTBA model in Fig.\u00a010. The main CTBA model parameter values used in validation are presented in Appendix\u00a02. In the roll simulation, the output values of each model are compared, including the roll torque of the suspension and axial force of the trailing arm" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0002521_ei250167.2020.9346917-Figure4-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0002521_ei250167.2020.9346917-Figure4-1.png", + "caption": "Fig. 4 Magnetic flux line distribution of the HS-PMSM", + "texts": [ + " Next, the machine\u2019s characteristics such as magnetic flux density distribution, magnetic flux line distribution, permanent magnet flux linkage, no load back electromotive force, electromagnetic torque, and loss are analyzed by finite element method and discussed. 2735 Authorized licensed use limited to: Miami University Libraries. Downloaded on June 15,2021 at 08:41:38 UTC from IEEE Xplore. Restrictions apply. The major parameters of HS-PMSM are listed in Table 1. While the distribution of magnetic flux density under load condition is shown in Figure 3, and the distribution of magnetic flux line under the same condition is shown in Figure 4. Figure 5 and 6 show the radial magnetic flux density in the machine gap. As shown in Figure 5, the red dot line and the yellow dash line represent for the flux density generated by the permanent magnets under no armature current condition and the flux density generated by the armature current under no permanent magnets exciting condition, respectively. While the blue solid line represent for the flux density generated by the armature current and permanent magnet together. It can been seen from Figure 5 that the flux density generated by the armature current and permanent magnet together is basically consistent with the superposition of the flux density generated by the permanent magnets under no armature current condition and the flux density generated by the armature current under no permanent magnets exciting condition" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0002759_b978-0-12-804560-2.00011-0-Figure4.3-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0002759_b978-0-12-804560-2.00011-0-Figure4.3-1.png", + "caption": "FIGURE 4.3 Cart-on-table model: the CoP location is determined by the horizontal acceleration component of the CoM.", + "texts": [ + " We have frx Mg = xg \u2212 xp z\u0304g (4.17) and hence, of the GRM, i.e. (xg \u2212 xp)Mg = z\u0304gfrx . Note that with this model, the vertical component of the GRF equals the gravity force; frz = Mg. Furthermore, using (4.8) with z\u0308g = 0, the equation of motion can be represented as x\u0308g = \u03c92xg + ma Mz\u0304g . (4.18) The LIP-on-cart model is used in numerous studies on balance control. Further details will be presented in Chapter 5. The LIP constraint is also used in the so-called \u201ccart-on-table\u201d model [55,54], shown in Fig. 4.3. The equation of motion of the cart-on-table model is identical to that of the IP models, (4.11). The difference lies only in the interpretation, when used in controller design. In fact, all these models demonstrate that balance control can be based on the relative CoP-CoM motion. In the table-cart model, the CoM motion trajectory is used as the input while the CoP position is subject to control. This is in contrast to the IP models, where the CoP is used as control input, as clarified above. CoP manipulation via CoM acceleration has also been exploited under the so-called \u201cangular momentum inducing IP model\u201d (AMPM) [62]" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0000125_amc44022.2020.9244424-Figure2-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0000125_amc44022.2020.9244424-Figure2-1.png", + "caption": "Fig. 2: 1D pendulum hardware and corresponding notations. This con guration corresponds to qb = \u2212\u03c0/4 and qc = \u03c0/2.", + "texts": [ + " This nonanthropomorphic robot has an auxiliary dynamic stabilization system which consists of two scissored pairs of control moment gyroscopes (CMG). The scissored pairs are orthogonal and thus the problem of vertical stabilization of the robot can be considered for each axis separately. Therefore, stabilization of the robot for one axis can be approximated with a simpli ed onedimensional prototype. In this paper we consider such a prototype, which is a control moment gyroscope inverted pendulum (Fig. 2). Note that the robot has a modular design: the biped is equipped with four identical CMG cubes. Control moment gyroscope is a widely used technological device that uses the reaction of a spinning wheel to external torques. Due to the advantages of a large ratio of produced torque to control torque and relatively low power consumption, CMGs have a wide range of applications, including vessel stabilisation [1], motorcycle and robot balancing [2], balancing aid for humans and This work was supported by the Russian Ministry of Education and Science, through the grant No", + " Downloaded on May 14,2021 at 02:01:16 UTC from IEEE Xplore. Restrictions apply. section III we present the state-feedback controller capable to stabilize the system if all states (including velocities) are measured. Then, in section IV we discuss three velocity observers. Hardware experiments and observers comparison are provided in section V. Finally, possible further research directions are discussed in the concluding section VI. II. Model Description The considered inverted pendulum is shown in Fig. 2. In this paper we follow the notations from the manual of the model 750 control moment gyroscope commercialized by Educational Control Products company [13]. The notations and the corresponding hardware parameters are summarized in the Table I. The pendulum consists of three bodies: the main body B, the gimbal C, and the wheel D. We associate a frame with each of the bodies: {O,~b1,~b2,~b3}, {O,~c1,~c2,~c3}, and {O, ~d1, ~d2, ~d3}, respectively. We assume that all the bodies are symmetric, and the centers of mass of all the bodies coincide at the point O. Due to the symmetry, the inertia matrices are diagonal. We denote the principal moments of inertia (w.r.t. the center of mass) of the bodies B, C and D as diag(Ib, Jb,Kb), diag(Ic, Jc,Kc) and diag(Id, Jd,Kd), respectively. The angle of the body B with respect to the vertical is denoted as qb, the angle of the body C with respect to the body B is denoted as qc, and the angle of the body D with respect to the body C is denoted as qd. The con guration shown in Fig. 2 corresponds to the angles qb = \u2212\u03c0/4 and qc = \u03c0/2. Note that the equilibrium position of an actual robot depends on the con guration of the legs and may be subject to external disturbances; therefore, for our cube we model the equilibrium point as qb = \u2212\u03c0/4\u2212 e, where e is the unknown (small) bias. The corresponding time derivatives are denoted as \u03c9b(t) := q\u0307b(t), \u03c9c(t) := q\u0307c(t), \u03c9d(t) := q\u0307d(t). If we apply torques \u03c4c and \u03c4d to the bodies C and D, respectively, we can nd the Euler Lagrange equations: 0 = ( J1 + J2 cos2 qc ) \u03c9\u0307b \u2212 J2 sin(2qc)\u03c9c\u03c9b + Jd\u03c9\u0307d cos qc \u2212 Jd sin(qc)\u03c9c\u03c9d \u2212mlg sin ( qb + \u03c0 4 + e ) , \u03c4c = (Ic + Id)\u03c9\u0307c + Jd\u03c9b\u03c9d sin qc + 1 2 J2 sin(2qc)\u03c9 2 b , \u03c4d = Jd\u03c9\u0307d + Jd cos(qc)\u03c9\u0307b \u2212 Jd sin(qc)\u03c9c\u03c9b, (1) where J1 = Id +Jb +Kc +ml2, J2 = Jc\u2212 Id +Jd\u2212Kc", + " More precisely, convergence to the vicinity instead of the nite-time convergence to the origin follows from the replacement of x2 and sin(x1) in f2(x, u) in (3) with x\u03022 and sin(y1) = sin(x1 +e), respectively. A more detailed analysis of the size of this vicinity can be performed by the means of Lyapunov function analysis in a similar way as in [16]; however, such a result is rather technical and is not presented here for brevity. V. Experiments and comparison The hardware for the tests1 (shown in Fig. 2) is assembled from o -the-shelf components. The hardware parameters are summarized in the Table I. A STM32F746 discovery board was chosen as the main computing unit. We have chosen a small brushless motor to drive the wheel 1We have lmed all the experiments, the video is available here: https://youtu.be/0xFanQ0QaEk. 49 Authorized licensed use limited to: Central Michigan University. Downloaded on May 14,2021 at 02:01:16 UTC from IEEE Xplore. Restrictions apply. D, the frame C is actuated by a Dynamixel MX106R servo motor", + " Three di erent velocity observers, namely the linear model-free di erentiator, the linear modelbased observer, and the nonlinear model-based di erentiator, have been implemented in the hardware and experimentally compared in the pendulum stabilization task. The experimental studies illustrate that the designs that utilize model knowledge outperform the model-free di erentiator. Moreover, the nonlinear model-based design provides smaller error than the linear one. At a further research direction, we intend to apply the designed nonlinear di erentiator for the biped system described in Section I, Fig. 2. At this step, we also intend to incorporate a friction model into the dynamics of the model-based velocity estimator to reduce the modeling errors in the system. 50 Authorized licensed use limited to: Central Michigan University. Downloaded on May 14,2021 at 02:01:16 UTC from IEEE Xplore. Restrictions apply. References [1] S. Inc., Anti-roll gyro, online: www.seakeeper.com/technology/. [2] S. D. Lee and S. Jung, Awakening strategies from a sleeping mode to a balancing mode for a sphere robot, International Journal of Control, Automation and Systems, vol" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0002830_978-94-007-6046-2_42-Figure5-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0002830_978-94-007-6046-2_42-Figure5-1.png", + "caption": "Fig. 5 Generalized 2D massless biped model [6]", + "texts": [ + " For example, we can use its analytical solution as the reference trajectory for walking with a big stride: xg.t/ D xg.0/ cosh.t=Tc/ C Tc Pxg.0/ sinh.t=Tc/ (18) Tc p yh=g Another useful equation is a conservation of the orbital energy E defined as 1 2 Px2 g g 2yh x2 g D constant E: (19) This can be obtained by multiplying (16) by Pxg and integrating it. Using the fact that E is kept constant for each step, we can design biped gait patterns by a simple calculation [8]. The potential energy conserving orbit was generalized in [6]. In this work, a generalized leg structure of Fig. 5 was introduced to avoid a specific mechanism. The leg posture is given by the ankle joint angle q1 and variables q2; q3 for an arbitrary mechanism to drive the body with respect to the first link. In the case of ordinary leg mechanism, q2 and q3 can be the knee angle and the hip angle, respectively. Taking .x; y; / as the body position/orientation, we can derive following general dynamics: m.y Rx x Ry/ C I R D u1 C mgx; (20) where m and I are body mass and moment of inertia and u1 represents the ankle torque" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0003837_aieepas.1960.4500711-Figure3-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0003837_aieepas.1960.4500711-Figure3-1.png", + "caption": "Fig. 3. Winding arrangements referring to loss calculations in Appendix 11", + "texts": [ + ")+OJ(t) 3 where sin 2t+sinh 2t cosh 2 -cos 2t and OW sinh t-sint 2t cosh +cosE and (12) (13A) (13B1) (14A) (14B)a=2ir(r,bf/ap'106)'/' 1/cm For windings with round conductors, the above ratio is given as: (15) 16A) k 1 +t'4(m'- 0.2)/15.25 I'a'd and '= 27r(df/d'p'10')1/2 1/cm (16B) Carter's Solution for D-C Machines3 The eddy-current loss in the active part of the conductor to the d-c loss is given as: 8 ' 1\"l sin PoV 2= 2( X 2 1 (sin -vnt~ - pt 2 2 1 1 ~~x pnt~ sin- i'6 (Yv+Z,, cos vy) (17) where (18A)Y=2(P-Q)-1 and Z=P-2Q In these equations, P kh(sinh 2kh+sin 2kh) cos 2kh-cos kh and (18B) (19A) (10) Q kh(cosh kh sin kh+sinh kh cos kh) cos 2kh-cos kh (19B) where k2= 472bfvn/pa (20) Solution Suggested by Stix36 (See Fig. 3) The factors ka denote the ratio of the additional losses in the winding to the d-c losses. 1. Windings with solid conductors; twolayer windings: Erdelyi-Calculation of Stray Load Losses in D-C MachineryJUNE, 1960 135 ka =L 38 T4w ,.f 21L+L90 7r \\T (2T) Four-layer windings: ka L 314W(- ' (22)LaL+L 180 rn-layer winding: kaL+L 15p -15p+4 , w F(a L+Le 45 T (23) 2. Two-layer windings, each conductor consisting of two laminations; winding as shown in Fig. 3(A): L 316 w a'\\ L+L, 180 T 20 T4w I'2a\" 64 r T Winding as shown in Fig. 3(B): L 316 4w Ia'\\ L+L 180 T 16 IT w F (25) 64 7 T/ Winding as shown in Fig. 3(C): k'L 316 wF a kaL+L 180 - + 4 w F(a ) (26) 64 T T/ 3. Two-layer windings, each conductor finally laminated; winding as shown in Fig. 3(D): ka= 1 rS r F( (27) 180 r Winding as shown in Fig. 3(E): 1a16Ow F- (28) 180 -r T) In these equations, a=(4rh2b)/(1071r2pa) (29A) 4a (29B) a\"=+ L (29C) L+Le and F(-) = (a/T)2 t1 (a/T)(1-T/6)} (30) ReFerences MUTATING MACHRINES IN GENERAL. ASA C50.41955, American Standards Association, Inc., New York, N. Y., 1955. 2. STRAY LOAD LossEs MEASURD IN D-C MOTORS, AIBE Committee Report. AIEE Transactions, vol. 68, pt. I, 1949, pp. 219-25. 3. STRAY LossEs IN ELECTRC MACHINES (in French), E. Roth. Bulletin, Soci&t6 Francaise des Electriciens, Malakoff, France, vol" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0002361_ffe.13405-Figure2-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0002361_ffe.13405-Figure2-1.png", + "caption": "FIGURE 2 Shape and dimensions (in mm) of the four-point bending test specimen: (A) rectangular specimen, (B) position from which rectangular plates were extracted from the round steel bar", + "texts": [ + " Type A defects were semicircular slits, simulating early stage flaking, with depths ranging from 0.05\u20130.15 mm (cf. Figure 1B). On the other hand, Type B defects were through-thickness incisions, replicating deeper cracks, with depths ranging from 1.0\u20131.8 mm (cf. Figure 1B). Both defect types were generated by electron discharge machining (EDM). Crack-growth testing was also conducted on JISSUJ2, the chemical composition of which was 1.02C0.27Si-0.39Mn-0.022P-0.008S-1.39Cr (mass %), balanced with Fe. Rectangular plate-specimens (cf. Figure 2A) were fabricated from a round bar of 75 mm in diameter, with care taken to avoid the central region (cf. Figure 2B). The plates were heat-treated at 1113 K for 1 h, then quenched in an oil bath and finally tempered at 473 K for 1.5 h. Subsequently, a parallel grinder was used to remove a surface oxidation layer, then a 3 mm deep notch was added via EDM as a crack-starter. To measure fatigue crack-closure behaviour during tension-compression fatigue tests, JIS-SUJ2 was also supplied, containing 1.00C-0.27Si-0.36Mn-0.014P-0.006S1.41Cr (mass %), balanced with Fe. Figure 3A exhibits the shape and dimensions of the specimen, fabricated from a round bar of 22 mm in diameter" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0003311_issnip.2015.7106943-Figure2-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0003311_issnip.2015.7106943-Figure2-1.png", + "caption": "Fig. 2. A simulation result showing how the flying cameras stabilize in the direction of the target\u2019s motion. (a) Cameras are initialized and head towards the goal locations; (b) cameras are about to reach the formation; (c) cameras reach their goal locations with a stable formation.", + "texts": [ + " Given a transitional control command, the camera\u2019s dynamics are controlled by the magnitude of the total thrust fi and by the control moment Mi = (M1,i,M2,i,M3,i) T for the desired attitude:\u23a1 \u23a2\u23a3 fi M1,i M2,i M3,i \u23a4 \u23a5\u23a6 = \u23a1 \u23a2\u23a3 1 1 1 1 0 \u2212d 0 d d 0 \u2212d 0 \u2212c\u03c4 c\u03c4 \u2212c\u03c4 c\u03c4 \u23a4 \u23a5\u23a6 \u23a1 \u23a2\u23a3 f1,i f2,i f3,i f4,i \u23a4 \u23a5\u23a6 , (12) where d is the distance from the center of the flying camera to the center of each rotor and c\u03c4 is a constant. Given a trajectory point xd,i(k), the total thrust fi and the desired direction of the third-body axis b3d,i(k) are selected to stabilize the translational dynamics. b2d,i(k) is then computed as a function of b1d,i(k) and b3d,i(k). Please refer to [15] for a detailed explanation of the control model. As an example, in Fig. 2 a team of flying cameras selfpositions around the moving target (black cuboid) by reaching their own goal locations (numbered black circles). The field of view of each camera is shown with cones. In Fig. 2a the flying cameras are initialized in random locations and their desired formation is an octagon. The flying cameras\u2019 dynamics make 1k is temporally omitted for simplicity of notation. the observation of a target difficult during the stabilization of the formation because of the limited field of view. When the cameras are about to reach the goal locations (Fig. 2b), the formation becomes more stable and the target is almost fully covered by the flying cameras\u2019 field of views. In Fig. 2b,c the cameras correctly reach their own goal locations by following and fully viewing the target. In this section we show how real-world flight dynamics of teams of flying cameras affect the viewing of a target during tracking. The starting locations of the cameras are randomly initialized for each simulation. We assess the stability and view performance using N = 6, 8, 10, 12. The parameters of the camera\u2019s dynamics are set as in [16], except those in Eq. 4 that we set as a = 4, b = 1, c = 0.6 in order to allow cameras to have a smoother repulsion when they interact" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0001117_kem.841.327-Figure12-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0001117_kem.841.327-Figure12-1.png", + "caption": "Fig. 12 Equivalent stress in Specimen.", + "texts": [ + " Also, the presence of the quasi- isotropic stacking order of the GFRP fibre enhances the load bearing capacity of the composite specimen. Graph 3. Stress vs Strain of tensile test from FEA Model. Fig. 9 Stress in GFRP. Fig. 10 Stress in Sugarcane. Flexural Stress Vs Strain. Fig. 11 shows the deflection obtained on the flexural loading on the FEA Model of the composite. A displacement of 7.5 mm was obtained in the experimental results, whereas in the FEA model it is only 3.5 mm. This indicates that the resistance to flexing of the material is better in FEA model. Equivalent Stress. Fig. 12 shows the stress distribution on the FEA model. It can be seen that the stress is minimum is the regions of support and maximum at the point of support. The stress is maximum at the outermost layers of top and bottom. Equivalent Stress in Sugarcane. Fig. 13 shows the stress distribution in the bottom Sugarcane fibre layer. A maximum stress of 19.21 MPa is obtained in the layer. It can be seen that the stress is maximum at the point of applying load. The stress distribution is gradual from minimum to maximum from the point of support to point of application of load" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0003239_icmtma.2015.150-Figure8-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0003239_icmtma.2015.150-Figure8-1.png", + "caption": "Figure 8. The trajectory of the robot driven by multi degree of freedom point", + "texts": [ + " Then establish a multipoint drive and driving deputy settings which are shown in figure 7. Set the X, Y, Z displacement function as: Tr X AKISPL(time,0,SPLINE_1, 0) Tr Y AKISPL(time,0,SPLINE_2, 0) Tr Z AKISPL(time,0,SPLINE_3, 0) Running the UP6 robot\u2019s simulation program in ADAMS, set the step to 1025 and set the time to 1205s. Then use the Create Trace Spline of the Review menu, after that choose Marker_42 and ground of work tools vertex orderly. Finally, the workpiece center operation curve[3-4] will be appeared on the screen which was shown in figure 8. IV. ANALYSIS OF THE KINEMATICS AND DYNAMICS SIMULATION Establish 3 measurement object for the marker Marker_42 on the end effector of a robot. Translation Displacement is the measuring method. The spatial position of markers are measured along the X, y and Z direction displacement component. Measurement data is obtained by ADAMS/Postprocessor post-processing module as shown in figure 10. Short dotted line represents the Z direction of the displacement. Due to the simulation trajectory in x-y plane, so the Z direction of the displacement not change" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0001921_s10010-020-00418-x-Figure2-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0001921_s10010-020-00418-x-Figure2-1.png", + "caption": "Fig. 2 Coordinate system for generating the conjugate gear pair, a the first enveloping motion, b the second enveloping motion", + "texts": [ + " Then, the meshing characteristics, such as the contact zones, contact ratio, sliding coefficients and induced normal curvature and the influence of those characteristics in varied geometrical design parameters is are analyzed. Furthermore, the contact stress is analyzed by using the finite element method (FEM) to validate the comparison result of the two cycloid drives according to the induced normal curvature analysis. Finally, three feasible design cases are developed based on the analysis mentioned above. Fig. 2 shows two movable coordinate systems S1, S2 and the fixed coordinate system Sf , which are rigidly connected to the conjugate envelope gear, the epitrochoid gear and the frame, respectively. In the initial position, the axes x1 and xf are coincident, x2 is parallel with xf . The axes y1, y2 and yf are coincident. The axes z1 and zf are coincident and z2 is parallel with zf . The distance between the axes of rotation of the epitrochoid gear and the conjugate gear is e. The epitrochoid curve is generated in S2 by point C that is rigidly connected to S1, at a distance a from its center o1. The conjugate envelope profile as the envelope to the family of the epitrochoid profile is also represented in S1. Point I is the instantaneous center of rotation. Based on the double-enveloping method [20\u201322], during the first enveloping motion as shown in Fig. 2a, the pin roller is used as the cutting tool, the internal ring gear rotates by angle 1 with the speed !1 about the axis z1 in counterclockwise direction, the external gear will rotate by angle 2 with the speed !2 about the axis z2 in the same direction according to the motion relation. When the pin roller radius is zero, the epitrochoid curve tooth profile can be obtained. Similarly, during the second enveloping motion as shown in Fig. 2b, the external epitrochoid gear rotates by angle 2 with speed !2 about the axis z2 in counterclockwise direction, the internal conjugate envelope gear will rotate by angle 1 with K speed !1 about the axis z1 in the same direction. Then, the rotation angles and speeds have the relationships: 2 1 = 2 1 = !2 !1 = n1 n2 = m21 (1) where m21 is the gear ratio, n1and n2 are the tooth number of the epitrochoid gear and the conjugate envelope gear, respectively. The representation of the epitrochoid profile is in two-parameter form" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0001896_icuas48674.2020.9213976-Figure1-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0001896_icuas48674.2020.9213976-Figure1-1.png", + "caption": "Fig. 1: Structure of the novel tilting tri-rotor system", + "texts": [ + " Clearly, this way of moving forward is inefficient and limited because there are fewer controllable variables than motion DOFs in the conventional model (underactuated system). In this study, each pair of coaxial rotors in the conventional tri-rotor have tilting degrees in two directions, which means an over-actuated system. Compared with the fully-actuated system, the actuators\u2019 status of the over-actuated system are not unique, which means a cost function can be set to optimize the flight task. The specific model structure is shown in Fig.1. Including the rotation speed of the six rotors that can be controlled, there are 12 controllable variables in the entire model. Therefore, in order to simplify the problem, the speed of the coaxial rotors is made equal, resulting in nine controllable variables after simplification. For the aircraft to move in the vertical direction, only three rotors are required. If horizontal flight is required, the rotors can be tilted in the desired direction, without needing to first tilt the fuselage. When more complex curved motion is required, the superiority of the proposed model over the underactuated model is dramatic. Before discussing the specific equations of motion, the coordinate system used in this study is defined and illustrated. The ground coordinate system OxEyEzE and the body coordinate system OxByBzB are both established as shown in Fig.1. We arbitrarily select a point on the surface of the Earth as the coordinate origin O, the direction of OxE points is parallel with the ground, the direction of OzE points is vertically downward, and the direction of OyE is determined by the right-hand coordinate law. Here, our model does not consider the influence of the Earth\u2019s rotation; thus, the ground coordinate system is an inertial coordinate system. Furthermore, in order to analyze the core problem and ignore the various subordinate factors, we employ the following assumptions: Assumption 1: The fuselage is a rigid body" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0000136_012076-Figure4-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0000136_012076-Figure4-1.png", + "caption": "Figure 4. Stress-strain state of the plate at different positions of the roll", + "texts": [ + " The concentration of such forces can provoke increased tenseness and deformations in these areas. The analysis of the plate stress-strain state is modeled in SolidWorks Simulation. During the simulation, the Finite Element Method was used and diagrams were obtained for the loads distribution and the stress-strain state of the plate. To study the stress-strain state, one of the working surfaces of the disk was divided into three segments, which describe three main areas of forces distribution. As it is seen from Figure 4 there are the stress diagrams in rolls for the material grinding. As a result of Reuleaux Triangle Profile rotation the working pressure area relative to the axis of roll rotation is shifted. Therefore, the forces operate on the roll surface nonuniformly. HIRM-2019 Journal of Physics: Conference Series 1353 (2019) 012076 IOP Publishing doi:10.1088/1742-6596/1353/1/012076 Within the analysis of the stress-strain state of the plates have been identified that the load is distributed evenly over the plate surface in order to reduce strain" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0003575_s00542-014-2136-5-Figure6-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0003575_s00542-014-2136-5-Figure6-1.png", + "caption": "Fig. 6 Finite element model and pressure distribution of FDBs, a FeM model, b pressure distribution", + "texts": [ + " This research estimates nrrO due to half-speed whirl (hSW) by applying the swept sine excitation in addition to the centrifugal force in force vector, as follows: where a, T , \u03c9s, and \u03c9e are sweep rate, sweep period, and starting and ending frequencies, respectively. 3.3 Simulation model We developed a finite element model of FDBs to numerically investigate the effect of hourglass-shaped sleeves. The finite element model of a 2.5\u2033 hDD spindle motor rotating at 7,200 rpm was developed with 3,380 isoparametric bilinear elements in four nodes. Figure 6a and b show the finite element model and its pressure distribution, respectively; Table 2 shows the major design variables. The FDB performance is investigated for an hourglass-shaped sleeve in which upper taper length (\u2206cu) of the upper grooved journal bearing and the lower taper length (\u2206cl) of the lower grooved journal bearing increases from 0 to 1 \u03bcm (8)F = p0 sin (a 2 t2 + \u03c9st ) , (9)a = \u03c9e \u2212 \u03c9s T , by 0.2 \u03bcm increments. The lower taper length of the upper grooved journal bearing and the upper taper length of the lower journal grooved bearing are assumed to be zero" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0000206_s00542-019-04659-x-Figure4-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0000206_s00542-019-04659-x-Figure4-1.png", + "caption": "Fig. 4 Controlled liquid metal catalyze the growth of silicon nanowires (a Schematic diagram of silicon nanowire growth catalyzed by controlled liquid metal droplets, b liquid metal wire is generated using a flexible direct write method, c silicon nanowires are generated using the SLS method)", + "texts": [ + " Si nanowires will growth from the eutectic due to the supersatyration of the eutectic (Sunkara et al. 2001). In this process, silicon is obtained from the solid substrate. A liquid phase eutectic is formed. Finally, the solid phase Si nanowires is gained. Therefore, this method is called as the SLS(solid\u2013liquid\u2013solid) mechanism (Yu et al. 2001). A low melting point liquid metal galinstan is used in this method to help form a eutectic with silicon at a lower temperature. It has the advantages of simple method and low cost. As shown in Fig. 4a. At first, liquid metal wires are direct written on a silicon substrate by a flexible nozzle. The silicon is then placed in a heating furnace at a high temperature of 600 C for 1 h. Silicon nanowires are available. The galinstan alloy can be removed by the hydrochloric acid from sample. Immerse the sample into 3 mol/L for 2 h, then take it out and rinse with deionized water for 2\u20133 times. As shown in Fig. 4b. It can be seen under the SEM that a silicon nanowire structure is formed at a position where liquid metal is present. The nanowires are clustered on a silicon substrate. The results show that low-temperature liquid metal can be used to gain silicon nanowires, which provides a new approach of the controlled growth of silicon nanowires. At the same time, the liquid metal processing based on the flexible nozzle direct writing processing method is small enough to be used in the growth of silicon nanowires" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0000492_tmag.2019.2954900-Figure3-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0000492_tmag.2019.2954900-Figure3-1.png", + "caption": "Fig. 3. The analysis model used in the validation of the proposed correction coefficient; (a) 2D FEA; (b) 3D FEA", + "texts": [ + "org/publications_standards/publications/rights/index.html for more information. CMP-539 4 III. VALIDATION OF PROPOSED CORRECTION COEFFICIENT THROUGH FEA ANALYSIS In order to verify the validity of the proposed correction coefficients, three types of Spoke type PMSM models are considered as shown in Table 1. The computer used in the performance analysis has two Intel Xeon CPU E5-2699 v4, 256GByte DDR4 RAM, and two Nvidia Titan X graphics cards. The 2D and 3D FEA analysis models used in the performance analysis are shown in Fig. 3. The 2D analysis model shown in Fig. 3 (a) has many meshes in the air gap for accurate analysis. As shown in Fig. 3 (b), only 1/2 of the stack is modeled using 3D symmetry. In addition, air dummy with sufficient thickness was added in the analysis model to accurately account for axial magnetic paths. While Fig. 4 shows the shapes of the analysis models, Fig. Fig. 5 shows the conventional 2D and 3D FEA analysis of the no-load back EMF of each analysis model in Fig. 4 and the 2D FEA analysis using the proposed correction coefficient. The analysis result waveform of 2D FEA using the proposed correction coefficient is consistent with the 3D FEA analysis result" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0000942_ilt-01-2020-0030-Figure3-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0000942_ilt-01-2020-0030-Figure3-1.png", + "caption": "Figure 3 Angle positions of elements", + "texts": [ + " Ac is the area associated with asperity\u2013 asperity contact, A0 is the total contact area, t fN is the Newtonian portion of the friction shear stress, andt flim is the maximum shear stress that can be sustained at the contact pressure. Note that for full-film lubrication, t f \u00bc t 1 fN 1 t 1 f lim -1 . Contact pressure s at any point (x, y) is determined from the following equation (20): s\u00bc 3Qc 2pab 1 x a 2 y b 2 \" #1=2 (20) where a and b are semi-major and semi-minor axis of the projected contact ellipse, respectively, as shown in Figure 2, andQc is the contact force. The angle positions of elements in the plane are shown in Figure 3. When the outer ring is fixed, under the combined action of contact force, centrifugal force and friction force, the equilibrium equations of mechanics of inner ring are given by the following equations: Fa \u00bc XZ j\u00bc1 Qci j\u00f0 \u00desinai j\u00f0 \u00de Ffi j\u00f0 \u00decosai j\u00f0 \u00de (21) Fr \u00bc XZ j\u00bc1 Qci j\u00f0 \u00decosai j\u00f0 \u00de Ffi j\u00f0 \u00desinai j\u00f0 \u00de cosc j (22) M \u00bc XZ j\u00bc1 Qci j\u00f0 \u00desinai j\u00f0 \u00de Ffi j\u00f0 \u00decosai j\u00f0 \u00de Rri 1Qci j\u00f0 \u00de D 2 cosc j (23) Fatigue life analysis Yue Liu Industrial Lubrication and Tribology The curvature center of the inner groove, outer groove and the ball center areOii,Oei andOj, respectively, after deformation, as O 0 ij,O 0 ej,O 0 j ", + " To confirm the above model and method, considering the angular contact ball bearing 7011C as an example, the result of spin to roll ratio by the method proposed in this study is comparedwith the raceway control theory. The outer ring is fixed, the geometric and material parameters are shown in Table I and Table II and operating conditions are shown inTable III. Fatigue life analysis Yue Liu Industrial Lubrication and Tribology When the speed is 5000, 20000, 30000, 40000 and 50000 r/min, the spin to roll ratio calculated by raceway control theory and the method provided by this paper are shown in Figure 5, and the position angle of the elements is shown in Figure 3. The red line represents the spin to roll ratio of inner race based on outer raceway control theory, while the blue line represents that of outer race based on inner raceway control theory. Comparing curves, for outer raceway control theory, the results calculated by two methods are basically consistent when the speed exceeds 20000 r/min; however, when the speed is 5000 r/ min, the results are quite different. For inner raceway control theory, the results calculated by twomethods are quite different except when the speed is 5000 r/min" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0002084_physrevfluids.2.033901-Figure4-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0002084_physrevfluids.2.033901-Figure4-1.png", + "caption": "FIG. 4. Frame (a) shows the body consisting of a sphere and a planar, elliptical appendage (with aspect ratio A = 0.7 and length L = 0.7 D) falling freely under the influence of gravity. Frame (b) shows the same body in a fixed framework, where the body is exposed to a constant free-stream velocity. Figure (b) also defines the drift (or lift) force Fdrift and the drag force on the body. The angle, \u03b1, between the total force and free stream direction in frame (b) corresponds to the drift angle in frame (a). The equilibrium angle \u03b80 in frame (b) corresponds to the turn angle in frame (a).", + "texts": [ + " The backflow region is modeled as an half-ellipse attached to the cylinder at angle \u03b8a = 55\u25e6 with the distance between the cylinder and the tip of the ellipse B(0) = 1.26 D. These coefficients, as described in Ref. [5], has been calibrated with respect to direct numerical simulations. The drag force Fdrag is not explicitly used, because it is canceled out, when 033901-5 taking ratio between the drift and drag forces. In the remaining part of this work, we explore the additional freedoms which a third dimension introduces by considering various shapes of the appendages. Figure 4(a) shows schematically the configuration of a sphere with an appendage falling freely under gravity in still fluid. We are interested in characterizing the drift and turn angles, as defined in the figure, for appendages of different shapes and sizes. We define turn as rotation around the y axis and drift as translation in the z direction. In order to take the first step in characterizing appendage-induced instabilities, we make two major simplifications. First, we limit ourselves to planar appendages that are shaped as ellipses with semimajor axis s1 and a semiminor axis s2 as shown in Fig", + " Therefore, the actual appendage shape is only the part of the ellipse that extrudes the sphere. Given these parameters, the length of ellipse semiaxis can be recovered as s1 = 1 2 ( D 2 + L ) , in the direction normal to the surface of the sphere and s2 = As1 in the direction tangential to the surface of the sphere. The thickness of the appendage [Fig. 5(b)] will be kept constant s = 0.02 D with respect to the diameter of the sphere. The second simplification is related to the fact that the general problem as illustrated in Fig. 4(a) is very challenging numerically, in particular at low density ratios [15] where wake-induced oscillations of the body exist [6]. In this work, we limit our investigations to sufficiently low Reynolds numbers such that the wake behind the sphere is steady but sufficiently large Reynolds numbers such that a significant recirculation region exists, i.e., Re = 200. In this way, we avoid nontrivial dynamic interactions between the motion of the body and the generated wake that may exist in time-dependent wakes. We can thus focus our attention on the instability generated by the appendage alone. Therefore the problem of a freely falling body with a constant velocity is replaced with a fixed body exposed to a constant free-stream velocity as shown in Fig. 4(b). It thus follows that all the degrees of freedoms of the rigid-body dynamics are constrained; the rotation around the y axis is modeled by considering the body at various turn angles. This approach is similar to the analysis carried out by Fabre et al. [16,17]. They also consider steady flow problems and relate the solution to the problem of freely falling body. In order to investigate flow structures responsible for oblique falling paths, they also carry out weakly nonlinear expansion in turn angle relative to the incoming flow velocity" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0003275_aim.2015.7222530-Figure3-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0003275_aim.2015.7222530-Figure3-1.png", + "caption": "Figure 3. Segment AJ follows a circular trajectory of radius r=40 mm.", + "texts": [ + " We aim to improve IBextl created by 12 arc-shaped permanent magnets, so that the highest magnetic torque can be imparted on the cubic IPM that is located at the center of the crank. To this end, we use the analytical model for the magnetic field created by arc-shaped permanent magnets [19]. We are interested in maximizing IBextl= JB\ufffd + Bf at the location of the IPM (note that Bz does not contribute to LZ) ' Thus, we need to understand the variation of the flux densities Bx and By at the center of the system produced by the radially and tangentially magnetized segments Ai, i= I ,2,3,4 when they follow a circular trajectory of radius r=40 mm (r = Rl +R Z) as shown in Fig. 3. 2 Fig. 4 shows the variation of Bx at the centre of the system produced by segment A3. Bx was maximum at y=90o and its maximum magnitude at this position was Bxrnax = 19.1 m T. In other words, the optimal location of the segment A3 along the circular trajectory is found when we place it at y=90o. We denote this as A\u00a7oo. If we placed A3 at y=60o (A\ufffdO\\ we would obtain Bx=16.S mT at the centre of the system. Similarly, A\ufffd200 produces Bx=16.S mT at the centre of the system (see Fig. 4). We also aim to find the maximum contribution to the magnetic flux density Bx at the centre of the system that the arc-shaped magnet A 1 can generate when it follows the same circular trajectory" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0003862_optim.2014.6850929-Figure3-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0003862_optim.2014.6850929-Figure3-1.png", + "caption": "Fig. 3. Distribution of flux density (in section) for 7.5kW-1000rpm three phase induction motor at t=0.5s", + "texts": [], + "surrounding_texts": [ + "In this stage is done the numerical modeling of parameters from induction motor with power 11kW -1000 rpm to establish the electromagnetic noise. In the first case is shows the magnetic flux density distribution at time t = 0.5 s. Distribution of field lines at t = 0.5 s is represented in the figure below. Normal component of magnetic induction in the air gap at moment t = 0.5s is shown in figure 5. From this graph it can be inferred the value of equivalent magnetic induction. The graphs that express the order of harmonics and their amplitude is shown in figure 6. Can observe that the higher amplitudes have them the odd harmonics, respectively number 2 and number 34 considering that a fixed supply frequency. If the induction motor is fed by an inverter, induction harmonics due to switching they had much higher amplitudes Acoustic noise level generated by harmonics (magnetic noise) is shown in the figure below. The average value of the continuous sound pressure is 39.8 dB. In order to determine the total acoustic noise there are, also, taken into account, the values of mechanical noise (produced by the bearings) and the aerodynamic noise (produce by fan). The electromagnetic torque of this motor is plotted comparatively in the next figure: Because of the high power of the motor (7.5kW) torque at startup varies between the limits -10 \u2013 10 Nm, and its stabilization takes place after, about, 1.5-1.6 seconds. The waveform of the supply voltage and of the current at the motor\u2019 terminals are presented in figure 9. Due to the transient regime which is registered at the startup of the motor, the values of the absorbed current are increasing since the motor power is very high. After the finishing of the transient regime, the current absorbed by the motor stabilizes at the value of 12 A in normal regime of functioning and without load shaft." + ] + }, + { + "image_filename": "designv11_71_0002923_9781119633365-Figure9.15-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0002923_9781119633365-Figure9.15-1.png", + "caption": "Figure 9.15 Zero-current turn-off lossless snubber with energy recovery.", + "texts": [ + " Nevertheless, the presence of the \u201ctail current,\u201d during their turn\u2010off process, induces significant switching losses in the semiconductor switches. Thus, a ZCS feature is mainly required when using IGBTs as switches. The switch turn\u2010off loss, which is usually the dominant switching loss in high\u2010 power applications, cannot be alleviated effectively with the ZVS technique. From Figure\u00a0 9.1, it can be seen that the ZCS technique can significantly reduce the switch turn\u2010off loss by forcing the outgoing switch current to zero prior to its turn\u2010 off. Figure\u00a09.15 shows the equivalent PWM converter with zero\u2010current turn\u2010off soft\u2010switching cell that can represent the basic active soft\u2010switching PWM con verter with a ZCS function. It differs from a conventional PWM converter by including additionally an auxiliary switch S1, a resonant inductor L1, a resonant Figure 9.14 Active soft-switching converter with ZVS (type 3). capacitor C1, and a freewheeling diode D1. The resonant branch is only active dur ing a relatively short switching time to create a ZCS condition for the main PWM switch without substantially increasing voltage or current stresses", + " The auxil iary switch S1 is turned on before the main switch SM is turned off so that the reso nant tank current can rise sinusoidally and allow zero\u2010current switching for the main switch SM. Thus, to implement this ZCS mechanism, an L\u2013C network is needed to add around the switch so that the switch current may be kept at zero during switching transition. When the soft\u2010switching transition is over, the con verter simply comes back to the basic PWM operating mode. According to the basic concept of Figure\u00a09.15, the ZCS\u2010PWM soft\u2010switching converters can be pos sibly implemented in a number of ways. From the circuit topological point of view, one should keep it in mind that every ZCS converter can be viewed as a vari ation of the equivalent circuit shown in Figure\u00a09.15. In this section, two ZCS schemes are proposed and investigated to further improve the ZCS technique. With modified control and topology, the main switches are switched off under ZCS conditions; thus, the switching losses and stresses are reduced significantly. Figures\u00a09.16 and 9.17 show two typical ZCS\u2010 PWM topologies. These converters have the following features: In the following, we use Figure\u00a09.16 as an example to illustrate the operational principle for a ZCS\u2010PWM converter. Referring now to Figure\u00a09" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0003275_aim.2015.7222530-Figure2-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0003275_aim.2015.7222530-Figure2-1.png", + "caption": "Figure 2. a) Arc-shaped permanent magnet. b) top view of different types of arc-shaped permanent magnets (i.e . \u2022 A,. A2. A1\u2022 and A4) used in this work.", + "texts": [ + " 1 shows a ring-shaped external magnet that produces a rotating magnetic field around the patient bed when it is physically rotated. A 3.17 S mm cubic IPM with magnetization grade of 1.4 T (i.e., NSO) is placed in the prototype of a capsule. The IPM is rotated by the external rotational magnetic field. This rotational movement is then converted into a translational movement by a slider-crank mechanism. In this way, a piston will push drug out of a reservoir when the external magnet is rotated around the patient. Arc-shaped magnets, as shown in Fig. 2 (a), have been used in applications where torque coupling between separated members is required [16]. Furthermore, multiple arc-shaped magnets have shown a high efficiency in power transmission where high torque densities are needed [17]. For these reasons, we propose in this work for the first time the use of multiple arc-shaped magnets as the source of an external rotating magnetic field. This external magnetic system consists of 12 arc-shaped magnets. The analysis presented in this paper can be used for any number of segments, but for practical reasons, we will use three segments of each type of arc-shaped permanent magnets. Specifications of each segment: R1=30 mm, R2=SO mm, thickness L of 30 mm and all have the same magnetization grade of 1.32 T (i.e., N4S) but are radially and tangentially magnetized as shown in Fig. 2 (b). With 12 segments, there are different possible configurations to place in a ring-shaped structure. We are interested in finding the optimal configuration to transmit the highest possible torque on a small cubic IPM. In the next section, we present the methodology used to achieve this goal by means of analytical functions that model the magnetic field created by arc-shaped permanent magnets. The magnetic torque exerted by an external magnetic field I Bextl on a cubic IPM with volume V and magnetization Iml is given by [18] LZ = " + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0000253_2019045-Figure6-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0000253_2019045-Figure6-1.png", + "caption": "Fig. 6. Left tooth profile after thermal deformation.", + "texts": [ + " Where, rbi (i=1, 2) is the radius of base circle, mki (i=1, 2) is the roll angle of point k in the involute, rki (i=1,2) is the radiusofpoint k in the involute, uki (i=1,2) is theexpansionangleofpoint k in the involute,gki (i=1,2) is the helix angle of ascent of point k in the involute, bki (i=1, 2) is the helix angle of point k in the involute. 3.2.1 Variation of tooth profile equation caused by thermal deformation Still take the left tooth surface as an example, the left tooth profile after thermal deformation is shown in Figure 6. The tooth profile equation after thermal deformation can be expressed as x 0 k \u00bc r 0 b sinm 0 k r 0 bm 0 k cosm 0 k y 0 k \u00bc r 0 b cosm 0 k \u00fe r 0 bm 0 k sinm 0 k ( \u00f06\u00de where r 0 b \u00bc rb\u00f01\u00fe Dtl\u00de, m0 k \u00bc tana 0 k, a 0 k \u00bc arcos r 0 b r 0 k , r 0 b is the radius of base circle after the thermal distortion, m 0 k is the roll angle of point k in the involute after the thermal distortion. The helix diagram of left tooth surface and right surface after thermal deformation is shown in Figures 7 and 8, respectively" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0001588_aim43001.2020.9158991-Figure2-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0001588_aim43001.2020.9158991-Figure2-1.png", + "caption": "Figure 2. TakoBot segment structure", + "texts": [ + " The segments are connected by universal joints and supported by four compressional springs to maintain a linear and bending stiffness between an adjacent segment. Universal joints are interconnected by linear guide shaft with 30mm of length, which passes through the linear 978-1-7281-6794-7/20/$31.00 \u00a92020 IEEE 460 Authorized licensed use limited to: University of Wollongong. Downloaded on August 11,2020 at 10:23:43 UTC from IEEE Xplore. Restrictions apply. bearing, fixed in the center of each spacer discs. The wires are aligned along with the wire eyelets on each disc (Fig. 2). Such a structure so-called sliding disc mechanism provides smart bending stress distribution along the slender part and improves manipulator bending features by decentralizing forces and bending torques [24]. The maximum traveling displacement of sliding discs is 10mm. The spacer disc diameter is 50mm, and the total length of the continuum part is 380 mm. The wire passes along the spacer discs via wire eyelets and passes through inside of the compressional springs (Fig. 3). In this prototype, we used a stainless-steel wire with a diameter of 1mm, and the eyelet hole size is about 1" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0003158_978-3-658-05978-1_8-Figure1-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0003158_978-3-658-05978-1_8-Figure1-1.png", + "caption": "Figure 1. System Outline of this control method", + "texts": [ + " For constant outside force like lateral slope, slight calibration between steering angle and tire angles runs to reduce the driver\u2019s effort to hold the steering wheel to keep the car along the lane. 2. To stabilize the car against fluctuating disturbance, the steering force and the tire angles are controlled precisely according to the yaw angle to the lane. In this approach, the steering force is only modulated in the latter case, therefore, the lateral stability is improved with a small amount of control, which does not spoil the natural steering feeling. Fig.1 shows a system outline of this control. This system consists of a steering system in which the steering force and the wheel angles can be controlled independently, a lane recognition camera unit, a yaw rate sensor, and a control unit. In this newly developed steering system, the mechanical linkage between the steering wheel and the tires is replaced with electric signals via three ECUs, and the tire angles and the steering force can be controlled independently. The outline of this system is shown in Fig" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0001060_j.ijimpeng.2020.103609-Figure5-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0001060_j.ijimpeng.2020.103609-Figure5-1.png", + "caption": "Fig. 5. (a) Gas gun used in experiment. (b) Penetrated hole in the sandbag.", + "texts": [ + " (1) because the equation does not include the effect of vacuum and pneumatically compaction on the calculated volume of sandbag. The projectile used for ballistic impact test was a 12 mm diameter chrome bearing steel (AISI E 52,100 steel) sphere weighing 7 g. This type of steel has an elastic modulus of 210 GPa, poisson's ratio of 0.30, hardness of 62\u201366 HRC, thermal expansion coefficient of 12.5 \u03bcm/m \u00b0C, specific heat capacity of 0.475 J/g \u00b0C and thermal conductivity of 46.6 W/mK. The projectile was propelled by high pressure gas gun (Fig. 5a). The gas gun utilized helium gas pumped from gas tanks to launch the projectile at desired impact velocities. The gas supply could be manually varied using a pressure regulator and a digital pressure gauge accurate to 0.1 bar. The spherical steel projectiles were selected because we had no access to a rifle barrel. Without rifling, no spin is produced to stabilize sharp nose projectiles around the longitudinal axis during shooting. The velocity of the projectile was measured by two laser photodiodes. The photodiodes were placed near the exit of the gun barrel, 0.1 m apart. When the projectile passed through the laser beam of the photodiodes, a peak voltage was registered by a digital oscilloscope connected to the two photodiodes. The experimental data was considered valid only when the perforation was at the middle of the sandbag where an \u2018X\u2019 was marked as shown in Fig. 5b. The ballistic limit is defined as the impact velocity of a given projectile at which 50% of the firings results in a complete penetration of the given target [18,19]. Through repeated experiments, the ballistic limit is taken to the average of the highest projectile velocity which did not result in complete penetration of the sandbag and lowest projectile velocity at which complete penetration occurred. The difference between these 2 velocities was less than 5 m/s. The kinetic energy of a projectile at the ballistic limit is taken to be the limit of energy that the sandbag specimen can gain, i" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0002816_thc-2010-0566-Figure5-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0002816_thc-2010-0566-Figure5-1.png", + "caption": "Fig. 5. Elastic and plastic deformation.", + "texts": [ + " Starting from zero, as the applied stress is increased the strain increases linearly up to a limit called the elastic limit. In this region I the material shows elastic behavior and follows Hooke\u2019s Law Eq. (3), where the strain is linearly proportional to the applied stress and the constant of proportionality E is the Young\u2019s modulus having the same units as stress i.e. Pascal. \u03c3 = E\u03b5 (3) If the applied stress is removed within the elastic limit then the material regains its original size meaning that the strain reduces to zero and there is no permanent deformation in the specimen (Fig. 5A). At the end of the elastic region, the plastic region starts where a slight increase in stress above the elastic limit i.e. yield stress results in deformation almost without an increase in stress, this process is called yielding (region II). In region III the applied stress again increases, resulting in an increase in strain. The point where the maximum amount of stress is supported by the specimen is called the tensile strength. In region IV the supported stress continuously decreases leading finally to failure. If the applied stress is removed in the plastic region, the strain does not reduce to zero but some residual strain remains in the form of permanent deformation (Fig. 5B). The start of the plastic region is difficult to determine, since the linear increase of the stress is gradually changing into the yielding phase. For practical reasons the start of the plastic region is defined by the point where if the applied stress is removed the specimen keeps a small amount of permanent deformation (\u03b5 = 0.002 = 0.2%); this stress is therefore called the yield stress \u03c30.2. Figure 6 shows two different type of materials; material A (Fig. 6A) is strong because it can support very high stresses but this material is also brittle, meaning that it cannot be given any permanent (plastic) deformation, just after reaching the maximum stress the material fails giving rise to a flat fracture surface (Fig" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0003338_jae-141850-Figure1-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0003338_jae-141850-Figure1-1.png", + "caption": "Fig. 1. Cross section of single phase LSPMM.", + "texts": [], + "surrounding_texts": [ + "Keywords: Permanent magnets, FEM, preisach modeling, LSPMM, stator coil magnetization\nThe line start permanent magnet motor (LSPMM) combines a permanent magnet rotor (which provides better motor efficiency during synchronous running) with an induction motor squirrel cage rotor. This combination enables the motor to start by direct coupling to an AC power source.\nThis solution may lead to important technical advantages, such as a reduction in manufacturing costs, and/or better performance. Also, if the space of the rotor is properly negotiated between the permanent magnets and the squirrel cage, it may result in reduced running costs.\nAs a magnetizing field is generated by the permanent magnets, no magnetizing current is drawn from the line. This can lead to a much higher power factor in full load operation, as well as a decrease in losses in the stator winding.\nIn many applications, common induction motors can be replaced by LSPMM. Due to their known torque-speed characteristic, one of the most appropriate examples would be centrifugal applications, such as fans or pumps.\nIn order to simplify the manufacturing process and reduce manufacturing costs, post-assembly magnetization, as in method [1] is necessary. However, the eddy current which occurrs in the rotor bar as a result of this magnetization also disturbs the magnetization of Nd-Fe-B magnets.\nThis paper deals with the analysis of characteristics evaluations for the PM magnetization process, using a stator coil, in a post-assembly LSPMM, using coupled finite element method (FEM) and Preisach\n\u2217Corresponding author: Young Hyun Kim, Department of Electrical Engineering, Hanbat National University, Daejeon 305- 719, Korea. E-mail: kimyh@hanbat.ac.kr.\n1383-5416/14/$27.50 c\u00a9 2014 \u2013 IOS Press and the authors. All rights reserved", + "modeling. The evaluation is presented in order to analyze the magnetic characteristics of permanent magnets.\nThe focus of this paper is the characteristics analysis, relative to magnetizing direction and the quantity of permanent magnets, due to the eddy currents which occurr in the rotor bar during post-assembly magnetization of Nd-Fe-B magnets. Coupled Finite Elements Analysis (FEA) and Preisach modeling have been used to evaluate this nonlinear solution [2\u20134].\nMaxwell\u2019s equations can be written as:\n\u2207\u00d7 H = J0 + Je (1)\n\u2207 \u00b7 B = 0 (2)\nB = 1\n\u03c50 H + M, B =\n1\n\u03c50 H + MPM (3)\nWhere, M and MPM are the magnetization of magnetic material and permanent magnets, respectively, with regard to the magnetic intensity, H . The magnetic vector potential, A, and the equivalent magnetizing currents, Jm and JPMm, are expressed as follows:\nB = \u2207\u00d7 A (4)\nJm = \u03c50(\u2207\u00d7 M), JPMm = \u03c50(\u2207\u00d7 MPM) (5)\nThe governing equation derived from Eqs (1)\u2013(5), is given by:\nJe = \u03c3 E = \u03c3\n( \u2212\u2202 A\n\u2202t + v \u00d7 B +\u2207\u03d5\n)\n\u03c50\n( \u2207\u00d7\u2207\u00d7 A ) = J0 + Je + Jm + JPMm\n(6)", + "When the moving coordinate system is used, the governing 2D is given as follows: \u2202\n\u2202x \u03c50\n( \u2202AZ\n\u2202x\n) + \u2202\n\u2202y \u03c50\n( \u2202AZ\n\u2202y\n) = \u2212JZ + \u03c3 \u2202Az\n\u2202t + \u03c3\n\u2202\u03d5 \u2202z \u2212 Jm \u2212 JPMm (7)\nJm = \u03c50\n( \u2202My\n\u2202x \u2212 \u2202Mx \u2202y\n) (8)\nJPMm = \u03c50\n( \u2202MPMy\n\u2202x \u2212 \u2202MPMx \u2202y ) Where, Az: z component of magnetic vector potential, Jz: current density, \u03c50: magnetic resistivity, Mx, My, MPMx, MPMy: the magnetization of magnetic material and PM with respect to the magnetic intensity, Hx, Hy, \u03c3: conductivity of the rotor bar, \u03c6: scalar potential.\nIn this model, as the periodic boundary condition is used, \u03c3 \u2202\u03c6 \u2202z should equal zero.\nThe circuit equation is written as follows:\n{V } = [R]{I}+ [L0] d\ndt {I}+ {E} (9)\nWhere, {E}: E.M.F. vector in the winding. {V }: supplying voltage vector. {I}: phase current vector. [L0]: leakage inductance.\nTo solve Eq. (7), the Galerkin finite element method is used. For the time differentiation in Eq. (9), a time stepping method is used with a backward difference formula. Coupling Eqs (7)\u2013(9), the system matrix is given as follows:[[\n\u03c50[S]\u2212[N ] [0] [R]\n] + 1\n\u0394t\n[ [0] [0]\n[LG]T [L0]\n]]{{A} {I} } t = 1 \u0394t [ [0] [0] [LG]T [L0] ]{{A} {I} } t\u2212\u0394t + {{M} {V } } t\n(10)\nThe magnetization, M , can be expressed as a scalar model. It can be supposed that the domain in a LSPMM is an alternating field with reference to an x and y axis [5,6]. It is natural that M and H , which are calculated on the same axis, have the same vector direction.\nM (t) = \u222b\u222b \u03b1 \u03b2 \u03bc (\u03b1, \u03b2) \u03b3\u03b1\u03b2 (H (t)) d\u03b1d\u03b2 = \u222b\u222b S+(t) \u03bc (\u03b1, \u03b2) d\u03b1d\u03b2 \u2212 \u222b\u222b S\u2212(t) \u03bc (\u03b1, \u03b2) d\u03b1d\u03b2\n(11)\nThrough certain function transforms, the area integration can be related to the hysteresis loops. A more convenient treatment of this model is to substitute the Everett plane for Preisach\u2019s model.\nE (\u03b1, \u03b2) = \u222b\u222b \u03b1 \u03b2 \u03bc (\u03b1, \u03b2) \u03b3\u03b1\u03b2 (H (t)) d\u03b1d\u03b2 (12) In the Everett plane, the distributions of M , which are accepted from experimental data (stator, rotor: S18, PM: Nd-Fe-B), are Gaussian. [7]." + ] + }, + { + "image_filename": "designv11_71_0002516_iros45743.2020.9341263-Figure2-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0002516_iros45743.2020.9341263-Figure2-1.png", + "caption": "Fig. 2. a) Diagram of the modified SLIP model used to study the effect of multiple springs. Each spring can utilize a different nominal length and stiffness. One spring is defined to be a virtual spring V , while the other is defined to be a physical spring P . The length and angle of the leg are \u03b6 and \u03c8, respectively. b) Diagram of the 5-bar leg simulated by the modified SLIP Model. The leg can include springs at 3 possible locations, labeled as k1, k2, and k3", + "texts": [ + " The simulated response of the hopper to changes in leg stiffness, ratio of mechanical to virtual springs, and ride height are outlined in Sec. III. The leg design and experimental setup for the Minitaur leg are described in Sec. IV, and Sec. V outlines the experimental results of running with different physical springs and payloads. Section VI describes the application of PEAs on the quadruped LLAMA. The paper concludes (Sec. VII) with directions for future work. To simulate the Minitaur leg shown in Fig 1a, a modified SLIP model is used. This model has been adapted to include two springs in parallel, as seen in Fig. 2a. One spring is assumed to be a physical element P , and the other a virtual element V . The nominal length and stiffness of each spring is dependent on its type and location on the five-bar leg. Potential types and locations are denoted by k1, k2, and k3 as a torsional spring at the hip, a pair of torsional springs at the knees, and a linear spring along the leg, respectively, as shown in Fig. 2b. The physical spring element is assumed to restore all energy, with no damping elements. The virtual spring is assumed to store no energy, such that any force generated is the result of actuation. The links, with lengths \u03bb1 and \u03bb2, are assumed to remain constant throughout this study and possess no mass. The robot body is defined as a point mass, while the foot is assumed to have sufficient friction to avoid sliding when contacting the ground. All motion is constrained within the sagittal plane", + " When moving away from the nominal leg length, having a small amount of physical stiffness kratio = 20-40% significantly improves the versatility of the leg, while still receiving benefits to energy efficiency. To validate these simulations, an experimental platform was developed with a single, direct drive 5-bar linkage leg from Minitaur [33] powered by two T-motor U8 brushless motors. To restrict the leg motion to the sagittal plane, the leg is attached to a 1.34 m boom arm. The touchdown point (toe) is made of 3D printed ABS plastic with an overmolded elastomer to increase friction between the ground and the leg as seen in Fig. 2c. Parameters for this design can be seen in Table I. To calculate motor power, Eq. 11, the desired motor torque \u03c4 from the leg and motor measured angular velocity \u03c9 from the motor\u2019s Hall Effect based absolute encoders are used. To measure horizontal speed, v, an Encoder Outlet model 15s rotary encoder operating in quadrature phase was used. Data is recorded at 1000 Hz with a Teensy 3.6 microcontroller. 3644 Authorized licensed use limited to: Central Michigan University. Downloaded on May 14,2021 at 08:01:01 UTC from IEEE Xplore", + "5 kg payload, the soft spring could no longer achieve stable running. The stiff spring not only supported the added payload for a small increase in COT, but also adapted to a gait similar to the soft spring without the payload, matching both touchdown angle and forward velocity. Simulation environments also allow us to study more complex robot designs outside of the SLIP model. In addition, we can evaluate other situations where versatility would be useful, such as walking up inclines. By using a different robot and linkage morphology from Figure 2b, we also show that different leg designs and platforms can benefit from PEAs. To extend our study of the tradeoffs of PEAs on the versatility and energetic efficiency of running we consider the quadruped robot LLAMA (Fig 1b) for which we have recently developed a multibody simulation using Simscape Multibody [28]. Like Minitaur, LLAMA currently relies on purely virtual springs in its legs. We modified the simulation to include a linear spring element attached from the hip to the toe. The nominal length of the spring is defined as the touchdown length of the trajectory, LSpring " + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0002869_ijhm.2018.094879-Figure6-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0002869_ijhm.2018.094879-Figure6-1.png", + "caption": "Figure 6 Comparison of maximum temperatures \u0394tmax between the simulation and experiment (tc = 50\u00baC), (a) simulation (b) experiment (see online version for colours)", + "texts": [ + " It should be noted here that the phenomenon can be observed in the experiment and modelled by the simulation where a location at \u0398 = 270\u00ba is cooled, which corresponded to the leading-edge side of the slipper pad. Figure 5 presents the temperature distributions \u0394t under a higher oil temperature (tc = 50\u00baC) for a wide range of speeds N (= 3.3\u201326.7 s\u22121) at a constant supply pressure ps = 35 MPa. The temperature \u0394t in the case of tc = 50\u00baC (Figure 5) was lower than \u0394t of tc = 30\u00baC (Figure 4) for both the simulated results and experimental data. In addition, regarding the coincidence between simulation and experiment, the differences of Figure 5 are smaller than those of Figure 4. Figure 6 shows contour plots of the maximum temperatures \u0394tmax for a the simulation b experiment (Kazama et al., 2015) for a wide range of supply pressures ps up to 35 MPa and rotational speeds N up to 26.7 s\u20131 at the oil temperature tc = 50\u00baC. As the pressure ps and speed N increased, the temperature \u0394tmax increased for both the simulation and experiment. This is because the total power becomes large. Although aspects of these plots are similar, the effect of ps on \u0394tmax of the experiment was more homogeneous than that of the simulation" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0000315_icems.2019.8921590-Figure3-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0000315_icems.2019.8921590-Figure3-1.png", + "caption": "FIGURE 3. (a) Polarizations at the input and the two outputs of the OMT. (b) Designation of the input and outputs ports of the OMT (top view).", + "texts": [ + " The concept of our OMT design is based on the following rules: 1) the matching element should not be split into four identical blocks intersecting on the axis of the input circular waveguide (as described in [16]); 2) the matching element should not be a mobile part to be glued, screwed, soldered or just VOLUME 1, 2013 481 inserted in the center of the turnstile junction (as described in [16] and [17]); and 3) the total number of blocks to assemble the OMT should be minimized. Respecting these rules is possible with the final structure of the OMT presented in Fig. 3(a) and 3(b). For each polarization, an E-plane 180\u02da out-of-phase power combiner is used in order to recombine the RF signals split by the turnstile junction through opposite waveguide outputs. From a practical point of view, the final structure of the OMT presented in Fig. 3(a) and 3(b) is realized by superimposing three aluminum blocks (see Fig. 4). The upper block consists of a circular waveguide input, a circular waveguide transition, two E-plane power combiners and four H-plane 90\u02da bends (as shown in Fig. 5(a)). The lower block consists of the base of the turnstile junction with two superimposed cylinders as a matching stub, four H-plane 90\u02da bends and four E-plane 90\u02da bends (as depicted in Fig. 5(b)). The middle aluminum block acts as an interface between the lower and upper blocks, connecting the outputs of the turnstile junction to the inputs of the power combiners" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0000812_cdc40024.2019.9029757-Figure3-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0000812_cdc40024.2019.9029757-Figure3-1.png", + "caption": "FIGURE 3. Work principle of plate-inclined plunger pump/motor.", + "texts": [ + " Results showed that the simulation model was corrected and the control strategy can improve the EHHT\u2019s control performance. II. PRINCIPLE ANALYSIS OF EHHT Electro-hydraulic servo plate-inclined plunger hydraulic transformer is a swash plate plunger-type hydraulic component, with several advantages such as compact structure, small radial size, low mass, low moment of inertia and so on. It achieves the function of oil suction and discharge through the volume change generated by the reciprocating movement of the piston against the cylinder bore. The schematic diagram is shown in Fig.3. As there is an angle \u03b9 in the swash plate, when the cylinder is rotated, the plunger will reciprocate in the cylinder cavity. The displacement equation is s = R tan \u03b9(1\u2212 cos\u03d5) (3) Where s is the displacement of plunger at the top dead center (m), R is the radius of plunger distribution circle (m), \u03b9 is the tilt angle of swash plate (\u25e6), \u03d5 is the angle of plunger versus the top dead center (\u25e6). Swash plate piston pump / motor is featured by compact structure, small radial size, low mass and low moment of inertia" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0002276_s12541-020-00431-8-Figure6-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0002276_s12541-020-00431-8-Figure6-1.png", + "caption": "Fig. 6 Components and hardpoints of CTBA model", + "texts": [ + " These errors are considered to occur because the Timoshenko beam elements cannot express the warping effect. 1 3 The torsion beam model presented in Sect.\u00a02.1 is combined with the remaining parts to create the CTBA model. First, parts such as the trailing arm, knuckle, and spring lower mount are configured as a rigid body model, such that the CTBA\u2019s hardpoints can be directly changed. Then, force elements such as the spring, damper, bush, and bump stopper are modeled based on the parameters in Table\u00a04. Figure\u00a06 shows the main configuration components and hardpoints of the CTBA model. Modeling is performed using Dymola (Dassault Syst\u00e8mes, France), which is a development environment based on the Modelica language. To verify the accuracy and perform application cases for the beam element based CTBA model proposed in this study, two test modes were performed: roll test and VTL (virtual testing lab). This section describes the test rig models for performing these tests. In roll simulation, vertical displacement is applied in the opposite direction to the left and right wheels to create roll torque" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0002989_978-3-642-16214-5_9-Figure9.12-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0002989_978-3-642-16214-5_9-Figure9.12-1.png", + "caption": "Fig. 9.12. Borg-Warner system single-cone synchronizer (ZF). 1 Idler gear running on needle roller bearings; 2 synchronizer hub with selector teeth and friction cone;", + "texts": [ + " The shift movement s at the gearshift sleeve is approximately 7.5\u221213 mm. The permissible wear \u2206Sperm is generally between 1 and 1.5 mm. The wear reserve of the synchronizer unit is calculated by subtracting the operating clearance from the permissible wear. The maximum wear \u2206Vmax is of the order of 0.15 mm per friction pairing for cone synchronizers. Single-cone synchronizers based on the \u201cBorg-Warner\u201d system are widely used in manual transmissions. The various phases of the synchronizing process are shown in Figure 9.13, based on the ZF-B synchronizer (Figure 9.12, \u201cB\u201d standing for Borg-Warner system). The synchronizer body 4 is fixed to the transmission shaft. The synchronizer ring 3 is guided by stop bosses in the synchronizer body. These are narrower than the grooves in the synchronizer body, which allows the synchronizer ring a certain amount of freedom to twist radially. 3 main functional element, synchronizer ring with counter-cone and locking toothing; 4 synchronizer body with internal toothing for positive locking with the transmission shaft and external toothing for the gearshift sleeve; 5 compression spring; 6 ball pin; 7 thrust piece; 8 gearshift sleeve with internal dog gearing 2 Synchronizer hub with selector teeth and friction cone; 3 main functional element, synchronizer ring with counter-cone and locking toothing; 4 synchronizer body with internal toothing for positive locking with the transmission shaft and external toothing for the gearshift sleeve; 5 compression spring; 6 ball pin; 7 thrust pieces; 8 gearshift sleeve with internal dog toothing Before the shifting process starts, the gearshift sleeve is held in the middle position by a detent" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0003895_icems.2014.7013984-Figure5-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0003895_icems.2014.7013984-Figure5-1.png", + "caption": "Fig. 5.", + "texts": [], + "surrounding_texts": [ + "The PSRM is a high nonlinear system due to the nonlinear magnetic characteristics. Since the two directions of the PSRM are almost decoupled, the mathematical model of each direction is the same. For the kth phase of l direction, the voltage equation is expressed as ( ( ), ( )) ( ) ( ( ), ( )) ( )( ) ( ) ( ) ( ) ( ( )) ( ) ( ) ( ) ( ) ( ( )) ( ) , , , , lk lk l l lk lk l lk lk lk lk l lk lk l l lk lk lk lk lk l l i t s t s t i t s t i t U t R i t s t t i t t L s t s t i t R i t i t L s t s t t t k A B C l X Y (1) where lkU , lki , lkR , lk and lkL are the phase voltage, phase current, phase resistance, the total flux linkage, and the phase inductance of the kth phase in l direction, respectively. ls is the relative position between stator and mover in l direction. The mechanical movement equation of l axis is given by 2 2 ( ) ( ) ( ( ), ( ), ( ), ( )) ( ) , l l l lA lB lC l l l lp d s t ds t f i t i t i t s t m b f t dtdt l X Y (2) where lf , lm , lb , and lpf are the generated electromagnetic thrust force, the mass of moving platform, the friction constant, and the external load force of l direction, respectively. In terms of the magnetic co-energy Wc, the electromagnetic thrust force of kth phase in l direction is derived by ( ) 0 ( )( ( ), ( )) ( ( ), ( ))= = ( ) ( ) , , , , lki t lk lklc l lk lk lk l l l d tW s t i t f i t s t s t s t k A B C l X Y . (3)" + ] + }, + { + "image_filename": "designv11_71_0002176_icsmd50554.2020.9261740-Figure2-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0002176_icsmd50554.2020.9261740-Figure2-1.png", + "caption": "Fig. 2 A dynamics model of the axle box bearing.", + "texts": [], + "surrounding_texts": [ + "To analyze the dynamic responses of the locomotive under the internal impact force excited by the localized defect of the axle box bearing, based on the locomotive-track coupled dynamics model with gear transmissions which was proposed by Chen et al. in Ref. [7], a locomotive-track coupled dynamics model with traction transmissions is proposed and the internal interactions between the components of the axle box bearing are considered in detail. As shown in Fig.1, the locomotive-coupled dynamics model consists of the vehicle subsystem, the track structure subsystem, the gear transmission subsystem and the rolling bearing subsystem. In the vehicle subsystem, one car body, two bogie frames and four wheelsets are contacted via the primary and the secondary suspensions, respectively. The traction torque is transmitted *Zaigang Chen is the corresponding author. (e-mail: zgchen@home.swjtu.edu.cn). 978-1-7281-9277-2/20/$31.00 \u00a92020 IEEE Authorized licensed use limited to: Rutgers University. Downloaded on May 19,2021 at 14:44:36 UTC from IEEE Xplore. Restrictions apply. 18 from the traction motor to the wheelset through the gear transmission and the longitudinal motion of the locomotive is driven by the tractive forces generated from the wheel-rail interface. An improved mesh stiffness calculation model of spur gear pair with tooth profile deviations [12] is applied to calculate the accurate time-varying mesh stiffness of the gear pair of the traction transmission. In addition, the classical ballasted track structure is used in this dynamics model, which is composed of the rail, rail pads, sleepers, ballasts and subgrade. In order to reflect the impact effect between the roller and the defect region, a dynamics model of rolling bearing considering contact forces between each component and the corresponding friction forces is proposed in this paper. Here, the inner ring is fixed on the wheelset and the outer ring rigidly connects with the axle box. The time-varying displacement excitation and the time-varying stiffness excitation of the failure bearing can be described by an improved analytical model proposed by Liu et al. in Ref. [1]. When a localized defect of the bearing is in the middle or late stage of plastic deformations at the propagated defect edges, the schematic of the contact relationship between the roller and the defect region are shown in Fig.3. It can be seen that there are some visible differences of the relative radial displacement and the contact stiffness between the roller and the outer race. Therefore, the impact interaction between the roller and the defect region can be described as the comprehensive effect of the time-varying displacement excitation and the time-varying contact stiffness excitation induced by the localized defect on the outer race. For a localized defect with the cylindrical surface edges, the time-varying displacement excitation of the localized defect can be represented as a piecewise function [13]: ( )( )( ) ( ) ( )( )( ) ( ) ( ) 1 o 0 o 1 1 o 2 2 o 2 o 3 sin 0.5\u03c0 / mod ,2\u03c0 mod ,2\u03c0 mod ,2\u03c0 sin 0.5\u03c0 / mod ,2\u03c0 mod ,2\u03c0 otherwise0 i i i i i D T D H D T = (1) where D is the depth of the localized defect. \u03b8oi represents the rotational angular displacements of ith roller. \u03b80\\\u03b83 and \u03b81\\\u03b82 are the angular positions of the start and end points of the localized defect and the roller-bottom surface contact area, respectively. \u0394T1 and \u0394T2 can be calculated by: ( )( )2 2 1 2 d r d r o/T T r R r R R = = + + \u2212 (2) where Rr and Ro are the radiuses of the roller and the outer race, respectively. rd is the radiuses of the cylindrical surface edge. According to the Hertz contact theory, the contact stiffness between the roller and the cylindrical surface at the localized defect edge can be deduced by [14]: rd rd d d Q K = (3) where Q is the external force. \u03b4rd is the relative elastic deformation between the roller and the edge, which can be calculated by: 2 r d rd 2 e 0 id 2 1 4 ln 0.814 \u03c0 Q R r l E b \u2212 = + (4) where le is the equivalent length, \u03c5 and E0 denote the Poission\u2019s ratio and equivalent elastic modulus. And bid is the semi-width of the contact surface. The contact stiffness between the roller and the bottom surface of the localized defect can be calculated by [15]: b b d d Q K = (5) where \u03b4b is the relative elastic deformation between the roller and the bottom surface, which can be calculated by: 0.5 2 2 r r o b r o 4 1 1 3 R E E \u2212 \u2212 = + (6) where \u03c5r\\\u03c5o and Er\\ Eo denote the Poission\u2019s ratio and equivalent elastic modulus of the roller and the outer ring, respectively. Authorized licensed use limited to: Rutgers University. Downloaded on May 19,2021 at 14:44:36 UTC from IEEE Xplore. Restrictions apply. Considering the large degrees of freedom of the dynamics model, an explicit-implicit hybrid method is proposed here to achieve the desired calculation speed and accuracy. The fastexplicit integration method [16] is utilized to solve the dynamic equations of the vehicle, the track structure and the gear transmission subsystems. While for the bearing subsystem, a fourth-order Runge\u2013Kutta method [17] is utilized to meet the precision requirements. The integration steps of two numerical methods are 3\u00d710-5 s and 5\u00d710-6 s, respectively. III. DYNAMIC INVESTIGATION AND RESULT DISCUSSIONS To reveal the vibration features of the locomotive axle box in presence of local defect in its outer race, a HX locomotive is used for dynamic simulation in this paper, and its major design parameters can be referenced to Ref. [7]. The locomotive running speed is set as 80 km/h and the traction torque of the motor is 1.5 kN/m. To consider the changes in the contact zones near the defect edges, the defect length is assumed to be larger than the roller length and its width is less than the roller diameter. Thus, the defect length, width and depth are assumed to be 14 mm, 5 mm and 0.1 mm, respectively. And the angular displacement of the start of the localized defect (\u03b80) is \u03c0/2. In addition, the characteristic frequencies of the axle box are listed in Table 1. To highlight the effect of the localized defect of the axle box, time histories and frequency spectrum of the vertical and longitudinal accelerations of the axle box are extracted from the simulation without considering the track irregularity. As shown in Figs. 4 and 5, it can be seen that the peak-peak value of the vertical acceleration increases evidently under the excitation of the defect, while the localized defect has an inconspicuous effect on the axle box in longitudinal direction in time histories. After analyzing the spectrum of the accelerations, the frequency of roller passing the outer race and its harmonics are more evident than the mesh frequency of the gear transmission while there is no obvious change in longitudinal direction. In addition, it should be noted that there are frequency bands of 16 Hz and 29 Hz around the passing frequency of roller (PFOR) and its harmonics. The similar phenomenon also can be observed in Fig. 5(b). Under the effect of the gravity of the locomotive and the interaction between the wheelset and the rail, the load region of the axle box bearing is formed on the top of the outer race. Therefore, the impact force between the roller and the defect region mainly affects the dynamic responses of the axle box in vertical direction while the changes of longitudinal acceleration are weak. 20 Figures 6 and 7 illustrate the effect of the localized defect of the axle box under the external wheel-rail interaction in actual operation of the locomotive. In order to truly reflect the railway conditions, the Association of American Railroads 5 class irregularity is adopted here. It can be seen that there are no significant changes in the time histories of the vertical and longitudinal accelerations. In addition, the frequency of roller passing the outer race and its harmonics become overwhelmingly obvious. The frequency bands of 16 Hz are covered and them of 29 Hz are not apparent under the external excitation from the wheel-rail interface. However, considering the effect of the track irregularity, there is no clear characteristic signal extracted from the dynamic represents of the axle box in the longitudinal direction, which means that it loses value for the localized defect diagnosis to analyze longitudinal acceleration. IV. CONCLUSIONS A locomotive-track coupled dynamics model with traction transmissions is established in this paper by considering the time-varying mesh stiffness and the internal excitation of the rolling bearing, where the impact effect of the localized outer race defect of the axle box bearing is represented by the timevarying displacement excitation and time-varying contact stiffness excitation. And then, dynamic simulations are performed to receive the vibration responses of the locomotive dynamic system, and the vibrations of the axle box are extracted for analysis under the complicated excitations. The results indicate that the localized outer race defect of the axle box bearing is mainly reflected in the vertical vibration accelerations. The passing frequency of rollers and its harmonics are the key characteristic signals, and the frequency bands around the passing frequency of rollers and its harmonics can be a good indicator for the axle box bearing fault detection." + ] + }, + { + "image_filename": "designv11_71_0003352_9783527673148.ch16-Figure16.5-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0003352_9783527673148.ch16-Figure16.5-1.png", + "caption": "Figure 16.5 Packaging of the second-generation biofuel cell implanted in a rat.", + "texts": [ + " Both bagged electrodes were then placed in a second, external dialysis bag and finally sutured inside an exPTFE coating for biocompatibility, leading to an implant of around 10 ml [13] (Figure 16.4). Even when this setup showed good appropriateness, improvements were targeted in terms of size of the implant and the diffusion of glucose and oxygen to the electrodes. The second generation packaging consisted in the coating of each bioelectrode with a dialysis membrane and their face-to-face placement in a perforated silicone tube (Figure 16.5). This not only allowed mechanical stabilization of the bioelectrodes but also fixed the distance between these electrodes. Further packaging was similar to the first-generation design except that instead of the exPTFE tissue a Dacron\u00ae sleeve was used for improved biocompatibility and wettability, enabling improved diffusion of the body fluids through this membrane. The final implant had a volume of 2.4 ml [20]. Both of these implants did not provoke any inflammations even after several months in the rat and a highly vascularized tissue covered these implants" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0000254_ecce.2019.8911883-Figure2-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0000254_ecce.2019.8911883-Figure2-1.png", + "caption": "Fig. 2. Flux barrier effect of a conventional FSPM (Zs=12, Zr=17, Pa=1).", + "texts": [ + " According to [8], the torque of the FSPM machine can be calculated as: 1,2,3... 3 2 r s g stk rms i i a Z T N r l I B P= = (1) where Zr and Pa are the rotor tooth number and stator armature winding pole pair number, respectively, and Bi is the amplitude of flux harmonics, which contributes to the fundamental back-EMF. From Eqn. 1, it can be seen that the higher pole ratio, the larger average torque of the FSPM machine. However, the output torques of the FSPM machines with high pole ratio are not as good as the expectation. As illustrated in Fig. 2, it can be seen that the main flux in the stator core is \u201ccompelled\u201d into the air-gap or outer air region because of the reversely excited PMs, which leads to a long flux path and high magnetic resistance. From the viewpoint of the flux path, the reversely excited PMs are flux barriers, so this phenomenon is called as \u201cflux barrier effect\u201d [21]. One can find that the higher pole ratio, the more reversely excited PMs, and thus the severer \u201cflux barrier effect\u201d and lower torque density. Therefore, this paper will aim to solve the \u201cflux barrier effect\u201d of FSPMs and increase the torque density at high pole ratios" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0001205_0954410020926769-Figure1-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0001205_0954410020926769-Figure1-1.png", + "caption": "Figure 1. Missile-target relative motion in the pitch plane.", + "texts": [ + " , n) and 8 05 p41, the following inequality holds x1j j \u00fe x2j j \u00fe \u00fe xnj j\u00f0 \u00de p4 x1j j p\u00fe x2j j p\u00fe \u00fe xnj j p \u00f03\u00de Missile-target relative motion model The skid-to-turn missile has axially symmetrical shape, and adopts the roll angle position stability design, so the three channels can be decomposed into the pitch plane motion and the yaw plane motion. The design method of pitch plane motion is similar to the yaw plane motion, therefore, this paper takes missile-target motion process in pitch plane as an example to analyze, which is illustrated in Figure 1. In Figure 1, M is the missile centroid and T is the target. Respectively, the missile-target relative distance and its change rate is represented by r and _r. Vt and Vm respectively respect the speeds of target and missile. The LOS angle is represented as q, and its derivative is the LOS angle rate _q. The missile ballistic angle and target flight-path angle are respectively written by t and m. amq and atq are respectively used to represent the acceleration components of missile and target in the normal direction of LOS in the pitch plane" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0001402_s12239-020-0083-y-Figure8-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0001402_s12239-020-0083-y-Figure8-1.png", + "caption": "Figure 8. Contact characteristics of roller-housing and ballinner/outer race. Figure 9. External and internal contact.", + "texts": [ + " Simulation Modeling of CV Joint In this study, a multibody dynamic simulation was performed using the CV joint model used on a C-segment vehicle. The size of the CV joint is TJ # 95 and BJ # 95. Based on the above model, the dynamic characteristics of the CV joint were evaluated using DAFUL, a multibody dynamic-analysis software package (Virtual motion Inc., 2008). During the simulation modeling of the TJ, the sphere-cylinder contact element is applied to the rollerhousing contact. As can be seen in Figure 8 (a), the contact of the roller-housing in the Y-Z plane generates an internal contact force. The roller and trunnion of the TJ are supported and rotated by the needle roller. However, in order to reduce the analysis speed of the simulation and overcome the convergence problem, the needle roller contact between the roller and the trunnion is simplified to a cylindrical joint. For a BJ, the contact in the X-Z plane is shown in Figure 8 (b), where it can be seen that the ball and inner race generate an external contact force, and the ball and outer race an internal contact force. As can be seen in Figure 8 (c), in the Y-Z plane both the inner and outer races generate contact forces from contact with ball. In the BJ simulation model, the contact of the inner race, cage, and outer race has been simplified to a spherical joint because the center of rotation has the same kinematic characteristics. The friction force, contact normal force, and penetration for the external and internal contacts are shown in Figure 9. When two objects are in contact, a normal contact force is generated in the normal direction on the contact surface, and a friction force is generated tangential to the contact surface" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0001432_acpee48638.2020.9136376-Figure2-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0001432_acpee48638.2020.9136376-Figure2-1.png", + "caption": "Fig. 2. Simplifilied diagram of the three-phase IPMSM", + "texts": [ + " The permanent magnet flux linkage and d- and q-axes inductances are simultaneously estimated using the MRAS technique, and these estimated parameters are used for the MTPA control method. Finally, the effectiveness of the proposed adaptive MTPA control method is verified by simulation studies. 978-1-7281-5281-3/20/$31.00 \u00a92020 IEEE 689 Authorized licensed use limited to: UNIVERSITY OF WESTERN ONTARIO. Downloaded on July 26,2020 at 15:19:15 UTC from IEEE Xplore. Restrictions apply. II. MTPA CONTROL OF IPMSM A. Mathematical Model of IPSMSM The simplified diagram of a three-phase IPMSM is shown in Fig. 2. In order to obtain the mathematical model of the IPMSM under the q- and d-axes, the following assumptions must be made on the motor system. First, ignoring the distribution of the air gap magnetic field by the cogging and the space magnetic field is sinusoidal. Then, the permanent magnet has no damper windings, the electrical conductivity is zero, and the magnetic permeability is the same as air. Finally, The magnetic permeability of the rotor core is infinite and the machine has negligible eddy currents and hysteresis losses" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0003629_itec-ap.2014.6940995-Figure3-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0003629_itec-ap.2014.6940995-Figure3-1.png", + "caption": "Fig. 3. Location of test points", + "texts": [], + "surrounding_texts": [ + "When permanent magnet motor is under inverter supply, the current loaded in the stator windings contains different harmonics then the current will have different degrees of distortion. In the three-phase bridge inverter-driven sinusoidal pulse width modulation (SPWM) circuit, because of using symmetric modulation waveform, the even harmonic currents will not be contained in the output voltage. The fundamental component of the output line voltage can be expressed as: )sin( 2 3 tmUU rdab (1) Where m is phase\uff0c dU is the bus voltage and r is the power supply frequency. The equation (1) can be expanded using the Fourier series and the Bessel function of the first kind as following equation, here are two kinds of situations. (1) When the harmonic component of output line voltage is n=1,3,5\u2026\uff0ck=2,4,6\u2026, 1 2 2 1 2) 2 () 4 ()1( n k k n ab mn J n H )] 3 (){sin[( ktnk cr ))]} 3 ()sin[( ktnk cr 3 sin k (2) It can be seen from equation (2) that the angle frequency of the harmonic component of output line voltage is cr nk and the amplitude of harmonic is ) 2 () 4 ( 2 3 mn J n U kd . But when k equals to a multiple of 3, 3 sin k is zero, the value of K is an even but not a multiple of 3. (2) When the harmonic component of output line voltage is n=2,4,6\u2026\uff0ck=1,3,5\u2026, 1 2 3 4 5 6 7 8 9 10 11 12 13 14 15 16 17 18 19 20 21 22 23 24 25 26 27 28 29 30 31 32 33 34 35 36 37 38 39 40 41 42 43 44 45 46 47 48 49 50 51 52 53 54 55 56 57 60 61 1 1 2 2) 2 () 4 ()1( n k k n ab mn J n H )] 3 (){cos[( ktnk cr ))]} 3 ()cos[( ktnk cr 3 sin k (3) From equation (3) can be seen that the angle frequency of the harmonic component of output line voltage is still cr nk , the amplitude of harmonic remain unchanged, the value is ) 2 () 4 ( 2 3 mn J n U kd . The value of K is an odd but not a multiple of 3. The phase of n-th harmonic currents are 120n \u00ba difference each other in symmetric system, show as equation(4): )cos(2 nnan tnIi ) 3 2 cos(2 nnbn n tnIi ) 3 2 cos(2 nncn n tnIi (4) As can be seen from the equation(4), different harmonic current has different law in circuits.When n=3k(k=1,2,3...), then the amplitude of the harmonic current of each phase are equal and the phase is same, so that the system does not exist harmonics in three-phase three-wire system which stator connected in star;when n=3k+1(k=1,2,3\u2026), the harmonic current amplitude of each phase remains equal, but the phase are 120 \u00ba difference each other which can be considered positive sequence harmonic current; when n=3k-1(k=1,2,3\u2026), he harmonic current amplitude of each phase remains equal, but the phase are 120 \u00ba difference each other which can be considered negative sequence harmonic current. As previously mentioned, the angular frequency of harmonic components of the line voltage in three-phase inverter bridge SPWM inverter circuit is cr nk through the derivation, modulation angular frequency is r and carrier angular frequency is c ,but the greater impact of harmonic to motor is refers to the frequency of harmonic current nearby rkf . Due to the prototype of this study is threephase three-wire system, the stator is connected in star, third harmonic and ninth harmonic current do not exist in motor winding, there is also no even harmonic current, the influence of low harmonic current to motor mainly refers to the fifth and seventh harmonic current, So this research is mainly studies the vibration and noise of permanent magnet motor under the fifth and seventh harmonic. III. SIMULATION OF ACOUSTIC FIELD AND VIBRATION In this paper, the finite element method(FEM) is used for prototype modeling, subdivision and simulation, mainly studies the vibration and noise of permanent magnet motor under the fifth and seventh harmonic. Studies have shown that the amplitude of harmonic current changes from 1% to 15% of the amplitude of fundamental current when the motor under inverter supply condition [7]. In this paper, 5%, 10% and15% of fundamental amplitude are taken as the harmonic amplitude of the fifth and seventh harmonic.Assuming that the initial phase of harmonic currents and fundamental current are same,study on the effects of different amplitude and different times of harmonic current to vibration and noise of permanent magnet motor. When the stator windings exist harmonic currents, the current of the permanent magnet synchronous motor can be expressed as following equation . )cos(2)cos(2 nna tnItIi (5) FEM is used to simulate on the motor to extract the magnetic flux density\uff0c then using FEM to simulate of acoustic field and vibration to extract Sound pressure datas and acceleration datas and acoustic field. Using the equation of (6),(7)and(8) to obtain sound pressure level(SPL), average SPL and total SPL, the method of obtaining vibration acceleration level and sound pressure level is the same. This method is used to calculate sound pressure level and vibration acceleration level of the motor under different amplitudes of 5th and 7th harmonics superimposed with fundamental. 1 2 3 4 5 6 7 8 9 10 11 12 13 14 15 16 17 18 19 20 21 22 23 24 25 26 27 28 29 30 31 32 33 34 35 36 37 38 39 40 41 42 43 44 45 46 47 48 49 50 51 52 53 54 55 56 57 60 61 motor cover under the superposition of different harmonics and fundamental, the three points are AX1, AX2 and AY1. As can be seen from the Figure 1(a,b,c), table1 and table2 , due to the existence of harmonic current, the sound pressure level and vibration acceleration level of the motor is decreased. Fig.1. Sound pressure curve a. The SPL curves of the motor under fundamental alone b\uff0e The SPL curves of the motor under 5th harmonic and fundamental superimposed c. The SPL curves of the motor under 7th harmonic and fundamental superimposed Fig.2. Sound pressure level curve a. The vibration acceleration curve of AX1 b. The vibration acceleration curve of AX2 c. The vibration acceleration curve of AY1 Fig .4. The vibration acceleration curve of three points 1 2 3 4 5 6 7 8 9 10 11 12 13 14 15 16 17 18 19 20 21 22 23 24 25 26 27 28 29 30 31 32 33 34 35 36 37 38 39 40 41 42 43 44 45 46 47 48 49 50 51 52 53 54 55 56 57 60 61 TABLE I THE SOUND PRESSURE LEVEL OF FUNDAMENTAL AND HARMONICS SUPERPOSED LP1/dB LP2/dB LP3/dB LPT/dB Lav/dB fundamental 70.72 70.73 70.16 77.70 70.71 fundamental+ 5th harmonic 68.14 67.93 68.33 75.09 68.10 fundamental+ 7th harmonic 69.52 69.65 69.81 76.62 69.62 TABLE \u2161 THE VIBRATION ACCELERATION LEVEL OF FUNDAMENTAL AND HARMONICS SUPERPOSED AX1/dB AX2/dB AY1/dB fundamental 85.90 86.56 87.38 fundamental+5th harmonic 85.47 86.01 87.42 fundamental+7th harmonic 84.63 85.38 86.30" + ] + }, + { + "image_filename": "designv11_71_0001617_icst47872.2019.9166418-Figure3-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0001617_icst47872.2019.9166418-Figure3-1.png", + "caption": "Fig. 3. Steering model of a vehicle", + "texts": [ + " ae , be , and ce are trapezoidal back EMFs. The electromagnetic torque of the motor can be found as: a a b b c c e m e i e i e i T \u03c9 + += or e m L m d T J T B dt \u03c9 \u03c9= + + (2) 978-1-7281-2369-1/19/$31.00 \u00a92019 IEEEAuthorized licensed use limited to: UNIVERSITY OF WESTERN ONTARIO. Downloaded on August 24,2020 at 01:33:48 UTC from IEEE Xplore. Restrictions apply. Fig. 2 demonstrates the three-phase voltage source inverter fed to the BLDC motor model. The Ackermann-Jeantand steering model of a vehicle is given in Fig. 3. This model was found by Rudolf Ackerman in the 19th century [8]. When the vehicle drives on a straight road, the rotating speed of each rear wheel is the same. However, on a curved road, the rotating speed of the outer wheel must be higher than the rotating speed of the inner wheel. Thus, the wheel speed mathematical equations are provided [9]: _ 60 . .D RL ref RL cg V R n R \u03c0 = ; _ 60 . .D RR ref RR cg V R n R \u03c0 = ; 60 ref V n D\u03c0 = (3) Where _ref RLn is the rotating speed at rear left side, _ref RRn is the rotating speed at the rear right side, refn is the rotating speed at the center of the vehicle and V is a linear speed of the vehicle" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0002047_ecce44975.2020.9235956-Figure1-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0002047_ecce44975.2020.9235956-Figure1-1.png", + "caption": "Fig. 1. Topologies of two AS-CP-HEFRMs.", + "texts": [ + " The key points, including the PM polarities, operating principles, the phenomenon and elimination of the phase shift between the flux linkages of PM and DC excitations, the relationship between the rotor pole number, the stator pole arc and PM ratio (which is defined as the ratio of PM pole arc to stator pole arc) for maximum average torque is derived and the influence of stator/rotor pole combinations on torque, torque ripple, and flux regulation capability is investigated by finite element analyses (FEA). II. MACHINE TOPOLOGY Topologies of the two machines are shown in Fig. 1. PMs are magnetized radially in the same direction (if PM number on each stator pole is over 1) on one stator pole, while in the opposite directions on adjacent stator poles. For two topologies, PM arrangements are different. In Type I, PMs and iron poles are alternately arranged, Fig.1(a), while Type II exhibits a PM-to-PM arrangement, Fig.1(b). As mentioned above, these two hybrid machine topologies are developed from the original CP-FRPMMs in [3] and [4], which have all-N-pole (or all-S-pole) PMs on each stator pole. Due to the introduction of DC excitation in the developed 978-1-7281-5826-6/20/$31.00 \u00a92020 IEEE 14 Authorized licensed use limited to: UNIVERSITY OF CONNECTICUT. Downloaded on May 19,2021 at 20:47:17 UTC from IEEE Xplore. Restrictions apply. hybrid machine topologies, the polarities of PMs on adjacent stator poles of AS-CP-HEFRMs are no longer the same compared with their original CP-FRPMM counterparts", + " Comparisons are made on these two PM polarity configurations for the two types of machines, as shown in Fig. 2. Figs. 2(a)-(b) and (e) show the flux distributions and flux-linkages of the two PM polarity configurations in DC-only excitation for Type I (PMs are removed, same for Type II), while Figs. 2(c)-(d) and (f) are for Type II. From Figs. 2(a) and 2(c), both Type I and II machines with all-N-pole PM configuration suffer from significant flux leakages between adjacent stator poles due to the same polarity of DC produced MMFs (see Fig.1 for coil excitation polarities). However, flux leakages in AS-CP-HEFRMs with alternate-NS-pole PMs are negligible, as shown in Figs. 2(b) and 2(d), and hence, obvious differences in flux-linkages produced by DC excitations of the two configurations can be obtained in Figs. 2(e) and 2(f). The contribution of DC excitation can almost be neglected in machines with all-Npole PMs. Consequently, the alternate-NS-pole arrangement is preferred in the two AS-CP-HEFRMs. III. OPERATING PRINCIPLE AND FLUX DISTRIBUTIONS From the topologies of AS-CP-HEFRMs, it can be seen that iron poles on stator poles provide flux paths for DC field excitation (both Types I and II), which make it possible for DC-produced magnetic field to pass through the air gap from ferromagnetic routes rather than go across PMs directly", + "), it can be seen that there is less saturation in stator yoke and big teeth than that of one excitation (DC or PM only), which verifies flux cancellation in AS-CP-HEFRMs. IV. PHASE SHIFT OF PM AND DC EXCITATIONS For AS-CP-HEFRMs, the phase shift between the flux linkages of PM and DC excitations may be caused because the central lines of magnetic fields produced by PM and DC excitations are not coincident, which will sacrifice the maximum average torque and the flux regulation capability. The illustration will focus on flux-linkages of two coils belonging to one phase. Take phase A as an example, which is made up of Coil 1 and Coil 4, Fig.1. Assume the central lines of magnetic field and poles (PM poles or iron poles) are the same. For Type I, the initial rotor position is shown in Fig. 1(a). It should be noticed that the number of rotor pole is odd. In only PM excitation operating mode, if only the DC bias and fundamental harmonic are considered, flux densities across Coil 1 (A1) and Coil 4 (A2) can be expressed as: = + cos( + ) (1) = \u2212 cos( + ) (2) Therefore, the flux-linkages of Coil 1 ( ) and 4 ( ) can be expressed as: = \u2219 d = ( + cos( + )) (3) = \u2212 \u2219 d= \u2212 ( \u2212 cos( + )) (4) where is the number of turns per coil, is the area of single coil, the initial phase angle of can be written as (5), where is the rotor pole number, and is the arc of each iron pole", + " In fact, there will be offsets between the PM (DC) produced flux and the geometric central line of PM (iron) pole. If the offsets are considered (shown in Fig. 6(a)), (10) will be changed to (16), where is the PM field angle offset and is the DC field angle offset. + + ( \u2212 ) = 180\u00b0 (16) The amplitudes of and are quite small, which can also be obtained in Fig. 5(d). Therefore, the method of phase shift elimination in (10) is applicable in Type I AS-CP-HEFRMs. For Type II, the initial rotor position is shown in Fig. 1(b). It should be noticed that the number of rotor pole is even and the offsets of central lines of magnetic field and poles are neglected. In only PM excitation operating mode, if only the DC bias and fundamental harmonic are considered, will 17 Authorized licensed use limited to: UNIVERSITY OF CONNECTICUT. Downloaded on May 19,2021 at 20:47:17 UTC from IEEE Xplore. Restrictions apply. be the same as (1) and flux density across Coil 4 (A2) can be expressed as: = + cos( \u2212 ) (17) Thus, the flux-linkage of Coil 4 will be: = \u2212 \u2219 d = \u2212 ( + cos( \u2212 )) (18) Therefore, the phase flux-linkage of PM excitation will be: = + = \u22122 sin sin (19) Similarly, in flux-enhancing operating mode, the phase fluxlinkage of DC excitation will be: = \u22122 sin sin (20) In hybrid operating mode (flux-enhancing), the flux-linkage of phase A ( ) will be: = + = \u22122 ( sin + sin ) sin (21) From (19) and (20), it can be seen that there will be no phase shift in the flux-linkages of PM and DC excitations in Type II inherently, which can be verified by FEA in Fig" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0003876_ipec.2014.6869595-Figure9-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0003876_ipec.2014.6869595-Figure9-1.png", + "caption": "Fig. 9. Waveforms of primary currentl.", + "texts": [], + "surrounding_texts": [ + "The ring specimen for evaluation of iron loss is composed of laminated electrical steel sheet (35H300). The specifications of this ring specimen are shown in Table II and three rings are made so that the specifications on each ring become the same as much as possible. To carry out the three phase inverter excitation, these three rings are connected in star type and in delta type as shown in Fig. 2 since these connection method are most general in three phase excitation. As a feature of star connection in Fig. 2(a), the neutral point exists. On the other hand, as a feature of delta-connection in Fig. 2(b), the zero-phase current circulating in circuit when the circuit is unbalanced exists. Usually, on motor driving system, the star-connection is more commonly used. C. Calculation of Iron Loss In order to calculate iron loss, the primary current I and the secondary voltage V on each ring are obtained by using AD converter. Magnetic field intensity H and magnetic flux density B in ring are calculated by following equations ( 1), (2). The iron loss W is obtained from integration of Hand B as following equation (3). H, = NlI , [Aim] ( 1) (2) (3) where NI, N2, L, Sand p in equation are the primary windings, the secondary windings, the magnetic path length, the cross section area and the density of electrical steel sheet, respectively. The subscript (i) indicates each ring number. In measurement, since the maximum flux density Bmax does not become just 1 T, the normalization of iron loss on just Bmax = 1 T is performed by using following equation (4) because iron loss is proportional to the square of magnetic flux density. Wn = \ufffd X 12 I Bma/ [W/kg] (4) where Wn and Bmax are the normalized loss and the measured maximum magnetic flux density, respectively. III. EXPERIMENTAL RESULTS The iron losses obtained from star-connection and delta-connection are shown in Table III. In addition, the iron losses on single ring by using u-v phase of three phase PWM inverter excitation are also shown for reference in Table III. First, it is obtained as theoretical that the applied DC voltage Vdc of star-connection is the square root of three times that of delta-connection. In Table III, although the iron losses of single ring and delta-connection are almost the same, the iron losses of star-connection are about 14 % larger than that of delta connection. Fig. 3 shows the BH curves obtained from star-connection and delta-connection on ring 1. When both BH curves are compared with, BH curve of star connection is expanded in direction of H than BH curve of delta-connection. Iron loss is calculated by equation (3) and this equation means that the iron loss is proportional to internal area of BH curve. Therefore, in star-connection with large internal area, iron losses become larger than delta-connection. In next chapter, the cause by which the BH curve is expanded on star-connection is discussed. -1.0 -150 -100 -50 o 50 100 Magnetic field intensity H [Aim] 150 First, waveforms of magnetic flux density B and magnetic field intensity H used for making BH curves are shown in Figs. 6 and 7, respectively. In Fig. 6, although magnetic flux density of delta-connection becomes sinusoidal waveform, that of star-connection is distorted. In Fig. 7, the waveforms of magnetic field intensity obtained from star-connection and delta-connection differ from each other. These factors cause the difference of BH curve. In particular, when point P in which Band H of delta-connection and star-connection are the same becomes basis, B of star-connection becomes the minimwn early than that of delta-connection. As a result, since H of star-connection becomes the minimum earlier than that of delta-connection, H of star-connection begins to change a little early. This phenomenon causes the expanding in direction of H. Next, reason which the distortion of magnetic flux density of star-connection is caused is verified. Figs. 8 and 9 show the secondary voltage waveform and the primary current waveform. Especially, voltage of star connection has third harmonic. Magnetic flux density B is calculated by using secondary voltage V as equation (2), so this third harmonics cause the distortion of B of star connection. The current waveforms of star-connection and delta-connection also differ due to difference of voltage waveform. We have described the relationship between iron loss and minor loop of BH curve [3]-[4], so the enlargement of minor loop near maximum flux density is shown in Fig. 10. However, the minor loop obtained from star connection is not clearly, and it seems that the iron loss increase by expanding in direction of H on star connection is dominant. Therefore, in this paper, only influence of waveform distortion is discussed. V. THEORETICAL VERIFICATION Third harmonics on voltage waveform of star connection is caused by potential on neutral point. In this chapter, by theoretical verification using a simple rectangular wave, the cause which potential on neutral point sways is explained. Fig. 1 1 shows the three phase inverter and the three loads connected in star type. It is assumed that impedances Z of each load are the same. As a switching signal, on-signal or off-signal is applied in each switching device Sj as shown in Fig. 12. The bar (-) means negation. At the time of (i) in Fig. 12, the electrical circuit becomes Fig. 13(a). In this case, the voltage Vmo between neutral point M and ground 0 becomes as following equation (5). On the other hand, at the time of (ii) in Fig. 12, the electrical circuit becomes Fig. 13(b) and the voltage Vmo becomes as following equation (6). Similarly, the waveform calculated to from (i) to (vi) is shown in Fig. 12. The waveform of Vmo has three times period in one period of switching waveform. As a result, the Vum becomes as shown in Fig. 12. This Vum corresponds to voltage of star-connection in Fig. 8. In Fig. 12, it is confirmed that the Vum does not sway by the third harmonics of neutral point potential theoretically. However, actually, impedances of three ring specimens are not the same because the characteristics differ in each ring as shown in Table III, so the electrical circuit must become unbalanced. By this unbalanced circuit, it is presumed that Vmo is distorted and Vum obtained from Vu and V mo has just third harmonics resulting from this unbalance. VI. CONCLUTION In this research, the iron loss characteristics in star connection and in delta-connection are evaluated by three phase PWM inverter excitation. The iron loss obtained from star-connection is about 14 % larger than the iron loss obtained from delta-connection. This difference is caused by expanding BH curve of star-connection due to potential of neutral point having the third harmonics on star-connection. Therefore, although the zero-phase current may flow, the driving by delta-connection should be carried out in order to reduce the iron loss." + ] + }, + { + "image_filename": "designv11_71_0001575_ls.1519-Figure2-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0001575_ls.1519-Figure2-1.png", + "caption": "FIGURE 2 Coordinate systems for the spindle vibration motion analysis", + "texts": [ + " Firstly, the small displacement perturbation in the lateral direction of the spindle is imposed, and each physical quantity can then be regarded as summation of a steady term and a time-varying term,6,19,20 for example, the pressure p in the air film can be expressed as p= p+ p0, \u00f01\u00de where p is the steady pressure and p 0 is the time-varying pressure. It is also supposed that the time-varying term is rather small compared with the steady one. The load carrying force FL is the integration of the pressure acting upon the spindle surface As in the air film as FL = \u00f0 As pdS= \u00f0 As pdS+ \u00f0 As p0dS= FL +F 0 L: \u00f02\u00de Hence, FL can also be decomposed into a steady part FL and a time-varying part F 0 L. In the analysis, the used coordinate systems are shown in Figure 2.where Ob is the center of the bearing, and Oj is the center of the spindle. ObXY is a fixed coordinate system, and Ojx0y0 is a moving coordinate system. Axis Ojx0 is along the direction of OjOb. G is an arbitrary point fixed in the air film, \u03b1 is its azimuth angle in ObXY and \u03d1 is its azimuth angle in Ojx0y0. The position of Oj in ObXY can be determined by the eccentricity e and the azimuth angle \u03c6. In ObXY, the motion equation of the spindle is expressed as m\u20acr=Fext +FL, \u00f03\u00de where m is the spindle mass", + " At the static equilibrium position, Fext is eliminated by FL, shown as FL xp,Ps, e, \u03c6 + Fext =0: \u00f06\u00de Hence, Equation (3) can be rewritten as m\u20acr0 =F 0 L +F 0 ext: \u00f07\u00de The calculation of the load carrying capacity, that is, FL = FL xp,Ps, e, \u03c6 , is the main task for the static analysis. In the dynamic analysis, the vibration of the spindle is of two degrees of freedom vibration.6,21-24 Hence, the following eight parameters, the stiffness Kxx, Kxy, Kyx, Kyy and the damping Dxx, Dxy, Dyx, Dyy, are required in illustrating the dynamic performances. As shown in Figure 2, the motion of Oj is described by e(t) and \u03c6(t) in ObXY, or equivalently by x(t) and y(t) in the rectangular form. Similarly, x(t) and y(t) can also be decomposed into a steady part and a time-varying part as formulated in Equations (4) and (5). At the static equilibrium position, e and \u03c6 or x and y are governed by Equation (6) and the vibration equations can be expressed by e 0 and \u03c6 0 or x 0 and y 0 , for example, by x 0 and y 0 , it can be expressed as: m\u20acr0 +D_r0 +Kr0 =F 0 ext, \u00f08\u00de where r = x0 y0 , D= Dxx Dxy Dyx Dyy , K = Kxx Kxy Kyx Kyy ", + " Once performances of each divided air film part are achieved, the journal bearing mechanical performances can be calculated by assembling performances of all the divided air film parts in global coordinate system ObXY. Different from analysis of the thrust bearing, in the dynamic calculation of the journal bearing, the two degrees of freedom vibration system should be considered. Hence, different way of CFD simulation should be explored. Firstly, the small displacement perturbation is imposed on the spindle center. The displacement excitations discussed in Equations (4) and (5) are usually defined in a moving coordinate system as shown in Figure 2, which can be described in the complex form6 as e0 = ece i\u03c9t \u00f010\u00de \u03c60 =\u03c6ce i\u03c9t, \u00f011\u00de where ec and \u03c6c are excitation amplitudes and \u03c9 is the excitation frequency. The air film thickness at any point is expressed in Equation (12). h= h0 1+ \u03b5cos\u03d1\u00f0 \u00de= h0 + e t\u00f0 \u00decos\u03d1: \u00f012\u00de According to the geometrical relation, and considering the fact that e and \u03d1 are composed by the steady and time-varying parts, h can be decomposed into the steady part (13) and time-varying part (14) as: h= h0 + ecos \u03b1\u2212 \u03c6\u00f0 \u00de \u00f013\u00de h0 = e0cos \u03b1\u2212 \u03c6\u00f0 \u00de+ e\u03c60sin \u03b1\u2212 \u03c6\u00f0 \u00de: \u00f014\u00de Hence, the time-varying air film thickness h 0 can be expressed as: h0 = hce i\u03c9t, \u00f015\u00de where the amplitude hc satisfies: hc = eccos \u03d1+ e\u03c6csin \u03d1: \u00f016\u00de The basic idea of the CFD simulation is that the steady part ofFL forms the static load carrying capacity, and the stiffness and the damping come from the squeeze film effect of the air film as discussed in Equation (9)", + " For the dynamic performances, theoretically, each divided air film part can be equivalently considered as a spring-damping system as shown in Figure 5. The vibration system of the journal bearing contains n groups of the spring-damping system. Once Ki and Di (i = 1,2,\u2026,n) are achieved, they can be transferred into the global coordinate system ObXY by the coordinate-transformation matrix T expressed as T = cos\u03b8 sin\u03b8 \u2212sin\u03b8 cos\u03b8 , \u00f018\u00de where \u03b8 is the included angle between OlocXloc axis of the ith part and ObX axis. Actually, \u03b8 equals \u03c0-(\u03d1i + \u03c6) as shown in Figure 2, while the eccentricity e is far smaller than the bearing diameter \u03a6, so the azimuth angle \u03b1i of the ith orifice in ObXY can be used and \u03c0-\u03b1i can be regarded as \u03b8 approximately. Ki and Di can be transferred into the global coordinate system ObXY as shown in Equations (19) and (20). Ki xx Ki xy Ki yx Ki yy \" # =TT Ki 0 0 0 T \u00f019\u00de Di xx Di xy Di yx Di yy \" # =TT Di 0 0 0 T: \u00f020\u00de Finally, the stiffness and damping of the journal bearing can be achieved by assembling stiffness and damping of all the divided air film parts, as formulated in equations (21) and (22)" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0003286_9781118869796.ch16-Figure16.4-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0003286_9781118869796.ch16-Figure16.4-1.png", + "caption": "FIGURE 16.4 Schematic image of a bipolar electrode: view of anode (a) and cathode (b) sides. (Reproduced with permission from Ref. [16]. Copyright 2010, The Electrochemical Society.)", + "texts": [ + " Even though the contactors are not in direct contact with the fuel/electrolyte, fuel contamination might occur (or any aerosol contamination from the external environment), leading to corrosion at the interface between the contactor and the bipolar plate, which in turn increases the impedance of the stack cell. Graphite composite materials provide high electronic conductivity, chemical stability, and sufficient mechanical properties at certain thickness for the contactor application. Alternatively, metal contactors (stainless steel or aluminum) might provide a thin and lightweight alternative for the contactors. A schematic image of a bipolar electrode is shown in Figure 16.4. The cathodic side has an integrated serpentine channel for feeding and distributing oxygen (oxidizer), and the anodic side has a rectangular compartment for anode integration. Two independent feed channels are also visible through the bipolar plate that allows for fuel filling and oxidized feeding in series throughout the entire stack. A schematic image and photograph of a five-EFC stack with bipolar electrodes is shown in Figure 16.3 (bottom). The stack has two integrated lines for fuel and oxidizer embedded in the bipolar plates and two pairs of ports for the lines are visible at the end plates" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0001057_0021998320920920-Figure1-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0001057_0021998320920920-Figure1-1.png", + "caption": "Figure 1. Dimensions of 3D textile model. (a) Analysis region. (b) Analysis region with the matrix removed to show tows.", + "texts": [ + " Orthogonally woven textiles are composed of layers of warps and wefts with binder tows weaving through the entire thickness. Binders undulate through the thickness of the textile and travel generally in the x-direction. Warps are relatively straight tows that are approximately aligned with the x-axis, and wefts are relatively straight tows that are approximately aligned with the yaxis. Without the binders, the warps and wefts are similar to a crossply laminate but with some space in between tows. Refer to Figure 1 for an example of a 3D orthogonally woven architecture. The binders provide the 3D component and provide the improved outof-plane properties and delamination resistance. However, the analysis of 3D textiles presents many challenges. Realistic models are difficult to create due to the complex tow architecture, meshes are often large and require significant computational resources, and stresses 1US Air Force Research Laboratory, USA 2Aerospace Department, Texas A&M University, USA Corresponding author: John D Whitcomb, Texas A&M University, 3141 TAMU College Station, Texas 77843, USA", + " Consequently, the cross-sections for each tow are shrunk towards the centroid by a small amount, resulting in some matrix material between each tow, removing any interpenetrations between tows. It should be noted that requiring a small amount of matrix between each tow is not realistic, but it is a straightforward method for creating a compatible mesh that is used by most of the textile modeling community, including very recent publications of Yan et al.27 For the final steps in VTMS, the surface geometry of each tow is discretized and clipped to a smaller region to avoid spurious effects near the edges of the initial model. Figure 1(a) shows the resulting model with the matrix added, along with the dimensions of the model. The model consists of four layers of warps and five layers of wefts through the thickness, and the binders are in a 2 2 pattern where they cross two wefts at the top or bottom of the weave before traversing through the thickness, which is known as a twill orthogonal weave.28 The region where a binder crosses the wefts before traversing the thickness is sometimes referred to as the z-crown. Figure 1(b) shows the tow architecture more clearly by removing the matrix and highlights the variation of cross-section shape within the tows. Because VTMS relies on contact to compact the textile geometry, the outermost tows experience more deformation than the tows nearest the midplane, which is shown in Figure 1(b). The extreme deformation of the outermost tows is reduced by compacting one layer at a time, but the outermost tows will still experience more deformation than the interior tows. The tow VF is important to consider for a textile model. As shown in Figure 2(a), there are four unitcells within the clipped region for postprocessing, and each unit-cell is expected to have some variation due to how VTMS simulates the compaction and relaxation of the digital chains. Table 1 shows the VF of each type of tow in each unit-cell shown in Figure 2, with the difference in the tow VF between unit-cells illustrating how much the tow architecture varies throughout the model", + " Boundary conditions and clipped analysis region Ideally, periodic boundary conditions would be applied to a textile unit-cell, but the geometry from VTMS is not periodic since it creates a model by virtually simulating textile processing. Consequently, a larger section of the textile is modeled, simple boundary conditions are applied to the larger analysis region, and a subregion is used for post-processing to reduce artifacts due to the boundary conditions. Three faces of the full analysis region (the x \u00bc 0, y \u00bc 0, and z \u00bc 0 planes) are assumed to be planes of symmetry. Refer to the coordinate system in Figure 1(a). This results in the following boundary conditions: u 0, y, z\u00f0 \u00de \u00bc 0, v x, 0, z\u00f0 \u00de \u00bc 0, and w x, y, 0\u00f0 \u00de \u00bc 0, where u, v, and w refers to the displacement along the global x-, y-, and z-axis respectively. In addition to these boundary conditions, the displacement along the warp direction at x \u00bc 13:4 mm is specified to result in 1% volume average strain, specifically specifying u 13:4 mm, y, z\u00f0 \u00de \u00bc 0:134 mm. All other boundaries are traction free. As mentioned before, the boundary conditions are applied to the full analysis region, and the analysis region is clipped for post-processing to reduce any artifacts near the boundaries due to non-periodic boundary conditions" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0000982_ifuzzy46984.2019.9066231-Figure2-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0000982_ifuzzy46984.2019.9066231-Figure2-1.png", + "caption": "Fig. 2. Variables in longitudinal model.", + "texts": [ + " 1 shows a photo of the flying-wing type UAV (the so called UEC-UAV I (unique, exciting, challenging unmanned aerial vehicle I) ) that will be used in our experiments. A main feature of the flying-wing type UAV is to have no horizontal and vertical stabilizers (tails) due to reducing the air resistance as much as possible, i.e., to enhancing its flight efficient as much as possible. Hence, the flying-wing type UAV realizes an extremely efficient flight in comparison with multirotor/multicopter type UAVs. In other words, flight control of the flying-wing type UAV is an extremely difficult and challenging nonlinear control problem. Fig. 2 shows variables in longitudinal model construction. Table I presents the list of the variables and parameters. As mentioned above, due to having no horizontal and vertical stabilizers (tails), the altitude and direction control are realized by manipulating the so-called elevon (mixing the elevator and aileron) on the main wing . From the six-degree-of-freedom equations of motion, we construct the following longitudinal dynamics by assuming that all the lateral variables are zero, i.e., (v(t) ,p(t) ,ret) , rLL, r < HL} is defined as: r = \u221a \u03c62 + p2 (24) and the lower and high limit ellipse radii are defined as (i \u2208 {LL,HL}): ri = \u03c6ipi\u221a pi2cos2 \u03b8 + \u03c6i 2sin2 \u03b8 (25) with angle \u03b8 measured from \u03c6 axis as: \u03b8 = arctan( p \u03c6 ) (26) A sample stability augmentation factor f as given by above equations is shown in Figure 2", + " With this fusion scheme the manual control is blended seamlessly with the autonomous control and the autonomous control ensures that the system remains in the stable flight regime. Similarly, a roll rate damper may be fused with the manual pilot control inputs for |\u03c6| < |\u03c6LL| if the inherent roll damping of the UAV is not sufficient for easier and smoother manual control. The stability of the fused control scheme given by equation (27) can be shown by observing the fact that for any initial conditions inside the roll angle control region (as shown in Figure 1) the autonomous control is fully applied for \u03c6 > \u03c6HL, i.e. f = 1. Since the autonomous control is stable, as shown in Section II-B, the system remains within one boundary layer of the transition region for commanded roll angle of \u03c6c = \u03c6LL if the manual input were zero. In the worst case of maximum adverse manual input, the system is stable around \u03c6HL since at \u03c6 = \u03c6HL the manual input is completely cut and system is controlled autonomously. The presented control fusion scheme was tested on the KSU testbed UAV" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0000628_mim.2020.8979522-Figure1-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0000628_mim.2020.8979522-Figure1-1.png", + "caption": "Fig. 1. Examples of soft robots developed by the authors.", + "texts": [ + " Chee Pin Tan and Dr. Surya Nurzaman, have expertise in state estimation and soft robotics, respectively. Together with their graduate research students, they share with us how state estimation can be used for indirect sensing in soft robots, thus contributing a solution to this long-standing challenge of enabling sensing in soft robots. T he interest in biological systems interacting with their environments has led to the development of a new class of machines called soft robots [1]. Some examples can be seen in Fig. 1. In contrast to robots built from rigid material, soft robots are at least partially made of soft materials for increased flexibility and adaptability that leads to several advantages [1], [2]. The first one is the ability to provide a safe working environment between humans and robots; in the event of a collision, soft robots can deform and yield easily, thus minimizing damage and injuries. Furthermore, soft robots have continuously deformable structures coupled with muscle-like actuations, and thus are well-suited for gripping delicate objects without damaging them, as well as holding objects of irregular shape" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0002196_icee50131.2020.9260649-Figure7-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0002196_icee50131.2020.9260649-Figure7-1.png", + "caption": "Fig. 7. Degree of membership for \u03c6", + "texts": [], + "surrounding_texts": [ + "One of the best ways to achieve stability in nonlinear systems is the sliding mode control method. This method is extensively used for controlling the systems in the presence of disturbances and uncertainties. Consider the following nonlinear model ( ) = ( , ) + ( , ) ( ) (6) Where (\u2219) and \u210e(\u2219) are the nonlinear functions of the selected system and The general form of the sliding surface is expressed as follows [13]. 1n d s e dt \u03bb \u2212 = + (7) Where Respectively show the error of the system and sliding surface. So the sliding surface equation for the MLS could be considered as follow 1 de x x= \u2212 (8) [ ][ ]2 1 1 T s e e e\u03bb \u03bb= (9) Where and are positive constant and is desired position of the system. A suitable control law is needed to drive trajectories from outside of the sliding surface to an area close to it, within a finite time, and make the system states to operate near the sliding surface. This control law is calculated using Lyapunov\u2019s theory. So the Lyapunov candidate function is considered as follow ( )21 0 2 s t\u03bd = > (10) this condition should perform = ( ) ( ) < \u2212 | ( )| (11) Where is a positive constant in (11). The appropriate control law for this system according to the sliding mode control could be selected as eq nu u u= + (12) Where ( ) [ ]1 2 1 equ e e g x \u03bb \u03bb\u2212= + (13) ( ) ( )sgnn k u s g x = \u2212 (14) is a part of the control law that is able to move trajectories toward the sliding surface. Also, is the part of the control law that makes the states to slide on the sliding surface to reach the desired performance. Although sliding mode control is capable of rejecting disturbances and achieving stability, but there is a destructive phenomenon which is known as chattering, which affects the control accuracy. Chattering is mainly caused by the use of function. that is shown sgn( ) in (14). Using a continuous function to approximate it, is a common solution. As an example, we can use tanh( ) or sat with boundary layer, but this method is not practical enough to solve chattering and also may import extra error to the system. Another solution is the use of intelligent control techniques such as fuzzy controller, which can reduce the effect of chattering as a considerable amount [14]. In the next subsection, the combination of two methods of sliding mode control and fuzzy control will be discussed." + ] + }, + { + "image_filename": "designv11_71_0002816_thc-2010-0566-Figure18-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0002816_thc-2010-0566-Figure18-1.png", + "caption": "Fig. 18. Sign convention.", + "texts": [ + " To handle these situations with combined loading we use Mohr\u2019s circle. Most of the materials have to be analysed 3-dimensionally but since they are geometric in nature, using the symmetry we can convert them to 2 dimensions like we did for the spherical storage tank (Fig. 16) Mohr\u2019s circle is a plot between the normal stress (\u03c3) on the horizontal or x-axis and the shear stress (\u03c4 ) on the vertical or y-axis (Fig. 17). Tensile normal stresses are positive and compressive are negative, for the shear stress we follow a sign convention shown in Fig. 18 where a very small volume element inside the bulk of the material is shown. Coming back to Fig. 17, the position of the center of the circle c is the average of the two normal stresses (\u03c3x and \u03c3y and reference point a can be plotted with \u03c4 xy on the y-axis and \u03c3x on the x-axis. Now we can join points c and a to get the radius and then the circle can be drawn with a radius given by Eq. (12). From this Mohr\u2019s circle for a given situation we can find out the maximum normal stress, point b, and shear stress, point e, active in the material" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0003757_amm.658.606-Figure5-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0003757_amm.658.606-Figure5-1.png", + "caption": "Fig. 5. Laser scanning system [10]", + "texts": [ + " The purpose is to reach and characterize all desirable points (table, ground, pieces, etc.) [7]. Another system considered consists of a 6 DOF KUKA KR 6 R900 robot arm (Fig. 4). A calibration method using stereovision uses two perpendicular CCD cameras which form a virtual robot tool center point (TCP). The two cameras have macro lenses and allow high magnification of the viewed object. A black sphere it was used for calibration purposes. One pixel is mapped to 0.05mm. The vision system allows high resolution and accurate readings of sphere center [8]. (a) (b) Fig. 5. Calibration method using a stereovision system: a) Schematic overview of the stereo vision system; b) Calibration procedure [8] The virtual TCP is manually guided to desired positions and orientations where the calibrated sphere is located. The machine vision algorithm adjusts robot position with respect to the calibration sphere in the camera reference frame. Robot joint states are recorded for later use in the parameter optimization algorithm. The recorded joint states are compared for each calibration point [8]", + " This method requires a camera that is rigidly attached to the end effector and a number of reference images. The purpose is to detect the corners from a chessboard, and then determined the pose of the robot. An automatic robot calibration is performed. The first step is automatic chessboard detection. The second step is calibration of the camera parameters. The final step is robot pose measurement [9]. A system with multiple 3D scanners consists of a portable laser scanner, an industrial robot and a turntable is another approach [10], Fig. 5. Information in 3D about the object surface placed on turntable can be obtained from multiple angles and directions. The calibration procedure is performed in three steps [10]: - camera and projector calibration; - robot TCP calibration; - turntable calibration. For the calibration of robot TCP, the robot is moving and the laser beam shoot at the sphere. This sphere is fixed in the position where the robot can easily reach. The center position of the sphere is obtained. A translation of Tool0 is the next step along y-axis for obtaining the next position of the center of the sphere" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0003569_s0960129512000527-Figure2-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0003569_s0960129512000527-Figure2-1.png", + "caption": "Fig. 2. The actual networked DC motor system setup", + "texts": [ + " Note that the CCL approach is quite standard, and it is easy to adapt the way we used it in Li et al. (2009) to solve the controller design problem in the current paper, but to save space and avoid repetition, we will omit the details of the CCL-based controller design procedure here. In this section, we will illustrate the applicability and effectiveness of the proposed approaches. To this end, we constructed a real-world networked DC motor system and subjected it to a comprehensive study to develop both simulation and experimental results. The system setup is shown in Figure 2, where the experimental apparatus consists of a PC controller, a local board and a DC motor with sensors. The PC controller is used to implement the networked controller. The local board is on the plant side and used for two functions: (1) to convert the control signal read from the buffer into a pulse width-modulation (PWM) signal, and then send the PWM signal to drive the DC motor; (2) to encapsulate the plant state and its timestamp into a packet and send it to the PC controller via the network" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0001821_j.matpr.2020.08.097-Figure2-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0001821_j.matpr.2020.08.097-Figure2-1.png", + "caption": "Fig. 2. The interaction of the roller, spindle and tube. Design scheme.", + "texts": [ + " The obtained relations are valid under the following assumptions: Loading the body is simple Tube material is incompressible hardening of the material when creating a load does not occur Such a model cannot be taken as the basis for studying the stress\u2013strain state of the tube and tubesheet during roller rolling, since the rolling process should be considered taking into account the diverse interaction of the contacting surfaces at each loading point in conjunction with the contact problem [16]. Fig. 1 shows the tube mount and the rolling scheme, and Fig. 2 shows the design scheme loading. As can be seen from them, such loading is complex. The tight connection of the pipe with the surface of the hole of the tubesheet is ensured by the spindle torque and indentation of the roller in a certain vicinity of the contact point of the tube and roller. In the zone of action of the roller, the metal of the tube becomes plastic when part of the layers is shifted in the direction of motion of the surface of the roller, and part is stretched and compressed, pressing against the wall of the hole", + " In the outer layers of the tube wall during rotation of the tool body, where in the grooves the rollers are placed, the stress intensity reaches a maximum near the contact point from which deformation begins, approximately in the middle of the arc between the locations of the rollers. On the inner surface of the tube in almost Fig. 4. Microhardness of tube outer surface: (a) before rolling; (b) after rolling. all modes, the intensity of the stress is equal to or exceeds the yield strength of the material. The tube deformation at the point of application of the radial force Frci (Fig. 2) reaches a maximum at an angle of rotation of the case of the tool ap = 20. . .25 . This means that for a given torque in each case in the interval 10 ap 35 , the tube will be pressed against the wall of the hole, and there is no contact outside the specified interval. From the literature [5,16,18] it follows that the value of the torque on the drive shaft, which ensures, when the spindle rotation is stopped, the necessary radial normal stress rk in the zone of contact between the tube surfaces and the hole of the tubesheet, should be equal M \u00bc k1rkls \u00f012\u00de where k1 is the coefficient characterizing the rolling and tool parameters; l is spindle contact length; s is the tube wall thickness" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0000306_icems.2019.8922343-Figure2-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0000306_icems.2019.8922343-Figure2-1.png", + "caption": "Fig. 2. LPTN modelling of active-winding.", + "texts": [ + " With respect to the convection phenomenon within the end-cap, the convection coefficients of the frame \u210e , the end-winding \u210e and the rotor \u210e are calculated by [5], [6]. Since the copper loss is one of the majorities of machine losses and the slot insulation is sensitive to over-temperature, the modelling of the active-winding region is particularly critical. Practically, the modelling of individual wires is unnecessary. In this paper, the active-winding region, including copper, impregnation and wire insulation, is built by the composite 3-dimensional hollow cylinder segment without any shape modification, as shown in Fig. 2 (a). The LP active-winding thermal network takes account into the heat flow path in axial, radial and circumferential directions, as shown in Fig. 2 (b). The equivalent thermal conductivity in radial and circumferential directions / , and in axial direction are calculated based on the adaptive H+S method in [19]. In Fig. 2 (b), the axial and radial thermal resistances are determined by the method in [5]. The circumferential thermal resistances are considered in this paper to describe the heat flow path between the stator teeth and the active-winding, the circumferential thermal resistances are expressed in (10): = ( \u2212 )2 \u2219 +\u2212 (10) , \u2212 5( \u2212 )24 \u2219 +\u2212 where is the circumferential thermal conductivity, the negative , is the \u201ccompensation resistance\u201d to set the midpoint temperature to be equal to the mean value in the circumferential direction" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0003851_e2014-02130-2-Figure1-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0003851_e2014-02130-2-Figure1-1.png", + "caption": "Fig. 1. Model of pendula coupled by a beam.", + "texts": [ + " Olssen has considered the dynamics of the spherical pendulum [18,19]. For the small pendulum\u2019s motion the derived equation of motion has been solved analytically using Lindstedt-Poincare method. In [20] different methods of solving Hamiltonian systems are presented. In the case of spherical pendulum it is advised to use Lagrange multipliers method instead of penalty method. Priest and Poth [21] have studied the dynamics of two spherical pendula mounted to the rigid beam which hang from the unmovable frame as shown in Fig. 1 focusing on the small oscillations along x-axis. In the present paper we consider the dynamics of the system of Fig. 1 but do not restrict ourselves to small oscillations, i.e., we consider large spherical displacements of the pendula. In the modeling we consider Cartesian, Cartesian with two angles and three angles descriptions of the system. In the first two cases the descriptions require more variables than there are degrees of freedom so one has to use Lagrange\u2019s multipliers. The bond equation has square functions and it is impossible to obtain one spatial configuration of the system. Only the third case leads to the unequivocal description of the system", + " The existence of synchronous states has been confirmed in the simple experiment. The paper is organized as follows. The considered model of the coupled spherical pendula is introduced in Sect. 2. Section 3 presents the results of numerical simulations and their experimental confirmation. Finally, we conclude our studies in Sect. 4. a e-mail: witkowskiblazej@gmail.com 2 General model for two spherical pendula coupled by the beam The considered system is composed of a beam and two spherical pendulums as presented in Fig. 1. The beam of mass M is attached to the ends of a weightless and inextensible strings, each of them has length l. At ends of the beam two identical mathematical pendula are suspended, each of length h and mass m. Model of following system is not straight forward to construct due its complexity. First we introduce separate models of the beam and the pendula. Them we construct the full model of the beam \u2013 pendula systems and derive formulas for its kinetic and potential energies. Using these expression one may obtain the second order ODEs describing the motion of the system, from Lagrange equations of the second kind", + " To obtain excitation for each variable let us consider virtual work introduced for each force: \u2202L = F1x\u2202(xD \u2212 xA) + F1y\u2202(yD \u2212 yA) + F2x\u2202(xE \u2212 xB) + F2y\u2202(yE \u2212 yB) = F1h\u221a (sin\u03b11)2 + (cos\u03b11 sin\u03b21)2 (sin\u03b21\u2202\u03b11 \u2212 sin\u03b11 cos\u03b11 cos\u03b21\u2202\u03b21) + F2h\u221a (sin\u03b12)2 + (cos\u03b12 sin\u03b22)2 (sin\u03b22\u2202\u03b12 \u2212 sin\u03b12 cos\u03b12 cos\u03b22\u2202\u03b22), where \u2202(xD \u2212 xA) and \u2202(yD \u2212 yA) and \u2202(xE \u2212 xB) and \u2202(yE \u2212 yB) are virtual displacments respectively for F1 and F2. Then one obtains: T\u03b11= F1h sin\u03b21\u221a (sin\u03b11)2 + (cos\u03b11 sin\u03b21)2 , T\u03b21= \u2212F1h sin\u03b11 cos\u03b11 cos\u03b21\u221a (sin\u03b11)2 + (cos\u03b11 sin\u03b21)2 , T\u03b12= F2h sin\u03b22\u221a (sin\u03b12)2 + (cos\u03b12 sin\u03b22)2 , T\u03b22= \u2212F2h sin\u03b12 cos\u03b12 cos\u03b22\u221a (sin\u03b12)2 + (cos\u03b12 sin\u03b22)2 . We assume that beam has no external excitation: T\u03b1 = 0, T\u03b2 = 0, T\u03b3 = 0. Appendix B Coupled second order ordinary differential equations, which describe the system of Fig. 1 can by written by matrices of: mass M 7x7, torque T 7x1, damping C 7x1 and rest RR 7x1: M(q, q\u0307)q\u0308 + C(q)q\u0307 +RR = T, q = (\u03b1\u03b2, \u03b3, \u03b11, \u03b21, \u03b12, \u03b22), M11 = 1 48d8 l2 ( 4d6(4l2M \u2212 b2(M \u2212 3m)) + 16d8(M + 3m) \u2212d2b2(32l4 \u2212 55l2b2 + 32b4)(M + 3m) + 4b2(8l6 + l4b2 + l2b4 + 8b6) \u00d7(M + 3m) + 4d4(4l4(M + 3m) + 2l2b2(M + 3m) + b4(7M + 9m)) + b ( 8b(d6M \u2212 2d2l2b2(M + 3m) + 6l2b2(l2 + b2)(M + 3m) +2d4l2(4M + 9m)) cos \u03b8 + 4l2b3(M + 3m) cos 4\u03b1 cos ( \u03b8 2 )2 (4(l2 + b2) \u221211d2(4(l2 + b2)\u2212 13d2) cos \u03b8) + b(\u22124d6 \u2212 7d2l2b2 + 12l2b2(l2 + b2) + 4d4(2l2 + b2)(M + 3m)) cos 2\u03b8 + 4b cos 2\u03b1(\u22123d6(M +m) \u2212d2(8l4 \u2212 5l2b2 + 8b4)(M + 3m) + 4(2l6 + l4b2 + l2b4 + 2b6)(M + 3m) + d4(\u22124l2(M + 3m) + b2(5M + 3m))\u2212 2(d6M + 6d4l2m+ 8d2l2b2(M + 3m) \u22128l2b2(l2 + b2)(M + 3m)) cos \u03b8 + (d6 \u2212 (d4 + 5d2l2 \u2212 4l4)b2 + 4l2b4) \u00d7(M + 3m) cos 2\u03b8) + 4l(4d6M + 6d2b2(l2 + b2)(M + 3m) + 2b2(4l4 \u2212 5l2b2 + 4b4)(M + 3m) + d4(b2(M \u2212 3m) + 8l2(M + 3m)) + 2(2d6M + 4l2b4(M + 3m) + 4d2b2(l2 + b2)(M + 3m) + d4(b2M + 4l2(M + 3m))) + cos \u03b8 + b2(d4 + 2l2b2 + 2d2(l2 + b2)) \u00d7(M + 3m) cos(2\u03b8)) sin\u03b1+ 4lb2(6d2(l2 + b2)(M + 3m) + (8l4 \u2212 7l2b2 + 8b4)(M + 3m)\u2212 3d4(M + 5m) + 2(6l2b2(M + 3m) + 4d2(l2 + b2)(M + 3m)\u2212 3d4(M + 4m)) cos \u03b8 \u2212 (3d4 \u2212 3l2b2 \u2212 2d2(l2 + b2)) \u00d7 (M + 3m) cos 2\u03b8) sin 3\u03b1+ 32l3b4(M + 3m) cos ( \u03b82 )4 sin 5\u03b1 )) , M12 = l3l2 6d5 sin \u03b8 cos\u03b13 (\u22123d2M + 2l2M \u2212 6d2m+ 6l2m+ 2ll(M + 3m) sin\u03b1 \u22122(M + 3m) cos \u03b8 (d2 \u2212 l2 \u2212 ll sin\u03b1)), M13 = l2b 6d6 sin \u03b8 cos\u03b1(lb2(M + 3m) cos\u03b12(\u22123d2 + 2l2 + 2b2 + 6lb sin\u03b1 +2 cos \u03b8(\u22122d2 + l2 + b2 + 3lb sin\u03b1)) + 12 (2d4lM + 4d2l3M + d2lb2M +2l3b2M + 2lb4M + 12d2l3m+ 3d2lb2m+ 6l3b2m+ 6lb4m + lb2(M + 3m) cos 2\u03b1(2(l2 + b2)\u2212 3d2 + 2(l2 \u2212 2d2 + b2) cos \u03b8) + 2d4bM sin\u03b1 +8d2l2bM sin\u03b1+ l2b3M sin\u03b1+ 24d2l2bm sin\u03b1+ 3l2b3m sin\u03b1+ l2b3M sin 3\u03b1 +3l2b3m sin 3\u03b1+ b(M + 3m) cos \u03b8((\u22124d4 + l2b2 + 4d2(l2 + b2)) sin\u03b1 + lb(2(2d2 + l2 + b2) + lb sin[3\u03b1])))), M14 = hlm (cos\u03b1 cos\u03c61 + sin\u03b1 cos (\u03b2 \u2212 \u03b81) sin\u03c61) , M15 = \u2212hlm sin\u03b1 sin (\u03b2 \u2212 \u03b81) cos\u03c61, M16 = hlm d4 (cos\u03c6" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0001928_s00170-020-06182-0-Figure3-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0001928_s00170-020-06182-0-Figure3-1.png", + "caption": "Fig. 3 a Camera positioning on 3D printer\u2019s carriage opposite to feedstock and b white dots marked on moving filament for motion profile measurement (camera view)", + "texts": [ + " Video frame analysis enables observation of extruder motor behavior during all stages of material extrusion of a single strand and correlation of these observations with the deposition results from specimen analysis. Especially Model R2 (%) Pure error (%) LOF (p value) Curvature (p value) Normality (p value) for factors whose impact lasts for a few milliseconds, such as DP, SR, and ESR, optical tracking is an effective and low-cost way to obtain information that is not otherwise measurable or indeed conceivable. A 3D printed jig was fabricated to mount the web camera on the 3D printer carriage. The web camera is secured with screws on the jig, facing the filament (Fig. 3a).White dots aremarked on the black PLA filament, to achieve high black-white contrast (Fig. 3b). Single particle tracking method is performed via TrackMate\u2122 [12], an intuitive plugin that is executed within Fiji ImageJ\u2122 software. White dot position is tracked and registered for each frame (Fig. 4); therefore, feedstock position and velocity are calculated with respect to time. To minimize the effect of vibrations that are transmitted to the web camera during carriage motion, a reference dot that is marked on a non-moving part is also tracked and its oscillation is used as correction factor to the feedstock position tracking" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0002048_ecce44975.2020.9235328-Figure9-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0002048_ecce44975.2020.9235328-Figure9-1.png", + "caption": "Fig. 9: Flux density distribution at 15.17 A, 1800 RPM.", + "texts": [ + " 8 shows the BEMF spectrum for uniform and nonuniform magnet axial temperatures at full load operation (15.17 A 1800 RPM). For better observation, the results are normalized with respect to fundamental harmonic of BEMF, and a logarithmic scale is used in the y-axis. The result shows that all harmonic spectrums are almost the same for the uniform and non-uniform PM temperature except the 3rd (at 180 Hz) and 7th (at 420 Hz) order harmonics. The 3rd and 7th order harmonics are increased by 3.5 dB and 2.3 dB, respectively, due to the uneven axial magnetization. Fig.9 shows the magnetic flux density (B) at full load condition. The maximum flux density is 2.61 T occurs in the rotor bridge and the stator teeth below the core material's saturation point. Hence, there is no short circuit of magnets by the bridge, which is a possible source of BEMF harmonics. Therefore, it is confirmed that the change in harmonic contents is due to the PM uneven magnetization only, which is caused by the magnet non-uniform axial temperature variation. Fig. 10 and 11, show the magnitude variation of fundamental and 3rd harmonics (for both uniform and nonuniform conditions) for different loading PM temperatures" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0000115_012066-Figure3-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0000115_012066-Figure3-1.png", + "caption": "Figure 3. Example of optimization towards the topology of the lower part of the laser head", + "texts": [ + " At present due to its design and overall size, the laser head can not be removed to the general tool magazine on a CNC machine, even if the collimator is disconnected. To store the IPDME 2019 IOP Conf. Series: Earth and Environmental Science 378 (2019) 012066 IOP Publishing doi:10.1088/1755-1315/378/1/012066 laser head, it must be removed manually, close the lens and the hole for the collimator with protective covers to prevent industrial dust and cutting fluid. Then, it must be removed to a special for it place. Fig. 3 and 4 show the examples on how the topology of the lower part of the laser head can be optimized and the collimator fastening and the back wall can be modernized. IPDME 2019 IOP Conf. Series: Earth and Environmental Science 378 (2019) 012066 IOP Publishing doi:10.1088/1755-1315/378/1/012066 Dividing the entire laser system into two main components (control unit and laser head), you can not only integrate it into most machining centers, but also increase its maintainability. In order to replace some element of the system, it will not need to be completely disassembled" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0002504_iceict51264.2020.9334200-Figure2-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0002504_iceict51264.2020.9334200-Figure2-1.png", + "caption": "Fig. 2. Bearing test rig.", + "texts": [ + " In the second step, the DFA method is employed to process the enhanced vibration signal, and the feature vector [H\u03b1, C] can be obtained. In the last, fault types can be identified by using the improved IKNN classifier to classify the feature vectors. 125 Authorized licensed use limited to: East Carolina University. Downloaded on June 16,2021 at 10:42:42 UTC from IEEE Xplore. Restrictions apply. IV. EXPERIMENTAL ANALYSIS The experimental data of bearing fault is obtained from the bearing test rig of the Xi\u2019an Jiaotong University [12]. The accelerated life test was conducted on a specially designed test bench (see Fig.2) and the full-life-cycle vibration data were collected. In Fig.2, the test rig consists of speed controller, AC motor, hydraulic loading system, and loading and bearing housing. The working conditions of the test rig mainly include rotational speed and loading force, which can be adjusted by the speed controller and the hydraulic loading system, respectively. The rolling bearings LDK UER204 were tested. During the testing, a constant load of 11 kN was provided by the hydraulic loading system, and a rotation speed of 2250 rpm was set by the speed controller. Three different bearing fault types were generated during the accelerated life test: inner race fault, cage fault, and outer race fault" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0001770_s00170-020-05701-3-Figure12-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0001770_s00170-020-05701-3-Figure12-1.png", + "caption": "Fig. 12 Representative standard part of a part family produced in a manufacturing cell", + "texts": [ + " 11 is obtained compiling the PPM editor program, which has similar characteristics to other modules proposed in the literature [20, 21]. One of the parameterized programming characteristics is the possibility of using algorithmic logic, such as, for example, relational and conditional conditions. Thus, it is possible for a parameterized program to contain several procedures or cycles in subroutines form and execute the cycles as it is triggered. This is especially interesting due to the commercial CNC module structure that is closed, which does not provide an approach to extend its functionality [22]. Within this context, Fig. 12 illustrates the longitudinal, transverse, and angle thinning turning operations of a standard part representative of a given part family produced in a manufacturing cell. A parameterized program can be developed to perform all operations or only those that are requested. Figure 13 illustrates the flowchart of the implemented algorithm to perform this operation. In order to demonstrate the methodology for programming the standard part of part family, a parameterized program was developed. Considering that developed routines\u2019 focus is on the tool positioning, tool compensation and finishing cycles were not added to the modules" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0002434_cdc42340.2020.9304075-Figure2-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0002434_cdc42340.2020.9304075-Figure2-1.png", + "caption": "Fig. 2. Unicycle mobile robot with handle.", + "texts": [ + " To illustrate the above result, we will apply it to a simple kinematic model of a unicycle, that can be seen as a simplified model of mobile robots subject to non-holonomic constraints. Here we make the simplifying assumption that the unicycle will not tilt away from the vertical. A state space model is then q\u0307d x = vcos\u03c6 \u2212\u03c9 d sin\u03c6 q\u0307d y = vsin\u03c6 +\u03c9 d cos\u03c6 \u03c6\u0307 = \u03c9 (12) where qd x (t) and qd y (t) denote the position of a point in front of the unicycle at a distanced d (a \u201chandle\u201d, see [26]), \u03c9 and v denote the angular and linear velocity, respectively, and the angle \u03c6 represents the orientation, see Fig. 2. In order to construct an LPV model of this system, it is convenient to first apply a dynamic coordinate transformation T (t) = cos\u03c6(t) sin\u03c6(t) 0 \u2212sin\u03c6(t) cos\u03c6(t) 0 0 0 1 , (13) which leads to the model \u02d9\u0303qd x = v+\u03c9 q\u0303d y \u02d9\u0303qd y =\u2212\u03c9 q\u0303d x +\u03c9d \u03c6\u0307 = \u03c9 (14) or \u02d9\u0303qd x \u02d9\u0303qd y \u03c6\u0307 = 0 \u03c9 0 \u2212\u03c9 0 0 0 0 0 q\u0303d x q\u0303d y \u03c6 + 1 0 0 d 0 1 [ v \u03c9 ] , (15) 100 Authorized licensed use limited to: Carleton University. Downloaded on June 04,2021 at 19:14:20 UTC from IEEE Xplore. Restrictions apply. F l Jl AlX +BuC\u0302l K Al +BuD\u0302l KCv Bl w +BuD\u0302l KDvw 0 0 0 I Y T \u2217 H l A\u0302l K YAl + B\u0302l KCv Y Bl w + B\u0302l KDvw 0 0 0 0 V T \u2217 \u2217 X +XT \u2212F j I +ST \u2212 J j 0 XTCl z T +C\u0302l K T Dzu T XT UT 0 0 \u2217 \u2217 \u2217 Y +Y T \u2212H j 0 Cl z T +Cv T D\u0302l K T Dzu T I 0 0 0 \u2217 \u2217 \u2217 \u2217 I Dl zw T +Dvw T D\u0302l K T Dzu T 0 0 0 0 \u2217 \u2217 \u2217 \u2217 \u2217 \u03b32I 0 0 0 0 \u2217 \u2217 \u2217 \u2217 \u2217 \u2217 \u03c3xI 0 0 0 \u2217 \u2217 \u2217 \u2217 \u2217 \u2217 \u2217 \u03c3xI 0 0 \u2217 \u2217 \u2217 \u2217 \u2217 \u2217 \u2217 \u2217 I 0 \u2217 \u2217 \u2217 \u2217 \u2217 \u2217 \u2217 \u2217 \u2217 I > 0 (11) y = [ 1 0 0 0 1 0 ] q\u0303d x q\u0303d y \u03c6 , which can be written as x\u0307(t) = A\u0303 ( \u03c9(t) ) x(t)+ B\u0303u(t), y(t) =Cyx(t) (16) We apply Euler discretization to the state equation (15) with sampling time Ts = 0" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0002111_s00170-020-06308-4-Figure1-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0002111_s00170-020-06308-4-Figure1-1.png", + "caption": "Fig. 1 Grain growth and morphology of an alloy in an AM process for a single track, b multiple tracks in a single layer, and c multiple tracks in", + "texts": [ + " During PBF and DED builds, a part is subjected to complex thermal changes due to rapid and repeated heating and cooling cycles [3]. The thermal history for each voxel influences the microstructure and mechanical properties of the part. With the current technologies, each layer may be programmed to incur different processing parameters or a complex geometry may encourage a thermal history difference between layers. High cooling rates and low thermal gradients usually result in fine equiaxed grains and low cooling rates with high thermal gradients form large columnar grains [4]. As shown in Fig. 1, the grain morphology resulting from a DED process is columnar with equiaxed grain structure around past layer bands; however, the addition of more layers introduces enough heat to partially re-melt former layers, allowing them to adopt the columnar shape with epitaxial growth [5, 6]. This type of grain growth can also be seen in PBF methods despite the difference in processing conditions between the two methods during a build [7]. The percentage and position of the equiaxed grains within the columnar structure are influenced by the parameters of the process like energy density and scan pattern, which determine the thermal history [4]", + " The orientation of the grain growth and the grain size are well displayed by EBSD images, while the OM images more clearly display the columnar growth of grains through layers. The EBSD maps in Fig. 8 represent inverse pole figure (IPF) maps of the face-centered-cubic (FCC) iron (gamma) or austenite phase where the build direction is used as the reference, and each pixel is displayed with a color value representing the crystal direction oriented along that reference build direction. It was observed that the sample microstructure was entirely austenitic iron (gamma). The microstructure in Fig. 9, like the DED phenomena depicted in Fig. 1, shows the columnar growth through multiple layer bands is present in SLM as well. Columnar grains grow through multiple layers due to the directional thermal gradient and epitaxial growth. There are various regions where the grains are finer and more equiaxed, which indicates that the heat from the melting of subsequent layers did not promote local columnar epitaxial growth. Furthermore, the software used for EBSD analysis creates a list of all grains identified and sorts them by area, allowing the median size to be easily identified for each mapped area" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0003209_chicc.2015.7260500-Figure1-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0003209_chicc.2015.7260500-Figure1-1.png", + "caption": "Fig. 1 Flying principle of the quadrotor UAV", + "texts": [ + " At the same time use the observer to estimate the disturbance, and compensate the disturbance in the control law, in order to improve the anti-interference ability of the aircraft. The control law and the observer are described in detail in this paper. *This work is supported by National Natural Science Foundation (NNSF) of China under Grant 61374032. 2 Quadrotor UAV model The movement of Quadrotor UAV can be seen as six degree of freedom rigid body motion, the motion includes three axes of rotation and three line movement. System diagram as shown in Figure 1 Assume the aircraft as a rigid body and ignored its elastic deformation, because the symmetries of the aircraft structure the quality of aircraft has a uniform distribution assume aircraft lift surface and center of gravity in the same plane, the mathematical model of the system is as follows. 3 ( ) m mge RF W J J 1 In the formula 3, R representing the position and speed of the aircraft under the ground coordinates. T 3 (0,0,1)e is the unit vector under ground coordinates, F is the force vector except gravity acting on the aircraft " + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0001471_j.matpr.2020.06.306-Figure2-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0001471_j.matpr.2020.06.306-Figure2-1.png", + "caption": "Fig. 2. Cross sectional view of a Bearing.", + "texts": [ + " The entire dataset includes SET A 3 datasets having no fault condition, SET B 3 datasets having outer race fault with constant load, SET C 7 datasets having outer race faults with various loads, SET D 7 datasets having inner race faults with various loads and SET E 3 industrial datasets are collected from an oil pump bearing, an intermediate speed bearing, and a planet bearing. Each dataset contains an acceleration signal collected from accelerometer, sampling rate (samples per second), shaft speed (Hz), load weight (lbs), and four critical frequencies (Hz) representing various fault locations: 1) ballpass frequency outer race (BPFO), 2) ballpass frequency inner race (BPFI), 3) fundamental train frequency (FTF), and 4) ball spin frequency (BSF). The cross sectional view of a bearing is shown in Fig. 2. The formulae to calculate critical frequencies are Ballpassfrequency; outerraceBPFO \u00bc n f r 2 \u00f01 d D cosh\u00de \u00f01\u00de Ballpassfrequency; innerraceBPFI \u00bc n f r 2 \u00f01\u00fe d D cosh\u00de \u00f02\u00de FundamentaltrainfrequencyFTF \u00bc f r 2 \u00f01 d D cosh\u00de \u00f03\u00de Ball\u00f0roller\u00despinfrequencyBSF \u00bc D 2d 1 d D cosh 2 \" # \u00f04\u00de Please cite this article as: V. R. Ravi, S. Aarthi, M. Aishwarya et al., Remaining s doi.org/10.1016/j.matpr.2020.06.306 The bearing parameters which are used in dataset are listed in Table 1. The further details of bearing dataset are listed below: SET A \u2013 comprising 3 datasets with no fault conditions" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0003921_coase.2014.6899477-Figure2-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0003921_coase.2014.6899477-Figure2-1.png", + "caption": "Figure 2. Robot equipped with the safety devices.", + "texts": [ + " In section III, we explain the structure and mechanism of the new safety device. In section IV, we show a wheeled robot equipped with the new safety device. In section V, we present experimental results to verify the effectiveness of the new safety device. Section VI concludes this paper. The velocity and contact force-based mechanical safety device consists of a contact force-based detection mechanism, a velocity-based detection mechanism, and a switch-off and shaft-lock mechanism, as shown in Fig.2. The characteristics of the safety device are as follows: (i) If the angular velocity of a drive-shaft (hereinafter referred to as \u201cshaft\u201d) exceeds a preset threshold level in a direction, then the safety device for the velocity's direction of the shaft switches off all motors of the robot and locks the shaft in the direction. We call the preset threshold level the \u201cdetection velocity level\u201d. (ii) The detection velocity level is adjustable. (iii) If the contact force acting on a contact force-based detection mechanism exceeds a preset threshold level, then the safety device for the contact force-based detection mechanism switches off all motors of the robot and locks the shaft in a direction opposite to the direction of the contact force" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0002189_icem49940.2020.9271053-Figure3-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0002189_icem49940.2020.9271053-Figure3-1.png", + "caption": "Fig. 3. IPMSM for EV traction.", + "texts": [ + " Hence, the actual current density of element i in the conductor can be derived as, ( ) ( ) ( )i 0 i e,i i 1 / n t t S t = = + J I J (8) Finally, the AC copper loss and the AC resistance can be calculated as follows, respectively, ( )2 AC copper e,i i i 1 1n avg P J t S L \u03c3= = \u22c5 \u22c5 (9) AC copper AC 2 input, rms P R I = (10) where, L is the axial length of conductor and Iinput, rms is the rms value of current input. An interior permanent magnet synchronous motor (IPMSM) with hairpin conductors was used to compare the results of proposed method with those of transient analysis. The IPMSM for EV traction is shown in Fig. 3 and the specifications are listed in Table I. The magnetic vector potential, the distribution of current density in conductors and the AC copper loss were compared in this section. When conducting the proposed method, the in-house code is used TABLE I SPECIFICATIONS OF IPMSM FOR EV TRACTION Contents Unit Value DC link voltage VDC 350 Line current limit Arms 380 Output power kW 115 Maximum torque Nm 287 Maximum speed RPM 12000 Authorized licensed use limited to: University of Gothenburg. Downloaded on December 19,2020 at 06:21:16 UTC from IEEE Xplore" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0003008_978-3-319-76138-1_1-Figure1.5-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0003008_978-3-319-76138-1_1-Figure1.5-1.png", + "caption": "Fig. 1.5 Comparison of the critical load case in the design of a robot. Serial robots have to be designed for bending, Stewart\u2013Gough-platforms are designed to withstand buckling of the legs, and cables are subject to pure tension", + "texts": [ + " Moreover, there are different designs where the actuators can be fixed on the machine frame and thus, only the passive machine structure must be lifted, balanced, and accelerated. The advantage of this design is twofold: One can use smaller actuators and lighter links. The weight of the robot\u2019s structure depends on the load case that has to be used for dimensioning. Serial robots are mostly driven by the torque in their joints and therefore the armsof serial robots are subject to bending (left in Fig. 1.5). Towithstand bending, large cross sections have to be used leading to heavy machine parts. This limits, in turn, the dynamic performance as well as the maximum payload. The structure of parallel robots like Stewart\u2013Gough-platforms orDelta robots can ease the load case used to dimension their parts. The critical load case changes from bending to buckling, which allows to reduce the weight of the robot\u2019s legs compared to serial robots. Delta robots impressively demonstrate the effectiveness of this approach and achieve very high accelerations [103]" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0003451_s00707-015-1403-6-Figure6-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0003451_s00707-015-1403-6-Figure6-1.png", + "caption": "Fig. 6 Motion of the NNC model with parallel velocities in a moving system x\u0304 O\u0304 y\u0304", + "texts": [ + " Finally, if x\u0307 = 0 (y\u0307 = 0) and \u03d5\u0307 = 0, the system is at rest; therefore, it will be that E = T = 0 \u2261 0 = const. A question can be posed: Can the absence of the reaction of a nonholonomic constraint (\u03bb = 0) be also realizedwhen themotion of the consideredmodel of the system is observedwith respect to amoving coordinate system? Further analysis will provide a positive answer. Let the considered model move within the system O\u0304 x\u0304 y\u0304 that, relative to the stationary system Oxy, moves according to the known (to start with, arbitrary) law xO\u0304 = B(t), Fig. 6. The equation of the nonholonomic constraint relative to the system Oxy reads F = ( x\u0307 \u2212 B\u0307 ) cos\u03d5 \u2212 y\u0307 sin \u03d5 = 0, (5.31) and the differential equations of motion are 2mx\u0308 = \u03bb cos\u03d5, 2my\u0308 = \u2212\u03bb sin \u03d5, Jz \u03d5\u0308 = 0, (5.32) Differential equations of constraint (5.31) with respect to time are obtained by dF dt = ( x\u0308 \u2212 B\u0308 ) cos\u03d5 \u2212 ( x\u0307 \u2212 B\u0307 ) \u03d5\u0307 sin \u03d5 \u2212 y\u0308 sin \u03d5 \u2212 y\u0307\u03d5\u0307 cos\u03d5 = 0, and substituting x\u0308 and y\u0308 from differential equations of motion (5.32), it follows that (where we have assumed that m = 1 2 ) dF dt = ( \u03bb cos\u03d5 \u2212 B\u0308 ) cos\u03d5 \u2212 ( x\u0307 \u2212 B\u0307 ) \u03d5\u0307 sin \u03d5 \u2212 (\u2212\u03bb sin \u03d5) sin \u03d5 \u2212 y\u0307\u03d5\u0307 cos\u03d5 = 0 wherefrom it is obtained \u03bb = ( x\u0307 \u2212 B\u0307 ) \u03d5\u0307 sin \u03d5 + y\u0307\u03d5\u0307 cos\u03d5 + B\u0308 cos\u03d5 = [( x\u0307 \u2212 B\u0307 ) sin \u03d5 + y\u0307 cos\u03d5 ] \u03d5\u0307 + B\u0308 cos\u03d5 = ( \u02d9\u0304x sin \u03d5 + \u02d9\u0304y cos\u03d5 ) \u03d5\u0307 + B\u0308 cos\u03d5" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0000472_s10556-020-00700-3-Figure1-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0000472_s10556-020-00700-3-Figure1-1.png", + "caption": "Fig. 1. Computational diagram: (a) \u2014 model of plant of cementation unit; (b) \u2014 pipeline of exhaust system.", + "texts": [ + " An analysis of the process in which a measuring container is heated as these vehicles travel with filled measuring containers was not considered in light of conditions that limit loads on highway pavement. The problem of heating the SIN-32 high-pressure pump and the ODN water-supply pump incorporated into the TsA-320 cementation assembly (produced by several Russian firms) and mounted on the KAMAZ 43118 chassis was selected for numerical investigation. A full-scale configuration plant model was developed in a computer-aided design environment (Fig. 1, a), comprising the KAMAZ 43118-46 chassis, SIN-32-00.400 pump with screen near the crankcase, ODN-120-100-63 pump, and a branch pipe for discharge of spent gases from the 4326-12000002 exhaust system without guiding units. Grooved pipe was produced to deliver hot spent gases to the heating units in the pipeline (cf. Fig. 1, b). The problem was investigated by a numerical method. A nonstationary nonlinear solver of gaseodynamic processes (Siemens PLM Software. STERCCM + Doc- umentation. Version 12, pp. 7038\u20137039, 7041\u20137044) was used as the instrument of the numerical method. Traveling speed of the station at the site \u2014 40 km/h; speed of oncoming flow of air \u2014 10 m/s; ambient (air) temperature \u2014 +5\u00b0C; midlength-section plate in the air flow used as the screening conditions of the chamber; the volume of the design air space was selected to be sufficiently high to exclude the influence of boundary conditions on the streamline parameters of the pumps" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0003208_9781118751992.ch4-Figure4.10-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0003208_9781118751992.ch4-Figure4.10-1.png", + "caption": "Figure 4.10 Schematic diagram showing the switching process in SSFLC.", + "texts": [ + " When the strength of the applied field is 1 V/\u03bcm = 106 V/m and P ! s is parallel to E ! , the electric energy density is \u2212P ! E! = 103 J=m3, which is much higher than the electric energy density of the dielectric interaction of nonferroelectric liquid crystals with electric field. This is one of the reasons for fast switching speed of ferroelectric liquid crystal devices. Now we consider the dynamics of the switching of SSFLC. We only consider the rotation around the cone (Goldstone mode), as shown in Figure 4.10. The electric torque is \u0393 ! e =P ! s \u00d7 E ! =PsE sin\u03d5x\u0302: \u00f04:47\u00de The viscosity torque is \u0393 ! v = \u2212\u03b3n ! \u22a5 \u00d7 \u0394 n ! \u0394t = \u2212\u03b3 sin\u03b8 sin\u03b8\u0394\u03d5 \u0394t x\u0302= \u2212\u03b3 sin2 \u03b8 \u2202\u03d5 \u2202t x\u0302, \u00f04:48\u00de where \u03b3 is the rotational viscosity coefficient. These two torques balance each other and the dynamic equation is PsE sin\u03d5\u2212\u03b3 sin 2 \u03b8 \u2202\u03d5 \u2202t = 0: \u00f04:49\u00de The solution is \u03d5 t\u00f0 \u00de= 2arctan tan \u03d5o 2 et=\u00f0\u03b3sin 2\u03b8=PsE : \u00f04:50\u00de The response time is \u03c4 = \u03b3 sin2 \u03b8 PsE : \u00f04:51\u00de The rotation around the (small) cone is another reason for the fast switching speed" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0001073_s12083-019-00858-5-Figure2-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0001073_s12083-019-00858-5-Figure2-1.png", + "caption": "Fig. 2 four level inverted pendulum mapped to humanoid robot", + "texts": [ + " In the falling process of a humanoid robot, in order to reduce the ground impact force, adding the cushioning material in the back, the buttock (back to the ground) or the chest and the abdomen (forward to the ground) can absorb the impact force. However, this passive method doesn\u2019t rely on the robot\u2019s own posture to provide protection, it\u2019s quite difficult to fundamentally solve the damage problem of ground collision. Optimizing robot\u2019s control parameters when touching ground is a good method to realize the self protection of the robot, so that the transition of the humanoid robot action is realized naturally and smoothly, and the ground impact, the landing location and the stability are chosen optimal. Figure 2 demonstrates a stable standing model of a four level link rod for a humanoid robot. mi , Ii , Li (1 <= i <= 4) are the mass, moment of inertia and length of the link rod respectively. Oi (1 <= i <= 4) is the rotating joint of the ith link rod, Ci (1 <= i <= 4) is the barycenter position of the link rod, li (1 <= i <= 4) is the distance between joint and barycenter of the link rod, \u03b8i(1 <= i <= 4) is the angle between the link rod and the vertical line of the rotating joint of the link rod. When the robot\u2019s joint motor is power down, the standing model of the connecting rod can be simplified as a quadruple inverted pendulum, the pendulum l1 is the connecting rod of the robotic calf, the l2 is the connecting rod of the robot thigh, and the l3 is the robotic torso The connecting rod l4 is the connecting rod of the robot arm" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0001856_elektro49696.2020.9130311-Figure7-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0001856_elektro49696.2020.9130311-Figure7-1.png", + "caption": "Fig. 7: back-EMF State filter.", + "texts": [], + "surrounding_texts": [ + "\u02c6 Estimated quantities * Commanded quantities s Stator reference frame r rotor reference frame" + ] + }, + { + "image_filename": "designv11_71_0001584_icced46541.2019.9161142-Figure2-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0001584_icced46541.2019.9161142-Figure2-1.png", + "caption": "Fig 2. Electronic Component Diagram", + "texts": [ + "11b/g/n \u2013 6 x Digital I/O, 3 x PWM Channels, 1 x ADC Channel \u2013 Full I/O control through WiFi network \u2013 GPIO with 15mA current drive capability \u2013 Supports Smart Link intelligent networking \u2013 Built in 32-bit MCU \u2013 Built-in TCP/IP protocol stack, and support multiple TCP Client connection \u2013 UART/GPIO data communication interface III. SYSTEM OVERVIEW At the design stage of minitoring system and rice mass controling in form of rice dispenser, divided into two parts of design that is hardware design part and software design part. In general the system is presented in Figure 1. The hardware rice dispenser design is divided into 2 parts: electronic component parts and rice storage parts. The electronic interface design for each component of the rice mass monitoring system is presented in Figure 2. Authorized licensed use limited to: University of Gothenburg. Downloaded on August 24,2020 at 02:28:44 UTC from IEEE Xplore. Restrictions apply. Power supply NodeMCU ESP8266 uses a 5VDC battery. All components are connected to the NodeMCU ESP8266 pins with the power supply on each component derived from the 3.3VDC output voltage NodeMCU ESP8266 which is arranged in parallel. The four half bridge load cells are assembled into a wheatstone bridge then converted from analog data to digital data on the HX711 module then in digital data it is converted into mass rice at nodeMCU" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0000662_icus48101.2019.8995953-Figure5-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0000662_icus48101.2019.8995953-Figure5-1.png", + "caption": "FIGURE 5. Antagonistically actuated VSA system with a reaction wheel.", + "texts": [ + " We employed two additional testbeds to benchmark the performance of our system: a single-link VSA only system and a single-link system actuated by a reaction wheel only. The dynamics of the latter system is identical to the combined system except the elastic torque \u03c4E,1 = 0. For the single-link VSA system, the 1\u00d7 1 dynamic matrices are M = [ml1(L l c,1) 2 + mbL2c,b + J l 1 + Jb], (20a) C = [0], (20b) B = [bl1], (20c) N = [g(ml1L l c,1 + mbLc,b) cos(\u03c81)]. (20d) B. EXPERIMENTAL SETUP The robot described in the previous section was designed in Solidworks and built as an experimenteal testbed (see Fig. 5). Most of the parts were 3D-printed using UP Plus2 rapid-prototyper. The joint shafts and the link were CNC-machined out of steel and textolite, respectively. Reaction wheel was realized using a bicycle brake disk and connected to the link via low-friction ball bearings. Due to their high torque and speed, two Dynamixel MX-28 servomotors were used to actuate the joint 1 through NEEs. VOLUME 4, 2016 4623 These servos were interconnected in a chain fashion and communicated with the computer via a USB2Dynamixel module" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0002216_icem49940.2020.9270977-Figure5-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0002216_icem49940.2020.9270977-Figure5-1.png", + "caption": "Fig. 5: Space vector diagram in the dxqx reference frame considering to apply the stator current along the d-axis.", + "texts": [ + " The rotor current can be expressed, from the second of (4): i x r = \u03bb x r Lr \u2212 Lm Lr i x s (7) then, substituting (7) in the first of (4): \u03bb x s = i x s ( Ls \u2212 L2 m Lr ) + Lm Lr \u03bb x r (8) and knowing the expression of the transient inductance (1), the stator flux linkage space vector is: \u03bb x s = i x s Lt + Lm Lr \u03bb x r (9) Further, knowing that Lm/Lr is the transformation ratio from rotor to stator quantity in the RFO reference frame, the rotor flux, referred to the stator can be indirectly measured using: \u03bb x rsd = \u03bb x sd \u2212 i x sdLt \u03bb x rsq = \u03bb x sq (10) Considering the rotor voltage space vector equation in the generic synchronous reference frame and neglecting the time derivative of the flux: u x r = Rri x r + j\u03c9sl\u03bb x r = 0 (11) the rotor current occurs only if \u03c9sl = \u03c9e \u2212 \u03c9me 6= 0. 227 Authorized licensed use limited to: UNIVERSITY OF WESTERN ONTARIO. Downloaded on May 25,2021 at 13:25:58 UTC from IEEE Xplore. Restrictions apply. Figure 5 shows an example of space vector diagram, considering the generic x reference frame adopted for measurement and the RFO reference frame indicated with \u03bb. Once the rotor flux space vector is computed, the angle \u03b1is\u03bbr between the rotor flux and the stator current space vector is obtained. The knowledge of \u03b1is\u03bbr allows to transform the space vector electrical quantities from the generic dxqx reference frame to the d\u03bbq\u03bb RFO reference frame. The stator current components in the RFO reference frame are: i\u03bbsd = |is| cos\u03b1is\u03bbr i\u03bbsq = |is| sin\u03b1is\u03bbr (12) instead, the transformed stator flux linkages are: \u03bb\u03bbsd = \u03bb xsd cos\u03b1is\u03bbr \u2212 \u03bb xsq sin\u03b1is\u03bbr \u03bb\u03bbsq = \u03bb xsd sin\u03b1is\u03bbr + \u03bb xsq cos\u03b1is\u03bbr (13) A common practice for the IM FEA is adopting the TimeHarmonic Analysis (THA) approach, in with is also possible to properly modify the effective material permeability to consider the iron saturation" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0002527_iros45743.2020.9341259-Figure1-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0002527_iros45743.2020.9341259-Figure1-1.png", + "caption": "Fig. 1: The diagram shows a moment of the collision situation which pictorially explains \u2220AUO and \u2220GUO.", + "texts": [ + " If \u02da A(tA)B(tB) is a path of a point moving on the periphery of the obstacle from A(tA) to B(tB), S\u02da A(tA)B(tB)G is the time optimal collision free path. Proof of Theorem 1: Let l(t) be the straight line connecting the UAV\u2019s current location and the goal\u2019s location, \u03c6 be the angle between l(t) and the velocity vector ~vu(t). If the UAV is travelling at the maximum speed vu, (5) becomes the cost function and should be minimized to travel to the goal\u2019s position. Let TS be the time when the UAV reaches S and TG be the UAV\u2019s goal reaching time. J(t) = \u222b TG TS vusin(\u03c6)dt (5) If \u2220AUO is \u03b8 and \u2220GUO is \u03b1 (as shown in Fig. 1), the collision constraint could be introduced as (6). If a certain \u03c6 angle is collision free, G(\u03c6) = 0. Where, 0 6 \u03c6 6 \u03c0 2 . G(\u03c6) = max [(\u03b8 \u2212 \u03b1)\u2212 \u03c6, 0] (6) According to (6) the collision free \u03c6 angles are \u03b8 \u2212 \u03b1 6 \u03c6 6 \u03c0 2 . At a given t = T time the minimum cost value will be given when \u03c6 = \u03b8\u2212\u03b1. Therefore, the total minimum cost could be given as in (7). Jmin(t) = \u222b TG TS vusin(\u03b8 \u2212 \u03b1)dt (7) According to Fig. 2, if (7) is satisfied, after t = tA the UAV should reach A(tA). SA(tA) should be a tangent to O(tA)" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0002150_j.engfailanal.2020.105123-Figure1-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0002150_j.engfailanal.2020.105123-Figure1-1.png", + "caption": "Fig. 1. Schematic representation of the sliding bearings used in turbine-generator units of hydroelectric power plants.", + "texts": [ + " The bearings used in turbine-generator units of hydroelectric power plants are called sliding bearing. Sliding bearings are divided into hydrodynamic and hydrostatic. The hydrodynamic bed oil film is formed spontaneously. On the hydrostatic bed, an oil pump from the outside supports the bearing oil. The bearing shells work together with the bearing pad. The oil film formed between the bearing shell and the bearing pad has very large radial or axial loads depending on the position of the rotating mass. Fig. 1 is a schematic view Y. Tas\u0327g\u0131n and G. Kahraman Engineering Failure Analysis 121 (2021) 105123 of the bearings in the Francis turbine of a hydropower plant with a vertical shaft. The lubrication must be perfect for a sliding bed to operate safely. In order to prevent warming and surface deterioration caused by the friction of the bearing surface, the sliding surfaces must be completely separated from each other by the oil film and the liquid friction event must take place. In order to realize the fluid friction in the sliding bearings, the conditions given below should be met by using hydrodynamic lubrication theory" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0003907_s0263574714000058-Figure6-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0003907_s0263574714000058-Figure6-1.png", + "caption": "Fig. 6. (Colour online) (a) eSwarBot platforms presented in Couceiro et al.13 (b) Electrical modification of XBee Series 2 from Digi International40 to provide the RSSI signal output.", + "texts": [ + " One of the most well-known reactive protocols is the AODV. The AODV routing protocol is one of the most adopted reactive MANET routing protocols.39 This protocol exhibits a good performance on MANETs, thus accomplishing its goal of eliminating source routing overhead. Nevertheless, at considerably high rates of node mobility, it requires the transmission of many routing overhead packets. Despite this limitation, the AODV has been extensively applied in most wireless equipments, such as the one used on the robotic platforms eSwarBots13 (Fig. 6(a)); and the Original Equipment Manufacturers RF (OEM-RF) Xbee Series 2 from Digi International40 (Fig. 6(b)). Under the AODV protocol, when a robot A needs to communicate with robot B, it broadcasts a route discovery message to its neighbors (i.e., local broadcast), including the last known sequence number for that destination.41 The route discovery is flooded through the network until it reaches a robot that has a route to the destination. Each robot that forwards the route discovery creates a reverse route for itself back to robot A. When the route discovery reaches a robot with a route to robot B, that robot generates a route reply that contains the number of hops necessary to reach robot B and the sequence number for robot B most recently seen by the robot generating the route reply", + ", communication reflection and refraction), thus resulting in http://journals.cambridge.org Downloaded: 01 Jul 2014 IP address: 155.97.178.73 the loss of packets with increase in inter-robot distance. In fact, preliminary experiments to test the XBee modules on the same scenario showed that the connectivity starts failing above 10 m (Fig. 9). Therefore, to allow a more realistic and conservative approach, the connectivity between robots was maintained using the received signal quality. To this end, the XBee modules were modified to provide the RSSI signal output (cf., Fig. 6(b)). This RSSI output is available as a pulse width modulation (PWM) signal of 120 Hz, where the duty cycle DC varies according to the signal level relative to the receiver sensitivity it follows: DC \u2248 38 + 0.1 \u00d7 RSSI, (7) in which the parameters of the straight-line equation were obtained in the equipment data sheet.13 For instance, a 30% duty cycle (i.e., 1.5 V) is equivalent to approximately the receiver sensitivity of \u201394 dBm. In order to choose a minimum signal threshold that would ensure MANET connectivity, Fig" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0001654_tmag.2020.3019082-Figure3-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0001654_tmag.2020.3019082-Figure3-1.png", + "caption": "Fig. 3. Mesh model of 24/18 SRM.", + "texts": [ + "036 Number of coil turns 17 Space factor of coil 45 % Coil connection Series Core material (Fig. 4) 35A300 0 0.5 1 1.5 2 2.5 10 100 1000 10000 100000 M ag n et ic f lu x d en si ty ( T ) Magnetic field intensity (A/m) Fig. 4. B-H curve of 35A300 (core material). C. Torque Comparison In order to evaluate the torque characteristics of 3 models, 2-D finite element analysis is conducted, where a commercial software (JMAG-Designer (JSOL Corporation)) is used. A half model shown in Fig. 2 is used, where the mesh shape is a triangle (Fig. 3), and slide meshes are applied to the air gap. The rotor position is applied so that the rotation speed will be at 1500 rpm, and pulse voltages are applied to each phase, where the DC-link voltage is 550 V. In order to apply pulse voltages on the coils, the electrical circuit shown in Fig. 1(c) is coupled with the electromagnetic analysis. The leading angle and application width of the applied voltage are adjusted to increase the torque. The maximum phase current is 270 A, and a conventional hysteresis control is applied to 3 models so that the phase current can be lower than 270 A" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0002200_022054-Figure2-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0002200_022054-Figure2-1.png", + "caption": "Figure 2. Direction of velocity vectors in Cartesian coordinate system: - vector of the pert periphery speed, - vector of the delivery speed, - vector of the tool vibration speed, - angle between the directions of the part speed and the tool vibration speed.", + "texts": [ + " In processing according to this scheme, there is a constant contact of the indenter with the processed surface, which significantly changes the stressed state in the contact area [7 \u2013 9]. In introducing vibrations tangentially the processed surface, the tool displaces relative to it at a speed , the value of which is determined by the equation: = + + (1) where - vector of the pert periphery speed, - vector of the delivery speed, - vector of the tool vibration speed. The character of the relative position of these speeds in ultrasonic surface deforming is presented in figure 2. ICMTMTE 2020 IOP Conf. Series: Materials Science and Engineering 971 (2020) 022054 IOP Publishing doi:10.1088/1757-899X/971/2/022054 The speed (m/sec) in processing the cylindrical parts is determined by the equation: = 1000 60 (2) where D - diameter of the processed part (mm), n \u2013 number of the part rotations per minute. The delivery speed (m/sec) is determined by the following equation = 1000 60 (3) where S - tool delivery (mm/rev). The tool vibration speed is presented by the equation: = sin( ) (4) where A0 - amplitude of ultrasonic vibrations, - radiant frequency ( = 2 ), f \u2013 vibration frequency, (sec-1), t - time (sec) The speed of the tool vibration motion will be equal to = cos( ) (5) Let us consider that in processing there is a tool point contact with the part surface, and the trace width, which is left by the tool on the part surface, is the width stipulated by this contact" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0000115_012066-Figure4-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0000115_012066-Figure4-1.png", + "caption": "Figure 4. An example on how the collimator fastening and the back wall can be modernized: before (a) and after (b) modernization", + "texts": [], + "surrounding_texts": [ + "In the modern age the tendency to an increase of the assortment and quality of manufactured metal products is becoming more noticable, as well as the improvement of the structural, operational and other technical characteristics of products and their surfaces. In this regard, the tasks are to replace obsolete production equipment with modern high-performance machines, in particular with computer numerical control (CNC), which can ensure production competitiveness of many enterprises by improving the quality of products and the production pace. The combination of various integrated systems and standard traditional metal processing equipment opens up the opportunities not only to expand production technological functions, but also to reduce production time, as well as to improve the technological, precision and functional characteristics of a product. An example is a regular micro- and macrorelief, regulation of roughness and microtopology formation on metal surface using a laser after a workpiece processing on the CNC machines. One of the most promising examples of integrated systems is pulsed fiber laser sources equipped with a scanning beam guidance system. The use of laser systems makes it possible to efficiently perform additional macro- and micro-processing of material surfaces directly at the stage of manufacturing parts, as well as to refine these surfaces to the necessary technological requirements." + ] + }, + { + "image_filename": "designv11_71_0000050_expat.2019.8876524-Figure6-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0000050_expat.2019.8876524-Figure6-1.png", + "caption": "Fig. 6. Applied pressure on the base bracket.", + "texts": [ + " The development of the base brackets, by means of 3D printing technology, instead of a standard commercial connections was substantiated in the fact that it supports the teaching of this type of technology. The stresses considered were focused on the pressure exerted by the compressed air of the network in the brackets, 7 bar, and the simulation has been performed in a 3D modelling program: SolidWorks\u00ae and SolidWorks Simulation. The results allowed to conclude that the component would be capable to sustain the efforts predicted. Some of the results through this analysis and simulation, can be seen in Fig. 6 and Fig. 7. With the components and systems defined, the laboratory base prototype that would serve as the basis for controlling and supervising the system was assembled, shown in the Fig. 8. The installation of the electropneumatic process was guided by a model developed with the use of the 3D modelling program, Solidworks\u00ae (see Fig. 3), in order to try to ensure the minimum tolerances necessary for the correct operation of the process. To program the PLC SIMATIC S7-1516-3 PN/DP developed by Siemens, different programs standards can be used" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0001046_ut.37.003-Figure1-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0001046_ut.37.003-Figure1-1.png", + "caption": "Fig 1: UUV coordinate systems", + "texts": [ + " Section 2 of the present paper describes the kinematic model for the unmanned underwater vehicle. Section 3 discusses the principle of adaptive cruise controller. Simulation experiments and their results are then presented for two cases: straightline and curve-line movement. In the last section, conclusions and discussion of future scope for improvement are presented. It is necessary to obtain a model of the UUV system with an assigned frame of reference. It is possible to build up two frames: inertial frame and body-fixed frame, as shown in Fig 1. There are six degrees of freedom (DOF): surge, sway, heave, roll, pitch and yaw. The present study defines q = [u v w p q r ]T as the UUV spatial velocity state vector with respect to its body-fixed fame, and h = [x y z f q Y]T as the position and orientation state vector with respect to the inertial frame. The specifications of the vehicle are illustrated in Table 1. According to the thruster distribution as shown in Fig 2, Sea-Kite II is an under-actuated UUV, and there are only four DOFs: surge, heave, pitch and yaw" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0001381_s1068798x20060210-Figure2-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0001381_s1068798x20060210-Figure2-1.png", + "caption": "Fig. 2. Sudnishnikov\u2019s test bench.", + "texts": [ + " One of the first was designed in the laboratory of the Institute of Mining, Western Siberian Branch, Academy of Sciences of the USSR, under the direction of B.V. Sudnishnikov; it is known in the literature as the UIPU universal test bench for jackhammers. This system was used to plot the air pressure in the jackhammer cylinder and the striker path, so as to determine the number of impacts and the work of each impact. Researchers strove to reproduce actual operating conditions on the test bench. 45 In Fig. 2, we show the Sudnishnikov test bench with a hammer drill [4]. Left column 1 supports pneumatic device 2 for the attachment of jackhammer 3 and its handle, as well as guide 4 for the hammer bit 5. Support 6 of drum 7 is mounted on the right column 8, where table 9 is mounted. Photographic-recording drum 7 sits on the table, together with the electric drive and housing 10 with an automatic photographic gate and a light source 11. These attachments rest on ring 12. Jackhammer 3, which rests on bit 5, is attached to piston rod 13 of the clamping unit" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0002075_s42417-020-00261-y-Figure2-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0002075_s42417-020-00261-y-Figure2-1.png", + "caption": "Fig. 2 Dynamic model of the compound planetary gear train with marine twin-layer gearbox case", + "texts": [ + " This finding may expand the literature on vibration isolators and rotating machinery equipment, and provide guidance for the optimization of the marine gearbox case structure. Figure\u00a01 shows a schematic diagram of the compound planetary gear train with a marine twin-layer gearbox case, which consists of a differential stage and an encased stage. The numbers in Fig.\u00a01 represent the elements of the compound planetary gear train with marine twin-layer gearbox case, and the names of their corresponding elements are listed in Table\u00a01. The\u00a0cylindrical vibration isolators are placed radially between the inner and outer shells. 1 3 Figure\u00a02 illustrates the dynamic model of the compound planetary gear train with the marine twin-layer gearbox case. The differential stage uses the rotational coordinate system. In Fig.\u00a02, HI, VI, Hp, and Vp are the rotating coordinate systems based on the rotation angular velocity \u03c9c of carrier H. The encased stage is fixed on the inner shell of gearbox. The encased stage uses the fixed coordinate system. HII, VII, Hm, and Vm are the fixed coordinate systems. X, Y, and Z in the global coordinate system mean axial, vertical, and horizontal directions, respectively. The Z and Y directions in the global coordinate system are the same as the HII and VII directions in the fixed coordinate system", + " Hs2, Hmj, and Hr2 represent the horizontal displacements of Zs2, Zmj, and Zr2, respectively. Vs2, Vmj, and Vr2 mean the vertical displacements of Zs2, Zmj, and Zr2, respectively. \u03c8rmj (\u03c8smj) is the position angles of the internal (external) mesh pair, which can be expressed by Eq.\u00a0(13): The resultant meshing force for elasticity and damping is defined by Eq.\u00a0(14): where S\u0307i(t) represents the first-order derivative of equivalent displacement. Ci denotes the meshing damping factor. According to Fig.\u00a02, differential equations of the compound planetary gear train are constructed by Newton\u2019s second law. (11) { rpi = pi+ rp spi = pi \u2212 sp . (12) \u23a7 \u23aa\u23aa\u23a8\u23aa\u23aa\u23a9 Srmj(t) = \ufffd xmj + Hmj sin rm \u2212 Vmj cos rm \ufffd \u2212 \ufffd xr2 + Hr2 sin rmj \u2212 Vr2 cos rmj \ufffd \u2212 ermj(t) Ssmj(t) = \ufffd xs2 \u2212 Hs2 sin smj + Vs2 cos smj \ufffd \u2212 \ufffd xmj + Hmj sin sm + Vmj cos sm \ufffd \u2212 esmj(t) , (13) { rmj = mj + rm smj = mj \u2212 sm . (14)Pi(t) = Ki(t)Si(t) + CiS\u0307i(t), i = spi, rpi, smj, and rmj, In the differential stage, differential equations can be expressed by 1" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0002690_ssrn.3281368-Figure1-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0002690_ssrn.3281368-Figure1-1.png", + "caption": "Fig. 1. The diagram of Maglev suspension system", + "texts": [ + " The input u = ucoil is the voltage, the disturbance d = D\u03b2zt is the rail vertical velocity, the output variable y = zt\u2212z is the variation of air gap. System matrices A,B,Bd, C are given, A = \u23a1 \u23a2\u23a2\u23a3 \u2212Rc Lc+KbNc Ap G0 \u2212KbNcApI0 G2 0 ( Lc+KbNc Ap G0 ) 0 \u22122Kf I0 MsG2 0 0 2Kf I2 0 MsG3 0 0 \u22121 0 \u23a4 \u23a5\u23a5\u23a6 , B = \u23a1 \u23a2\u23a3 1 Lc+KbNc Ap G0 0 0 \u23a4 \u23a5\u23a6 , Bd = \u23a1 \u23a2\u23a3 KbNcApI0 G2 0 ( Lc+KbNc Ap G0 ) 0 1 \u23a4 \u23a5\u23a6 , C = [0 0 1] . (33) The physical meanings and value of the parameters of Maglev suspension system are given in Table 1 [23]. The diagram of Maglev suspension system is shown in Fig. 1 [23]. The constraints for Maglev suspension system are listed as follows: the maximum air gap deviation is less than 0.075m, the maximum input coil voltage is less than 300V, and the setting time is less than 3s. Consider the tracking target of output is 0.016 at the start, and changes to 0.021 in 8th second. In the 13th second a step Electronic copy available at: https://ssrn.com/abstract=3281368 external disturbance d = 0.008 is introduced. The parameters of the controller and observer are listed as following:a1 = 320, a2 = 1, k1 = 180, k2 = 2, \u03b1 = 0.6, \u03b2 = 0.7, L = [75 0 0] . The simulation results of proposed FOSMC-DOB are depicted in Fig.2-Fig. 4. According the response curves of the states, output, input and sliding surface shown in Fig.1 (a)-(g), it is shown that the proposed FOSMC-DOB has excellent control performance. It satisfies the constraints of Maglev suspension system both in the dynamic and steady processes. It is obvious that this method has nice robustness to the mismatched disturbance from the 8th second of the curves. The good effect of tracking is achieved. It is embodied in the following parts. The response of system is very soon, steady tracking error is close to 0 and there is no overshoot. Moreover, the chattering problem is nearly eliminated" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0003617_ipec.2014.6870118-Figure1-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0003617_ipec.2014.6870118-Figure1-1.png", + "caption": "Fig. 1. Photo of the tested LIM.", + "texts": [ + " In addition, the authors propose an improved dc decay testing method suitable for determining the parameters of the asymmetric model. When the proposed method is employed, the proposed parameter measurement test is completed by only one simple standstill test. Moreover, troublesome tasks to pull out the neutral point, to lock a mover and to dismantle the secondary-side conductor are all unnecessary. The validity of the proposed method is verified with experimental and simulation results on a 3-phase, 4-pole and 12-slot single-sided LIM. II. Profile of Test Machine Table I, Fig. 1 and Fig. 2 show specifications, photo and structure of the tested LIM respectively. From Fig. 2, it can be seen that the coupling coefficient for a- and b-phase windings is the same as that for a- and c-phase windings, but they are larger than 978-1-4799-2705-0/14/$31.00 \u00a92014 IEEE 3044 the coupling coefficient for b- and c-phase windings. This is because the ends of the mover core exist in case of the LIM. Fig. 3 shows operational impedance loci obtained by the DC decay testing method for a-n, b-n, c-n, a-b, b-c nd c-a winding terminals" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0002156_icee50131.2020.9260646-Figure4-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0002156_icee50131.2020.9260646-Figure4-1.png", + "caption": "Fig. 4. The magnetic flux lines of partial and fully HTS-ISM. a) Partial HTS-ISM. b) Fully HTS-ISM", + "texts": [ + " Real power is actually the ability of a load to convert electrical energy into mechanical energy. Whereas apparent power is caused by the difference between voltage and current. Low power factor increases the current and results in high losses in the motor. The power factor relationship is given in 4: P PF = S (4) Where P is the real power (w) and S is the apparent power (V.A). The partial HTS-ISM power factor is 0.73 and fully HTS-ISM power factor is 0.76 which are obtained in steady state at full load. The magnetic flux lines of partial and fully HTS-ISM are shown in Fig. 4. As can be seen, both structures have created 8 magnetic poles. The intensity of magnetic flux line at fully HTS-ISM is lower than partial HTS-ISM, the reason behind this difference is the angle of radiation of magnetic flux into the surface of HTS tapes, since the geometry of HTS coil are flat, the HTS coil is exposed to perpendicular magnetic flux which decreases the critical current density significantly. In another words, critical current density of the HTS tape is dependent to flux density magnitude and angle of flux radiation, so that by variation of each of these components, critical current density will be changed" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0003146_978-3-662-45514-2_9-Figure2-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0003146_978-3-662-45514-2_9-Figure2-1.png", + "caption": "Fig. 2. Quadrotor with a 2 DOF manipulator", + "texts": [ + " The total thrust and the moments are given in Eq. (4, 5) where is the length of the arm ( ) = \u2212 \u03a9 + \u03a9 + \u03a9 + \u03a9 ( ) (4) = db(\u03a9 \u2212 \u03a9 )db(\u03a9 \u2212 \u03a9 )\u03b3 \u2212\u03a9 + \u03a9 \u2212 \u03a9 + \u03a9 (5) The overall system equations can be summarized as follows where and are symmetric inertia matrix at and propeller inertia, respectively. = \u2013 + + \u2013 \u2013 = 10 \u20130 (6) = \u2212\u2212\u2212 + \u2212 + 00\u2212 \u03a9 + \u03a9 + \u03a9 + \u03a9 = \u2212 ( ) 0 \u22120 \u2212\u2212 0 ( ) \u2212 ( ) (\u2212\u03a9 + \u03a9 \u2212 \u03a9 + \u03a9 ) \u22120 + ( ) db(\u03a9 \u2212 \u03a9 )db(\u03a9 \u2212 \u03a9 )\u03b3 \u03a9 \u2212 \u03a9 + \u03a9 \u2212 \u03a9 The structure of the proposed 2 DOF manipulator integrated to the quadrotor is shown in Fig. 2. The first link of the arm is horizontal to the quadrotor\u2019s body Z ( ( )) axis and can revolve with an angle . The second link\u2019s turning axis ( )is horizontal to first link\u2019s axis ( ( )) and can turn with an angle of \u03b1. The rotation sequence from body axis to link 1 and to link 2 are given in Eq. (7). ( ) = ( )\u2212 \u2212 \u2212 \u2212\u2212\u2192 ( ) = ( )\u2212 \u2212 \u2212 \u2212 \u2212\u2212\u2192 (7) The equations of motion of the quadrotor with 2 DOF arm can be derived by using Newton-Euler equations which express the force and moment balance on each body" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0001821_j.matpr.2020.08.097-Figure3-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0001821_j.matpr.2020.08.097-Figure3-1.png", + "caption": "Fig. 3. The approximate profile of the tube when it first touches the walls of the hole: 1 is tube; 2 is spindle; 3 is roller; 4 is tubesheet.", + "texts": [ + " Then the plastic limit of the pipe resistance p and the ultimate pressure pB are determined by the formulas p \u00bc rTffiffiffi 3 p 1 k 1 b b2 1 \u00fe k\u00fe 2klnb2 \u00f08\u00de pB \u00bc rTffiffiffi 3 p 2 1 a\u00f0 \u00delnb\u00fe b2 1 ;a \u00bc k\u00fe 1=b \u00f09\u00de Here b, b1, b2 are the parameters that determine the ratio of the inner radius, the depth of the plastic layer at the beginning and end of plastic deformation to the outer radius of the pipe, respectively; k is the coefficient of hardening of the material, which can be is determined by the formula k2 \u00bc 1\u00fe 0:08k; k \u00bc 1 2l 1\u00fe 2l \u00f010\u00de Here l is Poisson\u2019s ratio. In the process of rolling, the mechanism of pressing the pipe to the wall of the hole is as follows. The pipe is gradually distributed and at some point in time (hereinafter referred to as \u2018\u2018instant\u201d), at certain points, the pipe contacts the hole wall. The number of such contact points corresponds to the number of rolling rollers z [16]. In Fig. 3 shows the contact schema for a 3-roller tool (rolling). After the first contact and further rolling of the rollers, the contact is broken due to the elastic resistance of the tube. After the first contact and further rolling of the rollers due to the elastic resistance of the tube, the contact is broken, but the position of the tool body containing the thrust bearing relative to the tube does not change. In this case, the tube is usually, as a rule, secured against rotation. After a certain period of rotation, the rollers again fall into the zones located in the vicinity of the points of first contact" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0002543_iros45743.2020.9341403-Figure4-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0002543_iros45743.2020.9341403-Figure4-1.png", + "caption": "Fig. 4. Robot kinematics and notation.", + "texts": [ + " The control cost of a feasible path for the robot is defined as: J = \u222b \u03c4 0 (1 + uTt Rut)dt (4) where \u03c4 is the arrival time and R \u2208 Rm\u00d7m is a predefined positive-definite matrix used to weight the cost of the control inputs relative to each other. Finding a feasible path equals to calculating a control input sequence U = {u0, ..., un}. In this section, the nonholonomic property of the mobile robot is taken into consideration while constructing the control input sequence U = {u0, ..., un}. First, consider the kinematic equations describing the mobile robot: x\u0307 = v cos \u03b8, y\u0307 = v sin \u03b8, \u03b8\u0307 = \u03c9. (5) As shown in Fig. 4, let v and \u03c9 be the translational and angular velocity of the robot, \u03c1 the Euclidean distance between the center of the robot and the goal position, \u03b1 the angle between the y-axis of the robot reference frame and the vector connecting the robot and the goal position, \u03b8 the angle between the robot translational velocity v and the x-axis of the goal position reference frame, \u03c6 the angle between the x-axis of the robot reference frame and the x-axis of the goal position reference frame. This representation method makes a Cartesian-to-Polar coordinate transformation and the new kinematic equations can be described as: \u03c1\u0307 = \u2212v cos\u03b1, \u03b1\u0307 = v sin\u03b1 \u03c1 \u2212 \u03c9, \u03c6\u0307 = \u2212\u03c9" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0001304_j.matpr.2020.05.658-Figure6-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0001304_j.matpr.2020.05.658-Figure6-1.png", + "caption": "Fig. 6. Total deformation.", + "texts": [ + " The material chosen is stainless steel. Design structure analysis is also termed as directional uctural analysis of wheel chair using a rocker bogie mechanism, Materials Fig. 8. Normal stress distribution. Fig. 9. Equivalent stress distribution. deformation analysis, wherein, displacement of the system is checked by applying different forces along the Y axis. Here, one face of the system is fixed and a total force of about 100 N is applied on the system with a gravitational pull of 9.81 m/s along Y axis of the system. Fig. 6 shows the design structure analysis of the wheel chair using a rocker bogie mechanism. The total deformation in the rocker bogie arrangement is found to be negligible for the applied load for the static structural analysis. The total deformation of the chair is within the acceptable limits. The hydrostatic pressure analysis is carried Please cite this article as: S. Seralathan, A. Bagga, U. K. Ganesan et al., Static str Today: Proceedings, https://doi.org/10.1016/j.matpr.2020.05.658 out with a pressure of about 500 Pa applied on the fixed part of the rocker bogie system" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0001087_j.ijmecsci.2020.105636-Figure1-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0001087_j.ijmecsci.2020.105636-Figure1-1.png", + "caption": "Fig. 1. Illustrations of (a) largely deformable paper strips with extremely low rigidity, (b) stiff glass plates with low flexibility, and (c) extensible metabeams with high bending stiffness and well recoverability.", + "texts": [ + " The reported HBM accurately predict the mechanical behavior of extensible metabeams, which can be used to design plate-shaped advanced structures for applications requiring rigidity and well deformation recovery. 1 c t m r d p i r s t t d c i m w p s e T t o t M w s m t h R A 0 . Introduction Natural materials typically suffer severe limitations in mechanical haracteristics, which is one of the major obstacles in improving tradiional structures to obtain advanced functionalities. For example, soft aterials (e.g., paper strips in Fig. 1 (a)) are observed with complete ecovery from extreme deformations; however, they are insufficient in eformation resistance. On the other hand, brittle materials (e.g., glass lates in Fig. 1 (b)) have relatively large stiffness but undesirable flexibilty. Therefore, research efforts have been dedicated to developing mateials that perform opposite mechanical characteristics at the same time, uch as both rigid and robust. Other than achieving this goal through he material path (namely, series of nanotube-related materials, funcionally graded materials, etc.), metastructures have opened an exciting oor of solving the challenge through the structure path (i.e., periodially corrugated local structures)", + " Elastic chiral March 2020 m w i t ( p i i H s c m p o i l c a H b c s D e a [ t e t t r o o t t t t c t t a t c c S ( s o t t t S p H 2 m 2 c c t c p a g a u m s w o c l m m c t p a t c t t c m s s u r b e t i o \ud835\udc3c fi echanical metamaterials comprising pairs of hinged ridges squares ere reported in overall 1D shape [6] and 3D shape [5 , 11] . A majorty of studies in the literature have been conducted on bulk-like metasructures, only a few studies have focused on plate-like metastructures i.e., aspect ratio \u226b 1). To bridge this research gap, we recently reorted nanoscale metabeams with hexagonal corrugation (as illustrated n Fig. 1 (c)), and experimentally and analytically studied their promisng behaviors in sharp bending and deformation recovery [13 , 14 , 19] . owever, the metabeams were assumed to be unstretchable in those tudies. To investigate the influence of structure expandability on the mehanical behavior of metabeams, a size-dependent, large deformation odel based on modified couple-stress theory is to be developed in this aper. Although the mechanical response of the metabeams could be btained via theoretical method, it is worth pointing out that the solvng process requires high computational cost; let alone closed-form soutions may not be obtained for many cases with different boundary onditions [1] ", + " The periodic corrugation paterns significantly affect the mechanical response, and thus, numerical alibration is conducted to investigate the metabeams\u2019 effective material roperties (i.e., bending and compressive stiffnesses). Numerical models re built in Abaqus R2017x, and the dynamic/explicit solver with Nleom is particularly used. The length L and width W of the metabeams re fixed as 1 mm and 0.5 mm while the axial displacement ?\u0302? is given p to 85% of the length (i.e., ?\u0302? = 0 . 85 mm ). In order to capture the deforation caused by the face sheet (as shown in Fig. 1 (c)), the size of the hell elements (S4R) is chose as 1/4 of the rib width W hex . In particular, e have the convergence issue when the element size is larger than 1/3 f the rib width while reducing element size significantly increases the omputational cost. Table 1 summarizes the element size, type, and the oading conditions of the FE models. Fig. 2 displays the numerical modeling and results of the extensible etabeams subjected to an axial displacement. Fig. 2 (a) presents the esh, boundary and loading conditions of metabeams with hexagonal orrugation pattern" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0003544_cta.2127-Figure1-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0003544_cta.2127-Figure1-1.png", + "caption": "Figure 1. Wheatstone bridge for measuring self-inductance.", + "texts": [ + " Then in Section 3, we establish a measuring armature inductance method using a modified direct current inductance bridge. Analysis and discussion are presented for validating the measured results and impact of the measured inductance values in the model in Section 4. Section 5 shows the experimental setup and the numerical validation of DC machine model parameters. In the final section, a conclusion is given. Our proposed approach initially adopted a method in [9] which uses a Wheatstone bridge circuit for measuring a self-inductance L1, as shown in Figure 1. Here V1 is a DC supply to the Wheatstone bridge circuit. The bridge becomes balanced once it reaches V=0. If V1 is removed by opening the switch at t=0, the measured L1 in this circuit can then be expressed as follows: L1 \u00bc 1\u00fe R1 R2 I1 \u222b\u221e0 Vdt: (1) The derivation of (1) considers both of its steady and transient states in the circuit. The Wheatstone bridge reaches its steady state after the switch has been closed, and then R2, R3, and R4 are adjusted until V=0. Thus, a steady state can be obtained when the bridge is balanced by R1 R2 \u00bc R3 R4 " + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0002989_978-3-642-16214-5_9-Figure9.6-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0002989_978-3-642-16214-5_9-Figure9.6-1.png", + "caption": "Fig. 9.6. Example of a key lock and shift lock system", + "texts": [ + " On the one hand, the selector lever can only be moved from P if the ignition key releases it; on the other hand, the ignition key can only be removed if the selector lever is locked in P (parking lock engaged in the gearbox). In this way, rolling away and unauthorised movement of the vehicle is prevented. There are two common key lock systems on the market. The mechanical variant connects the locks in the ignition lock and gearshift system with a cable control, while the electric variant has one locking magnet in both the ignition lock and gearshift system. The locking magnet is operated by means of sensor signals (Figure 9.6). Since electric key lock systems have to lock without electricity, they are usually supplemented with an emergency release so that the vehicle can still be moved in case of emergency. The shift lock system in automatic gearshift systems has become the safety standard, serving as an additional safeguard against unintentional starting of the vehicle. Customarily, the selector lever is fixed in position P with an electric actuator, while release is only possible by activating the brake. In European vehicles, the shift lock function usually also takes hold in position N" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0002117_0954410020969319-Figure1-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0002117_0954410020969319-Figure1-1.png", + "caption": "Figure 1. The frame definitions of UAA path-following.", + "texts": [ + " The remainder of this paper is organized as follows: UAA modeling and problem formulation are presented in Section 2. In Section 3, the BLF-based finite-time adaptive NNs backstepping control approach is developed. The stability analysis is presented in Section 4. In Section 5, some simulations are provided to demonstrate the effectiveness of the control method. The conclusion is given in Section 6. Problem formulation and preliminaries Model description The 3-D path-following problem is shown in Figure 1, where {I}, {B}, and {SF} are denoting the inertial reference frame, body-fixed frame, and SF frame, respectively; P denotes the virtual reference target point in the SF frame. Q represents the center of gravity of an airship. h1\u00f0s\u00de and h2\u00f0s\u00de are given as the curvature and torsion of the path. To reduce the overall complexity, the effects of UAA rolling angle and rate are neglected. Thus a 5-DOF model can be established. Then, the kinematic equation of a UAA can be described as follows6 x \u00bc .1 (1) where x \u00bc \u00bdx; y; z; h;w T represents the position and attitude of the UAA in {I}; 1 \u00bc \u00bdu; v;w; q; r T denotes the velocities and angular velocities of the UAA expressed in {B}; " + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0003788_detc2014-34213-Figure4-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0003788_detc2014-34213-Figure4-1.png", + "caption": "Fig. 4 A six-bar linkage with three prismatic joints", + "texts": [ + " The singularity of the planar linkage happens when the three instant centers I24, I34 and I13 of the four-bar equivalent linkage I12I24I34I13 in Fig. 3(a) become collinear or the instant centers I25, I56, I16 of the four-bar equivalent linkage I12I25I56I16 (Fig. 3b) become collinear. That is, the link 5 is perpendicular to the corresponding ground. When the prismatic joint K is used as the input, the singularity of the linkage happens when the instant centers I12, I16 and I25 in Fig. 3(c) become collinear, i.e., the links AB, CD and EK intersect at the same point. The single-DOF planar linkage with three prismatic joints is shown in Fig. 4. If the joint A is used as the input joint, the singularity of the planar linkage happens when the instant centers I24, I34 and I13 in Fig. 4(a) become collinear or the three instant centers I25, I56 and I16 in Fig. 4(b) become collinear. In other word, the link 5 is perpendicular to the corresponding ground. When the prismatic joint K is used as the input joint, the singularity of the planar linkage occurs when the instant center I12, I16 and I25 in Fig. 3(c) become collinear, or the link AB, CD and EK intersect at the same point. \u03b3 45I 12I 25I 24I 13I 34I 56I 15I \u03b3 45I 12I 25I24I 13I 34I 56I 16I14I 15I 13I (a) (b) Fig. 2 A six-bar linkage with one prismatic joint K 1\u03b8 \u03b2 \u03b7 24I 34I 14I 12I (a) 1\u03b8 \u03b2 \u03b7 15I25I45I 24I 34I 12I (b) 2 Copyright \u00a9 2014 by ASME Downloaded From: http://proceedings" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0003878_s11434-014-0376-5-Figure8-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0003878_s11434-014-0376-5-Figure8-1.png", + "caption": "Fig. 8 The force diagram of the forelimb when horse standing", + "texts": [ + " The forelimb and hind limb both can be equivalent to plane open-chain five-link mechanisms. Although the length of links and the postures of two mechanisms are different, the characteristics of two mechanisms are similar. Therefore, the forelimb model is chosen as the object of force analysis, which could be also applied to the hind limb. The horse body is keeping stationary when standing. Connecting with the horse body permanently, the scapula of the forelimb is also stationary, as showing in Fig. 8. The ground reaction force to the forelimb foot is F. To keep the leg equilibrium, the drive moments added to each joint are denoted by M1, M2, M3 and M4. Each link in the five-bar mechanism must be equilibrium by external forces and external moments. Figure 9 shows the force diagram of the i-th link. Mi-1,i represents the i-th link moment exerted by the (i-1)-th link, Mi?1,i is the i-th link moment exerted by the (i ? 1)-th link. Fi - 1,ix and Fi - 1,iy refer to the x component force and y component force of the Fi-1,i exerted by the (i-1)-th link on i-th link, Fi" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0001949_j.mechmachtheory.2020.104139-Figure7-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0001949_j.mechmachtheory.2020.104139-Figure7-1.png", + "caption": "Fig. 7. Maximum creep coefficient Cr m of conformal ball CVT at reference condition: (a) 3D surface map, and (b) 2D contour map.", + "texts": [ + " In this respect, the traction capability is defined as the range that the ball CVT can drive without reaching the creep limit and is evaluated by the creep coefficient. Let the larger of Cr i and Cr o be Cr m . Because Cr is large when the torque is high, Cr m is Cr i for acceleration and Cr m is Cr o for deceleration, as determined at t m . Since Cr m reaching the creep limit value of the lubricant can be treated as a problem due to temperature rise, Cr m is designed to be below a certain limit in traction CVTs. A 3D surface map, Fig. 7 (a) shows the distribution of Cr m according to and the region of t m where Cr m is less than 0.06 at the reference condition. In the reference condition with the condition of s r ID = 1 . 0 , due to spin loss, the traction force at the output disk becomes smaller than the traction force at the input disk. Accordingly, in the case of s r ID = 1 . 0 , m becomes i . As shown in Fig. 7 (b), Cr m increases as t m or increases and, for any t m , there is a range of at which Cr m can exist below a certain value. The creep limit is the point where Cr m corresponds to the limit value, and appears as a curved line in the t m \u2212 coordinates. The creep limit value of this lubricant is 0.04 [ 23 , 27 ], as shown in Table 1 , so the creep limit line is indicated by the red billow dashed line (~~) in Fig. 7 (b). Because by definition t m is the amount of torque divided by the load and size, it yields the torque density itself. Accordingly, the maximum torque density according to becomes t m on the creep limit line. In Fig. 7 (b), the maximum torque density is when is about 41, and its t m value is about 0.0928. Fig. 8 (a) provides an efficiency contour map in the t m \u2212 coordinates of the reference condition. Most power transmis- sion devices are generally highly efficient. For example, gears that are very commonly used for power transmission usually have transmission efficiencies in excess of 95%. Therefore, it is natural for users that power transmission devices have high efficiency. Accordingly, we defined efficiency above 0" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0002861_978-3-030-55889-5-Figure6.20-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0002861_978-3-030-55889-5-Figure6.20-1.png", + "caption": "Fig. 6.20 Gravity load in a wall plate on two point supports, a principal stresses, b stresses \u03c3yy in horizontal sections (BE-solution)", + "texts": [ + " 6.16 and 6.17. No need to develop infinite stresses. 376 6 Singularities 6.7 Cantilever Wall Plate 377 Remark 6.2 Numerical tests prove that force couples do not produce singular stresses, 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 one really goes all the way in, as in Fig. 6.20, one notices the infinite stresses. In highly stressed mechanical parts as turbine blades such stress singularities can be relevant for the design (stress intensity factors). Also influence functions must cope with the singularities on the boundary, because these singularities pollute the FE-solution via an unavoidable \u201cboundary element\u201d mechanism: Also FE-solutions are potentials, since they are the superposition of the load with influence functions, see Sect. 9.18. To understand this, we consider a membrane, which covers an L-shaped opening, see Fig", + "36) Column #1 is the Kirchhoff shear of the Gi , column #2 are the moments, next come the normal derivatives and finally, in column #4, the Gi itself. So much for the theory. In practice, however, the effects of singularities are unlikely to be so dramatic, since engineering accuracy is not that demanding, and the experienced engineer has a well-developed sense of what is credible and what is not. Finite elements in structural analysis are both, modeling and \u201cslide rule\u201d, and the engineer is therefore very flexible\u2014not to say: indulgent\u2014in the interpretation of FE-results. In the case of the wall plate in Fig. 6.20 the tensile stress at the bottom increased from \u03c3xx = 168 kN/m2 to \u03c3xx = 220 kN/m2 after an adaptive refinement, see Fig. 6.23, which is an unusually large increase. Not in all cases will the difference be so large; unfortunately, no fixed rules can be established. We think stress peaks are \u201charmless\u201d, if the plastic zone, which may form, does not cause massive compensating motions. In the case of the wall plate in Fig. 6.23 it is probably different, since it is only held fixed by two \u201ctiny\u201d point supports" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0000435_022055-Figure1-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0000435_022055-Figure1-1.png", + "caption": "Figure 1. Rotary hydraulic machine (2\u00d74) [3]", + "texts": [ + " The reserve for efficiency upgrading for the considered hydraulic machines is the use of the teeth with half-round convex-concave side surfaces. Volumetric pumps and hydraulic machines are considered to be one of the widely spread and actively developed types of machines. Among them planetary rotary hydraulic machines (PRHM) with floating satellite wheels are known. This type of hydraulic machines has \u201crecord\u201d specific throughput in the conditions of sufficiently high pressure (about 200-250 atm.) [1, 2]. The important and understudied characteristic of PRHM is their mechanical efficiency. One of the schemes of PRHM [3] is shown in figure 1. The machine has rotor 1 with external teeth, stator 2 with internal teeth and floating satellite wheels 3 situated between them. Closed voids are made with mated gear links 1, 2, 3 and flat surfaces of end plates 5. Thanks to wavelike form of the rotor and stator, cavity space changes while rotor 1 spinning. Satellite wheels 3, rolling over between rotor 1 and stator 2, open and block corresponding channels 4. The numbers of waves of rotor \u041c and stator N are proportional to corresponding numbers of teeth Z1 and Z2", + " In table 2 there are also calculated loss coefficients of hydraulic machines, working in the regime of pump d d H 1 ; their efficiency in the regime of motor \u03b7d = 1- \u03c8d and efficiency in the regime of pump \u03b7H = 1- \u03c8H. Let us analyze the possible ways of increasing efficiency of PRHM: \u0410) The decrease of the module and the increase of the number of teeth Z2. However, in this situation, side force, affecting the satellite wheel, increases. In the end, this measure will limit maximum fluid pressure. B) The increase of angle \u03bb (see figure 1) of holding of satellite wheel as much as it is possible. The restrictions are connected in some cases with interference (co-occurrence condition) of central wheels, in other cases with the risk of satellite wheel \u201cfalling out\u201d. C) The change of the form of the teeth \u2013 transformation from involute to jointed engagement [6] \u2013 see figure 2. ICMTMTE IOP Conf. Series: Materials Science and Engineering 709 (2020) 022055 IOP Publishing doi:10.1088/1757-899X/709/2/022055 For PRHM with such engagements the ratio of sliding speed Vs to the calculated circular speed V\u0440 in the engagement: 32 ZV V p s " + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0003338_jae-141850-Figure5-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0003338_jae-141850-Figure5-1.png", + "caption": "Fig. 5. Magnetization process using stator coils.", + "texts": [ + " M \u2032(\u03b8) i+1 = M \u2032(\u03b8) i + r ( M \u2032(\u03b8) i+1 +M \u2032(\u03b8) i ) (17) Where r is the relaxation factor which, for this paper, is considered to be 0.5. The flux density and magnetic field intensity, which are composed of x-axis and y-axis components, are calculated from the FEM. Each component of magnetization is calculated independently from the x-axis and y-axis magnetic field intensity using Everett density distribution. Figure 4 shows the input voltage, current and back EMF characteristics. After the PM is magnetized, the voltage is removed and then, back EMF, current are also zero. Figure 5 shows the magnetization process using a stator coil. With this procedure, it is possible to investigate the influence of the rotor bar eddy current components on the overall magnetization. Figures 5(d) and 6 show the flux plots and magnetizing quantities (flux density By) of the PM after voltage removal. We can see that the flux density of the central points of a PM are smaller than those of the edge points, due to the rotor bar eddy current influence. We can also see that all points of flux density in the PM differ from each other, because the operating points of the PM are different due to the rotor bar eddy current influence, as shown in Figs 7\u20139" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0001701_0954405420949757-Figure3-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0001701_0954405420949757-Figure3-1.png", + "caption": "Figure 3. Discrete points on tooth crest line.", + "texts": [ + " Two endpoints of face line is denoted by A and B, whose coordinates are written as (xA, 0, zA) and (xB, 0, zB). Discrete points are obtained by evenly dividing face line, whose number is denoted by m. The i-th point is denoted by Pi, whose coordinates are written as (xi, 0, zi). xi and zi are calculated as xi = xA + (xB xA)(i 1)=(m 1) zi = zA + (zB zA)(i 1)=(m 1) \u00f012\u00de Discrete points on tooth crest line A circle is formed when point Pi rotates around Z1 axis, the circle intersects tooth surface of convex side and concave side at P0i and P00i in Figure 3, whose coordinates are denoted by (x0i, y 0 i, z 0 i) and (x00i , y 00 i , z 00 i ). P0i and P00i satisfy the following geometric relationshipsffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi x0i2 + y0i 2 p = ffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi x00i2 + y00i 2 p = xi z0i = z00i = zi \u00f013\u00de Let the coordinates of a point of tooth surface of convex side be (x, y, z), when t and u1 are given, (x, y, z) are calculated. The problem of solving (x0i, y 0 i, z 0 i) can be transformed into a numerical optimization problem min f= j ffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi x2 + y2 p xij+ jz zij s:t: p \\ t\\ p, p \\ u1 \\ p, s1 " + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0002811_978-94-007-6046-2_71-Figure6-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0002811_978-94-007-6046-2_71-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_71_0003921_coase.2014.6899477-Figure11-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0003921_coase.2014.6899477-Figure11-1.png", + "caption": "Figure 11. Switch-off and shaft-lock mechanism. (d)", + "texts": [ + " If the mass and moment of inertia of Claw A are so small that we can neglect the inertial torque and the gravitational torque, we can obtain the following motion equation of Claw A: c s k r xv \u2206\u03c9 = , (2) where c is the damping coefficient of Rotary Damper, \u03c9 is the detection velocity level, s is the gear ratio of Gear A to Gear B, kv is the spring constant of Linear Spring B, r is the distance between the shaft axis of Rotary Damper and the attachment position of Linear Spring B, and x\u2206 is the displacement from the natural length of Linear Spring B. From (2), the detection velocity level is represented as / ( )k r x csv \u2206\u03c9 = . (3) We can approximately set the detection velocity level on the basis of (3). 3) Switch-off and Shaft-lock Mechanism Fig. 11 shows the mechanism which mechanically switches off all motors of the robot and locks Shaft A. After Plate A is locked by Claw A or Claw D, each Claw B slides along each Guide Hole B of Plate A by the rotation of Plate B (Fig. 11(b)) and then one of three Claws B is hooked to the inner teeth and rotates Plate C (Fig. 11(c)). By the rotation of Plate C, Pin C switches off and Claw C moves along Guide Hole A (Fig. 11(d)). After that, Claw C meshes with Ratchet Wheel A and thus Shaft A is locked. The lock of Shaft A is released by rotating Shaft A in a direction opposite to the direction locked by the safety device. We developed a cylinder-shaped robot to verify the effectiveness of the new safety device. Fig.12 shows the developed robot. The diameter is 0.6[m] and the height is 0.7[m]. The mass is approximately 25[kg]. The robot has two wheels, two motors, one safety device, and two casters. Each wheel is controlled by each motor" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0001034_ec-06-2019-0272-Figure2-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0001034_ec-06-2019-0272-Figure2-1.png", + "caption": "Figure 2. Equivalent dynamics model of GTS", + "texts": [ + " In the same planetary stage: the three planet gears distributed evenly along the circumference direction and their rotate speeds are same at any time, namely, v j pi \u00bc v j p; the external working pressure angles between each of the planet gear and sun gear are same, namely, aj spi \u00bc aj w; the internal working pressure angle between each planet gear and ring gear are equal, namely, aj rpi \u00bc aj n. Based on lumped-parameter method and Lagrange general function (Hedlund and Lehtovaara, 2008; Velex and Ajmi, 2007), the equivalent dynamics model of GTS is established and shown in Figure 2, composed of mass (with inertia but no flexibility), spring (no inertia) and damper (no inertia). The part performance is transformed into four distributed components in four directions including translational degrees of freedom in x, y and z directions and rotational degree of freedom. The coordinate systems of a sun gear, ring gear are supposed to be fixed while coordinate systems of each planet gear can rotate along the center of sun gear: j along the line, which connects the center of planet gear and sun gear; h is the tangential direction of the base circle of the planet carrier" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0002830_978-94-007-6046-2_42-Figure3-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0002830_978-94-007-6046-2_42-Figure3-1.png", + "caption": "Fig. 3 A 2D biped robot simulation by Gubina, Hemami, and McGhee [2]. (a) 2D model of a massless biped. (b) Walk simulation by linear control with nonlinear compensation", + "texts": [ + " For example, in [19], an analog computer was used to confirm the control law. In 1971, Hall and Witt reported a detailed control of lateral stepping motion [3]; however, a report for further walking control has not been published. In the above work, the body posture control was not discussed since the model did not contain sufficient degrees of freedom. Gubina, Hemami, and McGhee [2] discussed the control of body posture as well as the gait stability using a strict equation of motion for a biped robot constrained in a sagittal plane as shown in Fig. 3a. Their model consists of a single rigid body with mass m and moment of inertia J supported by two massless legs represented by linear force generators. The support leg contacts with the ground at point and instantaneously exchanges with the swing leg. Each support phase is represented as an under-actuated system of three degrees of freedom with two inputs. Its behavior is described by three nonlinear equations derived from Lagrange\u2019s method: Rr C l R 2 sin. 1 2/ r P 1 2 l P 2 2 cos. 1 2/ C g cos 1 D F =m (8) r2 R 1 C rl R 2 cos. 1 2/ C 2r Pr P 1 C rl P 2 2 sin. 1 2/ gr sin 1 D M=m (9) l Rr sin. 1 2/ C lr R 1 cos. 1 2/ C .J =m C l2/ R 2 C 2l Pr P 1 cos. 1 2/ lr P 1 2 sin. 1 2/ lg sin 2 D M=m (10) It is impressive to observe such complicated equations that are necessary to simulate the simple mechanical system of Fig. 3. As an attempt to harness this complexity, the following feed-forward terms of the leg force F and the hip torque M were introduced to keep the body height and its upright posture: (The leg force F is taken along the \u201cvirtual leg\u201d which connects the ankle and the hip joint. F is positive when it acts to extend the virtual leg.) F D mg (11) M D rlF sin 1 r C l cos 1 (12) The gravity compensation (11) is specified to balance an equilibrium point in state space, whereas the hip torque of (12) can keep the robot body upright at all state under ideal conditions without disturbances. The dynamics of (8), (9), and (10) are linearized around the equilibrium point and transformed into the following standard linear system representation: Px D Ax C Bu; (13) where x is a state vector of six dimension and u is an input vector of twodimensional. Then a feedback controller was designed for this linearized system. Using the feedback controller with the feed-forward of (11) and (12), a successful 2D dynamic walking was simulated (Fig. 3b). As we observed, it is common to use approximations and linearization to design a suitable gait and controller, even if we have precise equations of the dynamics. Many researchers also presented the usefulness of various model simplifications to analyze global dynamics of biped locomotion [1, 15, 17]. In the next section, we explain the derivation of a linear model for biped walking which was inspired from these pioneering works. The first version of the linear inverted pendulum mode was derived by revisiting the 2D massless leg model of Gubina et al" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0000338_iecon.2019.8926794-Figure1-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0000338_iecon.2019.8926794-Figure1-1.png", + "caption": "Fig. 1. The system configuration", + "texts": [ + " In this section, the authors will explain the system configuration (hardware and software) of our one-tenth size model car created by them. In this study, in order to give versatility to future applications, we installed the Robot Operating System [4] (ROS) as a framework on the model car. ROS can control various sensors, actuators, and so on, as separate nodes, simultaneously. In addition, ROS is versatile in education, because it is compatible with programming languages such as C++, python and java. Figure 1 shows the system configuration of the car. A Raspberry Pi 3 model B is used for it. Because sources of various \u201cprograms and examples\u201d using a Raspberry Pi are published on the internet. In addition, Ubuntu which can be installed as Raspberry Pi\u2019s OS officially supports ROS. Also, essential sensors for autonomous driving such as Light Detected And Ranging (LiDAR). 978-1-7281-4878-6/19/$31.00 \u00a92019 IEEE 6871 Figure 2 shows the model car. Th model car is created based on the Tamiya chassis: \u201c58600 TT-02 Type S Chassis 4wd Kit" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0001846_eit48999.2020.9208283-Figure7-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0001846_eit48999.2020.9208283-Figure7-1.png", + "caption": "Fig. 7 Sprocket and spline coupling for chain drive", + "texts": [ + " Due to space restrictions and a mismatch between the differential and motor splines, they were connected using a chain drive rather than directly, as shown in Fig. 6. The chain used was an ANSI #40 chain. Special spline couplers were fabricated to mate with the output of the motor Authorized licensed use limited to: University of Glasgow. Downloaded on November 01,2020 at 13:18:06 UTC from IEEE Xplore. Restrictions apply. and the differential. They were fitted with chain sprockets to mount the chain, as shown in Fig. 7. differential (right) The accelerator and brake pedals were fabricated and assembled, after a couple of prototypes were tried, as shown in Fig. 8. The accelerator was connected to the MCOR unit (sends signal to the controller) via a small rod. The brake was connected to the drum brakes in the wheel hubs via a standard brake cable. The bench and pick-up bed (Fig. 9) were assembled from strut and aluminum sheet metal was used to comprise the surfaces. Seat belts were mounted to the bench. All these were fixed to the floor and/or the frame" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0003788_detc2014-34213-Figure18-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0003788_detc2014-34213-Figure18-1.png", + "caption": "Fig. 18 Eight-bar linkages with input joints A0, B0 and C0", + "texts": [ + " Therefore, the three-DOF eight-bar linkage is at the singular positions. Input joints not in the same six-bar loop In this group, the inputs are given through the joints not in the same six-bar loop of the eight-bar linkages. Two cases can be found according to whether one of the joints B, F and B0 is the input joint. When the input joints contain one of the joints B, F and B0, such as the input joints given through A0, B0 and C0, the eight-bar linkage can be decomposed into a Stephenson six-bar linkage with the input joints A0 and B0 held as shown in Fig. 18(a) or a two-DOF seven-bar linkage with the input joint B0 held as shown in Fig. 18(b). Thus, for the Stephenson six-bar linkages or the seven-bar linkage, the singularities happen when the links AE, B0B and CD intersect at a common point. Therefore, the three-DOF eight-bar linkage is at the singular positions. If the input joints do not contain one of the joints B, F and B0, such as the input joints given through A0, A and C0, the eight-bar linkage can be decomposed into a single-DOF six-bar linkage with the input joints A0 and C0 held as shown in Fig. 19(a) or a two-DOF seven-bar linkage with the input joint A0 held shown in Fig", + " Applying Sine-law for triangle EGB, there exists, BEG BG BGD \u2220 = \u2220 sin || sin |BD| (18) Similarly, applying Sine-law for triangle BGD, there exists, BDGsin |BG| BGDsin |BD| \u2220 = \u2220 (19) where 312BGE \u03b8\u2212\u03b8=\u2220 , 34EGB \u03b8+\u03b8\u2212\u03b7+\u03b2=\u2220 , \u03c0+\u03b8\u2212\u03b8=\u2220 74BDG , 47BGD \u03b8\u2212\u03b8=\u2220 . Therefore, if the three lines AE, BF and CD intersect at a common point G, the resultant equation (17) can be obtained from Eqs. (18) and (19). Thus, Eqs (17) is the condition that the three lines AE, BF and CD intersect at a common point G, as shown in Fig. 18(a). Therefore, the same result can be achieved while the proposed method can be use to obtain the singular configurations directly. This paper offers a degenerated method to analyze the singularity (dead center position) of multi-DOF multiloop planar linkages. The proposed method is based on the singularity analysis results of single-DOF planar linkages and the less-DOF planar linkages. For an N-DOF (N>1) planar linkage, it generally requires N inputs for a constrained motion. By fixing M (M 120 Pa and a shear stress > 71 Pa. From simple loading conditions Mohr\u2019s circle can be constructed and the highest normal and the highest shear stress can be determined and thus prediction of whether the construction will deform or totally collapse can be made" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0002222_icstcc50638.2020.9259658-Figure3-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0002222_icstcc50638.2020.9259658-Figure3-1.png", + "caption": "Fig. 3. Definition virtual steering axle and angle in grey.", + "texts": [ + " (7) The back-to-front velocity propagation for cases where at least one D = 0 is treated in different ways, e.g. [2], [15]. For the path-tracking task, it is customary to describe the movement with respect to the path, since it simplifies the control problem. Modeling of path-dependent relations is described e.g. in [2], [12], [16], [17]. In order to develop the offset relations, we introduce the equivalent virtual steering axles and their angles in the hitch points that describe the current body movement in Fig. 3. Note that from (3) and with the virtual steering axles, the relation \u03b8\u0307i = vi Li tan \u03b4i\u22121 (8) follows for each body. For the backward path tracking, it is convenient to take axle pm of the last trailer as the reference point of the vehicle. Fig. 4 shows the last trailer with the relations to the path. Newly introduced symbols are the curvilinear distance s along the path and the curvature \u03ba of the path. Subscript d denotes the desired values. Therefore, the desired path point is pd = pp(sd). 490 Authorized licensed use limited to: Univ of Calif Santa Barbara", + " . . 0 ) , i.e., the output of the system is set to the off-track distance yos = xos,1 = dos. The symbols f , g : M \u2212\u2192 Rn represent vector fields and h :M\u2212\u2192 R is a scalar field, with the open subsetM\u2286 Rn. For this system, a stable backward path tracking controller is designed by applying the exact linearization technique. Therefore, the task can be expressed as lim t\u2192\u221e xos(t) = 0. An equilibrium of the system is, for instance, a movement on a circle, following xos = 0. Geometrically, regarding Fig. 3, one obtains the following relations for that steady state regime: 1 \u03bai = 1 \u03bai+1 cos\u03c6i + Li+1 sin\u03c6i (24) and \u03c9i+1 = \u03c9i \u21d2 vi+1\u03bad,i+1 = vi\u03bad,i. (25) 491 Authorized licensed use limited to: Univ of Calif Santa Barbara. Downloaded on May 21,2021 at 03:33:25 UTC from IEEE Xplore. Restrictions apply. From the system equations for the steady state movement 0 = f(0) + g(0)uos, (26) the relation for the input in that state is uos = tan \u03b4m\u22121 = \u03bad,mLm, (27) and the relation for the relative angles follows as 0 = Li+1 sin\u03c6i \u2212 (Di + Li+1 cos\u03c6i) tan \u03b4i" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0003741_itec-ap.2014.6941162-Figure4-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0003741_itec-ap.2014.6941162-Figure4-1.png", + "caption": "Fig. 4. The simplified structure of the axial MFM-BDRM", + "texts": [ + " Therefore, the choice of pp and ps for the axial MFM-BDRM usually conforms to the principle that the greatest common divisor between the pp and ps is 1. The model with the pole-pair number of the stator being 3, the pole-pair number of the PM rotor being 19 and the polepair number of the ferromagnetic segments being 22 is picked to analyze the simplified model. As the axial MFM-BDRM is symmetrical along z=0 offset position versus xy-plane, the whole-structure model can be simplified to the semi-structure model, as shown in Fig. 4 a) and b). For the quarter-structure model, boundary conditions of the magnetic field produced by the stator and the PM rotor are both odd, which means that the model can be simplified to the quarter-structure model, as shown in Fig. 4 c). As the greatest common divisor between the pp and ps is 1, the model could not be simplified any more. With the PM rotor operating at the rotational speed of 2000 rpm, three models are simulated to assess the simplified method. The no-load and load air-gap flux density distributions under average radius are shown in Fig. 5 and 6. The flux density waveforms of the quarter-structure distribute Fig. 3. Variation of cogging torque and torque ripple with the polepair number of the PM rotor when the pole-pair number of the stator remains to be 4" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0000236_smc.2019.8914349-Figure3-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0000236_smc.2019.8914349-Figure3-1.png", + "caption": "Figure 3. Transportation/Assembling robot.", + "texts": [ + " CASE STUDIES The theoretical results provided in the previous section are validated through engineering applications, which confirm easiness of the design, simplicity of the realization, and effectiveness of the proposed controller. Example 1. Consider the case studies given in [15] and [18]. It is easy to verify that the several pseudo-PD controllers 2(2 2 ) cu G a e ae u designed in [15] and [18] can be replaced with the more efficient pseudo-PID controllers 2 36 4 3 .cu G a e a ed ae u Example 2. Consider the transportation/assembling robot shown in Fig 3, with 1 20.50 , 1 , 0.50 .H m L m L m Suppose that the aim of the robot is to pick up by the endeffector an object from the point (0,1,0) of the workspace and drop it off at the point (2.867, 0.50, 0) with the vertical final position of the end-effector, following the joint space reference trajectories shown in Fig. 4. Suppose that the operating values of the robot parameters are: M 10 (mass of the base),b Kg 2 (mass of the first link)vM Kg 20.1 (inertia moment of the first link)vI Kgm 1 1 (linear density of the second link)m Kg m 2 =1 (linear density of the third link)m Kg m [0" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0000161_012109-Figure2-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0000161_012109-Figure2-1.png", + "caption": "Figure 2. Adaptive grinding head control system", + "texts": [ + " As a tool for processing parts of complex aerodynamic shape the scientists used abrasive flap wheels, the processing of which is at high performance, low tool costs and the possibility of using both in single and in mass production [13-16]. An important task in grinding large-sized surfaces is the determination of the position of the flap wheel relative to the workpiece [17]. This condition ensures uniform removal of the stock from the surface to be treated. To solve this problem the scientists developed an adaptive control system for the position of the grinding head. It works using a pair of ultrasonic sensors installed in a seal at the ends of the working window. (fig. 2). In the figure 2, numerals denote the following elements: 1 \u2013 work surface; 2 \u2013 ultrasonic sensors; 3 \u2013 resilient seal; 4 \u2013 flap wheel. This system functions as follows (fig. 3). ICI2AE 2019 IOP Conf. Series: Materials Science and Engineering 632 (2019) 012109 IOP Publishing doi:10.1088/1757-899X/632/1/012109 Before processing, we fix the part (panel or fuselage shell) in a vertical position in the fixing device. Next, we set the required width of the abrasive flap wheel and control at the working position with the help of the corresponding positioning unit. After, at some distance, there is a coordination alignment of the center of the width of the flap wheel with the middle line of the processing band. Then, the grinding head is moving towards the surface to be treated, while using the sensor set 2 by the adaptive control system, the distances L1 and L2 to the surface 1 are continuously monitored (fig. 2). During processing, the sensors maintain constancy of distances, and, if necessary, correct it by tilting the housing from or to the surface at an angle \u03b1 (fig. 3\u0430). If necessary, there is the correction of the movements along the axis \u0394Y and \u0394Z (fig. 3b). The angle of inclination and corrective movements is calculated by the following formulas: H SS arctg 21 (1) 2 2 COS SinSinR Z (2) 2 2 COS COSSinR Y (3) where S1 \u0438 S2 \u2013 readings of the upper and lower ultrasonic sensors, H \u2013 distance between sensors, R, \u03b2 \u2013 polar coordinates of the center of seal" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0000023_012025-Figure4-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0000023_012025-Figure4-1.png", + "caption": "Figure 4. Instrument \"Orlik\" company DorAgroMash", + "texts": [ + " Manufacturers of this design machines: Strip-Cat, Orthman, B & H manufacturer, BighamBrothers, Hiniker, Carter, Yetter, Kongskilde, and Sunflower / Agco. This model is the most popular presented on the market. Type III: Based on its own system, Kuhn / Krause has developed a machine that, like in type I, has a cutting tool fixed to the machine frame. The working body, which purpose is loosening (for the equipment of this company type, Krause is the rack) is attached to the machine frame with a parallelogram (like type II), and the grinding working body leads the tool in depth (Figure 4). Such a functional principle is a distinctive model feature of the third type. Accordingly, the grinding working bodies should be made of wear-resistant materials. The exclusion from the grinding working body section will directly lead to a change in the loosening working body depth, which will complicate the regulating process such processing. Tools for band tillage, common in the Russian Federation: Orthman (1tRIPr), KRAUSE (Gladiator), as well as their domestic counterpart of DorAgroMash Orlik are quite expensive, which affects the final products price" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0003337_9781118886397.ch1-Figure1.19-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0003337_9781118886397.ch1-Figure1.19-1.png", + "caption": "Figure 1.19 Illustration of Amp\u00e8re\u2019s law for a resistor and capacitor.", + "texts": [ + "110) This can be thought of as an electric circuit composed of a loop of wire in combination with a time-varying magnetic field produced by a spinning magnet, the combination being an electric generator, Figure 1.18. In integral form, the moving magnetic field provides the work to move an electric charge along a specific path. \u222eL E \u22c5 dl = \u2212 d dt \u222bS D \u22c5 dS (1.111) Amp\u00e8re\u2019s law with Maxwell\u2019s displacement current correction states that a changing electric field produces a corresponding magnetic field, as in Figure 1.19, \ud835\udec1\u00d7H = Jf + \ud835\udf15D \ud835\udf15t (1.112) In point form, you imagine a moving electric charge producing a magnetic field according to the right-hand rule. In integral form, the work required for a magnet to move along a path L is provided by electric charges moving across the boundaries of a volume. \u222eL H \u22c5 dl = \u222bS Jf \u22c5 dS + d dt \u222bS D \u22c5 dS (1.113) Gauss\u2019s law for electric fields in volume form is, \u201cThe electric flux through any closed surface is proportional to the enclosed electric charge.\u201d A \u201cGaussian surface\u201d is an arbitrary closed three-dimensional surface (such as a sphere) which is chosen to ease the computation of volume integrals" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0001138_0954406220925836-Figure12-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0001138_0954406220925836-Figure12-1.png", + "caption": "Figure 12. Topological arrangements with internal articulated moving platforms.", + "texts": [ + " The full-cycle rotation of the end-effector realized through the revolute joint is directly actuated by the middle limb, and the rotation is decoupled from other motions. By adding serial kinematic chains to the above-proposed mechanisms, the GPMs with high rotational performance by combining different articulated moving platforms can be obtained. Therefore, the construction of this kind of GPMs is simplified as assembling decoupled serial limbs on the presented mechanisms. The available topological structures for GPMs with AMP3 are as drawn in Figure 12. Taking the 5-DOF GPMs with AMP1 and AMP2 as examples, the 3T2R-AMP2AMP3 GPMs is developed by assembling a 6-DOF RUPU serial kinematic chain and AMP3 on the existing 3T1R-AMP2 GPM. And the resulting GPM can output large rotational angles in two directions, as shown in Figure 13(a). By arranging the RUPU serial kinematic chain and AMP3 to the proposed 2T2R-AMP1AMP2 GPM, the 2T3R-AMP1AMP2AMP3 GPM with rotations about three directions is obtained, as drawn in Figure 13(b). Through similar analysis, the 6-DOF GPMs with AMP3 can be synthesized by adding middle limbs and AMP3 to the established 5-DOF GPMs with AMP2. Considering the structural symmetry of mechanisms, only the topological arrangement depicted in Figure 12(b) are sued to design GPMs with AMP3. While for 6-DOF mechanisms, six active joints are necessary to fully control the mechanisms. For the RUPU middle limb, the revolute joint and the prismatic joint are equipped with actuators. When constructing 6-DOF GPMs with AMP3, the articulated moving platform is working under the combined action of AMP2[AMP2 and AMP3. If two R3RMPaR3R 0 3 H2C-limbs are connecting to the intermediate moving platform, the directions of constraint-couples are parallel. To guarantee this requirement is satisfied during continuous motion, the axes of joints R3 are parallel to or coaxial, and two groups of parallel revolute joints are arranged on the intermediated moving platform" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0003499_muh-1404-6-Figure14-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0003499_muh-1404-6-Figure14-1.png", + "caption": "Figure 14. The first 6 mode shapes of the machine tool structure.", + "texts": [ + "7 \u00b5m, while the y deformation is equal to 7.1 \u00b5m. Generally, in the x direction (YZ plane), the static loop stiffness is found to be equal to 32.5 N/\u00b5m, while in the y direction (XZ plane), it is found to be equal to 70.4 N/\u00b5m, and therefore the static loop stiffness in the y direction is much better. Figure 13 shows the deformation all over the machine tool structure. Modal analysis has been performed on two cases: first, without including the bolt\u2019s prestress effect, and second, when the bolt\u2019s prestress effect is included. Figure 14 shows the first six mode shapes of the machine structure in the first case, while Table 2 represents the natural frequencies and the positions of maximum deformation at each mode. From the results obtained in both cases, the first natural frequency is 56 HZ in the first case, while it is 48 HZ in the second; therefore, the dynamic performance when putting the bolts before stress into account is lower, yet more realistic. they can be considered the weakest part in the structure. The base and bed show the minimum deformation in almost all mode shapes" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0003883_pomr-2014-0016-Figure4-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0003883_pomr-2014-0016-Figure4-1.png", + "caption": "Fig. 4. Definition of unit normal vector n and pressure distance \u0394s", + "texts": [ + " For a single panel the hydrodynamic force is calculated using the formula: (4) where: RHD \u2013 hydrodynamic resistance, CF, CP \u2013 friction and pressure force coefficients, VT, Vn, vT, vn \u2013 tangent and normal velocity vectors and absolute values, respectively. Definitions of VT and Vn, are given as: Vn = (V \u00b7 n)n; VT = V \u2013 Vn (5) On the hull of the lifesaving module, a number of bearing nodes were selected which were the objects of action of the ramp pressure forces. The value of the force is defined by the linear function of the crossing distance of the basic (undeformed) ramp surface by the bearing node, see Fig. 3 and Fig. 4. This linear function has been limited by the maximal pressure reaction, introduced to limit modelling of the bearing pressure forces to the range observed in real conditions (for instance due to the loss of stability of the structure, or exceeding the yield point). The pressure force acting on the ramp: (6) where: Espr \u2013 modulus of elasticity [N/m], \u0394s \u2013 distance between the bearing node and the base surface, see Fig. 2, n \u2013 unit normal vector Unauthenticated Download Date | 3/30/18 11:31 AM 37POLISH MARITIME RESEARCH, No 2/2014 The rolling friction force is a linear function of the pressure force" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0000802_gncc42960.2018.9018937-Figure3-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0000802_gncc42960.2018.9018937-Figure3-1.png", + "caption": "Figure 3. The definition of cgx", + "texts": [], + "surrounding_texts": [ + "( ) ( )1 2 1 1 1 1\n1 1 1 1\n[\n\u02c6 tanh( ) ] d\nT d\nx g f k e\nw v e y\u03c3\n\u2212= \u2212 \u2212\n\u2212 \u2212 +\nx x\n(13)\nwhere 1 1 Tw v is the output of the RBFNN for approaching 1\u0394 . And 1 1tanh( )e\u03c3 is robust item for eliminating approximation error 1\u03b5 brought by RBFNN. And 1 0k > , 1 0\u03c3 > .\nThe output of the RBFNN is described as\n( ) ( )n n= =T TN w, x w v w v x (14)\nwhere 1( ) ( ) ( ) p n n p nN N R = \u2208 N w, x x x is the output vector of RBFNN. k n R\u2208x is the input vector of RBFNN. 1 l p\np R \u00d7 = \u2208 w w w is the weight of RBFNN. l is the number of RBFNN nodes.\n[ ]1( ) ( ) ( ) l n n l n R= \u2208v x v x v x , ( )i nv x is the radial basis function, it is commonly taken to be Gaussian function (15). 2\n2( ) exp( ) ( 1, , )i n i v i l b\n= \u2212 =n ix - c x (15)\nwhere kR\u2208ic is the center vector of the Gaussian function,\nib is the base width of the Gaussian function. RBFNN is a linear parameterized neural network and network output is a linear combination of basis functions. Its structure is shown in the figure below.\nInvoking (13) and (12) yields\n( )1 1 1 1 2 1 1 1 1 1tanh( )Te k e g e w v e\u03b5 \u03c3= \u2212 + + + \u2212x (16)\nThe parameter 1k is used to adjust the dynamic characteristics of the system, and the parameter 1\u03c3 improves the robustness of the system. In this paper, the selection of parameters is only theoretically given constraints. In practice, these parameters are determined by means of simulation test.\nStep 2. The derivative of 2e can be described as\n( ) ( )2 2 2 2 2 2 2d de x x f g u x= \u2212 = + + \u0394 \u2212x x (17)\nAccording to equation (8), the uncertainty which approached by RBFNN can be designed as\n2 2 2 2 2 2 2 2\u02c6( )T Tw v w w v\u03b5 \u03b5\u0394 = + = + + (18)\nThe virtual control du is designed as\n( ) ( ) ( )1 2 2 1 1 2 2\n2 2 2 2 2\n[\n\u02c6 tanh( ) ] d\nT d\nu g f g e k e\nw v e x\u03c3\n\u2212= \u2212 \u2212 \u2212\n\u2212 \u2212 +\nx x x\n(19)\nwhere 2 2 Tw v is the output of the RBFNN for approaching 2\u0394 . And 2 2tanh( )e\u03c3 is robust item for eliminating approximation error 2\u03b5 brought by RBFNN. And 2 0k > , 2 0\u03c3 > .\nInvoking (19) and (17) yields\n( )2 2 2 1 1 2 2 2 2 2tanh( )T de k e g e w v e x\u03c3= \u2212 \u2212 + \u2212 +x (20)\nC. Stability analysis Consider the Lyapunov function as follow\n2 1 1 1\n1\n2 2 2 2\n2\n1 1 tr( ) 2 2 1 1 tr( ) 2 2 T V T L e w w e w w \u03bb \u03bb = + + +\n(21)\nTaking the time derivation of VL , yields\n1 1 1 1 1\n2 2 2 2 2\n1 tr( )\n1 tr( )\nT V\nT\nL e e w w\ne e w w\n\u03bb\n\u03bb\n= +\n+ +\n(22)\nSubstituting (16) and (20) into (22), yields\n2 1 1 1 1 1 1 1 1 1\n2 2 2 2 2 2 2 2 2 2\n1 1 2 2 1 11 2\n[ tanh( )]\n[ tanh( )] 1 1\u02c6 \u02c6\nT V\nT\np p\ni i i i i i\nL k e e w v e e k e e w v e e\nw w w w\n\u03b5 \u03c3 \u03b5 \u03c3\n\u03bb \u03bb= =\n= \u2212 + + \u2212\n\u2212 + + \u2212\n\u2212 \u2212\n(23)\nDesign weights adaptive law as\n1 1 1 1 1 1\n2 2 2 2 2 2\n\u02c6 \u02c6( ) \u02c6 \u02c6( ) i i i\ni i i\nw e v w w e v w \u03bb \u03b7 \u03bb \u03b7 = \u2212 = \u2212\nAnd consider the following facts\n1 1 1 1 1 1 1\n2 2 2 2 2 1\np\ni i i\np\ni i i\ne w e v\ne w e v\n=\n=\n=\n=\nT\nT 2\nw v\nw v\nThen equation (23) can be organized as\nAuthorized licensed use limited to: Western Sydney University. Downloaded on August 15,2020 at 21:56:59 UTC from IEEE Xplore. Restrictions apply.", + "2 2 1 1 2 2 1 1 1\n1\n2 2 2 1 1 1 1 1\n2 2 2 2\n\u02c6\n\u02c6 [ tanh( )]\n[ tanh( )]\np\nV i i i\np\ni i i\nL k e k e w w\nw w e e\ne e\n\u03b7\n\u03b7 \u03b5 \u03c3\n\u03b5 \u03c3\n=\n=\n= \u2212 \u2212 +\n+ + \u2212\n+ \u2212\n(24)\nSince\n2 22 2 1 1 1 1 1 1\n1 1\n2 22 2 2 2 2 2 2 2\n1 1\n1 1 1\u02c6 ( ) 2 2 2 1 1 1\u02c6 ( ) 2 2 2\np p\ni i i i i i p p\ni i i i i i\nw w w w w w\nw w w w w w\n= =\n= =\n\u2264 \u2212 = \u2212\n\u2264 \u2212 = \u2212\n(25)\nTherefore, formula (26) can be obtained.\n2 2 1 1 1 1 1 1 1 1\n2 2 1 1 2 2\n2 2 2 2 2 2 2 2 2 2 1( ) [ tanh( )] 2 1 1 2 2 1( ) [ tanh( )] 2 VL w k e e e w w w k e e e \u03b7 \u03b5 \u03c3 \u03b7 \u03b7 \u03b7 \u03b5 \u03c3 \u2264 \u2212 + \u2212 \u2212 \u2212 + \u2212 + \u2212\n(26)\n2 2 1 1 2 2 2 2 1 21 2\n1 2 1 2\n1 2\n, , 2 2\n, tanh( ) tanh( )\nM M\nm m\nM M\nm m\nk w k w e e\ne e\n\u03b7 \u03b7 \u03b5 \u03b5\u03c3 \u03c3\n> >\n> >\n(27)\nLet 1w and 2w have an upper bound, and 1 1Mw., to simplify the analysis. the time lag in activation of control jet is assumed negligible, except for limit-cycle operation. A simple example illustrating the basic aspects of time-optimal control is shown i n Fig. 2. Assume the vehicle initially at rest in the position shou-n in Fig. ?(a) and that a step angular position command is given requiring i t to assume the angular attitude (shown dotted) along the horizontal axis. When the step command is given, an initial angular error e is generated. The p+l jets will cause an angular acceleration about the controlled axis tending to decrease the error. These jets are used until a rotation equal to half the initial error angle is produced; then the jets are reversed and the p-1 jets are used to decelerate the rotation through the remaining angular error. If all jets are turned off the instant the commanded position is attained, the vehicle final velocity will be zero, and the desired attitude will have been obtained in a minimum of time. The results are shown graphically in Fig. 2(b), where e and i are plotted vs time for this example. -4s described in -Appendix I , the phase plane paths produced by jet utilization are parabolic, with the vertex always on the e axis. The theory shorn that timeoptimal control is achieved by turning on the proper jet with maximum torque in one direction until an intersection with the switching parabola is reached. (The switching parabola is defined as the one passing through the origin in each half-phase plane.) At this point, the opposite jet is turned on to carry the operating point to the origin in the phase plane where angle and angle rate errors are zero" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0002989_978-3-642-16214-5_9-Figure9.5-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0002989_978-3-642-16214-5_9-Figure9.5-1.png", + "caption": "Fig. 9.5. External gearshift system of an automatic transmission: an AT for standard drive", + "texts": [ + " Among existing designs, freewheels are used to support downshifts in the lower gears (e.g. Figures 6.32 and 12.23). Another significant area of application is the Trilok converter. In this case, the freewheel connects the reactor with the housing (see Section 10.4.6 as well as Figure 10.32). The external gearshift system is the interface between the driver and the transmission. In automatic transmissions, the wishes of the driver are transferred to the transmission usually by cable control (mechanical gearshifting, Figure 9.5), but also using electric signals (from tip-shifting to complete shift-by-wire). In passenger cars, the shift lever is customarily placed in the centre console, but also in the dashboard or on the steering column. The bases for the design of automatic shift activation systems are two important safety regulations of the NHTSA (National Highway Traffic Safety Administration) in the Unites States. FMVSS (Federal Motor Vehicle Safety Standard) 102 contains passages relevant to the shift pattern and display" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0003183_icra.2015.7139672-Figure1-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0003183_icra.2015.7139672-Figure1-1.png", + "caption": "Fig. 1. Combine harvester", + "texts": [ + " In Section IV, a solution is presented for a single grain cart that serves an arbitrary number of combines. In Section V, we present motion planning analysis for the grain cart. In Section VI, simulations are presented. In Section VII, we conclude with some future work. In this section, we describe the mathematical models for the motion of the combine harvester and the grain cart. Combine harvester is the machine for harvesting crops, for example, wheat, oats, rye, barley corn, soybeans and flax. Figure 1 shows a combine at work. In the active mode, the header cuts the crop and feeds it into threshing cylinder. Grain and chaff are separated from the straw when crop goes through the concave grates. The grain, after being sieved, will be stored in the on-board tank temporarily, and the waste straw is ejected. We use C to denote the maximum capacity of the on-board tank. Threshing loss is an important issue during any harvesting application. For any combine, the quantity of threshing loss primarily depends on the forward speed of the harvester", + " In this paper, we simplify the model and assume that all combines posses identical speeds which is assumed to be constant, and the field has a constant density of crop. Based on the above discussion, the filling rate of the tank, denoted as r f , can be regarded as constant. Since the tank does not have a large capacity, modern combine usually has an unloading auger for removing the grains from the tank to other vehicles. For most of the combines, the auger is mounted on the left side, as shown in Fig. 1. At this point, a vehicle has to be on the left side of the combine to empty the tank. Here we denote the unloading rate of the tank using auger as ru. Therefore, a combine involved in harvesting and unloading operation simultaneously unloads at a rate ru\u2212 r f (ru > r f ). A grain cart, also known as chaser bin, is a trailer towed by a tractor. In this paper, we use term grain cart to represent the system including both the tractor and the trailer. Fig. 2 (a) shows the appearance of a grain cart" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0003618_ijvsmt.2015.067521-Figure7-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0003618_ijvsmt.2015.067521-Figure7-1.png", + "caption": "Figure 7 (a) Free-body diagram of the wheel-tyre set (b) Free-body diagram of bar 1 (see online version for colours)", + "texts": [ + " The angles \u03a8 and \u0393, which define the location of the end-effector in the moving frame can be related to the angles \u03b3 and \u03b5, which define the camber and toe of the wheel in the fixed frame, respectively. The relationship between the angles \u03b3 and \u0393 is given by the equation (17) (see Figure 6) while the relationship between the angles \u03b5 and \u03a8 is given by the equation 18 (see Figure 5). The angles in the both equations are defined applying the right-hand rule. \u0393 2 \u03b3 = \u2212 + \u03c0 \u03c6 (17) ( )arctan y x\u03b5 w w= (18) ( ) ( )0 0 1 0 1 0 \u02c6 \u02c6\u02c6 \u02c6 \u02c6x y uw Ru i w Ru j u d \u2032 \u2032 \u2032 \u2032= \u22c5 = \u22c5 = (19) where 0 1 R is the rotation matrix of the moving frame with respect to the fixed frame. Figure 7(a) shows the free-body diagram of the wheel-tyre assembled to wheel carrier. In order to simplify the force actuators analysis, the following hypotheses are assumed: \u2022 only the mass of the wheel-tyre set is taken into account, hence, the other parts masses are neglected \u2022 the product of inertia of the wheel-tyre set is neglected \u2022 the longitudinal acceleration is not considered \u2022 the wheel vertical motion is only a mathematical function of the body roll angle (it is assumed there are no irregularities on the flat road) \u2022 the vertical and angular accelerations of the wheel-tyre set are neglected \u2022 the rolling resistance at the tyre is neglected", + " By applying the principle of d\u2019Alembert, the equations of motion for the rear left wheel-tyre set are: ( ) ( ) ( ) ( ) ( ) ( ) 22 1 2 3 1 3 1 1 1 0 xp A b b p Gy M A A F F A A F P A F G A mg G A ma + \u2212 \u00d7 + + \u2212 \u00d7 + \u2212 \u00d7 + \u2212 \u00d7 \u2212 \u2212 \u00d7 = (20) 1 2 2 3 0 xA A b b p GyF F F F F mg ma+ + + + + \u2212 = (21) where the vectors pF and pM are input variables associated to the tyre contact between the tyre and the road. The vectors in the equations (20), (21) are defined in the fixed frame. The forces 2bF and 3 ,bF applied by the bars 2 and 3, respectively, on the spherical joints of the end-effector are defined by equation (22) [Figure 7(a)]. ( ) ( )3 32 2 2 2 3 3 2 2 3 3 b b b b A BA B F F F F A B A B \u2212\u2212 = = \u2212 \u2212 (22) The forces applied on the points A1 and A2 are defined by the equation (23) [Figure 7(a)]. 1 1 0 1 0 1 0 2 2 0 2 0 2 0 2 0 2 \u02c6 \u02c6\u02c6 \u02c6 \u02c6 \u02c6 \u02c6 A A x A y A z A A x A y A z A x b F F i F j F k F F i F j F k F i F = + + = + + = + (23) The unknowns variables in equations (20), (21) are FA2x, Fb2, Fb3, FA1x, FA1y and FA1z. The forces Fb2, Fb3 and FA2x are obtained from equation (20). After that, the force components FA1x, FA1y and FA1z can be calculated from equation (21). For the bar 1, the application of the angular momentum theorem leads to [Figure 7(b)]: ( ) ( ) ( ) 11 1 1 1 1 0A SA B F D B F S B F\u2212 \u00d7 + \u2212 \u00d7 + \u2212 \u00d7 = (24) where the vector SF represents the spring force applied on the point S of the bar 1 [Figure 7(b)]. The required force for the actuator 1, F1, is given by equation (24), while the required forces for the actuators 2 and 3, F2 and F3, are given by equation (25). ( ) ( )1 1 2 0 2 1 3 0 3 1 \u02c6 \u02c6 b bF RF j F RF j= \u22c5 = \u22c5 (25) where 1 0 R is the rotation matrix of the fixed frame with respect to the moving frame. The required power for each actuator is obtained from equation (26). 31 2 1 1 2 2 3 3 dsds ds P F P F P F dt dt dt = = = (26) In order to evaluate how promising this mechanism is, two analyses are conducted" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0003223_978-3-319-08648-4_6-Figure6.2-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0003223_978-3-319-08648-4_6-Figure6.2-1.png", + "caption": "Fig. 6.2 Structure of a HOPG, b CNTs and c graphene sheets. The edge-plane-like sites responsible of a faster electron transfer rate are shown", + "texts": [ + " The first analyte is an \u201couter-sphere\u201d compound, insensitive to possible defects and impurities of the electrode, [Fe(CN)6] 3+/2+ is defect-sensitive but not oxide-sensitive compound and Fe3+/2+ is sensitive to both defects and oxygen-containing groups. For all the considered substances, the presence of graphene improves the electron transfer reaction rates of orders of magnitudes with respect to bare electrodes. It has been well established that the origin of the electron transfer for highly ordered pyrolytic graphite (HOPG) comes from the edge plane-like sites. The basal plane of HOPG is electrochemical inert (Fig. 6.2a). Recently, similar considerations have been extended to CNTs and graphene sheets (Fig. 6.2b, c). Both these materials show an anisotropic electron transfer. Nanotube and graphene peripheral ends exhibit an electrochemical behavior similar to the edge plane-like sites/defects of HOPG. On the other hand, a slow electron transfer characterizes both nanotube sidewalls and graphene sides electrochemically resembling the HOPG basal planes. In separated works, Nugent [2] and Bark [3] investigated the redox reaction of [Fe(CN)6] 3+/2+ and found enhanced currents and reduced peak-to-peak separations at MWCNT-based electrodes and at the edge plane of HOPG electrodes" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0001044_pesgre45664.2020.9070523-Figure6-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0001044_pesgre45664.2020.9070523-Figure6-1.png", + "caption": "Fig. 6 Flux distribution due to fundamental current.", + "texts": [ + " The order of the harmonics content is considered as \u2018k\u2019 while deriving r.m.s (4) and fundamental current (5) using Fourier series as per Fig. 5 (b) Hysteresis and Eddy current loss: Hysteresis and eddy current loss are the components of core loss. This hysteresis loss (Ph) is determined through steinmetz equation (6). The hysteresis loss is = = n i ihh fBKP 1 6.1 max (6) Where, Bmax is flux density, f1 is the line frequency and Kh is the loss coefficient. The flux distribution of BLDCM for fundamental current is shown in Fig. 6. Similarly, the eddy current loss including harmonics is = = n i iee fBKP 1 22 max (7) Where, Ke is the eddy current loss coefficient. Equations (6) & (7) indicate that hysteresis and eddy current loss are frequency dependent. Thus, current harmonics reduction ensures significant decrease of these losses. The proposed technique injects 7th harmonics in the system for performance improvement. The flux distribution without 7th harmonics injection is shown in Fig. 7 (a) where Figs. 7(b) and (c) shows the detailed flux distribution with 7th harmonics current 978-1-7281-4251-7/20/$31" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0000347_iecon.2019.8927574-Figure3-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0000347_iecon.2019.8927574-Figure3-1.png", + "caption": "Fig. 3. Schematic of the rotor structure of Spoke-type PM machine.", + "texts": [ + " 978-1-7281-4878-6/19/$31.00 \u00a92019 IEEE 1216 Authorized licensed use limited to: University of Exeter. Downloaded on May 07,2020 at 09:58:38 UTC from IEEE Xplore. Restrictions apply. Not only does the discussed problem exist in stator-PM machines, but also in other types of PM machines especially interior PM (IPM) machines. For instance, the spoke-type PM machine is another interior IPM machine under consideration for electric vehicle applications, but the same aforementioned issue will affect its performance. Fig. 3 shows the schematic of a spoke-type PM machine. It can be seen that the rotor structure has different parts which are kept together using glue. Generally, it is very challenging to construct such structures using traditional laminated steels because they usually have 3D flux paths. Thanks to the advancement of material science, nowadays it possible to realize very complex structure of electrical machines using Soft Magnetic Composite (SMC) material. SMC materials are ferromagnetic powder with a very particular insulated layer allows to reduce the eddy currents in a wide range of frequencies with in comparison with the laminated steels" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0002600_ieeeconf38699.2020.9389412-Figure1-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0002600_ieeeconf38699.2020.9389412-Figure1-1.png", + "caption": "Fig. 1. Line drawing of the SUP platform and thruster configuration.", + "texts": [ + " In the following section a maneuvering model of the system is developed and the control problem is formally stated. The design of a super-twisting path following controller is presented in Section III. The performance of the system is investigated in Section IV using simulations of the closed loop system following a representative lawn-mower survey path. The main conclusions are presented in Section V. Consider a second order dynamical system of the form \u03b7\u0307 = J(\u03b7 )v, (1) Mv\u0307 = N(v,\u03b7 )+ \u03c4 +dE . (2) The terms appearing in these equations are defined in TABLE I. As illustrated in Fig. 1, the equations of motion are expressed in the body-fixed frame of the vessel, where xb is the surge axis, yb is the sway axis and \u03c8 is the heading angle. In the vehicle\u2019s three degree of freedom (DOF) workspace, the pose vector \u03b7 = [x y \u03c8]T represents x (Northward) and y (Eastward) translations, as well as \u03c8 (the angular displacement from North) in a North-East-Down (NED) reference frame. The variable \u03c8 is commonly known as the heading angle and is taken to be positive when the sense of rotation corresponds to the positive z-axis of the NED frame", + "897 m and a draft of approximately T = 1 cm when fully loaded with instrumentation (total weight of the fully-loaded vehicle is m= 15.5 kg). The paddleboard is propelled using a pair of small thrusters (maximum magnitude of their bollard pull thrust is approximately 45 N each), which are positioned at a distance 1.35 m aft of, and \u00b10.15 m to either side of, the center of gravity so that r p := \u22121.35i\u0302\u22120.15 j\u0302, rs := \u22121.35i\u0302+0.15 j\u0302. Each thruster is oriented so that the thrust produced is parallel to the surge axis of the vehicle, i.e. T p = Tpi\u0302 and T s = Tsi\u0302 (see Fig. 1). Owing to their configuration, the propellers can only generate a control force along the surge direction and a moment about the yaw axis (using differential thrust) giving \u03c4 = \u23a1 \u23a3 Ts+Tp 0 T s\u00d7 rs+T p\u00d7 r p \u23a4 \u23a6= \u23a1 \u23a3 \u03c4x 0 \u03c4\u03c8 \u23a4 \u23a6 , (6) where Tp := |T p|, Ts := |T s|, \u03c4x is the total thrust force in the surge direction and \u03c4\u03c8 is the thruster-produced moment about the yaw axis. Thus, using (3)\u2013(4), the dynamic equations of motion (2) reduce to Mv\u0307 =\u2212C(v)v\u2212D(v)v+ \u23a1 \u23a3 \u03c4x 0 \u03c4\u03c8 \u23a4 \u23a6+d . (7) As shown in [9], the inertia tensor M and the Centripetal/Coriolis tensor C(v) can be decomposed into rigid body terms and added mass terms as M =MRB+MA and C(v) =CRB(v)+CA(v), respectively, where the subscript RB denotes a rigid body term and the subscript A designates an added mass term" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0000136_012076-Figure3-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0000136_012076-Figure3-1.png", + "caption": "Figure 3. Effort in the cutting zone", + "texts": [ + " HIRM-2019 Journal of Physics: Conference Series 1353 (2019) 012076 IOP Publishing doi:10.1088/1742-6596/1353/1/012076 The grinder design with combine rollers is implemented as follows: plates in a form of Reuleaux Triangle Profile are located on the shaft with an 120\u00b0 offset relative to each other. Figure 2 shows the material cutting areas. HIRM-2019 Journal of Physics: Conference Series 1353 (2019) 012076 IOP Publishing doi:10.1088/1742-6596/1353/1/012076 During cutting, the guillotine working principle is realized (with knives rotation) with non-linear distribution load [10]. Figure 3 shows directions of effort in the cutting zone along the radius of roll Reuleaux Triangle Profile. During roll rotation, the point of forces application shifts tangentially to the curve formed by the opposite edges of Reuleaux Triangle Profile. Furthermore, due to this design of the rolls, additional destructive forces arise in the transverse gap between the plates, such as crushing and cutting. Moreover, this rolls design allows grinding not only fragile materials, but also plastic, such as machining swaft of non-ferrous metals and alloys" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0000263_icems.2019.8922396-Figure5-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0000263_icems.2019.8922396-Figure5-1.png", + "caption": "Fig. 5. The equivalent rotor positions and axis of dual rotor", + "texts": [], + "surrounding_texts": [ + "Through the part of torque compensation, the optimized vector control and load torque observer (OVC-LTO) are combined in the overall control. Fig 6 shows the block diagram of optimized vector control with load torque compensation. IV. EXPERIMENTAL RESULTS" + ] + }, + { + "image_filename": "designv11_71_0000306_icems.2019.8922343-Figure1-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0000306_icems.2019.8922343-Figure1-1.png", + "caption": "Fig. 1. Structure of the studied PMSM.", + "texts": [ + " PROPOSED LUMPED-PARAMETER THERMAL MODEL In this section, the details of the prototype machine and the proposed LP thermal model are described. In this thermal model, the main parts of the electrical machine are considered. Afterwards, the determination of convectionheat-transfer coefficients of water jacket, end-cap are discussed. The winding region is modelled based on 3-D \u201cTtype\u201d thermal network in cylindrical coordinates and takes P 978-1-7281-3398-0/19/$31.00 \u00a92019 IEEE account into the heat transfer paths in radial, axial and circumferential directions. A 48-slot/8-pole TEWC PMSM, Fig. 1, is used for thermal analysis and its geometric parameters are listed in Table I. Water cooling via the housing jacket is one of the most widely used forced cooling approaches in many applications due to its high heat management capacity. The majority of heat generated within the stator is dissipated to coolant liquid, and a little quantity of the heat generated in the rotor is dissipated to the ambient through the convection heat transfer via airgap and end-cap. Convection is the transfer process due to fluid motion [10], and the convection thermal resistance R for the given surface is defined as: = 1\u210e (1) where \u210e is the convection-heat-transfer coefficient, and is the contact surface" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0003270_icpe.2015.7167862-Figure20-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0003270_icpe.2015.7167862-Figure20-1.png", + "caption": "Fig. 20. Demagnetization patterns of 10-pole 12-slot IPM-type BLDCM.", + "texts": [ + " In addition, in case of 70% demagnetization of N-pole, harmonics of 2nd and 4th are remarkably appeared. C. 10pole-12slot Fig. 19 shows the 10-pole 12-slot IPM-type BLDCM. 10-pole 12-slot IPM-type BLDCM has characteristic from fractional-slot unlike 6-pole 9-slot and 8-pole 12- slot. Namely, this model is different pole-slot combination ratio. In addition, reason of this model selection is to confirm equal BEMF harmonic characteristics from another model of different pole-slot combination ratio. Table IV shows the specification of 10-pole 12-slot IPM-type BLDCM. In addition, Fig. 20 shows the concept figure of 10-pole 12-slot model applying to equality, inequality, and weighted demagnetization patterns. Fig. 21 shows the analysis results of flux linkage and BEMF in the case of equality demagnetization. In case of equality demagnetization, variance of flux linkage and waveform of BEMF are symmetrical because demagnetization ratio of N-pole and S-pole make no difference. Therefore, distortion of waveform is not occurred. Fig. 22 shows the analysis results of flux linkage and BEMF in case of inequality demagnetization" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0002850_978-94-007-6046-2_41-Figure22-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0002850_978-94-007-6046-2_41-Figure22-1.png", + "caption": "Fig. 22 Inverted pendulum model with impacts at instantaneous exchange of support. By assuming the final velocity is zero and working backward to the initial velocity, v0 can be determined in closed form as a function of g, l0 , , and C", + "texts": [ + " These impulses may slow the system down, maintain its forward velocity, or speed it up depending on whether the impulses are supplied as a braking force from the front leg, as a driving force from the rear leg, or distributed between the front and rear leg to cause a net upward impulse on the CoM (Fig. 21). If the front leg acts as an impulsive braking force, the ability to regain balance after a disturbance can be significantly improved over maintaining a constant center of mass height. An inverted pendulum model (IPM) with constant leg length during stance, as shown in Fig. 22, can be used to determine where to step to come to a stop. The system has an initial horizontal velocity of v0 and leg length of l0. It maintains that leg length while rotating about the foot location until the swing leg lands. At exchange of support, the angle from vertical to the original support leg is , while the angle from vertical to the new support leg is C, where the angles are defined so that both values are positive. Due to the nonlinear equations, solving for step location as a function of initial velocity can only be done numerically", + " However, the reverse problem of solving for the initial velocity as a function of the step location can be solved in closed form by starting at the final conditions and working backward. During a stance phase, the total mechanical energy of the system remains constant for an inverted pendulum, E0 D 1 2 mv2 0 Cmgz0 D Ef D 1 2 mv2 f Cmgzf : (75) When the inverted pendulum comes to a rest, the final velocity vf is 0. Solving for the required velocity after the impact, we obtain vC 2 D 2glf 1 cos C ; (76) where lf D l0 cos. / cos. C/ The impact causes a decrease in the forward velocity magnitude. The angle between the preimpact and postimpact velocity vectors (Fig. 22 right) is C C. These vectors will lie on a right triangle if the impact comes from the front leg. This is because vC will be perpendicular to the front leg, requiring v to be parallel to it. This follows from the direction of the impulse being along the leg. Therefore, v D vC cos . C C/ (77) During the first part of swing, the total mechanical energy is conserved, resulting in v0 D q .v /2 2gl0 .1 cos . // (78) These three equations can be used to compute v0 as a function of v , l0, , and C. Additionally, if D CD , and hence l0D lf D l, then we have v0 D r 2gl " + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0001581_s00202-020-01082-8-Figure4-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0001581_s00202-020-01082-8-Figure4-1.png", + "caption": "Fig. 4 Magnetic flux density distribution over the crosssection area of the machine", + "texts": [ + " The stator phases are fed through a three-phase voltage source, and the aluminum-skewed squirrel-cage bars of the rotor are considered as solid conductors with a tilt of 30\u00b0/m and are connected in parallel inside the electrical circuit module. Also, the relative movement between the stator and the rotor is modeled by moving mesh techniques, because of change (1)z = Vvar Is (2)XM + Xls = \u221a z2 \u2212 R2 s in geometry due to the varying relative position between the stator and rotor slots. Similarly, the rotor is rotated at synchronous speed and the stator is connected to the three-phase voltage source. Figure\u00a04 shows the 2D distribution of the magnetic flux density in the cross section of the induction machine and more visibly around the air gap. The magnetizing flux is related to the radial component of the magnetic flux density in air gap that is perpendicular to the surface in the middle of air gap between stator and rotor. Figure\u00a05 shows the radial component of the magnetic flux density in air gap that has been calculated using COMSOL software package. Now, by integrating the flux density over the pole surface, the magnetizing flux m is computed, and then the amplitude of the air gap flux linkage m is calculated as [27] where Ns and are the number of turns in series per stator phase and winding factor for the fundamental wave, respectively", + " As mentioned earlier, the self-inductance of stator that is calculated by 2D FE method does not include the endwinding leakage inductance and Ls can be written as Now, by manipulating Eqs.\u00a0(8)\u2013(11), the following relation is derived by which the stator slot leakage inductance Llss is calculated as The total stator leakage inductance Lls is obtained as the difference between Eqs.\u00a0(2) and (6), and hence, the statorphase end-winding leakage inductance is found as The results are presented for two different motors as case I and case II, separately. Case I: The case I motor is a 1.1[kW], two-pole pairs, 230/400[V] and 50[Hz] induction machine. Figure\u00a04 shows a cross section of the case I machine with rounded trapezoidal rotor bars. Complete mesh consists of 60,000 domain elements and 2400 boundary elements as shown in Fig.\u00a06. The stator and rotor domains are meshed using three-node triangular finite elements, apart from the outer layers of the rotor bars where the special boundary layers are used to take into account more precisely the skin effect. There are 24 stator slots and 14 rotor bars. To model the magnetic characteristic of the iron core that is made of nonoriented 3% Si steel sheets, the corresponding saturation curve has been used [30]" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0003931_j.proeng.2014.12.635-Figure3-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0003931_j.proeng.2014.12.635-Figure3-1.png", + "caption": "Fig. 3. Variation region of bank angle value.", + "texts": [ + " Each input generated at random is kept constant during the prediction horizon , pt t T Different from receding-horizon optimization, solutions obtained using random searching are near-optimal solutions. The normal integration step size 10h s for random searching is adjusted such that searching horizon will not exceed the reference horizon. If altitude is less than or equal to 35 km, the step size set equal to the value in Eq. (13) [19]. finalV V h V (13) where finalV is a pre-specified final velocity. Before random searching starts, the bank rate constraint is used to shrink the searching space, which can increase the tracking accuracy considerably. As seen in figure 3, the variation region of bank angle value (0~180 deg) is split into three sub-regions, i.e. 1 , 2 , and 3 . Searching space at each guidance cycle is determined as 1 max 0 , :l past u past guid \u03c3 if \u03c3 \u03a9 \u03c3 \u03c3 \u03c3 T (14) max 2 max , :l past guid past u past guid \u03c3 \u03c3 \u03c3 T if \u03c3 \u03a9 \u03c3 \u03c3 \u03c3 T (15) max 3, : 180 l past guid past u \u03c3 \u03c3 \u03c3 T if \u03c3 \u03a9 \u03c3 (16) where l is the lower bound of the searching space, u is the upper bound of the searching space, past is the bank angle flown in last guidance cycle, and guidT is the time increment between current time and previous call" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0003753_amr.891-892.81-Figure6-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0003753_amr.891-892.81-Figure6-1.png", + "caption": "Figure 6. Photographs of the two failed air turbine starters.", + "texts": [ + " Thermal cycling generates temperature gradients within the component resulting in differential expansion between the hotter and cooler regions of the disc, resulting in compressive strain in the hotter parts and tensile strain in the cooler parts. In the long term, crazing cracks can merge and propagate as single large fatigue cracks. As a result of the investigation, procedures for inspecting the brake disc were revised to ensure that discs with craze-cracking are identified and removed from service. Two in-service failures of Air Turbine Starters (ATS) from ADF fixed wing aircraft were investigated. Both failures resulted in separation of the blade assembly from the hub, Figure 6. Investigation revealed that in both cases fatigue cracks were present at similar radial locations, Figure 7. The fatigue cracks appeared to have initiated from surface machining marks, Figure 8. The fracture surfaces of both cracks contained a combination of regularly spaced fatigue striations, likely the result of engine starts, and irregularly spaced larger progression marks, representing larger than typical stresses, Figure 9. Crack growth analysis indicated that both cracks were approximately 1" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0001103_s1064230720010037-Figure4-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0001103_s1064230720010037-Figure4-1.png", + "caption": "Fig. 4. Dividing the driving mode into two stages; D, B, start and end points of phase coordinates; 1, optimal mode on DB; 2, 3, optimal modes on and , respectively.", + "texts": [ + "2), it must be minimized with respect to parameters \u03c41, U, \u03c8, and \u03b7. In the problem to be solved, \u03c8 and \u03b7 are taken equal to zero: \u03c8 = \u03b7 = 0. Dividing the studied movement into two stages and determining the optimal mode at each of them does not guarantee optimality as a whole in two stages. However, this can be achieved by selecting the control actions \u03c41 and U. Indeed, if, in accordance with the optimality criterion introduced, the motion mode , which is conditionally represented in the phase coordinates x and in Fig. 4 by curve DB, then, by choosing arbitrary control parameters (point C), optimality can be ensured only on curves DC and CB. In the case that point C is selected on curve DB (point ), then the optimality of modes and will lead to optimality (in the problem being solved to the minimum of functional A) throughout mode DB. Therefore, after determining the optimal modes DC and CB such a point should be defined as (control parameters \u03c41 and U) for which the functional being studied will be minimal. The system of Euler\u2013Poisson equations at each stage of the movement has the form (2" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0003502_1.4933151-Figure1-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0003502_1.4933151-Figure1-1.png", + "caption": "Fig. 1. Square frame rotating about perpendicular axis through center.", + "texts": [ + "4935768 Conceptualizing Rolling Motion Through an Extreme Case Reasoning Approach The Physics Teacher 55, (2017); 10.1119/1.4976657 Statcast and the Baseball Trajectory Calculator The Physics Teacher 55, (2017); 10.1119/1.4976652 Bouncing Poppers The Physics Teacher 53, (2015); 10.1119/1.4933153 482 The Physics Teacher \u25c6 Vol. 53, NoVember 2015 DOI: 10.1119/1.4933151 symmetry. Finally, by the perpendicular axis theorem (see Ref. 3, p. 265), the moment of inertia I of a slender square frame about a perpendicular axis through its center (the zaxis in Fig. 1) satisfies the following relation I Iz = Ix + Iy, which together with (4) where a and m are respectively the length and mass of a side (uniform rod), and Irod is its moment of inertia about a perpendicular axis through midpoint, to be derived in the next section. Armed with the above four formulas, we are now ready to head forward. Consider a uniform rod of length L and mass M = lL (l being the constant linear mass density) rotating about an axis perpendicular to the rod through its midpoint. Let I denote the moment of the inertia of the rod about the named axis" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0001845_icra40945.2020.9196746-Figure8-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0001845_icra40945.2020.9196746-Figure8-1.png", + "caption": "Fig. 8 Detection of product height change.", + "texts": [ + " By driving the drive portion, the first rack and the second rack are moved in opposite directions from each other along the horizontal direction. Gear A1, which meshes with and is driven by the first rack, rotates counterclockwise. Furthermore, Gear B1, which meshes with and is driven by the second rack, rotates clockwise. As a result, the third rack that meshes with Gear A2 and the fourth rack that meshes with Gear B2 move in opposite directions from each other in the horizontal direction. Therefore, the first finger mechanism and the second finger mechanism also move away from each other. Figure 8 shows an example of how the change in product height is detected using each displacement sensor. As shown in Fig. 8(1), product G1 is placed horizontally on product G2. The closed parallel gripper is lowered by the manipulator. When the height information of product G1 is unknown, the first to fourth nails are brought into contact with the upper surface of product G1 so that the first to fourth sensors are displaced when contact with product G1 is made. The control device thus detects that the first to fourth nails have come into contact with product G1, and it immediately stops the descending operation of the parallel gripper. Next, as shown in Fig. 8(2), the parallel gripper drives the first and second finger mechanisms to gradually open while tracing the surface of product G1 with the first to fourth nails. When each nail no longer is in contact with product G1, the response of the displacement sensor changes significantly. A method for deriving the product width Lobj will be described here as an example of the usefulness of the developed gripper. The parallel gripper is set to closed as the initial state. L1 is the amount of movement from the initial state of the first finger mechanism until the response of the displacement sensor of the first and second nails changes significantly" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0001138_0954406220925836-Figure8-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0001138_0954406220925836-Figure8-1.png", + "caption": "Figure 8. 3-DOF GPMs with AMP2: (a) 2T1R-AMP2; (b) 1T2R-AMP2AMP2.", + "texts": [ + " When the R1R1R1 1F2C-limb and the R3RMPaR3 2C-H2limb serve as the connected limbs, the plane determined by the constraint-couples in the 1F2C-limb is parallel to the constraint-couples of the 2C-H2limb. As a result, the revolute joint R1 is parallel to the revolute joint R3. Then the R3RMPaR3 2CH2limb evolves into the R1RMPaR1 2C-H2limb. For the articulated moving platform, to perform large rotational output, one group of parallel revolute joints is required to connect the end-effector. Then the 2T1R GPM with AMP2, signified by 2T1RAMP2, in one direction is obtained, as shown in Figure 8(a). The wrench system of the 1T2R mechanisms consists of two constraint-forces and one constraintcouple. The available groups of two limbs to achieve the expected wrench system can be formed by 2F1C and D, 2F1C and F, 2F1C and C, 2F1C and 1F1C, 1F1C and 1F1C, 1F1C and F, F and 2F, C and 2F. Taking the R1R1R2 2F1C-limb and R3RMPaR3R 0 3 CH2limb as an example, the constraint-couple of the 2F1C-limb is perpendicular to the axes of the revolute joints in the C-H2limb. Hence, the axis of R1 is parallel to the axis of R3. Furthermore, two groups of parallel revolute joints are connecting to the end-effector in turn to output rotational angles. As a result, the R3RMPaR3R 0 3 C-H2limb is transformed into the R1RMPaR1R3 C-H2limb. Where R3 is parallel to the direction of the R2 joint in the 2F1C-limb. Then the 1T2R-AMP2AMP2 GPM that can conduct high rotational performance about two directions is derived, as drawn in Figure 8(b). 4-DOF GPMs with articulated moving platforms When synthesizing 4-DOF mechanisms, four active joints are necessary to fully control the GPMs. Then the available topological arrangement of the structures can be seen in Figure 7(b) and (c). Hence, one serial limb and one hybrid limb with three active joints, or two hybrid limbs with two active joints are organized on the two sides of the end-effector. For 3T1R mechanisms, the wrench system is constituted by two constraint-couples. The combination of two limbs to provide the wrench system can be derived as follows: 2C and D, 2C and C, 2C and 2C, C and C" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0000603_iccas47443.2019.8971558-Figure9-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0000603_iccas47443.2019.8971558-Figure9-1.png", + "caption": "Figure 9. Mid-section kinetic structure", + "texts": [], + "surrounding_texts": [ + "221\nThe first segment is located by the distal part and the second segment is located by the proximal part. The first segment is operated by 2 pairs of 2 wires; total 4 wires. One pair of 2 wires is controlled by one motor that pulls one wire and push the other wire in the same length by using the pulley. While, the second segment is operated by 4 pair of 4 wires, total 8 wires. Therefore the second segment is controlled by 4 motors The second segment has m units and the first segment has n-m units. 4 pairs of wires are labeled by \ud835\udc4e\ud835\udc4e and ?\u0302?\ud835\udc4e, \ud835\udc4f\ud835\udc4f and ?\u0302?\ud835\udc4f, \ud835\udc50\ud835\udc50 and ?\u0302?\ud835\udc50, d and ?\u0302?\ud835\udc51,\nEquilibrium in moments at Un belonging to the first segment is (Sa,n-fa)(an-an-1)\u00d7(an-un)+(Sa\u0302,n-fa\u0302)(a\u0302n-a\u0302n-1)\u00d7(a\u0302n-un)+\n(Sc,n-fc)(cn-cn-1)\u00d7(cn-un )+(Sc\u0302,n-fc\u0302)(c\u0302n-c\u0302n-1)\u00d7(c\u0302n-un)++\nmw(pn-un)\u00d7g= ( 0 0 0 )\n( (9)\nwhere, an-an-1= an-an-1\n|an-an-1| , etc. m\ud835\udc64\ud835\udc64 is a payload applying at the\nend-point and g is the gravity acceleration vector. Equilibrium in moments at Ui , (i=m+1, \u22ef, n-1), belonging to the first segment is (Sa,n-fa)(an-an-1)\u00d7(an-un)+\n(Sa\u0302,n-fa\u0302)(a\u0302n-a\u0302n-1)\u00d7(a\u0302n-un)+ (Sc,n-fc)(cn-cn-1)\u00d7(cn-un )+ (Sc\u0302,n-fc\u0302)(c\u0302n-c\u0302n-1)\u00d7(c\u0302n-un)+\nmp \u2211 (pk-ui) n\nk=i+1\n\u00d7g= ( 0 0 0 )\n(10)\n(10)\nwhere. fa, fa\u0302, fc, fc\u0302 are wire tensions, Sa,i, Sa\u0302,i, Sc,i, Sc\u0302,i , (i=m+1, \u22ef, n) are spring tensions of the ith unit. \u201c\u00d7\u201d means a cross product and \u201c|*|\u201d, means the modulus of a vector \u2217. \ud835\udc5a\ud835\udc5a\ud835\udc5d\ud835\udc5d is the mass of one unit including the plate, the rod and the universal joint.\nThe spring tensions are obtained as, Sa,i\uff1dk(L-|ai-ai-1|), Sa\u0302,i\uff1dk(L-|a\u0302i-a\u0302i-1|),\nSc,i=k(L-|ci-ci-1|), Sc\u0302,i=k(L-|c\u0302i-c\u0302i-1|),)\n(11)\nwith spring coefficient k. Equations (9) and (10) contain 3(nm) equations including 4(n-m) -1variables of the n-m universal joints angles \u03b8xi, \u03b8yi, \u03b8zi, (i=m+1, \u22ef, n) and slide length of plates \ud835\udc59\ud835\udc59\ud835\udc56\ud835\udc56 (i=m+1, \u22ef, n-1). Equilibrium in force at ith plate (i=m+1, \u22ef, n-1) is,\n[-Sa,i+1(ai+1-ai) +Sa,i(ai-ai-1)+-Sa\u0302,i+1(a\u0302i+1-a\u0302i)+ Sa\u0302,i(a\u0302i-a\u0302i-1)-Sc,i+1(ci+1-ci)+Sc,i(ci-ci-1)-Sc\u0302,i+1(c\u0302i+1-c\u0302i))\nSc\u0302,i(c\u0302i-c\u0302i-1)+(n-i)mpg ]\u2219 (pi-ui)=0\n(12)\n(12) provide n-m-1equations. Combined it with (9) and (10), we obtain 4(n-m)-1 equations, which suffices in number to solve for 4(n-m)-1 variables; \u03b8x,i, \u03b8y,i , \u03b8z,i (i=m+1,\u22ef,n) and li (i=m+1, \u22ef, n-1) for a given set of wire tensions fa, fa\u0302, fc, fc\u0302.\nEquilibrium in moments at Um , the universal joint located at the most distal position belonging to the second segment is\n-Sa, m+1(am+1-am)\u00d7(am-um)+(Sb,m-fb)(bm-bm-1)\u00d7(bm-um)\n-Sa\u0302,m+1(a\u0302m+1-a\u0302m)\u00d7(a\u0302m-um)+(Sb\u0302,m-fb\u0302) (b\u0302m-b\u0302m-1) \u00d7(b\u0302m-um) -Sc, m+1(cm+1-cm)\u00d7(cm-um)+(Sd,m-fd)(dm-dm-1)\u00d7(dm-um)\n-Sc\u0302,m+1(c\u0302m+1-c\u0302m)\u00d7(c\u0302m-um)+(Sd\u0302,m-fd\u0302) (d\u0302m-d\u0302m-1) \u00d7(d\u0302m-um)\n+ (mw(pn-um)+mp \u2211 (pk-um) n-1\nk=m+1\n) \u00d7g= ( 0 0 0 )\n(13)\nFor the second segment, we can derive similar equations as (10), (11) and (12) by replacing {ai, a\u0302i, ci, c\u0302i} with {bi, b\u0302i, di, d\u0302i}, {Sa,i, Sa\u0302,i, Sc,i, Sc\u0302,i} with {Sb,i, Sb\u0302,i, Sd,i, Sd\u0302,i} for i=1, \u22ef, m-1 in (10) and for i=1, \u22ef, m in (11) and (12).\nAs a result, we obtain 4m equations included by (13), which suffices in number to solve for 4m variables; \u03b8x,i, \u03b8y,i, \u03b8z,i and li (i=1, \u22ef,m) for a given set of wire tensions \ud835\udc53\ud835\udc53\ud835\udc4f\ud835\udc4f, \ud835\udc53\ud835\udc53?\u0302?\ud835\udc4f, \ud835\udc53\ud835\udc53\ud835\udc51\ud835\udc51, \ud835\udc53\ud835\udc53?\u0302?\ud835\udc51.\nWire tensions fa, fa\u0302, fc, fc\u0302, fb, fb\u0302, fd, fd\u0302. are determined according to 4 motors\u2019 angles \u03d5a, \u03d5b,\u03d5c,\u03d5d As\nfa=kp ( \u03bb (\u03d5p+\u03d5a) 2\u03c0 -nL+ \u2211|ai-ai-1| n\ni=1\n) ,\nfa\u0302=kp ( \u03bb (\u03d5p-\u03d5a) 2\u03c0 -nL+ \u2211|a\u0302i-a\u0302i-1| n\ni=1\n) ,\nfc=kp ( \u03bb (\u03d5p+\u03d5c) 2\u03c0 -nL+ \u2211|ci-ci-1| n\ni=1\n) ,\nfc\u0302=kp ( \u03bb (\u03d5p-\u03d5c) 2\u03c0 -nL+ \u2211|c\u0302i-c\u0302i-1| n\ni=1\n) ,\nfb=kp ( \u03bb (\u03d5p+\u03d5b) 2\u03c0 -nL+ \u2211|bi-bi-1| m\ni=1\n) ,\nfb\u0302=kp ( \u03bb (\u03d5p-\u03d5b) 2\u03c0 -nL+ \u2211|b\u0302i-b\u0302i-1| m\ni=1\n) ,\nfd=kp ( \u03bb (\u03d5p+\u03d5d) 2\u03c0 -nL+ \u2211|di-di-1| m\ni=1\n) ,\nfd\u0302=kp ( \u03bb (\u03d5p-\u03d5d) 2\u03c0 -nL+ \u2211|d\u0302i-d\u0302i-1| n\ni=1\n) ,\n(14)\nwhere, \ud835\udf19\ud835\udf19\ud835\udc5d\ud835\udc5d is a motor rotation angle to generate a pretension, \ud835\udf06\ud835\udf06 is a lead of the screw rod and \ud835\udc58\ud835\udc58\ud835\udc5d\ud835\udc5d is the spring constant of the pretension spring.\nAuthorized licensed use limited to: Auckland University of Technology. Downloaded on October 25,2020 at 08:37:21 UTC from IEEE Xplore. Restrictions apply.", + "222\nC. Inverse kinematic solution\nAccording to given set of variables \ud835\udf03\ud835\udf03\ud835\udc65\ud835\udc65,\ud835\udc56\ud835\udc56, \ud835\udf03\ud835\udf03\ud835\udc66\ud835\udc66,\ud835\udc56\ud835\udc56 , \ud835\udf03\ud835\udf03\ud835\udc67\ud835\udc67,\ud835\udc56\ud835\udc56 (\ud835\udc56\ud835\udc56 = 1, \u22ef , \ud835\udc5b\ud835\udc5b) and \ud835\udc59\ud835\udc59\ud835\udc56\ud835\udc56 (\ud835\udc56\ud835\udc56 = 1, \u22ef , \ud835\udc5b\ud835\udc5b \u2212 1), we calculate the end-point position by Eq. (6),\n(pn 1 ) =H0,n ( 0 0 ln 1 ) (in jn kn rn 0 0 0 1 ) ( 0 0 ln 1 ) = (knln+rn 1 )\n(15)\nTaking a total differentiation of pn=knln+rn with respect to \u03b8x,i, \u03b8y,i , \u03b8z,i (i= 1, \u22ef, n) and li (i=1, \u22ef, n-1) and also motor angles \u03d5a, \u03d5b, \u03d5c, \u03d5d,\n\u2206pn= \u2202pn \u2202v \u2206v+ \u2202pn \u2202\u03d5 \u2206\u03d5 (16)\nWhere, v=(\u03b8x1,\u03b8x2, \u22ef,\u03b8xn, \u03b8y1,\u03b8y2, \u22ef,\u03b8yn ,\u03b8z1,\u03b8z2, \u22ef,\u03b8zn, l1 , l2 , \u22ef,ln-1 )\n\u2208R4n-1 and \u03d5=(\u03d5a, \u03d5b, \u03d5c, \u03d5d ). \u2202pn \u2202v \u2208R3\u00d74n-1 and \u2202pn \u2202\u03d5 \u2208R3\u00d74. Whereas, let w=(w1, w2, \u22efw4n-1)T=04n-1 represents the 4n-1 equations provided by Eqs. (9)(10)(12)(13), which also includes \u03b8x,i, \u03b8y,i , \u03b8z,i (i=1, \u22ef, n) , li (i=1, \u22ef, n-1) and also motor angles \u03d5a, \u03d5b, \u03d5c, \u03d5d. Taking a total differentiation for w=04n-1 as well, we have,\n\u2206w= \u2202w \u2202v \u2206v+ \u2202w \u2202\u03d5 \u2206\u03d5=04n-1 (17)\nwhere, \u2202w \u2202v \u2208R(4n-1)\u00d7(4n-1) and \u2202w \u2202\u03d5 \u2208R(4n-1)\u00d74 . Since \u2202w \u2202v is a square matrix, we can solve (19) with respect to the vector \u2206\ud835\udc63\ud835\udc63 as,\n\u2206v=- (\u2202w\n\u2202v ) -1 \u2202w \u2202\u03d5 \u2206\u03d5 (18)\nSubstituting (18) into (16), we have\n\u2206pn=\u2202pn \u2202v (\u2202w \u2202v ) -1 \u2202w \u2202\u03d5 \u2206\u03d5+ \u2202pn \u2202\u03d5 \u2206\u03d5=\n( \u2202pn \u2202\u03d5 - \u2202pn \u2202v (\u2202w \u2202v ) -1 \u2202w \u2202\u03d5 )\u2206\u03d5=J\u2206\u03d5\n, which can be solved for \u2206\ud835\udf19\ud835\udf19, by using a generalized inverse of the Jacobian J\u2208R3\u00d74\n\u2206\u03d5=J\u2020\u2206pn+P\u22a5(J)\u03a8 (19)\nwhere, J\u2020\u2208R4\u00d73 is a generalized inverse of \ud835\udc71\ud835\udc71 and P\u22a5(J)\u2208R4\u00d74 is a null projection operator of J, and \u2206\u03d5N\u2208R4 is a correction of \ud835\udf53\ud835\udf53 so as to minimize a positive scalar potential \ud835\udf11\ud835\udf11 by making use of a redundant actuation.\nWe use J\u2020=JT(J JT)-1 and P\u22a5(J)=I-J\u2020J. Eq.(19) provides a variation of motor angles \u2206\ud835\udf53\ud835\udf53 for a given position and direction variation \u2206pn . Applying the Euler method, we have the following variational equation, \u03c6+(\u2202\u03c6/\u2202\u03d5)\u2206\u03d5N=0\n,which is solved by\n\u2206\u03d5N=\u03c6 (\u2202\u03c6/\u2202\u03d5)(\u2202\u03c6/\u2202\u03d5)T (\u2202\u03c6 \u2202\u03d5)\nT\n( 20) As a candidate of \u03c6 , we take \u03c6=knz\n2 , where, k\ud835\udc5b\ud835\udc5b\ud835\udc67\ud835\udc67 is the z component of kn: the unit vector of the end-point orienting an axial direction. It means that the axial direction the endpoint takes on a horizontal plain as far as possible while keeping a designated position.\nAuthorized licensed use limited to: Auckland University of Technology. Downloaded on October 25,2020 at 08:37:21 UTC from IEEE Xplore. Restrictions apply.", + "223\nIV. EXPERIMENTS AND SIMULATION TakoBot 2 kinematics computed by Wolfram Mathematica software. As a robot controller, Arduino UNO was utilized, and for stepping motors, we used SilentStick TMC2208 motor drivers to provide sufficient power to the motors. During the experiments, manipulator demonstrated basic motions and helical motion in both directions. According to the conducted experiments, passive sliding mechanism increased robot bending stress tolerance and acted smoother than the previous prototype. Likewise, the mechanism discovered new horizons for continuum robot applications and imagination about continuum manipulators.\nV. REFERENCES\n[1] A.Yeshmukhametov, K. Koganezawa, Y. Yamamoto, \u201cDesign and kinematics of cable driven continuum robot arm with universal joint backbone\u201d,IEEE ROBIO 2018 conference proceeding, Kuala \u2013Lumpur, Malaysia. 2018. [2] V.C Anderson, R.C. Horn: Tensor Arm Manipulator Design, Mech. Eng 8998), 54-65. 1967 [3] R. Buckingham, A. Graham, \u201cNuclear snake-arm robots\u201d, Industrial Robot: An International Journal, Vol.39 Issue: 1,pp 6-11, https://doi,org/10.1108/01439911211192448, 2012. [4] M.W. Hannan and I.D.Walker, \u201cKinematics and the\nImplementation of an Elephant\u2019s Trunk Manipulator and other Continuum Style Robots\u201d, Journal of Field Robotics, pp 45-63,February 2003. https://doi.org/10.1002/rob.10070\n[5] Graham, R., Bostelman, R., \u201cDevelopment of EMMA Prototype suspended from the 6 m RoboCrane Prototype.\u201d Proc. ANS Seventh Topical Meeting on Robotics and Remote Systems,\u201d Augusta, GA, April 27-May 1, 1997. [6] I.A. Gravagne ; C.D. Rahn ; I.D. Walker, \u201cLarge deflection dynamics and control for planar continuum robots\u201d, IEEE/ASME Transactions on Mechatronics, Volume: 8 , Issue: 2 , June 2003. [7] Camarillo, D., Milne, C., Carlson, C., Zinn, M., Salisbury, J.: Mechanics modeling of tendon driven continuum manipulators. IEEE Transactions on Robotics (accepted for publication) (2008) [8] S. Neppalli, B. Jones, W. McMahan, V. Chitrakaran, I. Walker M. Pritts, M. Csencsits, C. Rahn, M. Grissom, \u201cOctArm - A Soft Robotic Manipulator\u201d, Proceedings of the 2007 IEEE/RSJ International Conference on Intelligent Robots and Systems San Diego, CA, USA, Oct 29 - Nov 2, 2007 [9] Jones, B., Walker, I.: Kinematics for multisection continuum robots. IEEE Transactions on Robotics 22(1), 43\u201355 (2006) [10] Haitham E., Usman M., Zain M.,Sang-Goo J., Muhammad U., Elliot W.H., Alisson M.O.,Jee-Hwan R., \u201cDevelopment and Evaluation of an Intuitive Flexible Interface for Teleoperating Soft Growing robots\u201d, 2018 IEEE/RSJ IROS, Madrid,Spain, October 1-5,2018. [11] A. Benouhiba, K. Rabenorosoa, P. Rougeot, M. Ouisse, N. Andreff, \u201c A multisection Electro-active Polymer based milli-Continuum Soft Robots. 2018 IEEE/RSJ IROS, Madrid,Spain, October 1-5,2018. [12] B. Ouyang, H. Mo, H. Chen, Y. Liu, D. Sun, \u201cRobust Model-Predictive Deformation Control of a Soft Object by Using a Flexible Continuum Robot, 2018 IEEE/RSJ IROS, Madrid,Spain, October 1-5,2018. [13] R. Kang, D.T. Branson, T. Zheng, E. Guglielmino, D.G. Caldwell, \u201cDesign, modelling and control of a pneumatically actuated manipulator inspired by biological continuum structures\u201d, Bioinspiration&Biomimetics 8 (2013) 14pp. [14] Bryan A. Jones, Ian D. Walker, \u201cA New Approach to Jacobian Formulation for a Class of Multi-Section Continuum Robots\u201d, Proceedings of the 2005 IEEE, ICRA, Barcelona, Spain, April 2005. [15] Han Yuan and Zheng Li, \u201cWorkspace analysis of cabledriven continuum manipulators based on static model\u201d, Robotics and Computer \u2013 Integrated Manufacturing. 49 (2018) 240-252. [16] X. Dong, M. Raffles, S. Gobos-Gusman, D. Axiente, J. Kell, \u201c A Novel Continuum Robot Using Twin-Pivot Compliant Joints: Design, Modelling and Validation, Journal of mechanism and robotics, February 16, 2015.\n[17] T. Liu, Z. Mu, H. Wang, W. Xu, Y. Li, \u201c A Cable-Driven Redundant Spatial Manipulator with Improved Stiffness and Load Capacity\u201d, 2018 IEEE/RSJ IROS, Madrid, Spain, October 1-5,2018\n[18] J. Starke, E. Amanov, M.Taha Chihaoui, J. Burgner-Kahrs, \u201cOn the Merits of Helical Tendon Routing in Continuum Robots\u201d, 2018 IEEE/RSJ IROS, Vancouver, BC, Canada, September 24-28, 2017.\nAuthorized licensed use limited to: Auckland University of Technology. Downloaded on October 25,2020 at 08:37:21 UTC from IEEE Xplore. Restrictions apply." + ] + }, + { + "image_filename": "designv11_71_0003880_scis-isis.2014.7044812-Figure3-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0003880_scis-isis.2014.7044812-Figure3-1.png", + "caption": "Fig. 3. Overview phenomena of the proposed APF and force distribution", + "texts": [ + " As a result of the cancelling out of the attractive force with the repulsive force, dead-lock takes place. To cope with this issue, APF based motion planning algorithm is modified by taking into account the front-face obstacle-velocity information of the robot. The new information added to the algorithm generates an additional controlling force to the motion planning algorithm which is called obstacle-velocity repulsive potential force, . The basic overview of the proposed phenomena is explained in the Fig.3. The obstacle-velocity repulsive force has a direct relationship to the angle between line of the robot to obstacle and the velocity vector, besides the distance to the obstacles. This is acting perpendicular to the original repulsive force and its magnitude varies with the angle of . When the robot detects an obstacle within its sensory range, primary repulsive force acts to avoid the collision with the detected obstacle. Secondary repulsive force component behaves to bring the robot away from its original path directing to the obstacle but towards the goal" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0000925_9781119477891.ch8-Figure8.14-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0000925_9781119477891.ch8-Figure8.14-1.png", + "caption": "Figure 8.14 Blown film extrusion line.", + "texts": [ + " For the manu facturing of these type of films, the three main processes that can be used are blown film extrusion, cast film extru sion, and extrusion lamination or extrusion coating. Blown film extrusion is the process where the molten polymer that is coming out of the extruder is pushed into a die which as a thin annular opening at the top. Through this opening, the material is pushed out forming a tube of molten material. This tube is pinched off at the top of the equipment, by nip rolls, to form a closed bubble (see Figure\u00a08.14). Air will then be blown into the bubble to inflate it to the required diameter. This diameter will determine the final width of the extruded film. The ratio between the bubble diameter and the die diameter is called the blow\u2010 up ratio. When the molten material is coming out of the die it needs to be cooled down to take its rigid form as film. This is done by blowing cooled air against the molten material through an air ring that is sitting around the bubble on top of the die (see Figure\u00a08.15). At the nip rolls, the bubble is, through the web takeoff unit, collapsed into a flat tube, which is going finally to the winding station, where it is winded into a roll of film" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0003492_0309324715598916-Figure3-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0003492_0309324715598916-Figure3-1.png", + "caption": "Figure 3. Overview of the test rig with indication of the applied measurement techniques (in italics).", + "texts": [ + " In contrast, this taper mismatch will most likely alter the available thread clearance near the pin tip, having an effect on the sealing capabilities. Since this mismatch induces higher at University College London on June 5, 2016sdj.sagepub.comDownloaded from levels of deformation on the pin tip, the connection is likely to have a sealing capacity exceeding the commonly accepted 27.5MPa. After initial make-up, the thread compound was dried out before applying internal pressure and axial tension. A schematic overview of the setup for the TLE and failure tensile tests is illustrated in Figure 3. Both tests were performed on a 1000 kN universal tensile test rig. In addition to the ability to apply tensile loads, an external pump was used to control the internal pressure of the connection, allowing pressures of up to 60MPa using water as a pressurizing fluid. Strain measurement. With the intention of performing a varied evaluation of the applied numerical model, various measurement techniques were used. Strains were measured at the outside of the box using two different methods which are illustrated in Figure 3. At first, 12 biaxial strain gages (type FCA-3-11) are distributed over three different sections (A, B and C) and are spaced 90 apart. The location of the sections approximates the start of the thread, the transition between complete and vanishing threads and the end of the thread. In addition, the optical measuring technique DIC is applied at one side. Accurate DIC calculations require the application of a non-uniform high-contrast speckle pattern on the specimen\u2019s area of interest. This pattern was achieved by projecting black spray droplets upon a uniform, dry white bottom layer of paint", + " In addition to the TLE test, a tensile test up to failure was conducted to examine the ability of the numerical model to predict the location of failure when loads exceeding the yield strength of the connections are applied. Because the amount of heat generated is proven to be dependent on the applied strain rate,19 a constant axial displacement rate of 0.02mm/s without internal pressure was applied until fracture. As mentioned earlier, strains were measured using both strain gages and DIC. The experimentally obtained at University College London on June 5, 2016sdj.sagepub.comDownloaded from axial strains in section C (see Figure 3) are given in Figure 5 together with the numerically predicted strain values. The plotted axial strains using DIC were calculated by averaging 24 values evenly distributed over the visible surface within the applicable section. For the strain gages, a similar approach using all four measurements distributed over the circumference was applied. The black dots represent the load combinations that were effectively measured and/or calculated. The other combinations are estimated by means of interpolation using a thin plate spline smoothing algorithm" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0000742_ssci44817.2019.9003002-Figure8-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0000742_ssci44817.2019.9003002-Figure8-1.png", + "caption": "Fig. 8. Miscellaneous contact detectors. (a) Iron\u2013mercury autocoherer reported in 1899 by Bose and rediscovered in 1901 by Castelli. (b) Galena (lead sulfide) detector patented in 1901 by Bose. (c) Silicon detector, patented in 1906 by Pickard. (d) Carborundum (carbid of silicon) detector patented in 1906 by Dunwwody. (e) The carborundum detector was the only contact reliable enough to be packaged in a cartridge. (d) Perikon detector developed by Pickard in 1907 and used in the U.S. Navy. (Sources: http://earlyradiohistory.us/ and http:// www.clarkmasts.net.au/.)", + "texts": [ + " 1) Cat Whisker, Point-Contact Rectifiers: When the Branly\u2019s coherer started to be used commercially, nonlinear resistances, provoked by two different materials in contact, had already been reported for a while. The coherer had been exactly what was necessary to stimulate research on an RF detector: 1) sensitive enough to be used commercially as well as to show the potential applications of radiowaves to a grand public; and 2) so imperfect that significant research was necessary to find a better detector. In search for viable RF rectifier, researchers started to tryand-error different material combinations. In less than a decade, the main rectifier devices were found (Fig. 8). To achieve a good contact between two solids, they soon realized that much stain should be applied to the contact so not to leave any \u2018\u2018gap\u2019\u2019 between materials at the atomic level. At a given force, the pressure could be increased by reducing the point contact area. Therefore, most of the designs moved toward a very thin metal wire (\u2018\u2018cat whisker\u2019\u2019) pushed by a spring upon the crystal surface. But the contact was built from raw crystal materials, which had a nonuniform surface property. The devices were thus usually fabricated with some mechanical degree of liberty: the operator would modify (more or less by luck) the arrangement to find a suitable spot for rectification", + " Actually, the first metal\u2013semiconductor contact-based rectifier did not use the cat whisker but a liquid instead. This device, the iron\u2013mercury \u2018\u2018autocoherer\u2019\u2019 (also called \u2018\u2018Italian navy coherer\u2019\u2019), has a place of choice in the textbooks of history because it was used in 1901 by Marconi for the first transatlantic communication. Although Marconi claimed its invention in 1901, it soon appeared that he had taken the work of somebody else. Paolo Castelli was rapidly acknowledged for the invention [64] [Fig. 8(a)], but it was eventually found that Jagadish Chandra Bose [65] was the first to report results on this detector in 1899. 1674 Proceedings of the IEEE | Vol. 102, No. 11, November 2014 Detecting signal was possible due to the formation of an oxide film on the interface of mercury and iron [66]. Since no current\u2013voltage data have been found by the authors, this device is not collected in Fig. 5. The mechanical stability of the oxide layer proved to be quite unreliable, and Bose directed his researches toward Cat Whisker rectifier based on lead sulfide (Galena). This device was the first metal\u2013semiconductor solid-state rectifier, and was patented in 1901 [Fig. 8(b)]. It was an inexpensive detector, easy to build from its early days, and became widely used, especially among amateurs. This detector was very sensitive but the contact became so unreliable that the operator had to constantly search for a \u2018\u2018spot.\u2019\u2019 Five years later, Greenleaf W. Pickard proposed to use silicon as crystal [Fig. 8(c)], for yielding better reliability. High reliability was accomplished during that same year (although at the expense of sensitivity) in 1906 when Henry H. C. Dunwwody made use of carborundum [Fig. 8(d)], a manmade material (silicon carbide) that was used at that time as a polishing abrasive. The great advantage of the carborundum material was its density. It was so hard that the cat whisker could be firmly pressed against it, thus resulting in a very repeatable point contact [67]. This detector almost did not necessitate a new arrangement of the point contact, and could be packaged [Fig. 8(e)]. The U.S. Navy was prone to the Perikon detector (\u2018\u2018PERfect pIcKard cONtact\u2019\u2019). In the configuration shown [Fig. 8(f)], multiple zincite crystals were provided because of their fragility. In 1909, George Pierce found that the molybdenite\u2013copper contact [68] was very sensitive, but still not reliable. The reliability was then found in vacuum tubes, and commercial interest in crystal rectifier progressively vanished. Responsivity of the devices is shown in Fig. 5 in a chronological perspective. Data are derived from a comprehensive comparison of I\u00f0V\u00de characteristic curves reported by Coursey [69]. The Perikon has been chosen to start the line of point-contact metal\u2013semiconductor diodes because it had the best sensitivity\u2013reliability tradeoff" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0000925_9781119477891.ch8-Figure8.15-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0000925_9781119477891.ch8-Figure8.15-1.png", + "caption": "Figure 8.15 Air cooling of the bubble with blown film extrusion.", + "texts": [ + " This tube is pinched off at the top of the equipment, by nip rolls, to form a closed bubble (see Figure\u00a08.14). Air will then be blown into the bubble to inflate it to the required diameter. This diameter will determine the final width of the extruded film. The ratio between the bubble diameter and the die diameter is called the blow\u2010 up ratio. When the molten material is coming out of the die it needs to be cooled down to take its rigid form as film. This is done by blowing cooled air against the molten material through an air ring that is sitting around the bubble on top of the die (see Figure\u00a08.15). At the nip rolls, the bubble is, through the web takeoff unit, collapsed into a flat tube, which is going finally to the winding station, where it is winded into a roll of film. The setup as it is shown here, is a mono\u2010extrusion instal lation. This means that we have one extruder that will feed one material or one blend of materials to the die, generating a mono\u2010layer film. Depending on the final film structure that we want, a co\u2010 extrusion installation can be used. In a co\u2010extrusion instal lation, a multi\u2010layer film can be produced by having multiple extruders feeding molten material to the die" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0000253_2019045-Figure5-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0000253_2019045-Figure5-1.png", + "caption": "Fig. 5. Helix diagram of right tooth surface.", + "texts": [], + "surrounding_texts": [ + "The tooth surfaces of double helical gear is involute helicoid. Figures 4 and 5 are its left and right tooth surface helix diagram, respectively. The left tooth surface equation is expressed as xk1 \u00bc rb1 sinmk1 rb1mk1 cosmk1 \u00fe rk1 sin uk1 yk1 \u00bc rb1 cosmk1 \u00fe rb1mk1 sinmk1 \u00fe rk1 cos uk1 rb1 zk1 \u00bc rk1uk1 tan gk1 8< : \u00f04\u00de The right tooth surface equation is expressed as xk2 \u00bc rb2 sinmk2 rb2mk2 cosmk2 \u00fe rk2 sin uk2 yk2 \u00bc rb2 cosmk2 \u00fe rb2mk2 sinmk2 \u00fe rk2 cos uk2 rb2 zk2 \u00bc rk2uk2 tan gk2 : 8< : \u00f05\u00de Equations (4) and (5) are all the equation of involute helicoid. Where, rbi (i=1, 2) is the radius of base circle, mki (i=1, 2) is the roll angle of point k in the involute, rki (i=1,2) is the radiusofpoint k in the involute, uki (i=1,2) is theexpansionangleofpoint k in the involute,gki (i=1,2) is the helix angle of ascent of point k in the involute, bki (i=1, 2) is the helix angle of point k in the involute." + ] + }, + { + "image_filename": "designv11_71_0002683_012034-Figure14-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0002683_012034-Figure14-1.png", + "caption": "Figure 14. Shapes of (a) basic cell of specimen (b) tension specimen (c) compression specimen.", + "texts": [ + " We decided to test both tension and compression as the lattice structure behavior may vary depending on the sign of the applied load [15]. As it was mentioned before, uniaxial tension and compression tests were carried out. Since structures have complex geometry, it is not possible to mount them in the test machine directly without any additional gear. So the specimens were adapted to the test machine by adding special blocks for installing. The shapes of the specimens and a basic cell are presented in Fig. 14. IOP Conf. Series: Materials Science and Engineering 986 (2020) 012034 IOP Publishing doi:10.1088/1757-899X/986/1/012034 Tests were carried out with Instron 8850 Axial-Torsion System. The photographs of the specimens during tests are shown in Fig. 15. FE modeling of uniaxial tension and compression was performed in SIMULIA Abaqus implicit solver with the same finite element models both for tension and compression. Constructed model is shown in Fig. 16. IOP Conf. Series: Materials Science and Engineering 986 (2020) 012034 IOP Publishing doi:10" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0002989_978-3-642-16214-5_9-Figure9.26-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0002989_978-3-642-16214-5_9-Figure9.26-1.png", + "caption": "Fig. 9.26. Locking-pin synchronizer. 1 Idler gear with dog gearing; 2 synchronizer ring; 3 gearshift sleeve; 4 locking pin; 5 compression spring", + "texts": [ + "21, the opening angle \u03b2 can be reduced relative to the Borg-Warner system, with the same safety factor S; shifting comfort is improved. The friction surfaces are located outwards, as compared to synchronizers based on the Borg-Warner system. In accordance with Equation 9.2, this arrangement results in reduced gearshift effort, and in lower specific stresses because of the increased friction surface area AR. Because of the larger effective diameter d, the friction speed v increases in accordance with Equation 9.12, and the synchronizable speed difference falls. The Spicer or Tompson synchronizer shown in Figure 9.26 is a locking-pin synchronizer. The gearshift sleeve 3 has six drill holes and is rotationally fixed to the transmission shaft, but axially linked by a sliding connection. The locking pins 4 engage in the drill holes parallel to the axis. They are each rigidly connected to a synchronizer ring 2. The conically countersunk drill holes are larger than the conical part of the locking pin, enabling the synchronizer pin to turn by a certain amount. As long as there is a speed difference, the opening torque at the conical surfaces of the locking pins and the drill holes is smaller than the locking friction torque" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0001019_j.precisioneng.2020.04.003-Figure4-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0001019_j.precisioneng.2020.04.003-Figure4-1.png", + "caption": "Fig. 4. Two 25.4-mm [1-in] diameter cylinders forming a vee can be set at inclination angles from 25\ufffd (top) to 80\ufffd (bottom). They are guided by circular slots in the plastic plate.", + "texts": [ + " It is worth noting to the potential designer of an adjustable mount that the direction of actuation may be chosen, for example, along the centerline of the vee, to make friction on that surface favorable to motion and unlikely for the mechanism to lock. That is, the direction of friction on the moving surface flips direction, tending to cancel friction on the nonmoving surface. The 3-V kinematic coupling designed for this experiment has adjustable inclination angles which cover a broad practical range from 25\ufffd to 80\ufffd as Fig. 4 shows. Each vee is formed by two cylinders secured with threaded fasteners to a plastic plate, which allows an electrical conductivity test to confirm contact (although it was not used nor very reliable using DC). Five triangles having different angles were machined from plastic sheet and used to set the inclination angle \u03b1 of each vee in 5\ufffd increments, as Fig. 5 demonstrates. Referring to Fig. 6, cylinder pairs that form vees are arranged with their axes in three vertical planes that are tangent to and equally spaced on a 203" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0003721_1.4879277-Figure2-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0003721_1.4879277-Figure2-1.png", + "caption": "Fig. 2. Improved demodulator preserves good autocorrelation properties down to one cycle per code bit. The circles with crosses through them are called mixers and are mathematically equivalent to multiplication (and not to be confused with the exclusive-or operators).", + "texts": [ + " 1(b), (i) the weight force ~W ; (ii) the upward buoyant force ~Fb due to the vertical pressure gradient, (iii) the force ~FH responsible for the accelerated motion of that portion of matter due to a horizontal gradient of pressure, and (iv) the fictitious inertial force ~Fi that results from being in a non-inertial reference frame and is directed opposite to the acceleration vector. All of these forces must be in dynamical equilibrium in order for the volume V to remain at rest relative to the rest of the fluid. We now replace the fluid in that region with an object of the same volume and whose mass mo is smaller than mf and attach it to the bottom of the container with a string (for reasons that will soon become clear), as shown in Fig. 2(a). The forces that are applied by the surrounding fluid\u2014the buoyancy and the horizontal pressure-gradient forces\u2014will remain the same, but both the weight and the inertial forces will change due to the change in the object\u2019s mass. An additional force due to the string will now also be present. We thus end up with the situation depicted in Fig. 2(b), in which the object can be thought of as having an effective (negative) mass given by [see Fig. 2(c)] meff \u00bc \u00f0mf mo\u00de: (1) By introducing such an effective mass, the problem simplifies considerably; instead of having to deal with four forces we only have to deal with two effective forces. The effective weight of the object ~W eff is given by the effective mass times the gravitational field strength, and in this case it will be negative (directed upward). Meanwhile, the effective mass times the acceleration gives an effective inertial force Fi;eff that will now have the same direction as the acceleration", + " For a dynamical equilibrium, these two forces plus the force applied by the string must sum to zero\u2014the object will 997 Am. J. Phys. 82 (10), October 2014 http://aapt.org/ajp VC 2014 American Association of Physics Teachers 997 This article is copyrighted as indicated in the article. Reuse of AAPT content is subject to the terms at: http://scitation.aip.org/termsconditions. Downloaded to IP: 129.89.24.43 On: Sun, 05 Oct 2014 09:30:52 thus tilt towards the acceleration. In the (normal) situation of an object with a positive effective mass [see Fig. 2(d)], the two effective forces change their sign and point in the opposite direction as in Fig. 2(c). Thus, a pendulum bob and a helium balloon inside an accelerating container of air will have opposite behaviors; the former will lean backward while the latter will lean forward. In either case (positive or negative effective mass), the tilt angle is given by tan h\u00f0 \u00de \u00bc Fi;eff Weff \u00bc a g : (2) We note that the dimensions of the container do not come into play in this discussion. Of course, the container will be important if, in the process of tilting, the object impacts the walls of the container", + " The correlations are most efficiently computed using a Fast Fourier Transform (FFT) and take the form R ref; data\u00f0 \u00de \u00bc 1 N XN 1 m\u00bc0 ref m\u00f0 \u00de data m\u00fe n\u00f0 \u00de \u00bc DFT 1 DFT ref\u00f0 \u00deDFT data\u00f0 \u00de\u00f0 \u00de; (2) where the bars denote the complex conjugate, N is the length of oversampled ML-sequence, \u201cref\u201d is the reference PN code expressed as 2Z-1, \u201cdata\u201d is the received signal, and \u201cDFT\u201d is the digital Fourier transform (preferably computed as an FFT). Note that here we are using the normalized version of the correlation function, which is somewhat different than our original implementation. The first improvement is the quadrature correlator demodulator,6 as shown schematically in Fig. 2. The main advantage to using this demodulator is that it has the capability of demodulating the signal down to 1 cycle/code-bit while preserving good autocorrelation properties\u2014it produces a nice clean pulse without side-lobes or other artifacts. As seen in Fig. 2, this demodulator is very similar to a quadrature demodulator except that a correlator has replaced the lowpass-filter section;7 it is also one of the standard demodulators used in communications8 for demodulating binary PSK (BPSK) signals. In this demodulator, the reference sine and cosine signals have the same frequency as the carrier and are stored in memory at the same length as the frame size of the modulation; they need not be of any particular phase. As in the original implementation, the reference PN code does need to have the same phase as the outgoing signal", + "3,6 This type of signal uses a two-state phase of 0; p\u00f0 \u00de. Creating a BPSK signal is very simple and takes the form 1001 Am. J. Phys., Vol. 82, No. 10, October 2014 Notes and Discussions 1001 This article is copyrighted as indicated in the article. Reuse of AAPT content is subject to the terms at: http://scitation.aip.org/termsconditions. Downloaded to IP: 129.89.24.43 On: Sun, 05 Oct 2014 09:30:52 n n\u00f0 \u00de \u00bc 2Z n\u00f0 \u00de 1\u00bd sin 2p n 1=2\u00f0 \u00deP M : (3) As stated earlier, BPSK modulation results in better SNR compared to AM. Using the demodulator in Fig. 2 and the same reference kernel 2Z n\u00f0 \u00de 1, the correlation peak will have twice the height and therefore twice the SNR. A potential alternative to BPSK would be linear or non-linear swept frequency,9 which is also a standard modulation technique used in both radar and sonar. Let us take an example where we have a sound card with a sample rate of 44.1 kHz. With M\u00bc 16 and P\u00bc 4, the carrier frequency is 4=16\u00f0 \u00de 44:1 kHz\u00bc 11.025 kHz. The code bitrate is 44.1/16\u00bc 2.756 kbps. With a speed of sound of c\u00bc 340.29 m/s, the resolution is c=2br \u00bc 340:29= 2 2756\u00f0 \u00de m \u00bc 6:17 cm, where br is the bitrate. The unambiguous range, defined as the maximum measurable distance before wraparound, is N c=2br \u00bc 255 340:29= 2 2050\u00f0 \u00de m\u00bc 15:74 m. Figure 3 shows a theoretical plot of the resulting autocorrelation function for both BPSK and AM modulation using the quadrature correlator shown in Fig. 2. As can be seen from Fig. 3, the amplitude of the resulting synthetic pulse is greater by a factor of two using PSK modulation, and therefore has twice the SNR. A mathematical proof of this can be found elsewhere.6 If more range is needed, one can potentially use a longer code. A table for generating these is shown in Table I. a)Electronic mail: joel.f.campbell@nasa.gov 1J. F. Campbell, M. A. Flood, N. S. Prasad, and W. D. Hodson, \u201cA low cost remote sensing system using PC and stereo equipment,\u201d Am" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0003347_9781119016854.ch7-Figure15-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0003347_9781119016854.ch7-Figure15-1.png", + "caption": "Figure 15: Prediction of bulk residual stresses as a result of electron beam (EB) welding of Alloy 718 ring components.", + "texts": [ + " These have been applied to the heat treatment of Alloy 718 components to support subsequent component machining and application performance prediction and optimization. [23] A significant program has been initiated to establish standards for the application of bulk residual stress modeling using Alloy 718 as the demonstration material. [10] This effort involves establishment of standard protocols for sharing of requirements, predictions and uncertainty quantification so bulk residual stress can be effectively incorporated into the design function of new components. Residual stresses can also be of concern and issue during and after welding processes. Figure 15 shows the principle stresses developed in the electron beam welding of two rings of Alloy 718 material. Designing with or mitigating the residual stresses during the manufacture of fabrications is very important is becoming a common element of integrated computational engineering. Advanced material and process modeling applications: As Alloy 718 has evolved, myriad manufacturing methods have also evolved to exploit this material and related alloys to their fullest. Modeling tools and methods are supporting this advancement such as the application of DA718 material" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0003808_s00542-015-2524-5-Figure2-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0003808_s00542-015-2524-5-Figure2-1.png", + "caption": "Fig. 2 Coordinate system of the journal bearing with stationary herringbone grooves", + "texts": [ + " The FDBs are composed of lower and upper grooved journal bearings and lower and upper grooved thrust bearings, in which the grooves are inscribed in a stationary sleeve. The governing equations of the stationary grooved journal bearing and the thrust bearing can be written in a coordinate system fixed to the sleeve as shown in Eqs. (1) and (2), respectively. where R, \u03b8\u0307, h, p and \u00b5 are the radius, rotating speed, film thickness, pressure and coefficient of viscosity, respectively. And r, \u03b8 and z are the radial, circumferential and axial coordinates, respectively. Figure 2 shows the coordinate system of the journal bearing with stationary herringbone grooves. Perturbation equations are obtained by substituting a first-order Taylor series of pressure, film thickness and rate change of film thickness with respect to small displacement and velocities into the Reynolds equation, and by applying the separation of variables with respect to each perturbed displacement or velocity. The perturbation equations are solved by using the finite element method and the dynamic coefficients can be obtained by integrating the perturbed pressure in bearing regions (Jang and Kim 1999)" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0002442_icarcv50220.2020.9305376-Figure3-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0002442_icarcv50220.2020.9305376-Figure3-1.png", + "caption": "Fig. 3: Several examples of printhead movements paths.", + "texts": [ + " It then progressively adds 855 Authorized licensed use limited to: Cornell University Library. Downloaded on May 24,2021 at 13:12:17 UTC from IEEE Xplore. Restrictions apply. 1.5 layer-upon-layer and expands the print radially outwards. Each layer has multiple trajectories and printing starts from the bottom and ends at the top. While printing a single layer, the robot arm gradually moves along a vertical rail from bottom to top in a vertical line and the central column rotates to achieve the helical shape. Fig. 3 shows a topdown view of the printer and the dashed lines shows these concentric vertical layers or cylinders. The problem that this paper addresses is discovering the path that should be taken when moving outwards from the central column. Since the location of the robot arm in the print space changes in order to cover a large print volume with a small robot arm, then the potential dexterity of the robot arm is also constantly changing. This ultimately affects the quality of the print and can cause major failures in the print. Fig. 3 shows a few example paths printhead can move in each layer. Theoretically, there are a massive number of possible paths that can be generated, some of which are undesirable or even infeasible, and this paper addresses the problem of finding the optimal path. 1) Manipulability: The ability of an industrial robotic arm to move easily in any arbitrary direction is called its manipulability [9]. Yoshikawa\u2019s manipulability measure describes an arm\u2019s distance to a singular configuration [26]. This approach is based on an analyses of the velocity ellipsoid that is spanned by the singular vectors of the Jacobian [26]" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0001701_0954405420949757-Figure8-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0001701_0954405420949757-Figure8-1.png", + "caption": "Figure 8. Extended upper and lower edges.", + "texts": [ + "01mm, the value of f is calculated, calculating process is ended if f\\ 0:05mm, then the current coordinates of A0i and S0i are obtained. By the same method, coordinates of A00i and S00i which ensure chamfering width of concave side can also be obtained. In order to ensure integrity of tooth crest chamfering, edges are extended over inner end and outer end along tooth width direction, it is realized by extending tooth width at inner and outer end when mathematical model of tooth surface is established. Extended upper and lower edges are shown in Figure 8. Chamfering tool path for both sides chamfering phase Due to the spatial asymmetry of convex and concave side, when cutter moves from inner end to outer end, chamfering process is divided into three phases in the sequence of concave side, both sides and convex side. In order to process tooth crest chamfering surface accurately, cutter surface should contact with specific edges in different phase of chamfering process. In concave side chamfering phase, cutter surface contacts with upper edge and lower edge of concave side" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0002059_ecce44975.2020.9236256-Figure1-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0002059_ecce44975.2020.9236256-Figure1-1.png", + "caption": "Fig. 1. Different HESFM topologies. (a) HE1 [17]. (b) HE2 [20]. (c) HE3 [19]. (d) HE4 [21]. (e) HE5 [22].", + "texts": [ + " Moreover, concentrated non-overlapping DC field coils are adopted in these machines for short end-windings, good faulttolerant capability and compact volume. For each phase of armature windings, different armature coils are seriesconnected. Similarly, different DC field coils are seriesconnected as well in the field winding of each HESFM. The field windings of HE1, HE3 and HE5 contain 12 DC field coils (f1, f2, \u2026, f12), while the field windings of HE2 and HE4 contain 6 DC field coils (f1, f2, \u2026, f6), as shown in Fig. 1. Basic geometric parameters of each HESFM can be found in Table I. As a matter of fact, the operation principles of these HESFMs are very similar. The air-gap field is excited by the PMs and DC field coils simultaneously. The PM excited airgap field is fixed while the DC excited one is variable. The direction and strength of the DC excited field, and consequently, the resultant air-gap field and the output torque, can be easily adjusted by controlling the polarity and magnitude of the current in DC field coils. In the HESFMs in Fig. 1, the PM flux and DC excitation flux have different trajectories, i.e. the DC flux does not go through the PMs. Therefore, they exhibit good flux-controllability, since the reluctance of the DC flux path is low. In addition, the PMs do not suffer from the risk of irreversible demagnetization, as the DC flux does not pass them. III. VOLTAGE PULSATION INDUCED IN DC FIELD WINDING In HEMs, a constant DC excitation current is preferred in DC field winding under steady state from the perspective of both machine design and drive", + " Downloaded on June 14,2021 at 22:41:20 UTC from IEEE Xplore. Restrictions apply. ratios of the HESFMs from HE1 to HE5 are 3.4%, 153.0%, 7.3%, 340.7%, and 13.9%, respectively. The HESFM with Ecore stator (HE4) has the highest no-load EIV and VPR among those HESFMs. VPR= EIV VDC \u00d7100% (1) In order to explain the cause of no-load induced voltage pulsation in the DC field winding of the HESFMs, the HESFM with E-core stator (HE4) is taken as an instance. HE4 has 6 field coils in the DC winding as illustrated in Fig. 1(d). As shown in Fig. 4, the flux-linkages of 6 DC field coils are variable versus the electric position (\u03b8e) of rotor and they contain both DC components and other harmonics. By way of example, the minimum value of the flux-linkage of DC field coil f1 (\u03a61) occurs at \u03b8e=156o, while the maximum value of \u03a61 occurs at \u03b8e=336o. By referring to the flux distributions illustrated in Fig. 5(a), it can be observed that the maximum value of the reluctance in the magnetic circuit of DC field coil f1 (Rm1) occurs at \u03b8e=156o, since the middle tooth of the E-core stator that accommodates f1 does not align with a rotor tooth but align with the middle of a rotor slot and hence very few fluxes pass through it", + " 6, and likewise, they possess the same voltage spectrum and the phase displacement between two adjoining pulsating voltage waveforms is 120o. At no-load condition, the phase displacement for the HESFMs from HE1 to HE5 is 60o, 120o, 60o, 120o, and 60o, respectively. For the voltage pulsation in the DC field winding of HE4, it reveals that the harmonic orders of 3M (M is a positive integer) in different DC field coil flux-linkages shown in Fig. 4(a) cannot be counteracted in the DC field winding as shown in Fig. 2(b) and so do the voltage pulsations illustrated in Fig. 3(b). For the HESFMs in Fig. 1, the harmonic orders (kM) of the flux-linkage and induced voltage pulsation in each DC field winding are different at no-load, i.e. 6M, 3M, 6M, 3M, and 6M for the HESFMs from HE1 to HE5, respectively, which can be described as (2), where LCM is the least common multiple. The number of stator/rotor pole (Ns/Nr) for the HESFMs from HE1 to HE5 is 12/10, 6/10, 12/10, 6/10, and 12/10, respectively. Therefore, it can be concluded that the no-load DC field winding induced voltage pulsations in different HESFMs are resulted from the variation of the permeance in the air-gap with the change of the rotor position, owing to the utilization of doubly salient machine structures in all of these HESFMs" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0001576_aim43001.2020.9158924-Figure1-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0001576_aim43001.2020.9158924-Figure1-1.png", + "caption": "Fig. 1: A MAGLEV platform enabling robot-based high precision in-line measurements on free form surfaces. a) When the robot is repositioning and approaching the sample surface section to be inspected, the platform needs to be stabilized in a free-floating position with respect to its supporting frame using the internal position sensors (IPS). b) To actively isolate both the measurement tool and the sample, the platform is required to maintain a constant distance and orientation relative to the sample surface by use of the tracking sensors (TS).", + "texts": [ + " In Section III the prototype system is presented. Based on an identification of the system dynamics, the internal and external position control is designed in Section IV The proposed transition scheme is described in Section V, while the system performance is evaluated in Section VI. Finally, Section VII concludes the paper. The system concept of a MAGLEV platform enabling robot-based inline measurements on free forms comprises two different sensor systems for two different control tasks and is illustrated in Fig. 1. Initially, the MAGLEV platform needs to be stabilized in a desired free-floating position with respect to the supporting frame (internal position control), when the robot arm is repositioning and approaching a sample surface (Fig. 1a). Once the external tracking sensors (TS) are in range, disturbing environmental vibrations are actively compensated during a surface measurement by tracking the sample surface and maintaining a constant distance and orientation between the measuring tool and the sample (external position control), as depicted in Fig. 1b. Therefore, the requirement of an efficient handing over between the internal and external position control task in Fig. 1 arises. However, two major aspects need to be considered. On the one hand, during the transition from internal to external position control and vice versa, unintended transients in the platform position need 978-1-7281-6794-7/20/$31.00 \u00a92020 IEEE 1943 Authorized licensed use limited to: Cornell University Library. Downloaded on August 30,2020 at 06:16:05 UTC from IEEE Xplore. Restrictions apply. to be avoided to prevent a sensitive measurement tool from damage. On the other hand, a fast transition is desired for reasons of efficiency, keeping the overall measurement time at the surface section of interest as low as possible", + " Each of the three external position controller is designed for a closed-loop control bandwidth of about 130Hz. Now, the system prototype presented in Fig. 2 is capable of tracking a sample surface in the three out-of-plane DoFs z, \u03c6x and \u03c6y , while the platform can be stabilized inplane (x, y and \u03c6z). Considering the desired system concept with the MAGLEV platform carrying a measurement tool and mounted as an end effector to an industrial robot arm, an efficient transition between the internal and external position control is required, once the TS are in range (see Fig. 1). This is done by inverting the sign of the external position signal and using the same PID parameters for control of the internal and external out-of-plane position pairs {zi, ze}, {\u03c6x,i, \u03c6x,e} and {\u03c6y,i, \u03c6y,e} (see Section IV). The transition is implemented of a time-dependent cross-fading error gain \u03b3 \u2208 [0, 1], as exemplary shown for one DoF in Fig. 4. Let \u03b3 be defined as \u03b3 = 0 if t \u2264 0 1 if t \u2265 T \u03b3(t) otherwise, (3) with T being the transition time between internal and external position control" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0001380_s1068798x20060076-Figure3-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0001380_s1068798x20060076-Figure3-1.png", + "caption": "Fig. 3. Saw module of M2005 multisaw system: (1) saw blades; (2) upper hinge; (3) elastic elements; (4) lower hinge with corrective mass; (5) upper shaft; (6) lower shaft; (7) cam; (8, 9) corrective masses for upper and lower hinges, respectively; (10) pin; (11) nut; (12) bearing; K\u2013K, boundary between upper and lower parts of module.", + "texts": [ + "eywords: saw module, saw assembly, dynamic stability, corrective masses DOI: 10.3103/S1068798X20060076 A fundamentally new saw system is undergoing development and plant testing at present; versions for cutting logs and double-edge beams are under design (Fig. 1). The machine consists of two main mechanisms: one for introducing the workpiece in the cutting zone; and the other for removing the finished product from the cutting zone (Fig. 2). The system includes six identical saw modules (Fig. 3). Sharpening the tool in the saw module decreases its mass. In addition, in order to decrease costs, long blades are replaced by short blades. In other words, the number of teeth is decreased, since coating with hard alloy of stellite type to increase tooth hardness considerably increases the cost of the blade. 45 With variation in mass of the saw modules, their dynamic characteristics are impaired, and the operational stability of the blades is decreased. Dynamic stability of the saw blades may be ensured by means of corrective masses at the upper (8) and lower (9) hinges (Fig. 3). Therefore, it is important to understand the influence of the corrective masses. When the saw module moves, the forces F1 and F8 of the corrective masses (Fig. 4) balance the inertial 3 Fig. 4. Inertial forces in saw module when \u03c8 = 180\u00b0. F1 F F m1 m2 F2 F3 Mfr Mfr F4 Zmax A B q b e e F5 F6 2 1 1 = 2 F7 O2 O2 L 2L 1 L 4 L 3 L 8 L 7 L 6 L 5 L 9 L 10 L L /4 L /4 O1 O1 F8 forces F4 and F6 arising in the blade. However, on installing one or two additional blades and on switching between short and long blades, it is important to know the range of corrective masses capable of ensuring blade stability in the saw module", + "33 depends on the ratio b/h; b is the blade width; and h is its thickness. Then (2) The distributed load q corresponding to the blade\u2019s inertial forces and the following torques acts on the saw blade: the blade\u2019s tensile torque \u041cF; the torques \u041cu.cm and \u041cl.cm corresponding to the corrective masses at the upper and lower parts of the saw module; and the frictional torque \u041cfr. The limiting equilibrium condition for the system may be written for the upper part of the saw module (above line K\u2013K in Fig. 3) as follows (3) Here \u041cu.b is the torque corresponding to the reduced inertial forces of the mass of the upper free part of the blade (with center of mass at point \u0410) (4) where mb is the mass of the free part of the blade of length L (Fig. 3); e is the eccentricity of the circular motion of any point of the saw module; n is the rotary speed of the saw module; and L4 is the distance at which force F4 acts relative to point \u041e1. The maximum tensile torque when \u03c8 = 180\u00b0 is \u041cF = FZmax, where Zmax = e + b/2 + ht is the maximum ( )1/2 cr fl cr / .\u041c EJ GJ l= \u03c0 ( )[ ]( ){ }1/23 3 cr /12 ' / 2 1 / .\u041c \u0415 h b h b \u0415 l = \u03c0 \u03b2 + \u03bc cr u.b u.cm u.fr.F\u041c \u041c \u041c \u041c \u041c= \u2212 + + \u2212 \u2212= \u03c01 2 2 u.b b 40.5 4 ,\u041c m g en L RUSSIAN ENGINEERING RESEARCH Vol. 40 No. 6 eccentricity of the blade\u2019s tensile force F with respect to the tooth-tip line when sawing with negative feedback", + " The inertial torque due to the upper corrective mass m1 is (6) The frictional torque of the upper hinge is disregarded in calculating the frictional torque of the upper and lower parts of the saw module, because it is small (\u041cu.fr \u2248 0.0045\u041ccr). Taking account of the torques associated with the corrective mass, the blade inertia at point \u0410, blade tension, and friction, we may write the limiting stability condition for a plane saw blade in the form (7) Note that the opposing torques \u041cu, \u041clo, and \u041cF act in the plane of the blade and are in equilibrium at the line K\u2013K (Fig. 3). From Eqs. (2)\u2013(7), we obtain the critical inertial torque and tensile torque in the saw module for half the free length of the blade, disregarding the frictional force in the hinge (8) From Eq. (8), we obtain an expression for the critical rotary speed ncr of the saw module, above which stability loss of the blade is observed In the calculations, we want to find ncr, the critical rotary speed of the saw module. The variables are F, the tensile force of the saw blade with eccentric tension Zmax = e + b/2 + ht; and m1, the corrective mass of the upper hinge" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0002368_icicm50929.2020.9292221-Figure1-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0002368_icicm50929.2020.9292221-Figure1-1.png", + "caption": "Figure 1. Internal UHF partial discharge detection method of GIS.", + "texts": [ + " Through the on-site basin-type insulator tightening screw inspection, the judgment is consistent with the test conclusion, which verifies the accuracy of on-site live detection. II. PARTIAL DISCHARGE DETECTION AND POSITIONING TECHNOLOGY OF GIS In recent years, UHF detection technology has received extensive attention and application in the detection and positioning of partial discharges in GIS. In the current GIS state detection technology, partial discharge detection has the best effect. [3]. The basic principle of the UHF detection method is shown in Figure 1. When a partial discharge occurs in the GIS, the excited UHF electromagnetic wave signal of several GHz will be transmitted through non-metallic parts such as epoxy materials. we can detect partial discharge signals and analyze the severity of partial discharge through the built-in or external UHF sensor to receive UHF electromagnetic waves [4-5]. UHF detection method can judge the type of insulation defect according to the characteristics of the UHF signal amplitude, quantity, phase distribution, and frequency 14 978-1-7281-8978-9/20/$31" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0003491_ilt-01-2015-0005-Figure1-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0003491_ilt-01-2015-0005-Figure1-1.png", + "caption": "Figure 1 The linear rolling guide with squeeze film damper", + "texts": [ + " The effect of surface roughness on squeeze film damping characteristics of the damper is analyzed and discussed. The dynamic behaviors of the linear rolling guide as the important functional component of a CNC machine tool can have an influence on the machining accuracy of parts. The damping performance of the linear rolling guide needs to be further improved due to the finite points or lines contacts of the rollers between slider and rail. One way is to install the squeeze film damper on the linear rolling guide based on squeeze film damping effect. Figure 1 shows the schematic drawing of the linear rolling guide with a squeeze film damper. Two sliders are connected together with the centered damper by the bench. In this case, the sliders have contact with the rail through rollers, while the thin film of lubricant (about 15-32 m) between the damper and rail is formed when the sliders and the damper are assembled through the bench. The excitation force acting on the bench can bring about deformation of the rollers, which leads to the damper moving toward the rail" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0002386_012019-Figure3-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0002386_012019-Figure3-1.png", + "caption": "Figure 3 \u2013 Delineation of UAV motion in 3 dimensional space", + "texts": [], + "surrounding_texts": [ + "Feedback linearization is a common approach used in controlling nonlinear system. In addressing the UAV environment a nonlinear dynamic motion employed the concept of dual controller. For example, feedback linearization is applied within an inner controller and also with linear outer controller to achieve robust navigation capabilities. Linearity is not invariant under nonlinear state feedback. Therefore, it may be possible to convert nonlinear system into linear system under this transformation. This is called feedback linearization. Adaptive feedback linearization is used for altering the incorrect estimation of the nonlinearity. This utilizes lyapunov stability criteria and produces a parameter update rule that treats with time for the stability of the system. This chapter gives a brief overview of some of the most commonly used methods for navigation and tracking for Unmanned Aerial Vehicles. Modes of UAV navigation using vision sensors, inertial sensors, integrated sensor systems, hierarchical control and feedback linearization are discussed with proper examples." + ] + }, + { + "image_filename": "designv11_71_0000487_bibe.2019.00186-Figure3-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0000487_bibe.2019.00186-Figure3-1.png", + "caption": "Fig. 3. Photographs of the robot\u2019s top (a) and bottom (b). The 3D virtual model (c) was captured from the Unity3D console.", + "texts": [ + " To address the limited computational power and memory of the Microsoft HoloLens and enable real-time interaction, all processing of the MRI data sets are handled on the Host PC. Figure 1(b) illustrates the interaction between the HoloLens and the Host PC. The operator requests data by using hand gestures or voice commands, to which the PC responds accordingly. The holographic scene in Figure 2(a) includes an MRI slice, segmentations of segmented structures (blood vessels) extracted from those sets, a model of the manipulator, and a needle. Figure 2(b) illustrates the workspace and Figure 2(c) displays the 2D virtual window to show individual slices. Figure 3 shows the physical prototype and the model of the three-DoF MRI-compatible robot [14], [15]. The manipulator has two actuated DoF: a rotating ring and a prismatic joint that carries the tool with a ball-and-socket A. With the combined actuation of DoF-1 and DoF-2, point A can be positioned inside a circle of radius 25 mm. Since the distal end of the tool is carried by ball-and-socket B, which is permanently anchored onto the frame of the robot, the tip of the tool can be at any location of a spherical cone workspace" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0000327_j.mechmachtheory.2019.103729-Figure6-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0000327_j.mechmachtheory.2019.103729-Figure6-1.png", + "caption": "Fig. 6. Kinematic diagram of the outer loop mechanism in the aligned state.", + "texts": [ + " C I 1 = \u23a1 \u23a2 \u23a2 \u23a2 \u23a2 \u23a2 \u23a2 \u23a2 \u23a2 \u23a2 \u23a2 \u23a2 \u23a3 2 0 3 0 4 0 9 0 2 \u2032 0 3 \u2032 0 4 \u2032 0 9 \u2032 S 1 0 0 0 0 0 0 0 S 1 0 0 0 0 0 0 0 S 1 0 0 0 0 0 0 0 S 1 0 0 0 0 0 0 0 S 1 0 0 0 0 0 0 0 S 1 0 0 0 0 0 0 0 S 1 S 1 S 1 S 1 S 1 S 1 S 1 S 1 0 \u23a4 \u23a5 \u23a5 \u23a5 \u23a5 \u23a5 \u23a5 \u23a5 \u23a5 \u23a5 \u23a5 \u23a5 \u23a6 (5) Thus the adjacency matrix C O 1 of the outer loop mechanism can be obtained. C O 1 = C 1 \u2212 C I 1 = \u23a1 \u23a2 \u23a2 \u23a2 \u23a2 \u23a2 \u23a2 \u23a2 \u23a2 \u23a2 \u23a2 \u23a2 \u23a3 2 0 3 E 1 4 0 9 E 1 2 \u2032 E 1 3 \u2032 0 4 0 9 \u2032 0 E 1 0 E 1 0 0 E 1 0 0 0 E 1 0 E 1 0 0 E 1 0 E 1 0 E 1 0 0 0 0 R 1 E 1 0 0 0 0 E 1 0 R 1 0 E 1 0 0 E 1 0 E 1 0 0 0 E 1 0 0 E 1 0 R 1 0 0 0 R 1 R 1 0 R 1 0 \u23a4 \u23a5 \u23a5 \u23a5 \u23a5 \u23a5 \u23a5 \u23a5 \u23a5 \u23a5 \u23a5 \u23a5 \u23a6 (6) A mechanism diagram of the corresponding outer loop mechanism is shown in Fig. 6 . 2) Kinematic diagram of second configuration of Rubik\u2019s Cube unit mechanism The particularity of Rubik\u2019s Cube mechanism relies not only on the complexity of its structure, but also on the intensification of its invariable components during its movement. Thus, when the Rubik\u2019s Cube rotates in a single layer, the position relationship between all moving components remains unchanged and the DOF remains constant. When the Rubik\u2019s Cube mechanism is in motion, the intensification exhibited by the component occurs in the non-aligned state", + " (6) Each three-loop mechanism is equivalent to a serial branch and the outer loop mechanism is equivalent to a three- chain parallel mechanism to solve its DOF. (7) The DOF of the cube unit is obtained by intersection of the motion-screw systems of the outer loop mechanism and the inner common spherical pair. The Rubik\u2019s Cube mechanism has a plurality of configurations, the most important of which is the first one that is the aligned state. The adjacency matrix (4) of the first configuration is established in Section 2 and the mechanism diagram of the Rubik\u2019s Cube in the first configuration is shown in Fig. 6 . Moreover, the loop connection is shown by the hierarchical directed graph of the outer loop mechanism in aligned state which is Fig. 8 . On this basis, the DOF of the first configuration (aligned state) will be analyzed in this section. In the above, a mechanism diagram of the unit mechanism of the first configuration of the Rubik\u2019s Cube is shown in Fig. 6 . The mechanism in this configuration includes eight components 9 \u2032 ,9,2 \u2032 ,4 \u2032 ,2,4,3 \u2032 and 3, in which 9 \u2032 is 0-rank; 9,2 \u2032 and 4 \u2032 are 1-rank; 2,4 and 3 \u2032 are 2-rank; 3 is 3-rank. The focus of solving the DOF is to analyze the DOFs of components at all ranks. In order to analyze this complex mechanism, a method based on loop layer decomposition and superimpose is adopted. Firstly, the loop to be analyzed and the order of analysis need to be determined. As can be seen from the hierarchical directed graph, Fig", + " Its external motion-screw system { S/ O3 } is expressed as: { S/ O3 } = \u23a7 \u23aa \u23aa \u23aa \u23aa \u23aa \u23aa \u23aa \u23aa \u23a8 \u23aa \u23aa \u23aa \u23aa \u23aa \u23aa \u23aa \u23aa \u23a9 S/ O3 1 = ( 1 0 0 ; 0 0 0 ) S/ O3 2 = ( 0 1 0 ; 0 0 0 ) S/ O3 3 = ( 0 0 1 ; 0 0 0 ) S/ O3 4 = ( 0 0 0 ; 1 0 0 ) S/ O3 5 = ( 0 0 0 ; 0 1 0 ) S/ O3 6 = ( 0 0 0 ; 0 0 1 ) \u23ab \u23aa \u23aa \u23aa \u23aa \u23aa \u23aa \u23aa \u23aa \u23ac \u23aa \u23aa \u23aa \u23aa \u23aa \u23aa \u23aa \u23aa \u23ad (23) Because of { S/ O3 } \u2283 { S/ I } , the final motion-screw system of the corner piece is the internal common motion-screw system. { S/ 3 } = { S/ O3 } \u2229 { S/ I } = \u23a7 \u23aa \u23a8 \u23aa \u23a9 S/ 2 1 = ( 1 0 0 ; 0 0 0 ) S/ 3 2 = ( 0 1 0 ; 0 0 0 ) S/ 3 3 = ( 0 0 1 ; 0 0 0 ) \u23ab \u23aa \u23ac \u23aa \u23ad (24) Therefore, the final DOF of corner piece is 3, with three rotational DOFs. The mechanism diagram of the outer loop mechanism in the second configuration of the Rubik\u2019s Cube mechanism is shown in Fig. 7 . Comparing with the outer loop mechanism diagram of the first configuration of Fig. 6 , it can be known that the Rubik\u2019s Cube mechanism from the aligned state to non-aligned state has the following changes: New kinematic pairs have been generated, the orientation characteristic of the kinematic pair has been changed and Structural symmetry of unit mechanism has been broken. Therefore, the loops become more complicated. This section takes the second configuration of rotation about the Z axis as an example to analyze the DOF of the cube in the non-aligned state. Since the structural symmetry of the unit mechanism is broken, components with the same rank will be differentiated and different components at the same rank will have different DOFs" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0002827_ijvas.2017.087124-Figure1-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0002827_ijvas.2017.087124-Figure1-1.png", + "caption": "Figure 1 Four-wheel non-linear vehicle dynamics model", + "texts": [ + " Then, the tyre model and tyre properties related to the estimation are introduced in detail. Correspondingly, the EKF-based observer is designed and validated through simulation in Matlab/Simulink/ CarSim. In order to make the full use of the available ESC sensors (including four wheel speeds, yaw rate, longitudinal and lateral acceleration), the proposed observer adopts a 7-DOF vehicle model as the reference model (accordingly, including four wheels rotation dynamics, vehicle yaw motion, vehicle longitudinal and lateral motions), as shown in Figure 1. The vehicle roll dynamics is neglected since it will significantly increase the coupling property of the movement of the centre of gravity (c.g.) leaving no significant effect on the estimation algorithm in this paper. The equations for vehicle body dynamics are given by equations (1)\u2013(3): cos cos sin1 sin xfl fl xfr fr yfl fl x y yfr fr xrl xrr res F F F v v F F F Fm (1) sin sin cos1 cos xfl fl xfr fr yfl fl y x yfr fr yrl yrr F F F v v F F Fm (2) sin sin cos cos cos cos 2 sin s 2 2 1 in f xfl fl xfr fr f yfl fl yfr fr r yrl yrr xfl fl xfr fr f xrl xrr yfl fl yf fr z r l F F l F F df l F F F F ddr F F F F I (3) where Fres is the total resistance consisting of tyre rolling resistance and air resistance" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0003311_issnip.2015.7106943-Figure1-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0003311_issnip.2015.7106943-Figure1-1.png", + "caption": "Fig. 1. Coordinate systems of a flying camera and a target with respect to the global coordinates. The cone models the field of view of the camera with view angle \u03c6 and capture distance h.", + "texts": [ + " Ri(k) is the rotation matrix from the local (body) to the global coordinate system that defines 1 978-1-4799-8055-0/15/$31.00 \u00a92015 IEEE the flying camera\u2019s attitude. The body is defined by three main directions b1,i(k) = Ri(k)e1, b2,i(k) = Ri(k)e2 and b3,i(k) = \u2212Ri(k)e3 where e1 = (1, 0, 0)T , e2 = (0, 1, 0)T and e3 = (0, 0, 1)T . The flying camera\u2019s motion is determined by a controller that follows a desired trajectory Xd,i = {xd,i(k)}Kk=0, with xd,i(k) \u2208 R 3, and a desired direction of the first body direction b1d,i(k) [15]. Fig. 1 summarizes the variables. The target trajectory is defined as Xt = {xt(k)}Kk=0, where xt(k) \u2208 R 3 is the position of the target in the global coordinate system. The desired (goal) locations of each flying camera are defined by the set G(k) = {gi(k)}Ni=1, where gi(k) \u2208 R 3. The relative goal locations are known to the flying cameras [14]. The target state is t(k) = (xt(k), Rt(k), Qt), where Rt(k) defines the target attitude and Qt defines the 3D bounding volume occupied by the target. Let us assume that the target location xt(k) is known by each flying camera with respect to its local coordinate system", + " (10) The dynamics of the camera-equipped quadrotors are determined by four identical propellers, which are equidistant from the centre of body, and generate a thrust and torque orthogonal to the plane defined by b1,i(k) and b2,i(k) [15]. The motion of the ith camera can be expressed as1 x\u0307i = vi miv\u0307i = mige3 \u2212 fiRie3 R\u0307i = Ri\u03a9\u0302i Ji\u03a9\u0307i +\u03a9i \u00d7 Ji\u03a9i = Mi, (11) where vi \u2208 R 3 is the velocity in the global coordinate system, g = 9.8m/s2 is the gravity acceleration, fi = \u22114 n=1 fn,i \u2208 R 3 is the total thrust generated by the four propellers in b3,i(k) direction (Fig. 1), \u03a9i \u2208 R 3 is the angular velocity, Ji \u2208 R 3\u00d73 is the inertia matrix, Mi \u2208 R 3 is the total moment (the last three terms are calculated in the local coordinate system). The hat operator is defined as x\u0302y = x\u00d7y for all x, y \u2208 R 3, where \u00d7 is the cross product [15]. Given a transitional control command, the camera\u2019s dynamics are controlled by the magnitude of the total thrust fi and by the control moment Mi = (M1,i,M2,i,M3,i) T for the desired attitude:\u23a1 \u23a2\u23a3 fi M1,i M2,i M3,i \u23a4 \u23a5\u23a6 = \u23a1 \u23a2\u23a3 1 1 1 1 0 \u2212d 0 d d 0 \u2212d 0 \u2212c\u03c4 c\u03c4 \u2212c\u03c4 c\u03c4 \u23a4 \u23a5\u23a6 \u23a1 \u23a2\u23a3 f1,i f2,i f3,i f4,i \u23a4 \u23a5\u23a6 , (12) where d is the distance from the center of the flying camera to the center of each rotor and c\u03c4 is a constant" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0001077_ab8ef0-Figure2-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0001077_ab8ef0-Figure2-1.png", + "caption": "Figure 2. Schematic diagram of computing domain.", + "texts": [ + " r r r r r t \u00b6 \u00b6 + \u00b6 \u00b6 + \u00b6 \u00b6 - \u00b6 \u00b6 - \u00b6 \u00b6 = - \u00b6 \u00b6 \u00b6 \u00b6 + \u00b6 \u00b6 + \u00b6 \u00b6 + \u00b6 \u00b6 T t U T X V T Y q H T z H T Z t h a T T P t U P X V P Y h a H U Z Pr Re 1 Ec Y Ec Yt 11 2 2 2 0 2 0 * \u239c \u239f\u239c \u239f \u239c \u239f \u239b \u239d\u239c \u239e \u23a0\u239f \u239b \u239d \u239e \u23a0 \u239b \u239d \u239e \u23a0 \u239b \u239d \u239e \u23a0 \u00b7 \u00af \u00af \u00af \u00af \u00af \u00af \u00af \u00af \u00af \u00af \u00b7 \u00af \u00b7 \u00af \u00af \u00b7 \u00af \u00af \u00af \u00b7 \u00af \u00b7 ( ) where \u00f2 \u00f2r r= \u00b6 \u00b6 \u00a2 + \u00b6 \u00b6 \u00a2q t H dZ X H UdZ , Z Z 0 0 \u239b \u239d\u239c \u239e \u23a0\u239f \u239b \u239d\u239c \u239e \u23a0\u239f\u00af \u00af \u00af \u00af h r h = = = = = c k h a u a u cT cap u k Pr , Re Re , Re , Ec , Yt . 0 0 2 0 0 0 0 2 0 H 0 * \u239c \u239f\u239b \u239d \u239e \u23a0 4. Numerical methodology A multi-grid method was applied to compute the pressure solution. A multilevel integration was implemented to speed up the calculation of surface elastic deformation. For both steady-state and transient cases, the solution domain was determined as- X2.5 1.5 and- Y1.5 1.5, and the schematic diagram of computing domain is presented in figure 2. In z direction, there are 20 nodes in lubricating film layer and 12 nodes in surface 1 and 2 layers, respectively. To sum up, there are 44 nodes in z direction, and it is assumed that the temperature of the coincident node between lubricant and solid surface is identical. To note, the z-coordinate isometric grid spacing is adopted in film layer and unequal grid spacing is applied in solid surface layers. The lubricant and two solid surface flash temperatures were computed by means of column by column scanning technology" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0001314_1464419320933382-Figure6-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0001314_1464419320933382-Figure6-1.png", + "caption": "Figure 6. Three-dimensional model of the PTL representing the conductors on quad bundled using ACSR. ACSR: aluminium conductor steel reinforced.", + "texts": [ + " From the above, the Newton\u2013Euler dynamic equation of the robot, which reflects the relationship between the resultant external wrench and acceleration twist of the robot, is given by Ir CT r Cr 0 \" # ar kr \u00bc Fr ur \u00f016\u00de in which ar \u00bc aTb aT1 aTn T is the acceleration twist of the robot, fcr \u00bc CT r kr is the constraint wrench of the robot where kr is the coordinate vector of constraint force. ANCF is used in PTL dynamics because the approach has better handling of cable problem which has the characteristics of dynamic flexibility and high nonlinearity. As shown in Figure 6, ACSR for type LGJ-400/35 is used by the conductors on quad bundled where the PTL maintenance robot works. The ACSR physical parameters are listed in Table 2. ACSR can be divided into the following structures: the inner layer (or the core wire) is a single-stranded or multi-strand galvanized steel stranded wire bearing force load; the outer layer is a single-layer or multi-layer hard aluminium stranded wire being conductive. The cable mainly bears axial tensile stress, bending stress and torsional shear stress under external load and self-gravity" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0003542_syseng.2015.7302733-Figure4-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0003542_syseng.2015.7302733-Figure4-1.png", + "caption": "Fig. 4. Heterogeneous system model for vibro-acoustic analysis of a two-stage electro-mechanical drive train", + "texts": [ + " From a system level point of view and with regard to the model requirements, [18] provides a study of three appropriate approximation methods: \u2022 Equivalent radiated power \u2022 Method based on volume velocities \u2022 Lumped parameter model for sound radiation One should note that these approximation methods assume unidirectional coupling. Hence, fluid fluctuation has no influence on the vibrating structure. However, for many applications this is a valid assumption. IV. MODEL VALIDATION, RESULTS AND LIMITATIONS In accordance with the proposed generic simulation method the model of an electro-mechanical drive train is generated as depicted in Fig. 4. The periphery includes power supply, motor controller, position sensor, power electronics, clutch, magnetic particle brake, and electric current dynamics of the motor. Transmission gearing is a twostage gear with 22 mechanical degrees of freedom. Though dampers are not illustrated due to clarity reasons they exist in parallel to each spring. The model is generated in Simscape including an embedded SimMechanics multi-body simulation. As Simscape does not cover all the previously defined modeling techniques, the simulator requires two functional extensions: \u2022 Gear mesh model (lumped parameter model) \u2022 Flexible body superposition (state space form) Both functionalities are implemented following literature: According to [16] mechanical vibration excitation is accounted for within the gear mesh" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0000231_s00285-019-01455-z-Figure4-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0000231_s00285-019-01455-z-Figure4-1.png", + "caption": "Fig. 4 Arm model with one degree-of-freedom at the elbow joint", + "texts": [ + " In the case of Hatze\u2019s activation dynamics and four muscle groups, the complete system dynamics is given by \u239b \u239c\u239c\u239c\u239c\u239c\u239c\u239c\u239c\u239c\u239c\u239c\u239c\u239c\u239c\u239c\u239c\u239c\u239d \u03b3\u03071 \u03b3\u03072 \u03b3\u03073 \u03b3\u03074 l\u0307CE1 l\u0307CE2 l\u0307CE3 l\u0307CE4 \u03d5\u0307h \u03d5\u0307 \u239e \u239f\u239f\u239f\u239f\u239f\u239f\u239f\u239f\u239f\u239f\u239f\u239f\u239f\u239f\u239f\u239f\u239f\u23a0 = \u239b \u239c\u239c\u239c\u239c\u239c\u239c\u239c\u239c\u239c\u239c\u239c\u239c\u239c\u239c\u239c\u239c\u239c\u239c\u239c\u239c\u239c\u239c\u239c\u239d ma(u1 \u2212 \u03b31) ma(u2 \u2212 \u03b32) ma(u3 \u2212 \u03b33) ma(u4 \u2212 \u03b34) \u2212a1 F isom(lCE1 )Fmax 1 +FSEE(lCE1 ,\u03d5)+d l\u0307MTU 1 (\u03d5\u0307,\u03d5) a1 F isom(lCE1 )\u03bc1 Fmax 1 +d \u2212a2 F isom(lCE2 )Fmax 2 +FSEE(lCE2 ,\u03d5)+d l\u0307MTU 2 (\u03d5\u0307,\u03d5) a2 F isom(lCE2 )\u03bc2 Fmax 2 +d \u2212a3 F isom(lCE3 )Fmax 3 +FSEE(lCE3 ,\u03d5)+d l\u0307MTU 3 (\u03d5\u0307,\u03d5) a3 F isom(lCE3 )\u03bc3 Fmax 3 +d \u2212a4 F isom(lCE4 )Fmax 4 +FSEE(lCE4 ,\u03d5)+d l\u0307MTU 4 (\u03d5\u0307,\u03d5) a4 F isom(lCE4 )\u03bc4 Fmax 4 +d \u22114 i=1 ri (\u03d5) FMTU(ai ,lCEi ,l\u0307CEi ) J \u03d5h \u239e \u239f\u239f\u239f\u239f\u239f\u239f\u239f\u239f\u239f\u239f\u239f\u239f\u239f\u239f\u239f\u239f\u239f\u239f\u239f\u239f\u239f\u239f\u239f\u23a0 . (18) The formulation of muscle activation was introduced by Hatze (1977) and slightly restated by Kistemaker et al. (2006) (Fig. 4). Herein, the activity ai is calculated depending on the free calcium ion concentration \u03b3i as well as on the length of the contractile element lCEi ai (\u03b3i , l CE i ) = a0 + ( \u03c1(lCEi )\u03b3 )3 1 + ( \u03c1(lCEi )\u03b3 )3 , (19) where the monotonically increasing function \u03c1 is calculated as Table 3 Activation dynamics parameters for the formulation of Zajac \u03c4act (s) \u03c4deact (s) a0 (\u2212) 0.02 0.06 1e\u22123 \u03c1(lCEi ) = c\u03b7 k \u2212 1 k \u2212 lCEi lCE,opt lCEi lCE,opt . (20) The values of the constants ma, c, \u03b7, a0 and k are based on the work of Kistemaker et al" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0001971_icsgrc49013.2020.9232583-Figure3-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0001971_icsgrc49013.2020.9232583-Figure3-1.png", + "caption": "Fig. 3. CAD Design", + "texts": [ + " The most common printing specimen for FDM is Polylactic Acid (PLA), because it is one of the most common thermoplastics that is used to making common patron products. It can endure heat, chemical compounds and mechanical stress which makes it best material for 3D printing. In the slicing software, there are many types of parameters. This different type of parameters is shown in the Table 1. The researchers have utilized the software AutoCAD in order to have a clear design that would be inputted into the FDM machine. As seen in Figure 3, the specimen has a horizontal dimension such as a 30x30mm square, an inner diameter of 7.5mm and 15mm while the vertical ones have a total height of 30mm, half height for the square and cylinder measuring 15mm, a square of 7.5x7.5mm and finally an inner diameter of 10mm. The researchers have measured the length, width, height, in addition to the diameter and the depth of the holes. The build orientation has started from bottom to top since the area from the bottom is larger than the top because if the researchers started from top, they will need a support on the specimen to make it stand vertically" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0003146_978-3-662-45514-2_9-Figure1-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0003146_978-3-662-45514-2_9-Figure1-1.png", + "caption": "Fig. 1. Quadrotor model", + "texts": [ + " The original complex and uncertain plant dynamics is reduced to a simple cascade integral plant which can be easily controlled. Quadrotor consists of a cross frame equipped with four rotors which are powered by electric motors as shown in Fig. 4. Vertical motion is achieved by increasing or decreasing the thrust of the propellers at the same amount. Motor pairs 1, 3 and 2, 4 rotate in the opposite direction and yaw command can be achieved by increasing or decreasing the motor pair speeds. Roll command can be achieved by changing the speed of the motors 2 and 4 conversely and for pitch 1 and 3. As shown in the Fig. 1. c represents the quadrotor center of gravity without 2 DOF arm. denotes the inertial reference frame whose origin o. denotes the body reference frame whose origin is c where and are intermediate reference frames between and , , are Euler angles in yaw, pitch, roll directions respectively. Rotation sequence is given in Eq. (1). ( )\u2192 ( )\u2192 ( )\u2192 (1) The states variables of the quadrotor are given below. \u2022 {X,Y,Z} inertial positions and altitude along { ( ), ( ), ( )} in \u2022 { , , } body frame velocity components measured along { ( ), ( ), ( )} in \u2022 { , , } roll, pitch and yaw angles with respect to { , , } \u2022 { , , } roll, pitch, yaw rates measured along{ ( ), ( ), ( )} in It\u2019s assumed that the generated force and drag moment is proportional to the square of the propeller speed" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0001274_0142331220930623-Figure4-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0001274_0142331220930623-Figure4-1.png", + "caption": "Figure 4. The closed-loop experiment of the EHA. (a) Schematic diagram; (b) Prototype.", + "texts": [ + " After a careful evaluation, the amplification coefficient KF is set to 10. A sliding resistance displacement sensor (Type: SUKENKON KTC-75, Linear:60.05%) is used to detect the displacement signal of the hydraulic cylinder. The triangular wave signal with amplitude of 5V is taken as the input signal by a signal generator (Type: Tektronix AFG3051C). To plot the Bode diagram, the displacements of piston at different frequencies are recorded by a recorder (Type: YOKOGAWA DL350). The schematic diagram and prototype of the closed-loop experiment are shown in Figure 4(a) and Figure 4(b), respectively. According to the requirements of the exoskeleton\u2019s dynamic response, the frequency response should be above 2Hz. Therefore, the input signal with a larger frequency range (0-10Hz) is selected for testing. However, the amplitude frequency of the prototype decreased to -3dB at about 1Hz. The speed of decrease is faster and faster with the increase of the frequency. In order to make the measurement more accurate, the frequency range of the input signal is adjusted to 0- 2Hz. Finally, the displacements of piston are recorded when the frequency of the input signal increases from 0" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0003973_amm.659.515-Figure1-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0003973_amm.659.515-Figure1-1.png", + "caption": "Fig. 1. Forces and torques acting on a system tractor-trailer during braking.", + "texts": [ + " Mathematical model to calculate the braking forces and the force at the hitch point The study on the braking dynamics of the tractor-trailer system calls for the general principles concerning the dynamics of mechanical systems. The complexity of the study consists in establishing the calculation diagram. This paper analyzes the rectilinear motion of the tractor-trailer system under the conditions of intense braking. The diagram of forces and torques actuating upon the tractor-trailer during braking is presented in Fig. 1. Generally, the following forces actuate upon the tractor-trailer system, which leads to speed reduction: braking force, rolling resistance force, aerodynamic drag, frictions in the bearings of the rolling system etc. The value of braking forces, under the conditions of intense braking, represents 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", + " 1 i redAredT F mm x (6) The braking force at tractor\u2019s wheels is considered to vary according to an exponential law: ( )( ) ( )( ) ( )\u2211 ++\u2212\u2212+\u2212\u2212== \u03b1\u2212\u03b1\u2212 ,11)( maxmax fGGefZFefZFtFF AT t AA t TTi (7) where: maxTF denotes the maximum braking force of the tractor; maxAF \u2013 maximum braking force of the trailer; ZT \u2013 weight distributed on the trailer\u2019s braked wheels; ZA \u2013 weight distributed on the trailer\u2019s braked wheels; GT \u2013 tractor\u2019s total weight; GA \u2013 trailer\u2019s total weight; f \u2013 rolling resistance coefficient; \u03b1 = 6/t0 \u2013 exponent depending on the time t0 of brakes complete application. Making use of the bond insulation principle, the tractor-trailer system interaction is analyzed during the braking process. In Fig. 1 the binding forces are noted with F. Balance equations are written separately for both tractor and trailer: +\u2212= \u2212\u2212= , ; FFxm FFxm AA TT (8) where: FT denotes the total braking force of the tractor: 21 TTT FFF += , (for a tractor with brakes on the rear axle only, FT1 = 0); FA \u2013 denotes the total braking force of the trailer: 21 AAA FFF += . It is considered that the tractor and the trailer are not braked simultaneously: the trailer is braked in advance. In order to analyze this case, the braking forces are determined separately for both tractor and trailer" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0001961_j.matpr.2020.09.086-Figure4-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0001961_j.matpr.2020.09.086-Figure4-1.png", + "caption": "Fig. 4. Shell fin protrusions.", + "texts": [ + " Processing of parts is carried out when they rotate in the same plane, with a radius of rotation of the centers of the machined surfaces, equal to 140 mm. rotation speed nvr = 100 rpm and compressed air pressure from 0.2 to 0.4 MPa. To process the fins and bottom of the cooling channels, you can use beads with a diameter of 50\u2013100 mm, based on structural and technological considerations. Design considerations are determined by the small size of the channel, the required depth of the turbulent imprint, and the rigidity of the shell. Technological requirements are to remove burrs from the edges of the ribs (Fig. 4) without blunting the angle (Rmax = 0.1 mm) for the subsequent brazing operation with a smooth shell forming closed channels. The main problem in our case is the warpage of the finned finned fin shell due to the hardening of the surface layer caused by shot peening. Ways to reduce unwanted hardening are as follows: the low kinetic energy of the microspheres does not in itself allow hardening of more than 2\u20133% on plastic materials to a depth of not more than 0.03\u20130.05 mm, but even such hardening of the surface layer is undesirable; To remove part of the hardened layer, it is proposed to use it in conjunction with the mechanical effect of the effect of anodic dissolution of the material" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0000603_iccas47443.2019.8971558-Figure2-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0000603_iccas47443.2019.8971558-Figure2-1.png", + "caption": "Figure 2. CAD view of TakoBot 2", + "texts": [], + "surrounding_texts": [ + "218\nI. INTRODUCTION\nDue to a great contribution to the investigation on continuum robots made them more applicable for industry and medicine. Today, bio-inspired robots from the elephant trunk, octopus arm and other animals with tentacles represent the great potential to be used almost in every aspect of our life. The slender design and hyper-redundant structure make continuum manipulators to adapt an unstructured environment and reachable in confined workspace easily. Traditional serial manipulators with a rigid structure and motorized joints cannot meet these requirements[1].\nAccording to the backbone structure, continuum manipulators can be divided into three groups: discrete, hard and soft continuum robots. Discrete continuum structure consists of a hard spacer disc and interconnected by universal joints or ball joints, and drives by wires. For instance, Tensor arm, Snake robot from OCRobotics, Elephant trunk and Emma type robots [2-5]. Hard continuum type of robots presents manipulators with an elastic support backbone bending feature, which made of rubber or other soft materials and separated by hard spacer discs. For example, Backbone, Catheter, OctArm and Air-OCTOR robots [6-9]. Such kind of robots actuates by cables, dielectric elastomer and by a pneumatic actuator. Among all actuation system, only cabledriven actuation generates high output force, which dramatically increases robot payload capacity.\nAnd the third type of continuum manipulator is soft continuum robots, which are made of hundred per cent soft\nmaterials such as rubber or silicone. These robots are safest and softest robots which can able to change the body size as well. There are many contributed investigations on these type of robots [10-13].\nDue to such great potential, wire-driven continuum manipulators attracted wide attention in a couple decades. There are many scholars and engineers brought great contribution to the development of continuum robotics field. For instance, I.D. Walker and Jones established a new kinematic model based on improved DH method [14]. Moreover, Han Yuan and Zheng Li also proposed a kinematic analysis of the continuum robot based on a static model [15]. Furthermore, X. Dong et.al. from Nottingham University proposed continuum robot design with twin pivot compliant joint and it moves by twin cable actuation which provides motion and high cable tension. Dong proposed kinematics based on cable length variation[16]. Recently, Tiangliang and et.all. developed a cable-driven redundant spatial manipulator with improved stiffness and load capacity by actuating fewer numbers of cables [17]. Furthermore, Julia Starke and et.al. developed a continuum robot arm with a helical tendon routing system, such kind actuation increased robot dexterity and obstacle avoidance capabilities. However, during the robot motion, torsional stress increased [18].\nThis paper presents a novel discrete wire-driven continuum manipulator with a sliding disc mechanism named TakoBot 2. This is a modified prototype of TakoBot 1 [1]. Based on gained experimental results from TakoBot 1 prototype, a modified prototype should have additional redundant degrees of freedom such as torsional motion and twist motion as well. Moreover, this paper explains forward and inverse kinematic formulation with validated simulation\n978-89-93215-18-2/19/$31.00 \u24d2ICROS\nAuthorized licensed use limited to: Auckland University of Technology. Downloaded on October 25,2020 at 08:37:21 UTC from IEEE Xplore. Restrictions apply.", + "219\nresults. The new prototype supposes to improve robot dexterity and reachability capabilities. In chapter II will explain the concept design, III chapter will explain kinematic kinetic formulation, IV chapter will present experiments and simulation results.\nII. CONCEPT DESIGN\nA. Application Analysis The intended application of the robot is shown in Fig.1. The manipulator is designed to be used in the agriculture field for harvesting, weeding and inspection operations if necessary. The intended working environment is highly constrained and by the time workspace will be different because of growing plants.\nTo perform necessary tasks the robot should have the following functions:\n(1) Flexible dexterity: the robot should work in the confined workspace\n(2) Obstacle avoidance capability: the robot should avoid contact with solid surfaces and not collide.\n(3) Reachability: the robot should get the required position in spite of narrow space and obstacles.\n(4) Safety: the robot should be safe enough to avoid breaking any sticks of the plant.\n(5) Portability: the robot should be compact to be used as a tool for farming CNC platform (Fig.1).\nB. System Design Based on application analysis, the robot should have a slender structure with redundant DOFs. After consideration of previous TakoBot 1 prototype capabilities, we made changes in the design and actuation system as well. After numerous experiments with TakoBot 1, demonstrated some limitations related to the design. For example, a robot could not perform torsional motion during the work which led to the accumulation of strain energy inside of manipulator. Secondly, in case of bending it could not perform a pure bending shape. In the new prototype, we added a passive sliding mechanism (Fig.4). Proposed sliding mechanism demonstrated the following benefits:\n1) Smart bending stress distribution: This ability helped to the manipulator to bend purely and it distributed spring potential energy to the whole segment of the manipulator. In comparison with previous, it required less torque and less load to the cables. 2) Torsional motion: passive sliding disc mechanism provided better dexterity ability in case bending helical motion. Helical motion increases torsional stress of the structure unless a disc does not generate yawing around the z-axis. In spite of the heavier weight of the TakoBot 2 was able to perform better and more accurate.\nPassive sliding disc mechanism works for all discs attached to the backbone except base and end-effector discs, which means total length always remains constant. Fig.4. demonstrates a sliding mechanism structure. According to the design, travelling distance of the disc is 10mm, while the disc diameter is 50mm, an average distance of the single section (between discs) about 35 mm. 3D printed discs connected by coil compression springs, such design provides stiffness to the manipulator. The spring constant could be variable depends on motor torque, in this prototype we used spring with constant 0.63 N/mm. Each section consists of four segments, end segments discs connected by four springs, while mid-section segments connected by 8 springs between discs.\nAuthorized licensed use limited to: Auckland University of Technology. Downloaded on October 25,2020 at 08:37:21 UTC from IEEE Xplore. Restrictions apply.", + "220\nC. Actuating mechanism Manipulator's tendons actuate by a linear lead shaft\nconnected to the stepper motor by a coupler. Tendons attached to both sides of specially designed inlets on screw housing and screw with housing travels along the shaft. 1mm steel cable utilized as a tendon.\nIII. KINEMATICS AND KINETIC FORMULATION\nA. Forward kinematic formulation Coordinate systems are set at every universal joint. The homogeneous coordinate transform matrices:\n\u03a30\u2192\u03a31 , H0,1= ( R u0,1 0 0 0 1 ), u0,1=( x0 y0 l0 )\n(1)\n\u03a3i-1\u2192\u03a3i, Hi-1, i= ( R ui-1,i 0 0 0 1 ) , ui-1,i=( 0 0 L ),\n(i=2, \u22ef, n)\n(2)\nR= Rz(\u03b8zi)Rx(\u03b8xi)Ry(\u03b8yi) (3)\nwhere, \ud835\udc65\ud835\udc650 and \ud835\udc66\ud835\udc660 are an initial position of the base. \ud835\udc45\ud835\udc45\ud835\udc65\ud835\udc65(\ud835\udf03\ud835\udf03\ud835\udc65\ud835\udc65\ud835\udc65\ud835\udc65) and \ud835\udc45\ud835\udc45\ud835\udc66\ud835\udc66(\ud835\udf03\ud835\udf03\ud835\udc66\ud835\udc66\ud835\udc65\ud835\udc65) are rotation matrices of ith universal joint that has two rotation angles \ud835\udf03\ud835\udf03\ud835\udc65\ud835\udc65\ud835\udc65\ud835\udc65 and \ud835\udf03\ud835\udf03\ud835\udc66\ud835\udc66\ud835\udc65\ud835\udc65, \ud835\udc45\ud835\udc45\ud835\udc67\ud835\udc67(\ud835\udf03\ud835\udf03\ud835\udc67\ud835\udc671) is a rotation matrix of the ith disk with a rotation angle \ud835\udf03\ud835\udf03\ud835\udc67\ud835\udc671 along the axial axis and L is a length between neighbouring universal joints. Three rotation matrices have:\nMultiplying the H-matrices successively, we get unit\nvectors and the position vector of the ith coordinate system; H0,i=H0,1H1,2\u22efHi-1, i= (\nii ji ki ui 0 0 0 1 ) (4)\nwhere, \ud835\udc62\ud835\udc62\ud835\udc65\ud835\udc65 is the position of the ith universal joint Ui (\ud835\udc56\ud835\udc56 =\n1,\u22ef , \ud835\udc5b\ud835\udc5b \u2212 1). The position vector \ud835\udc5d\ud835\udc5d\ud835\udc65\ud835\udc65 of the end-point P\ud835\udc5b\ud835\udc5b and position of sliding plates Pi (i=1, \u22ef, n-1) of the manipulator are obtained by, (pi\n1)=H0,i(0 0 li 1)T, (i=1, \u22ef, n) (5)\nwhere, \ud835\udc59\ud835\udc59\ud835\udc5b\ud835\udc5b is a fixed length between the nth universal joint and the most distal plate. Position vectors of 8 hole A0, B0, C0, D0 , A\u03020, B\u03020, C\u03020, D\u03020 at the base plate are determined as,\na0=( ax ay 0 ) , b0=( bx by 0 ) , c0=( cx cy 0 ) , d0=( dx dy 0 ) ,\na\u03020=( a\u0302x a\u0302y 0 ) , b\u03020=( b\u0302x b\u0302y 0 ) , c\u03020=( c\u0302x c\u0302y 0 ) , d\u03020=( d\u0302x d\u0302y 0 ) ,\n(6)\nPosition vectors of 4 hole A\ud835\udc65\ud835\udc65, A\u0302\ud835\udc65\ud835\udc65, C\ud835\udc65\ud835\udc65, C\u0302\ud835\udc65\ud835\udc65 at the ith plate\n(\ud835\udc56\ud835\udc56 = 1,\u22ef , \ud835\udc5b\ud835\udc5b) are obtained as\n(ai 1)=H0,i( ax ay li 1 ) , (a\u0302i 1)=H0,i( a\u0302x a\u0302y li 1 ) ,\n(ci 1)=H0,i( cx cy li 1 ) , (c\u0302i 1)=H0,i( c\u0302x c\u0302y li 1 ) , (i=1, \u22ef, n)\n(7)\nwhere, \ud835\udc59\ud835\udc59\ud835\udc65\ud835\udc65 is an axial length between the ith universal joint and the ith plate, which varies as the plate slides along rods, except \ud835\udc59\ud835\udc59\ud835\udc5b\ud835\udc5b.\nIn the same way, position vectors of 4 hole B, B\u0302\ud835\udc65\ud835\udc65, D\ud835\udc65\ud835\udc65, D\u0302\ud835\udc65\ud835\udc65 at the ith plate (\ud835\udc56\ud835\udc56 = 1,\u22ef ,\ud835\udc5a\ud835\udc5a) are obtained as,\n(bi 1)=H0,i( bx by li 1 ) , (b\u0302i 1 )=H0,i ( b\u0302x b\u0302y li 1) ,\n(di 1)=H0,i( dx dy li 1 ) , (d\u0302i 1 )=H0,i ( d\u0302x d\u0302y li 1) , (i=1, \u22ef, m)\n(8)\nB. Kinetic formulation Our continuum manipulator is divided by two segments.\nAuthorized licensed use limited to: Auckland University of Technology. Downloaded on October 25,2020 at 08:37:21 UTC from IEEE Xplore. Restrictions apply." + ] + }, + { + "image_filename": "designv11_71_0000052_3341215.3356984-Figure2-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0000052_3341215.3356984-Figure2-1.png", + "caption": "Figure 2: Setup of the wooden prototype", + "texts": [ + " The tube has a diameter of 90 cm, an internal diameter of 45 cm, and a tire width of 22,5 cm. The tube is inflated to a certain extent, allowing it to balance out the leaning of the player standing on the board. The tube aligns itself to the weight of the player. In order to secure the board on the tire and have a more modular design, we used a round board. We added some thresholds in the distance of the inner circle to prevent the board from slipping. The sensor was fixed in the middle of the circle (Figure 2). This module can be screwed onto various other boards. These other boards can redistribute the weight of the player in different ways. This type of modular design was chosen as it offers the opportunity to create input for different games. We developed two different types: The first one consists of a wooden plank (87cm x 87cm x 3cm). It allows the player to lean in every direction to some extent. The second setup is a slimmer plate fixed on two long strips, which is screwed to the first plate. The setup redistributes the weight of the player in different ways, making it easier to lean to the right and left than to the front and back" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0000327_j.mechmachtheory.2019.103729-Figure12-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0000327_j.mechmachtheory.2019.103729-Figure12-1.png", + "caption": "Fig. 12. Three-loop mechanism equivalent to generalized kinematic pair.", + "texts": [ + " Then, the sub-loop can be regarded as a motion pair, called the generalized kinematic pair. It can be replaced by serial chains with the same number and DOF. In the main loop analysis, replacing the original sub-loop with a generalized kinematic pair, the number and nature of the mechanical DOF will not be changed. The DOF of this generalized kinematic pair is that of the 2-rank component relative to the fixed platform which is the DOF of the highest rank of the components in the front-end loop. Obviously the equivalent mechanism is a PRREERRP single loop mechanism, as shown in Fig. 12 . In order to obtain the DOF of the highest rank component in this loop, the 3-rank component is regarded as a fixed platform, so that, the mechanism can be replaced by a 2-PRRE parallel mechanism. The motion-screw system of one of the PRRE branches 9 \u2032 - \u22122\u20133 can be expressed as: S/ 1 1 = ( 0 0 0 ; 0 1 0 ) S/ 1 2 = ( 1 0 0 ; 0 0 0 ) S/ 1 3 = ( 0 0 1 ; 0 0 0 ) S/ 1 4 = ( 0 1 0 ; 0 0 0 ) S/ 1 5 = ( 0 0 0 ; 1 0 0 ) S/ 1 6 = ( 0 0 0 ; 0 0 1 ) (15) These screws are independent of each other. The rank of the screw system is 6, which is a six-dimensional screw system with no reciprocal screw" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0002175_s00170-020-06312-8-Figure1-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0002175_s00170-020-06312-8-Figure1-1.png", + "caption": "Fig. 1 Basic configuration of the AGC system", + "texts": [ + " The HAGC system of cold rolling mill uses the merits of a hydraulic servo system. This hydraulic servo system is typically nonlinear because of pressure-flow characteristics, threshold, hysteresis, friction, servo amplifier, nonlinearities of A/D and D/A converters, and so on. The rolled strip thickness is mainly determined by the gap between the two work rollers that are initially set by the HAGC. The work rollers are the rollers used to deform the steel plate, while the backup rolls serve to support the work rolls (Fig. 1). This avoids excessive bending of the work rolls. It is possible to modify the exit thickness by moving the upper roll during the passage, that is to say, hydraulic systems are actuators of the thickness control system. In practical HAGC thickness control process, two main factors will lead to an uncertain time delay for the controlled system. One is the certain distance between the thickness gauge and rolling gap of the stand, and the other is the essentially existing time delay in network transmission" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0000526_icra40945.2020.9197293-Figure1-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0000526_icra40945.2020.9197293-Figure1-1.png", + "caption": "Fig. 1: (a) A non-planar area contact is created when a compliant gripper jaw surface contacts a non-planar object\u2019s surface. (b) An enlarged view of the deformed jaw and the contact profile obtained by the REACH model [11]. The non-planar contact area consists of triangles and the redder colors represent higher pressure due to larger deformation of the jaw pad at that point. (c) A projection of the 6D friction cone that constrains the wrenches that can be applied at the contact. Each ellipsoid represents a projection of the friction limit surface for a given gripper closing force and its center corresponds to the wrench created by the contact pressure.", + "texts": [ + " These models rely on accurate estimation of the forces and torques applied at the contacts between the gripper and object to form the Grasp Wrench Space (GWS), or the set of wrenches that can be applied to the object by a set of gripper jaws. By computing the GWS, we can determine which wrenches the grasp can resist. In previous work, area contact models made up of multiple points or regions have been considered, but these either assume a planar contact area [9, 10, 16] or can be inefficient for fine surface geometries [11]. In this paper, we consider non-planar soft area contacts from a compliant gripper and formulate constraints for the wrenches that can be applied at each contact, as shown in Figure 1. We extend the 6D ellipsoidal model proposed by 1The Autolab at University of California, Berkeley. 2Technical University of Munich, Chair of Media Technology. {jingyi xu, mdanielczuk, goldberg}@berkeley.edu, eckehard.steinbach@tum.de Xu et al. [43] by more efficiently sampling wrenches on the 6D friction limit surface (FLS) without Finite Element Analysis (FEA) and combining the 6D FLS with the normal wrench imposed at the contact. We present a novel 6D friction cone (6DFC) that fully constrains the normal and frictional wrenches that can be applied at a contact by varying the grasp force", + " We propose that the total wrench applied at a contact with a grasp force fC can be modeled with a 6D ellipsoid that is centered at the 6D normal wrench, so that varying the value of fC results in a 7893 Authorized licensed use limited to: Carleton University. Downloaded on October 04,2020 at 15:25:44 UTC from IEEE Xplore. Restrictions apply. 6D friction cone, whose center lies along the vector fN \u2208 R 6 and has 6D frictional ellipsoids as contours for each value of fC , similar to the one shown in Figure 1. We can express this cone as: (w \u2212 fCfN) T A (w \u2212 fCfN) \u2264 f2 C , 0 \u2264 fC \u2264 fC,max (IV.2) Figure 2(c) shows a 3D projection of the 6D friction cone that is produced from a non-planar area contact. For the 3D frictional point contact or 4D soft point contact case discussed above, the ellipsoid that approximates the 2D or 3D friction limit surface can be easily found, since it is axis-aligned. Then, by finding the maximum values for fx, fy , and \u03c4z (in the soft point contact case), we can construct the ellipsoid", + " For each algorithm, parameters such as the friction coefficient, elasticity coefficient, and robustness sample standard deviation were chosen using leave-one-out cross validation. The results for each algorithm are shown in Table I. These results suggest that formulating the 6DFC algorithm can increase the number of successful grasps on the physical system that can be recalled by as much as 17% over existing algorithms while maintaining a similar number of false positives predicted. We find that grasps found only by 6DFC are often of the kind shown in Figure 1, where the contact area is small with high surface curvature, and the grasp may include slight dynamic effects such as a bowing of the jaws. We hypothesize that allowing for reduced closing forces in this scenario could more accurately model the changing contact profile that occurs. We also evaluate the reliability of each algorithm as part of a grasp planning policy via a second experiment. We hypothesize that our algorithm can find grasps in scenarios where a point contact algorithm would not return any highquality grasps due to predicted collisions with other objects, environmental constraints, or motion-planning constraints" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0003638_indin.2014.6945534-Figure1-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0003638_indin.2014.6945534-Figure1-1.png", + "caption": "Fig. 1. Cluster space state variables.", + "texts": [ + "00 \u00a92014 IEEE 332 The benefit such approach is to promote levels of abstraction for the control, allowing specifying the desired motion of the formation (high-level control) without specifying the motions of each individual robot (low-level control). These two control levels are related through kinematic transformations, detailed in [10], [12] and in the sequel. In this work, a formation of three robots is considered, so that the geometric figure dealt with is a triangle. The variables used to represent the cluster thus composed are those illustrated in Fig. 1. Groups of more robots can also be dealt with, using the strategy reported in [12], for instance. The position of the cluster is defined by PF = [xF yF ], the coordinates of the centroid of the triangular formation. Its shape is defined by the vector SF = [\u03d5F pF qF \u03b2F ], whose elements represent, respectively, the orientation of the cluster, the distance between the robots Rb1 and Rb2, the distance between the robots Rb1 and Rb3, and the opening angle of the cluster \u0302Rb2Rb1Rb3. Thus, the triangular cluster is totally defined by Q = [PF SF ] T " + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0003275_aim.2015.7222530-Figure22-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0003275_aim.2015.7222530-Figure22-1.png", + "caption": "Figure 22. Slider-crank mechanism. The centre of the system is located at the centre of the crank.", + "texts": [ + " The peak torque of 10 mNm on the [PM was possible when the external magnetic field was created by the array of magnets set in Configuration2 (see Fig. 19). For this reason, we used the Configuration2 to actuate the drug release mechanism. Its details are presented in the following subsection. We fabricated the slider-crank mechanism in a plastic material (ABS) with a 3D printer. All its components are depicted in Fig. 21. Since the IPM is connected to the crank of the slider-crank mechanism, the piston will release the drug from the reservoir when the IPM is rotated around the z axis by the external magnetic field (Fig. 22). The crankshaft torque T to balance the discharge force F can be expressed as follows [20]: T = FRsina(l +\ufffdcosa) (2) L By using the law of cosines, we can also express the crank angle, a, as a function of x (i.e., the position of point B) as _ (R Z +X Z _L Z ) a = cos 1 (3) 2Rx Due to practical reasons, we placed a helical spring to measure the piston force F by using the Hooke's law: F = Kilx (4) ilx represents the displacement of the spring and K is the stiffness of the spring. K was measured as l" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0001320_j.mechmachtheory.2020.103953-Figure6-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0001320_j.mechmachtheory.2020.103953-Figure6-1.png", + "caption": "Fig. 6. A tricycle: scheme and notations.", + "texts": [ + " 4 As consequence, it can be concluded that, in general, the IMT is \u201cefficient enough\u201d to be implemented in a general multibody-dynamics software and that its real efficiency can be greatly improved by combining it with suitable notations that depend on the set of mechanism types to analyze. The identification of these notations is out of the scope of this paper. 4. Case studies In this section, the IMT is applied to three case studies to better illustrate its effectiveness. Even though the technique is general, the three cases have been intentionally chosen simple enough for the reader to be easily able to follow and check all the steps. The first case ( Fig. 1 ) refers to a mechanism with mobile frame (i.e., with time-dependent constraints). The second case ( Fig. 6 ) refers to a mechanism with non-holonomic constraints. The third case ( Fig. 11 ) refers to a spatial mechanism usable as a pointing system, for instance, to orientate antennas [ 35 , 36 ]. 4.1. A 2-DOF mechanism with mobile frame Fig. 1 shows a two-DOF planar mechanism whose frame (link 1) is mobile with respect to an observer fixed to the Cartesian reference Ox 0 y 0 . This situation occurs when a mechanism is mounted on a ground or air vehicle. The mechanism of Fig. 1 controls the position of point P, which is fixed to link 5, with respect to link 1", + " The analysis of these diagrams reveals that - the obtained time-histories of the secondary variables respect the differential-calculus rules regarding the relationships among the time histories of a variable, of its first time derivative and of its second time derivative; - the same time-histories numerically (i.e., the maximum numerical difference is 1.5 \u00d710 \u221215 in Matlab R2019b, which uses 16 digits as default precision) coincide with those obtained by using formulas independently deduced with an algebraic manipulator. All these verifications prove the correctness of the formulas deduced with the proposed method. 4.2. A mechanism with non-holonomic constraints Fig. 6 shows a tricycle with a castor steering wheel that moves on a plane. It could be either the mobile base of a robot or a vehicle driven by somebody. The radii of the wheels have been chosen all equal for the sake of simplicity; r will denote wheels\u2019 radius. The Cartesian reference Ox 0 y 0 is fixed to the motion plane. With reference to Fig. 6 , link 5 carries the useful load; the pose of link 5 is located by the coordinates (x P , y P ) of point P, in Ox 0 y 0 , together with the angle \u03d5. \u03b3 is the steering angle. u 1 and u 2 are the unit vectors of the rear-wheels\u2019 axis and of the steering-wheel\u2019s axis, respectively; whereas, \u03b8 i , i = 1,2,3 is the rotation angle of the i th wheel, and it is positive if counterclockwise with respect to its axis\u2019 unit vector. w 1 ( w 2 ) is a unit vector parallel to the motion plane and perpendicular to u 1 ( u 2 )", + " (9) and (10) , provide \u239b \u239d \u02d9 x P \u02d9 y P \u02d9 \u03d5 \u239e \u23a0 1 = \u02d9 \u03b8c a 2 c \u03b3 + a 3 \u239b \u239c \u239d rc \u03d5 ( a 2 c \u03b3 + a 3 ) r s \u03d5 ( a 2 c \u03b3 + a 3 ) r s \u03b3 \u239e \u239f \u23a0 , \u239b \u239d \u02d9 x P \u02d9 y P \u02d9 \u03d5 \u239e \u23a0 2 = \u02d9 \u03b3 a 2 c \u03b3 + a 3 \u239b \u239d 0 0 \u2212a 3 \u239e \u23a0 (42) and ( \u02d9 x P \u02d9 y P \u02d9 \u03d5 ) = ( \u02d9 x P \u02d9 y P \u02d9 \u03d5 ) + ( \u02d9 x P \u02d9 y P \u02d9 \u03d5 ) = 1 a 2 c \u03b3 + a 3 \u239b \u239d \u02d9 \u03b8c r c \u03d5 ( a 2 c \u03b3 + a 3 ) \u02d9 \u03b8c r s \u03d5 ( a 2 c \u03b3 + a 3 ) \u02d9 \u239e \u23a0 (43) 1 2 \u03b8c r s \u03b3 \u2212 \u02d9 \u03b3 a 3 6 It is worth noting that, if the rear wheels are driven through a differential gear train [1] as it happens in the vast majority of ground vehicles, \u03b8 c is the rotation angle of the carrier and it is the variable directly controlled by the engine. 7 In ground vehicles, \u03b8 c and \u03b3 are usually the two variables directly driven by the actuators, and the pose of the chassis (i.e., link 5 in Fig. 6 ) has to be controlled. In general, wheels\u2019 angular velocities are not of interest. Anyway, they can be computed immediately by exploiting Eqs. (37c) , ( 37d ), and ( 37e ) after the pose parameters (i.e., x P , y P , and \u03d5) as functions of time have been computed. 4.2.1. Acceleration analysis In the acceleration analysis, formulas (41) immediately allow the computation of Eqs. (14a) , which yield \u239b \u239d x\u0308 P y\u0308 P \u03d5\u0308 \u239e \u23a0 1 = \u03b8\u0308c a 2 c \u03b3 + a 3 \u239b \u239c \u239d rc \u03d5 ( a 2 c \u03b3 + a 3 ) r s \u03d5 ( a 2 c \u03b3 + a 3 ) r s \u03b3 \u239e \u239f \u23a0 , \u239b \u239d x\u0308 P y\u0308 P \u03d5\u0308 \u239e \u23a0 2 = \u03b3\u0308 a 2 c \u03b3 + a 3 \u239b \u239d 0 0 \u2212a 3 \u239e \u23a0 (44) In this case, Eqs", + " \u03d5 \u239e \u239f \u23a0 5 = 1 ( a 2 c \u03b3 + a 3 )3 \u00d7 \u239b \u239c \u239c \u239c \u239d 0 0 ( a 2 c \u03b3 + a 3 )2 { r ( 2 \u0308\u03b8c \u02d9 \u03d5 + \u02d9 \u03b8c \u0308\u03d5 ) c \u03b3 ( a 2 c \u03b3 + a 3 ) + ( 2 \u0307 \u03b3 \u03b8\u0308c + \u02d9 \u03b8c \u0308\u03b3 ) r ( a 3 c \u03b3 + a 2 ) \u2212 3 \u0307 \u03b3 \u03b3\u0308 a 2 a 3 s \u03b3 + r \u0307 \u03b8c s \u03b3 \u02d9 \u03d5 2 } + +r \u0307 \u03b8c s \u03b3 \u02d9 \u03b3 2 ( a 2 a 3 c \u03b3 + 2a 2 2 \u2212 a 2 3 ) \u2212 \u02d9 \u03b3 3 a 2 a 3 ( a 2 + a 2 s 2 \u03b3 + a 3 c \u03b3 ) \u239e \u239f \u239f \u239f \u23a0 (55b) Eventually, Eq. (20) yields \u239b \u239c \u239d ... x P ... y P ... \u03d5 \u239e \u239f \u23a0 = \u2211 i=1 , 5 \u239b \u239c \u239d ... x P ... y P ... \u03d5 \u239e \u239f \u23a0 i (56) After the velocity, the acceleration, and the jerk analyses, the higher-order analyses with order greater than 3 follows the same scheme. 4.2.3. Numerical example The above-deduced formulas have been applied to a tricycle geometry defined as follows (see Fig. 6 ; the linear sizes are measured in a generic length unit (l.u.)): a 1 = 1 l.u., a 2 = 1 l.u., a 3 = 0.3 l.u., r = 0.25 l.u. . The initial conditions of motion, referred to the instant of time t = 0 s, have been assigned as follows: x P0 = 3 l.u., y P0 = 0 l.u., \u03d50 = ( \u03c0 /2) rad; whereas, the motion imposed by the input variables has been assigned as follows: \u03b8 c = \u03c9t rad, \u03b3 = ( \u03c0 /6) rad with \u03c9= (20/r) rad/s. The motion simulation has been implemented in Matlab R2019b and the results are reported in the diagrams of Figs" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0001931_j.jmmm.2020.167474-Figure14-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0001931_j.jmmm.2020.167474-Figure14-1.png", + "caption": "Fig. 14. The distribution of magnetic flux density obtained by frequencydomain FEM.", + "texts": [], + "surrounding_texts": [ + "The time-domain hybrid FEM-BEM was proposed for modeling of hysteresis motors by which, when the rotor rotates, the permeability discontinuities are avoided in the FEM equations, which includes the J-A hysteresis model of the rotor ring. Thereby, the numerical model takes into account the asynchronous operation of the hysteresis motor. The history state of the rotor magnetization can be observed, and the dynamic operation of the hysteresis motor was analyzed during the short period of over-excitation. Also, the steady-state behavior of the motor was studied using the proposed transient-time model and compared with those obtained by the frequency-domain FEM. The frequencydomain model uses the complex effective reluctivity to consider the magnetic hysteresis. Both models provide comparable results regarding the steady-state operation of the motor. However, the time-domain model requires much higher computational cost than the frequencydomain model. CRediT authorship contribution statement Behrooz Rezaeealam: Investigation, Conceptualization, Methodology, Software, Validation, Writing - original draft, Writing - review & editing." + ] + }, + { + "image_filename": "designv11_71_0000269_978-3-030-35699-6_43-Figure17-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0000269_978-3-030-35699-6_43-Figure17-1.png", + "caption": "Fig. 17. 3D diagram of a system with multiple coils [3]", + "texts": [ + " Moreover, in case of a non-deformable contact, the non-magnetic object would have to be very resistant to the very important instantaneous acceleration during the shock. In a further work, the structure of our coil gun will be optimized dividing the coil in several smaller coils activated sequentially. For example, scenarios using 2, 3 or 4 coils (each one having a half, a third or a quarter of the total turns of the initial coil), triggered sequentially by software or using a position sensor instead of a single coil as shown in Fig. 17 will be evaluated as in [2,7]. This will lead to have successively a maximum current on each coil when the rod is optimally placed in the coil, leading to a increased projectile speed. Eighteenth International Middle East Power Systems Conference (MEPCON), pp. 506\u2013511, December 2016. https://doi.org/10.1109/MEPCON.2016.7836938 2. Bencheikh, Y., Ouazir, Y., Ibtiouen, R.: Analysis of capacitively driven electromagnetic coil guns. In: The XIX International Conference on Electrical Machines - ICEM 2010, pp" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0001303_s00202-020-01035-1-Figure1-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0001303_s00202-020-01035-1-Figure1-1.png", + "caption": "Fig. 1 The geometrical representation of the investigated machine", + "texts": [ + " Vernier machines are generally classified into two discrete-pole and hybrid-pole categories according to the type of modulation. In the hybrid type, permanent magnets are used in the stator to modulate the magnetic field of the air gap. While in the discrete-pole design, the stator slots are used to modulate the magnetic field of the air gap. The design proposed in this study is a combination of these two types, which then introduces the characteristics of the structure and its parameters. A 2-D geometric model of the proposed internal rotor three-phase VPM machine is shown in Fig.\u00a01. The rotor of this design consists of a series of slots and teeth. Each rotor pole also includes two flux modulation poles (FMP). That is, the rotor has 16 FMP in total. Also within each rotor slot, there is a permanent magnet with radial magnetization orientation arranged N and S in succession. The stator structure also includes several slots for wiring and several slots for the installation of permanent magnets. Permanent magnets are mounted as spoke type, and their magnetic orientation is peripheral" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0003617_ipec.2014.6870118-Figure8-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0003617_ipec.2014.6870118-Figure8-1.png", + "caption": "Fig. 8. Wire connection of a DC decay testing to fJ-axis.", + "texts": [ + " The test procedure is as follows. First, a DC decay testing method is carried out under the connection of Fig. 7(a), then Za-hc(W) is calculated. Since magnetomotive force shown in Fig. 7(b) is alined with a axis, a-axis operational impedance XJjs) is determined as follows: x (\u00b7s)=\ufffd Za_bc(OJ)-ra_bc a J 3 . JS (22) where s is the slip (=w / utJ), utJ is the angular frequency of the power source (rad/s), and ra-bc is the winding resistance between a and bc terminals. Next, after changing the connection as shown in Fig. 8 (a), a DC decay testing is carried out, then Zh_c(W) is calculated. In this case, magnetomotive force occurs as Using the results mentioned above, the equivalent circuit parameters lb md, and rz of Fig. 9 (a) are identified by a least-squares method [6] [7] . The secondary side of the tested LIM consists of a secondary conductor of an aluminum plate with an iron plate. Therefore, one can determine as Iz = 0 from its physical image. A similar technique can be applied to identify the parameters I], mq, and rz of Fig" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0003020_gt2009-59108-Figure1-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0003020_gt2009-59108-Figure1-1.png", + "caption": "Fig. 1 Turbocharger rotor assembly with semi-floating ring bearing (SFRB) support", + "texts": [ + " Predicted total shaft motion is also in good agreement with test data for all engine loads and over the operating TC shaft speed range. The comparisons validate the rotor-bearing model and will aid in reducing product development time and expenditures Turbochargers enable smaller and more fuel-efficient passenger vehicle internal combustion (IC) engines with power outputs comparable to those of large displacement engines. In a turbocharger (TC), exhaust gases drive a turbine wheel which is connected to a compressor wheel by a thin shaft, as shown in Figure 1. The compressor forces a denser charge of air (i.e. ed From: http://proceedings.asmedigitalcollection.asme.org/ on 01/30/2016 Ter larger air mass) into the engine cylinders, improves combustion efficiency, thus producing more power. While simple in concept, TC design and operation are complex. As a TC rotor spins, it vibrates due to several causes (imbalance, engineinduced vibrations, fluid film bearing self-excitations, aerodynamic loading, pressure fluctuations, etc.). The vibration amplitudes need to be controlled to ensure TC reliability", + " In addition, unstable operation results when the external shock frequency is close to that of the natural frequency of the RBS. Lee et al. [19] also demonstrate both analytically and experimentally that vibration amplitudes are amplified when the external shock frequency is close to a system natural frequency. This paper presents progress on the refinement of the linear and nonlinear rotordynamic models developed in Refs. [9-13] to predict the response of a TC supported on a SFRB due to specified base excitations that simulate engine induced vibrations. Figure 1 displays the TC rotor assembly with the SFRB as a one-piece design integrating the compressor and turbine bearings at its ends. A button pin prevents ring rotation. Figure 3 Copyright \u00a9 2009 by ASME s of Use: http://www.asme.org/about-asme/terms-of-use Dow 2 shows the TC rotordynamic model consisting of 43 finite elements for the rotor, including the thrust collar and spacer, and 13 finite elements for the SFRB. Fig. 2 Turbocharger rotor and semi-floating ring bearing structural models Childs [20] details the theoretical aspects employed for rotor lateral bending and associated finite element modeling" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0000866_1475090220903217-Figure1-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0000866_1475090220903217-Figure1-1.png", + "caption": "Figure 1. General arrangement of gyrostabilizer.2", + "texts": [], + "surrounding_texts": [ + "Gyroscopic ship stabilization, HN performance, linear matrix inequality optimization, time-domain constraints\nDate received: 16 July 2019; accepted: 8 December 2019\nReduction of roll motion in marine vehicles is sought to prevent damage to cargo, to allow the crew to work efficiently, and to provide comfort for the passengers. The systems that have been developed to control undesired lateral movements can be classified as either external or internal systems. To reduce the ship movements, they can respond to the disturbance by either using a closed-loop feedback control system (active control) or relying only on the performance of the stabilization system in open-loop (passive control). There are several active devices for ship roll reduction, including fin-roll and rudder-roll stabilizers, gyroscopes, moving mass, and activated tanks. External motion control systems generate forces and torques outside the hull of the ship and usually rely on hydrodynamic interactions. Internal systems are most often used to generate forces and moments in the whole body, so as to stabilize mass moving moments.1\nThe main disadvantage of external systems such as fin-roll is their ineffectiveness at low speeds. For lowspeed or zero-speed state, internal devices such as gyrostabilizers are used. Gyrostabilizers use the property of the rotating flywheel to create a counteracting stabilizing torque against disturbance. Gyrodynamics established by the specific gimbaling arrangement means\nthat without any intervention, the vessel rolling motion combines with the flywheel angular momentum to cause oscillating precession motion (nutation). This motion combines with the flywheel angular momentum to create stabilizing torque, which directly opposes the wave-induced rolling motion. The earliest designs for ship gyrostabilization used passive devices. Dr. Schlick proposed such a passive control device for in 1904,2 where it was implemented with an electric motor and a brake to control the precession of the gyroscope. In this arrangement, the rate of precession was made proportional to the rate of roll of the ship. Further development involved twin gyros with precession in opposite directions to cancel the cross-torque that can be generated by a single gyro in maneuvers or pitching.3\nGyrodynamics requires no further intervention in order to function. However, in the case where the stabilizing torque created is less than the wave-induced rolling torque and hence rolling motion is not completely\nDepartment of Mechanical Engineering, Marmara University, Istanbul, Turkey\nCorresponding author: Sina Kuseyri, Department of Mechanical Engineering, Marmara University, Go\u0308ztepe, Istanbul 34722, Turkey. Email: sina.kuseyri@marmara.edu.tr", + "attenuated, active control is required in closed loop to manage the precession motion.4 When used as a passive stabilization device, one of the gyroscopic axes is fixed to the axis of the vessel about which the rolling motion occurs. The other axis of the gyroscope rotates freely and the moment causing the undesirable motion is reacted by the gyroscope resulting in a precession of the gyroscope about its free axis. Alternatively, using an active gyrostabilizer system, the gyroscope is forced to rotate about its free axis and this forced motion is referred to as nutation. Active gyrostabilization was first implemented in the marine context at model-scale by Sperry in 1908, and he was granted a patent for his active gyrostabilizer. Today, developments in materials, bearings, mechanical design techniques, and digital control systems have revived the interest in the gyrostabilizers. In fact, these provide improved bearing materials and custom electric motors of higher prices and the use of smaller gyrostabilizers to produce roll moments for high reduction of the torques created by the sea waves.5 However, in active stabilization, the precession oscillation range must be managed to within desirable mechanical design limits. Therefore, the challenge in design of active control system is to incorporate the precession angle and possibly control torque constraints in the closed-loop control synthesis.6\nIn this article, we suggest a constrained H\u2018 control scheme for gyroscopic marine vehicle stabilization systems with output and control constraints. The H\u2018 performance is used to measure the roll angle reduction of the vessel relative to wave disturbances in regular beam seas. Time-domain constraints, representing requirements for precession angle of gyroscopes and for actuator saturation, are captured using the concept of reachable sets and state-space ellipsoids. A state feedback solution to the constrained H\u2018 stabilization control problem is proposed in the framework of linear\nmatrix inequality (LMI) optimization and multiobjective control.\nThe article is organized as follows: In \u2018\u2018System dynamics and control\u2019\u2019 section, we first present the dynamic model of vessel rolling motion coupled with the gyrostabilizer. Then, we introduce the suggested constrained H\u2018 controller synthesis. In \u2018\u2018Numerical simulations and results\u2019\u2019 section, we provide numerical simulation results on stabilization performance. In \u2018\u2018Conclusion\u2019\u2019 section, we present our final comments.\nAssuming the surge forces and responses are small in comparison to other motions with a perfectly symmetrical port/starboard architecture, and further assuming that the sway and yaw effects on the roll motion are small, the roll motion can be uncoupled from the general ship motion dynamics and it can be expressed by a one-degree-of-freedom model. The incorporation of dual gyrostabilizer to this model leads to the following coupled nonlinear differential equations of motion7\n(I44 +A44)\u20acf+B44 _f+C44f=tw nIsvs _acosa \u00f01a\u00de It\u20aca+Bg _a+Cg sina= Isvs _fcosa+ tc \u00f01b\u00de\nwhere variables f and a are vessel\u2019s roll angle and gyrostabilizer\u2019s precession angle, respectively. Parameters I44,A44,B44, and C44 denote vessel\u2019s lateral moment of inertia, added mass moment of inertia, roll damping, and roll restoring, respectively. Parameters Is, It,Bg,Cg, and vs denote gyrostabilizer\u2019s polar moment of inertia, transverse moment of inertia, gyroscopic damping, gyroscopic restoring, and spin velocity, respectively. Forcing terms tw and tc are the rolling moment due to", + "waves and the control torque generated by the gyrostabilizer. Using a dual gyrostabilizer system (n=2), closed-loop control system dynamics is shown in Figure 2, where Kg = Isvs denotes the angular momentum of the gyrostabilizer in the spin direction.\nIf the magnitude of the precession angle jaj is small, equation (1) can be linearized and can be represented in state-space form\n_x=Ax+B1w+B2u \u00f02\u00de\nwith states x :\u00bc (fa _f _a)T, and with system matrices\nA=\n0 0 1 0\n0 0 0 1 C44\nI44 +A44 0 B44 I44 +A44 2Isvs I44 +A44\n0 Cg\nIt\nIsvs It Bg It\n0 BBBBBB@\n1 CCCCCCA ,\nB1 =\n0\n0 1\nI44 +A44\n0\n0 BBB@\n1 CCCA, B2 = 0 0\n0 1 It\n0 BBB@ 1 CCCA\nIn wave modeling, it is often assumed that sea states are statistically stationary, that is, neither growing nor decaying.4 By simulating a Joint North Sea Wave Project (JONSWAP) wave spectrum, a wave forcing disturbance can be approximated by the superposition of a collection of sinusoidal forcing terms. The JONSWAP spectrum describes a sea state by two parameters: the average significant wave height H1=3, and average wave period T. It is expressed as8\nSJz(ve)=0:658CSBz(v) g\ng 2vUcos(h) \u00f03\u00de\nwhere C=3:3J and J=exp\u00bd( 1=(2g2)) \u00bd(vT0=2p) 1 2 . In the exponent J of C, g =0:07 for v42p=T0 or g =0:09 for v . 2p=T0. SBz(v) represents the Bretschneider or International Towing Tank Conference (ITTC) two-parameter spectrum and T0 is\nthe modal period. Parameter h represents the heading angle and U is the vessel\u2019s forward speed.\nThe amplitude of an individual wave within the spectrum can be expressed as zn = \u00bd2SJz(vn)dv 1=2, and an irregular wave amplitude can be simulated as\nz(t)= XN n=1 zn cos(vnt+cn) \u00f04\u00de\nwhere vn and cn represent the frequency and phase angle of the nth wave component. Then, the forcing amplitude can be approximated by the Gaussian function\ntwn\nzn = tmffiffiffiffiffiffiffiffiffiffiffi 2ps2 p e (vn m)2=2s2 \u00f05\u00de\nwhere twn, zn, and vn represent the wave forcing (torque) amplitude, the wave amplitude, and the wave frequency of the nth wave component, respectively. Parameters tm, m, and s2 represent a forcing coefficient (representative of the wave torque per wave amplitude) and the mean and variance of wave frequency. As the wave torque per wave amplitude of each component wave can be determined by equation (5), an irregular disturbance torque can be determined by the superposition of individual sinusoidal forcing terms that can be expressed as\ntw(t)= XN n=1 twn sin(vnt+cn) \u00f06\u00de\nOur control objective is to find an internally stabilizing controller such that the H\u2018 norm from sea wave disturbance torque to ship roll angle is minimized while observing time-domain hard constraints arising from actuator saturation and gyrostabilizer\u2019s precession angle within bounds. Proposed state feedback control system can be formulated as follows" + ] + }, + { + "image_filename": "designv11_71_0001928_s00170-020-06182-0-Figure9-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0001928_s00170-020-06182-0-Figure9-1.png", + "caption": "Fig. 9 LPR and DP interaction effect onMDD for a low SPR and low T, b high SPR and low T, c low SPR and high T, and d high SPR and high T", + "texts": [ + " Apparently, this LPR value was large enough for the solid part of the feedstock to cause loss of contact with the melt, and besides, it cooled down due to its displacement towards cooler nozzle region during retraction. Thus, when priming occurs, there is no sufficient time for the filament to absorb the thermal energy required to return to extrudable state in time, due to low temperature setting in the hot end, and due to high SPR, too. To sum up, the interacting effect between LPR and DP is depicted in Fig. 9. Generally, low LPR with high percentage of DP or high LPR with low percentage of DP should be combined in order to avoid unfilled region at the beginning of the deposition. This is indicated by the darkest regions in the plot. Furthermore, it is shown that increasing SPR (Fig. 9b, d) and decreasing temperature (Fig. 9a, b) limit the effective range (darkest regions) between LPR and DP pointing to avoidance of material discontinuity. The rest of the factors were kept at middle setting for all plots in Fig. 9. Once uninterrupted deposition is ensured, the next objective is to minimize SWD. To measure SWD, strand area and length calculations from image analysis were used to define mean strand width W as follows: W \u00bc A L \u00f04\u00de where A is the strand area and L is the strand length. Then, two localized measurements at the beginning and at the end of the strand were performed in Fiji ImageJ\u2122 in order to evaluate start-stop artifacts (Fig. 10). Standard deviation of all three measurements was registered as the SWD response that must be minimized in order to achieve uniform material deposition", + " Their prediction ability may be limited, but they still serve as a reliable roadmap with respect to the factors and their effect on material deposition quality. Therefore, specific improvement steps have been identified and are used as guidelines: & LPR is set to minimum since it is proven beneficial both for deposition continuity and uniformity. & Because of the low LPR setting, SPR effect is minimized in practice, since there is not enough time for the extruder motor to reach its target angular speed. & According to Fig. 9a, c, DP efficient range (black region) widens as T increases. For T = 195 \u00b0C, DP = 50 is a reasonable starting point. For higher temperature, reduced DP may also suffice. & Although high SR setting favors overfills, it cannot be reduced to zero because of the stringing that occurs during nozzle travel towards the next start point. For T = 195 \u00b0C, a mid-range SR value is a good starting point. & When DtE is set to maximum tested distance, continuous deposition is aided, especially at lower melt temperature" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0000763_s00202-020-00967-y-Figure1-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0000763_s00202-020-00967-y-Figure1-1.png", + "caption": "Fig. 1 Single-phase stack: a TFM; b HFM", + "texts": [ + " This HFM has the advantages of TFM (geometrical decoupling of magnetic and electric circuits and low ohmic losses) and the advantage of LFM (simple rotor construction with single magnet layer per stack, better magnet utilization). The proposed HFM topology is a structurally modified version of TFM topology. So, the performance of HFM is compared with that of TFM. In the literature, two topologies of TFM, without pole shaping [7,9,18] and with pole shaping [16,19], are reported. However, the comparison study between these two topologies is not available. Hence, both TFM topologies (without and with pole shaping) are considered to compare with HFM. The single-phase stack of TFM and the proposed HFM is shown in Fig. 1, and its exploded view is given in Fig. 2. Three-phase topology with outer rotor configuration is considered in this paper. The complete machine can be formed by combining three single-phase stacks with 120\u25e6 electrical angle between them. Each phase stack consists of p number of pole pairs. A 6-pole pair design is considered in this section to explain the structure of machines. Soft magnetic composite (SMC) is assumed as the core material to avoid the lamination difficulties in the manufacturing of TFM and HFM. In a single-phase stack of TFM, the stator has 12 poles, and the rotor has 24 poles. The stator is a hollow-shaped cylinder with 6 poles projected outwards at each side of the cylinder as shown in Fig. 2a. A ring-shaped coil is placed between the projected poles. The rotor has two layers of surface-mounted magnets with 12 magnets in each layer. The span between the stator poles is twice of the span between the rotor poles as shown in Fig. 1a. Due to this, only half the number of magnet\u2019s flux can link with the stator coil at maximum. In HFM, the number of poles in stator and rotor is 12. The stator is made of a hollow cylinder core with 6 outwards projected poles (I-core) at themiddle of the cylinder as shown in Fig. 2b. At both sides of the projected poles, a ring coil is inserted, and in between two I-cores, a C-core is placed. Also, both the coils are enclosed by C-cores. This C-core can be supported at both ends by non-magnetic frames" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0000241_012039-Figure2-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0000241_012039-Figure2-1.png", + "caption": "Figure 2. The design scheme for the representation of a hydraulic manipulator with a recuperative hydraulic actuator in a mathematical model.", + "texts": [ + " Within the framework of the mathematical model of the mechanical subsystem of a hydraulic manipulator with a regenerative hydraulic drive, the main physical processes occurring in its mechanical subsystem are described. The structure of the hydraulic manipulator is presented in a mathematical model by a system of algebraic and differential equations. The solution of this system of equations is carried out on the basis of numerical integration. Despite the fact that the boom, as a mechanism, consists of two links, in the mathematical model, the assumption is that the boom is a single link and is a single solid body (figure 2). In this case, in the Forestry 2019 IOP Conf. Series: Earth and Environmental Science 392 (2019) 012039 IOP Publishing doi:10.1088/1755-1315/392/1/012039 mathematical model, it is considered that the assortment is lifted only by means of a hydraulic cylinder located between the column and a long link of the boom, and the hydraulic cylinder between the long and short links of the boom holds them in a given angular position relative to each other. In the mathematical model of the mechanical subsystem of a hydraulic manipulator using the equations of classical dynamics are described: the movement of the boom relative to the column, the rotation of the column (absolutely solid) relative to the vertical axis Z, the swinging of assortment D relative to the attachment point of the grapple on the arrow of the hydraulic manipulator C", + " The recuperative hydraulic actuator can significantly reduce the sharp fluctuations in the pressure of the working fluid due to its direct direction to the pneumatic-hydraulic accumulator from the braking hydraulic cylinder. The moment created by a pack of assortments is determined by: cossin GYGXGG FFLM , (8) where FGX and FGY \u2013 \u0441artesian components of the force acting from a bundle of assortments on the Forestry 2019 IOP Conf. Series: Earth and Environmental Science 392 (2019) 012039 IOP Publishing doi:10.1088/1755-1315/392/1/012039 boom of a hydraulic manipulator; LG \u2013 the distance between the attachment point on the grapple boom and the axis of rotation of the hydraulic manipulator column (figure 2). As a result of this, the equation of rotational motion of the hydraulic manipulator column will take the following form: 22 2 2 2 1 sin cos 1 1 4 3 2 G gw L P G GX GY fr vfr sl w C G cyl yl Dd d R P P L F F M k M M dt dtm L m R . (9) Similarly, the rotation of the boom \u041e\u0421 in a vertical plane relative to the hinge O is described. The main difference is the calculation of the moment relative to the hydraulic cylinder and the absence of the moment from the slope of the bearing surface" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0000659_978-981-15-1616-0_21-Figure15-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0000659_978-981-15-1616-0_21-Figure15-1.png", + "caption": "Fig. 15 Velocity contours", + "texts": [ + " 12 Zoomed locations of max VONM stress in concrete (case 2) Case (3): InternalWaterPressure+Steel Shell+Concrete+RockSurrounding Figures 13 (cut section) and 14 (zoomed view) show that the maximum von Mises stress of 177.7 MPa in the stiffener provided on the steel pipe. The exact location of maximum von Mises stress in the steel pipe is at the internal surface of the pipe near the junction of branching of pipes as also observed in cases (1) and (2). Themaximum von Mises stress is well below the yield stress limiting value of 550 MPa mainly due to enormous rock mass providing both the required strength and stiffness. Figure 15 shows the cut section view, and Fig. 16 shows the zoomed view of the stress contours in the concrete region and themaximumvonMises stress is 16.0MPa. Figure 17 show the cut section view, and Fig. 18 shows the zoomed view showcasing the location of maximum von Mises stress contour in the rock region. The maximum von Mises stress is found to be 9.98 MPa in the rock region. The location of maximum stress in concrete and rock regions is occurring in the neighbouring region of steel pipe where the maximum stress is observed. Fig. 15 Locations of max VONM stress in concrete (case 3) Fig. 17 Locations of max VONM stress in rock (case 3) Fig. 18 Zoomed locations of max VONM stress in rock (case 3) Table 1 An overview of the location of and magnitude of maximum von Mises stresses S. No. Location Von Mises (MPa) Case (1) Case (2) Case (3) 1 Internal surface of piping at the branching junction 2741.99 934.17 177.7 2 Internal surface of concrete at the branching junction \u2013 68.07 16.0 3 Internal surface of rock at the branching junction \u2013 \u2013 9" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0003177_aim.2015.7222793-Figure1-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0003177_aim.2015.7222793-Figure1-1.png", + "caption": "Fig. 1. Coordinate frame arrangement of an AUVMS", + "texts": [ + " The paper is organized in the following sequence. Section II describes dynamic modeling of an AUVMS. Section III presents a proposed robust nonlinear controller design scheme. Performance analysis of proposed control scheme to an AUVMS are presented in Section IV, followed by the conclusion in section V. 978-1-4673-9107-8/15/$31.00 \u00a92015 IEEE 1713 In this work, dynamic modeling of an AUVMS is derived using Newton - Euler and recursive Newton - Euler formulation schemes [8], [9], [10]. The frame assignment of the AUVMS is presented in Fig. 1, where [xt yt zt] is the endeffector frame, [xI yI zI ] is the Earth-fixed reference frame and [xb yb zb] denotes the body-fixed frame attached to body of AUV. The dynamic equations of motion of an AUVMS can be expressed as follows: M(q)q\u0308 + C(q, q\u0307)q\u0307 +D(q, q\u0307)q\u0307 +G(q) + F (q, q\u0307) = \u03c4 + d (1) where, M(q) = [ Mv HT (\u03b6) H(\u03b6) Mm(\u03b6) ] , (2) C(q, q\u0307) = [ Cv(\u03b7, \u03b7\u0307) 0 0 Cm(\u03b6, \u03b6\u0307) ] , (3) D(q, q\u0307) = [ Dv(\u03b7, \u03b7\u0307) 0 0 Dm(\u03b6, \u03b6\u0307) ] , (4) F (q, q\u0307) = [ Fv(q, q\u0307) Fm(q, q\u0307) ] , (5) G(q) = [ Gv(\u03b7) Gm(\u03b6) ] , \u03c4 = [ \u03c4v \u03c4m ] , (6) q = [\u03b7 \u03b6] with \u03b7 = [x y z \u03c6 \u03b8 \u03c8] is the vector of absolute positions and Euler angles (roll, pitch and yaw)" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0001579_0954407020945823-Figure5-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0001579_0954407020945823-Figure5-1.png", + "caption": "Figure 5. (a) Structural diagram of join keys and (b) force analysis of join keys.", + "texts": [ + " The gear actuator successfully completing the gearshifting operation is related to the driven gear and the driven shaft being smoothly connected. Therefore, the join keys are the design focus of the automatic gear shifting actuator. Design and analysis of join keys The join keys on the driven shaft are the key component in transmission and shifting. The structural design determines that gear shifting can be smoothly performed. The force condition of the join key, which is in steady contact transmission state, is analyzed. The structural diagram and the force analysis of the connecting join keys are shown in Figure 5(a). Let join keys and the driven shaft be the design objects, and analyze force F1 T1 =9550 P n \u00f01\u00de where T1 is the input torque, P is the input power, and n is the input speed. Thus, the following equations are obtained T1 = 9550P n T2 = F1r 1000 T1 T2 = n2 n1 = i T2 = 9550P ni \u00f02\u00de where T2 is the torque subjected to the object, i is the transmission ratio, F1 is the force of the driven gear on the join keys, r is the distance from the point of action to the center of rotation, and a is the angle between the plane and the top plane of the key. After the equations are solved, the formula for calculating the reaction force of the bond block can be obtained N1 =F1sina= 9550Psina3103 nir \u00f03\u00de Static analysis is performed with the join keys as the object. The force is N1, N2, as shown in Figure 5(b). The friction is f1, f2. The centrifugal force is F1, and the pressure is PS. These equations meet the following relationship F1 +PS= f1sina+N1cosa+ f2 N2 =N1sina+ f1cosa F1 = mp2n2r 0 93105 f1 =N1m f2 =N2m \u00f04\u00de After these equations are solved, pressure P on the bottom of the join key is obtained as follows P= 1 S 95503103Psina nir (2msina+ cosa+m2cosa) mp2n2r1 93105 \u00f05\u00de From the above equations, it is obtained that when the force value at the bottom of the join key is constant, the size of the bottom cylindrical surface is inversely proportional to the pressure value it receives" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0001701_0954405420949757-Figure5-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0001701_0954405420949757-Figure5-1.png", + "caption": "Figure 5. Auxiliary circles.", + "texts": [ + " Unit normal vectors of face cone at P0i and P00i are denoted by c0i and c00i , which are calculated as c0i = f0i 3 (r03 f0i)= f 0 i 3 (r03 f0i)j j c00= f00i 3 (r03 f00i)= f 00 i 3 (r03 f00i)j j \u00f015\u00de Let t0i and t00i be unit tangent vectors of tooth crest line at P0i and P00i , which are calculated as follow t0i = c0i 3 n0i= c 0 i 3 n0ij j t00i = c00i 3 n00i= c 00 i 3 n00ij j \u00f016\u00de c0i, c 00 i , n 0 i, n 00 i , t 0 i and t00i are shown in Figure 4. Upper and lower edges calculation for chamfering Auxiliary circles for edges calculating Let the plane through P0i perpendicular to t0i be planeP0 i, the plane through P00i perpendicular to t00i be planeP00 i . Auxiliary circle R0i is drawn on P0 i with P0i as its center and r0i as its radius, and auxiliary circle R00i is drawn on P00 i with P00i as its center and r00i as its radius, as shown in Figure 5. Let the intersection of R0i and convex side of tooth surface be S0i, R 0 i and face cone be A0i,R 00 i and concave side of tooth surface be S00i , R 00 i and face cone be A00i . E 0 i is a point on R0i when angle between P0iE 0 i ! and c0i is f0i, E00i is a point on R00i when angle between P00iE 00 i ! and c00i is f00i . Let the unit vectors of P 0 iE 0 i ! and P00iE 00 i ! be e0i and e00i , which are calculated as e0i = c0i cosf0i + i(t 0 i 3 c0i) sinf0 e00i = c00i cosf00i + i(t 00 i 3 c00i) sinf00 \u00f017\u00de O1E 0 i " + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0002759_b978-0-12-804560-2.00011-0-Figure4.4-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0002759_b978-0-12-804560-2.00011-0-Figure4.4-1.png", + "caption": "FIGURE 4.4 Linear RWP models. (A) Left: Ankle joint as passive joint. (B) Middle: The linear RWP-foot model is identical to the LIP when the RW torque is zero. (C) Right: A nonzero RW moment mC alters the direction of the GRF f r so that its line of action deviates from the CoM. The centroidal (reaction) force f C is in the same direction as the GRF but its line of action passes through the CoM. The application point of f C defines the centroidal moment pivot (CMP) [40,120].", + "texts": [ + " Controllable angular acceleration/deceleration of the RW induces a variation in the joint rate of the passive pivot joint, via inertial coupling. In this way, the stability of the system can be improved. The RWP was introduced in the beginning of the century [138] and studied extensively as a landmark underactuated example in nonlinear control [12]. The RWP has been adopted in humanoid robotics as well, e.g. as a means to determine appropriate foot locations for a reactive-step strategy, the so-called \u201cCapture point\u201d [121]. As with the LIP model, the leg length of the RWP is made variable (cf. Fig. 4.4A). The general form of the equation of motion for such system is [ Mp Mpa MT pa Ma ][ q\u0308p q\u0308a ] + [ cp ca ] + [ gp ga ] = [ 0 Fa ] . (4.19) Here, qp = \u03b8, qa = [ l \u03c6 ]T denote the generalized coordinates, Fa = [ fs mc ]T is the generalized force, Mp = Ml2, Mpa = [ 0 I ] , Ma = diag [ m I ] are the inertia matrix compo- nents, I denotes the RW moment of inertia, cp = 2Mll\u0307\u03b8\u0307 , ca = [ Ml\u03b8\u03072 0 ]T are the nonlin- ear velocity-dependent forces, and gp = \u2212Mgl sin \u03b8, ga = [ Mg cos \u03b8 0 ]T are the gravitydependent forces", + " As already explained, according to the partial-feedback linearization approach, the (linear) inertia coupling for the passive coordinate (pendulum rotation \u03b8 ) is essential for stability. As apparent from the coupling inertia matrix (row matrix Mpa), the coupling is realized via the RW moment I \u03c6\u0308. To impose the task-based LIP constraint, rewrite the equation of motion in Cartesian coordinates, i.e. Mx\u0308g = |fs | sin \u03b8 \u2212 mc l cos \u03b8 = fcx, M(z\u0308g + g) = |fs | cos \u03b8 + mc l sin \u03b8 = fcz, (4.20) I \u03c6\u0308 = mC. (4.21) The vector f C = [ fcx fcz ]T , thus defined, acts in the same direction as the GRF vector f r (cf. Fig. 4.4C). The application points however are different. This follows from the fact that the line of action of the GRF does not pass anymore through the CoM; the direction of f r depends on the centroidal moment, mC . The line of action of f C , on the other hand, always passes through the CoM. The point of application of this vector will be clarified in short. Next, apply constraint z\u0307g = z\u0308g = 0 to reduce the dimension of the system, i.e. x\u0308g = \u03c92xg \u2212 1 Mz\u0304g mC, (4.22) \u03c6\u0308 = 1 I mC, where \u03c9 = \u03c9 = \u221a g/z\u0304g ", + " The above model facilitates the design of control laws for increased balance stability of the LIP-on-foot model since the CoP can be manipulated with the rate of change of the centroidal angular momentum l\u0307cy = I \u03c6\u0308, in addition to the CoM acceleration. This becomes apparent from the CoP equation. We have xp = xg \u2212 x\u0308g \u03c92 \u2212 l\u0307cy Mg (4.23) = xg \u2212 frx frz z\u0304g \u2212 mC frz . When the RW is nonaccelerating, i.e. stationary or spinning with constant angular velocity, then \u03c6\u0308 = 0 and the relations are the same as with the LIP-on-cart model (cf. Fig. 4.4B). An accelerating RW, on the other hand, alters the direction of the reaction force f r (cf. Fig. 4.4C). This means that the point on the ground where the moment of the reaction force is zero may be outside the BoS. This point cannot be referred to as the CoP (or ZMP) anymore since there will be a contradiction with the definition of the CoP or ZMP [158]. To alleviate the problem, new terms were introduced as follows. In [40], the point was referred to as the zero rate of change of angular momentum point, while in [118] the zero spin center of pressure was used. Another term has become commonly accepted, though: centroidal moment pivot (CMP) [120]" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0003139_j.ymssp.2015.07.017-Figure1-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0003139_j.ymssp.2015.07.017-Figure1-1.png", + "caption": "Fig. 1. Tested mechanical system.", + "texts": [ + " In the case of quasi-SEA systems or for general ED systems the situation becomes more intricate and it should be checked at each occurrence that Xijo0, to guarantee that a TPA is feasible. Yet it is to be noted that in the case of low modal overlap and global modes X could hardly be an M-matrix. Actually, for the limiting case of energy equipartition A becomes singular so that X does not even exist [20]. In this section, an energy transmission path analysis is performed for the ED model of a simple mechanical system inspired on that in [33]. The system is depicted in Fig. 1, and consists of 6 steel plates (density \u03c1\u00bc 7800 kg=m3, Young modulus E\u00bc 2:1 1011 Pa, Poisson's ratio \u03bc\u00bc0.3125), whose dimensions are listed in Table 1. The damping loss factor of every plate has been taken as 0.05. An ED model of the system has been built using a finite element approach with the commercial software VA-ONE. Subsystems have been identified with the physical plates in Fig. 1. To obtain each EIC Aij in A, the average energy at plate i resulting from rain-on-the-roof excitation of plate j has been calculated. A minimum of 6 elements per wavelength have been considered for the computations in all subsystems. The analysis has been carried out for 20 bands ranging from 100 Hz to 2000 Hz, with a constant bandwidth of 100 Hz. For illustrative purposes and to avoid unnecessary data overload, only the results for two representative frequency bands will be presented. In particular, we have chosen the bands with central frequencies 500 Hz and 2000 Hz", + " In fact, it turns out for the present case that not only SEA cannot be applied to it, but that the ED model is not even a quasi-SEA model (it can be checked that indirect coupling exists between the entries of the inverse EIC matrix X, and that X does not satisfy the CPP relation). However, it follows that Xijo0 so X turns to be anM-matrix. Therefore, a transmission path analysis can be carried out for this ED model and we can associate an ED graph to it, whose adjacency matrix is given by the transpose of the generating matrix F . The ED graph corresponding to the mechanical system in Fig. 1 is presented in Fig. 3. For simplicity we have only plotted one edge between linked subsystems but each edge has two side end arrows indicating that the connection exists in both directions (ED and SEA graphs are strongly connected graphs). It can be observed that several indirect couplings exist, which have been indicated using dashed lines. In order to perform a transmission path analysis making use of this ED graph, we have applied the algorithm in [3] to compute and rank all dominant paths between the input subsystem #1, and the target subsystem #6" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0001219_icemi46757.2019.9101620-Figure3-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0001219_icemi46757.2019.9101620-Figure3-1.png", + "caption": "Fig. 3 UR5 structure diagram", + "texts": [ + " By translating and rotating, one coordinate system can be transformed to the next. According to the transformation order of i i i id a , the corresponding transformation matrix is expressed as (1). 1 ( ) Rot ( ) ( ) Rot ( cos sin cos sin sin cos sin cos cos cos sin cos = 0 sin cos 0 0 0 1 ) i i i i i i i i i i i i i i i z i z i x i x i i i ia a a T Trans d Tra d ns (1) Where, i-1 i T represents the transformation matrix between adjacent coordinate systems. In this paper, the UR5 manipulator is used. Its physical map and structural diagrams are shown in Fig.2 and Fig.3. 435 Authorized licensed use limited to: UNIVERSITY OF WESTERN ONTARIO. Downloaded on July 26,2020 at 11:43:10 UTC from IEEE Xplore. Restrictions apply. The D-H parameters of UR5 are shown in table . Table . The D-H parameters of UR5 Joint i id (mm) ia (mm) i (\u00b0) i (\u00b0) 1 1 89.46d 0 1 2 0 2 425.00a 0 2 3 0 3 392.25a 0 3 4 4 109.15d 0 4 5 5 94.65d 0 - 5 6 6 82.30d 0 0 6 From the D-H parameters of UR5 manipulator, the transformation matrix i-1 i T of each joint of the manipulator can be obtained. Therefore the transformation matrix from the end to the base of manipulator is as (2)" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0001821_j.matpr.2020.08.097-Figure1-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0001821_j.matpr.2020.08.097-Figure1-1.png", + "caption": "Fig. 1. Scheme of roll-forming of tube attachment assembly into the tubesheet: 1 is rolle", + "texts": [ + " The obtained relations are valid under the following assumptions: Loading the body is simple Tube material is incompressible hardening of the material when creating a load does not occur Such a model cannot be taken as the basis for studying the stress\u2013strain state of the tube and tubesheet during roller rolling, since the rolling process should be considered taking into account the diverse interaction of the contacting surfaces at each loading point in conjunction with the contact problem [16]. Fig. 1 shows the tube mount and the rolling scheme, and Fig. 2 shows the design scheme loading. As can be seen from them, such loading is complex. The tight connection of the pipe with the surface of the hole of the tubesheet is ensured by the spindle torque and indentation of the roller in a certain vicinity of the contact point of the tube and roller. In the zone of action of the roller, the metal of the tube becomes plastic when part of the layers is shifted in the direction of motion of the surface of the roller, and part is stretched and compressed, pressing against the wall of the hole" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0003750_j.ast.2014.04.013-Figure2-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0003750_j.ast.2014.04.013-Figure2-1.png", + "caption": "Fig. 2. Characteristics velocities and reference frame.", + "texts": [ + " This section describes the full nonlinear six-degree-of-freedom model of the glider. The glider, shown in Fig. 1, is an ASH-26 E, a self-launching powered sailplane whose main characteristics are W = 5150 N, b = 18 m, S = 11.68 m2 [13]. The equations of motion of the rigid vehicle are V\u0307 = F m + g \u2212 \u03c9 \u00d7 V (1) \u03c9\u0307 = J\u22121(M \u2212 \u03c9 \u00d7 J\u03c9) (2) where F and M are the aerodynamic force and moment F = 1 2 \u03c1V 2 S \u23a1 \u23a3 C X CY C \u23a4 \u23a6 ; M = 1 2 \u03c1V 2 S \u23a1 \u23a3 Clb Cmc C b \u23a4 \u23a6 , (3) Z n I is the unit tensor, and, with reference to Fig. 2, V and \u03c9 are aircraft velocity and angular velocity, respectively, both expressed in body frame as V = V V\u0302, V\u0302 = \u23a1 \u23a3 u\u0302 v\u0302 w\u0302 \u23a4 \u23a6 \u2261 \u23a1 \u23a3 cos\u03b1 cos\u03b2 sin\u03b2 sin\u03b1 cos\u03b2 \u23a4 \u23a6 , \u03c9 = \u23a1 \u23a3 p q r \u23a4 \u23a6 (4) where u\u0302, v\u0302 and w\u0302 are the direction cosines of V, whereas p\u0302 = pb 2V , q\u0302 = qc 2V , r\u0302 = rb 2V (5) denote the dimensionless components of the angular velocity. The kinematics of the rigid vehicle is described, as usual, by the relationship between the rates of the Euler\u2019s angles and the angular velocity [9] \u03d5\u0307 = p + sin\u03d5 tan\u03d1 q + cos\u03d5 tan\u03d1 r \u03d1\u0307 = cos\u03d5 q \u2212 sin\u03d5 r \u03c8\u0307 = sin\u03d5 sec\u03d1 q + cos\u03d5 sec \u03d1 r (6) and by the navigation equations x\u0307 = (cos\u03d1 cos\u03c8) u + (sin\u03d5 sin\u03d1 cos\u03c8 \u2212 cos\u03d5 sin\u03c8) v + (cos\u03d5 sin\u03d1 cos\u03c8 + sin\u03d5 sin\u03c8) w y\u0307 = (cos\u03d1 sin\u03c8) u + (sin\u03d5 sin\u03d1 sin\u03c8 + cos\u03d5 cos\u03c8) v + (cos\u03d5 sin\u03d1 sin\u03c8 \u2212 sin\u03d5 cos\u03c8) w z\u0307 = (\u2212 sin\u03d1) u + (sin\u03d5 cos\u03d1) v + (cos\u03d5 cos\u03d1) w (7) As the dynamical model needs the value of J, the measurements of the moments of inertia in roll, pitch and yaw and of the inclination of the principal inertia axis are carried out", + " This shows that the inertia moments could be measured with an absolute error of 3%, 3% and 2% for the roll, pitch and yaw axes, respectively, whereas the inclination of the principal axes could be determined with an error less than 0.7 deg. This section summarizes the main results obtained in Ref. [6], where the vehicle aerodynamics is studied through the continuity equation for incompressible flow. This equation, linear with respect to the local fluid velocity v, states that, if the flow structure about the aircraft remains unaltered as \u03b1 and \u03b2 change, v is a linear function of the characteristic velocities of the rigid vehicle [15], which are V and \u03c9 (see Fig. 2). Aerodynamic force and moment are expressed according to the Tait\u2013Kirchhoff equations F = \u2212 d dt ( \u2202T \u2202V ) \u2212 \u03c9 \u00d7 \u2202T \u2202V , M = \u2212 d dt ( \u2202T \u2202\u03c9 ) \u2212 V \u00d7 \u2202T \u2202V \u2212 \u03c9 \u00d7 \u2202T \u2202\u03c9 (8) where T is the fluid kinetic energy T = 1 2 V \u00b7MV + V \u00b7 S\u03c9 + 1 2 \u03c9 \u00b7J\u03c9 + Tw (9) in which Tw is the part of fluid kinetic energy due to the wakes and responsible for the lift, and M, J and S are the so called aircraft apparent mass tensors, second order tensors proportional to the fluid density. When dV/dt and d\u03c9/dt are both assigned, M, J and S are given tensors" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0003943_icra.2014.6907115-Figure1-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0003943_icra.2014.6907115-Figure1-1.png", + "caption": "Fig. 1. Illustration of the 2D FPE [12].", + "texts": [ + " These different approaches are explained in further detail below to help justify the choice of the 3D FPE approach to calculate the FPE in 3D, to be used as a control input to a high level state-machine based controller with online tasklevel trajectory generation and a prioritized Jacobian-based feedback loop, as described in the next section. The original 2D FPE concept was developed for planar bipeds in [10], with a point mass at the system\u2019s Center of Mass (COM), point feet, and massless legs. The 2D FPE is illustrated in Fig.1. The FPE is computed using the conservation of angular momentum in the 2D plane, by finding the angle \u03c6 where the total energy of the biped after swing foot impact is equal to the peak potential energy. In both [13] and [14], extensions of the 2D FPE concept are made to take 3D dynamics into account. In both approaches, the 3D dynamics are projected into a vertical plane which passes through the COM, but the computation of the plane orientation about the vertical axis differs between the two methods" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0001055_012011-Figure3-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0001055_012011-Figure3-1.png", + "caption": "Figure 3. Magnetic Flux Distribution of Model of PMSG studied a fractional slot of 8 slot /18 PMSG", + "texts": [ + " Series: Materials Science and Engineering 807 (2020) 012011 IOP Publishing doi:10.1088/1757-899X/807/1/012011 IOP Conf. Series: Materials Science and Engineering 807 (2020) 012011 IOP Publishing doi:10.1088/1757-899X/807/1/012011 The simulation results [9], provide that the slotting design in the magnet edge can effectively reduce the cogging torque of the inset permanent magnet synchronous machine. The peak value of the cogging torque was obtained as much as 92.20% compared with the experimental model. From the brief analysis of the Fig.3, 4 and 5 a substantial improvement of the peak cogging torque in fractional system combined with smaller slot opening width and higher of stator shoe in the PMSG is evident. For fractional machine (8 pole/18 slot) the cogging torque can be reduced by using any higher stator shoe with smaller slot opening width. However, the wave of cogging torques for all design are unbalance, it means that peak of negative and positive of the cogging torque are different value, except for PMSG structure with slot opening width of 1", + "3 mm it is can be accepted as the best structure. The simulation of fractional structure of 8 pole/18 slot has been presented in this paper. It shows that by using factional structure of PMSG can reduce the cogging torque effectively. However, most of the PMSG structure simulated in this research show that their peak of both negative and positive values of the cogging torque are different. These refer to the effect the static unbalance magnetic pull (UMP). In the simulation result (as shown in Fig.3), the peak value of cogging torque in the structure is 0.00062 N-m for positive value and about -0.00064 N-m for the negative value. And other most important is that structure is not deteriorated with core saturation, the fact that is very important for the smooth of voltage waveform of the generator. It can be concluded that this typical of slot opening width of 1.55 mm and shoe height of 2.3 mm is the best structure of the fractional 8 slot /18 poles of PMSGs, because it has any significant low of cogging torque and least of unbalance magnetic pull affect" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0003286_9781118869796.ch16-Figure16.3-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0003286_9781118869796.ch16-Figure16.3-1.png", + "caption": "FIGURE 16.3 (Top) Modular EFC research hardware. (Bottom) Optimized EFC prototype with bipolar electrodes. (Left) Exploded drawing. (Right) Photograph of EFC stack assembly. (Reproduced with permission from Ref. [16]. Copyright 2010, The Electrochemical Society.)", + "texts": [ + " Additionally, in conventional fuel cells, the oxygen feed is sometimes humidified for the best catalytic performance. However, feeding oxygen to EFCs imposes critical technical, safety, and cost burdens on theEFCpower system.As a result, an energetically passive air supply leveraged from other systems provides better technical justification. A flooded cathode, which uses dissolved oxygen in the fuel solution, should be avoided due to lowoxygen supply and thus lowpower performance. An exploded view schematic image of a modular EFC research cell is shown in Figure 16.3 (top left) and features (i) an-air breathing oxygen reduction half-MEA cathode, (ii) a GOx-based anode fabricated on a carbon felt internal current collector, and (iii) a stainless steel plate anode and cathode external current collector. The EFC is shown in a three-electrode configuration with a port for the reference electrode and a pair of ports for filling the anode with fuel. Inmicrofluidics-basedEFCs, the fuel andoxidizer are supplied in twoparallel colaminar streams in a single channel, thus avoiding the need for a membrane", + " As described, a reference electrode might be integrated with the EFC so that a cutoff voltage limit can be defined precisely and independently for both electrodes. In general, the application circuitry can be modified to accommodate a lower cutoff voltage, and the initial voltage conditioning and the cutoff limit may be handled and controlled with a DC/DC converter at the penalty of increased system cost and a substantial power loss. An EFC stack composed of five modular research hardware single-cell EFCs is shown inFigure 16.3 (top left). The anode andcathode contactors of each cell are visible on the top and the individual cells are electrically insulated; the intercell electric connection is provided with external U-shape leads. This allows for series or parallel connection of the cells in the stack as well as ease of monitoring of the individual cell voltage. The individual cells in an EFC stack can also be electrically connected in parallel. This may seem counterintuitive in relation to the low operating voltage of a single EFC cell, and to our knowledge no parallel-connected EFC stack has been reported as of this writing", + " Alternatively, metal contactors (stainless steel or aluminum) might provide a thin and lightweight alternative for the contactors. A schematic image of a bipolar electrode is shown in Figure 16.4. The cathodic side has an integrated serpentine channel for feeding and distributing oxygen (oxidizer), and the anodic side has a rectangular compartment for anode integration. Two independent feed channels are also visible through the bipolar plate that allows for fuel filling and oxidized feeding in series throughout the entire stack. A schematic image and photograph of a five-EFC stack with bipolar electrodes is shown in Figure 16.3 (bottom). The stack has two integrated lines for fuel and oxidizer embedded in the bipolar plates and two pairs of ports for the lines are visible at the end plates. Design optimization and application of bipolar electrodes provides significant volumetric reduction and increased electrode size in comparison with an EFC stack composed of five researchmodular single-cell EFCs (refer to the top image in Figure 16.3). The EFC stack with bipolar plates, however, does not allow for cells with parallel connections and does not provide ports for reference electrode integration. As described previously, despite the maximum power performance for a cathode fed directly with oxygen, the design is not practical because of its technical requirements, safety, cost, and its need for the user to service the feed system. It may be more feasible, however, to supply a low-pressure, low flow rate compressed air feed leveraged from other parts of the system, in a passive form" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0001846_eit48999.2020.9208283-Figure2-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0001846_eit48999.2020.9208283-Figure2-1.png", + "caption": "Fig. 2 Rolling chassis of the BUV", + "texts": [ + " The frame was a simple ladder shape with multiple rungs and out-runners in order to be able to affix the axles, suspension, floor, etc. The footprint was to be 4\u2019 x 8\u2019, standard golf cart dimensions. Steel box tubing was used. The appropriate members were cut and welded and then ground and spray-coated to prevent rust. Leaf springs design for trailers, rated for 1,200 lb each, were affixed to the frame. Various holes were drilled to mount cables and the body. The wheels were added at the end. The rolling chassis is depicted in Fig. 2. A Yamaha G23 rear axle with a built-in differential was used as the rear axle. Due to a mismatch in the sizes (metric to US), wheel adapters were purchased and mounted to the hubs as shown in Fig. 3. The off-road wheels (Honda TRX500, radius 12.5\u201d or 0.32 m) were mounted on top of these. The brake cable was connected to the drum brakes in the wheel hubs via a standard brake cable, which can be seen in Fig. 2 (cable) and Fig. 3 (brake pad). The front axle that can be seen in Fig. 2 and Fig 4. It was specially designed and fabricated from the box tubing to accommodate the wheels and the steering mechanism. The steering mechanism was mounted using the clamps as shown. The steering assembly was similarly fabricated using stock metal and universal couplings that mated with the rack-andpinion steering mechanism. It was mounted using Unistrut C-channel as shown in Fig. 5. Due to space restrictions and a mismatch between the differential and motor splines, they were connected using a chain drive rather than directly, as shown in Fig" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0002441_ismr48331.2020.9312940-Figure6-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0002441_ismr48331.2020.9312940-Figure6-1.png", + "caption": "Figure 6: Guide template\u2019s points projection on the ultrasound imaging plane.", + "texts": [ + " where FP , Fd , and Fe are applied force by the robot to the tissue that is measured by a force sensor, the desired contact force and force error, respectively. Va is the velocity control input along ZPd . III. GRID POINTS PROJECTION The goal is to project the guide template\u2019s grid points coordinates that are given in the base frame on the US imaging plane by considering the US probe-affixed frame. First, the intersections between the axes of holes in the guide template and the imaging plane are obtained as it is shown in Fig. 6. The intersection points are represented in the base frame {B}. Therefore, the transformation matrix from the metric base frame {B} to the US image\u2019s pixel domain is required. In this paper, we are using a flat rectangular ultrasound probe that generates a rectangular image. Therefore, an affine transformation matrix, P BT , is enough to do the point registration. The general form of P BT is P BT =P B A \u00b7PB B \u00b7PB C (7) where A is a translation matrix from the origin of one frame to another, B is a rotation matrix that corresponds to the angles between the frames, and C is a scaling matrix that converts the units of one frame to another" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0001402_s12239-020-0083-y-Figure7-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0001402_s12239-020-0083-y-Figure7-1.png", + "caption": "Figure 7. Multibody dynamic simulation condition of CV joint.", + "texts": [ + " As can be seen in Figure 5, the maximum S/N ratios of the factors are level 3 for R1, level 4 for R2, and level 1 for R3. The values of R1, R2, and R3 optimized through ANOM and the values of existing variable are presented in Table 5. The values of R1, R2, and R3 were modified through HyperMorph, and a contact analysis was performed. The contact analysis of the rollers and housings is shown in Figure 6, and the optimization results showed a stress reduction of 7.41 %. 3.1. Simulation Conditions for GAF Evaluation As seen in Figure 7, the GAF of the CV joint is observed according to the joint angle. To observe the shudder phenomenon as shown in Figure 7, 150 rpm is applied to the TJ and 600 Nm torque is applied to the BJ. The spider of the TJ rotates with the joint angle \u03b8. In this case, the roller rotates through the groove track of the housing because of the rotation of the needle roller. However, as the direction of the roller is tilted when rotating with the joint angle \u03b8, it is periodically moved in the axial direction of the drive shaft, resulting in great sliding resistance. When the sliding resistance increases, three axial forces are generated per revolution" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0000291_icems.2019.8922118-Figure16-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0000291_icems.2019.8922118-Figure16-1.png", + "caption": "FIG. 16. FLUX DENSITY DISTRIBUTIONS.", + "texts": [ + " In addition, the 5th and 7th component of the magnetic flux density of the CW-IM are greater than those of the DW-IM. This reason is that those component is the stator slot harmonics on the CW-IM with 6 stator slots per pairs of poles [5]. That is, the abnormal starting phenomena of the CW-IM at a vicinity of s = 0.8 is caused by the asynchronous torque of the 7th harmonic. Therefore, the starting characteristics appear to have been improved the rotor has by a skew slot which can reduce the 7th component. Figure 16 shows the magnetic flux density distributions of the DW-IM and CW-IM at any time at a s = 0.1. As shown this figure, the magnetic flux density at the stator back-yoke of the CW-IM is approximately 1.5 T while that of the DWIM is approximately 1.0 T. In addition, the magnetic flux density at top of the stator teeth of the CW-IM is approximately 1.8 T. Thus, the magnetic flux density of the CW-IM is high overall, and it is seemed that the magnetic saturation occurs at the top of the stator teeth. Consequently, it was seemed that the efficiency of the CW-IM was reduced" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0003300_iicpe.2014.7115859-Figure2-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0003300_iicpe.2014.7115859-Figure2-1.png", + "caption": "Figure 2 Properly oriented qd synchronously rotating frame", + "texts": [], + "surrounding_texts": [ + "Keywords: Induction Motor, closed loop control, motor drive, PI controller\nI. INTRODUCTION\nWith the availability of faster and less expensive processors and solid state switches, alternating current induction motor drives can be compared favorably to dc motor drives on considerations such as power to weight ratio, acceleration performance, maintenance, operating environment, and higher operating speed without the mechanical commutator. Costs and robustness of the machine, and perhaps control flexibility, are often reasons for choosing induction machine drive in small to medium power range applications. Even without simulating the switching waveforms in detail that is with only the fundamental component of inverter outputs simulated, we can still gain useful insights into the dynamic performance attainable by common control strategies used in today\u2019s adjustable speed induction motor drives.\nThe main aim of this paper is to study the field oriented control strategy as the torque control can be obtained in it and thus allowing the torque to act as a torque transducer.[1] Irrespective of the robust direct filed oriented control, the problem regarding the sensing of the air gap flux linkage and expensive approach behind it motivated to develop a study of indirect filed oriented control technique which are more sensitive to knowledge of the machine parameters but do not require direct sensing of the rotor flux linkages.\nFor study purpose, 20HP induction motor data was used, the details of which are given in the TABLE I.\nTABLE1 DATA OF INDUCTION MOTOR\nSymbol Quantity Value\nP Poles,P 4 HP Capacity 20 V Voltage rating 220 V f rated Rated frequency (Hz) 50 Hz P.F Rated power factor 0.853 s rated Rated slip 0.0287 Nrated Rated speed(Rev/Min) 1748.3 iasb Rated armature current 49.68 Xls stator leakage reactance (ohms) 0.2145 ohms XM Magnetizing Reactance 5.8339 ohms Xplr Rotor leakage reactance 0.2145 ohms J Rotor Inertia 2.8 kg/m2\nBase electrical frequency can be known by the formula wb = 2*pi*frated (1) where frated is the rated frequency in Hertz(Hz)[2].\nII. FIELD OREINTATION METHOD:\nField orientation is a technique for controlling the flux and electromagnetic torque of an induction motor by using electrical parameters estimated in closed-loop; the estimated parameters are used to tune the flux and torque controllers.\nThis control technique is divided into two subcategories i) Open Loop Indirect vector control ii) Closed Loop Indirect vector control\n978-1-4799-6046-0/14/$31.00 \u00a92014 IEEE", + "The field-oriented control has allowed extending the use of induction motors in high performance applications. Direct field-oriented control (DFOC) includes a closed loop rotor flux controller and requires the calculation of the amplitude and position of the rotor flux [8]. This is standard solution for high performance motor drives but requires relatively complex algorithms. Indirect-field oriented control (IFOC) does not need a closed loop rotor flux controller and only requires the angular position of the rotor flux vector which is calculated by integrating the vector angular speed; this can be computed using the rotor speed and the stator current measurement.\nIII. INDIRECT FIELD OREINTATION CONTROL\nIndirect Field Oriented Control is the standard regulation method for high performance induction motor drives that computes the rotor flux angular frequency by integration of the sum of the rotor speed and the slip angular frequency, respectively provided by a suitable transducer, and a simple expression obtained from the mathematical model of the induction motor. [3] The IFOC technique is essentially a predictive approach in that it estimates the angular position of the rotor flux vector by exploiting the model of the machine.\nA commonly used IFOC technique uses the following equations to satisfy the condition for proper orientation:\nThe developed torques Tem reduces to\n= \u2137\u2019edr ieqs (2)\nRelation between slip speed and the ratio of the stator qd current components for the d axis of the synchronously rotating frame to be aligned with the rotor field :\n2 = e - r =(r\u2019r/L\u2019r)(ie qs/ie ds)elect.rad/s (3)\nGiven some desired level of rotor flux, the desired value of ie*\nds dmay be obtained from \u2137\u2019*dr = ((r\u2019rLm)/(r\u2019r+L\u2019rp))ie*ds Wb.turn (4) If the above conditions is satisfied, ensure the decoupling of the rotor voltage equations. To what extent this decoupling is actually achieved will depend on the accuracy of motors parameters used. Indirect field oriented control scheme for a current controlled PWM induction motor drive can be shown as", + "In this part, we will implement the SIMULINK simulation of a current regulated PWM induction motor drive with indirect field oriented control. The objectives are to become familiar with the implementation of a type of field oriented control and to examine how well such a control keeps the rotor flux constant during changes in load torque. For simulation purposes, we will use the same 20-hp, 220-V , four pole induction machine whose parameters are already defined in Table 1.\nThe overall SIMULINK simulation diagram for a current regulated PWM induction motor drive with IFOC is given as :\nOn the left side of the model, a proportional-integral (PI) torque controller converts the speed error to a reference torque T*em. Going into the field orientation block are the reference torque, T*em , the d axis rotor flux \u2137\u2019*dr and the rotor angle, \u03b8r..\nThe inside of the field oriented block is shown as\nEquations (2), (3) and (4) are used inside the field oriented block to compute the values of , \u2137\u2019*dr and *2..Simulation for the generation of gate pulses and PWM inverter is given in figure 7 and figure 8 respectively.\nVoltages\nThe generated gate pulses are fed accordingly to the inverter for switching and thus these generated voltages after being converted to abc reference frame are fed to the induction motor." + ] + }, + { + "image_filename": "designv11_71_0002478_cac51589.2020.9327315-Figure1-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0002478_cac51589.2020.9327315-Figure1-1.png", + "caption": "Fig. 1. Two-wheel differential drive WMR model.", + "texts": [ + " In section IV, it mainly introduces the entire formation algorithm of the WMR-SBOS in detail. Several simulation results are shown in section V. Finally, concluding remarks are given in section VI. 20 20 C hi ne se A ut om at io n C on gr es s ( C A C ) | 9 78 -1 -7 28 1- 76 87 -1 /2 0/ $3 1. 00 \u00a9 20 20 IE EE | D O I: 10 .1 10 9/ C A C 51 58 9. Authorized licensed use limited to: San Francisco State Univ. Downloaded on June 25,2021 at 16:57:47 UTC from IEEE Xplore. Restrictions apply. 3793 The kinematics model of two-wheel differential drive WMR is shown in Fig. 1. This kind of WMR can change the speed and the direction of movement by adjusting the rotation speed difference between the left and right drive wheels. The whole driftless nonlinear system has two degrees of freedom, the kinematics of this system is as follows: \u03b7\u0307 = \u23a1\u23a3 x\u0307 y\u0307 \u03b8\u0307 \u23a4\u23a6 = \u23a1\u23a3 cos \u03b8 0 sin \u03b8 0 0 0 \u23a4\u23a6[ v \u03c9 ] (1) where v = vR + vL 2 , (2) \u03c9 = vR \u2212 vL L . (3) In the above fomula, \u03b7 = [x, y, \u03b8] T represents the pose of the WMR, and u = [v, \u03c9]T represents the speed at which the WMR moves forward and turns" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0000254_ecce.2019.8911883-Figure13-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0000254_ecce.2019.8911883-Figure13-1.png", + "caption": "Fig. 13. Comparison of on-load flux contour plots: (a) Conventional FSPM machine. (b) Proposed FSPM machine.", + "texts": [ + " 0 60 120 180 240 300 360 -150 -100 -50 0 50 100 150 200 P ha se b ac k -E M F Rotor position (deg) Conventional FSPM machine Proposed FSPM machine (a) 0 2 4 6 8 10 0 25 50 75 100 125 150 P h as e ba ck -E M F ( V ) Order Conventional FSPM machine Proposed FSPM machine (b) Fig.10. Comparison of phase back-EMFs: (a) Waveforms. (b) FFT spectrums. 0 60 120 180 240 300 360 -200 0 200 C og gi n g to rq u e (m N m ) Rotor position (deg) Conventional FSPM machine Proposed FSPM machine Fig.11. Comparison of cogging torques The on-load flux lines and flux contour plots of the compared two FSPM machines are presented in Fig. 12 and Fig. 13. It can be found that the flux density in the stator yoke of the proposed FSPM machine is larger than that of the conventional one, because the flux barrier effect of the proposed topology becomes lower. Moreover, the rated torque waveforms of two FSPM machines are compared in Fig.14 (a). It reveals that the proposed FSPM machine has 17.6% higher average torque than the conventional FSPM machine. Moreover, the PM usage of the proposed FSPM is also 35.5% smaller. Therefore, the torque per PM usage of the proposed FSPM machine is about 1" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0001921_s10010-020-00418-x-Figure1-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0001921_s10010-020-00418-x-Figure1-1.png", + "caption": "Fig. 1 Comparison between the two conjugate gear pairs", + "texts": [ + " Yan and Lai [18] applied the theory of envelope to derive the surface equation of a ring gear of an elementary planetary gear train, the planet wheel of which having cylindrical teeth. Chen et al. [20] proposed a new double-enveloping cycloid drive. The main characteristic of this drive is that in some meshing area there are double contact lines simultaneously between one meshing tooth pair. Liu et al. [21, 22] developed a method to calculate the torsional stiffness of double-enveloping cycloid drive, and they also proposed a novel cycloid drive with a cycloid-arc gear based on the double-enveloping method. As shown in Fig. 1a, the conventional cycloid drive (dashed line) consists of cycloid planet gears that have an equidistant curve to the epitrochoid and multiple ring gear rollers placed at equal intervals on the circumference. While the novel one proposed in this paper is a type of non-pin wheel cycloid drive (solid line) with epitrochoid tooth profile and its conjugate envelope, as shown in Fig. 1a. The conventional cycloid drive has been adopted that not only requires less space but also provides a large ratio with only one stage, whose typical structure is shown in Fig. 1b. However, from Fig. 1b, due to the sliding contact between the ring gear roller and ring gear pins as well as the sliding contacts between the ring gear rollers and cycloid planet gears with large sliding coefficients, power loss often occurs unavoidably in the meshing gear pair, which leads to a low transmission efficiency. Furthermore, even though all the teeth of the cycloid planet gear are theoretically in contact with the ring gear rollers and half of them transfer loads, in K practice, such a meshing does not exist due to the modification of the tooth profiles and assembly and manufacturing errors" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0003935_wcica.2014.7053392-Figure1-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0003935_wcica.2014.7053392-Figure1-1.png", + "caption": "Fig. 1 Planar engagement geometry of missile-target", + "texts": [ + " For the convenience of study, there are some assumptions: for missile and target, they\u2019re both regarded as point mass objects; there\u2019s no acceleration in longitudinal axis and the acceleration is perpendicular to line-of-sight (LOS) just to change the velocity direction rather than its magnitude. Of course the assumptions are also suitable for the following threedimensional guidance problem. Based on above assumptions the relative motion equations of planar guidance are written as cos costz t mz mr V q V q (1) sin sintz t mz mrq V q V q (2) Y mO tO Z X mV tV SY SZ SX LOS m m s sq s t Missile CG Target CG Fig.2 Three-dimensional engagement geometry As it\u2019s shown in Fig.1 and Fig. 2, the three-dimensional engagement problem can be decoupled into longitudinal and lateral planar engagements. By reducing the system order three-dimensional guidance becomes a superposed planar guidance which is easier to handle. By reference to the planar method three-dimensional relative motion equations are written as below. cos cos cos sin sin cos cos cos sin sin t t t s s t s m m m s s m s R V q q V q q (3) cos cos sin sin cos cos cos sin sin cos s m m m s s m s t t t s s t s Rq V q q V q q (4) cos cos sin cos sin s s t t t s m m m s R q V V (5) where R is the relative distance from missile CG to target CG; tV , t , t are velocity, trajectory angle and trajectory deflection angle of the target respectively, while mV , m , m are used to depict the missile; sq and s are LOS angles in the longitudinal and lateral planes, respectively" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0000958_ei247390.2019.9062204-Figure3-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0000958_ei247390.2019.9062204-Figure3-1.png", + "caption": "Fig. 3 Circle diagram of voltage and current limit for surface mounted PMSM", + "texts": [ + " Formula (2) can be written as follows 2 2 2( ) ( )s r f r d d r q qu L i L i= + + (3) Assuming that maxu is the maximum value su can be reached, the expression of voltage limit is as follows: 2 2 2 max( ) ( )r f r d d r q qL i L i u+ + \u2264 (4) According to the above analysis, the voltage limit equation of permanent magnet synchronous motor can be written as follows: 2 2 2max( ) ( ) ( )f d d q q r u L i L i+ + \u2264 (5) Similarly, the winding current of the motor is limited by the maximum current of the motor design and the capacity of the inverter. If the maximum value of the stator current is expressed as maxi , the stator current limit equation is as follows: maxsi i\u2264 (6) When the stator current is decoupled into excitation current and torque current, the expression (6) can be expressed as 2 2 2 maxd qi i i+ \u2264 (7) The stator voltage limit ellipse and stator current limit circle are composed of formulas (5) and (7), as shown in Figure 3 below [6]. In the figure, the solid line circle is the motor current limit circle, and the dotted line ellipse is the motor voltage limit ellipse. In particular, for surface mounted PMSM, the ellipse's two axes are equal in length, so the voltage limit is circular. The stator current vector control trajectory of surface mounted PMSM is introduced as an example. Formula (5) shows that the radius of the voltage limit circle decreases with the increase of speed. The stator current vector must satisfy both the voltage limit condition and the current limit condition, so when the motor runs at a certain speed, the trajectory of the stator current vector must fall in 1216 Authorized licensed use limited to: University of Exeter" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0002436_icamechs49982.2020.9310088-Figure1-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0002436_icamechs49982.2020.9310088-Figure1-1.png", + "caption": "Fig. 1. Coordinates system of lower body", + "texts": [], + "surrounding_texts": [ + "This phase occurs when only one foot is in contact with the ground as shown in Fig. 2, so the whole lower body can be considered as a 12 DOFs robot manipulator. By applying Euler-Lagrange method, the dynamic model for a manipulator is given by [17]: (16) where is the vector of joints. is the inertia matrix. Authorized licensed use limited to: UNIVERSITY OF CONNECTICUT. Downloaded on May 16,2021 at 11:40:21 UTC from IEEE Xplore. Restrictions apply. is the Coriolis and centrifugal force matrix. ) is the vector of gravitational forces. is the vector of joint torques. 2m 3m 4m 5m fm hm O x z Swing leg Stance leg The integration of and over are zero, we have: (19) Since the contact duration is infinitesimal, the joint positions remain unchanged, therefore: Then the final result when integrating Eq. (17) is: (20) The differential of the joint angle can be calculated using the following equation [12]: (21) where is the velocity of the foot just before the contact with the ground. is Jacobian matrix. III. SLIDING MODE CONTROLLER DESIGN To realize the biped walking process, pre-computed walking trajectories with constraints for the robot should be given first. Then by using inverse kinematic, a set of joint trajectories is deduced. Finally, a suitable controller is used to make the joints follow the prescribed trajectories. The basic work flow to realize biped walking is shown in Fig. 4. Inverse Kinematic Constraints SMC System Forward Kinematic ref error u + - realCOM refCOM Fig. 4. Simulation workflow To implement the sliding mode controller, the dynamic model (17) must be in the form: (22) where is the system state. are known non-linear functions. is control inputs vector. is external disturbances vector (24) By defining , , , Eq. (24) can be summarized as follows: (25) Defines a joint trajectories vector as , the difference between the current joint angles and the desired angle is defined as the error: (26) Defines a vector as the sliding surfaces which is given in the following form: (27) where is a constant symmetric positive-definite matrix. Taking the first derivative of Eq. (27), it yields : (28) The control input designed for the system of Eq. (25) is as followed: (29) where is a constant diagonal matrix. Each element in the main diagonal of should satisfy the following: (30) where The candidate Lyapunov function (CLF) is given by: (31) By taking first time derivative of CLF, it yields: (32) Let us define : (33) Taking the infinity norm of , it yields : (34) Recall Eq. (23) and Eq. (30) we obtain: (35) IV. SIMULATION RESULTS To evaluate the effectiveness of the proposed controller, a simulation was done by MATLAB. In the simulation, 5 steps walking trajectories were created, the sampling time is 0.01 seconds, the pre-computed trajectories take 10 seconds to complete. The physical parameters are given by Table 2. First, center of mass trajectories was calculated beforehand with the constraint: the motion of the center of mass does not occur much. Then, the joint trajectories were calculated using inverse kinematic. a. trajectory and tracking error b. trajectory and tracking error c. trajectory and tracking error d. trajectory and tracking error e. trajectory and tracking error Authorized licensed use limited to: UNIVERSITY OF CONNECTICUT. Downloaded on May 16,2021 at 11:40:21 UTC from IEEE Xplore. Restrictions apply. The simulation of joint trajectories as shown in Fig. 5 showed that the SMC tracked the joint trajectories. At the time: 2 seconds, 4 seconds, 6 seconds and 8 seconds of the pre-computed trajectories in Fig. 5 b, c, d, h, j, the joint angles had unusual peaks, leading to large errors (up to 6 degrees). This is due to one main reason, the changes in the calculated angle values happened in a 0.01s interval, which is too quick for the controller to respond, this can be fixed by adjusting the pre-computed trajectories. The COM trajectories was calculated as shown in Fig. 6, 7, 8. The COM trajectories showed that the real COM has successfully tracked the pre-computed trajectories. During the initial time, the trajectory dropped to due to the initial errors in the joints. However, the z motion maintained at during the steady Authorized licensed use limited to: UNIVERSITY OF CONNECTICUT. Downloaded on May 16,2021 at 11:40:21 UTC from IEEE Xplore. Restrictions apply. state. The assumed distance of the center of mass of UXA-90 to the ground is approximately . So this satisfied the constraint proposed before: the movement does not occur much. V. CONCLUSION This paper proposed a more general dynamic model for UXA-90 than previous studies [15, 16] and use sliding mode controller to control it. The simulation results show that the joint trajectories tracking is quite good, the errors can be driven to zero. The COM successfully tracked the desired trajectory, satisfying the constraint: the movement does not occur much. This validates that the SMC can be used to control the 12 DOFs dynamic system of humanoid robot. ACKNOWLEDGMENT This research is supported by DCSELAB and funded by Vietnam National University HoChiMinh City (VNU-HCM) under grant number TX2020-20B-01. REFERENCES [1] Yoshiaki Sakagami et al (2002), The intelligent ASIMO: System overview and integration, in Intl. Conference on Intelligent Robots and System, pp. 2478-2483. [2] T. Jung, J. Lim, H. Bae, K. Kyu Lee, H.-M. Joe and J.-H. Oh, \"Development of the Humanoid Disaster Response Platform DRCHUBO+,\" IEEE Transactions on Robotics, vol. 34, no. 1, pp. 1-17, Feb. 2018. [3] K. Kaneko, M. Morisawa, S. Kajita, S. Nakaoka, T. Sakaguchi, R. Cisneros and F. Kanehiro, \"Humanoid Robot HRP-2 Kai - Improvement of HRP-2 towards Disaster Response Tasks,\" in 2015 IEEE-RAS 15th International Conference on Humanoid Robots (Humanoids), Seoul, 2015. [4] Y. Hwang, E. Inohara, A. Konno and M. Uchiyama, \"An Order n Dynamic Simulator for a Humanoid Robot with a Virtual SpringDamper Contact Model,\" in International Conference on Robotics & Automation, Taipei, 2003. [5] K. Atsushi, K. Noriyoshi, S. Satoshi, F. Tomoyuki and U. Masaru, \"Development of a Light-Weight Biped Humanoid Robot,\" in International Conference on Intelligent Robots and Systems, 2000. [6] J. H.Park and K. D.Kim, \"Biped Robot Walking Using GravityCompensated Inverted Pendulum Mode and Computed Torque Control,\" in International Conference on Robotics & Automation, Leuven, 1998. [7] J. Han, Bipedal Walking for a Full-sized Humanoid Robot Utilizing Sinusoidal Feet Trajectories and Its Energy Consumption, 2012. [8] M. Raibert, S. Tzafestas and C. Tzafestas, \"Comparative Simulation Study of Three Control Techniques Applied to a Biped Robot,\" in Systems Man And Cybernetic - SMC, Le Touquet, 1993. [9] F. Plestan, J. W.Grizzle, E. R.Westervelt and G. Abba, \"Stable Walking of a 7-DOF Biped Robot,\" in Transactions on Robotics and Automation, 2003. [10] E. R.Westervelt, J. W.Grizzle, C. Chevallereau, J. H. Choi and B. Morris, Feedback Control of Dynamic Bipedal Robot Locomotion, CRC Press, 2007. [11] I.-W. Park, J.-Y. Kim and J.-H. Oh, \"Online Walking Pattern Generation and Its Application to a Biped Humanoid Robot - KHR-3 (HUBO),\" Advanced Robotics, 2008. [12] T. Spyros, R. Mark and T. Costas, \"Robust Sliding-mode Control Applied to a 5-Link Biped Robot,\" Journal of Intelligent and Robotic Systems, pp. 67-133, 1996. [13] S. A.Moosavian, M. Alghooneh and A. Takhmar, \"Stable trajectory planning, dynamics modeling and fuzzy regulated sliding mode control of a biped robot,\" in 7th IEEE-RAS International Conference on Humanoid Robots, Pittsburgh, 2007. [14] I.-G. Park and J.-G. Kim, \"Robust control for dynamic walking of a biped robot with ground contacting condition,\" in 2001 IEEE International Symposium on Industrial Electronics Proceedings, Pusan, 2001. [15] A. Nguyen Van Tien, H. Quoc Le, T. P. Tran and T. T. Nguyen, \"Design of Biped Walking Gait on Biped Robot,\" in 14th International Conference on Ubiquitous Robots and Ambient Intelligence (URAI), 2017. [16] N. D. K. Nguyen, B. L. Chu, V. T. A. Nguyen, V. H. Nguyen and T. T. Nguyen, \"Optimal Control for Stable Walking Gait of a Biped Robot,\" in 14th International Conference on Ubiquitous Robots and Ambient Intelligence (URAI), 2017. [17] M. W.Spong, S. Hutchinson and M.Vidyasagar, Robot Modeling and Control, 1989. Authorized licensed use limited to: UNIVERSITY OF CONNECTICUT. Downloaded on May 16,2021 at 11:40:21 UTC from IEEE Xplore. Restrictions apply." + ] + }, + { + "image_filename": "designv11_71_0002683_012034-Figure7-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0002683_012034-Figure7-1.png", + "caption": "Figure 7. The statement of a torsion problem.", + "texts": [ + " The value of shear modulus is calculated just as a ratio of shear stress to shear strain in a linear part of stress-strain curve. But if we take into account the assumption about orthotropic behavior the stress-strain state will be noticeably different. That\u2019s why in case of torsion problem it is necessary to perform an additional research. Calculation submitted below are based on the anisotropic models of continuum mechanics [10]. First of all, we consider that we get the amount of values of torque and twist angle as a result of torsion tests. We assume the uniaxial torsion is performed about Z axis (Fig. 7). IOP Conf. Series: Materials Science and Engineering 986 (2020) 012034 IOP Publishing doi:10.1088/1757-899X/986/1/012034 Let\u2019s assume the following relation between torque and twist angle: \ud835\udf11 = \ud835\udc40\ud835\udc59 \ud835\udc36 (5) Second thing to mention is that there are two independent shear moduli which affect the stress-strain state during the uniaxial torsion along proposed axis. In case of torsion about Z axis these two moduli are G_xz and G_yz. This means that the relation between torque and twist angle includes both these moduli" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0000471_physreve.101.013002-Figure1-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0000471_physreve.101.013002-Figure1-1.png", + "caption": "FIG. 1. (a) Edge-on view of a buckled thin viscoelastic sheet when it touches ground. V0 and H represent the feeding velocity and feeding height, respectively. (b) Schematic of the particle system:ka represents the axial stiffness; kb represents the stiffness of the rotational spring; ca represents the axial damping coefficient; cb represents the rotational damping coefficient. (c) Schematic of elastic bending forces in the local bending system formed by particles (i \u2212 1) \u2212 (i) \u2212 (i + 1). (d) Schematic of viscous bending forces: f v1,i is due to the angular velocity \u03c91,i, f v2,i is due to the angular velocity \u03c92,i, and f v,i is the negative sum of f v1,i and f v2,i.", + "texts": [ + " The sheet is treated as a series of discrete masses connected by viscoelastic (Kelvin-Voigt) elements which can capture the large deformation and the viscoelasticity. In this paper we use this particle approach to investigate the periodic folding of a viscoelastic sheet. The remainder of this paper is organized as follows. In Sec. II the discrete particle model is presented. In Sec. III the model is applied to simulate the folding process of a viscoelastic sheet. First, the quasistatic results are reproduced. Then the effects of inertia and viscosity are discussed in the dynamic case. Section IV summarizes the main conclusions drawn from this study. Figure 1(a) shows a falling viscoelastic solid sheet as it starts to buckle following contact with a steady horizontal plane due to the compression. Note the choice of the KelvinVoigt model given our interest in textile and paper-folding processes. V0 and H represent the vertical feeding velocity and feeding height. To simulate the folding process, the model needs to capture the large buckling deformation and the contact between the sheet and the ground. A. Particle approach We assume that deformation along the width direction is uniform during folding of the sheet, thus twisting can be neglected. The sheet can be modeled as a series of discrete particles connected by axial and rotational viscoelastic elements, shown as Fig. 1(b). The axial spring-damper element known as the Kelvin-Voigt model has been shown to be suitable to model the viscoelasticity of paper products [33]. a0 = L0/(N \u2212 1) and m = \u03c1ha0w are the spacing and mass of the particles, respectively, where L0, w, \u03c1, and h represent the initial length, width, density, and thickness of the web, respectively, and N denotes the total number of particles. The axial stiffness is given by ka = Ewh/a0 where E is Young\u2019s modulus. After deformation the axial elastic force acting on particle i in vector form is f a,i = ka(|ri\u22121 \u2212 ri| \u2212 a0)ei,i\u22121 + ka(|ri+1 \u2212 ri| \u2212 a0)ei,i+1, (1) where ei,i+1 denotes a unit vector pointing from particle i to particle i + 1 and ri is the position vector of particle i", + " The in-plane viscous force is related to the axial velocity difference between two adjacent particles: f d,i = \u2212ca[(r\u0307i \u2212 r\u0307i\u22121) \u00b7 ei,i\u22121]ei,i\u22121\u2212ca[(r\u0307i\u2212r\u0307i+1) \u00b7 ei,i+1]ei,i+1, (2) where r\u0307i represents the velocity of particle i and ca denotes the phenomenological viscous damping coefficient with ca = \u03bcwh/a0. Here \u03bc is the dynamic viscosity of the sheet. Since the particles are not endowed with a rotational degree of freedom, we cannot directly apply the bending moment. Instead, we introduce bending forces to the particles to achieve the same bending effect. We consider a local bending system formed by three consecutive particles (i \u2212 1), (i), and (i + 1), as shown in Fig. 1(c). Due to the bending effects, the system tends to recover to the initial flat state. Thus, we can imagine there are two restoring forces f b1,i and f b2,i acting on the left particle (i \u2212 1) and the right particle (i + 1) to achieve that effect. The restoring forces are perpendicular to the axial direction and related to the local bending angle \u03b3i, given by [30] | f b1,i| = | f b2,i| = kb(\u03b30 \u2212 \u03b3i ) a2 0(1 + \u03b5i )2 , (3) where kb = Ewh3/12 is the bending stiffness, \u03b30 = \u03c0 is the initial bending angle, and \u03b5i = (|ri+1 \u2212 ri| \u2212 a0)/a0 is the axial strain", + " Thus, the total elastic bending force acting on particle i is f b,i = f b0,i + f b2,i\u22121 + f b1,i+1. (4) During buckling, as the local bending angle \u03b3i changes, the sheet is also subject to viscous dissipation that slows down this process. The viscous bending moment is proportional to \u03b3\u0307i and can be derived as \u03bc\u03b3\u0307iwh3/12a0. The detailed derivation can be found in Appendix A. We implement the viscous bending moment in a similar fashion as the elastic bending moment, by introducing the corresponding viscous forces as shown in Fig. 1(d). The magnitude of the viscous bending force can be written as | f v1,i| = | f v2,i| = ca\u03b3\u0307ih2 12a0 , (5) where ca = \u03bcwh/a0 is the phenomenological damping coefficient. f v1,i and f v2,i are perpendicular to the segments (i \u2212 1) \u2212 (i) and (i) \u2212 (i + 1) respectively, and the exact directions depend on whether the local bending angle \u03b3i is increasing or decreasing. For example, in Fig. 1(d) \u03b3i is increasing, so f v1,i and f v2,i both point \u201cinward\u201d and slow down the increase of \u03b3i. In a discrete particle system, \u03b3\u0307 i is related to the angular velocities \u03c91,i and \u03c92,i as \u03b3\u0307 i = \u03c91,i \u2212 \u03c92,i. Here \u03c91,i can be calculated based on the velocity difference between the particle i \u2212 1 and the particle i and on the length of the segment (i \u2212 1) \u2212 (i). \u03c92,i can be derived following the same rule. When \u03c91,i is equal to \u03c92,i, \u03b3\u0307 i becomes zero and the sheet undergoes rigid rotation. As before, to maintain the overall force balance, we apply a viscous bending force applied on particle i, given by f v,i = \u2212 f v1,i \u2212 f v2,i", + " The elastic bending model and the viscous damping model are both validated in Appendix B and Appendix C, respectively. III. RESULTS AND DISCUSSION A. Energy analysis Considering a fold with an arc length l f formed as the falling viscoelastic sheet hits the ground, before folding the initial kinetic energy scales as \u03c1hwl f V 2 0 and the gravitational potential energy scales as \u03c1ghwl2 f . Due to the viscous effect and extensibility of the sheet, the folding process is associated with both bending and in-plane stretching dissipations, represented by the axial and rotational dampers in Fig. 1(b). The axial stress caused by the bending dissipation is proportional to the strain rate, given by \u03c3 \u223c \u03bc(z\u03c6s)t \u223c \u03bcz\u03c6st . Here z represents the distance to the midplane in the cross-direction and \u03c6s represents the curvature. Integrating \u03c3 zdA over the cross section and using the approximation \u03c6st \u223c V0/l2 f , the viscous bending moment can be derived as M\u03bc \u223c \u222bh/2 \u2212h/2 \u03c3 zw dz \u223c \u222bh/2 \u2212h/2 \u03bcz2wV0/l2 f dz \u223c \u03bcIV0/l2 f . Finally integrating M\u03bc\u03c6s along the length of the sheet, the total energy dissipated by viscous bending scales as \u222bl f 0 M\u03bc\u03c6sds \u223c \u222bl f 0 M\u03bc/l f ds \u223c \u03bcIV0/l2 f ", + " The ratio of the viscous bending energy and the elastic bending energy (\u03bcV0/El f ) can be interpreted as a ratio of the viscoelastic relaxation time (\u03bc/E ) and the time required for a single fold to form (l f /V0). This ratio remains small for the folding of a viscoelastic solid for the range of parameters in Table I (relevant for a viscoelastic solid). Arguably, the same ratio can be interpreted as a Deborah number for a viscoelastic fluid folding problem. However, the choice of Kelvin-Voigt viscoelastic solid (see Fig. 1) does not apply to a Maxwell fluid, because under a constant applied stress (creep test), a Kelvin-Voigt viscoelastic solid does not exhibit an unbounded strain (characteristic of a fluid), whereas a Maxwell (dashpot and spring in series) viscoelastic fluid does exhibit fluid-like response after the initial elastic response. This distinction between a viscoelastic solid and viscoelastic fluid is worth noting. Due to the energy balance, the sum of the initial gravitational energy and the kinetic energy should be equal to the sum of the dissipated energies and the bending energy of the fold: c1\u03c1ghwl2 f + c2\u03c1hwl f V 2 0 = c3 EI l f + c4 \u03bcIV0 l2 f + c5\u03bchw \u221a gl f , (10) where c1 \u2212 c5 are the coefficients for each term" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0000356_iecon.2019.8927121-Figure10-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0000356_iecon.2019.8927121-Figure10-1.png", + "caption": "Fig. 10. The helical (screw) PM RLM [28]", + "texts": [ + " The rotor has 12\u00d716 poles (96 PMs and 96 made of iron core). It was demonstrated, that for the proposed structure, even if the quantity of the used PMs was reduced to the half, the power density was not compromised. The comparison with the PM RLM of similar sizes proposed in [18] revealed that back-emf, torque and axial force are only diminished by round 15\u00f717% [26]. Upon a totally other approach, a helical PM RLM was proposed by Y. Fujimoto et al. in [27]. The so-called screw motor has itself a unique spiral structure, as it can be seen in Fig. 10. Both of its iron cores have a special helical shape with equal lead length. The stator has surface slots in which two distributed three-phase windings are placed. These can generate the axial direction magnetic flux. The mover has two iron core layers having among them helical-shaped alternatively and axially magnetized PMs. Even if the iron cores of this RLM can be manufactured of easy-to-machine free-cutting steel with high permeability or of soft magnetic composites (SMCs) [29], the structure (including the specially shaped PMs) seems complicated and requires special fabrication technology" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0001046_ut.37.003-Figure3-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0001046_ut.37.003-Figure3-1.png", + "caption": "Fig 3: Tracking error in body coordinate frame of underwater vehicle", + "texts": [ + "5 (m/s) Weight 75 kg Depth rating 300 m Power supply 150 VDC, 1.5 Kwh Video cameras 1/3\u201d DSP CCD Thrusters 4 total: 2 horizontals and 2 verticals Lighting LED lights Deng et al. Collision avoidance with control barrier function for target tracking of an unmanned underwater vehicle 6 In the trajectory tracking task, the UUV must follow a desired path in Cartesian workspace with a specified timing law. The desired path satisfies the UUV nonholonomic constraint. Equivalently, it tracks a path generated by a reference UUV (Fig 3). For the horizontal-plane motion control of the UUV, the desired reference state of a vehicle is described as hd = [xd yd Yd]T, qd = [ud rd]T, where qd = [ud rd]T is the desired velocity in the body-fixed frame, hd = [xd yd Yd]T is the desired state of the UUV in the inertial frame, qc = [uc rc]T is the output of velocity controller in the body-fixed frame, and the actual state of the UUV is represented by h = [x y Y]T and q = [u r ]T. The state errors are denoted by e = [ex ey eY]T in the body frame" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0001352_j.mechmat.2020.103523-Figure2-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0001352_j.mechmat.2020.103523-Figure2-1.png", + "caption": "Fig. 2. Material model", + "texts": [ + " The second one describes theoretical formulas for predicting the temperature variation under adiabatic condition. The last one proposes a method to consider the effect of heat transfer on temperature variation. The temperature variation will be predicted from the stress-strain relationship in this study. Since it is inconvenient to directly employ measured stress-strain data to predict temperature variation, a three-elements model with two rubber elastic elements and a linear viscous element as shown in Fig. 2 is employed to reproduce the stress-strain data. Based on the model, the total nominal stress is given as sum of the nominal stresses in Spring 1 ( ) and Spring 2 ( ). In addition to this, the nominal stress of Spring 2 is equal to that of dashpot . Thus, the relations of these stresses are shown as: . (1) Herein, non-linear behavior of both springs is characterized by using three-chain network model [19] as follows: [{ ( ) ( \u2044 )} ] (2) The parameters and are additionally introduced to consider residual strain and stress softening due to mechanical accommodation, respectively" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0000019_chicc.2019.8865756-Figure4-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0000019_chicc.2019.8865756-Figure4-1.png", + "caption": "Fig. 4: The desired formation configuration of the four UAVs", + "texts": [ + " Therefore, from (36) one knows that V\u0307 (t) \u2264 0 if \u03beT (t)\u03a6i\u03be(t) + \u03b32\u03b2i(\u03c3,\u0393) \u2264 0, which implies that \u03bei(t) will converge into the set S1 = {\u03bei : |\u03beTi \u03a6i\u03bei| \u2264 \u03b32\u03b2i(\u03c3,\u0393)}, and this yields that the error {e\u0304i(t) : i = 1, 2, \u00b7 \u00b7 \u00b7 , N} will converge to a small ball around the origin. This completes the proof. To illustrate the effectiveness of the proposed FTC ap- proach, a simulation study on a practical UAVs formation is provided in this section. A swarm consisting of four Kyosho 260 unmanned helicopters (see Fig. 1 and Fig. 2) have been used for simulation testing and analysis. The communica- tion topology graph and desired formation configuration of the four UAVs are given in Fig. 3 and Fig. 4, respectively. trol problem has been investigated for a class of leader- follower MASs subject to actuator faults. A decentralized state observer and fault estimator based formation control protocol is proposed, together with the sufficient conditions under which the desired formation can be reached. It is proved that formation errors of all the following agents can converge to a small set around the origin. The numerical ex- ample has verified the proposed theoretical results. Future research efforts will be devoted to the fault tolerant coopera- tive formation problem of MASs with stochastic dynamics" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0000005_s11771-019-4184-6-Figure4-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0000005_s11771-019-4184-6-Figure4-1.png", + "caption": "Figure 4 Experimental setup (a), fault bearing (b) and defect (c)", + "texts": [ + " Although the amplitudes of demodulated frequencies under different noise levels vary slightly, the demodulation precisions have nearly unchanged. However, though some peaks appear in the spectra via the traditional method, these peaks deviate from the theoretical frequencies and are surrounded or even overwhelmed by the interfering noise. This confirms the noise robustness of the proposed method. In this section, the effectiveness of the proposed method is validated using the experimental signal measured by a vibration test rig. The experimental setup and the tested bearing are shown in Figure 4. The vibration signals are collected by an accelerometer mounted on the top of bearing support casing extremely close to the bearing. The sampling rate is set to 24000 Hz. An encoder is mounted at the drive end to measure the corresponding rotating speed. To mimic the bearing faults, the manually cutting defects are introduced to the inner and outer races of the tested bearing. The rotating frequency increases from 13.1 to 56.2 Hz with an approximate parabola pattern. The data is collected for 5 s" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0001308_1077546320936901-Figure2-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0001308_1077546320936901-Figure2-1.png", + "caption": "Figure 2. Pendulum time-delay system with vibration vertical to the pivot point.", + "texts": [ + " The IHTGA improves the HTGA by using the generation-based crossover rate \u03b1\u00f0~g\u00de \u00bc \u03b1min \u00fe \u00f0\u00f0~gmax ~g\u00de=~gmax\u00de\u00d7 \u00f0\u03b1max \u03b1min\u00de and the generation-based mutation rate \u03b2\u00f0~g\u00de \u00bc \u03b2min \u00fe \u00f0\u00f0~gmax ~g\u00de=~gmax\u00de \u00d7 \u00f0\u03b2max \u03b2min\u00de, where \u03b1max \u00bc 0:9, \u03b1min \u00bc 0:7, \u03b2max \u00bc 0:03, \u03b2min \u00bc 0:01, as well as ~g and ~gmax are the current generation and the maximum of generation, respectively. In this study, the maximum generation number is 250. On the other hand, the Taguchi method is a well-known local optimization approach and has been merged into the GA by Tsai et al. (2004) for obtaining both better and robust results. To reduce the length of the article, for the detailed explanation of the hybrid Taguchi-genetic algorithm (HTGA), the readers may refer to the work by Tsai et al. (2004), where the Taguchi method has been explained in detail therein. Figure 2 shows the illustrative example in which Y is the amplitude, \u03c9 is the oscillation frequency, and y\u00f0t\u00de is the displacement of the pivot point in the vertical direction (Rao, 1995). The pendulum pivot point vibrates as y\u00f0t\u00de \u00bc Y cos\u00f0\u03c9t\u00de (27) The nonlinear equation for the motion of the timevarying pendulum system with time-delay torque applied to the system mass can be derived as ml2\u20ac\u03b8\u00f0t\u00de \u00fe ml g \u20acy\u00f0t\u00de sin\u00f0\u03b8\u00f0t\u00de\u00de \u00bc u\u00f0t \u03c4\u00de (28) where m is the system mass of the pendulum system, l is the length of the pendulum system, g is the acceleration of gravity, \u20acy\u00f0t\u00de is the acceleration of the pivot point in the vertical direction, \u03b8\u00f0t\u00de is the angle of the pendulum system and changes between \u03c0=2 and \u03c0=2, \u20ac\u03b8\u00f0t\u00de is the angular acceleration of the pendulum system, u\u00f0t \u03c4\u00de is the torque with time delay applied to the system mass, and \u03c4 is the known time delay" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0001145_s40430-020-02387-2-Figure4-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0001145_s40430-020-02387-2-Figure4-1.png", + "caption": "Fig. 4 Cross section of twin-rod MR damper", + "texts": [ + " The piston is connected to piston rods on both sides, and the terminals of the coil are extracted from the hole in one of the piston rods. The piston is contained inside the inner cylinder which is filled with an MR fluid. To prevent leakage of the fluid, piston and cylinder seals are accommodated on the cylinder cap. The outer cylinder is simply an attachment used for further characterization of the damper, while it also helps provide free space for one of the piston rods to move. The geometric variables and the cross section of the MR damper are shown in Fig.\u00a04. A damper is fabricated using the parameters listed in Table\u00a02. A total of 500 turns of copper coil of wire gauge 26 were wound on the web of the piston, and the sleeve was press-fit. The outer dimensions of the damper are selected based on the normal human knee size and shape constraints, and the dependent variables such as the number of turns of coil, wire gauge of coil and gap size are determined based on the outer geometry of damper and also using the previously published literature. The fabricated damper is shown in Fig" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0002107_iecon43393.2020.9255202-Figure2-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0002107_iecon43393.2020.9255202-Figure2-1.png", + "caption": "Fig. 2. The BDFRG primary voltage vector oriented phasor diagram.", + "texts": [ + " Introducing the measured secondary current stationary frame components as the reference model output, coupled with the primary real and reactive power based adaptive model design, has allowed a much higher accuracy and fewer parameter dependence than the MRAS observers in [13], [14]. The implications of parameter mismatches are investigated and supported by realistic simulation studies. The rotor angular speed of the BDFRG is related to the supply frequency of the primary (\u03c9p) and secondary (\u03c9s) windings as follows [6], [7], [12]: \u03c9rm = \u03c9p + \u03c9s pr = \u03c9p pr ( 1 + \u03c9s \u03c9p ) = \u03c9syn ( 1 + \u03c9s \u03c9p ) (1) \u03b8rm = \u03b8p + \u03b8s pp + ps = \u03b8p + \u03b8s pr (2) where pp and ps denote the windings pole pair numbers, and pr = pp+ps is the number of the rotor poles. Fig. 2 illustrates the meanings of \u03b8p and \u03b8s reference frames angles in (2). The BDFRG operates in synchronous mode when the secondary is DC (i.e \u03c9s = 0). Note that the corresponding speed (\u03c9syn) is half that of a conventional pr-pole DFIG given (1). Thus, the BDFRG can be classified as a medium-speed machine requiring a two-stage gearbox unlike the vulnerable three-stage counterpart with DFIG wind turbines [16]. The active power balance in steady-state can be expressed using (1) as: Pm = Te \u00b7 \u03c9rm = Te \u00b7 \u03c9p pr \ufe38 \ufe37\ufe37 \ufe38 Pp + Te \u00b7 \u03c9s pr \ufe38 \ufe37\ufe37 \ufe38 Ps = Ps \u00b7 ( 1 + \u03c9p \u03c9s ) (3) where the electro-magnetic torque (Te) and the primary power (Pp) are both negative in sign for the BDFRG with the adopted motoring (BDFRM) convention in (3)", + " The bi-directional flow of the secondary power (Ps) allows the machine operation above and below the synchronous speed i.e. in supersynchronous mode (\u03c9s > 0) when Ps < 0, and at subsynchronous speeds (\u03c9s < 0) for Ps > 0. The variable speed range of a BDFRG wind turbine around \u03c9syn i.e. [\u03c9min, \u03c9max] can be defined as: r = \u03c9max \u03c9min = \u03c9p + \u03c9s \u03c9p \u2212 \u03c9s =\u21d2 \u03c9s \u03c9p = r \u2212 1 r + 1 (4) For a usual r = 2, the secondary frequency is limited to \u03c9s = \u03c9p/3 and Ps \u2248 0.25Pm from (3). This implies that a partiallyrated converter can be used by analogy to DFIG. The dynamic model of the BDFRM in rotating reference frames (Fig. 2) and using standard space vector notation can be represented by the following set of equations [6], [7]: vp = Rpip + d\u03bbp dt + j\u03c9p\u03bbp (5) vs = Rsis + d\u03bbs dt + j\u03c9s\u03bbs (6) \u03bbp = Lpip + Lmi\u2217sm = Lp(ipd + jipq) + Lm(imd \u2212 jimq) (7) \u03bbs = Lsis + Lmi\u2217pm = \u03c3Lsis + Lm Lp \u03bb\u2217 p = \u03c3Lsis + \u03bbm (8) Te = 3 2 pr (\u03bbpdipq\u2013\u03bbpqipd) = 3 2 pr (\u03bbmdisq\u2013\u03bbmqisd) (9) J d\u03c9rm dt = Te \u2212 Tt (10) where J denotes the moment of inertia, Lm is the magnetising inductance, Lp and Ls are the primary and secondary selfinductances, \u03c3 = 1 \u2212 L2 m/(LpLs) is the leakage coefficient, and \u03bbm is the mutual flux linkage magnitude [17]. It is important to emphasise that \u03c9p rotating ism is the frequency (but not amplitude) modulated secondary current vector (is) running at \u03c9s as shown in Fig. 2. Furthermore, the relative angular displacements and magnitudes of ism and is from the respective flux vectors \u03bbp and \u03bbm are the same so the following relationship applies in the corresponding reference frames under FOC conditions (i.e. with dp-axis aligned with \u03bbp, and ds-axis with \u03bbm) [6], [7], [12]: ism = ismej\u03b3 = is = ise j\u03b3 (11) In a MRAS observer, the rotor speed information is retrieved by comparing some actual state variables and their rotor position dependent estimates generated by the reference and adaptive models, respectively", + " For a star-connected winding with a positive phase sequence and an isolated neutral point, the latter can be obtained as: is = ise j\u03b4 = is\u03b1 + jis\u03b2 = isa + j isa + 2isb\u221a 3 (12) The adaptive model acquires the magnetically coupled secondary currents in the \u03c9p frame, i\u0302mq and i\u0302md, using the primary active (Pp) and reactive (Qp) power measurements, respectively. The general expressions, Pp = 1.5 Re{vpi\u2217p} and Qp = 1.5 Im{vpi\u2217p}, can be simplified by aligning the qp-axis with the primary voltage vector (i.e vpq = vp, vpd = 0) as shown in Fig. 2. Consequently, Pp = 1.5vpipq and Qp = 1.5vpipd. The key relations between the primary powers and secondary currents are developed by using (5) and (7) and applying a valid approximation of \u03bbpd >> \u03bbpq \u21d2 \u03bbpq \u2248 0: Pp = 3 2 \u03bbpq + Lmimq Lp vp \u2248 3 2 Lm Lp vpimq (13) Qp = 3 2 \u03bbpd \u2212 Lmimd Lp vp \u2248 3 2 ( v2p \u03c9pLp \u2212 Lm Lp vpimd ) (14) Under the resulting near FOC circumstances, (11) is applicable and the actual secondary current estimates in the \u03c9s frame, i\u0302sq and i\u0302sd, can be obtained as follows: i\u0302mq = i\u0302sq \u2248 L\u0302p vpL\u0302m (vp\u03b1ip\u03b1 + vp\u03b2ip\u03b2 \ufe38 \ufe37\ufe37 \ufe38 ) 2 3 Pp (15) i\u0302md = i\u0302sd \u2248 vp \u03c9pL\u0302m \u2212 L\u0302p vpL\u0302m (vp\u03b2ip\u03b1 \u2212 vp\u03b1ip\u03b2 \ufe38 \ufe37\ufe37 \ufe38 ) 2 3 Qp (16) where L\u0302m and L\u0302p are inaccuracies prone inductance values obtained by off-line testing, vp = (v2p\u03b1 + v2p\u03b2) 1/2 and: vp = vp\u03b1 + jvp\u03b2 = vab + vac 3 + j vbc\u221a 3 (17) A common ds\u2212qs to \u03b1\u2212\u03b2 frame conversion transformation, i\u03b1\u03b2 = idqe j\u03b8s , can now be applied to derive: i\u0302s\u03b1 = i\u0302sd cos(\u03b8\u0302r \u2212 \u03b8p)\u2212 i\u0302sq sin(\u03b8\u0302r \u2212 \u03b8p) (18) i\u0302s\u03b2 = i\u0302sd sin(\u03b8\u0302r \u2212 \u03b8p) + i\u0302sq cos(\u03b8\u0302r \u2212 \u03b8p) (19) where \u03b8\u0302s = \u03b8\u0302r \u2212 \u03b8p (Fig" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0003714_esda2014-20232-Figure5-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0003714_esda2014-20232-Figure5-1.png", + "caption": "Figure 5: Conceptual water brake CAD model", + "texts": [], + "surrounding_texts": [ + "To design this product, it is essential to understand the complete function, acquisition and maintenance process. It is also wisely to consider the possibility of varying part characteristics without the need to change the whole design cycle. This would lead to a modular design hierarchy. The product is also divided into sub products that are accessible and can be easily assembled. This would lead to a design tree that is shown in figure (4). To the authors' knowledge and vast experience in practice, this is the best entry-level product design. As shown, Rotor Assy with WB-100-00 code is the main sub-assembly for this product. It includes rotating parts such as shaft, rotor discs, spacers, and sleeves. Housing Assy or WB200-00 code includes casted nacelle, stator ring, discharge and charge cavities. Cabs and related attachments to make preload on bearings and sealing sets are also known as Cab Assy, WB300-00. In order to keep the water brake system suspended, two retainers named as Pedestal Assy or WB-400-00 are also designed with cylindrical bearings. The skid named as Stand Assy or WB-500-00 is also designed in order to settle the mechanical core. Other subparts that are connected to the mechanical core are known as accessories with code name WBA00-00. Using this approach makes it possible to check the quality of the product in every section with more accuracy. It leads assuring quality in every step of the design, development, and manufacturing of the product. It fulfills the pure meaning of quality control and quality assurance [15]. In addition, because the design tree divides the product to sub-products that are simpler, the development process is easier. Also, maintenance is easier and more justified because in the design process the function of each sub assembly is the basis of product development and manufacturing process. In addition, primary calculations are based on the database that is predicted from othermanufacturers\u2019 data like Phoenix, Kahn, Taylor and Power Link. CAD model would also help visualizing the product before the costly prototype is built as seen in figure (5). The proposed algorithm for the preliminary design process is presented in figure (6). Downloaded From: http://proceedings.asmedigitalcollection.asme.org/ on 01/26/2016 Terms of Use: http://www.asme.org/about-asme/terms-of-use 4 Copyright \u00a9 2014 by ASME" + ] + }, + { + "image_filename": "designv11_71_0003878_s11434-014-0376-5-Figure12-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0003878_s11434-014-0376-5-Figure12-1.png", + "caption": "Fig. 12 (Color online) The directions of ground reaction force corresponding to the minimum of joint moments", + "texts": [ + " The variations of the joint moments M1, M2, M3, M4 and the overall evaluation M are showed in Fig. 11. In Fig. 11, the minimum values of four joints are zero, it shows that the single joint moment could get to the minimum when the ground reaction force along the corresponding directions of the ground reaction force hF. Therefore, Table 5 shows the different values of hF corresponding to the minimum of joint moments. As shown in Table 5, the values of hF are different when each joint comes to the minimum value, and Fig. 12 indicates that the line where the direction of ground reaction force lies goes through the joint when the joint moment is minimum. For example, when hF is equal to hF2, M2 is minimum, and the line determined by hF2 is almost through the wrist joint. Obviously, the arm of the ground reaction force to the wrist joint reaches the shortest. Making one joint moment reach minimum is not the desired one, which will make other joints moments large, and it is just the local optimum. In order to keep the balance between the four joints, the overall evaluation M is used to measure the degree of comfort for the horse" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0002078_j.matpr.2020.09.582-Figure2-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0002078_j.matpr.2020.09.582-Figure2-1.png", + "caption": "Fig. 2. The three points bend test s", + "texts": [ + " of nodes BB-1 Shell 1158 1338 BB-2 Shell 1825 1967 BB-3 Shell 3541 1790 From the various material data obtained during a tensile test, several three-point bend tests were simulated using FEA modeling. Table 3 shows the meshing details for all the three bumper beams. The meshes for all the three bumper beams are selected after the mesh convergence study (where several meshes are simulated. The final mesh is selected shows less than 5% variation in maximum stress and a maximum 3% variation in deformation value as compared to the finer mesh. However, FEA tests corresponding to the final mesh generated on the proposed three bumper beam are discussed in this paper. Fig. 2 (a) shows the experimental three points bend test set up for the BB-1 bumper beam. A very similar arrangement is provided for the bumper BB-2 and BB-3; hence their figure is not shown here. For numerical simulations, two rollers applied at the end for roller support while one roller with the same diameter used to apply the force in a downward direction. Fig. 2 (b) shows the numerical setup for the BB-1 bumper beam. Fig. 3 (a) shows the Load (kN) versus displacement (mm) graph generated during a three-point bend physical test for bumper beam BB-1. Whereas Fig. 3 (b) shows the graph of Resultant force (kN) versus time in second obtained by numerical simulations. From these figures it can be seen that initial the displacement changes linearly with a load applied, then plastic deformation occurs and at maximum load value of 1.167 kN the bumper beam deforms continuously to a maximum value of 176" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0001588_aim43001.2020.9158991-Figure5-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0001588_aim43001.2020.9158991-Figure5-1.png", + "caption": "Figure 5. PtM holder design", + "texts": [], + "surrounding_texts": [ + "The general view of TakoBot demonstrated in Fig.1. TakoBot consist of the three main components: continuum part, pretension mechanism, and actuating part. A. Continuum part design TakoBot continuum part consists of ten segments divided by two actuating sections: first section (distal end) and the second section (proximal). The segments are connected by universal joints and supported by four compressional springs to maintain a linear and bending stiffness between an adjacent segment. Universal joints are interconnected by linear guide shaft with 30mm of length, which passes through the linear 978-1-7281-6794-7/20/$31.00 \u00a92020 IEEE 460 Authorized licensed use limited to: University of Wollongong. Downloaded on August 11,2020 at 10:23:43 UTC from IEEE Xplore. Restrictions apply. bearing, fixed in the center of each spacer discs. The wires are aligned along with the wire eyelets on each disc (Fig. 2). Such a structure so-called sliding disc mechanism provides smart bending stress distribution along the slender part and improves manipulator bending features by decentralizing forces and bending torques [24]. The maximum traveling displacement of sliding discs is 10mm. The spacer disc diameter is 50mm, and the total length of the continuum part is 380 mm. The wire passes along the spacer discs via wire eyelets and passes through inside of the compressional springs (Fig. 3). In this prototype, we used a stainless-steel wire with a diameter of 1mm, and the eyelet hole size is about 1.5mm. Based on previous experiences, the hole size was determined to be close to the wire diameter. Otherwise, it might cause twist deformation in the continuum part. B. Pretension mechanism design The pretension mechanism (PtM) located in the middle part of the robot. Driving wires passes through the PtM device. In this robot, we drive eight wires, four wires for each section. A linear actuating unit drives wires; in this robot, one motor drives two paired wires with a strain and ease manner, Total four motors actuate eight wires (Fig 4). We modified the previous design of the passive pre-tension mechanism [22], with equipping linear potentiometers (Fig.6) The new PtM device consists of PtM holder and center holder, where all components are fixed, and PtM slider, which slides along the guide shaft and moves linear potentiometer. The compressional springs are equipped along the slider pass. The PtM yields wire-tension and provides tension information that is obtained by measuring the compression length of springs by the linear potentiometers. (Fig.6). Figure 6. Fabricated PtM device All of the components are made of ABS. The guide shaft and slide potentiometer length are 60mm. We used 50mm helical compression springs with constant 1.8 N/m and with a 35% deflection rate. This means the traveling distance of the PtM slider is about 15-20 mm (Fig 7). Spring linear potentiometer PtM slider 461 Authorized licensed use limited to: University of Wollongong. Downloaded on August 11,2020 at 10:23:43 UTC from IEEE Xplore. Restrictions apply." + ] + }, + { + "image_filename": "designv11_71_0003771_sisy.2014.6923592-Figure1-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0003771_sisy.2014.6923592-Figure1-1.png", + "caption": "Figure 1. Stick diagrams of the primitives: a) leg bending, b) leg stretching, c) shifting the trunk forward", + "texts": [ + " Secondly, each primitive is parameterized, so that a form of the realization is adjusted by choosing the appropriate parameters. To obtain the complex motion it is only necessary to determine the parameters of the primitives and the sequence of their realization. There are no limitations in the defining of new primitives, either in the number of the parameters or in the number of joints that will be activated in its realization. Several illustrative examples of the primitives which are used for the walk synthesis are presented in Fig. 1. Fig. 1 a) shows the initial and final postures for leg bending, Fig. 1 b) leg stretching., and Fig. 1 c) shifting the trunk forward. Therefore, characteristics of primitives can be summarized as follows: \u2022 A primitive defines a simple movement in which more joints can be involved. \u2022 Primitives are parameterized, so that the movement shape the can be altered during its realization. The number of parameters can vary. \u2022 Realization of a primitive does not depend on the current posture, but only on whether the robot is in such a posture from which is possible to perform the given primitive. For example, for leg stretching, the stretching leg must not be in contact with the ground; for leg bending, the ZMP should be inside the support area formed by the foot which will remain in contact with the ground" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0003752_2786545.2786551-Figure2-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0003752_2786545.2786551-Figure2-1.png", + "caption": "Figure 2 Behavior of the tunnel rescue robot", + "texts": [ + " Suppose that we develop an automatic control based on the previous manual rescue robot. The robot provides multiple services: investigating at the emergency and rescuing something. Additionally, the robot contains two motion mechanism services: the ground mode and the flight mode. Some people does not recognize motion mode as service. However almost people will think that airplane service is clearly different from vehicle service. In this paper, service means that it looks independent software. Figure 2 shows a behavior of the robot. The mode service changes will occur dependent on the local terrain. Thus, the requirement of the above is as follows. Requirement: (1) Prepare unexpected abnormal situation (2) Support the ground mode service and the flight mode service (3) Provide the investigation service and the rescue service Those requirements are helpful for considering our three problems for multi-purpose robots. To implement the above requirement, we consider a simple structure of the robot in Figure 3" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0003878_s11434-014-0376-5-Figure7-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0003878_s11434-014-0376-5-Figure7-1.png", + "caption": "Fig. 7 (Color online) The skeletal structure of the hind limb and the plane five-link mechanism [14]", + "texts": [ + " All major bones of the leg are supposed to be in the longitudinal plane of the horse; 4. The mass of the leg, compared with the horse weight, is too light and ignored. According to the assumptions above, the skeleton-joint leg structure could be regarded as the link-revolute pair structure. Taking the skeletons as rigid bodies and joints as the revolute pairs with active drives, the forelimb and hind limb of the horse turn into a plane open-chain five-link mechanism. The corresponding relations are shown in Table 2 and Table 3. In Fig. 6 and Fig. 7, A is the forelimb foothold on the ground and B is the hind limb foothold. H represents the height of the shoulder joint in forelimb and the hip joint in hind limb. L refers to the horizontal distance from foothold A to the shoulder joint, or from foothold B to the hip joint. The forelimb and hind limb both can be equivalent to plane open-chain five-link mechanisms. Although the length of links and the postures of two mechanisms are different, the characteristics of two mechanisms are similar. Therefore, the forelimb model is chosen as the object of force analysis, which could be also applied to the hind limb" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0001381_s1068798x20060210-Figure4-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0001381_s1068798x20060210-Figure4-1.png", + "caption": "Fig. 4. Device described in Soviet Inventor\u2019s Certificate 81662 (1948).", + "texts": [ + " In the Soviet Union, in 1948, S.I. Dobroborskii recorded his invention of a device for testing jackham- ( ) , 2 hA nG H \u03a3\u0394\u03a3 = + \u2212 \u03a3\u03b1 n G H \u0394h a ( ) ( )2\u0394 1 , 1 G H h k a m M + \u2212 = + k m M 2020 mers (Inventor\u2019s Certificate 81662). He proposed the testing of jackhammers to determine their operational parameters. In this approach, the work performed by the jackhammer is determined by the displacement of a measuring piston in a cylinder, with the action of air pressure on the piston. The system is presented in Fig. 4. The jackhammer (not shown) is mounted in chuck 1, which is rigidly connected to piston 2 of cylinder 3. Compressed air is supplied to the cylinder through valve 4, which keeps the cylinder (below the piston) at constant pressure; the pressure is monitored by manometer 5. The amplitude of the piston oscillations under the action of the hammer face is measured by means of a contact head with electrocontact probe 6, which moves vertically as manual lever 7 rotates. The position of probe 6 is regulated by microscrew 8 so that the position of lever 7 at the zero of scale 9 corresponds to contact of the probe 6 and the tip 10 of piston 2 in its upper position. At contact, the electrical circuit with battery 11 is completed (shown by dashed lines) and neon lamp 12 is turned on. Under the action of spring 13, lever 7 takes a vertical position (as shown in Fig. 4), while probe 6 moves from tip 10 to a distance greater than the maximum possible motion of the tip from hammer impact. Each division of the instrument\u2019s scale 9 corresponds to 1 mm of vertical motion of probe 6. When the jackhammer is turned on, tip 10 (together with piston 2) performs reciprocating action of specific amplitude. With clockwise rotation of lever 7, the tip 10 touches probe 6 at a particular lever position on scale 9. The electric circuit is completed, and lamp 12 is turned on. To avoid sparking and damage to the contact and to ensure safe operation, the voltage in the primary circuit supplied by battery 11 is 6 V, while the voltage in RUSSIAN the secondary circuit of lamp 12 is 110 V, supplied by transformer 14" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0001117_kem.841.327-Figure11-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0001117_kem.841.327-Figure11-1.png", + "caption": "Fig. 11 Directional Deformation.", + "texts": [ + " 10 shows the stress in the Sugarcane fibre on tensile loading. A maximum stress of 31.34 MPa is obtained. GFRP layer exhibits 2.65 times better stress distribution than Sugarcane fibre. This is due to the better mechanical properties of the GFRP. Also, the presence of the quasi- isotropic stacking order of the GFRP fibre enhances the load bearing capacity of the composite specimen. Graph 3. Stress vs Strain of tensile test from FEA Model. Fig. 9 Stress in GFRP. Fig. 10 Stress in Sugarcane. Flexural Stress Vs Strain. Fig. 11 shows the deflection obtained on the flexural loading on the FEA Model of the composite. A displacement of 7.5 mm was obtained in the experimental results, whereas in the FEA model it is only 3.5 mm. This indicates that the resistance to flexing of the material is better in FEA model. Equivalent Stress. Fig. 12 shows the stress distribution on the FEA model. It can be seen that the stress is minimum is the regions of support and maximum at the point of support. The stress is maximum at the outermost layers of top and bottom" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0002485_cac51589.2020.9327292-Figure3-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0002485_cac51589.2020.9327292-Figure3-1.png", + "caption": "Fig. 3: The symbol definition of H", + "texts": [ + " In [11], the cascade PID controller is used to conduct the flight experiment of Bi-copter UAV, and the kinetic equations of Bi-copter\u2019s model are established in [12]. Compared with the tiltable quad-rotor UAV established in [8], the influence of the position of opi along the axis zb in the body coordinate system is more considered. Therefore, (3) is revised as: ob pi = Rz((1 \u2212 i) \u03c0 2 ) [ L 0 H ]T , i = 1 . . . 4. (13) where H is the distance between the center of gravity ob to the axis of rotation opi in the zb direction, and H is positive while ob below opi (as Fig. 3 shown). Then, in the condition of P2 and P4 have failed, the roll channel of the aircraft needs to balance the moment generated by the component force of the propeller thrust Fpi in the tangential direction of the obopi axis along zb (as Fig. 4 shown). Similarly, when P1 and P3 have failed, the pitch channel needs the moment generated by the component force of the Fpi in the tangential direction of the obopi axis to balance. III. Controller design Based on the classical control theory, this paper linearizes the tilting quad-rotor model proposed in Section II" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0002119_icaccm50413.2020.9212874-Figure8-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0002119_icaccm50413.2020.9212874-Figure8-1.png", + "caption": "Fig. 8. Equivalent stress for CFRP propeller(case1).", + "texts": [], + "surrounding_texts": [ + "A Quadcopter\u2019s propeller is subjected to horizontal bending due to collision of the propeller with obstacles or a wall. To calculate the bending deformation under the applied rotational moment about the rotation axis, the propeller has been fixed by both the tip as shown in Figure 10.In addition, it is fixed at the point where it is connected to the motor shaft (Figure 11).Figure 12 shows the applied rotational velocity of 897 rad/s [8].Figure 13, 14 and 15, 16 show the total deformation and equivalent stress for CFRP and GFRP materials, respectively. The comparison of the obtained results is presented in Table III. 61 Authorized licensed use limited to: Carleton University. Downloaded on May 30,2021 at 15:29:13 UTC from IEEE Xplore. Restrictions apply. Fig. 14. Total deformation of GFRP propeller (case2). Fig. 15. Equivalent stress for CFRP propeller (case2). Table III. Results due to rotational velocity CFRP GFRP Max. Deformation 0.2012 mm 0.268 mm Max. stress 807.54 MPa 1127 MPa C. Vibration analysis The propeller has been tested for vibrational effect to find out the resonance frequencies for both the materials. The modal analysis has been performed in Ansys, Figure 17 and 18 show the frequency variation with first six mode shapes for CFRP and GFRP material, respectively. The color contour describes the range of deformation from minimum to maximum. Table 4 and figure 19 present the frequency comparison for CFRP and GFRP materials and figure 19 also present the comparison with previously publish work. Fig.17. CFRP based propeller\u2019s vibration frequencies and mode shapes (case3). 62 Authorized licensed use limited to: Carleton University. Downloaded on May 30,2021 at 15:29:13 UTC from IEEE Xplore. Restrictions apply. IV. CONCLUSION A Quadcopter\u2019s propeller has been analyzed for the structural and vibrational aspects. The propeller has been analyzed under the thrust (case1) and rotational loading (case2) conditions to find out the stress and deformation produced in two materials. It is found that for case1 both the materials show almost same results, while for case2 GFRP shows the high values of maximum deformation and stress compared to CFRP. Moreover, from the results of the modal analysis, it is found that CFRP is having high values of frequencies compared to GFRP. From these observations, it can be concluded that CFRP perform better than GFRP material under static and vibration loading condition. Furthermore the natural frequency of propeller is validate with publish literature [10] in figure 19. REFERENCES [1] Ahmad, F., Kumar, P., Bhandari, A., & Patil, P. P. (2020). Simulation of the Quadcopter Dynamics with LQR based Control. Materials Today: Proceedings, 24, 326-332. [2] Ahmad, F., Kumar, P., & Patil, P. P. (2018). Modeling and simulation of a Quadcopter with altitude and attitude control. Nonlinear Studies, 25(2). [3] Wei, P., Yang, Z. and Wang, Q. (2015). The Design of Quadcopter Frame Based On Finite Element Analysis,In 3rd International Conference on Mechatronics, Robotics and Automation. Atlantis Press. [4] Jaouad, H., Vikram, P., Balasubramanian, E., & Surendar, G. (2020). Computational Fluid Dynamic Analysis of 63 Authorized licensed use limited to: Carleton University. Downloaded on May 30,2021 at 15:29:13 UTC from IEEE Xplore. Restrictions apply. Amphibious Vehicle. In Advances in Engineering Design and Simulation (pp. 303-313). Springer, Singapore. [5] Kumar, S., & Mishra, P. C. (2016). Finite element modeling for structural strength of quadcoptor type multi modevehicle. Aerospace Science and Technology, 53, 252- 266. [6] Ahmed, M. F., Zafar, M. N., & Mohanta, J. C. (2020, February). Modeling and Analysis of Quadcopter F450 Frame. In 2020 International Conference on Contemporary Computing and Applications (IC3A) (pp. 196-201). IEEE. [7] Martinetti, A., Margaryan, M., & van Dongen, L. (2018). Simulating mechanical stress on a micro Unmanned Aerial Vehicle (UAV) body frame for selecting maintenance actions. Procedia manufacturing, 16, 61-66. [8] Singh, R., Kumar, R., Mishra, A., & Agarwal, A. (2020). Structural Analysis of Quadcopter Frame. Materials Today: Proceedings, 22, 3320-3329. [9] Xiu, H., Xu, T., Jones, A. H., Wei, G., & Ren, L. (2017, December). A reconfigurable quadcopter with foldable rotor arms and a deployable carrier. In 2017 IEEE ROBIO (pp. 1412-1417) [10] Ahmad, F., Bhandari, A., Kumar, P., & Patil, P. P. (2019, November). Modeling and Mechanical Vibration characteristics analysis of a Quadcopter Propeller using FEA. In IOP Conference Series: Materials Science and Engineering (Vol. 577, No. 1, p. 012022). IOP Publishing. 64 Authorized licensed use limited to: Carleton University. Downloaded on May 30,2021 at 15:29:13 UTC from IEEE Xplore. Restrictions apply." + ] + }, + { + "image_filename": "designv11_71_0000762_icsai48974.2019.9010588-Figure3-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0000762_icsai48974.2019.9010588-Figure3-1.png", + "caption": "Figure 3. Three-phase magnetic latching relay model diagram", + "texts": [ + " Since the model contains magnetically permeable materials, the use of the three-dimensional vector method will generate a large calculation error, so the magnetic scalar method is used. Using the current source module SOURC36 unit, unlike the other solid model, the SOURC36 unit does not need to be split, but defines the unit type, defines the real constant, and then directly generates. The SOURC36 unit consists of three shapes, a ring unit, a rod unit and an arc unit. The contact system is built by the rod-shaped and arc-shaped units, and the integrated models of the two relay contact circuits and the electromagnetic system are respectively established, as shown in FIG. 3 and FIG. 4. III. INFLUENCE OF CONTACT LOOP ON MAGNETIC LATCHING With the wider use range of magnetic latching relays, the current requirements for the contact loop load are also higher and higher. Under large current conditions, the magnetic field generated by the contact loop will cause electromagnetic interference to the electromagnetic system. Since the relay is required to remain stable for a certain period of time under the condition of short-circuit current, this paper simulates the contact loop under steady-state current and short-circuit current respectively, and analyzes the influence of the contact loop on the magnetic latching force" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0001028_s1052618820020053-Figure1-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0001028_s1052618820020053-Figure1-1.png", + "caption": "Fig. 1", + "texts": [ + " This work aims to develop a procedure for calculating mixed oscillations in systems with limited excitation based on direct linearization methods. This procedure for applying direct linearization can be used in the presence of a delay and limited power of the energy source in a nonlinear system. The procedure is quite simple, which is especially important from a practical point of view for calculating the parameters of mechanisms, machines, and equipment at the design stage. 2. Consider the well-known model (Fig. 1) of a self-oscillating system [1, 3, 17, 18]. Taking into account the delayed elastic force, the motion of the system is described by the equations (1) where is an external force with constant amplitude and frequency; T(U) is a nonlinear friction force dependent on the relative velocity , causing self-oscillations; is the radius of the application point of the friction force at the contact point of a body of mass m with the ribbon; is the damping coefficient; and are the linear and nonlinear parts of the spring + + + = + \u2212 = \u2212 0 0 1 \u03c4 0\u03bb sin \u03bd ( ) ( ), \u03c6 (\u03c6) ( ),mx k x c x c x t T U f x J M r T U \u03bb sin \u03bdt = \u2212 U V x = 0 \u03c6rV =0 constr ( )T U 0k +0 1 \u03c4c x c x ( )f x 105 force, respectively; , , , is the delay; J is the total moment of inertia of the rotating parts; is the difference between the torque of the energy source and the moment of the forces of resistance to rotation, and is the rotation velocity of the engine" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0000886_012117-Figure1-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0000886_012117-Figure1-1.png", + "caption": "Figure 1. Deviation of the position of the flat two-link manipulation system", + "texts": [ + " The best value of the fitness function (Best fitness) is determined by the direction of optimization, by finding the maximum or minimum of fitness function. When analyzing the convergence process, the average value of fitness function by population (Mean fitness) is often used. As a fitness function for solving inverse kinematics problem of manipulating robots, we will use a function of calculating the square of position deviation of the robot\u2019s output link characteristic point from selected point on a motion path. We use the genetic algorithm to search for values of hinge coordinates minimizing fitness function using example of a flat two-link (figure 1) and three-link spatial manipulation systems (figure 2). For a two-link manipulation robot, the chromosome will be a vector X=[x1 x2]T which elements are the hinged coordinates, x1 \u2013 is rotational, \u0430 x2 \u2013 is translational. For these coordinates, we introduce restrictions , . We make up a fitness function where Y=[y1 y2]T \u2013 vector of Cartesian coordinates of point selected on the trajectory. We set coordinates of selected point y1= 1.0000 m \u2013 rectangular coordinate along x axis, y2= 0.0000 m \u2013 rectangular coordinate along y axis" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0000347_iecon.2019.8927574-Figure1-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0000347_iecon.2019.8927574-Figure1-1.png", + "caption": "Fig. 1. Schematic of the stator structure of flux switching PM machine.", + "texts": [ + " Although some of the presented structures are really potential for the aforementioned applications, their complex structures cause them not to be widely produced; some examples of these structure are stator-PM machines, AxialField PM (AFPM) machines, Transverse Flux PM Machine (TFPM), Claw Pole machines, Spoke-Type PM machines. On the other hand, these structures can have higher efficiency, as well as compact sizes in comparison with the traditional PM machines. Stator-PM machines are an attractive structure due to placing both magnets and coils in the stator structure only. Flux Switching PM (FSPM) machines (Fig. 1) and Flux Reversal PM (FRPM) machines (Fig. 2) are two common stator-PM structures [2]-[4]. Placing the PMs in the stator structure makes these structures an appropriate candidate for transportation applications because they have salient and robust rotor without PMs or windings; good thermal dissipation as the magnets located on stator; and the armature field has little influence on the PM due to the PM field and the armature field are in parallel [5]. For instance, linear FSPM machines could be an interesting solution for the railway application because both PMs and windings are positioned on stator and translator does have a very simple and cheaper structures in compared to the other typical types [5], [6]" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0003632_0954406214549786-Figure1-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0003632_0954406214549786-Figure1-1.png", + "caption": "Figure 1. Configuration of the tilting table driven by worm and worm wheel.", + "texts": [ + " In this paper, the variation of natural frequency in the tilting direction of the tilting table driven by worm and worm wheel with tilting angle is studied with a lumped parameter model. In addition, the effect of three stiffnesses on the system natural frequency is investigated, respectively, including the equivalent tangential meshing stiffness of worm and worm wheel, the torsional stiffness of the wormwheel shaft, and the axial stiffness of worm supporting bearings. Dynamic model of the tilting table with tilting angle The tilting table driven by worm and worm wheel is shown in Figure 1. The tilting base, assembled with a rotary table and supported by two groups of angular contact ball bearings, is connected to a worm wheel by a shaft. The worm links the servo motor with a coupling, which is constrained by two groups of supporting bearings. A global Cartesian coordinate system O(x,y,z) is established in the center of the rotary table. The position of the center of mass G of the titling table will change when the tilting base rotates an angle s around x-axis (right hand rule). Equivalent dynamic model in the tilting direction The dynamic characteristics of the tilting table driven by worm and worm wheel are mainly decided by the stiffness and damping of the worm wheel shaft, worm gear pair, supporting bearings, worm shaft, and the coupling", + "comDownloaded from System natural frequency and sensitivity at different tilting angles The natural frequency f s\u00f0 \u00de of the tilting table in the tilting direction is given by equation (17) f s\u00f0 \u00de \u00bc 1 2 ffiffiffiffiffiffiffiffiffiffi K s\u00f0 \u00de Jeq s \u00f017\u00de Taking partial derivation of natural frequency f s\u00f0 \u00de with respect to part stiffness, we can get the sensitivity Si of f s\u00f0 \u00de to changes in the part stiffness as follows29 Si s\u00f0 \u00de \u00bc @f s\u00f0 \u00de @Keq \u00bc 1 4 ffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi JeqK s\u00f0 \u00de p @K s\u00f0 \u00de @Keq \u00f018\u00de where Keq is the equivalent torsional stiffness of each part of the transmission system as shown in equations (2) to (5). The tilting table in this study is driven by dual-lead ZN type worm and worm wheel as shown in Figure 1. The worm is dual-lead with one tooth, 3mm module, 15 pressure angle, and 52.4mm diameter of the pitch cylinder. The different lead angles on the pitch cylinder are 3:191 and 3:828 . The worm is made of steel with an elasticity modulus of 2:06 1011Pa, a density of 7850 kg=m3, and a Poisson ratio of 0.30. The worm wheel has 72 teeth with an elasticity modulus of 1:08 1011Pa, a density of 8760 kg=m3, and a Poisson ratio of 0.35. The total mass ms and the moment of inertia about x-axis Jeq of tilting parts are 201" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0000662_icus48101.2019.8995953-Figure2-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0000662_icus48101.2019.8995953-Figure2-1.png", + "caption": "FIGURE 2. Schematic drawing of multi-DOF planar robot with revolute joints and with a single reaction wheel attached to each link.", + "texts": [ + " Since the controller needs to gather the measurements from the sensors, simulate the system for one-step ahead prediction, compute the control sequence and transmit it to the actuators in one loop, the following condition must hold true: Tsens + Tsim + Tsolv + Ttran \u2264 Ts (8) 4620 VOLUME 4, 2016 However, even if at a given time step this condition is violated, the system will still be fault-tolerant to a limited extent, since it has simulated and generated controls forKNMPC steps ahead in the previous time step; therefore, it can skip one step and apply the next control inputs in the queue. For a more detailed description of this framework the reader is referred to [19]. III. REACTION WHEEL INTEGRATED VSA ROBOTS A. SYSTEM MODELING In this subsection we describe the dynamics of VSA robots with reaction wheels. For simplicity, we limit our treatment to planar VSA robots which have one reaction wheel attached to every link. We also assume that robot links are moved by rotational joints only. The schematics of a such VSA robot is shown in Fig. 2. Without loss of generality, it is assumed that the center of mass of the momentum wheels are coincident with their axis of rotation. The described procedure can be extended to non-planar multi-DOF VSA robots with rotational joints. Angular position and velocity of link i are described by \u03c8i and \u03c8\u0307i, respectively. Angular position and velocity of the wheel, attached to the link i, are denoted by the generalized coordinates \u03b6i and \u03b6\u0307i. A joint i is connected to the proximal end of the link i and causes its motion" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0001138_0954406220925836-Figure2-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0001138_0954406220925836-Figure2-1.png", + "caption": "Figure 2. Articulated moving platforms with two directions: (a) AMP1[ AMP2; (b) AMP2[ AMP2; (c) AMP1[ AMP3; (d) AMP2[ AMP3.", + "texts": [], + "surrounding_texts": [ + "Parallel manipulator, articulated moving platform, rotational capability, type synthesis, constraint synthesis Date received: 6 March 2020; accepted: 19 April 2020" + ] + }, + { + "image_filename": "designv11_71_0003451_s00707-015-1403-6-Figure13-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0003451_s00707-015-1403-6-Figure13-1.png", + "caption": "Fig. 13 Special case of motion of the NNC model with orthogonal velocities", + "texts": [ + " Both multipliers of constraints will simultaneously equal zero when \u03d5\u0307 = 0, \u03bb1 = \u03bd1\u03d5\u0307 = 0 and \u03bb2 = \u2212\u03bd2\u03d5\u0307 = 0. If \u03d5\u0307 = 0 it follows that \u03d5 = \u03d50 = const, which means that the direction M1M2 is fixed by the angle \u03d50, and it is the only possible direction along which the particles M1 and M2 can move. However, the particle M1 must then be at rest due to a blade positioned in it perpendicular to the possible direction of M1M2 motion, while the particle M2 can move along the direction M1M2 because the blade is positioned in it exactly in the direction of the possible motion of M1M2\u2014Fig. 13. It is then that \u03bb1 = \u03bd1\u03d5\u0307 = 0 on account of \u03d5\u0307 = 0, and due to the constraint in M1, it follows also that \u03bd1 = 0, \u03bb2 = \u2212\u03bd2\u03d5\u0307 = 0 on account of \u03d5\u0307 = 0, and due to the constraint in M2, it follows that \u03bd2 = 0! If it were \u03bd2 = 0, it is then \u03bb2 = \u2212\u03bd2\u03d5\u0307 = 0, and the particle M1 would be able to move along the circle of radius \u03be0 ( M20M1 ) and then it would be \u03bd1 = 0 and \u03d5\u0307 = 0 and \u03bb1 = 0, respectively! The next case is: if it were \u03bb1=0 and \u03bb2=0, it would be the case when the velocities of both particles equal zero, i" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0000967_0954406220916504-Figure5-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0000967_0954406220916504-Figure5-1.png", + "caption": "Figure 5. Ground-wheel contact model.", + "texts": [ + " For the vehicle traveling forward with a constant speed v, the sum of side forces must be equal to the centrifugal force, i.e. FY \u00bc FC. Considering a moment equilibrium of the vehicle about the Z-axis, it becomes clear that RY \u00bc 0. Therefore, the sum of the tyre side forces w\u0302Y, as shown in Figure 4, can be rewritten as w\u0302Y \u00bc FC 0 1 0 0 0 0 T . Consequently, the sum of the wrenches w\u0302C and w\u0302Y becomes w\u0302 \u00bc w\u0302C \u00fe w\u0302Y \u00bc FC 0 0 0 Zcg 0 0 T \u00f012\u00de and acts on the vehicle body during cornering. The wheel of the vehicle maintains contact with the road surface at the contact point P, as shown in Figure 5. Since the wheel can slide and rotate about P, the contact may be modelled as a kinematic pair consisting of a plane pair (E-joint) and a spherical pair (S-joint). E-joint can be represented by two intersecting P-joints with two twists S\u0302Ex \u00bc 1 0 0 0 0 0 T and S\u0302Ey \u00bc 0 1 0 0 0 0 T . S-joint can be replaced by three orthogonally intersecting R-joints whose motions are expressed by S\u0302Sx \u00bc rP sx sx , S\u0302Sy \u00bc rP sy sy and S\u0302Sz \u00bc rP sz sz , where sx \u00bc 1 0 0 T , sy \u00bc 0 1 0 T , sz \u00bc 0 0 1 T and rP is the position vector from origin O to the contact point P" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0000923_ibcast47879.2020.9044556-Figure8-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0000923_ibcast47879.2020.9044556-Figure8-1.png", + "caption": "Fig. 8. Faulty ball bearings with a single hole: (a) IRF; (b) ORF; and (c) BBF.", + "texts": [], + "surrounding_texts": [ + "Equation (10) indicates that, by using the absolute value in the PCA, the total SSE distances are always minimized regardless of the bases signs. Rotor speed-based BFD using the proposed AVPCA is described in the following section.\nThe presence of bearing faults influences characteristic parameters, which causes the parameters to vary. A sample\nformed with the variations of the faulty parameters will contain specific information related to the fault itself. AVPCA extracts a number of characteristic vectors from these samples that are mutually orthogonal. To develop the AVPCA model for BFD, the study is divided into two procedures: 1) an offline training procedure; and 2) an online fault detection and diagnosis procedure. Hence, the collected speed signal data set is divided into two parts: training data (known data) and testing data (unknown data). The training data are regarded as the historic data and used to build the training database, whereas the testing data are used to verify the performance of the proposed RSB-BFD.\nTable I describes the offline training procedure, which contains the AVPCA steps.\nThe overall offline training procedure of the proposed AVPCA is summarized in Fig. 5.\nAfter the offline training of the system in both healthy and faulty cases, the algorithm is ready for the online fault detection and diagnosis procedure, which is composed of the steps described in Table II.\nThe total SSE is given by \u039e = \u03be1 + \u03be2 + \u00b7 \u00b7 \u00b7+ \u03ben, which will be minimized at each stage of the cluster analysis. Considering a predefined threshold \u03c1, the fault can be detected. The overall online fault detection and diagnosis procedure for the proposed RSB-BFD is summarized in Fig. 6.", + "Once the new test data set has been projected onto the AVPCA subspace, the fault has been detected. The fault diagnosis procedure is then used to diagnose a detected fault by projecting the new set of the test data into the subspaces spanned by the principal component vectors of the training set. If the similarity between the set and subspace is greater than the similarity with any other class, then the membership of the new set in a class can be determined using the SSE distance because this distance can achieve better extraction results with less training compared with other distances [32]. The fault is assigned to a class under the given hypothesis of the minimum distance SSE, which is used as a basis for testing and further investigation of the fault diagnosis task.\nTo evaluate the performance of the proposed RSB-BFD method, an experimental setup was built with a 750-W brushless direct current (BLDC) motor powered by a TMC-7 BLDC motor driver. The rotor speed was the only quantity measured. The schematic diagram of the experimental setup is illustrated in Fig. 7. Two ball bearings (NSK 6204) with eight balls were integrated in the experimental setup. The tested ball bearings were artificially damaged to produce flaws on the outer race, inner race, or ball. The flaws consisted of 1-mm holes that were drilled axially through the outer and inner raceways and the ball (see Figs. 1 and 8). A flywheel was added to the experimental setup to generate a constant load torque. All measured quantities were collected with a national instruments (NI) cDAQ-9178 eight-slot universal serial bus (USB) chassis. The NI9411 module was used for the rotor speed signal. The measured data were sampled at 17.06 kHz and processed using MATLAB R2012a.\nThe rotor speed signal for the proposed algorithm was measured with an incremental encoder type E60H NPN open collector output at 1024 pulses per revolution with a maximum allowable revolution of 6000 r/min.\nTwo different environments were considered: 1) constantspeed environment; and 2) variable-speed environment. First, a constant motor speed of wr = 2500 r/min was considered while the bearing fault detection and diagnosis algorithm was running, as shown in Fig. 9.\nFor the offline training procedure, a set of historical process data from the motor rotor speed signal under different faulty bearing conditions (ORF, IRF, and BBF) and the healthy condition (BFF) was first measured and sampled to form the training data vectors Wm\nrj (k) given by (11).\nThe SCREE [33] plot indicates that the number of PCs can be chosen as one (p = 1). Using the CPV approach to obtain a precision of 99.9%, the number of principal components is fixed at three (p = 3), as shown in Table III. The AVPCA projection reduces the original n-space to a 3-D subsystem in p-space.\nFig. 10 shows the variation of the residual weights rhi of the AVPCA subspace in 3-space for the healthy case and the three", + "TABLE III CPV PRECISION OF THE EIGENVALUES\nFig. 10. Variation of residual weights.\nAfter performing the training procedure in the healthy and faulty cases, a new set of measurements from the rotor speed signal was considered to form the data test vectorWnew\nr (k) that contains the unknown rotor speed signal measurements. Faults can thus be detected by evaluating the SSE distance between the data test vector and training vectors, as described in (22). By evaluating the SSE distance variation, computing the total SSE, and considering its minimum, the fault was detected and then diagnosed. The results of the total SSE distance with the proposed AVPCA for RSB-BFD at a constant-speed condition are summarized in Table IV.\nBearing fault detection and diagnosis under variable speed is more challenging though more realistic. Thus, as shown in Fig. 11, four scenarios with different profiles of variable motor speed wr were considered while the RSB-BFD algorithm was running.\nFig. 11 shows the variation of the SSE distance between the unknown data test vector Wnew\nr (k) on the nonfaulty case and the training vectors Wm\nr (k) (where m = 1, 2, 3, 4,) using\nthe proposed AVPCA as the RSB-BFD fault detection method under the four scenarios with different speed profiles and fault cases. In this figure, the SSE distance for the proposed AVPCA can clearly be used to detect any of the faults in all of the measurement sets. The SSE distance variation in the BFF case was always less than the SSE distance variation for all the fault cases (ORF, IRF, and BBF).\nFig. 12 and Table V depict respectively the SSE distance variation and the total SSE distance in tested cases (2), (3), and (4) by the proposed RSB-BFD for fault diagnosis. Fig. 12 shows" + ] + }, + { + "image_filename": "designv11_71_0000896_0954406220912005-Figure4-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0000896_0954406220912005-Figure4-1.png", + "caption": "Figure 4. Kinematic scheme of the ith leg of a RSRR wheel.", + "texts": [ + " Forward kinematic analysis consists in determining the position, velocity, and acceleration, which are described by vectors ti, _ti, and t :: i respectively, of the ith leg when the input movement of the actuators is known, given by vectors q, _q, and q :: . On the other hand, inverse kinematic analysis aims at calculating the motion of the actuators given the desired motion of the leg. The following subsections are devoted to solve the direct kinematics of the mechanisms under study. First, the procedure to carry out this analysis using the RSRR mechanism as an example is presented, and then the mathematical expressions for the remaining cases are obtained. Figure 4 presents a kinematic scheme and the considered vectors for the analysis of the ith leg of a RSRR HeIse wheel. Two reference frames are defined for the analysis. The first one, denoted by 1, is attached to the chassis of the vehicle and its origin is placed on the axis of the wheel, and its plane x1\u2013y1 contains the motion of the traction link. It should be noted that reference frame 1 will not rotate with the wheel due to a change in the variable qr. On the other hand, the reference frame li is attached to the wheel and oriented in such a way that the motion of the rod is contained in the plane xli\u2013yli , and axis zli is parallel to axis z1 of the frame 1, but in the opposite direction. The obtained expressions from the analysis of the ith leg are valid for the rest of the legs by just replacing the value of angle i, which defines the location of each leg around the rim. To simplify the position analysis, please note that the leg can be separated into two sub-mechanisms. From Figure 4, it is possible to write the following closed-loop equation ai \u00fe r2i \u00fe p \u00bc si \u00f01\u00de From this equation, it is possible to calculate the angle between the connecting rod and the sliding shaft, 1, and the magnitude of vector lisi denoted by Sp as follows 1 \u00bc arctan ffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi L1 lp qp 2 1 s0@ 1A \u00f02\u00de and Sp \u00bc Rc2 ffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi L2 1 \u00f0lp qp\u00de 2 q \u00f03\u00de where L1 \u00bc kaik is the length of the connecting rod, Rc2 \u00bc kr2ik is the distance from the wheel axis to the position of joint R1, and lp is the distance from the plane x1\u2013y1 to the reference point from which the value of variable qp is measured", + " Moreover, it is possible to write another closed loop equation such that bi \u00fe ci \u00fe r1i \u00bc si \u00f04\u00de From (4) it is possible to compute the angle of the traction link 2, and the angle of the proximal link 3, both with respect to the axis x1 2 \u00bc 2 arctan Q2 ffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi Q2 1 \u00feQ2 2 P2 1 q P1 \u00feQ1 0@ 1A \u00f05\u00de and 3 \u00bc 2 arctan Q2 ffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi Q2 1 \u00feQ2 2 P2 2 q P2 \u00feQ1 0@ 1A \u00f06\u00de where Q1 \u00bc Sp cos i Rc1 cos 1i, Q2 \u00bc Spsen i Rc1sen 1i, P1 \u00bc \u00f0S 2 p \u00fe R2 c1 2SpRc1 cos\u00f0 i 1i\u00de \u00fe L2 2 L2 3\u00de=2L2 and P2 \u00bc \u00f0S 2 p \u00fe R2 c1 2SpRc1 cos \u00f0 i 1i\u00de \u00fe L2 3 L2 2\u00de=2L3. The geometrical parameters of the mechanism are L2 \u00bc kbik is the distance from the joint R2 to the joint S, L3 \u00bc kcik is the distance from the joint R2 to the joint R3, Rc1 \u00bc kr1ik is the distance from the wheel axis to the joint R3. The angles 1i \u00bc qr \u00fe i and i \u00bc qr \u00fe i \u00fe indicate the orientation of the corresponding vectors and are measured from the x1 axis of the 1 frame as shown in Figure 4. d is the angle between vectors r1i and r2i. Angle i is a parameter that defines the orientation of the vector r1i for each leg. We define a vector c \u00bc 1 Np T to indicate the orientation of each leg. Finally, the position of the tip of the ith leg is obtained as follows ti \u00bc \u00f0lT=L2\u00deR# \u00fe I2\u00bd 1bi\u00fe 1ci\u00fe 1r1i \u00f07\u00de where lT is the length of the traction link and I2 is an identity matrix. Symbol over vectors means that only the first two elements are considered, and the left superscript means the reference frame in which they are measured" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0000642_s12206-020-0134-3-Figure2-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0000642_s12206-020-0134-3-Figure2-1.png", + "caption": "Fig. 2. Set of finite poses to be reached.", + "texts": [ + "B Moreover, a prismatic joint allows a translation of body 3 with respect to link 2 along center line CQ of the corresponding slot. Finally, it should be noted that the guided body, namely, body 4, is assumed to be rigidly attached to and move with the coupler of the linkage, namely, link 2. The input data of the problem under study is a given set of finitely separated poses of a rigid body on a plane. On this regard, the designer has the freedom to define the prescribed poses with respect to an arbitrary coordinate system X Ylocated at ,O as shown in Fig. 2. Thus, the poses are given in terms of a set of position vectors, ,jr and orientation angles, ,j\u03b8 as follows: { } ( )0 , , , , for 0,1,2, , . N T j j j j j\u03b8 X Y j N\u00ba =r r L (1) However, in order to reduce the complexity of the synthesis approach, it is more useful to define the set of prescribed poses with respect to the first pose, that is: { } ( ) ( )0 0 01 0 , , , , , , 1,2, , . N T T i i i i i i i i i i x y X X Y Y for i N f f q q \u00ba - = = - - \u00ba - = p p r r L (2) Thus, for a set of four prescribed poses 3,N = whereas 4N = if we have five poses to be visited" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0003339_wemdcd.2015.7194534-Figure1-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0003339_wemdcd.2015.7194534-Figure1-1.png", + "caption": "Fig. 1: Representation of a three phase induction motor with interturn faulty windings.", + "texts": [ + " The model performance with a detailed description about the MPO is also given. Section III describes the inverse problem formulation, followed by numerical and experimental validation results in section IV. Finally, a conclusion is drawn in section V. In this section, the dynamic state-space model of a distributed-winding three phase induction motor under interturn fault condition is presented. This model computes the three phase currents of the stator and rotor, the electromagnetic torque, and rotor speed for a given supply voltage. Fig. 1 shows the three phase windings of the stator and rotor in the abc reference frame, where the axis of stator phase a is chosen the reference axis. The inter-turn fault is assumed to occur only in phase a, and the other two phases, i.e. b and c, are kept healthy. The mutual inductances (L) between any two phases i, j of stator and/or rotor can be computed as: Li, j = L\u25e6 cos(\u03b8d(i, j)), (1) where \u03b8d(i, j) is the difference between ith and jth phases angles with respect to the reference frame and L\u25e6 is the maximum mutual inductance. As it is clear from Fig. 1, phase a is splitted into two phases as1 and as2. Hence, the inductance matrix can be computed as: L = [ Lss Lsr Lrs Lrr ] , (2) 978-1-4799-8900-3/15/$31.00 \u00a92015 IEEE 226 where Lss represents the mutual inductances between the stator windings and it is given by: Lss = Lsl (1\u2212\u00b5)2 0 0 0 0 \u00b5 0 0 0 0 0 0 0 0 0 0 + +L\u25e6 (1\u2212\u00b5)2 \u00b5(1\u2212\u00b5) \u22120.5(1\u2212\u00b5) \u22120.5(1\u2212\u00b5) \u00b5(1\u2212\u00b5) \u00b52 \u22120.5\u00b5 \u22120.5\u00b5 \u22120.5(1\u2212\u00b5) \u22120.5\u00b5 1 \u22120.5 \u22120.5(1\u2212\u00b5) \u22120.5\u00b5 \u22120.5 1 , (3) with \u00b5 being the percentage of the shorted turns to the total turns of the faulty phase" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0001301_saci49304.2020.9118823-Figure3-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0001301_saci49304.2020.9118823-Figure3-1.png", + "caption": "Fig. 3. Geometrical variables and parameters of UnTrans.", + "texts": [ + " 2 outlines the hardware diagram of UnTrans, and it is built up using the following elements: \u2022 two wheels, driven independently by one DC motor each using pulse width modulation (PWM) controllers, with incremental encoders measuring their rotation angles, \u2022 a mass located on the rotation axis, \u2022 the Inertial Measurement Unit (IMU) containing triple axis accelerometer and gyroscope sensors to measure the coordinates of UnTrans, \u2022 two 12 V, 2.3 Ah batteries, \u2022 an interface board serving as intermediary between the Zynq board and the rest of the system \u2022 and an RT-DAC/Zynq 7000 board containing a wireless card for using Wireless local area network (LAN), several general-purposed input/output (GPIOs) and an SD card slot to store the operating system. The board runs the real-time application and communicates with the external computer [5]. Fig. 3 displays the schematic diagram of UnTrans with the generalized coordinates: \u03b8 \u2013 the translational position of the wheel axis, \u03c8 \u2013 the tilt of the vehicle as difference to the vertical position and \u03c6 \u2013 the rotation around the vertical axis. Table I lists the main parameters of UnTrans system shown in Fig. 3. TABLE I. IMPORTANT PARAMETERS OF THE MODEL Symbol Units Values Description R m 0.075 radius of one wheel W m 0.4 distance between wheels (breadth) L m 0.102 height of the mass m kg 0.32 weight of one wheel M kg 5.41 weight of the body JW kgm2 0.0013 moment of inertia of one wheel J\u03c8 kgm2 0.104 moment of inertia in tilt axis 000212 Authorized licensed use limited to: University of Exeter. Downloaded on June 17,2020 at 13:24:44 UTC from IEEE Xplore. Restrictions apply. Symbol Units Values Description J\u03c6 kgm2 0" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0002273_j.fusengdes.2020.112102-Figure3-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0002273_j.fusengdes.2020.112102-Figure3-1.png", + "caption": "Fig. 3. Simplified environment for collision checking.", + "texts": [ + " Therefore, to reduce the complexity of collision checking without compressing the free space too much, we simplify VV as the space enclosed by two cylindroids and two planes. To ensure the enclosed space is totally inside VV, the radii of the inner and cylindroids and the heights of two planes should be chosen carefully. One reasonable set of these parameters are shown in Table 3. Since the inner space of CASK is a cuboid, we do not make any simplifications. The overall system used for collision checking is shown in Fig. 3. To make EAMA and VV both free of collision, the minimum distance dmin between EAMA and VV should be greater than a predefined safety threshold dsafe. And considering that all links are cylinders with the same radius r, if we modify the safety threshold as dsafe + r, then the original collision checking problem can be converted to the minimum distance between segments and VV should be greater than the modified threshold. And the two end points of segment i are the origins of consecutive joint frames i and i+ 1, where frame 9 is the end effector frame" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0002361_ffe.13405-Figure7-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0002361_ffe.13405-Figure7-1.png", + "caption": "FIGURE 7 Fractured outer ring having Type B defect: (A) schematics of fracture surface, (B) picture of the fracture surface; (C) picture of the cross section", + "texts": [ + "0 mm deep defect and was tested with a TABLE 1 Crack-depth values and minimum mesh-sizes (MMS) resulting from the error-verification models and the resultant KI errors Crack-depth (CD) Minimum mesh-size (MMS) MMS/CD KI error mm mm % % 6 0.024 0.4 0.46 6 0.10 1.7 0.86 6 0.48 8.0 2.4 10 0.10 1.0 0.67 14 0.10 0.7 1.1 0.2 mm fitting-gap. It is noted that, in the two tests denoted by asterisks in Table 3, specimens were classified as broken since crack lengths reached 0.6\u20131.0 mm as measured from the notch root after the fatigue tests. Figure 7 displays an example of the fracture surface and cross section of a broken specimen with a Type B defect (h = 1.0 mm, fitting-gap = 0.5 mm). Figure 7C features a crack that initiated near the notch-bottom, later propagating obliquely. Similarly, in specimens with Type B defects, cracks originated in the vicinity of the notch-bottom, rather than specifically at the notchbottom itself, propagating obliquely, with a random scattering in the crack-path. On the other hand, in specimens with Type A defects, cracks started at defect-tips, later growing in a roughly radial direction. To investigate the mechanistic reason for seemingly peculiar crack-path for Type B defect, an FEM stress analysis was performed using a two-dimensional, planestrain model, with the notch subjected to rolling-contact loading", + " The variation of normal stress along the notchroot surface was calculated at assorted roller positions. Figure 8 reproduces the distribution of maximum stress, minimum stress, and stress range in a circumferential direction along the notch-root surface; \u03c3max, \u03c3min and \u0394\u03c3, where the position is represented by the angle, \u03b8 (cf. illustrative definition in Figure 8). The difference between \u03c3max and \u03c3min represents the stress range, \u0394\u03c3, a dominant parameter in fatigue crack-initiation. The \u0394\u03c3 slightly varied with \u03b8, peaking at \u03b8 \u2248 \u00b140 . However, considering the broken bearing presented in Figure 7, the crack-initiation site was at \u03b8 \u2248 75 , where the \u0394\u03c3 value was 11% lower than the maximum. It is difficult to accurately predict the crack-initiation site, since crackinitiation can also be dominated by the stress ratio that varies with \u03b8. Furthermore, the crack-initiation site can be influenced by a microstructural irregularity near the notch-bottom. Consequently, as previously showcased, under gentle distribution of stress range and stress ratio in the notch-bottom region, the crack-initiation site was not always exactly at notch-bottom, leading to scatter of the crack-growth direction early in the fracture process" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0001568_aim43001.2020.9158880-Figure2-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0001568_aim43001.2020.9158880-Figure2-1.png", + "caption": "Fig. 2 Mecanum wheel platform diagram", + "texts": [ + " EKF is an incremental process that considers only the relationship between t and t-1 estimating the state variable, and observation measurement is determined by the previous moment. The main content of this section is the wheel odometer and the vision odometer methods that estimate the robot pose using an asynchronous data fusion formed by a wheel odometer and a visual odometer in EKF. Wheel odometer refers to a model that calculates the current robot\u2019s movement velocity and position by the encoder of the mobile robot. The mobile robot in our system is a Mecanum wheel structure (Fig. 2). 1991 Authorized licensed use limited to: UNIVERSITY OF ROCHESTER. Downloaded on August 30,2020 at 06:36:24 UTC from IEEE Xplore. Restrictions apply. Let the center of mobile be O, velocity and angular velocity in the world coordinate system are xv , yv and \u03c9 , the center of the n-th wheel be nO , the rotation velocity is n\u03c9 , and the velocity of n-th wheel center in the world coordinate system is nxv and nyv , the velocity of the contact point between the wheel and the ground is gnv , and the direction of travel of the wheel angle is \u03b1 " + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0003943_icra.2014.6907115-Figure6-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0003943_icra.2014.6907115-Figure6-1.png", + "caption": "Fig. 6. Simulation results when the robot tries to walk by leaning forward until it enters an unstable state, then recovering.", + "texts": [ + " The series of images in Figure 4 show the progression of the simulation at the transitions between the various states in the controller\u2019s state machine. The second simulation included a desired COM velocity, to dictate how fast and in what direction the robot should attempt to walk. In this simulation, the robot is statically stable until it reaches the Swing state, as the Lift state is augmented to move the swinging foot in the direction of desired motion before switching to tracking the FPE in the Swing state. A series of images, in Figure 6, shows the simulation at the transition points between states for one step of the walking cycle. In the final set of simulations, the robot was subjected to external disturbances from various angles while dynamically walking. For these simulations, the disturbance lasted for 0.1 seconds and acted on the center of the torso, as in the standing simulation. The external force was applied during the Swing state, to demonstrate the robot is capable of responding to disturbances while in the dynamic portion of its walk cycle" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0003350_9781119016854.ch38-Figure1-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0003350_9781119016854.ch38-Figure1-1.png", + "caption": "Figure 1. Basic mechanism of UNSM treatment.", + "texts": [], + "surrounding_texts": [ + "Ultrasonic nanocrystalline surface modification (UNSM) treatment UNSM is a method of metal improvement that utilizes ultrasonic energy. This technology has aroused considerable interest due to its economical effectiveness, possible fine adjustment of effects upon the workpiece and because it is a safe, simple and effective method for application in production of machine components and machinery of various purposes. The ultrasonic device consists of piezo-transducer (vibrator), booster (amplifier and transmitter of vibration), and horn. At the end of horn, there is a ball that transmits mechanical vibration to the material. The booster, horn and ball gripper are made of Ti alloy with good wave propagation properties and fatigue resistance. The ball is made of tungsten carbide (WC) with good wear resistance and hardness. In this UNSM technology process, not only the static load, but also the dynamic load is exerted to a workpiece surface. A generator and piezoelectric transducer which are shown in Fig. 2 emit ultrasonic waves at 20 kHz. The waves are amplified when they travel through an acoustic booster. The dimension of the vibrating part which contact surface, allows vibration amplitudes of 10 to 100 p,m to be attained. This UNSM technology provides a uniform treatment on the treated surface. The principle of UNSM is based on the instrumental conversion of harmonic oscillations of an acoustically tuned body into resonant impulses of ultrasonic high-frequency. The acoustically tuned body is brought to resonance by energizing an ultrasonic transducer. The energy generated from these high-frequency impulses with the total force a workpiece surface from 20,000 to 40,000 times per second and from 1,000 to 4,000 shots per square millimeter. The UNSM treatment set-up and procedure have been described in details in our previous study [6]. The important and attractive phenomena of UNSM treatment from other surface treatments are that this treatment have led to the development of a homogenous and uniformly distributed microstructure as well as smooth treating process due to setting on an NC machine. Specimen preparation In this study, disk and rotary bending specimens made of annealed Inconel 718 alloy were prepared for tribological and fatigue test484respectively. Table I shows the chemical w8e8re8 8 p8re8p8a8re8d8 8 f8o8r 8 tr8ib8o8lo8g8ic8a8l 8 a8nd8 8 f8a8tig8u8e8 8 tests. 8 R8o8ta8ry8 8 b8e8n8d8in8g8 8 f8a8tig8u8e8 8 s8p8e8c8im8e8n8s8 8 w8e8re8 tre8a8te8d8 8 b8y8 8 p8r8e8c8ip8ita8tio8n8 8 h8e8a8t 8 tre8a8tm8e8n8t 8 a8t 8 720 8 0C8/8h8rs8 8 a8nd8 8 f8u8r8n8a8c8e8 8 co8ol 8 to8 8 620 8 0C 8 at 38 8 0C8/hou8r. 8 In 8th8is8 8 study, 8 th8e8 8U8N8S8M8 8te8c8h8n8iq8u8e8 8w8a8s8 8 a8p8p8lie8d8 8to8 8 th8e8 8 s8p8e8c8im8e8n8s8 8u8n8d8e8r8 8 strictly c8o8n8tro8lle8d8 8p8a8ra8m8e8te8rs8 8a8t 8d8iff8e8r8e8n8t 8a8m8p8litu8d8e8s8 8as 8sh8o8w8n8 8in8 8T8a8b8le8 8II. T8a8b8le8 8I. 8C8h8e8m8ic8a8l 8c8o8m8p8o8s8itio8n8 8o8f8 In8c8o8n8e8l 8718 8allo8y8 8(in8 w8t.%8). Element8 Ni8 C8r8 Nb + Ta8 Mo8 Al8 C8 Si8 Ti8 Fe W8t.%8 580-558 17-218 48.7858-58.58 28.8-38.38 08.28-08.8 0.08 0.3858 08.6858-18.1858 B8a8l Il I$ '> lq. In Fig. 7(b), the minimum saliency point of the q axis was simulated, i.e. id=0 and iq = i\u2032qT0, being i\u2032qT0 the current value on q axis corresponding to the maximum lq. As can be seen, in this condition the structural ribs are not saturated, and so the magnetic flux can flow almost linearly either in d and q direction. As a result, the reluctance along d and q axes is nearly the same and so ld \u2248 lq. In Fig. 7(c) the negative iq was further increased (iq=-7 A). It can be noted that the ribs are saturated again, so ld >> lq, but differently from Fig. 7(a) the saturation is due to the current, instead of the PM, so the magnetic induction in the ribs has opposite sign. VI. DETERMINATION OF iqT0 BASED ON MINIMUM SALIENCY The basic assumption behind the method proposed here is that the current iqT0 is nearly equal to the current correspond- ing to maximum lq, called i\u2032qT0. In other words, the curve \u03bbq(iq) presents its maximum slope approximately at \u03bbq(iqT0). It must be noted that this condition also corresponds to the minimum local saliency along the q axis" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0002146_icem49940.2020.9270956-Figure2-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0002146_icem49940.2020.9270956-Figure2-1.png", + "caption": "Fig. 2. 2D cross section of one pole of the benchmark machine.", + "texts": [ + "15 mm, but still rotates around its own center. As shown in Fig. 1 (c), the rotor is similarly shifted from the center point to the stator for 0.15 mm and rotates around the stator center. Both two cases are having an uneven airgap, that there is a smaller airgap on the positive x-axis direction. The healthy rotor is also considered in the case study as a comparison, which is shown in Fig. 1 (a) [7]. A 4 pole/36 slots PMaSynRel machine with three layers of flux barriers per pole is adopted here as shown in Fig. 2, and its main dimensions are list in Tab. I [8]. TABLE I MAIN DIMENSIONS OF STUDIED MACHINE Parameters Value Parameters Value External stator diameter 200 mm Number of slots 36 Internal stator diameter 125 mm Number of poles 4 Airgap 0.35 mm Rated current 5.29 A Stack length 40 mm Rated torque 12.5 Nm Impact on Vibration of Eccentric Permanent Magnet Assisted Synchronous Reluctance Machine Jiaqi Li1, Hanafy Mahmoud2, Michele Degano1,3, Anuvav Bardalai1, Xiaochen Zhang3, Chris Gerada1,3 P Authorized licensed use limited to: Carleton University. Downloaded on May 31,2021 at 17:56:31 UTC from IEEE Xplore. Restrictions apply. Fig. 2 shows a PMaSynRel machine in healthy condition, which means it has a concentric rotor and the machine is rotating around the center. Then, the rotor is moved 0.15 mm along the d- axis to simulate the static and dynamic eccentric conditions. In addition, the three cases shown in Fig. 1 are analyzed and compared in the following sections for their electromagnetic forces, airgap flux densities, and finally their vibrational level. For the radial flux permanent magnet machines, the vibration is mainly caused by radial force" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0000256_icems.2019.8921568-Figure2-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0000256_icems.2019.8921568-Figure2-1.png", + "caption": "Fig. 2. The specific composition of the motor", + "texts": [ + " Then the characteristics of demagnetization of AlNiCo are introduced, and the simplified mathematical model of AlNiCo is established to explain the principle of field regulation. In order to obtain the largest magnetic field linear control range, the combination scheme of the two permanent magnet thicknesses is analyzed by finite element method. Finally, the rule of influence of permanent magnet thickness on permanent magnet working point, magnetic field regulation ability, gap flux density linear control range, no-load back EMF and rated load torque is summarized. II. STRUCTURE OF THE FVPSCPM The structure of FVPSCPM is shown in Fig. 1 and Fig. 2. Memory permanent magnet material AlNiCo and high performance permanent magnet material NdFeB are used in This work was supported by Major Project of Application Technology Research and Development Plan of Heilongjiang Province of China under Grant GA15A401. 978-1-7281-3398-0/19/$31.00 \u00a92019 IEEE series as excitation part to ensure that the motor has good regulation ability of gap flux density as well as high energy density. The rotor adopts claw pole shape. Since the presence of magnetic regulation coil, in order to get rid of brush and slip ring, a partitioned stator structure is adopted, and the excitation structure is separated from claw pole rotor as partitioned inner stator" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0002902_978-94-007-6046-2_43-Figure12-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0002902_978-94-007-6046-2_43-Figure12-1.png", + "caption": "Fig. 12 Mapping simple gait models to legged robots. (a) Virtual leg control with physical legs producing imagined action of a virtual one-legged hopper [45]. (b) Explicit mapping of the SLIP into more complex leg model in simulation [49]. (c) Generalization to virtual model control [44]", + "texts": [ + " They have in common that the template behavior is mapped onto the degrees of freedom of the more complex legged robot with techniques similar to operational space control. Simplified approaches consider the legged robot as a rigid body with massless legs and map an imagined wrench F des body acting on this body into the joint torques D J T F des body; (11) using the Jacobian J . Example approaches include virtual leg control and, more generally, virtual model control. In virtual leg control, the net forces and torques generated by two or more physical legs on the center body of a robot are made equal to the wrench applied by a virtual leg (Fig. 12a). Raibert and colleagues [45] used this approach to map the control developed for their one-legged hopper onto quadrupedal machines. The control algorithms reviewed in Sect. 3.4 operated on a virtual one-legged hopper, and the resulting net wrench applied by the hopper\u2019s leg was distributed to the physical legs of the quadrupedal machine under the constraint that the axial forces of the legs on the ground in synchrony are equal. Saranli and colleagues [49] applied the same quasi-static torque mapping to specifically embed the SLIP model acting between the toe and the hip of a simulated segmented leg (Fig. 12b). Later, the concept of quasi-static mapping has been generalized to arbitrary model systems. Using virtual spring and damper models acting on the center body, Pratt and colleagues [44] devised an intuitive walking control for their bipedal robot \u201cspring turkey\u201d (Fig. 12c). Recent approaches to embedding the SLIP take the full rigid body dynamics of legged robots into account. Sreenath and colleagues [62] applied control techniques from hybrid zero dynamics to design a feedback controller for the robot dynamics model: M Rq C h D S T C J T F ext; (12) that embeds a virtual spring behavior in the stance leg. They demonstrated with this controller running at 3 m/s of the MABEL robot constrained to a boom (Fig. 5a). More often, the SLIP is embedded as the target for the translational dynamics of the CoM of the entire robot using direct operational space control methods" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0001345_j.promfg.2020.05.058-Figure1-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0001345_j.promfg.2020.05.058-Figure1-1.png", + "caption": "Fig. 1. Weld spatter test sample geometry", + "texts": [ + "85 the weld arc length reduces and in doing so reduces the welding voltage. Similarly, when the trim is increased from 1 to 1.15, the weld arc length increases and therefore increases welding voltage. Accordingly, in Table 4 two numbers in the column for Voltage/Trim are for Pulse welding process and just one number for the CV process. The test samples were made from \u00bc\u201d thick ASTM A572 Grade 50 steel. Two plates of dimensions 10\u201dx3\u201dx1/4\u201d and 10\u201dx2\u201dx1/4\u201d were tag welded at the ends to form a \u201cT\u201d junction as shown in the schematic in Figure 1. A total of 80 such test samples were made for weld spatter testing to cover 64 test pieces for DOE, and the rest for use in preliminary trial experiments. Iqbal Shareef et al. / Procedia Manufacturing 48 (2020) 358\u2013371 3614 Iqbal Shareef and Christopher Martin/ Procedia Manufacturing 00 (2019) 000\u2013000 Initially three spatter collection trays were designed and fabricated from a 1/16\u201d sheet metal. A \u00bc\u201d diameter steel rod bent in the shape of a shallow U, and welded at the ends of the trays, served as handles for easy placement and handling of spatter collected in the trays", + " Various methods were employed for collection and analysis of weld spatter data, to express the loss due to spatter in terms of cost. Several materials and tools were used, first to establish a baseline, and subsequently to test and analyse the responses Even before the start of experiments, the current state of weld spatter generation had to be quantified and a baseline established. The process of determining the baseline included quantifying spatter in production. Consequently, a highvolume robot was identified running normal settings and three trays shown in Figure 1 were placed at the base of the welding fixture. These remained for an entire shift after which they were removed, and the trays were emptied, and spatter accumulated in the trays was collected. Subsequently the spatter was transferred into a Ziploc bag and weighed. The costs associated with weld spatter were also explored and summarized. A summary of the standard settings are listed in Table 5 below, and the weight in pounds of spatter collected on each day from the collection trays is shown in Table 6", + " \ud835\udc4a\ud835\udc4a\ud835\udc4a\ud835\udc4a\ud835\udc4a\ud835\udc4a\ud835\udc4a\ud835\udc4a\u210e\ud835\udc61\ud835\udc61 \ud835\udc5c\ud835\udc5c\ud835\udc5c\ud835\udc5c \ud835\udc64\ud835\udc64\ud835\udc4a\ud835\udc4a\ud835\udc56\ud835\udc56\ud835\udc4a\ud835\udc4a \ud835\udc49\ud835\udc49\ud835\udc5c\ud835\udc5c\ud835\udc59\ud835\udc59\ud835\udc61\ud835\udc61 \ud835\udc4a\ud835\udc4a\"2 = \ud835\udc4a\ud835\udc4a\" \u2212 \ud835\udc4a\ud835\udc4a,- (4) \ud835\udc43\ud835\udc43\ud835\udc4a\ud835\udc4a\ud835\udc56\ud835\udc56\ud835\udc43\ud835\udc43\ud835\udc4a\ud835\udc4a\ud835\udc43\ud835\udc43\ud835\udc61\ud835\udc61 \ud835\udc49\ud835\udc49\ud835\udc5c\ud835\udc5c\ud835\udc59\ud835\udc59\ud835\udc59\ud835\udc59 \ud835\udc5c\ud835\udc5c\ud835\udc5c\ud835\udc5c \ud835\udc64\ud835\udc64\ud835\udc4a\ud835\udc4a\ud835\udc49\ud835\udc49\ud835\udc4a\ud835\udc4a \ud835\udc64\ud835\udc64\ud835\udc4a\ud835\udc4a\ud835\udc56\ud835\udc56\ud835\udc4a\ud835\udc4a \ud835\udc4a\ud835\udc4a\"2% = STFU TVW TF X100 (5) Number of weld spatter particles (Nsp): A GoPro camera and four stationary video cameras were installed in the robotic cell shown in Figure 5. These cameras were used to monitor welding robots and manual welding processes. During each test videos were recorded and transferred to a video editing software for analysis. Each video from the welding cell was analyzed and screenshots for each part were taken. While welding was performed on both sides of the 10 inch long inverted-T section test specimen shown in Figure 1, data from only one side of each test piece was processed due to the robotic welding arm covering some of the spatter on the other side, leaving only one side of each test piece to be analyzed. Using the weld on the back side of each test piece to capture spatter optically, screenshots were taken at regular intervals from the videos. Those images from each piece were ran through a particle counting software called ImageJ. The ImageJ is an open source image processing program that was used to count the spatter particles coming off of the weld", + " Then taking the average of each of the four repeat experiments, an overall average (Nsp) of each test listed in Table 4 was determined. Intensity of sound during welding (Is): The GoPro camera installed in the robotic welding cell recorded both audio and video information during each test. The audio data was imported from the GoPro camera into the open source audio software called Audacity. In Audacity the stereo data was converted to mono at a sample rate of 48,000 Hz. From each weld five samples each of two second duration were exported into text files. Each test specimen shown in Figure 1 had two welds. Each two second data packet contained 96,000 data points. The text file contained the normalized decibel waveform data, and a decibel scale range from -1 to 1. A value of zero is quietest while -1 or 1 indicates the loudest relative sounds. The exported text file is inserted into Excel where the relative minimum and maximum values were determined. The maximum values are averaged. The larger the average of maximum values, the louder the weld is. A louder weld is an indication of spatter being produced" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0000891_cdc40024.2019.9029431-Figure3-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0000891_cdc40024.2019.9029431-Figure3-1.png", + "caption": "Fig. 3. Trajectories of leader and agents with time as the vertical axis. The agents converge to the desired relative positions at approximately t = 70 s.", + "texts": [ + "4, which implies that the speed dynamics are a low-pass filter with unity gain at dc. The time constant 0.4 is estimated from the closedloop step response of the UAV with the middle-loop speed controller. The desired positions are d1 = [\u221210 10 0]T m, d2 = [\u221210 \u221210 0]T m. Let \u03b1i = 0.35, \u03b212 = \u03b221 = 0.1, \u03b3i = 1, \u03b7i = 0.001, and let Ts = 0.1 s, and \u03c6g(t) \u2261 0. Example 3. The leader moves to a location away from the launch site and loiters in a 120m radius circle. The agents start from circling the launch site and enter formation. Figure 3 shows the trajectories of the leader and agents in the e1\u2013e2 plane plotted with time as the vertical axis, and Figure 4 shows Rgdi and qi \u2212 qg. The relative positions converge to the desired values by approximately t = 70 s. After t = 70 s, the RMS magnitude \u2016\u03b6i\u2016 of the position errors are 4.36 m and 4.49 m for agent 1 and 2, respectively. 4 Example 4. The leader moves in a square pattern. The agents start from circling the launch site and enter formation. Figure 5 shows the trajectories of the leader and agents in the e1\u2013e2 plane plotted with time as the vertical axis, and Figure 6 shows Rgdi and qi\u2212qg Note that the desired position is approximately maintained through the abrupt turns in the 8234 Authorized licensed use limited to: Carleton University" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0002318_cacs50047.2020.9289703-Figure1-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0002318_cacs50047.2020.9289703-Figure1-1.png", + "caption": "Fig. 1. Kinematics of mobile Robot", + "texts": [ + " Section II defines the model derivation and the fault-motion analysis of Mecanum wheel omni-directional vehicle. Section III introduces the fault detection of signal-based method and model-based method. Section IV summarizes the setup and the result of the experiment. Section V draws conclusions. II. MODELS A. Kinematics In this research, there are four Mecanum wheels in the omnidirectional Vehicle, and each Mecanum wheel is composed by several free small wheels which are in 45 degree. The configuration of Mecanum wheel omni-directional vehicle is shown as Fig. 1. Where \u03a3! is the Global coordinate, and \u03a3\" is the coordinate system of the centroid of the omni-directional Authorized licensed use limited to: London School of Economics & Political Science. Downloaded on May 17,2021 at 13:35:26 UTC from IEEE Xplore. Restrictions apply. vehicle, and \u03a3#$ (\ud835\udc56 = 1\u30012\u30013\u30014) is the coordinate system of the center of each wheel, and \u03a3\" and \u03a3#$ are the Local coordinate. The angle between (\ud835\udc4b% , \ud835\udc4c% , \ud835\udf03%) and \u03a3\"(\ud835\udc4b, \ud835\udc4c, \ud835\udf03) is \ud835\udf03& , and these two coordinate is converted by applying rotation matrix \ud835\udc3d%\" and inverse matrix \ud835\udc3d\"%" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0003239_icmtma.2015.150-Figure5-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0003239_icmtma.2015.150-Figure5-1.png", + "caption": "Figure 5. UP6 robot model is added constraint vice", + "texts": [ + " Through MicroStation software, the various components of the UP6 robot are output into IGES format. After running the ADAMS software, the various components of the robot in IGES format were imported correctly by choosing Import item under the File menu. UP6 robot only needs to add two kinds of motion pair which are the fixed side (Fixed) and rotation (Revolute). The fixed vice is applied between the base and the earth and the other five joints were applied rotation vice. The completion is shown in Figure 5. You can right-click the Information button in the lower-right corner of the ADAMS/View window to check whether the constraints between various components is correct or not. The data extracted function of AutoCAD can generate EXCEL file by dividing the number of fixed points of X, Y, Z coordinate values, for example, the operation trajectory which is drawn by AutoCAD in figure 6. Running ADAMS, selecting file type for the Test Data, clicking on the Create Splines, finding the imported text document, after that, you can see three curves that were generated automatically" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0003929_scis-isis.2014.7044649-Figure3-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0003929_scis-isis.2014.7044649-Figure3-1.png", + "caption": "Figure 3. Differential drive robot model", + "texts": [ + " The MCU controls the entire system of the riding robot, including functions such as analyzing a communication packet and controlling a motor. In addition, a Bluetooth wireless communication module is equipped for communication with the smartphone and an inertial matrix unit is built into a compass sensor for measuring the heading orientation of the riding robot. 978-1-4799-5955-6/14/$31.00 \u00a92014 IEEE 18 In this study, the implemented riding robot is a differential drive robot model, as shown in Fig. 3. The differential drive robot model is described by three vectors. The three vectors are the current position ( ),c cx y and orientation angle c\u03b8 with respect to a reference frame. Assuming that the wheels of a robot are against the ground with no slippage, the equation for the position and orientation angle is given by sin cos 0c c c cx y\u03b8 \u03b8\u0394 \u2212 \u0394 = (1) where P is the center between both sides of a wheel and c\u03b8 is the orientation angle of the robot. Because of natural constraints, integration is impossible" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0003714_esda2014-20232-Figure11-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0003714_esda2014-20232-Figure11-1.png", + "caption": "Figure 11: Equivalent von-Mises stress on shell \u2013 static analysis (Pa)", + "texts": [], + "surrounding_texts": [ + "The water brake specification is presented in table 1. In the first step, an initial estimation of loading about the rotor is made. Considering the mechanism of power absorption which is based on passing a water flow with an approximate pressure of 3.5 bar through the mechanical structure which produces friction between moving and stationary part, the initial loading is estimated. By measuring the amount of the torque, the turbine torque can be calculated precisely. In this regard, a torque load is considered on the load cell arm. Bearings components should be chosen carefully to handle the loading. Consequently, angular contact bearings are used for their ability to deal with high rotational speed and their widely use in precise applications. All connections are bolted and lock nuts are used to create preload on the bearings. Twelve anchor bolts are used tohold the \u201cCabAssy\u201don the housing and maintain bearings preload on their position. To transmit torque from the engine to the dynamometer, three options can be considered including: flexible coupling, diaphragm coupling and floating quill shaft (splined). Here, floating quill shaft is used. Anti-seize lubrication is applied to the splines. Water brake shaft splines are carefully designed and machined to be mounted on the shaft and engine spline. To prevent premature wear of the splines, the shaft is hardened. Of course, if the mechanical analysis is performed on all parts, more accurate results will be achieved. However, the analysis has been carried out on some important parts according to the need and importance which is more time and cost effective. The effect of the other parts are only considered in boundary conditions of the important components during the analysis. In some cases, transient analysis is needed to assess critical situations. The static load analysis in is done in 1 second period of time which is equivalent to a time step. Mode shape analysis is also performed for the structural components for predicting critical damage locations." + ] + }, + { + "image_filename": "designv11_71_0001057_0021998320920920-Figure10-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0001057_0021998320920920-Figure10-1.png", + "caption": "Figure 10. Contours of maximum Hashin failure index, Hmax, for binder in unit-cell 4 with warps and wefts shown for context.", + "texts": [ + " The severity of 0zz appears inversely proportional to the distance between a binder and the weft, as expected. It should be also noted that the binders seem unaffected by the cross-sectional shape of the neighboring wefts. For the warps and wefts, multiple tows lie within a single unit-cell, which can provide some insight into how much variation occurs between tows of the same type for the textile model used in this study. However, since there is only one wavelength of a single binder within a unit-cell, Figure 10 shows contours of the maximum index of Hashin\u2019s criteria for the binder in unitcell 4, which can be compared to Figure 5 that shows the same results for unit-cell 2. The locations of high Hmax matched well between the two unit-cells. However, the highest concentrations were more severe within the volume of the binder in unit-cell 2 compared to unit-cell 4, and the cross-sectionally averaged Hmax in the regions of concentrations in the binders differed by up to 16% across all four unit-cells shown in Figure 2(a)" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0000327_j.mechmachtheory.2019.103729-Figure15-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0000327_j.mechmachtheory.2019.103729-Figure15-1.png", + "caption": "Fig. 15. Front-end loop L(2) mechanism.", + "texts": [ + " L(7) is a newly added back-end loop, and L(8) is an intermediate loop without 0-rank and highest rank components. The mechanism of the front-end loop L(1) and the orientation of the kinematic pair have not changed relative to the aligned state, so the DOF of the 1-rank component 2 \u2032 and 4 \u2032 in the L(1) is 1, and 2-rank component 3 \u2032 has a DOF of 3. Next, the DOFs of the components in L(2) and L(3) where the orientation changes have been analyzed. Taking the loop L(2) as an example, the mechanism diagram is shown in Fig. 15 . The two axes of rotation are collinear with the Y and Z axes, respectively. The planar pair E 2 \u2032 2 is perpendicular to the Z axis, and the planar pair E 92 is parallel with the Z axis, and the angle with the XOZ plane is \u03b8 (0 \u25e6 < \u03b8 \u2264 30 \u25e6). In this loop, the highest rank component 2 is used as the moving platform. And the mechanism can be regarded as a 2-RE parallel mechanism. The motion-screw system of branch 9 \u2032 -2 \u2032 -2 can be expressed as S/ 1 1 = ( 1 0 0 ; 0 0 0 ) S/ 1 2 = ( 0 0 1 ; 0 0 0 ) S/ 1 3 = ( 0 0 0 ; 1 0 0 ) S/ 1 4 = ( 0 0 0 ; 0 1 0 ) (26) These screws are independent of each other, and their constraint-reciprocal-screw is S/ r 11 = ( 0 0 1 ; 0 0 0 ) S/ r 12 = ( 0 0 0 ; 0 1 0 ) (27) The motion-screw system of another branch 9 \u2032 -9\u20132 is expressed as S/ 2 1 = ( 0 0 1 ; 0 0 0 ) S/ 2 2 = ( cos \u03b8 sin \u03b8 0 ; 0 0 0 ) S/ 2 3 = ( 0 0 0 ; \u2212 sin \u03b8 cos \u03b8 0 ) S/ 2 4 = ( 0 0 0 ; 0 0 1 ) (28) Its constraint-reciprocal-screw is S/ r 21 = ( cos \u03b8 sin \u03b8 0 ; 0 0 0 ) S/ r 22 = ( 0 0 0 ; \u2212 sin \u03b8 cos \u03b8 0 ) (29) According to two constraint-screw systems: four reciprocal-screw linear independent, no common constraint, d = 6 , no redundant constraint, v = 0 " + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0003208_9781118751992.ch4-Figure4.9-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0003208_9781118751992.ch4-Figure4.9-1.png", + "caption": "Figure 4.9 Schematic diagram of the bookshelf cell structure of the surface-stabilized ferroelectric liquid", + "texts": [ + " The larger the tilt angle, the more biased the rotation around the long molecular axis becomes, and therefore the larger the spontaneous polarization. In electro-optical devices, it is usually required that the liquid crystal director is unidirectionally oriented. In the smectic-C* phase, however, the liquid crystal director twists from layer to layer. This problem is overcome by Clark and Lagerwall in their invention of the surface-stabilized ferroelectric liquid crystal (SSFLC) device [16], shown in Figure 4.9. The liquid crystal is sandwiched between two parallel substrates with the cell gap, h, thinner than the helical pitch, P, of the liquid crystal. The inner surface of the substrates is coated with alignment layers which promote parallel (to the substrate) anchoring of the liquid crystal on the surface of the substrate. The smectic layers are perpendicular to the substrate of the cell, while the helical axis is parallel to the substrate. Now the helical twist is suppressed and unwound by the anchoring" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0001824_j.procir.2020.09.036-Figure1-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0001824_j.procir.2020.09.036-Figure1-1.png", + "caption": "Fig. 1. Sectional view along the axis of the simulation model \"1.5\", A: Shift in x-direction, B: symmetry plane, C: Area with adjusted element size", + "texts": [ + " The known anisotropy of PBF-LB/M components was not considered in order to keep model complexity and computational time low. In this case the results of the simulation are dependent on the building direction of the examined tensile rods and can therefore not be transferred to other sample orientations. A CAD model of the faulty and the reference tensile specimens was created for each simulation carried out. Tetrahedron geometries in size of 3 mm were used to mesh the model. Starting from the symmetry plane (B), where the defects are located, mesh density was enhanced to an element size of 0,15 mm with C = 3.5 mm (Fig. 1). In order to optimize the calculation time and due to symmetry, only half of the symmetrical tension rods were simulated. As in the tensile testing, a constant shift (A = 3,36 mm) was applied on the thread of the tensile rod in X-direction (see Fig. 1). The shift was calculated using the elongation at break which was determined in the tensile testing. Thus, the applied displacement corresponds to the real movement until failure in the tensile test. For the simulation of the equivalent stress and the main stress vector the sparse direct solver (SDS) was used. Fig. 2 shows a section of a prepared test body with a defect height z = 0.2 mm compared to the underlying CAD model. The lower defects, up to a width of x = 0.06 mm, are remelted by the further layer structure and are no longer detectable in the component structure (not shown in Fig" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0001808_tie.2020.3022502-Figure1-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0001808_tie.2020.3022502-Figure1-1.png", + "caption": "Fig. 1. The cross section of a 12/8-pole DSEM.", + "texts": [ + " The electromagnetic characteristic and basic operation principle of DSEM are presented in Section II. In Section III, the initial rotor sector estimation method is introduced. In Section IV, the novel sensorless startup strategy is presented. Section V provides the simulation and experimental evaluations of this method. Section VI concludes the contribution of this paper. II. BASIC OPERATION PRINCIPLE OF DSEM The DSEM has a salient-pole structure with armature and field windings on the stator and without any magnet or coils on the rotor, as shown in Fig.1. The unsaturated electromagnetic characteristic of DSEM is shown in Fig.2. The reluctance of each phase varies with the rotor position, and the back EMF is induced by the variation of the phase flux linkage. The flux linkage of each phase can be derived as , , , , ,x x x y xy z xz f xfi L i M i M i M x y z a b c = + + + = (1) where ix is phase current, if is the field current, Lx is phase self-inductance, Mxy is the mutual inductance between phase windings, Mxf is the mutual inductance between phase winding and field winding" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0001279_piicon49524.2020.9112890-Figure2-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0001279_piicon49524.2020.9112890-Figure2-1.png", + "caption": "Fig. 2 Healthy Condition of Induction motor", + "texts": [], + "surrounding_texts": [ + "A 3HP, 415V, 50 Hz, 1440rpm , 4 pole three phase Induction motor is designed using the basic equations of motor and it has been modelled in RMxprt in Ansys Maxwell. All the parameters such as stator diameter, length, number of stator slots and specifications, Type of winding, dimensions, Type of steel, Number of conductors/slot, No of Parallel branches, rotor length, diameter, number of slots are given from the calculated values. The performance characteristics is analysed for both healthy and inter-turn short circuit fault of three phase Induction motor. The inter-turn fault is modelled for the same motor by short circuiting 16% and 25% of the turns in one phase. The Transient analysis for healthy as well as faulty induction motor is done in Ansys Maxwell 2D. Table I Induction motor Specification" + ] + }, + { + "image_filename": "designv11_71_0002600_ieeeconf38699.2020.9389412-Figure3-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0002600_ieeeconf38699.2020.9389412-Figure3-1.png", + "caption": "Fig. 3. Relation between \u03b1k , \u03c7 and \u03c7d .", + "texts": [ + " (31) The control objective is to drive the cross-track error to zero by steering the vessel. The Lookahead-Based Steering Method is used [9], [15]\u2013[17]. In order to drive the cross track error to zero, a look-ahead distance \u0394 is defined along the path between s and the next waypoint pk+1. The vessel is then steered so that its velocity vector is parallel to the LOS vector (Fig. 2). The resulting velocity vector will have a component perpendicular to the path, driving the vessel towards the path until the LOS vector is parallel to the path and e\u2192 0. From Fig. 3 it can be seen that when the velocity vector v and LOS vector are aligned, so that \u03c7 = \u03c7d , the angles Authorized licensed use limited to: University of Prince Edward Island. Downloaded on May 29,2021 at 18:57:11 UTC from IEEE Xplore. Restrictions apply. (\u03b1k\u2212\u03c7) and \u03be will be the same. The change in angle required to achieve this is given by \u03c7\u2212\u03c7d =\u2212(\u03b1k\u2212\u03c7)+\u03be =\u2212(\u03b1k\u2212\u03c7)+ tan\u22121 ( e \u0394 ) . (32) Solving for \u03c7d , gives \u03c7d = \u03b1k\u2212\u03be = \u03b1k+ tan \u22121 ( \u2212 e \u0394 ) . (33) Define the slide-slip angle to be \u03b2 = sin\u22121 ( v U ) = sin\u22121 ( v\u221a u2+ v2 ) . (34) where U is the magnitude of the velocity vector v, which can be expressed in the body-fixed reference frame as U =\u221a u2+ v2 and in NED reference frame as U = \u221a x\u03072n+ y\u03072n For the purposes of control, it is simpler to define the control input in terms of a desired heading angle \u03c8d , instead of the desired course angle \u03c7d . From Fig. 3 it can be seen that \u03c7 = \u03c8+\u03b2 \u21d2 \u03c7d = \u03c8d+\u03b2 . (35) The desired course angle and desired heading can then be related using (33), as \u03c7d = \u03b1k+ tan \u22121 ( \u2212 e \u0394 ) = \u03c8d+\u03b2 , (36) so that \u03c8d = \u03b1k+ tan \u22121 ( \u2212 e \u0394 ) \u2212\u03b2 . (37) Let the heading error be defined as \u03c8\u0303 := \u03c8 \u2212\u03c8d . Taking the time derivative of the heading error gives \u02d9\u0303\u03c8 = \u03c8\u0307\u2212 \u03c8\u0307d = \u03c8\u0307\u2212 d dt [ \u03b1k+ tan \u22121 ( \u2212 e \u0394 ) \u2212\u03b2 ] . (38) Let tan\u03be := ( e \u0394 ) . (39) Then d dt tan\u03be = 1 cos2 \u03be \u03be\u0307 and d dt ( e \u0394 ) = e\u0307 \u0394 , so that \u03be\u0307 = cos2 \u03be e\u0307 \u0394 . (40) From Fig. 3 it can be seen that cos\u03be = \u0394/ \u221a e2+\u03942. Substituting this expression into (38) above yields d dt [ tan\u22121 ( \u2212 e \u0394 )] =\u2212\u03be\u0307 =\u2212 \u0394 (e2+\u03942) e\u0307 (41) Thus, the time derivative \u02d9\u0303\u03c8 is \u02d9\u0303\u03c8 = (r\u2212 rd)+ rd+ \u0394 (e2+\u03942) e\u0307+ \u03b2\u0307 , = r\u0303+ rd+ \u0394 (e2+\u03942) e\u0307+ \u03b2\u0307 , (42) where r\u0303 := r\u2212 rd is the yaw rate error and rd is the desired yaw rate. Thus, let rd :=\u2212 \u0394 (e2+\u03942) e\u0307\u2212 \u03b2\u0307 , (43) be a virtual control input such that the heading error dynamics become \u02d9\u0303\u03c8 = r\u0303. (44) Next, the definition of the yaw rate error r\u0303 := r\u2212 rd and (7)\u2013 (14) can be used to rewrite the equation of motion for the yaw rate as (Iz\u2212Nr\u0307) \u02d9\u0303r = \u2212(Iz\u2212Nr\u0307)r\u0307d+(Yv\u0307v+Yr\u0307r)u\u2212Xu\u0307uv +Nvv+Nrr+(Nv|v||v|+N|r|v|r|)v +(N|v|r|v|+Nr|r||r|)r+ \u03c4\u03c8 +d\u03c8 " + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0002182_j.sna.2020.112472-Figure1-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0002182_j.sna.2020.112472-Figure1-1.png", + "caption": "Fig. 1. Schematic of theoretical calcula", + "texts": [ + " Single sphere Considering the simplest case, a non-magnetic sphere rises in vertical tube filled with ferrofluids. The sphere passes through n induction coils wrapping around the outer wall of the tube, and errofluids in the tube are magnetized by the uniform magnetic eld. Thus, the magnitude of the induced electromotive force and ts relation with the radius of the sphere and the speed of rising can e calculated. Schematic of a theoretical calculation model of the induced elecromotive force is revealed in Fig. 1. The inner diameter of the ircular tube is R. Ignoring the wall thickness of the tube, a nonagnetic sphere (radius a) rises at a constant velocity (v) in a ertical circular tube, and the width along the z-axis of an induction oils is also R. The radius of the induction coils is assumed to be the ame as the vertical tube, and the number of coils is n. A uniform magnetic field H is imposed in the vertical direction. tant) and satisfies the theory of the equilibrium magnetization, .e., M= H. Ferrofluids remain stationary during the ascent process f the sphere, and the disturbance of the sphere to ferrofluids is F", + " This process is equivalent to the magnetic flux varies in the induction coils caused by the motion of a uniform magnetic sphere (radius a, volume V, magnetization M). The distribution of the magnetic field outside the magnetic sphere is equal to the magnetic field generated by the magnetic dipole located at the spherical core with a magnetic moment of m=V\u00d7M: B (r) = 0 4 3n (n \u00b7 m) \u2212 m r3 ( n= r r ) (2) where r is the distance from the spherical center to the point of calculation. To facilitate the calculation of the magnetic flux passing through the induction coils, a spherical surface \u2211 with a radius r is selected for integration, as shown in Fig. 1, which can be obtained: \u030a = \u02d9dS \u00b7 B = 0 4 r3 \u222b 0 0 [ 2 r2 sin ( ) d ] \u00b7 ( 2m cos ( )) = 2 0MR2a3( )3/2 (3) 3 R2 + z2 where 0 is the angle between the z-axis and the radius of the spherical surface. \u02da \u02db odel of induced electromotive force. Thus, the expression of induced electromotive force generated n the induction coils can be written as: = \u2212n d\u02da dt = \u2212n 2 0MR2a3zz\u0307( R2 + z2 )5/2 (z = z0 + vt, z\u0307 = dz dt = v) (4) .2. Gas-liquid two-phase flow If the gas-liquid two-phase flow is flowing in the tube, the olution of induced electromotive force also needs to know the evoution of interface with time" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0000671_ictai.2019.00036-Figure1-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0000671_ictai.2019.00036-Figure1-1.png", + "caption": "FIGURE 1. One stack-up of the piezoelectric film based capacitive touch screen panel.", + "texts": [ + " It is an approach to implement force sensing in capacitance TSPs by employing piezoelectric materials [4], [5] which generate charges on the surface when a force load is applied. The amount of charge produced is proportional to the strength of the force and the piezoelectric film thickness. Thus when the piezoelectric material is used as substrate of the touch sensors, the induced charge can be read by the touch sensors for force sensing. Four stack ups widely used in industry are investigated, and one of them is depicted in Fig. 1. A thin layer of the piezoelectric film (\u223c10\u00b5m) is underneath the touchscreen glass (\u223c0.5mm). The electrodes are much thinner compared to the piezoelectric film and cover glass so they are not shown in the figure. When a force is applied to the glass surface, the stress transmitted to the piezoelectric film layer will result in the induction of charges, which will be measured to interpret the force level. Among the four stack-ups, the highest responsivity and signal-to-noise ratio (SNR) are 0.42 V/N and 59" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0002964_ijcvr.2019.098008-Figure4-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0002964_ijcvr.2019.098008-Figure4-1.png", + "caption": "Figure 4 Schematic diagram showing the push recovery applied on the biped robot, (a) sagittal view (b) frontal view (c) instability criteria", + "texts": [ + " Moreover, if the external forces applied on the robot are small in magnitude, then a resisting torque can be applied at the ankles of the robot that can prevent the falling of the robot. If the forces are large in magnitude, the robot has to generate resisting torque by moving its joints. The amount of resisting torque that the robot can generate and the maximum force that it can withstand depend on the joint toque limit and physical rotation limits of the joints. If forces are further large, then the robot can take a step forward/backward to avoid fall. Hip and ankle joint correction strategy is implemented [ref. Figure 4(a)] on the robot when the robot is stationary and about to recover from an external disturbing force. The robot was made to apply maximum body torque irrespective of the force experienced. It is important to note that the servo motors used in the robot are position controlled motors and it is not possible to apply a torque directly at the joints. In order to accommodate this, the algorithm is designed to supply the necessary change in the joint angle at an appropriate time. The body has a moment of inertia of 0", + " It is assumed that the external disturbing force is applied from the back and no additional disturbances are applied on the robot. Under such scenario, the set of values that produce the highest resisting force has been identified. Highest resisting force implies a force that is sufficient to make the robot fall. In such case, the robot is provided with hip and ankle joint angles at the appropriate time so that the robot can avoid fall. In addition to the body of the robot moving forward, the ankle joint moves backward during this process. The schematic view of hip-ankle strategy is shown in Figure 4(a). The dotted line represents the posture after execution of hip-ankle strategy or prior to execution of hip-ankle strategy. After applying hip-ankle strategy, the robot maintains in that posture for some time at about 500 ms and then returns back to its original posture shown by solid line. The amount of ankle joint traversal is determined based on the position of COM of the end posture (immediately after execution of hip-ankle strategy). Push recovery can also be applied for lateral forces [ref. to Figure 4 (b)]. During double support phase (DSP), the robot is generally stable to lateral forces due to its wider support region. However, while in single support phase (SSP) the robot is very unstable and hence push recovery can be applied at this configuration. This strategy is helpful to the robot in motion, if some external forces act on the mechanism particularly during walking or stair climbing. During SSP, the entire weight of the robot is taken by the ankle motor and the weight of the body is taken by the single hip motor", + " This method does not work if the force pushes the robot outwards (that is, towards right when balancing on right leg or towards left when balancing on left leg). In the present study, a swinging pendulum of mass 450 g and a length of 20 cm are used to apply external disturbing force. First the force needed to topple the robot has been calculated by utilising the inverted pendulum model with a mass of 1.85 Kg (that is, actual mass of robot) at a distance of 22 cm from ground (that is, actual height of COM from the ground) shown in Figure 4(c). It is to mention that the force is assumed to be applied from the back of the robot. The robot foot is represented by line op. When the force is applied, it causes the robot to rotate about the point o and causes the COM to move onto the edge of the foot (that is, which is the condition for toppling of the robot). At this point the principle of conservation of energy can be applied to find the initial velocity of COM during impact. The change in the height (\u2206h) of the COM of the robot has been used to determine the initial velocity of the pendulum" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0002861_978-3-030-55889-5-Figure2.52-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0002861_978-3-030-55889-5-Figure2.52-1.png", + "caption": "Fig. 2.52 Influence function for the average value of \u03c3xx in an element, a the \u201cDirac delta\u201d consists of horizontal forces along the vertical edge and (small, \u03bd-fold) vertical forces along the horizontal edge of \u03a9e, b the horizontal displacements of the Green\u2019s function, here plotted in z-direction", + "texts": [ + "125) but since \u03b5xx = ux ,x and \u03b5yy = uy,y , the domain integral can be replaced by a boundary integral over the edge \u0393e of the element 2.17 Influence Functions for Integral Values 139 \u03c3\u2205 xx = E |\u03a9e| \u222b \u03a9e (\u03b5xx + \u03bd \u03b5yy) d\u03a9 = E |\u03a9e| \u222b \u0393e (ux nx + \u03bd uy ny) ds , (2.126) and because the influence function for the displacement ux or uy of a boundary point x is generated by a single force Px = 1 or Py = 1 placed at x, the influence function for the boundary integral is the displacement field generated by horizontal and vertical line forces, E/|\u03a9e| \u00b7 nx and E/|\u03a9e| \u00b7 ny , along the edge of the element \u0393e, see Fig. 2.52. An equations says it directly: Let G\u2205 the reaction of the plate to the edge forces E/|\u03a9e| \u00b7 nx and \u03bdE/|\u03a9e| \u00b7 ny in Fig. 2.52, which we write as vector t , and let u the displacement field of the plate then W1,2 = \u222b \u03a9 G\u2205 \u2022 p d\u03a9 y = \u222b \u0393 t \u2022 u ds y = W2,1 , (2.127) and the second integral is identical with (2.126). The mean stresses in a plate held fixed at its edge are therefore zero, since the edge forces which generate the influence function cannot displace the fixed edge. The same is true with slabs: The mean values of the moments of a slab, clamped on all sides, are zero. In the 1-D case we have encountered this phenomenon already in Chap" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0000763_s00202-020-00967-y-Figure12-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0000763_s00202-020-00967-y-Figure12-1.png", + "caption": "Fig. 12 Magnetic field distributions at no load and 10 A load", + "texts": [ + " With an increase in the electrical loading (current), the inductive voltage drop in the winding increases and the power factor decreases. For a given current, the power factor of M3 is better than that of M1 and M2, as the M3 has low phase inductance than others. From the plots shown in Figs. 10 and 11, there is a trade-off between the torque and power factor. Selection of rated current decides the torque, power factor and voltage of the machine. Based on the cooling method considered, the rated current is chosen as 10 A (corresponds to current density 5 A/mm2) for further analysis. In Fig. 12, the magnetic field distributions at no load and load are shown for all topologies. This figure is prepared using LATEX animation package to include various rotor positions in onepicture.Bydefault, themagneticfield distribution is shown for the rotor position \u03b8e = 0\u25e6. The field distribution of other rotor positions can be viewed frame by frame using the control buttons provided below the each sub-figure. This file should be opened with Adobe PDF reader to see this animation. At \u03b8e = 0\u25e6, the rotor poles are completely unaligned from the stator poles" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0000967_0954406220916504-Figure20-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0000967_0954406220916504-Figure20-1.png", + "caption": "Figure 20. Compliance in the bushings and the elasticity of the links.", + "texts": [ + " It generates the wrench w\u0302B acting on the vehicle body through the suspension mechanism. The wrench w\u0302B can be determined from equation (20) w\u0302B \u00bc fb r\u0302 T b S\u0302H r\u0302 T HS\u0302H r\u0302H In the suspension, the elastomeric bushing transmits force and torque between its outer and inner sleeve with corresponding elastic deflection. The suspension links can be assumed to include an equivalent longitudinal compliance as a consequence of radial compliance of the elastomeric bushings. The linear stiffness rate of each link may be expressed by ki (i \u00bc 1, . . ., 5) as shown in Figure 20. The wrenches generated by the linear stiffness of the links can be determined as w\u0302li \u00bc flis\u0302li where s\u0302li are the unit line vectors along the links. Since the unit line vectors s\u0302li are the constraint wrenches of the suspension mechanism, the wrenches w\u0302li cannot generate the motions of the suspension mechanism. These wrenches can be viewed as the external forces acting on the vehicle body directly." + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0002478_cac51589.2020.9327315-Figure2-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0002478_cac51589.2020.9327315-Figure2-1.png", + "caption": "Fig. 2. The separation-bearing-orientation scheme of two WMRs.", + "texts": [ + " vR and vL respectively represent the speed of the left and right driving wheels. L is the length of the central axis. r is the radius of two wheels. The characteristic of this model is that it assumes that the wheels purely rolling without sliding on the ground. So, there exists a non-holonomic constriant which is x\u0307 sin \u03b8 \u2212 y\u0307 cos \u03b8 = 0 (4) This formation method can more flexibly control a single WMR in the multi-robot sceanario of leader-follower (LF) mode. The detailed description of WMR-SBOS is shown in Fig. 2. The whole system includes two leaders and two followers. The \u201cLeader\u201d needs to track the \u201cVirtual Leader\u201d, and the \u201cFollower\u201d needs to track the \u201cVirtual Follower\u201d. The trajectory of \u201cVirtual Leader\u201d is a preset path after motion planning in advance. The trajectory of the \u201cVirtual Follower\u201d needs to be calculated in real time in combination with the position of the \u201cLeader\u201d. d represents separation which means the distance between leader and follower. \u03c6 represents bearing between leader and follower", + " (5) Utilizing the actual position coordinate [x, y] as the origin to define the tracking error, the specific form is as follows:\u23a1\u23a3 xe ye \u03b8e \u23a4\u23a6 = \u23a1\u23a3 cos \u03b8 sin \u03b8 0 \u2212 sin \u03b8 cos \u03b8 0 0 0 1 \u23a4\u23a6\u23a1\u23a3 xr \u2212 x yr \u2212 y \u03b8r \u2212 \u03b8 \u23a4\u23a6 (6) Combining with (1) and (5), it can be rewritten as\u23a1\u23a3 x\u0307e y\u0307e \u03b8\u0307e \u23a4\u23a6 = \u23a1\u23a3 ye\u03c9 \u2212 v + cos \u03b8e(x\u0307r cos \u03b8r + y\u0307r sin \u03b8r) \u2212xe\u03c9 + sin \u03b8e(x\u0307r cos \u03b8r + y\u0307r sin \u03b8r) \u03b8\u0307r \u2212 w \u23a4\u23a6 (7) In the system of WMR-SBOS, the tracking error of follower and leader could be represented as (7). However, it should be noted here that the tracking path of the follower needs to be derived in conjunction with the actual motion path of the leader. From the geometric relationship described in Fig. 2, we can see that in order to maintain the d\u2212 \u03c6\u2212 \u03b2 formation state, the virtual path that the follower needs to track is\u23a7\u23a8\u23a9 xv = xl \u2212 dvl cos(\u03c6vl + \u03b8l \u2212 \u03c0) yv = yl \u2212 dvl sin(\u03c6vl + \u03b8l \u2212 \u03c0) \u03b8v = \u03b8l + \u03b2vl. (8) Therefore, on the basis of nonlinear control theory, the tracking controllers are designed as follows for the error Authorized licensed use limited to: San Francisco State Univ. Downloaded on June 25,2021 at 16:57:47 UTC from IEEE Xplore. Restrictions apply. 3794 dynamic models of leader and follower in the WMR-SBOS system: Follower : \u23a7\u23a8\u23a9 vf = (x\u0307v cos \u03b8v + y\u0307v sin \u03b8v) cos e\u03b8f +Kf xexf wf = \u03b8\u0307v +Kf \u03b8 e\u03b8f +Kf y (x\u0307v cos \u03b8v+ y\u0307v sin \u03b8v)eyf sin e\u03b8f/e\u03b8f (9) Leader : \u23a7\u23a8\u23a9 vl = (x\u0307r cos \u03b8r + y\u0307r sin \u03b8r) cos e\u03b8l +Kl xexl wl = \u03b8\u0307r +Kl \u03b8e\u03b8l +Kl y(x\u0307r cos \u03b8r+ y\u0307r sin \u03b8r)eylsin e\u03b8l/e\u03b8l (10) where [xv, yv, \u03b8v] T represents the pose vector of the virtual follower and [xr, yr, \u03b8r] T is the preset trajectory of the virtual leader" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0001917_icuas48674.2020.9214023-Figure5-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0001917_icuas48674.2020.9214023-Figure5-1.png", + "caption": "Fig. 5. Arc with the center at Ct and end positions \u03b1s and \u03b1t", + "texts": [ + " Let Pe be the set containing \u03b1s, \u03b1t and the angular positions corresponding to the extrema of \u2206T , and the entries in Pe are ordered in ascending order of the time of arrival of the target. To find the intercepting position, we systematically perform a bisection search in the intervals defined by the successive positions in Pe. Let Ct represent the center of the circular arc, and rt represent the radius of the arc. The circular arc and the angular positions of the ends of the arc, \u03b1s and \u03b1t , are shown in Fig. 5. A Dubins path from the initial position of the pursuer, pi, to a position \u03b1 on the arc is also shown in the figure. The algorithm that searches for intercepting position on the arc is presented in Algorithm 2. The inputs to this algorithm are the initial position of the 803 Authorized licensed use limited to: University of New South Wales. Downloaded on October 18,2020 at 17:32:00 UTC from IEEE Xplore. Restrictions apply. pursuer, the time of arrival of the target at the start of the arc, the set Pe, and the algorithm outputs an intercepting position if there exists one" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0000215_ecce.2019.8912674-Figure1-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0000215_ecce.2019.8912674-Figure1-1.png", + "caption": "Fig. 1: Cross-sectional view of machine.", + "texts": [ + "00 \u00a92019 IEEE 6147 where the subscripts d and q represent the components of the signal vectors in the synchronous reference frame, s donates the derivative operator, V , I and Te represent the voltage, current and torque respectively, R, L, \u03bbm represent the inductances, resistance, and permanent magnet (PM) flux linkage of the machine respectively and p is the number of magnetic pole pairs. A 3-phase, 9-slot, 6-pole (9S-6P) PMSM with surface mounted magnets (SPMSM). The cross-sectional view of the motor is shown in Fig. 1. The geometric details and lumped model parameters of the machine are given in Table (I). The SPMSM is non-salient pole in nature, i.e., the d and qaxis inductances are approximately equal across the operating space. Further, the effect of magnetic saturation is virtually negligible. In this paper, magnetic saturation is accounted for by scheduling the PM flux linkage and the machine inductances as a function of currents. In particular, the PM flux linkage is assumed to be a function of q-axis current only, while the inductances are assumed to vary with both currents" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0001494_ur49135.2020.9144981-Figure4-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0001494_ur49135.2020.9144981-Figure4-1.png", + "caption": "Fig. 4: The thruster configuration", + "texts": [ + " A water temperature sensor is attached for environmental data collection because water temperature is 496 Authorized licensed use limited to: Macquarie University. Downloaded on July 25,2020 at 19:00:28 UTC from IEEE Xplore. Restrictions apply. one of the typical environmental data for water monitoring application. All the parts are contained in a plastic the cylinder case. The aluminum frame guard protects cylinder case and actuators. The device control system is important for autonomous navigation. We designed the control system, the thruster configuration and the localization method. \u2022 Thruster allocation Fig. 4 shows proposed sensing device\u2019s thruster configuration. Four thrusters are attached to the sensing device in order to move in any arbitrary direction or keep its position. Two sets of thrusters facing each other enables translation and rotation. Fig. 5 indicates the sensing device in fixed coordination. In Fig. 5, \u03b8 indicates thruster angle and \u03b8 = 45\u25e6 in the proposed prototype. l is the length between the center of the device and a thruster. \u2022 Localization and pose estimation method The proposed sensing device has a GPS receiver and two IMU sensors for localization and pose estimation" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0001019_j.precisioneng.2020.04.003-Figure13-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0001019_j.precisioneng.2020.04.003-Figure13-1.png", + "caption": "Fig. 13. Schematic of the coupling similar to Fig. 7. Length dimensions are indicated by lines with endpoints and motions are indicated by lines with single arrows.", + "texts": [ + " 6) For each vee, the distance between the instantaneous center and the points of contact is r or in dimensionless form \u03c1 \u00bc r/R. 7) All the geometry defined above applies for the coupling virtually close to full engagement so that small angle approximations may be used and the shapes of contact surfaces do not make a substantial difference. The derivation of (1) begins by identifying the one-degree-of-freedom motion of the coupling and the sliding directions at the five engaged constraints. Two axes of rotation admissible for constraints 3\u20136 are apparent in Fig. 13: axis A is out of the page and identical to the instantaneous center demonstrated in Fig. 7; and axis B passes through the instantaneous centers for constraints 3\u20134 and 5\u20136. A fifth constraint at 1 or 2 ties rotation about axes A and B to a fixed ratio. Fig. 14-a shows the fifth constraint undergoing a unit displacement along its sliding direction and its components about axes A and B. Combining this with lever arms from Fig. 13, Equations (2) and (3) give corresponding rotations about axes A and B. \u03b8A\u00bc cos \u03b11 2 3 R (2) \u03b8B\u00bc 2 sin \u03b11 2 3 R (3) L.C. Hale Precision Engineering 64 (2020) 200\u2013209 The sliding directions at constraints 3\u20136 are less apparent, involving components in the plane of the vee and out. Upon careful study, constraints 3\u20136 have slightly different sliding directions and resultant forces. The simplifying approach taken here uses those displacements and forces that act in common among the constraints and discards those that are ordinarily small and tend to cancel out of the total. Furthermore, the geometry is simplified by using the instantaneous centers rather than each of the nearby contact points. Referring to Fig. 13, rotation about axis A produces out-of-plane motion common to constraints 3\u20136 as expressed in Equation (4). Rotation about axis B produces both in-plane and out-of-plane motion. For small rotation, \u03b8 B may be treated as an angle vector with components as shown. The larger one produces in-plane motion common to constraints 3\u20136 as expressed in Equation (5). The smaller one produces out-of-plane motion but constraints 3\u20134 and 5\u20136 have canceling directions and are discarded. As these two motions \u03b4 A and \u03b4 B are orthogonal, the resultant motion along the sliding direction is expressed by Equation (6)" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0003481_jae-141829-Figure5-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0003481_jae-141829-Figure5-1.png", + "caption": "Fig. 5. Measurement of magnetic flux density.", + "texts": [ + " Spring constant of variable rate spring. compression spring in the center. Magnetic springs are placed 175 mm to the right and left. The constant of the compression spring is 9.8 \u00d7 104 N/m. The upper parts of the two magnetic springs and the top of the compression spring are linearly installed and move only up and down. A magnetic spring should meet the requirements of Eq. (1). However, it is preferable to use as strong a magnet as possible. Thus, the Halbach array is used to create the magnet for the magnetic spring. Figure 5 shows the measurement results of the flux of magnetic induction by the Halbach array. The magnetic induction was obtained approximately 1.5 times compared with the single magnet. The rotation mechanism of the magnetic spring is shown in Fig. 6. The belt pulley on both sides of the magnetic spring is rotated by using the belt from the belt pulley on the motor side. The excitation mechanism used in this research is installed in the upper part of the variable rate spring. The vibration is induced that addresses the self-weight of the exciting part" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0001448_stab49150.2020.9140656-Figure1-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0001448_stab49150.2020.9140656-Figure1-1.png", + "caption": "Fig. 1. Two coupled inverted pendulums.", + "texts": [ + " At the second step, when matrix K1(t) was determined, the following optimization problem is Authorized licensed use limited to: Carleton University. Downloaded on August 06,2020 at 08:36:56 UTC from IEEE Xplore. Restrictions apply. solved: trace(Q1(t)) + trace(Q2(t))\u2192 min with constraints (11) and Q(t0) \u2265 (1 + \u03bb)R0, Q(t) \u2264 R(t), ( Q2(t) I I P2(t) ) \u2265 0. Here optimization is performed on matrix variables Q1(t), Q2(t), P2(t), Y2(t), K2(t) and scalar variable \u03b22(t), \u03b22(t). As a result, matrices L(t) = P\u221212 (t)Y2(t) and K2(t) are defined. A system consisting of two inverted pendulums is considered (see Fig. 1). Each of them has mass m, length l and is set in motion by its engine, which creates moments T1, T2, respectively. The pendulums are connected to each other at a height using an elastic spring with a stiffness coefficient \u03ba. The position of the pendulums is determined by the angles \u03b81, \u03b82 of their deviations from the vertical axis. The dynamics of the parallel movement of the inverted pendulums is determined by the following nonlinear differential equations ml2\u03b8\u03081\u2212mgl sin \u03b81+\u03baa2 [cos \u03b81 (sin \u03b81 \u2212 sin \u03b82)] = T1\u2212w11, ml2\u03b8\u03082\u2212mgl sin \u03b82+\u03baa2 [cos \u03b82 (sin \u03b82 \u2212 sin \u03b81)] = T2\u2212w21, Neglecting electromagnetic transients in the armature winding of a DC motor the expression for the moment is represented as Ti = KgKm r ui \u2212 (KgKm)2 r \u03b8\u0307i, i = 1, 2, where Kg,Km, r are coefficients of reduction, proportionality and active resistance of the motor winding, ui is control voltage" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0002782_0954410019872116-Figure3-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0002782_0954410019872116-Figure3-1.png", + "caption": "Figure 3. Illustration of Pirouette maneuver.", + "texts": [ + " Initial state of the helicopter is assumed to be in hovering position with a certain altitude. The initial states and inputs of the hover motion of AAU Bergen Industrial Twin RC are given in Table 3. Pirouette maneuver is a circular path with a fixed altitude at which the heading is always directed towards the center of the circle during the complete simulation. Helicopter rotates around a circumference of a fixed radius while rotating about its own yaw axis to keep the north axis pointing to the center of the circle. Pirouette maneuver is illustrated in Figure 3. Nonlinear simulations are initiated at the hover trim condition. The SDC matrices are computed at the hover trim condition to obtain the linear model of the helicopter. Then this linearized system is utilized to design the linear SMC for the cascaded autopilot structure given in Figure 1. The system matrices of the linear system are given as follows ASAS \u00bc AATT \u00bc 0 0 0 0 0 0 0 0 0 0 0 0 0 0 0 0 0 0 1 0 0 0 0 0 0 1 0 0 0 0 0 0 1 0 0 0 2 66666664 3 77777775 Table 4. Percentage of parameter uncertainties" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0002540_iccr51572.2020.9344388-Figure1-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0002540_iccr51572.2020.9344388-Figure1-1.png", + "caption": "Figure 1. Flywheel inverted pendulum model: is the fulcrum of the pendulum and is set at robot\u2019s ankle as a virtual contact point, is the position coordinate of the COM, and are respectively the x-axis and zaxis components of the position of COM, is the centroidal angular momentum, is the external torque acting on the virtuall contact point,", + "texts": [ + " By introducing the cumulative task, the walking controller no longer concentrates on the tracking of the COM trajectory at every control cycle of the entire stepping process but focuses on the completion of the cumulative task at the landing time, to reduce the dependence of the controller on the COM trajectory. In our control framework, furthermore, a hybrid foot placement trajectory with a partial reference trajectory calculated by the feedback state is adopted to reduce the dependence on the precalculated footstep trajectory. II. CONTROL MODEL To analyze the overall motion of bipedal robots, a flywheel inverted pendulum (showed in Fig. 1) is adopted to describe the relationship between the center of mass and angular momentum. The mathematical description of the model is as follows: \u0307 = \u2212 \u0308 + ( \u0308 + ) + (1) is the acceleration of gravity, is the total mass of the robot. In order to obtain the numerical solution of the motion, the robot is considered to walk in the LIPM-like walking pattern, that is, the walking pattern in which the height of the center of mass remains constant and angular momentum has little influence on the motion", + " The joint configuration of the simulation model is showed in the Fig. 2. The floating frame is set in the middle of the two \u210e 1 joints and is attached to the torso. The position and posture of the whole robot are described based on the floating coordinate and joint coordinate. Due to the walking study is carried out in the sagittal plane, the floating coordinate and actuated joint coordinate are respectively established as =[ , , ] (representing the x-axis position, z-axis position and rotation of the floating base within the sagittal plane as shown in Fig. 1) and = , , , , , , , , , , , , while the remaining joints are considered to be fixed. The size, mass and inertia parameters of model are described as = [ , , , , , , , ], = [ , , , , , , , ], = [ , , , , , , , ]. In the simulation model, each joint is attached to a corresponding rigid body with mass and inertia. Through the Euler-Lagrange method, the dynamics of the mode can be expressed as the following equation: ( ) \u0308 + ( , \u0307 ) \u0307 + ( ) = ( ) + ( ) ( ) \u0308 + (\u0307 ) \u0307 = 0 (20) where = , is the generalized coordinate of whole model, \u2208 \u211d \u00d7 is the torque of the actuated joint, =, , , , is the vector of the generalized ground reaction force equivalent to the ankle of the standing leg, ( ) \u2208 \u211d \u00d7 is the inertia matrix, ( , \u0307 ) \u2208 \u211d \u00d7 is the Coriolis and centrifugal matrix, ( ) \u2208 \u211d \u00d7 is the gravity vector, ( ) \u2208 \u211d \u00d7 is the mapping from the joint torque to the generalized torque, ( ) is the jacobian of coordinate of the virtual contact point set at ankle joint" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0003342_978-3-319-07398-9_9-Figure3-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0003342_978-3-319-07398-9_9-Figure3-1.png", + "caption": "Fig. 3 Contest table for 2.007 in 2005", + "texts": [ + "007 work in a lab that is divided into two parts: a machine shop with six lathers and six milling machines as well as sheet metal equipment and large bandsaws; and a 6 m 9 20 m assembly area with benches and six sets of drill presses belt sanders and small bandsaws. There is also a soldering bench and a fasteners cabinet. Table 2 shows the supporting staffs assigned for teaching 2.007 in 2005. The contest table for year N is newly designed every year by students and staff who completed the course in year N - 1. The 2005 table (MIT 2.007 2005) was inspired by the new undergraduate dormitory at MIT called Simmons Hall, and designed in the summer of 2004 as shown in Fig. 3. Each student\u2019s machine must begin the competition within an 18 inch by 18 inch square. The starting zone extends to 26 inches above the table surface. Robots can be located anywhere within this volume. However, they still must fit into the sizing box (16 9 16 9 24) when in the starting zone. The starting zone is demarcated by a 1 inch line. The score is the sum of the total points scored by placing colored foam blocks into the various scoring bins as shown in Fig. 4. A student can only score with the foam blocks of the color of their side of the table, but they can place a block in any scoring bin (Fig" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0001748_j.matpr.2020.07.354-Figure6-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0001748_j.matpr.2020.07.354-Figure6-1.png", + "caption": "Fig. 6. Max. Principle stress for the casting and fabricated planet carrier noted 143.27 N/mm2 (Mpa).", + "texts": [ + " Please cite this article as: G. Vignesh, M. Prakash and M. Arun, Numerical analy tion stage in industrial gearbox, Materials Today: Proceedings, https://doi.org/ Average element dimension 1.5 mm Note: in the areas where maximal stress values are expected, it\u2019s been used a finer mesh in order to get a better analysis result. Fig. 5 shows the mesh geometry. This theory fails when the maximum main stress of the system reaches the value of the maximum resilience in simple tension at the elastic limit as shown in Fig. 6. The calculated stress values on the planet carrier in correspondence of the peak of torque have a magnitude lower than Yield stress of material. According with the obtained results using the extreme torque 1,45,000 Nm, the safety factor against yield stress sis on replacement of integrated planet carrier at maximum torque reduc10.1016/j.matpr.2020.07.354 Fig. 4 (continued) As we mentioned SG500 for casting of planets and FE510 for the manufactured planet carriers have been used. The cumulative theory tension is increased from 143", + "org/ safety of fabricated planet carrier was lesser then casting planet carrier, the safety of fabricated one was less high than factor of safety required, whereas the safety of casting one was much higher and give the criteria over design for more safety. The Production cost of the Fabricated planet carrier was lesser than casting planet carrier, as the casting planet carrier requires separate foundry, Pattern and core preparation, Labour charge, Time for casting to solidify and much investment for all the above, whereas Fabrication was done in simple method whereas all the plates should be welded with a single welder and very quick process. sis on replacement of integrated planet carrier at maximum torque reduc10.1016/j.matpr.2020.07.354 Fig. 6 (continued) The cost of the SG500 is Rs 110/ Kg and the cost of FE510 is 90/ Kg. Since the per Kg cost of SG500 is more than the FE510 and casting process takes more lead time and investment than fabricated, the casting planet carrier can be replaced with the fabricated planet carrier for the unit quantities of gearboxes to be manufactured which involves higher torque. The calculation and analysis procedure remain same as discussed, for all the torque capacities. Declaration of Competing Interest The authors declare that they have no known competing financial interests or personal relationships that could have appeared to influence the work reported in this paper" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0001148_1077546320927598-Figure1-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0001148_1077546320927598-Figure1-1.png", + "caption": "Figure 1. Three dimensional model of robotic manipulator.", + "texts": [ + " The reminder of this article is organized as follows. In Section 2, the system dynamic model is established via the second Lagrange equation. In Section 3, the PD-like continuous time-varying feedback control law is presented and the system stability is proved. In Section 4, experiments under different conditions are conducted and results are analyzed and discussed. Finally, in Section 5, this article is concluded. 2. System modeling and identification The model of the robotic manipulator studied in this article is shown in Figure 1. The transfer arm is driven by a direct current (DC) motor through two reducers. A gas spring is linked to the transfer arm to provide a balance torque. The entire robotic manipulator system is mounted on a profiled frame structure, which is mounted on a 3- degree-of-freedom vibration platform. The vibration platform is composed of three same servo motors and three same electric cylinders. It can produce vibration in three directions: (1) in vertical direction (linear vibration along axis y), (2) in pitching direction (rotational vibration relevant to axis z), and (3) in swing direction (rotational vibration relevant to axis x)", + " Figure 4 shows the convergence process of objective function. The value of objective function converges to 1.87 at the end of iteration. The specific identification values of parameters are shown in Table 1. It can be concluded that the identification result is reasonable. 2.3. Balance torque modeling and identification The robotic manipulator in this article has two working conditions: with and without load, and the load is mounted at the end of the transfer arm. Besides, under the condition of with load, the mass of load varies. From Figure 1, we can see that while the mass of load varies, the gravity moment of the transfer arm differs greatly. To keep the motor working at a relatively constant load, thereby improving the motion smoothness of the robotic manipulator, it is particularly important to give the transfer arm a proper balance torque during its working process. Under the conditions of working with and without load, the system gravity moments are modeled as follows MG \u00bc m1gL1 cos\u00f0\u03b8 \u00fe \u03b8b\u00de \u00fe m2gL2 cos\u00f0\u03b8 \u00fe \u03b8b \u03b1\u00de (7) MG \u00bc m1gL1 cos\u00f0\u03b8 \u00fe \u03b8b\u00de (8) By comparing equations (7) and (8), the difference of gravity moments under two loading conditions is m2gL2 cos\u00f0\u03b8 \u00fe \u03b8b \u03b1\u00de", + " In this group, comparison experiments, control with and without compensation, are conducted. The second group considers the influence of load uncertainty under the condition of with compensation. The case of load mounted at the end of the transfer arm in this group of experiments is: (1) case 1: no load; (2) case 2: one mass block, 0:88 kg; and (3) case 3: two mass blocks, 1:76 kg. First, the vibration of the robotic manipulator installation platform is tested. The platform works in the mode of vibrating in both vertical and pitching directions as shown in Figure 1. Figure 7 shows the sensor installation diagram in the base vibration test. Angular velocity sensor, laser displacement sensor, and acceleration sensor are used here to measure the base vibration data. And Figures 8\u201310 give out the vibration curves tested by the three sensors. The amplitude of angular velocity of the base pitching vibration is 0:92\u00f0rad=s\u00de, the amplitude of displacement of the base vertical vibration is 32:6mm, and the amplitude of acceleration of the base vertical vibration is 4\u00f0m=s2\u00de" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0001186_3334480.3383139-Figure5-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0001186_3334480.3383139-Figure5-1.png", + "caption": "Figure 5: Pressure Touch Matrix.", + "texts": [ + "de/rio/ Arranging: Functional tapes can be combined with patches that are placed at the required position, ironed-on and then electrically connected in one stroke (Figure 3). While tapes can be arranged side by side manually (Figure 2, C), there are cases where a certain arrangement is needed. Layering: The RIO approach also supports versatile layering options to create new functionalities (e.g., shielded traces, multi-layer circuits or custom-shaped sensors by ironing multiple material layers on each other as shown in Figure 5) or to combine existing functions by ironing functional modules one above the other (e.g., combining visual output and touch tapes at the same position). Since our materials are ultra-thin and flexible, the capability of layering is one of the major advantages over existing approaches that use yarn or threads to isolate or extend conductors. Utilizing Textile Accessories: Finally, the approach also enables the digital enhancement of existing textile accessories like zippers (see Figure 4) that serve as a physical connection or shirt buttons (Figure 7, C) that act as joystick by ironing a four-electrode capacitive patch underneath", + " Moreover, traces are characterized by various additional properties for customization: level of conductivity, number of parallel wires (e.g., as required for bus systems), presence or absence of electrical shielding, strechability, and its connector pitches. We demonstrate how our iron-on device works in concert with iron-on trace spools to create basic conductive traces or traces with advanced multi-wire, elastic, and shielded properties. We further present methods to create, connect (see Figure 6) and delete complex trace designs and realize custom-shaped sensor matrices (see Figure 5). INT014, Page 3 Elastic Smart Cuff button at the cuff (A), Doctor\u2019s White Coat with identification and e-ink patch (B) and a Messenger Bag (C) with solar cells, touch controls, interior lights as well as moisture sensors. To demonstrate the feasibility of integrating various functional textile materials, printed electronics components and flexible PCBs inside iron-on tapes and patches, we present a repertoire of RIO components (see Figure 1, F) that allows us to leverage on a rich set of established principles for building functional iron-on modules" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0000372_s0036024419120355-Figure2-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0000372_s0036024419120355-Figure2-1.png", + "caption": "Fig. 2. Dependences of heat of adsorption , kJ/mol, on molecular polarizability \u03b1M, \u00c53 in the series of mono derivatives of thiophene (s) and benzene (d).", + "texts": [ + "185 nm [22]) being greater than the half-thickness of the benzene ring (0.177 nm) [23]. The CH groups in the thiophene molecule in close proximity to the S atom (ring positions 2 and 5) are thus elevated above the f lat surface of graphite, weakening their interaction with the adsorbent and reducing ln K1,c and . The conclusion that the sulfur atom raises the plane of the thiophene ring above the graphite surface confirms the presence of two separate dependences = f(\u03b1\u041c) for mono derivatives of benzene and thiophene (Fig. 2): the dependence for derivatives of thiophene in the graph coordinates = f(\u03b1\u041c) lies below the corresponding dependence for benzene derivatives. Similar effects were observed in [24, 25] while studying the adsorption of other aromatic compounds with bulky substituents on a surface of graphite. Differences in the adsorption of the main structural fragments (benzene and thiophene rings) also influence the condif,1q dif,1q dif,1q RUSSIAN JOURNAL OF PHYSICAL CHEMISTRY A Vo Table 3. Contributions of various functional groups to the heats of adsorption ( , kJ / mol) of mono derivatives of thiophene and benzene on the surface of Carbopack C HT Derivatives of Cl\u2212 Br\u2212 I\u2212 Thiophene 6" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0003261_aim.2015.7222734-Figure3-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0003261_aim.2015.7222734-Figure3-1.png", + "caption": "Figure 3. Roll, pitch, yaw, and rotation motions.", + "texts": [ + "00 \u00a92015 IEEE 1394 compass modules were used as sensors for detecting the location and direction of the hexarotor, respectively. In general, adjustments to the roll, pitch, yaw, and elevation are made to allow a helicopter to move to a certain three-dimensional coordinate in a global coordinate system. For a hexarotor-type helicopter, these commands are achieved by controlling the direction of movement through adjusting the rotational speed of each rotor, and by implementing the desired motion characteristic through the velocity feedback control from the gyro sensor in the sub-control MCU, as shown in Fig. 3 [5][7]. First, as shown in Fig. 3(a), the roll direction control is achieved by setting the X axis passing through the hexarotor as the symmetry axis, and then creating a difference in the rotational speed of the right and left rotors in order to control the thrust that is generated from this difference. The thrust generated from a difference in the rotational speed, compared to that of the roll, is used to control the pitch. However, a more complicated thrust method is required for an asymmetrical multi-rotor helicopter such as the BM-01 hexarotor in order to achieve roll and pitch. In this case, the pitch motion is generated by raising the speed of rotors 1, 2, and 6 in Fig. 3(b) to increase thrust, and lowering the speeds of rotors 3, 4, and 5. For the directional control of the yaw (azimuth), the reverse torque generated by each rotor is used to control the vehicle. This torque is the reaction force of the moment for rotating the rotor, and occurs in the opposite direction of each rotor\u2019s rotational direction. The azimuth is controlled through the moment generated by the tail rotor cancelling the reverse torque that is created by the main rotor. However, for multi-rotor-type helicopters, the reverse torque is cancelled by reversing the rotational direction using rotors on the opposite half of the vehicle, as shown in Fig. 3(d). To rotate the vehicle in the yaw direction, a rotational difference is generated by the rotors in the clockwise and counter-clockwise directions, as shown in Fig. 3(c). In Fig. 3(c), the speeds of rotors 1, 3, and 5 rotating in the counter-clockwise direction are increased, and the speeds of rotors 2, 4, and 6 are decreased. By doing so, the torque in the opposite direction of the clockwise direction increases, causing the vehicle to rotate in the clockwise direction [8][10]. This section introduces the mathematical model of the hexarotor-type multi-rotor helicopter for flight tests. This model is basically obtained by representing the hexarotor as a solid body evolving in three dimensional space" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0002108_s40799-020-00405-5-Figure3-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0002108_s40799-020-00405-5-Figure3-1.png", + "caption": "Fig. 3 Schematic drawings of six-axis load head (all dimensions shown are in inches): a side view; b plan view (top and bottom plates)", + "texts": [ + " As mentioned in the previous section, a Stewart platform consists of a base plate, and a top platform, connected by six struts. A SOLIDWORKS [6] model, and a picture, of the actual devices discussed here are shown in Fig. 2. The load heads consist of two hexagonal plates, connected by six commercially available 1000 lb uniaxial load cells (LC1011K [26]); other types of individual load cells may be used depending on the requirements of an application. In the interest of making the load heads as light as possible, the plates are made out of aluminum and, as shown in Fig. 3, circular holes are cut at their center. The hole in the bottom plate is smaller merely to fit a bolt to connect it to an electrical bushing (see Fig. 1). The six uniaxial load cells are connected to both plates with spherical joints (2458K161 [21]) and shoulder bolts (91273A316 [22]). Strictly, a universal joint should have been provided at one end to completely prevent rotation of the load cells about their axes. However, accommodating such joints would have made the load head bigger and heavier", + " The Stewart platform has been widely used as a motion simulation base in various applications ranging from road simulators and flight training to seismic simulators. However, its use as a sensor is less common. In past sensor implementations outlined in \u201cIntroduction\u201d, the concept has been used in application-specific configurations, and typically in smaller scale in the context of robotics. The load head design presented in this paper is generic and scales readily; the only design change necessary is to change the thickness of the top and bottom plates (Fig. 3) to accommodate shoulder bolts with the necessary strength. In fact, even in the geometry exactly as presented, individual load cells of capacity up to 2,500 lb can be employed, thus without changing the construction, but varying the range and sensitivity. The aspect ratios of the base geometry presented can also be easily modified to obtain non-isotropic sensitivity if desired. Furthermore, the use of high-sensitivity uniaxial load cells as constituents results in both high stiffness and uniform sensitivity in all axes" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0001896_icuas48674.2020.9213976-Figure2-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0001896_icuas48674.2020.9213976-Figure2-1.png", + "caption": "Fig. 2: Illustration of tilting angles \u03b1i and \u03b2i", + "texts": [ + " Assumption 3: The origin of the body coordinate system coincides with the centroid of the fuselage and the moment of inertia tensor is a diagonal matrix. Assumption 4: The gyro effect is not considered because the rotation motion equations will be more complex if the gyro moment items are included in the equations. Furthermore, the fuselage mass and inertia moment are relatively small. Assumption 5: To ensure that the tilting rotors do not hit the shafts, we limit the tilt angles. Thus, the tilting angles range from: \u03b1i \u2208 [\u03b1min, \u03b1max] , \u03b2i \u2208 [\u03b2min, \u03b2max]. These angles are defined in the nomenclature and depicted precisely in Fig.2. The specific values of \u03b1min, \u03b1max, \u03b2min, and \u03b2max are given in SectionIV. Based on these assumptions, classical mechanics theory can be employed. According to the Newton-Euler formalism of rigid body and mechanics theories, we obtain the following formula [28]: [ mI3\u00d73 O3\u00d73 O3\u00d73 I ] [ V\u0307 B \u03c9\u0307B ] + [ \u03c9B \u00d7 ( mV B ) \u03c9B \u00d7 ( I\u03c9B ) ] = [ FB MB ] (2) where m, I = diag(Ix, Iy, Iz), \u03c9B = [p, q, r]T , and 330 Authorized licensed use limited to: Middlesex University. Downloaded on October 18,2020 at 08:03:39 UTC from IEEE Xplore" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0001845_icra40945.2020.9196746-Figure1-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0001845_icra40945.2020.9196746-Figure1-1.png", + "caption": "Fig. 1 Developed parallel gripper.", + "texts": [ + " Therefore, in developing our finger mechanism, we decided to arrange the elastic parts and the contact sensor such that the configuration has a finger mechanism that is thin. In this study, we aimed to expand the applications of parallel grippers, and thus we proposed and verified the operation of a displacement-magnification mechanism and a finger mechanism with fingers that are both extendable and thin, as well as attachable on a commercially available parallel gripper. The developed parallel gripper is shown in Fig. 1. The displacement-magnification mechanism has a stacked rack-and-pinion system that doubles displacement. The extendable finger mechanism has two nails that extend and contract, reducing impact force and detecting changes in product height from expansion and contraction amounts. In this paper, we report on the design process, the specific mechanism and system configurations, and results of verifying the mechanisms\u2019 operations. II. DEVELOPMENT CONCEPT The development concept of the parallel gripper in this study is to double the displacement while preventing damage to products by mounting a modular displacement-magnification mechanism and an extendable finger mechanism on a commercially available parallel gripper that horizontally opens and closes" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0002342_rcar49640.2020.9303289-Figure1-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0002342_rcar49640.2020.9303289-Figure1-1.png", + "caption": "Figure 1. Rope drive schematic diagram", + "texts": [ + " Flexible rope drive uses the motor drive, and the two ends of the rope are respectively connected to the foot wearing end and the motor steering wheel end. The prototype realizes the tightening force of the rope through the winding device. But the rope can only provide tension to achieve dorsiflexion. It is unable to be subjected to the opposing pressure so that the plantarflexion movement could not be achieved. So we add a suitable tension spring between the heel and the wear of the calf for the plantar return movement of the assisting equipment. A schematic diagram of the mechanical structure is shown in figure 1. Kefan Xing, Yinghan Wang, Diansheng*Chen Min Wang and Sitong Lu are with Institute of Robotics, Beihang University, Beijing, 100191, China. chends@163.com. 978-1-7281-7293-4/20/$31.00 \u00a9 2020 IEEE 464 Authorized licensed use limited to: Univ of Calif Santa Barbara. Downloaded on May 19,2021 at 02:45:08 UTC from IEEE Xplore. Restrictions apply. The flexible ankle-assist robot prototype mainly includes a rope drive mechanical module, a spring recovery module, a foot wearing module and a calf wearing module" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0003241_ceit.2015.7233021-Figure1-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0003241_ceit.2015.7233021-Figure1-1.png", + "caption": "Fig. 1. Inverted pendulum system", + "texts": [ + " In this paper, state constraints are integrated in predictive control where constraints are translated in linear equalities and control vector is obtained by solving a linear quadratic cost function . The extracted control law is applied to cart with inverted pendulum system in order to walk the cart in predefined trajectory while respecting state constraints those of: pendulum angle, pendulum velocity, cart position and cart velocity. Note here that state constraints are proposed in a way which can help to keep the pendulum in balanced position. II. STATE SPACE MODEL OF INVERTED PENDULUM Forces and moments acting in the system were analyzed using figure fig. 1 where represents respectively the angle of pendulum rod, M and m stands for the weight of the cart and pendulum, lS is a distance between centre of gravity of the pendulum and the centre of rotation of the pendulum and g is the gravity acceleration constant. Symbol F represents the force [15]. The mathematical model describing the inverted pendulum is most easily derived by using Lagrange\u2019s equations [16]. In this example it is natural to use the cart position x and the angle of deflection of the pendulum arm as the coordinates describing the system" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0001117_kem.841.327-Figure8-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0001117_kem.841.327-Figure8-1.png", + "caption": "Fig. 8 Equivalent stress in Specimen.", + "texts": [ + " The increases in surface area of crack propagation, increases the impact toughness of the material, which enhances the Impact energy of the material, when compared to unidirectionally arranged fibres in the same stacking sequence. Stress \u2013 Strain Graph. Graph 3 shows the stress vs strain relation of the hybrid composite, with the displacement till the Yield from experimental results. The maximum tensile strength obtained is 56.038 MPa. This value is higher than the experimental tensile strength by 17.94 %. The perfect linearity of the stress to strain is due to absence of defects such as voids, fabrication errors in the FEA model. Equivalent Stress. Fig. 8 shows the equivalent stress distribution in the hybrid composite. It can be seen that the maximum stress is obtained in the middle layer of the composite, which is GFRP. A symmetricity exists in the stress distribution about the Z-plane, as noted form the figure. Equivalent Stress in GFRP. Fig. 9 shows the stress in the GFRP layer on tensile loading. A maximum stress of 83.267 MPa is obtained. This indicates that the distribution of the longitudinal loading on the GFRP is higher. The maximum stress of the entire composite is enhanced by the presence of three layers of GFRP" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0003544_cta.2127-Figure9-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0003544_cta.2127-Figure9-1.png", + "caption": "Figure 9. (a) Establishing 7-coil equivalent circuit. (b) Establishing 5-coil equivalent circuit.", + "texts": [], + "surrounding_texts": [ + "The physical meaning of the 7-coil and the 5-coil equivalent circuits [1, 2] is best visualized with the help of Figures 8 and 9, which is based on the commutating process on the DC machine over one revolution. The 7-coil and 5-coil equivalent circuits represent the operation of the rotating DC machine during a certain period. Namely, as a 7-coil, the armature winding is presented by seven coils all together, as both the positive and negative brushes are contacting three segments at the same time. Soon after the rotor moving over the 7-coil circuit, alternatively, another one is presented by the 5-coil armature winding of the machine when only two segments are under commutation. Consequently, the process of the DC machine commutation involves only these two equivalent circuits, either the 7-coil or the 5-coil from time to time." + ] + }, + { + "image_filename": "designv11_71_0003275_aim.2015.7222530-Figure7-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0003275_aim.2015.7222530-Figure7-1.png", + "caption": "Figure 7. Optimal configuration with 4 segments in the region 900 < y < 1800.", + "texts": [ + " We can use Table I to find the optimal configuration for the array of magnets that enables us to obtain the maximum Bx at the centre of the system. For instance, at y=900, A3 contributes to Bx with 19.1 mT while the contribution of Al at this position is nil. Similarly, at y=1200, A3 contributes to Bx with 16.5 mT while the contribution of Al at this position is slightly higher. Following this methodology, we find that the optimal configuration to accommodate magnets A 1 and A3 in the region of interest (i.e., 90\u00b0 < y < 180\u00b0) is as shown in Fig. 7. By considering the symmetry of the magnetic system with respect to the x and y axes, we fmd the optimal configuration in the entire circular trajectory to be the one shown in Fig. 8 that consists of 10 segments radially magnetized (5 segments of the type Al and 5 segments of the type Az) and two segments tangentially magnetized. By following the same procedure with the variation of By, we find the same results. We can also use a curve fitting process and obtain the same optimal configuration if we note that the variation of Bx and By can be expressed as sinusoidal functions with peak amplitudes of 19" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0001285_icieam48468.2020.9112086-Figure2-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0001285_icieam48468.2020.9112086-Figure2-1.png", + "caption": "Fig. 2. The design of the valve with electrical drive", + "texts": [ + " With reference to the problem of automatic control of the gate position the choice of the drive mechanism (the controlled object) is determined by the following requirements: ensuring the necessary multiplicities of the minimum, starting, nominal and maximum (critical) moments of the drive in the appropriate operating modes of the shut-off valves; corresponding the traction characteristic of the drive to the load characteristic of the gate; ensuring the possibility of controlling the position of the gate and the force applied to the shut-off element; ability to automate controlling processes; high reliability and noise immunity. With this in mind we chose a hermetic electromechanical converter with the linear movement of the executive element connected to the external network through a frequency converter as the driving mechanism [5\u20137]. The design of the gate with the electrical drive is conventionally shown in Fig. 2. The electromechanical drive of this gate valve consists of a magnetic circuit with the network winding 1 placed on it, the rotating short-circuited secondary winding 2 on which inner surface a thread is applied, the rod 3 connected to the gate 4, the rings 5 and 6 of an antifriction material performing the function of radial-thrust sliding bearings. There is the threaded connection between the rotating winding 2 and the rod 3. At least two strain and/or piezoelectric sensors 7 connected in the differential measuring circuit are integrated into the ring 6. The pressure cover 8 connected to the base of the gate 9 provides the structural unity of the device. The sealing element 10 is located between the cover 8 and the base 9. The magnetic core with a network winding is made in the form of a hermetic module with a non-conductive shell (Fig. 2 the shell is not shown due to the small thickness of the coating). At the synthesis of the system, the controlled object is replaced by a modified mathematical model based on a generalized electromechanical converter. This model in the \" , \" axes has the form [8, 9]: s s r r s s ssssr ssssr rsrsrr p r pp ppp rrr a rsrs srsss srsss s s r r s s i i i i i i L dt drM dt dM dt d M dt dL dt drM dt d M dt dM dt dL dt drLMM MMLL dt drM dt dM dt d M dt dL dt drM dt d M dt dM dt dL dt dr u u u u u u 2 1 1 1 2 1 222121 121111 12111111211 12111111211 212221 111211 2 1 1 1 2 1 000 000 000 000 The windings have the designation indicating their belonging to the axes \" \"or \" \", the serial number" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0003788_detc2014-34213-Figure10-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0003788_detc2014-34213-Figure10-1.png", + "caption": "Fig. 10 Seven-bar linkages with inputs not in the same loop", + "texts": [ + "org/pdfaccess.ashx?url=/data/conferences/asmep/82119/ on 04/05/2017 Terms of Use: http://www.asme.org/about-asme/terms-of-use collinear or the joints D, C and C0 become collinear. Therefore, the two-DOF seven-bar linkage is at singular positions. Input joints not in the same five-bar loop In this group, the inputs are given through the joints not in the same five-bar loop of the seven-bar linkage, such as joints A0 and C0, the seven-bar linkage can be decomposed into a Stephenson six-bar linkage shown in Fig. 10(a) with the input joints A0 held or Stephenson six-bar linkage shown in Fig. 10(b) with the input joint C0 held. Thus, for the Stephenson six-bar linkages, the singularity happens when links AE, B0B and CD intersect at a common point. Therefore, the two-DOF seven-bar linkage is at singular positions. (a) (b) Fig. 6 Seven-bar linkage with input joints A and A0 (a) (b) Fig. 7 Seven-bar linkage with input joints A0 and C0 (a) (b) Fig. 8 Seven-bar linkage with input joints A0 and B 3.2 Two-DOF Planar Linkages with Prismatic Joints The proposed method can be also applied to the two-DOF planar linkage with prismatic joints" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0003880_scis-isis.2014.7044812-Figure5-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0003880_scis-isis.2014.7044812-Figure5-1.png", + "caption": "Fig. 5. Representation of new repulsive force component: (a) On the Right side, (b) On the left side", + "texts": [ + " The rate of change of the depends on the parameter value of that should be properly defined to limit the variation within the range of . The matrix defines the direction for new repulsive force to appear and its right-side ( ) and left-side ( ) matrices are expressed as in (8). This is defined in such a way to move the robot always towards the goal while moving away from the obstacles. (5) The obstacle-velocity repulsive force acts on the robot when there are obstacles within the detectable range of the robot sensors as explained in the Fig.5. This shows the generation of the new repulsive force component that varies according to the direction to the obstacles. The secondary repulsive force should always act on the goal side to have the total repulsive force that helps to bring the robot towards the goal. As shown in the Fig.5 when the goal is located in the right hand side from the obstacle and the robot, the secondary repulsive force should be in the right side of the robot. When the goal is in the left hand side, force should act on the left side. Therefore the matrix should be selected properly. As a result of the two components of the repulsive forces with different magnitudes for different values of , total repulsive force has a rotation from original repulsive force. This will support robot to move away from the local minima, and for a smooth and dead-lock free motion" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0003788_detc2014-34213-Figure16-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0003788_detc2014-34213-Figure16-1.png", + "caption": "Fig. 16 Eight-bar linkages with input joints A0, A and B0", + "texts": [ + " It requires three inputs to conduct a constrain motion. If no specific relation between the three inputs is required and the three inputs can be arbitrarily placed, the linkage (Fig. 14) can be regarded as a two-DOF planar seven-bar linkage by holding one input joint or a single-DOF planar six-bar linkage by holding two input joints. For example, if the inputs are given through the joints A0, A and B, the eight-bar linkage degenerates into a Stephenson six-bar linkage by holding joint A and B0 (Fig. 16a). Therefore, a three-DOF planar linkage can be decomposed into a twoDOF planar linkage and an additional input joint or a singleDOF planar linkage and two additional input joints. If any of the two-DOF planar linkage or single-DOF linkage is at singular positions, the whole linkage should be at singular positions. Hence, the following criterion can be used to analyze the singularities of the three-DOF eight-bar linkages. Criteria 1: For a three-DOF planar eight-bar linkage, it requires three inputs for a constrain motion", + " The singularity of the six-bar linkages or the two-DOF seven-bar planar linkage happens when the joints E, F and B0 are in the common line or the joints D, C and C0 are collinear, which is also the singular condition for the three-DOF eight-bar linkage. These two composition methods can lead to the same conclusion. According to the above classification, Type-112 is the second case, in which the inputs given through the joints A0A B0, or A0FB0, or BFB0. Let the inputs be given through the joints A0, A and B0 of the six-bar loops in Fig. 16(a) as an example. There are also two decomposition ways. One is the three-DOF eight-bar linage is decomposed into a Stephenson 5 Copyright \u00a9 2014 by ASME Downloaded From: http://proceedings.asmedigitalcollection.asme.org/pdfaccess.ashx?url=/data/conferences/asmep/82119/ on 04/05/2017 Terms of Use: http://www.asme.org/about-asme/terms-of-use planar six-bar linkage with the input joints A0 and B0 held, as shown in Fig. 16(a). The other is the three-DOF eight-bar linage is decomposed into a two-DOF planar seven-bar linkage with the input joint B0 held, as shown in Fig. 16(b). Thus, for the Stephenson six-bar linkages or the two-DOF seven-bar planar linkage, the singularity happens when the joint E, B and F are collinear or the joints D, C and C0 are collinear. Therefore, the three-DOF eight-bar linkage is at the singular positions. For the Type-222 input condition, let the inputs be given through the joints A, B and B0 as an example, as shown in Fig. 17(a). There are also two decomposition ways. The three-DOF eight-bar linkage can be decomposed into a Watt planar sixbar linkage with the input joints B and B0 held, as shown in Fig" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0001408_j.ifacol.2020.12.2546-Figure2-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0001408_j.ifacol.2020.12.2546-Figure2-1.png", + "caption": "Fig. 2. Geometry of a 4-mecanum wheeled robot.", + "texts": [ + " After that, the proposed time delay compensation approach is detailed, followed by an analysis allowing to predict if the delay will be totally compensated or not, and to estimate the non compensable delay in the end of the mission. Next, simulation results are presented for different possible scenarios to show how effective is the proposed approach. Finally, a conclusion and some perspectives are given to conclude the paper. Let\u2019s assume that the robot is placed on a plane surface where ((O,\u2212\u2192x ,\u2212\u2192y ) is the inertial reference frame and (G,\u2212\u2192xR, \u2212\u2192yR) is a local coordinate frame fixed on the robot at its center of mass and geometric center G, (see Fig. 2). The following assumptions are made to consider the robot dynamic model: \u2022 Disturbances are neglected due to the robot evolving environment (low velocity, no slippage, no slops, ...). \u2022 Three measurements are available: x and y positions provided by a positioning system, and the rotation angle \u03b8 returned by a gyroscope. \u2022 Obstacles are detected using laser distance sensors, allowing to know the distance between the robot and the obstacle from each side. \u2022 Measurement noises are modeled by taking into account the sensors accuracy. The following notations in Table 1 are used throughout this article (see Fig. 2). Dynamic model Neglecting the model uncertainties and frictions and denoting l = lx + ly, the dynamic model is given in (G,\u2212\u2192xR, \u2212\u2192yR) by the following equations: (see Sahoo et al. [2018] for more details) x\u0308R = 1 2mRw (\u03c41 + \u03c42 + \u03c43 + \u03c44) y\u0308R = 1 2mRw (\u03c41 \u2212 \u03c42 + \u03c43 \u2212 \u03c44) \u03b8\u0308 = l 2IzRw (\u03c41 \u2212 \u03c42 \u2212 \u03c43 + \u03c44) (1) This model can be expressed in the inertial reference frame (O,\u2212\u2192x ,\u2212\u2192y ) using the following transformation matrix: (Vlantis et al. [2016]) x\u0307 y\u0307 \u03b8\u0307 = R(\u03b8) x\u0307R y\u0307R \u03b8\u0307 , R(\u03b8) = [ cos\u03b8 \u2212sin\u03b8 0 sin\u03b8 cos\u03b8 0 0 0 1 ] (2) Continuous-state space representation (CSSR) Using (1) and (2), the robot CSSR model is given as follows: { X\u0307 = AX +B\u03b8u Y = CX + w (3) where X = [x, y, \u03b8, x\u0307, y\u0307, \u03b8\u0307]T , u = [\u03c41, \u03c42, \u03c43, \u03c44] T , w denotes the sensor noises, assumed to be uncorrelated Gaussian white noises with known variances linked to the sensors accuracy" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0001701_0954405420949757-Figure12-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0001701_0954405420949757-Figure12-1.png", + "caption": "Figure 12. Integral chamfering tool path.", + "texts": [ + " In order to avoid interference, a constraint is added into tool path solving process of convex side chamfering, which is written as ffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi (r cos u x0C) 2 + (r sin u y0C) 2 + (zj z0C) 2 q . rc \u00f035\u00de Cutter can not approach the region near point P0C after the constraint above is added, a part of segment PCPD near point PC in Figure 10 is chosen to process the region near point P0C in Figure 11. When solving processes for concave side, both sides and convex side chamfering are completed, integral chamfering tool path for concave side, both sides and convex side is obtained, as shown in Figure 12. There are two discontinuities on borders of three phases, cutter moves up and then moves down in those positions to ensure processing correctly. Cutter axis direction calculating Cutter axis is set to be perpendicular to pinion axis Z1, therefore chamfering can be realized on four-axis CNC machine tools. Cutter center is denoted by point E and cutter axis direction angle is denoted by uc, as shown in Figure 13. When the position of point E keeps unchanged, variation of uc has no effect to chamfering shape, but contact position on cutter surface changes with uc, which relates to tool wear" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0003921_coase.2014.6899477-Figure6-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0003921_coase.2014.6899477-Figure6-1.png", + "caption": "Figure 6. Case where batteries in the robot are dead.", + "texts": [ + " Also, by (iii), we can expect that the safety device prevents the robot from providing unexpected large forces to humans (Fig. 4). Additionally, by (iv), we can adjust the detection contact force level according to the requirements for the given task. Furthermore, by (v), if a human is pressed against a wall by the robot locked by the safety device, we can easily rescue the human by moving the robot in a direction opposite to the direction in which the human is pressed (Fig. 5). Finally, by (vi), even if the batteries in the robot are dead, the safety device can act because it requires no power supply (Fig. 6). As shown in the next section, each shaft can be locked in clockwise and counterclockwise directions by using two safety devices for each shaft. III. VELOCITY AND CONTACT FORCE-BASED MECHANICAL SAFETY DEVICE In this section, we explain the structure and mechanism of the velocity and contact force-based safety device (see [16] for more information about the velocity-based detection mechanism and the switch-off and shaft-lock mechanism). A. Structure Fig.7(a) shows an example of robots equipped with the safety devices" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0000021_etfa.2019.8868950-Figure5-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0000021_etfa.2019.8868950-Figure5-1.png", + "caption": "Fig. 5. The Pioneer3Dx mobile base dimensions [26].", + "texts": [ + " Moreover, in order to ensure that the resulting path tracks as precisely as possible the desired algebraic curve, the knowledge of the LP about the GP long term path is reduced, influencing the short-term plan generation, since a wider overview on the tracked plan would result in local path optimization. In order to test the supervisory algorithm with path following and obstacle avoidance we have decided to employ a Pioneer 3DX mobile robot [26] equipped with a SICK LMS200 laser range finder [27] with 10-meter range and scanning angle of 180 degrees, a Raspberry Pi [28] 3 Model B mounting an ARM Cortex-A53 (x4 core) CPU (1.2 GHz) and 1-GB RAM, which takes the role of a processing unit for receiving data and controlling the robot (Figure 4). Note that the physical specifications of the robot (Figure 5) are crucial for modelling the kinematic equations necessary for executing the supervisory algorithm, to suitably adapt ROS Algorithm 1: Proposed algorithm with path following. 1 Input: map, start position, goal position 2 Output: path /* Set-up cell expansion lists */ 3 openList ; // Declare the open list 4 closedList ; // Declare the closed list /* Insert starting cell in the open list */ 5 openList.insert(start) 6 while openList is not empty do /* Pop cell c with lowest f(c) off the openList */ 7 c = openList" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0002861_978-3-030-55889-5-Figure5.25-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0002861_978-3-030-55889-5-Figure5.25-1.png", + "caption": "Fig. 5.25 Stiffness change in a spring", + "texts": [ + "120b) and extending this to any functional J (w) means J (e) = \u2212\u0394k1 w\u2032 c(x1)G \u2032(x1, x) \u2212 \u0394k2 w\u2032 c(x2)G \u2032(x2, x) , (5.121) where G(y, x) is the influence function for J (w) and the G \u2032(xi, x) are the slopes at the two springs. 328 5 Stiffness Changes and Reanalysis Displacement Spring To be complete, we also discuss a change k \u2192 k + \u0394k in a displacement spring, and we let J (w) = w(x) be the deflection at a point x. According to Mohr\u2019s equation we have w(x) = \u222b l 0 M MG EI dy + k \u00b7 w(l) \u00b7 G(x, l) , (5.122) whereMG is the moment of the influence function G(x, y), see Fig. 5.25. To evaluate J (e) = wc(x) \u2212 w(x) = \u2212\u0394k wc(l)G(x, l) (5.123) we need wc(l) and so we apply the formula at the point x = l wc(l) \u2212 w(l) = \u2212\u0394k wc(l)G(l, l) , (5.124) where G is now the influence function for w(l) and we solve for wc(l) = w(l) 1 + \u0394k G(l, l) , (5.125) and so (5.123) becomes J (e) = \u2212\u0394k w(l)G(x, l) 1 + \u0394k G(l, l) (5.126) This formula allows to track the change in any functional J (w), if we know by how much the spring (stiffness k) is compressed (= G(x, l)) by the Dirac delta acting at the distant source point x" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0002759_b978-0-12-804560-2.00011-0-Figure4.6-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0002759_b978-0-12-804560-2.00011-0-Figure4.6-1.png", + "caption": "FIGURE 4.6 Model of a 3D reaction wheel pendulum with a telescopic massless leg. The spherical IP is connected to the foot via a 2-DoF ankle joint. The three RWs are mounted on three orthogonal axes. RW accelerations/torques produce a controlled variation of centroidal angular momentum. This results in a variation of the three reaction moments at the stance foot. The tangential reaction moments mx and my induce displacements of the CoP. The vertical reaction moment mz is used to compensate the gyroscopic moment [64].", + "texts": [ + " As shown in [20], with the help of the nonlinear control theory, it is possible to stabilize the IP to its equilibrium manifold. These models have been embodied recently in the form of self-balancing cubes [36,83,76]. The cubes comprise an assembly with three RWs mounted on mutually orthogonal axes, thus resembling the arrangement used in the three-axis attitude stabilization of spacecraft. In its essence, this system represents a 3D RWP. The idea can be applied in humanoid robotics according to the conceptual representation in Fig. 4.6. The figure depicts the model of a 3D RWP supported by a telescopic massless leg. The spherical IP is connected to the foot via an actuated universal joint. The three RWs are mounted on three orthogonal axes. The RW accelerations/torques produce a controlled variation in the centroidal angular momentum, in accordance with the Euler equation for rigid-body rotation, i.e. I \u03c9\u0307 + \u03c9 \u00d7 I\u03c9 = mRW . (4.37) Here \u03c9 is the angular velocity in the inertia frame, mRW = l\u0307C = I \u03c6\u0308, and I denotes the diagonal inertia matrix of the RW assembly" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0003544_cta.2127-Figure7-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0003544_cta.2127-Figure7-1.png", + "caption": "Figure 7. Flux directions in the direct current machine as a generator.", + "texts": [ + " The series-connected coils 9C\u20133C and 11C\u20131C have 1A current flowing through them from the supply V1. Thus, the DCIB measures the self-inductance of two sets of the six sub-coils in parallel. In this diagram, the field circuit is not included, but it carries a 0.2A DC field current during the measurement. In order to establish the paralleled inductances being measured in Figure 6, it is necessary to first look at the direction of the flux produced by a current flowing over the sub-coils, referring to Figure 7 Using a similar concept to the convention in a generalized machine, the m.m.f. of a coil is along its axis. Therefor, the currents in the sub-coils 9C\u20133C and 11C\u20131C both produce m.m.f. and flux vertically, and they are of course mutually coupled. The physical meaning of the 7-coil and the 5-coil equivalent circuits [1, 2] is best visualized with the help of Figures 8 and 9, which is based on the commutating process on the DC machine over one revolution. The 7-coil and 5-coil equivalent circuits represent the operation of the rotating DC machine during a certain period", + "2A current in the field coil, it was found that changing the armature currents from 1 A to 3A did not vary the measured values of the inductance significantly (i.e., only with a few percent difference), and the difference was considerably close to the inherent error attributed in this method. With the use of a constant field current If = 0.2 A, any sub-coil inductance can be measured under the same flux condition, which would exist when the motor is running. However, if the measurement is under the unsaturated condition (i.e., given a zero field current), the self-inductance of the total armature winding varies with its angular position. Referring to Figure 7, the self-inductance measured between the number 2 and number 10 segments was found to be varying almost sinusoidally between 0.15 and 0.31H while turning the rotor. By contrast, if the field current is set at 0.2A, the measured value of the armature self-inductance remains 0.07\u00b1 0.01H, almost independent of the rotor position. Thus, the self-inductance and mutual inductance of the armature sub-coils can be measured to be constant and independent of its armature current and regardless of the rotor\u2019s position" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0001472_s40032-020-00599-y-Figure1-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0001472_s40032-020-00599-y-Figure1-1.png", + "caption": "Fig. 1 Position geometry of 3-PRS manipulator", + "texts": [ + " But the procedure of multi-position 3-PRS manipulator is based only on constraints of revolute joints but the range of motion of mobile platform of manipulator will be influenced by the spherical constraints of the 3-PRS manipulator which is not studied earlier. Therefore, this study attempts the synthesis procedure of 3-PRS manipulator based on cone angle of spherical constraints on range of rotational motion of spherical joints and can be applied to multi-number of prescribed positions and orientations of the mobile platform of 3-PRS manipulator. The mobile platform, i.e., DleS1S2S3 of 3-PRS manipulator is connected to a fixed base platform by three similar supporting limbs PRj Sj j\u00bc1;2;3 is shown in Fig. 1. The prismatic actuators are actuated along three limbs, i.e., BjAj j\u00bc1;2;3 and the actuated prismatic actuator of each limb is inclined to the moving limb (lj) by an angle aji; where aji j\u00bc1;2;3 and i\u00bc1;2;3;4 etc is the angle between the ith position of the moving limb PRj Sj ! j\u00bc1;2;3 and ith actu- ation of actuator PRj j\u00bc1;2;3 : Let dji ! j\u00bc1;2;3 and i\u00bc1;2;3;4 etc is the position displacement vector which is the distance between Bj j\u00bc1;2;3 and PRj j\u00bc1;2;3 ; where Bjjj\u00bc1;2;3 are the fixed base vertex and PRj j\u00bc1;2;3 are the positions of actu- ated prismatic joint" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0000408_s00707-019-02584-8-Figure4-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0000408_s00707-019-02584-8-Figure4-1.png", + "caption": "Fig. 4 Elastic deformation obtained when using only the first eigenmode in a contact simulation is visualized at 0.5 ms with the elastic deformation magnified 1000-fold. Thanks to the magnified deformation, the elliptic shape of the first eigenmode is made visible. The circles highlight the contact area", + "texts": [ + " Detailed convergence analysis with different projection bases has been performed in several contributions [9,19,34], and thus this analysis is not repeated in full detail in the present contribution. In the example shown in Fig. 3, using 100 eigenmodes, the duration of the impact is in high agreement. With just very few modes, the contact force is slightly over-estimated which is expected since the gears can only deform in directions of modes that do not allow the loaded teeth to bend. Instead, the gear deforms globally, see Fig. 4, which makes it stiffer. The same simulation was also performed with the FE program Abaqus, using linear tet4, linear hex8, and quadratic tet10 elements. As can be seen from Fig. 5, the tet4 mesh results in an overly stiff response. The fact that tet10 elements do not exhibit this behavior is a very important reason to advocate their use instead of tet4 elements. For the next simulation case, the gear bores are modeled as rigid and the rigid-body degrees of freedom of the gears are constrained with exception of rotation around the gear axes" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0003499_muh-1404-6-Figure13-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0003499_muh-1404-6-Figure13-1.png", + "caption": "Figure 13. Total deformation all over the machine tool due to static loads Fx = 1000 N and Fy = 500 N at TCP.", + "texts": [ + " Finally, cutting load values are transferred in a tabulated form to the transient structure module where they are subjected to the TCP through a specified number of time steps so as to obtain the deformation of the TCP during one tooth interval. The static analysis performed on the machine tool when subjected to loads previously mentioned shows that the x deformation at the TCP relative to the worktable is equal to 30.7 \u00b5m, while the y deformation is equal to 7.1 \u00b5m. Generally, in the x direction (YZ plane), the static loop stiffness is found to be equal to 32.5 N/\u00b5m, while in the y direction (XZ plane), it is found to be equal to 70.4 N/\u00b5m, and therefore the static loop stiffness in the y direction is much better. Figure 13 shows the deformation all over the machine tool structure. Modal analysis has been performed on two cases: first, without including the bolt\u2019s prestress effect, and second, when the bolt\u2019s prestress effect is included. Figure 14 shows the first six mode shapes of the machine structure in the first case, while Table 2 represents the natural frequencies and the positions of maximum deformation at each mode. From the results obtained in both cases, the first natural frequency is 56 HZ in the first case, while it is 48 HZ in the second; therefore, the dynamic performance when putting the bolts before stress into account is lower, yet more realistic" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0000942_ilt-01-2020-0030-Figure4-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0000942_ilt-01-2020-0030-Figure4-1.png", + "caption": "Figure 4 Geometrical relationship of the race centers", + "texts": [ + " When the outer ring is fixed, under the combined action of contact force, centrifugal force and friction force, the equilibrium equations of mechanics of inner ring are given by the following equations: Fa \u00bc XZ j\u00bc1 Qci j\u00f0 \u00desinai j\u00f0 \u00de Ffi j\u00f0 \u00decosai j\u00f0 \u00de (21) Fr \u00bc XZ j\u00bc1 Qci j\u00f0 \u00decosai j\u00f0 \u00de Ffi j\u00f0 \u00desinai j\u00f0 \u00de cosc j (22) M \u00bc XZ j\u00bc1 Qci j\u00f0 \u00desinai j\u00f0 \u00de Ffi j\u00f0 \u00decosai j\u00f0 \u00de Rri 1Qci j\u00f0 \u00de D 2 cosc j (23) Fatigue life analysis Yue Liu Industrial Lubrication and Tribology The curvature center of the inner groove, outer groove and the ball center areOii,Oei andOj, respectively, after deformation, as O 0 ij,O 0 ej,O 0 j . Figure 4 shows the geometric relationship between before and after deformation. Therefore, there is a mathematical expression that related axial displacement d a, radial displacement d r and angular displacement u to the deformation between elements and inner, outer race d i(j), d e(j). Furthermore, Ri is the distance between the center of inner race and central axis of bearing. The following relationship exists between normal load and deformation: Qci j\u00f0 \u00de \u00bc cPd i j\u00f0 \u00de (24) Qce j\u00f0 \u00de \u00bc cPd e j\u00f0 \u00de (25) For the multi-bearing shafting system, the method provided by Yue et al" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0000058_icmae.2019.8881011-Figure2-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0000058_icmae.2019.8881011-Figure2-1.png", + "caption": "Figure 2. Geometry of the base and the mobile platform.", + "texts": [ + "00 \u00a92019 IEEE of the hexagon, the six vertices are located as shown in Table I, considering that the lengths of the major and minor edges are denoted as and , respectively; and thus, forming a vector of positions denoted as whose in the 6 points is at zero. The vertices of the mobile platform are located in the points arranged in table 2, which are rotated with respect to the reference hexagon since the HUNT type configuration specifies the union of the joints 1-2, 3-4, 5-6, at the lower edge of the platform geometry, as shown in figure 2. Since the mobile platform is subject to translation and rotation movements established by the desired position and orientation, the multiplication of the rotation matrices must be performed at , , , the sum of the desired positions in X, Y, Z, according to equations (1) and (2). (4) (5) where are the points of the mobile platform obtained from Table II. B. Inverse Kinematics In the previous section it found expressions for the location of the fixed platform in , and for the mobile platform in in terms of the dimensions of the edges that make up the hexagon, as well as the desired position and orientation", + " From the above, it obtain vectorially the expression that defines from the triangle OPL that, being a non-rectangle triangle, its sides are related by the cosine law together with the angle and mathematically described in the equation (3), (1) (1) (3) and given that the value of is unknown, we apply the definition of the product point as it is shown in the equation (4), to finally obtain the equation (5) that mathematically expresses . Where, y . In addition, it noted that the coordinate system for each arm rotates with angles described in (6) and defined in figure 2. To obtain the vector , it must be considered that the vertices 1 - 6 are in different quadrants with respect to points 2 to 5. Then for 1 - 6 the value of is described in equation (7) and for the others in equation (8). Then the magnitude of is redefined by the Pythagorean theorem in equation (9). Having two equations for in (5) and (9), equal terms are used to find the equation of each arm, taking the form of equation (10), according to figure 4. Where vary depending on the point to work so equations (11) are defined for points 1 and 6 and equation (12) for points 2 to 5" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0003618_ijvsmt.2015.067521-Figure5-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0003618_ijvsmt.2015.067521-Figure5-1.png", + "caption": "Figure 5 (a) Rotation angle (\u03a8) of the vector u around the king-pin (b) Toe angle (\u03b5) (see online version for colours)", + "texts": [ + " Figure 3 shows the fixed frame Ox0y0z0 and the moving frame Ox1y1z1, which is attached to the car body. Owing to the fact that there are no irregularities on the flat road, the wheel vertical motion with respect to the moving frame is only a mathematical function of the body roll angle, which is represented by the angle .\u03c6 The actuators 1 and 2, and the bars 1 and 2 are in the Oy0z0 plane. The goal of the kinematic position analysis is to obtain the mathematical transformation between the end effector location, defined by the vector [\u0398,\u0393,\u03a8]T\u03c7 = [Figure 4(a) and Figure 5(a)] and the displacements provided by the actuators, defined by the vector 1 2 3[ , , ]TS s s s= [Figure 4(a) and Figure 4(b)]. Both vectors are defined in the moving frame. Regarding to the limb 1 [Figure 4(a)], the centres of the spherical joint (point D1) and the universal joint (point E1) that connect the actuator 1 are given by the equation (3): 11 1 1 1[0, cos\u0398, sen\u0398] 0, , TT ED E z= \u2212 \u2212 = \u2212\u23a1 \u23a4\u23a3 \u23a6 (3) where 1 1| 0 | .D= \u2212 In Figure 4(a), if the angle is null, the distance between points D1 and E1 is 1 ", + " 1 [0, cos\u0398, sin\u0398]TA = (6) With respect to the limb 2, the origin of the displacement of the actuator 2, C2, is defined by equation (7) [Figure 4(a)]: 2 [0,0, ]TC h= (7) The centre of the revolute joint that connects the bar 2 to the actuator 2 (point B2) and the centre of the spherical joint that connects the bar 2 to the end-effector (point A2) are obtained by equation (8) [Figure 4(a) and Figure 4(b)]: [ ] [ ]2 2 20, , 0, cos\u0398 cos\u0393, sin\u0398 sin\u0393 TTB s h A h h= = + + (8) Regarding to the limb 3, the origin of the displacement provided by the actuator 3, C3, is defined by equation (9) [Figure 4(b)]: 3 [ ,0, ]TC d h= (9) The point representing the centre of the universal joint that connects the bar 3 to the actuator 3 (point B3) and the point that represents the centre of the spherical joint that connects the bar 3 to the end-effector (point A3) are obtained by equation (10) [Figure 4(b) and Figure 5(a): [ ]3 3 3 2, , TB d s h A A u\u2032= = + (10) The vector u\u2032 is calculated by applying the equation (11) which comes from the quaternion algebra (Yang and Freudenstein, 1964). ( ) ( )2\u02c6 \u02c6 \u02c6sin(\u03a8) 2sin (\u03a8 / 2)u u p u p p u\u2032 \u23a1 \u23a4= + \u00d7 + \u00d7 \u00d7\u23a3 \u23a6 (11) where the vector u is defined by equation (12), \u03a8 is the rotation angle of the vector u around the king-pin [ 1 2A A in the Figure 5(a)] and d is the distance between points A3 and A2. 1u dl= (12) The unit vector \u02c6 ,p defined by equation (13), has the direction of the king-pin. ( )2 1 2 1 \u02c6 [0,cos\u0393,sin\u0393]TA A p A A \u2212 = = \u2212 (13) In addition, the coupling between the location of the end-effector and the displacements of the actuators 2 and 3 is obtained by the equation (14): ( ) ( ) 2 ( 1, 2,3)T i i i iA B A B i\u2212 \u2212 = = (14) Developing the equation (14) for the bars 2 and 3, we obtain the equations (15) and (16), respectively: 2 2 2 2 22( cos\u0398 cos \u0393) 2 cos(\u0398 \u0393) sin\u0398 sin \u0393 0s h s h h h h\u23a1 \u23a4\u2212 + + + \u2212 \u2212 \u2212 =\u23a3 \u23a6 (15) 2 2 2 3 3 2 2 2 2 2( cos\u0398 cos \u0393 sin \u0393 sin\u03a8) (1 cos\u03a8) ( sin\u0398 sin \u0393 cos \u0393 sin\u03a8 ) ( cos\u0398 cos \u0393) ( sin \u0393 sin\u03a8) 2 sin \u0393sen\u03a8( cos\u0398 cos \u0393) 0 s h d s d h d h l h d d h \u2212 + + + \u2212 + + \u2212 \u2212 \u2212 + + + + = (16) The equations (5), (15) and (16) are second-degree polynomial, independent and decoupled in the variables s1, s2 and s3", + " These equations represent the mathematical transformation between the end-effector location and the displacements provided by the actuators. The angles \u03a8 and \u0393, which define the location of the end-effector in the moving frame can be related to the angles \u03b3 and \u03b5, which define the camber and toe of the wheel in the fixed frame, respectively. The relationship between the angles \u03b3 and \u0393 is given by the equation (17) (see Figure 6) while the relationship between the angles \u03b5 and \u03a8 is given by the equation 18 (see Figure 5). The angles in the both equations are defined applying the right-hand rule. \u0393 2 \u03b3 = \u2212 + \u03c0 \u03c6 (17) ( )arctan y x\u03b5 w w= (18) ( ) ( )0 0 1 0 1 0 \u02c6 \u02c6\u02c6 \u02c6 \u02c6x y uw Ru i w Ru j u d \u2032 \u2032 \u2032 \u2032= \u22c5 = \u22c5 = (19) where 0 1 R is the rotation matrix of the moving frame with respect to the fixed frame. Figure 7(a) shows the free-body diagram of the wheel-tyre assembled to wheel carrier. In order to simplify the force actuators analysis, the following hypotheses are assumed: \u2022 only the mass of the wheel-tyre set is taken into account, hence, the other parts masses are neglected \u2022 the product of inertia of the wheel-tyre set is neglected \u2022 the longitudinal acceleration is not considered \u2022 the wheel vertical motion is only a mathematical function of the body roll angle (it is assumed there are no irregularities on the flat road) \u2022 the vertical and angular accelerations of the wheel-tyre set are neglected \u2022 the rolling resistance at the tyre is neglected" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0001588_aim43001.2020.9158991-Figure4-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0001588_aim43001.2020.9158991-Figure4-1.png", + "caption": "Figure 4. Wire routing schematics", + "texts": [ + " Based on previous experiences, the hole size was determined to be close to the wire diameter. Otherwise, it might cause twist deformation in the continuum part. B. Pretension mechanism design The pretension mechanism (PtM) located in the middle part of the robot. Driving wires passes through the PtM device. In this robot, we drive eight wires, four wires for each section. A linear actuating unit drives wires; in this robot, one motor drives two paired wires with a strain and ease manner, Total four motors actuate eight wires (Fig 4). We modified the previous design of the passive pre-tension mechanism [22], with equipping linear potentiometers (Fig.6) The new PtM device consists of PtM holder and center holder, where all components are fixed, and PtM slider, which slides along the guide shaft and moves linear potentiometer. The compressional springs are equipped along the slider pass. The PtM yields wire-tension and provides tension information that is obtained by measuring the compression length of springs by the linear potentiometers", + " The virtual work principle gives an identity; \u0394\ud835\udc40\ud835\udc54\ud835\udc47\u2206\ud835\udc5d\ud835\udc5b = \ud835\udc47\ud835\udc5a \ud835\udc47 \u2206\ud835\udf19 (2) where \u0394M is a perturbed mass loaded on the end-point and \ud835\udc54 = (0 0 \ud835\udc54)\ud835\udc47 is the gravitation acceleration vector. If a Jacobian that stipulates the infinitesimal displacement of the end-point and the infinitesimal rotation of the motors; \ud835\udc3d \u225c \ud835\udf15\ud835\udc5d\ud835\udc5b \ud835\udf15\ud835\udf19 \u2208 \u211c3\u00d74 is available, Eq.(2) becomes, \u0394\ud835\udc40\ud835\udc54T\ud835\udc3d\u2206\ud835\udf19 = \ud835\udc47\ud835\udc5a \ud835\udc47 \u2206\ud835\udf19 (3) It provides an identity on two vectors, \u0394\ud835\udc40\ud835\udc3dT\ud835\udc54 = \ud835\udc47\ud835\udc5a \u2208 \u211c4 (4) Torque loaded about motor axis is related to the wire tension such as, \ud835\udc47\u03c3 = \ud835\udf06 2\ud835\udf0b (\ud835\udc53\ud835\udf0e \u2212 \ud835\udc53?\u0302?), (\ud835\udf0e = \ud835\udc4e, \ud835\udc4f, \ud835\udc50, \ud835\udc51), (\u03c3\u0302 = ?\u0302?, ?\u0302?, ?\u0302?, ?\u0302?) (5) Where, \u03bb is a lead of the screw (see Fig.4). Substituting (1) into (5), we have \ud835\udc47\u03c3 = \ud835\udc58\ud835\udc5d \ud835\udf06 4\ud835\udf0b (\ud835\udc62\ud835\udf0e \u2212 \ud835\udc62?\u0302?), (6) Hence \ud835\udc7b\ud835\udc5a in (4) is obtained by, 462 Authorized licensed use limited to: University of Wollongong. Downloaded on August 11,2020 at 10:23:43 UTC from IEEE Xplore. Restrictions apply. \ud835\udc7b\ud835\udc5a = \ud835\udc58\ud835\udc5d \ud835\udf06 4\ud835\udf0b \u2206\ud835\udc96 (7) where \u2206\ud835\udc96 = (\ud835\udc62\ud835\udc4e \u2212 \ud835\udc62\ud835\udc4e \ud835\udc62\ud835\udc4f \u2212 \ud835\udc62?\u0302? \ud835\udc62\ud835\udc50 \u2212 \ud835\udc62?\u0302? \ud835\udc62\ud835\udc51 \u2212 \ud835\udc62?\u0302?)\ud835\udc47 Substitution of (7) into (4) gives \u0394\ud835\udc40\ud835\udc3dT\ud835\udc54 = \ud835\udc58\ud835\udc5d\ud835\udf06 4\ud835\udf0b \u2206\ud835\udc96 (8) Since \ud835\udc88 = (0 0 \ud835\udc54)\ud835\udc7b and \ud835\udc71 = \ud835\udf15\ud835\udc91\ud835\udc5b \ud835\udf15\ud835\udf53 = ( \ud835\udf15\ud835\udc5d\ud835\udc5b\ud835\udc65 \ud835\udf15\ud835\udf53 \ud835\udf15\ud835\udc5d\ud835\udc5b\ud835\udc66 \ud835\udf15\ud835\udf53 \ud835\udf15\ud835\udc5d\ud835\udc5b\ud835\udc67 \ud835\udf15\ud835\udf53 ) \ud835\udc7b , \ud835\udc3d\ud835\udc47\ud835\udc54 = \ud835\udc54 ( \ud835\udf15\ud835\udc5d\ud835\udc5b\ud835\udc67 \ud835\udf15\ud835\udf53 ) \ud835\udc47 = \ud835\udc54 ( \ud835\udf15\ud835\udc5d\ud835\udc5b\ud835\udc67 \ud835\udf15\ud835\udf19\ud835\udc4e \ud835\udf15\ud835\udc5d\ud835\udc5b\ud835\udc67 \ud835\udf15\ud835\udf19\ud835\udc4f \ud835\udf15\ud835\udc5d\ud835\udc5b\ud835\udc67 \ud835\udf15\ud835\udf19\ud835\udc50 \ud835\udf15\ud835\udc5d\ud835\udc5b\ud835\udc67 \ud835\udf15\ud835\udf19\ud835\udc51 ) \ud835\udc47 ", + " If a vertical deviation of the end-point \u2206\ud835\udc5d\ud835\udc5b\ud835\udc67 is obtainable by using a gyro-sensor, e.g., attached at the end-point, Eq.(12) directly determines the rotation angle of motor- \ud835\udf0e. However, it is sometimes harmful to put an electronic device at an end-point in some severe environments. Thus, we approach to control the end-point motion with a simple and empirical C-driving only by using tension information as follows. The control diagram is shown in Fig.13. Based on robot actuating design, one motor drives a pair of wires (see Fig.4) that go out the actuating unit and enter the PtM device. Inside the PtM, each wire pushes the corresponding spring (see Fig.10) accompanied by sliding the linear potentiometer of which voltage (tension information) goes to the Arduino. Fig. 14 demonstrates slider arrangements, and we noted paired wires as A1, A2 ( wires a, \u0203) and C1, C2 (c, \u0109) for the first section and B1, B2 (b,b\u0302) and D1, D2 (d,d\u0302) for the second section (Fig.1). Based on the wire tension level determined by the data obtained with the linear potentiometer, we use them to predict a vertical deviation of the end-point and compensate it Driving motor angles are defined by the obtained value from the tension sensors" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0003381_icdcs.2015.46-Figure4-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0003381_icdcs.2015.46-Figure4-1.png", + "caption": "Fig. 4. Illustration of rotating time of a camera sensor", + "texts": [ + " Camera sensors are deployed above region R to monitor those selected subareas. So, a 3-D space is considered. As mentioned above, each sub-area is covered entirely by one camera sensor, thus we just need to assign one camera sensor to cover one subarea. As camera sensors can pan and tilt, the total rotation time of camera sensor rotating from sub-area to another includes panning time and tilting time. To provide a balanced angle of view, we assumed that the axis of camera Ci\u2019s FoV, i.e.,\u2212\u2212\u2212\u2192 Hi 1Op in Fig. 4, always goes through the center of sub-area Rp. We denote panning angle velocity by \u03c9h and tilting angle velocity \u03c9v . As shown in Fig. 4, when camera Ci changes its FoV from sub-area Rp to sub-area Rq , it first pans and then tilts, i.e., the axis of camera\u2019s FoV first pans from \u2212\u2212\u2212\u2192 Hi 1Op to \u2212\u2212\u2212\u2212\u2192 Hi 1O i pq by an angle of \u03b2ihpq , and then tilts from \u2212\u2212\u2212\u2212\u2192 Hi 1O i pq to\u2212\u2212\u2212\u2192 Hi 1Oq by an angle of \u03b2ivpq . As a result, the rotating time of camera sensor Ci from Rp to Rq t i pq equals to: tipq = \u03b2ihpq/\u03c9h + \u03b2ivpq/\u03c9v. In Fig. 4, \u2212\u2212\u2212\u2192 Hi 1H i 2 is perpendicular to the plane region R and plane R is the base plane. We assume that the locations of the center of sub-areas are known to cameras. This assumption is consistent with the prior works, where sub-areas\u2019 locations could be determined by using a complementary mechanism such as using a supporting network of low-resolution wide-angle cameras[19], or camera calibration[2]. Given the coordinate(xp, yp, 0), (xq, yq, 0) of the center of sub-area Rp, Rq respectively, and the coordinate (xCi , yCi , zCi) of camera sensor Ci, we can figure out \u03b2ihpq and \u03b2ivpq " + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0002989_978-3-642-16214-5_9-Figure9.28-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0002989_978-3-642-16214-5_9-Figure9.28-1.png", + "caption": "Fig. 9.28. Multi-plate synchronizer", + "texts": [ + " This axial force, which presses the synchronizer plate towards the direction of the gearwheel to be shifted, speeds up the shifting process, shortens the time needed for synchronizing and reduces the force that must be applied by the driver to the gearshift lever. There is thus an increase in force at the friction surface between the synchronizer body 1 and the synchronizer ring 2. As soon as the speeds of the sliding sleeve 8 and the gearwheel 6 are synchronized, the selector teeth 9 of the sliding sleeve 8 are positively engaged with the selector teeth 9 of the gearwheel 6 [9.6]. The multi-plate synchronizer in its present form has been developed from the multi-plate clutches used in powershift transmissions (Figure 9.28). Because of its large power transmission surface AR, it is suitable wherever there is a requirement for very high synchronizer performance. The cone angle \u03b1 of a multi-plate synchronizer is 90\u00b0. To be operated with the same gearshift effort as a single-cone synchronizer with \u03b1 = 6.4\u00b0, according to Equation 9.2 a multi-plate synchronizer of the same effective diameter must have j = 9 friction surfaces. The lengths of the two synchronizers are then roughly equal. Multi-plate synchronizers are complex and costly" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0001057_0021998320920920-Figure5-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0001057_0021998320920920-Figure5-1.png", + "caption": "Figure 5. Contours of maximum Hashin failure index, Hmax, for binder in unit-cell 2 with warps and wefts shown for context. (a) Most severe stresses where binders transition. (b) Second most severe stresses where binders are near a weft.", + "texts": [ + " Additionally, the stress states at an applied 1% volume average \"xx are used, which can result in a failure index >1. The failure indices are used for relative comparison, and if the expected point at which initial damage occurs is needed, the stresses can be linearly scaled with the applied load, since the elastic response is considered. The results herein describe the stresses as predicted by a linear elastic analysis with the aim to understand which features of the tow architecture have a strong influence on creating stress concentrations that can lead to damage. Figure 5 shows contours of the maximum index of Hashin\u2019s criteria, Hmax, for the binder in unit-cell 2 (refer to Figure 2 for the unit-cell enumeration). The maximum failure index provides a metric of severity of the stress state. The regions of severe stresses tended to form near two types of features of the tow architecture. First, the most severe stresses in the binder developed where the binder transitions between travelling along the x-axis and travelling through the thickness, such as label A in Figure 5. In these regions where the tow path of the binder transitions and the most severe stresses occur, transverse tension is the mode corresponding to the maximum failure index, as shown in Figure 6. The high failure index for transverse tension is due to a severe 0xz that occurs. Where the binders travel along the x-axis, the binders carry a significant amount of load, since the fibers align with the applied load. This can be illustrated by contours of the stress in the direction of the applied load, xx, which is in the global coordinate system, as shown in Figure 7", + " Figure 8 shows the 0xz concentrations that result in the binders in these regions, and the location the 0xz concentrations in the binders match the locations where xx in the binder dropped significantly. It should also be noted that the most severe 0xz occurs within the volume of the binder, not on the surface, as seen in the cutaway of the binder in Figure 8. The high HTT in these regions are dominated by the large magnitude of 0xz. The regions with the second most severe stresses in the binders occur where z-aligned binders come close to wefts, such as indicated by label B in Figure 5. For all of these regions, the failure index corresponding to transverse tension, HTT, is highest, as shown in Figure 6. The stress component that contributes the most to HTT where the binders traverse the textile thickness is 0zz. Figure 9 shows 0zz in the binder, and the locations of 0zz concentrations can be seen to match the locations of local maxima of Hmax, shown in Figure 5. When the binders are most closely aligned with the z-axis, the local z-axis of the binders most closely aligns with the global x-axis, which is the direction of load. Consequently, the 0zz concentrations form where load is transferred from the binders to nearby wefts. The transfer of load is clearly shown by contours of xx in Figure 7, such as at label B in the figure. The severity of 0zz appears inversely proportional to the distance between a binder and the weft, as expected. It should be also noted that the binders seem unaffected by the cross-sectional shape of the neighboring wefts. For the warps and wefts, multiple tows lie within a single unit-cell, which can provide some insight into how much variation occurs between tows of the same type for the textile model used in this study. However, since there is only one wavelength of a single binder within a unit-cell, Figure 10 shows contours of the maximum index of Hashin\u2019s criteria for the binder in unitcell 4, which can be compared to Figure 5 that shows the same results for unit-cell 2. The locations of high Hmax matched well between the two unit-cells. However, the highest concentrations were more severe within the volume of the binder in unit-cell 2 compared to unit-cell 4, and the cross-sectionally averaged Hmax in the regions of concentrations in the binders differed by up to 16% across all four unit-cells shown in Figure 2(a). Figure 11 shows contours of the maximum index of Hashin\u2019s criteria, Hmax, for the wefts in unit-cell 2" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0002845_ijcaet.2019.102500-Figure5-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0002845_ijcaet.2019.102500-Figure5-1.png", + "caption": "Figure 5 Target situations, (a) target classes (b) target neural network architecture", + "texts": [ + " These latter are modelled from the robot\u2019s possible movements [the steering angle is bounded by (\u201318\u00ba, 18\u00ba)] and from the distance to the nearest obstacle in each area of the divided environment (Figure 4): right area \u2013 ROD \u2013 (in the direction of \u201318\u00ba), front area \u2013 FOD \u2013 (in the direction of 0\u00ba) and left area \u2013 LOD \u2013 (in the direction of 18\u00ba) [Figure 3(a)] Azouaoui et al. (2011). The obstacles classes correspond \u2013 from the human being perception \u2013 to corridors, right turns, left turns, the possibility to go left or right only, left or forward only and right or forward only (for example, enter a room or continue onwards). The target class recognition is based on the neural network shown in Figure 5(b). Trained by fuzzy Artmap method, the network reckons the right goal category from five situations [Figure 5(a)] built using the temperature field process (Sorouchyari, 1989). This process, utilises the robot\u2019s heading angle and the robot-target distance to compute the temperature in three environment zones: right (TR), front (TF) and left (TL) (inputs for the network) as follows (where \u03b3 is the target angle) (Azouaoui et al., 2008; Chohra et al., 1998): If 45 81 (Class T1), Then TR=12sin( ), TF 6cos( ), TL 6cos( ), If 81 99 (Class T2), Then TR=6 cos( ) , TF 12sin( ), TL 6 cos( ) , If 99 135 (Class T3), Then TR=6 cos( ) , TF 6 sin( ) , TL 12sin( ), If 135 < 27 < = = < = = < = = \u03b3 \u03b3 \u03b3 \u03b3 \u03b3 \u03b3 \u03b3 \u03b3 \u03b3 \u03b3 \u03b3 \u03b3 \u03b3 0 (Class T4), Then TR = 12 sin( ) , TF=6 sin( ) , TL 12 sin( ) , If 270 < 405 (Class T5) If <315 , Then TR = 12 sin( ) , TF 6 sin( ) , TL 6cos( ), If 315 < 360 , Then TR=12cos( ), TF=6sin( ), TL 6 sin( ) , If 360 , Then TR=12 = = = = \u03b3 \u03b3 \u03b3 \u03b3 \u03b3 \u03b3 \u03b3 \u03b3 \u03b3 \u03b3 \u03b3 \u03b3 \u03b3 cos( ), TF 6cos( ), TL 6sin( )" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0003883_pomr-2014-0016-Figure2-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0003883_pomr-2014-0016-Figure2-1.png", + "caption": "Fig. 2. Launching of the Lifesaving Unit [4]", + "texts": [ + " One of these systems, shown in Fig. 1 and described in detail in [3], provides opportunities for fast and safe evacuation of people using only gravitational forces. The rescued people get simultaneously into all lifeboats situated on different decks, after which a system is started to lower the boats at a controlled speed. When the boats reach the level of the ship\u2019s slip, they are hooked in, one by one, and freely slide down the ramp to the water. In the second system, developed by Fassmer and schematically shown in Fig. 2, the watercraft units are situated in rows on the lower deck. People get into all boats at the same time, and then using the already accumulated energy, or that delivered from an additional electric power generator, the boats are rolled to the ramp from which they slide down to the water. The ramp has a three-segment structure, but when the boats with people slide down to water all three segments are blocked in a one-plane position. The last ramp segment has an air tank connected to its bottom side, to allow the ramp to adapt its inclination angle to the water level overboard, which changes mainly to ship rolling or pitching motions and sea waving", + " The value of the force is defined by the linear function of the crossing distance of the basic (undeformed) ramp surface by the bearing node, see Fig. 3 and Fig. 4. This linear function has been limited by the maximal pressure reaction, introduced to limit modelling of the bearing pressure forces to the range observed in real conditions (for instance due to the loss of stability of the structure, or exceeding the yield point). The pressure force acting on the ramp: (6) where: Espr \u2013 modulus of elasticity [N/m], \u0394s \u2013 distance between the bearing node and the base surface, see Fig. 2, n \u2013 unit normal vector Unauthenticated Download Date | 3/30/18 11:31 AM 37POLISH MARITIME RESEARCH, No 2/2014 The rolling friction force is a linear function of the pressure force. (7) where: \u03bc \u2013 rolling friction coefficient, eT \u2013 unit tangential vector The unit tangential vector is defined as: (8) The damping force was introduced to the model to take into account the kinetic energy loss caused by the deformation of constructional elements of the module during collision. It was assumed in the presented model that the damping force is proportional to the normal velocity during collision: RDAMP = \u2013 cDAMPVN (9) where: cDAMP \u2013 damping coefficient, VN \u2013 normal velocity vector" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0000316_icems.2019.8921951-Figure9-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0000316_icems.2019.8921951-Figure9-1.png", + "caption": "Fig. 9 Second-order natural frequency of rotor system", + "texts": [ + "3 times greater than the first-order critical speed and less than 0.7 times the secondorder critical speed [7]. Modal analysis and calculation of natural frequency are the key steps to determine the critical speed of the rotor system. The calculation of the frequency and the analysis of the vibration mode can make the motor avoid the resonance phenomenon. The finite element method is used to analyze the modality of the rotor system, and the first four natural frequencies and vibration modes of the rotor system are obtained as shown in Fig. 8 to Fig. 9. From Fig. 8 to Fig. 11, the natural frequencies of the various stages of the rotor are known, and the critical speeds corresponding to the various circles are calculated as shown in Table \u2163. It can be seen from Table 4 that the first-order critical speed of the rotor of the motor is much larger than the rated working speed of the rotor, and the motor does not resonate during normal operation, which proves that the rotor design is reasonable and can ensure safe and reliable operation. In the design of high-speed motors, the rotor's operating speed needs to avoid the first-order critical speed" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0003532_icl.2015.7318107-Figure2-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0003532_icl.2015.7318107-Figure2-1.png", + "caption": "Fig. 2. Sample macine models used for the CAD team project", + "texts": [ + " Each team member is assigned to create four to five components/parts of a machine assigned to the team. All team members are expected to measure the geometric dimensions of the parts/components of the machine, create hand sketches (with orthographic projection views) of the components, and create 3D solid models and document drawings. Sample machines used for this team CAD project in the past years were clausing lathe, band and grab saw, trak k3 mill, bicycle, floor scrubber, fork lift, drill press, single and double cylinders as well as other machines. Figure 2 shows some of the complete assembly models created by CAD teams within about four weeks of work. In order for the design team/group to efficiently work together, depending on the assigned machine, each team selects two team leaders, called captains, and four to six small Proceedings of 2015 International Conference on Interactive Collaborative Learning (ICL) 978-1-4799-8707-8/15/$31.00 \u00a92015 IEEE 20-24 September 2015, Florence, Italy Page 675 group leaders, called lieutenants. The role of the small group leaders (i" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0000327_j.mechmachtheory.2019.103729-Figure7-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0000327_j.mechmachtheory.2019.103729-Figure7-1.png", + "caption": "Fig. 7. Kinematic diagram of the outer loop in the second configuration.", + "texts": [ + " C 2 = \u23a1 \u23a2 \u23a2 \u23a2 \u23a2 \u23a2 \u23a2 \u23a2 \u23a2 \u23a2 \u23a2 \u23a2 \u23a3 2 0 3 E 0 4 0 9 E 0 2 \u2032 E 1 3 \u2032 E 1 4 \u2032 0 9 \u2032 S 1 E 0 0 E 0 0 0 E 1 E 1 S 1 0 E 0 0 E 0 0 0 E 1 S 1 E 0 0 E 0 0 0 0 0 S 1 + R 1 E 1 0 0 0 0 E 0 0 S 0 + R 0 E 1 E 1 0 0 E 0 0 E 0 S 0 0 E 1 E 1 0 0 E 0 0 S 0 + R 0 S 1 S 1 S 1 S 1 + R 1 S 0 + R 0 S 0 S 0 + R 0 0 \u23a4 \u23a5 \u23a5 \u23a5 \u23a5 \u23a5 \u23a5 \u23a5 \u23a5 \u23a5 \u23a5 \u23a5 \u23a6 (7) A mechanism diagram of the unit Rubik\u2019s Cube can be drawn by combining the above mentioned sub-adjacency matrix with the actual azimuth feature, then after removing the common constraint spherical pair, a mechanism diagram of the outer loop as shown in Fig. 7 is obtained. It can be seen from the mechanism diagram of the Rubik\u2019s Cube mechanism that each component of that (namely, the sub-piece of Rubik\u2019s Cube) is connected to a plurality of components, which is a typical multi-loop coupled mechanism [42] . But each of its components is a coupling node, and the connection relationship is very complicated. In order to solve this problem, this paper firstly proposes a hierarchical directed graph to realize the successive layered decomposition of the Rubik\u2019s Cube mechanism, analyzes gradually and to determine the constraint problem of it", + " Its external motion-screw system { S/ O3 } is expressed as: { S/ O3 } = \u23a7 \u23aa \u23aa \u23aa \u23aa \u23aa \u23aa \u23aa \u23aa \u23a8 \u23aa \u23aa \u23aa \u23aa \u23aa \u23aa \u23aa \u23aa \u23a9 S/ O3 1 = ( 1 0 0 ; 0 0 0 ) S/ O3 2 = ( 0 1 0 ; 0 0 0 ) S/ O3 3 = ( 0 0 1 ; 0 0 0 ) S/ O3 4 = ( 0 0 0 ; 1 0 0 ) S/ O3 5 = ( 0 0 0 ; 0 1 0 ) S/ O3 6 = ( 0 0 0 ; 0 0 1 ) \u23ab \u23aa \u23aa \u23aa \u23aa \u23aa \u23aa \u23aa \u23aa \u23ac \u23aa \u23aa \u23aa \u23aa \u23aa \u23aa \u23aa \u23aa \u23ad (23) Because of { S/ O3 } \u2283 { S/ I } , the final motion-screw system of the corner piece is the internal common motion-screw system. { S/ 3 } = { S/ O3 } \u2229 { S/ I } = \u23a7 \u23aa \u23a8 \u23aa \u23a9 S/ 2 1 = ( 1 0 0 ; 0 0 0 ) S/ 3 2 = ( 0 1 0 ; 0 0 0 ) S/ 3 3 = ( 0 0 1 ; 0 0 0 ) \u23ab \u23aa \u23ac \u23aa \u23ad (24) Therefore, the final DOF of corner piece is 3, with three rotational DOFs. The mechanism diagram of the outer loop mechanism in the second configuration of the Rubik\u2019s Cube mechanism is shown in Fig. 7 . Comparing with the outer loop mechanism diagram of the first configuration of Fig. 6 , it can be known that the Rubik\u2019s Cube mechanism from the aligned state to non-aligned state has the following changes: New kinematic pairs have been generated, the orientation characteristic of the kinematic pair has been changed and Structural symmetry of unit mechanism has been broken. Therefore, the loops become more complicated. This section takes the second configuration of rotation about the Z axis as an example to analyze the DOF of the cube in the non-aligned state" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0000893_msf.982.75-Figure7-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0000893_msf.982.75-Figure7-1.png", + "caption": "Fig. 7. The rupture of the solidified blank in the FML during the forming (140 and 150 mm): sidewall fracture, and fracture in the cup corner", + "texts": [ + " However, as the cavity pressure was more augmented outside the optimum value, rupture increased because of less flow of material from the flange to the die and is detected as a failure domain. Fig 6. Effect of the Curing Condition. As mentioned before, failure strain in Al alloys is larger while composites have very small formability. This is precisely why that is forming a cured GLARE is not possible, when a cured blank was used for forming all the experiments failed, As it is shown, the achieved forming depth is too small, and the rupture occurs a little after beginning the forming process Fig. 7. The solution-dependent state variables (SDVs) track constitutive quantities of interest at each integration point in the finite element model. The most useful of the MCT (Multicontinuum Theory) state variables is SDV1 which tracks the discrete failure state of the composite material at each integration point in the finite element model. SDV1 signifies the failure state of the GLARE laminate, the color code is signifying, Blue: No damage in either matrix or fiber, Green: Matrix Damage, Red: Fiber and Matrix Damage, the failure starts occurring in just 10 mm of depth much before the ultimate load is reached" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0002282_khpiweek51551.2020.9250158-Figure1-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0002282_khpiweek51551.2020.9250158-Figure1-1.png", + "caption": "Fig. 1. General design scheme of a multi-support spindle unit.", + "texts": [ + " 403 - the spindle unit housing, the cutting tool and the workpiece are considered absolutely rigid; - the loads acting on the spindle shaft consist of the cutting forces, forces on the drive side and reaction forces from the supports. In comparison with them, the gravitational and centrifugal forces are considered negligible and the influence of thermal effects is not taken into account. In the design scheme of a multi-support spindle shaft, each bearing is represented by an elastic concentrated support of rigidity nC and producing reaction nR , Nn ,...2,1= (Fig. 1). In addition to the support reactions, in the bending plane yOz the shaft is also loaded with the cutting force and moment yrF , xrM at location 0=z and the force and moment of the drive yprF , xprM at location przz = . In the general case a similar system of forces can cause bending of the shaft in the orthogonal plane xOz as well. The use of the Timoshenko beam model is justified for multi-bearing shafts because their bending stiffness is comparable to the shear stiffness. According to this theory, the equation of the bent axis of the shaft has the form [18]: dz zdQ GA K EI zM dz yd t )()( 2 2 += , (1) where z \u2013 axial coordinate, )(zy \u2013 shaft deflection function, E \u2013 elasticity modulus, G \u2013 shear modulus, I , A \u2013 the axial moment of inertia and the cross-sectional area of the shaft, respectively, )(zM , )(zQ \u2013 bending moment and shear force in section with coordinate z , respectively, tK \u2013 cross-sectional shape factor", + " In accordance with [17], the solution of equation (1) by the initial parameter method consists of the general solution of the homogeneous equation and particular solutions of the heterogeneous equations for individual segments of the shaft. The general solution contains the deflection 0y and angle of rotation 0\u0398 of the shaft at the origin, whereas particular solutions depend on the external forces forming the right side of the equation and are defined through universal functions )(z\u03a6 , )(* z\u03a6 . For the design scheme shown in Fig. 1, the solution of equation (1) for the deflections )( izy and angles of rotation )( iz\u0398 of the shaft sections in the i -th segment of the shaft can be represented in the form [17]: , )( )( 00 EI z zyzy i ii \u03a6 +\u0398+= (2) EI z z i i )( )( * 0 \u03a6 +\u0398=\u0398 , (3) whereby = \u2212 = \u2264\u2264 i j ji i j j LzL 1 1 1 , (4) and functions )(z\u03a6 and )(* z\u03a6 are defined as follows: [ ],)( )( 6 )( 6 )( 2 )( 62 )( 1 1 1 0 1 1 3 3232 ipryipy t i j i j ji t i j ji j i pry i prpy i px zFzF GA EIK Lz GA EIK Lz R z F z M z F z Mz \u03b4 \u03b4\u03b4 \u2212+ \u2212\u2212 \u2212 + ++\u2212=\u03a6 \u2212 = \u2212 = \u2212 = (5) [ ]" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0002867_ijiei.2018.096579-Figure1-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0002867_ijiei.2018.096579-Figure1-1.png", + "caption": "Figure 1 Typical quadrotor configuration (see online version for colours)", + "texts": [ + " In Section 6, we present the simulation results and we demonstrate the efficiency and the robustness of PD-PSO controller for variable altitudes and in presence of wind disturbance. Finally, in the last section, we give our conclusions. The quadrotors aircrafts are complex flying machines, strongly nonlinear, fully coupled, and their aerodynamic is affected by many physical effects, including gravity, gyroscopic effects, friction and moment of inertia. Therefore, their modelling is a delicate task. From the configuration shown in Figure 1, the quadrotor consists of four rotors situated at the ends of a cross, and the control electronics is situated in the centre of the cross. The two pairs of propellers must rotate in opposite directions to prevent the vehicle\u2019s overturning. In fact, the energy spent to oppose the rotational motion contributes in the pushing force. In this work, the developed model assumes the following working assumptions: \u2022 Rigid and symmetrical quadrotor structure (diagonal inertia matrix). \u2022 Rigid propellers (negligible effect of deformation during rotation)" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0000081_978-3-030-32710-1_10-Figure14-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0000081_978-3-030-32710-1_10-Figure14-1.png", + "caption": "Fig. 14 Organization of simultaneous impact and vibrational effects on the end wall by slotting tools (a) and one time impact (b)", + "texts": [ + " The examples of applications of OD and its new functional capabilities are established below, e.g. [10, \u2022 Example 1. Possibility of control over of physical-mechanical properties, geometrical shape of contact surface and displacement trajectory (Fig. 10). \u2022 Example 2. Travel of long objects in pipe (Fig. 11). \u2022 Example 3. Travel of object in pipe with possibility of vibroprotection and positioning (Fig. 12). \u2022 Example 4. Possibility of hole drilling in the end of pipe wall (Fig. 13). \u2022 Example 5. Possibility of battering in pipe of wall end (Fig. 14). \u2022 Example 6. OD can connect together, forming some novel mobile selfreconfigurable structures (swarm systems) for various applications (Fig. 15). Example 1 Example of control of geometrical shape of contact surface by the OD is given in Fig. 10. In this mode, in the motion of the OM 1 (ABCDEF) (Fig. 4) of the adaptivemobile parallel spatial robot (Fig. 10), each longitudinal displacement of the rear face ABC and frontal face DEF is preceded by alternating discrete rotations in both directions, with specified increment, relative to the direction of motion", + " In tool rotation, coordinated increase in length of the rods at the lateral faces ensures its longitudinal supply with specified force; the generation of impact and vibration effects is possible in combination with tool rotation. In that case, the spatial position, cutting force, and impact and vibration effects are monitored by means of sensors 9 and 10 (Fig. 4) at the radial limiters of the frontal face and sensors 3, 5, and 6 (Fig. 4) at the rods in the lateral faces. The axis of tool rotation may be coaxial with the symmetry axis of OM 1 (Fig. 13a) or eccentric (Fig. 13b). Example 5 Organization of impact and vibration effects by a slotting tool on the end surface (Fig. 14) of a tubular profile by means of OM 1 (Fig. 4). In thismode, the slotting tools are established at each point of the frontal face. (The clamping of the slotting tool\u2019s tailpiece is not shown in Fig. 14). The radial limiters of the rear face are fixed at the internal contact surface as in previous modes. Then, coordinated change in the length of the rods at the lateral faces brings the working sections of the slotting tools into contact with the end surface at the machining site, and machining begins at specified frequency, amplitude, and force. In one time or double action,where necessary, the sequence of actionmay bemodified. In each case, the spatial position, cutting force, and magnitude of the tools\u2019 impact and vibration effects are monitored by means of sensors 9 and 10 (Fig" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0001574_s11668-020-00938-2-Figure2-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0001574_s11668-020-00938-2-Figure2-1.png", + "caption": "Fig. 2 Thermal\u2013solid coupling analysis of driving gear in shearer cutting section", + "texts": [ + " Figure 1 shows the comparison between the gear multi-field coupling simulation results and the experimental results under the conditions of oil distribution, long-term temperature distribution and oil flow. The results show that the simulation results of gear multi-field coupling are close to the experimental results to some extent on the basis of the new meshing method and prediction model. The multi-field coupling simulation process of gear is inseparable from the establishment of the parametric gear model and the selection of reasonable coupling methods. Figure 2 shows the simulation process of a thermal\u2013solid coupling analysis method for a certain type of shearer transmission gear. Through the thermal\u2013solid coupling analysis of the simplified finite element model of the transmission gear of the shearer shearing section, we can obtain the cloud map of the overall temperature distribution of the shearer rocker at ? 45 and 45 . With more demanding gear operating requirements being proposed by industry, the present work reviews the relevant research on the influencing factors of gear multi-field coupling simulation analysis: the gear accurate modeling method and multi-field coupling simulation analysis method" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0001224_s1068798x20050184-Figure1-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0001224_s1068798x20050184-Figure1-1.png", + "caption": "Fig. 1. Dependence of the tangential load (\u0420) on the path length L\u0445 for long (1) and short (2) projections.", + "texts": [ + " Optimizing the surface relief entails modifying the structure. The traditional relief\u2014a set of extended wedge-shaped projections\u2014must be replaced by a set of projections of limited length. In traditional relief, which is observed after cutting, the length of the projections is at least twice their width. For projections of limited length (such as spherical segments, pyramids, and cones), the length and width are comparable. For relative motion of the tool and workpiece in a plastic state, short projections offer benefits in terms of tangential loading (Fig. 1). For long projections, the tangential load depends on their orientation with respect to the direction of motion. It is greatest when the projections are perpendicular to the direction of motion. With increase in path length, it increases to a critical value. This is not the case for short projections, for which the tangential load does not depend on the path length. The metal in the workpiece f lows laterally away from the projections and does not create a dead zone. The maximum normal stress at the frontal surface of spherical microprojections depends only on the plastic constant of the workpiece, the contact-friction coefficient, and the frictional angle" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0001693_052026-Figure2-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0001693_052026-Figure2-1.png", + "caption": "Figure 2. Pinching of tubers vertically, horizontal knives and counter.", + "texts": [ + " The tuber, falling from the loading hopper onto the disk and, when interacting with the contradiction, is evenly distributed on its end surface and is processed. Horizontal and vertical knives exert a force on the product, resulting in its separation into pieces in the shape of a parallelepiped, one of the faces of which corresponds to the profile of the windows formed by the surfaces of the parts of the cutting apparatus. Consider the moment of tuber pinching between horizontal, vertical knives and the counter, corresponding to the beginning of the cutting of the tuber figure 2. The quality of the cutting process is significantly influenced by the angle of inclination of the counter \u03b1, which determines the moment of jamming of the tuber. When touching raw materials with a contradiction, a normal pressure reaction P of the contradiction to the tuber occurs. At this moment, the total cutting forces from the horizontal Rg and vertical knives Rv and the friction forces arising on the surface of the contradiction Ftr and the vertical knife Ftr1 also act on the tuber. In addition, in this case, the tuber gravity G and the normal reaction of the N disk, figure2. As a result, we get an arbitrary flat system of forces acting on the tuber. Let us analyze the process of cutting solid raw materials on the example of potatoes, the shape of the tuber of which, as an assumption, we will take in the form of a ball of radius r. According to the theorems of statics [8], we compose the equation of the sum of the moments of all the forces acting on the tuber, relative to point O, the tangent point of the horizontal tuber knife: ,0)cos(sincos)sin(cossin )cos(sinsin)sin(coscoscos t rt r 1 =+++\u2212 \u2212++++\u2212= = r\u0420r\u0420 rF\u0441osrFrNrGm n k \u041ek (1) AGRITECH-III-2020 IOP Conf", + " At the moment the tuber begins to move, the reaction force of the support N will be infinitesimal, and the friction forces determined by the equalities ,, t rt r BfRFPfF == (2) where f - coefficient of friction, reach the maximum value. In the limit equilibrium [9], when the force of the contradiction is many times greater than the gravity of the tuber P >> G, equation (1) takes the form .0)cos(sincos )sin(cossin\u0420)cos(sinsin)sin(cos t rt r =++ ++\u2212+++ r\u0420 rrF\u0441osrF After dividing both sides of the resulting equation by P, taking into account equalities (2), we will have .0)cos(sincos )sin(cossin)cos(sinsin)sin(cos =++ ++\u2212+++ r rrf\u0441osrf (3) From the geometry of figure 2 the equalities are true: , 2 2cos ,sin hrhrhrr \u2212=\u2212= (4) where h is the take-off height of the horizontal knife. Taking into account expressions (4) and the reduction of such terms, equation (3) is written 0)( 2 2sin 2 2)(cos =\u2212\u2212\u2212+\u2212+\u2212+ hrhrhfhrhhrfrf (5) If we introduce the notation AGRITECH-III-2020 IOP Conf. Series: Earth and Environmental Science 548 (2020) 052026 IOP Publishing doi:10.1088/1755-1315/548/5/052026 ,)( 2 2 , 2 2)( hrhrhf\u0412hrhhrf\u0410 \u2212\u2212\u2212=\u2212+\u2212= (6) then equation (5) is reduced to the trigonometric equation 0sincos =++ BArf (7) functions of the same argument \u03b1, corresponding to the optimal value of the incline angle of the contradiction of the considered chopper, at which the tuber will not roll and be destroyed by collapse" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0001775_0278364920955242-Figure10-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0001775_0278364920955242-Figure10-1.png", + "caption": "Fig. 10. (a) Collision between a closed-loop articulated body B and a fixed surface can be treated in the CSR model by estimating the projected dynamics at the contact point through the augmented object model. (b) Belted ellipsoid of effective mass for the articulated body B at the contact point (shown as a red point). (c) Belted ellipsoid of effective mass for just the rigid body RF at the contact point without considering the articulation.", + "texts": [ + " More generally, the CSR model reduces the problem of resolving a multi-surface articulated-rigid-body collision into a multisurface free-rigid-body collision. A solution to (20) is guaranteed to conserve linear and angular momentum of the rigid bodies in the system, in the dynamics subspace complementary to the joint and contact constraints (19). The above treatment of articulated-body collision applies uniformly to bodies with open-loop (or treestructure kinematic) mechanisms as well as to closed-loop (or parallel kinematic) mechanisms. Figure 10(a) shows a 6-bar mechanism, body B in contact with a fixed surface, and the effective mass is visualized at the contact point in Figure 10(b). The projected dynamics at the contact point is computed using the augmented object model (Khatib, 1995). The augmented object model has previously been applied to the control of contact forces at multiple points on the end-effector of a robotic manipulator (Park and Khatib, 2008). According to this model, the contact-space projected dynamics for the whole articulated body is equal to the sum of the projected dynamics of the two open-loop articulated bodies connected to the rigid body RF in contact, and the dynamics of RF itself at the contact point. Instead, if one only considered the effective mass of the rigid body RF (Figure 10(c)), the inertia of the colliding body would be underestimated. Lastly, we note that although the previous inferences were drawn for the multi-point collision scenario, identical inferences can be drawn for the multi-point steady contact scenario regarding the generalized accelerations and change of generalized momenta. To test the accuracy of the CSR model, we devised several setups consisting of pendulums with serial kinematic chain structures as shown schematically in Figure 11. The fabricated setups are shown in Figure 12" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0000896_0954406220912005-Figure2-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0000896_0954406220912005-Figure2-1.png", + "caption": "Figure 2. Components of the RSRR HeIse wheel.", + "texts": [ + " The motion of the links that are in contact with the ground is made in a plane orthogonal to the rotation axis of the wheel to prevent excessive forces produced by the dragging of the limbs against the ground during their extension, as it happens in the case in which the end of the legs move out of such plane.8,15 . The hybrid wheel can have one or any number of limbs, by replicating the kinematic chain of each deployable limb. Figure 1 shows a 3RSRR HeIse wheel in open and closed position. The mechanism consists of a sliding shaft, a rim, and a series of deployable limbs (see Figure 2). The kinematic chain of one limb consists of binary links connected by the following sequence of kinematic pairs: Revolute (R1), Spherical (S), Revolute (R2), and Revolute (R3). The sliding shaft is connected with the central rim through a prismatic joint (P) that allows the shaft to move axially on the rim. On the other hand, the shaft is also connected to a link (which we will refer to it as rod, and appears in the rest of the mechanisms) of the limb through the rotational joint R1. The link in contact with the ground (the traction link) is connected to the rod by a spherical joint (S)" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0000852_012079-Figure2-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0000852_012079-Figure2-1.png", + "caption": "Figure 2. Kinematic scheme of the planar parallel mechanism.", + "texts": [ + " Series: Materials Science and Engineering 747 (2020) 012079 IOP Publishing doi:10.1088/1757-899X/747/1/012079 As the introduction said, a workspace is one of the most important characteristics of the mechanism. For the given mechanism, it is easy to find the workspace dimension along the Z direction, because it is fully defined by the limits of the corresponding linear guides. The main difficulty is to evaluate the working area of the planar parallel mechanism, and the next section presents this procedure. Figure 2 shows a kinematic scheme of the planar parallel mechanism. For convenience, the instrument is fixed in a vertical position, and it is considered that the instrument and the platform form a single solid body. Other instrument orientations will be addressed in the next section. One can distinguish the following design constraints for the mechanism: \u2022 permissible ranges of the kinematic chains\u2019 lengths q1 \u2013 q4 (figure 2); \u2022 permissible rotation angles in the base joints A1 \u2013 A4; \u2022 permissible rotation angles in the platform joints B1 \u2013 B4. Though the real prototype has the points B1, B2 and B3, B4 being coincident, here we consider the general case. The mechanism parameters according to the prototype are presented in Table 1. In the current work we will focus on the three following types of the workspaces [12]: 1) a maximum workspace \u2013 a set of positions which the output link can place without any restrictions on its orientation; 2) a constant orientation workspace \u2013 a set of positions which the output link can place for one fixed orientation; ToPME-2019 IOP Conf", + "1088/1757-899X/747/1/012079 3) a dextrous workspace \u2013 a set of positions which the output link can place for any orientation in the given range. An optimization chord method [9\u201311] allows finding the boundaries for all these types of working areas. Corresponding MATLAB programs were written, and the next few figures present the obtained results. The mechanism scheme in an arbitrary position also presents on these plots for better clearness. First, the maximum workspace was obtained. In this case, there were no restrictions on the output link orientation angle, i.e. on the angle \u03c6 (figure 2). Figure 3 demonstrates the boundary of the working area. The pink arrows indicate output link orientation on the workspace boundary. The next figure shows the constant orientation workspace for \u03c6 = 0. ToPME-2019 IOP Conf. Series: Materials Science and Engineering 747 (2020) 012079 IOP Publishing doi:10.1088/1757-899X/747/1/012079 Finally, the dextrous workspaces for three ranges of \u03c6: [\u20135\u00b0, +5\u00b0], [\u201315\u00b0, +15\u00b0], and [\u201330\u00b0, +30\u00b0], \u2013 are presented on figure 5. As we can see, the optimization chord technique allows building different types of workspaces", + " The chord method also includes several optimization procedures, which are non-convex [13] in a general case. This results in a local solution, which sometimes can be not the correct one, and can even get stuck in this solution. The dense points at left and right boundary corners on figures 4 and 5 are examples of this situation (they are also the bifurcation points). At the beginning of the 3rd section, there was an assumption that the instrument had a vertical orientation. One can also deal with other orientations. Changing the instrument orientation will lead to a change in the length of CD (figure 2). The shape of the resulting working area will not change, but the whole workspace will move up or down. In conclusion, this paper presented an effective approach for workspace evaluation of the real robot prototype. The results of this work were successfully used in the experiments and were applied in the robot control system at the trajectory planning stage. The future work and the potential possibilities are to use the presented optimization chord method for the optimal dimensional synthesis with respect to the given workspaces" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0002222_icstcc50638.2020.9259658-Figure4-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0002222_icstcc50638.2020.9259658-Figure4-1.png", + "caption": "Fig. 4. Offset relations backward motion of last trailer w.r.t. path.", + "texts": [ + " Modeling of path-dependent relations is described e.g. in [2], [12], [16], [17]. In order to develop the offset relations, we introduce the equivalent virtual steering axles and their angles in the hitch points that describe the current body movement in Fig. 3. Note that from (3) and with the virtual steering axles, the relation \u03b8\u0307i = vi Li tan \u03b4i\u22121 (8) follows for each body. For the backward path tracking, it is convenient to take axle pm of the last trailer as the reference point of the vehicle. Fig. 4 shows the last trailer with the relations to the path. Newly introduced symbols are the curvilinear distance s along the path and the curvature \u03ba of the path. Subscript d denotes the desired values. Therefore, the desired path point is pd = pp(sd). 490 Authorized licensed use limited to: Univ of Calif Santa Barbara. Downloaded on May 21,2021 at 03:33:25 UTC from IEEE Xplore. Restrictions apply. The offset parameters are the off-track-distance dos and the orientation errors \u03b8os and \u03c6os,i, dos = (pm,1 \u2212 pd,1) sin \u03b8d + (pm,2 \u2212 pd,2) cos \u03b8d (9) \u03b8os = \u03b8m + \u03c0 \u2212 \u03b8d (10) \u03c6os,i = \u03c6i \u2212 \u03c6d, i = 1, " + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0002268_j.apples.2020.100032-Figure2-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0002268_j.apples.2020.100032-Figure2-1.png", + "caption": "Fig. 2. Boundary conditions and constraints for the FE analysis.", + "texts": [ + " Generative anufacturing processes, on the other hand, offer the potential to com- letely reproduce the existing geometric situation at the adaptations, ith simultaneous use of highly developed alternative materials and ptimization methods for weight minimization. A further advantage is the possible integration of functions, so that he sub-assembly consisting of the inner part, offset adapter and conecting elements can be designed as an integral component. The opti- ization goal is to reduce the mass while maintaining the operational afety. This applies both to the component strength and its rigidity. The umerical investigations are based on the bending cycle test, SAE J328005 . The resulting boundary and constraint conditions are shown n Fig. 2 . The wheel is fixed over the rim flange. A transverse force is applied at defined distance a , resulting in a bending moment. Due to the structure esulting from the five load cells and the rotating load (bending force F B ) f the measuring wheel, a cyclic symmetry for the optimized structure hould be aimed at. Accordingly, only a fifth of the structural compoent has to be considered for the further optimization process, since the urther geometry is obtained by mirroring. In order to map the circulatng load, the measuring wheel component is analyzed in 4\u00b0 rotating load teps" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0003617_ipec.2014.6870118-Figure13-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0003617_ipec.2014.6870118-Figure13-1.png", + "caption": "Fig. 13. System configuration of lock test.", + "texts": [ + " While a three-phase voltage (1 OOV 150Hz) was applied in this state, phase voltages and currents were measured by the analyzing recorder (Fig. 11). Fig. 12 (a) shows the result. On the other hand, Fig. 12 (b) shows simulation results of phase currents by (10) using the value of Table II. From Fig. 12, it can be confirmed that the simulation results agree well with the measurement results. Next, after returning the secondary conductor, the authors locked the mover with a load-sell and measured steady-state thrust at standstill using the test circuit shown in Fig. 13. Fig. 14 (a) shows the simulated thrust obtained by (10) and (20) using the values of Table II. From Fig.I4, it can be seen that the simulated thrust ripple components (amplitude and phase) agree well with the measured result. Thus, the results of Fig. 13 and Fig. 14 suggest the validity of the proposed two-phase model and its parameter measurement method. The obtained average thrust was Sl.lN. On the other hand, measured the calculated one was 96.SN. It is considered that the difference is caused by neglecting iron loss effect in (l0) and (20). VI. Conclusions The results are summarized as follows. 1) A new simple two-phase circuit model of LIM was proposed. This model is simpler than conventional model, but can correctly calculate both asymmetric primary current and thrust ripple performances" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0002067_0954410020965094-Figure1-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0002067_0954410020965094-Figure1-1.png", + "caption": "Figure 1. Engagement geometry of missile and target.", + "texts": [ + " Next, the Singularity-free fast TSM structure is introduced and some preliminary concepts relevant to the fixed-time stability theory is presented. Later, the process of designing the smooth fixed-time guidance law with terminal angle constraint is described and its fixed-time stability is proved. Finally, various maneuvering targets simulation scenarios are shown to verify the superior performance of the smooth fixedtime guidance law, and conclusions are provided in the final section. Considering the scenario of missile interception maneuvering target, the two-dimensional engagement geometry is described in Figure 1. We assume the missile and the target are regarded as mass points. The engagement dynamics equations are as follows _r \u00bc Vtcos\u00f0k ct\u00de Vmcos\u00f0k cm\u00de (1) r _k \u00bc Vtsin\u00f0k ct\u00de \u00fe Vmsin\u00f0k cm\u00de (2) _ct \u00bc at Vt (3) _cm \u00bc am Vm (4) where k and r denote respectively the line of sight (LOS) angle and relative distance from missile to target; _k and _r denote respectively the LOS angular rate and relative velocity. V, c, a denote respectively the tangential velocities, flight path angles, and normal accelerations" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0000115_012066-Figure2-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0000115_012066-Figure2-1.png", + "caption": "Figure 2. The scheme of the laser head integration into the machining center", + "texts": [ + " That is, under these conditions, in order to realize the possibility of using a laser head on a CNC machine, it is necessary to partially or completely modernize the laser head frame and adapt it for manual installation in the machine spindle by an operator or a preset person. Fig. 1 shows an example on how to use a tool cone to install the laser head in a CNC machine. A diagram of the laser head integration into the machining center, developed with using a circuit diagram of a fiber laser setup, presented in [8], is shown in Fig. 2. IPDME 2019 IOP Conf. Series: Earth and Environmental Science 378 (2019) 012066 IOP Publishing doi:10.1088/1755-1315/378/1/012066 When solving the problem of integrating the laser head into the machining center, it is necessary to determine the installation location of the main system unit. Since the control unit of the laser system has a large number of components sensitive to electromagnetic effects, it can not be installed directly into the control cabinet of the CNC system. It can only be placed directly next to the CNC cabinet" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0001115_ab9036-Figure2-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0001115_ab9036-Figure2-1.png", + "caption": "Figure 2. Schematic diagram of the optical system.", + "texts": [ + " When linearly polarized light emitted from the laser enters the LC, the polarization plane rotates by 90\u25e6 due to optical rotatory power of the LC molecules and passes through the polarizing plate in a state where no voltage is applied to the LC. On the other hand, when a voltage is applied to LC, the optical rotatory power of the LC molecules is lost, and the light is blocked by the polarizing plate. Therefore, the transmission and blocking of light can be controlled by controlling the application of voltage. The principle of laser beam selection is shown in figure 1. A schematic diagram of the optical system is shown in figure 2. This system is based on the principle of the optical interferometer with some additions. The laser beam is diverged by a concave lens. The diverged light enters the transmissive LC and the polarizing plate, and then enters a BL attached to the measuring object. Four BLs were attached to the measuring object and used as target mirrors (TMs). BLs with a high refractive index close to 2.0 were used, and they were coated with silver film hemispherically. As a result, it is expected that the incident light on the BL is retroreflected [19]. The retroreflected light from the four BLs is originally a single light beam emitted from the laser. Thus, these lights converge to a single point. Here, as shown in figure 2, by setting up a light-receiving element at a position where the retroreflected light converges after entering the beam splitter, it is possible to detect the light reflected from four BLs with one photodetector (PD). When the displacement in the yaw direction is detected, the change in the phase difference between TM1 and TM3 is detected by transmitting only the light entering TM1 and TM3 by the LC and, conversely, blocking the light entering TM2 and TM4. These operations are performed by controlling the color image output to one LC screen" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0003758_2014-01-2520-Figure2-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0003758_2014-01-2520-Figure2-1.png", + "caption": "Figure 2. Brake judder road load acquisition schematic diagram", + "texts": [ + " The paper ignore the effect of road vibration on the generation of brake rotor thickness variation when simulating brake road testing on the dyno. Repeated brake application is taken as the only factor to study the effect on RTV generation. Brake pad wear, brake energy, and temperature are the three key factors in brake application. Based on the above analysis, specific attention is given to brake energy intensity, temperature and wear. Road load data was recorded with LINK 3801, taking the brake light signal as the braking acquisition trigger signal. The road test sketch map is shown in Figure 2. The brake speed, brake pad temperature, deceleration, pressure, and ambient temperature were collected. These data were used to study the relations between brake pad wear and brake energy and temperature, taken as the input conditions of the road simulation. The vehicles were equipped with the most powerful engine specified for brake testing in the Huangshan tests. The vehicle test mass represents the heaviest vehicle version, loaded with the mass of two people. A temperature profile of a day of testing is shown in Figure 3" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0000613_pesgm40551.2019.8973400-Figure3-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0000613_pesgm40551.2019.8973400-Figure3-1.png", + "caption": "Fig. 3. Technical drawing of a HELIAS-5B power plant concept [36]. Shown are the plasma vessel, the five different coil types, the coil support structure, the magnet system support, and the outer cryostat vessel. The horizontal and vertical ports are indicated to demonstrate that the coils and the coil support structure permit large-scale access to the plasma vessel.", + "texts": [ + " Concerning further upgrades of the steady-state heating power, ECRH launchers, parts of the transmission optics and the gyrotron building, already foresee two additional gyrotrons, increasing the number of microwave beams from 10 to 12. Combined with developing new gyrotrons, which, similar to the ITER gyrotrons, are designed for a power of 1.5 MW, this would increase the power level by almost a factor of two. A direct extrapolation from W7-X to a power plant is the helical advanced stellarator (HELIAS) [33], [34]. Recent studies focus on the HELIAS-5B [35]\u2013[37], which like W7X has fivefold symmetry (Fig. 3). The aspect ratio is similar to W7-X, the major radius is 22 m, and the average minor radius is 1.8 m. The average magnetic field lies in the range of 5 to 6 T with a maximum at the coils ranging from 10 to 12 T. From the engineering point of view, 6 T on axis seem feasible. However, to ease the requirements for the support structure and to save costs, lower values are desirable. The size of the coils and the magnetic field are similar to the ITER values, enabling the use of the ITER coil technology [38] including the superconductor material and the magnetic field values" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0001858_icra40945.2020.9197549-Figure3-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0001858_icra40945.2020.9197549-Figure3-1.png", + "caption": "Fig. 3: The additional fiber of length L0 adhered to an extending-type FREE permits no extension along its right side. As the volume of the FREE increases from its relaxed volume (left) to a larger volume (right), so does the arc length along its central axis L. However, the arc length along its right side remains constant, which induces a curvature of radius R and angle \u03b8.", + "texts": [ + "74\u25e6 can be calculated using trigonometry. As previously stated, the maximum length Lmax occurs when a FREE has been deformed to the point where its fiber angle equals the \u201cneutral angle\u201d of \u2248 54.74\u25e6. As pictured in Fig. 2, this length corresponds to the height of a triangle with hypotenuse B and angle 54.74\u25e6, which is given by the following expression, Lmax = B cos54.74\u25e6 = L0 cos54.74\u25e6 cos\u03b10 . (4) Extension can be converted into bending by adding a strain-limiting fiber along the body of a FREE, as shown in Fig. 3. As the volume of the FREE increases, the fiber permits no extension of one of its sides, which results in a bent geometry. This bending can be quantified in terms of a radius of curvature R and angle of curvature \u03b8, as shown in Fig. 3. Assuming that the arc length of the central axis of a bent FREE is equal to the extended length of an identical FREE without the strain-limiting fiber and that the inner arc length is equal to L0, the following expressions hold true: L = R\u03b8, (5) L0 = (R\u2212R)\u03b8. (6) Combining Equations (5) and (6), solving for R, and taking its reciprocal yields an expression for the curvature of the FREE, denoted K , K = 1 R = 2\u03c0N ( 1\u2212 L0 L ) ( B2\u2212L2 )\u2212 1 2 . (7) The maximum curvature Kmax is given by substituting Lmax into Equation 7, with the following result, Kmax = 1 R0 ( sin\u03b10 sin54" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0002923_9781119633365-Figure1.13-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0002923_9781119633365-Figure1.13-1.png", + "caption": "Figure 1.13 Converters with multiple pairs of active\u2013passive switches: (a) half-bridge and (b) full-bridge configurations.", + "texts": [ + " However, the current flow in LR can be bidirectional and has higher resonant frequency, while the one in inductor L1 is unidirectional only. How to construct this type of quasi\u2010resonant converters is worth further discussing. In literature, there are similar converters, such as zero\u2010current switching 1.3 Pween-noon PWM onnwrtwre 15 quasi\u2010resonant converters and multi\u2010resonant converters. Can they be developed with a systematical approach? PWM converters can have more pairs of active\u2013passive switches, such as the half\u2010bridge and full\u2010bridge converters shown in Figure\u00a01.13. Figure\u00a01.13a shows a half\u2010bridge configuration, which has two pairs of switches. The two switches take turn conducting, and each one takes care of one\u2010half switching cycle. In each half cycle, the switch is pulse\u2010width modulated to control power flow from the input to the output. If the natural frequency of the L1C1 network is designed to be far below the switching frequency, the converter is just like a conventional PWM converter. On the other hand, if the frequency is close to the switching frequency, the current and voltage waveforms are sinusoidal\u2010like, and it is called a resonant converter. In fact, it is still belonged to a PWM converter but just with variable frequency operation. In general, it is also classified as a PWM converter, because its power transfer is still limited by an LC network. Figure\u00a0 1.13b shows a full\u2010 bridge converter, in which there are four switches and they form two pairs, S1&S4 and S2&S3. When these two pairs of switches take turn conducting or are in bipolar operation, the converter is the same as the half\u2010bridge one. Again, it can act as a conventional PWM or a resonant converter depending on the order of the LC network natural frequency. This is also classified as a PWM converter. All of the converters discussed above are non\u2010isolated. By introducing transformers into the non\u2010isolated versions of PWM converters, they can be transformed to their isolated counterparts" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0003370_chicc.2015.7260834-Figure1-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0003370_chicc.2015.7260834-Figure1-1.png", + "caption": "Fig. 1: Relative coordinates", + "texts": [ + " 2 Problem Formulation Consider a collection of mobile agents modeled as unicy- cles: x\u0307i = vi cos(\u03b8i) y\u0307i = vi sin(\u03b8i) \u03b8\u0307i = wi, i = 1, \u00b7 \u00b7 \u00b7 , n (1) where col(xi, yi) \u2208 R 2 denotes the position of vehicle\u2019s center of mass in the plane, \u03b8i \u2208 R is the vehicle\u2019s orientation, and ui = col(vi, wi) \u2208 R 2 is the control input with vi as the linear speed and wi as the angular speed. Associated with system (1) is a so-called digraph1 G(t) = (V , E) where V = {1, ..., n} with i = 1, \u00b7 \u00b7 \u00b7 , n associated with the ith subsystem of (1), and E \u2286 V \u00d7 V . The set V is called the node set of G and the set E is called the edge set of G. Like in [10, 17], we first consider a special pursuit graph with edge set given by E = {(1, 2), \u00b7 \u00b7 \u00b7 , (i, i+1), \u00b7 \u00b7 \u00b7 , (n\u2212 1, n), (n, 1)}. To facilitate analysis, we introduce the relative coordi- nates as shown in Fig. 1. Let \u23a1 \u23a3 x\u0304i y\u0304i \u03b8\u0304i \u23a4 \u23a6 = R(\u03b8i+1) \u23a1 \u23a3 xi \u2212 xi+1 yi \u2212 yi+1 \u03b8i \u2212 \u03b8i+1 \u23a4 \u23a6 where R(\u03b8) = \u23a1 \u23a3 cos(\u03b8) sin(\u03b8) 0 \u2212 sin(\u03b8) cos(\u03b8) 0 0 0 1 \u23a4 \u23a6. Let ri = \u221a x\u03042 i + y\u03042i , \u03b1i = arctan( y\u0304i x\u0304i )+\u03c0\u2212 \u03b8\u0304i, \u03b2i = \u03b8\u0304i \u2212\u03c0 with ri denoting the distance between vehicle i and vehicle i + 1, \u03b1i denoting the bearing angle, i.e., the angle difference between vehicle i\u2019s heading and the line of sight which would take directly towards vehicle i + 1, and \u03b2i being the heading difference minus \u03c0. Note that we identify index n+ 1 with 1 and set \u03b1i \u2208 [\u2212\u03c0, \u03c0) and \u03b2i \u2208 [\u2212\u03c0, \u03c0)" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0000742_ssci44817.2019.9003002-Figure10-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0000742_ssci44817.2019.9003002-Figure10-1.png", + "caption": "Fig. 10. Silicon point contact diode in its 1941 form (from [75], copyright (2014) by Alcatel-Lucent). It was the first diode fabricated at a large scale and was primarily used in the Allies\u2019 radar systems during World War II.", + "texts": [ + " As an electrochemist by education, he understood that crystals used in the detectors were not uniform, and that only a right kind of impurity would give the best responsivity. Thus, he first pushed for ultrapure crystal [73]. By 1939, silicon crystal ingot purity attained 99.8%. When ingots were polished prior to be contacted by the cat whisker, resulting rectifiers showed a behavior so reliable that any part of silicon wafer could be used to yield the same result. Searching for the right spot was no longer necessary and the point contact could be packaged, ready for rugged applications (Fig. 10). As can be seen in Fig. 5, responsivity did not get improved in nearly 30 years. The Bell Laboratories diodes of 1939 were not more sensitive than molybedine\u2013copper contacts of 1910. But as it gained in reliability, the point contact could return to commercial/military applications. Vol. 102, No. 11, November 2014 | Proceedings of the IEEE 1675 The diode rapidly became a key building element in the front\u2013end mixer within radar systems, so precious in the wartime. Crystals were reliable but far from being perfect" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0001696_9781119679455-Figure10.15-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0001696_9781119679455-Figure10.15-1.png", + "caption": "Figure 10.15 Delay calculation of AND plane: (a-b) delay propagation path of a gate and a cross-point and (c) overall delay propagation path for AND plane.", + "texts": [ + "6 Reversible Write-Enabled Master\u2013Slave flip-flop 132 Figure 10.7 Block diagram of a reversible RAM 132 Figure 10.8 A reversible logic element of Plessey FPGA 133 Figure 10.9 3 \u00d7 3 Reversible MUX gate 135 Figure 10.10 Different uses of a Feynman gate 135 Figure 10.11 One template of toffoli gate 135 Figure 10.12 Two templates of MUX gate 135 xxii List of Figures Figure 10.13 Ex-OR plane realization for the function F based on the Algorithm 10.2.1.1 136 Figure 10.14 Design of reversible PLAs for multi-output function F 138 Figure 10.15 Delay calculation of AND plane: (a-b) delay propagation path of a gate and a cross-point and (c) overall delay propagation path for AND plane 140 Figure 10.16 Delay calculation of Ex-OR plane: (a-b) delay propagation path of a gate Figure 11.2 Quantum realization of 3 \u00d7 3 reversible FS gate 144 Figure 11.3 2 \u00d7 22 Reversible decoder 145 Figure 11.4 3 \u00d7 23 Reversible decoder 146 Figure 11.5 n \u00d7 2n reversible decoder 146 Figure 11.6 Reversible D flip-flop 147 Figure 11.7 Reversible write-enabled master\u2013slave D flip flop 147 Figure 11", + "1 l inputs and TDOT cross-points in the AND plane, the total number of garbages in the AND plane of reversible PLA is l + q \u2212 TDOT. In this subsection, the delay of reversible PLAs is calculated in a greedy approach. The calculation is divided into two phases: (i) AND plane delay (APD (pi)) and (ii) Ex-OR plane delay (XPD (pi)) in terms of product lines (horizontal lines). Then both of the delays are merged with respect to both planes. The multi-output function F is used to calculate the delay. First, the delay of AND plane is calculated and then the delay of Ex-OR plane is calculated. Figure 10.15 and Figure 10.16 show the delay calculation of AND plane and Ex-OR plane, respectively. To calculate the delay, the following things are considered: 1. Gate (Via) is represented as circle (DOT). 2. Delay of any gate is 1 and via (DOT) denotes 0. 3. Decimal values show the delay of corresponding circle. For AND plane, every gate updates its delay by comparing the delay of neighboring gates at left (L) and top (T) and then, it propagates the updated delay to neighboring gates placed in right (R) and bottom (B) sides, as shown in Figure 10.15. Each circle in Figure 10.15 represents the delay of particular point and arrows show the path of delay propagation. For Ex-OR plane, every gate updates its delay by comparing the delay of neighboring gates at right (R) and bottom (B) and then, it propagates the updated delay to neighboring gates placed in left (L) and top (T) sides as shown in Figure 10.16. After calculating the delay of both planes, the delay of product lines having maximum value is the final delay of the overall design of reversible PLAs. Algorithm 10.5 is introduced here 10" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0003988_15599612.2014.916016-Figure10-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0003988_15599612.2014.916016-Figure10-1.png", + "caption": "Figure 10. Experimental condition of tracking a target. A sheet was set perpendicular or with an angle to the endoscope.", + "texts": [ + " Moreover, the extracted feature amount and the correctly tracked number are almost two times of those for the phantom, which is the main reason of the increased processing time. The tracking performance reduces greatly for the 0.5-1mm pig vessel. There is one time that the tracking performance is about 55% during 5 experiments. 3.2.1. Target on paper vessels 3.2.1.1. Target on paper phantom. To evaluate the tracking affectivity, several spots in the view of field are assessed. In the front of the endoscope, the observation object is placed and then we select the tracking target. The observation object is perpendicular to, or 30 degrees with the endoscope, shown as Figure 10. During surgery for the treatment of TTTS, to ensure effective laser coagulation, the tip of the endoscope must be close to the object vessel point. In our experiment, therefore, we set the distance between the tip of the endoscope and the observation image paper to 10mm, 15mm, 20mm for the perpendicular condition, and for the 30 degrees we do the experiment of 10mm to obtain the influence of the degree of angle. Then the points shown in the Figure 11 are assessed, the point A is the center point, point B the left-up point and point C the left point" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0003337_9781118886397.ch1-Figure1.11-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0003337_9781118886397.ch1-Figure1.11-1.png", + "caption": "Figure 1.11 Three-dimensional vectors in spherical form.", + "texts": [ + " is normal to the cone C, generated by sweeping the vector B around the vertical z-axis, and pointing away from the vertical z-axis. The angular unit vector ?\u0302? is normal to the vector B. The angular unit vector ?\u0302? is identical to the same vector in the cylindrical coordinate system, normal to the plane A, defined by the vector B and the vertical z-axis. The angular unit vector ?\u0302? is tangential to the sphere S, in the direction of increasing angle, and normal to the other two unit vectors. The spherical coordinate system is shown in Figure 1.11. The spherical unit vectors can be expressed in terms of the Cartesian unit vectors: r\u0302 = sin \ud835\udf03 cos\ud835\udf19i + sin \ud835\udf03 sin\ud835\udf19j + cos \ud835\udf03k\u0302 (1.57) ?\u0302? = cos \ud835\udf03 cos\ud835\udf19i + cos \ud835\udf03 sin\ud835\udf19j \u2212 sin \ud835\udf03k\u0302 (1.58) ?\u0302? = \u2212 sin\ud835\udf19i + cos\ud835\udf19j (1.59) The components of the vector (a, b, c) converted to spherical coordinates (r, \ud835\udf03, \ud835\udf19) are r = \u221a a2 + b2 + c2 (1.60) \ud835\udf03 = cos\u22121 c r (1.61) \ud835\udf19 = tan\u22121 b a (1.62) The cylindrical coordinates of the example vector of B = (4, 5, 6) shown in Figure 1.11 are r = \u221a 42 + 52 + 62 = 8.775 \ud835\udf03 = cos\u22121 6 8.775 = 46.86\u2218 \ud835\udf19 = tan\u22121 5 4 = 51.34\u2218 The Cartesian unit vectors are expressed in spherical form as i\u0302 = sin \ud835\udf03 cos\ud835\udf19r\u0302 + cos \ud835\udf03 cos\ud835\udf19?\u0302? \u2212 sin\ud835\udf19?\u0302? (1.63) j\u0302 = sin \ud835\udf03 sin\ud835\udf19r\u0302 + cos \ud835\udf03 sin\ud835\udf19?\u0302? + cos\ud835\udf19?\u0302? (1.64) k\u0302 = cos \ud835\udf03i \u2212 sin \ud835\udf03j (1.65) The components of the vector (r, \ud835\udf03, \ud835\udf19) converted to rectangular coordinates (a, b, c) are a = r sin \ud835\udf03 cos\ud835\udf19 (1.66) b = r sin \ud835\udf03 sin\ud835\udf19 (1.67) c = r cos \ud835\udf03 (1.68) In the three-dimensional spherical coordinate system, a differential vector dl may be defined from the differential lengths dr, d\ud835\udf03, and d\ud835\udf19 as dl = dr r\u0302 + rd\ud835\udf03" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0003268_chicc.2015.7259719-Figure1-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0003268_chicc.2015.7259719-Figure1-1.png", + "caption": "Fig. 1: Definition of body axis and stable axis", + "texts": [ + " However, the normal acceleration of aircraft is closely coupled to the angle of attack \u03b1, which plays a vital role in the equation of motion in longitude. For nonlinear approach, \u03b1 is selected as controlled variable here. In the lateral direction, roll rate p and sideslip \u03b2 are also considered as command control. For convenient analysis, the body-fixed axis selected as the basic of motion axis. However, the angle of attack \u03b1 and sideslip \u03b2 appear in the stable axis in general. For that, the rotation axis can instead be selected as the stable axis, see Fig.1. The objective of attitude control of aircraft is to design a controller, which guarantee controlled variables (\u03b1, \u03b2, pw) tracking a desired trajectory (\u03b1d, \u03b2d, pdw). \u03b1 \u2192 \u03b1d \u03b2 \u2192 \u03b2d pw \u2192 pdw (1) The motion equations of aircraft have been defined in [14], and the results in wind axis can be expressed as following: V\u0307T = 1 m (\u2212D + FT cos\u03b1 cos\u03b2 +mg1) \u03b1\u0307 = 1 cos \u03b2 (qw + 1 mVT (\u2212L\u2212 FT sin\u03b1+mg2)) \u03b2\u0307 = \u2212rw + 1 mVT (Y \u2212 FT cos\u03b1 sin\u03b2 +mg3) (2) where, VT ,m are velocity and mass of aircraft, respectively. pw, qw, rw are roll rate, pitch rate and yaw rate in stable axis, respectively" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0003788_detc2014-34213-Figure3-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0003788_detc2014-34213-Figure3-1.png", + "caption": "Fig. 3 A six-bar linkage with two prismatic joints", + "texts": [ + " In other words, link 5 is perpendicular to the corresponding ground. When the prismatic joint K is used as the input joint, the singularity of the planar linkage happens when the three instant centers I12, I16 and I25 in Fig. 2(b) become collinear, i.e., the links AB, CD and EK intersect at the common point J. The result is similar to the case when the input joint is given through joints A and G (Fig. 1) for Stephenson Type III six-bar linkage. The single-DOF planar linkage with two prismatic joints is shown in Fig. 3. In Fig. 3, let the joint A as the input joint. The singularity of the planar linkage happens when the three instant centers I24, I34 and I13 of the four-bar equivalent linkage I12I24I34I13 in Fig. 3(a) become collinear or the instant centers I25, I56, I16 of the four-bar equivalent linkage I12I25I56I16 (Fig. 3b) become collinear. That is, the link 5 is perpendicular to the corresponding ground. When the prismatic joint K is used as the input, the singularity of the linkage happens when the instant centers I12, I16 and I25 in Fig. 3(c) become collinear, i.e., the links AB, CD and EK intersect at the same point. The single-DOF planar linkage with three prismatic joints is shown in Fig. 4. If the joint A is used as the input joint, the singularity of the planar linkage happens when the instant centers I24, I34 and I13 in Fig. 4(a) become collinear or the three instant centers I25, I56 and I16 in Fig. 4(b) become collinear. In other word, the link 5 is perpendicular to the corresponding ground. When the prismatic joint K is used as the input joint, the singularity of the planar linkage occurs when the instant center I12, I16 and I25 in Fig. 3(c) become collinear, or the link AB, CD and EK intersect at the same point. \u03b3 45I 12I 25I 24I 13I 34I 56I 15I \u03b3 45I 12I 25I24I 13I 34I 56I 16I14I 15I 13I (a) (b) Fig. 2 A six-bar linkage with one prismatic joint K 1\u03b8 \u03b2 \u03b7 24I 34I 14I 12I (a) 1\u03b8 \u03b2 \u03b7 15I25I45I 24I 34I 12I (b) 2 Copyright \u00a9 2014 by ASME Downloaded From: http://proceedings.asmedigitalcollection.asme.org/pdfaccess.ashx?url=/data/conferences/asmep/82119/ on 04/05/2017 Terms of Use: http://www.asme.org/about-asme/terms-of-use For the two-DOF seven-bar planar linkage with all rotational joints in Fig" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0003451_s00707-015-1403-6-Figure4-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0003451_s00707-015-1403-6-Figure4-1.png", + "caption": "Fig. 4 A model of a nonlinear nonholonomic constraint (NNC) with parallel velocities", + "texts": [ + " Example 2 Consider the models of a nonlinear nonholonomic constraint whose physical realization is given in papers [11] and [12]. In those models, a nonlinear nonholonomic constraint has a clear physical meaning and can be equally treated with respect to a linear nonholonomic constraint. The first model consists of two particles, connected by a light-weight rod,movingwith parallel velocities. Physical realization of this constraint is accomplished by positioning the Chaplygin blade in the center of a light-weight rod, perpendicular to the rod\u2014Fig. 4a. Let us choose a set of quantities (x, y, \u03be, \u03d5) for the generalized coordinates, represented in the Figure, but they are not \u03be \u2212 2l = 0, x\u0307 cos\u03d5 \u2212 y\u0307 sin \u03d5 = 0, (5.20) where the second constraint results from the condition that the particle M of the Chaplygin blade does not have the velocity perpendicular to the blade, i.e., perpendicular to the direction AB but has the velocity only along the blade, i.e., along the direction AB\u2014Fig. 4b. Accordingly, the given system has two degrees of freedom. However, a set (x1, y1, x2, y2) can be also chosen for the generalized coordinates, where there exist the constraints (x2 \u2212 x1)2 + (y2 \u2212 y1)2 = 4l2, v1|| v2 \u2192 x\u03071 y\u03071 = x\u03072 y\u03072 \u21d2 x\u03071 y\u03072 \u2212 y\u03071 x\u03072 = 0, (5.21) with the second constraint expressing the parallelism condition of the particles M1 and M2 and representing a nonlinear nonholonomic constraint. Nonlinear tracking of a nonholonomic constraint emerges in the Cartesian coordinate system on account of relating the velocities of two particles within those coordinates (parallelism of velocities of the particles M1 and M2), and in the coordinate system the constraint tracking is linear because it regards the restriction of the velocity of only one particle, i" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0003670_2014-01-1727-Figure2-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0003670_2014-01-1727-Figure2-1.png", + "caption": "Figure 2.", + "texts": [ + " The effective reduction ratio is then 200:1 and output torque will be 200 times greater than the input torque. The torque on the CVT (variable-belt) is equal to 1/2 the output torque. Therefore, the torque limitation is still determined by the CVT (variable-belt). It will be shown later how the use two variable-inertia-flywheels (VIF) overcome this limitation. There are numerous examples of this technique being used to form an IVT. Among the best known is the TorotrakTM system, which incorporates a toroidal variator similar to the NSK Half Toroidal CVT (Figure 2). There are two inputs to an epicyclic differential: One is variable using a torodial variator and the other is fixed. The output (ring) is the sum of each input divided by two. Figure 3 shows an Input coupled split-power transmission with re-circulated power [11]. All of these configurations can be shown diagrammatically as in Diagram A. This is a motor vehicle using a standard ICE (internal combustion engine) that is set to run at a fixed speed of 2200 RPM. The engine output is coupled to one side of a differential with a 1:2 step-up gear box and the other side with a conventional CVT operated through a range of 1 to 2 through 0" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0002355_j.egyr.2020.11.179-Figure1-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0002355_j.egyr.2020.11.179-Figure1-1.png", + "caption": "Fig. 1. Topologies of BEESMs. (a) BEESM with four field winding coils (BEESM1). (b) BEESM with two field winding coils (BEESM2).", + "texts": [ + " The rest of this paper s organized as follows. In Section 2, the topology and operation principle of BEESM are analyzed. In Section 3, he electromagnetic performance indexes, including the air-gap flux density, flux linkage, back electromotive force back-EMF), cogging torque, output torque and losses of the BEESM are fully analyzed. In Section 4, the proposed EESM is optimized to suppress the torque ripple. In Section 5, the conclusions of this paper are drawn. . Topology and operation principle .1. Topology Fig. 1(a) and (b) illustrate the topologies of BEESMs with four and two field winding coils, which are named s BEESM1 and BEESM2, respectively. As shown, the rotor of BEESMs are made of silicon steel sheet without ny auxiliary parts and are designed as arc-shape to weaken the harmonic of air-gap flux density. Two kinds of indings are placed on the stator, one is dc field winding, and the other is ac armature winding. The field winding s concentrated and the armature winding is distributed. One field winding is corresponding one salient pole and two alient poles for the BEESM1 and BEESM2, respectively" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0001273_ab9212-Figure3-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0001273_ab9212-Figure3-1.png", + "caption": "Figure 3. Incorporating an adjustable pivot into the pencil holder design.", + "texts": [ + " The author could see how the student found a sense of purpose to learn about this upper secondary physics concept to make decisions for his design, which would eventually enhance the user-experience. The student\u2019s sketches and annotations incorporating physics into his pencil holder design as well as how he conceptualized this idea from the toy shown to him are shown in his design sheet and prototype in figure 2. The terms \u2018lever\u2019, \u2018pivot\u2019 and \u2018principle of moments\u2019 in relation to his design are also evident in the design sheet. July 2020 3 Phys. Educ. 55 (2020) 045019 Another creative design is shown in figure 3. The design highlights how a student incorporated an \u2018adjustable pivot\u2019 into his prototype. The pencil holder consists of two parts\u2014themain portion and a rectangular wooden piece which is connected to the main portion by a wooden dowel that acts as the \u2018adjustable pivot\u2019. The student has described how this \u2018adjustable pivot\u2019 can be plugged in and out of the main portion, allowing the pencil holder to hold both short and long pencils. The mechanism for the pencil holder is highlighted in figure 4" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0001511_jsyst.2020.3006990-Figure27-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0001511_jsyst.2020.3006990-Figure27-1.png", + "caption": "Fig. 27. Quansar unmanned ground vehicle [31].", + "texts": [ + " The MPC minimizes an objective function that can be of the following form: \u03a6(k) = N\u2211 j=1 x\u0303T (k + j|k)Qx\u0303(k + j|k) + u\u0303T (k + j \u2212 1|k)Ru\u0303(k + j \u2212 1|k) (28) whereN is the prediction horizon, and Q and R are positive definite and positive semidefinite weighting matrices, respectively. The notation a(j|k) is the value of the quantity a at the instant j predicted at the time instant k. Hence, the optimization problem can be expressed as finding u\u0303\u2217 such that u\u0303\u2217 = minu\u0303{\u03a6(k)}. (29) The Quanser unmanned ground vehicle (QGV) (see Fig. 27) is a testbed designed and manufactured by Quanser [31], and is available at the Network Autonomous Vehicle Lab in the Department of Mechanical, Industrial and Aerospace Engineering, Concordia University. The QGV is a differential drive ground robot that has a 4-DOF robotic manipulator with gripper. To use on-board sensors\u2019 measurements and drive the motors, the QGV utilizes Quanser\u2019s embedded data acquisition card (DAQ), a microcontroller board named (HiQ), and the embedded Gumstix computer. The HiQ is a high-resolution input/output (I/O) card, which can be used effectively in path tracking applications" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0000940_ilt-10-2019-0418-Figure1-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0000940_ilt-10-2019-0418-Figure1-1.png", + "caption": "Figure 1 Schematics of the geometry of misaligned GFB", + "texts": [ + " The hydrodynamic pressure causes the deformation of foils, which in turn distorts the pressure distribution; this interaction continues until it reaches a steadystate. From a mathematical point of view, the calculation of GFBs is based on the Reynolds equation. In this paper, the ideal gas isothermal assumption is adopted without considering the rarefaction effect. The equation can be simplified according to the following parameters: x \u00bc Ru ; z \u00bc RZ;U \u00bc vR; p \u00bc PaP; h \u00bc CH; K \u00bc 6mv P0 R C 2 ; t \u00bc tv Hence, the dimensionless Reynolds equation can be expressed as: @ @u PH3 @P @u 1 @ @Z PH3 @P @Z K @ PH\u00f0 \u00de @u 2K @ PH\u00f0 \u00de @t \u00bc 0 (1) FromFigure 1, the film thickness (H) can be described as: H \u00bc 11 ffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi \u00abXZ\u00f0 \u00de2 1 \u00abYZ\u00f0 \u00de2 q cos arctan \u00abXZ=\u00abYZ\u00f0 \u00de1 w\u00bd 1 u (2) where \u00abXZ \u00bc Z C tanb 1 e0 C sinw 1 L 2RC tanb \u00abYZ \u00bc Z C tana1 e0 C cosw 1 L 2RC tana Here, the X-axis is defined as the load direction in this model. The pressure distribution can be obtained by introducing equation (2) into equation (1), and the foil deformation (u) should be updated at each step until the iteration is completed", + " Performance of gas foil bearing with misalignment Hao Li, Peng Hai Geng and Hao Lin Industrial Lubrication and Tribology To obtain the pressure distribution, equation (1) is discretized using the finite difference method (FDM), the foil deformation is solved by FEM, and the Newton\u2013Raphsonmethod is used to iterate those equations. The obtained pressure distribution is then used to estimate the static and dynamic performance. The bearing load capacity (W) is calculated as: W \u00bc ffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi FX 2 1F2 Y p FX \u00bc 2 \u00f0L R 0 \u00f02p 0 Psinu du dZ (4) As shown in Figure 1, theFy should be 0 approximately. The restoringmoments (MX,MY) can be described as: MX MY \u00bc Z \u00f0 L 2R 0 \u00f02p 0 P cosu sinu du dZ \u00f0L R L 2R \u00f02p 0 P cosu sinu du dZ ( ) (5) The friction force (T) can be written as: T \u00bc R \u00f0L R 0 \u00f02p 0 H 2 @P @u 1 K 6 1 H du dZ (6) The small perturbation method is used to analyze dynamic coefficients of gas film. The steady position is perturbed by small displacements (DX, DY) and small misalignments (Da, Db ). Therefore, Taylor expansion can be carried out: W \u00bc F1 @F @X 0 DX1 @F @Y 0 DY 1 @F @a 0 Da1 @F @b 0 Db 1 @F @X 0 0 DX 0 1 @F @Y 0 0 DY 0 1 @F ", + " Performance of gas foil bearing with misalignment Hao Li, Peng Hai Geng and Hao Lin Industrial Lubrication and Tribology When only displacement perturbations are considered, eight equations are obtained, which stand for stiffness coefficients (KXX, KXY, KYX and KYY) and damping coefficients (CXX, CXY, CYX and CYY). When misalignments appear, eight perturbation equations are added, which mean dynamic coefficients (Kaa, Kab , Kba, Kb b , Caa, Cab , Cba and Cb b). Once the perturbation pressure distribution is obtained, the above 16 dynamic coefficients can be calculated with @F=@X, @F=@Y, @F=@X 0 , @F=@Y 0 and @M=@a, @M=@b , @M=@a 0 and @M=@b 0 in equation (7). The program scheme is shown in Figure 2. In the simulation of this paper, the GFB, shown in Figure 1, is composed of 1 top foil and 3 bump foils. The No.1 and No.3 bump foil have 15 arches, while the No.2 bump foil has 20 arches. The 3 bump foils have the same width, and two 1mm gaps are manufactured between each bump. The foils are clamped in the base in trailing edges, and the leading edges are free. To prevent lamination during installation and operation, the wrap angles of all foils are 350\u00b0 only. The geometrical and operational parameters of the studiedGFB are given in Table I. In the present researches, the nominal clearance ranges from 0" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0002203_icem49940.2020.9271015-Figure4-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0002203_icem49940.2020.9271015-Figure4-1.png", + "caption": "Fig. 4. The studied cooling channel topologies.", + "texts": [ + " Over half of the core losses are localized in the stator teeth. Above base speed, this share increases with increasing speed. Two bearings are used, one on each side on the shaft just inside the housing. The bearings used for loss modelling are the 61820 deep groove ball bearings from SKF [12], each with a mass of 0.3 kg. The losses in each of the two bearings are estimated using a constant coefficient of friction of 0.0015 as in [8, p.140]. The studied cooling channel topologies are presented in Fig. 4. The used wave channel makes 16 passes along the 861 Authorized licensed use limited to: University of Prince Edward Island. Downloaded on May 17,2021 at 03:11:04 UTC from IEEE Xplore. Restrictions apply. axial extent of the machine (not 12 as in Fig. 4), the spiral makes four laps around the circumferential extent and the squiggly channel makes two laps inside each end bell, with two different average diameters. With the wave and spiral cooling channels the frame thickness outside the stator back is 15 mm, whereas it is 7 mm with the squiggly channel. The wave channel size is 6.8\u00d738.8 mm and the spiral channel is 5\u00d730 mm. The width and height of the squiggly duct is 6 mm inside the 12 mm thick end bells. The thermal lumped parameter network of the reference machine with the wave and spiral cooling jackets, as shown in Fig" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0000838_012104-Figure6-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0000838_012104-Figure6-1.png", + "caption": "Figure 6. Turbine stage diagram: 1", + "texts": [ + " The DDL_IMPO function, responsible for the number of degrees of freedom, xes the movement of the high and low bounds along the Oy axis. Further, taking into account previously created material concepts and boundary conditions, the analysis type is determined MECA_STATIQUE. Using the CALC_CHAM function, the temperature stresses in the mesh nodes SIGM_NOEU are calculated. Figure 4 shows the radial, axial and hoop stresses in the cross section of the pipe. ToPME-2019 IOP Conf. Series: Materials Science and Engineering 747 (2020) 012104 IOP Publishing doi:10.1088/1757-899X/747/1/012104 The steel disk of a turbine (Fig. 6) [5] located in a stationary temperature eld is considered. The radial section of the disk is shown in Fig. 7 [6]. nozzle diaphragm, 2 turbine rotor blade, 3 shaft, 4 disk. The geometric model of the disc is built using AutoCAD's computer-aided design and drafting system. The symmetry of a rigid body with respect to the axis Ox makes it possible to construct only the upper half of the section. The result in .iges format is imported into the Salome-Meca. The MESH module allows to build a mesh using two-dimensional triangular quadratic elements (Fig" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0000896_0954406220912005-Figure1-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0000896_0954406220912005-Figure1-1.png", + "caption": "Figure 1. Variable geometry legged wheel in two configurations: (a) closed (circular wheel) and (b) open (legged wheel).", + "texts": [ + " electric,12 pneumatic,17 and hydraulic18) and power transmission mechanisms can be used to control the wheel. . The motion of the links that are in contact with the ground is made in a plane orthogonal to the rotation axis of the wheel to prevent excessive forces produced by the dragging of the limbs against the ground during their extension, as it happens in the case in which the end of the legs move out of such plane.8,15 . The hybrid wheel can have one or any number of limbs, by replicating the kinematic chain of each deployable limb. Figure 1 shows a 3RSRR HeIse wheel in open and closed position. The mechanism consists of a sliding shaft, a rim, and a series of deployable limbs (see Figure 2). The kinematic chain of one limb consists of binary links connected by the following sequence of kinematic pairs: Revolute (R1), Spherical (S), Revolute (R2), and Revolute (R3). The sliding shaft is connected with the central rim through a prismatic joint (P) that allows the shaft to move axially on the rim. On the other hand, the shaft is also connected to a link (which we will refer to it as rod, and appears in the rest of the mechanisms) of the limb through the rotational joint R1" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0002009_icma49215.2020.9233755-Figure1-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0002009_icma49215.2020.9233755-Figure1-1.png", + "caption": "Fig. 1. the structure of the snake-like manipulator.", + "texts": [ + " Firstly, the paper introduces the mechanical structure of the snake-like manipulator. Then the forward kinematics of the snake-like manipulator is analyzed by modified D-H method. Next, the improved backbone mode method is used to solve the inverse kinematics of the snake-like manipulator. Finally, the reachable space size and computational efficiency of the improved algorithm were compared with the original backbone mode method by MATLAB simulation. In this paper, the structure of the snake-like manipulator is shown in Figure 1. The snake-like manipulator is driven by ropes. The joint of the manipulator is the Cardan joint. A Cardan joint has two degrees of freedom in the manipulator. 1774978-1-7281-6416-8/20/$31.00 \u00a92020 IEEE Proceedings of 2020 IEEE International Conference on Mechatronics and Automation October 13 - 16, Beijing, China Authorized licensed use limited to: San Francisco State Univ. Downloaded on June 20,2021 at 06:21:12 UTC from IEEE Xplore. Restrictions apply. The whole manipulator is connected by 8 joints" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0002861_978-3-030-55889-5-Figure4.16-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0002861_978-3-030-55889-5-Figure4.16-1.png", + "caption": "Fig. 4.16 In 1-D problems the error is zero in any element which carries no load", + "texts": [ + " Let us study a beam, since in a beam the influence functions Gh 0 of the nodal displacements agree with the exact influence functions G0. In addition, the following applies: (i) the shape functions \u03d5i are homogeneous solutions, (i i) the influence functions are homogeneous solutions at all points\u2014except the source point, (i i i) homogeneous solutions are defined by their nodal values. So, the error in an FEinfluence function Gh 0 is zero outside of the element with the source point x , and a 1-D FE-solution4 is exact at all points x lying on elements with zero load, see Fig. 4.16. What happens, if x lies on an element, which carries a load p ? In each element the exact deflection w can be split into a homogeneous solution w0 and a particular solution wp corresponding to fixed ends 4Before we add the local solutions. w(x) = w0(x) + wp(x) E I w I V 0 (x) = 0 E I w I V p = p , (4.79) and the homogeneous solution is just the FE-solution,wh = w0, since the load is only applied at the nodes. The error e(x) = w(x) \u2212 wh(x) = wp(x) of the FE-solution is therefore identical with the particular solution and so the error in the bendingmoment is identical with the bending moment Mp of the particular solution, (= the moment if the element were clamped on both sides, le = element length) Mp(x) = \u2212E I w\u2032\u2032 p(x) = p l2e 2 ( 1 6 \u2212 x le + x2 l2e ) " + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0001889_icra40945.2020.9196995-Figure10-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0001889_icra40945.2020.9196995-Figure10-1.png", + "caption": "Fig. 10: SEA-based 3-DOF manipulator.", + "texts": [ + " The end effector of the SEA will overshoot the reference position as long as the input energy of the current cycle is equal to the output energy of the previous cycle: EO = EI = \u2206T n \u2211 k=0 (\u03c4(k)+\u2206O)(q\u0307re f (k)\u2212 q\u0307(k)) (11) Therefore, larger values of \u2206O will result in smaller overshoots and vice versa. Even for higher Kp, if a feed-forward offset is introduced in the pressing path of each cycle, the system would exhibit a faster convergence. This phenomenon is shown experimentally in the following section where it is also compared to a best tuned impedance controller and TDPA. Fig. 10 shows the SEA-based 3-DOF manipulator that we have used for our experimental study. The SEA joint consists of torsion-spring with 1700 Nm/rad spring constant after 1:100 gear reduction, which can be converted into approximately 10.6 kN/m spring constant in a neutral Cartesian space. Please note that Cartesian space impedance controller is designed and implemented to the manipulator position tracking control. Fig. 11 shows the position tracking capability of the normal impedance controller with different stiffness gain (Kp) of desired Cartesian impedance" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0001210_j.matpr.2020.04.491-Figure5-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0001210_j.matpr.2020.04.491-Figure5-1.png", + "caption": "Figure 5: Finite element model of deep groove ball bearing and new bearing design", + "texts": [ + "1 \u2022 The model is developed with zero radial clearance. Geometrical inaccuracies are not simulated. Bearing cage is eliminated for this simulation. \u2022 Lubrication, contamination and temperature effects are not considered. The 3D CAD model of the bearings are as shown in Figure 4. Finite element modelling is carried out using hexahedral and tetrahedral elements. At the contact zone, mesh density is maintained with element size of 20 micrometres, and other portion has much lower density to optimize calculation run time. These models are as shown in Figure 5. Figure 7. Contact Pressure plots of DGRB H207 Frictional contacts with friction value of 0.1 (Yellow coloured) Bonded contacts (Green coloured) Description Inner Ring Rolling Element Outer Ring Radial deformation (mm) 0.0646 0.048 0.0166 Stress Von Mises MPa 2183 1966 1801 Contact Pressure MPa 3601 - 3502 Contact ellipse \u201ca\u201d (Semi major axis) mm 3.40 - 2.5 Contact ellipse \u201cb\u201d (Semi minor axis) mm 0.25 - 0.349 The loads applied on each of the bearing are different and arrived based on Bearing calculation program" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0002321_012009-Figure7-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0002321_012009-Figure7-1.png", + "caption": "Figure 7. Change in the angular velocity of the shaft from the angle of rotation when the shaft is decelerated by the cam mechanism", + "texts": [ + " The rotation angle is measured by changing the magnetic field around the transducer. Using the model of the device the experimental data were obtained. The device was tested in two modes. In the first mode, the shaft was rotated at a constant speed without the influence of artificially created defects. In the second mode, the rotation of the shaft was artificially slowed when it was rotated through a predetermined angle using the cam mechanism. The result of measurement is shown in Figures 6 and 7. The graph in Figure 7 shows a decrease in angular velocity during rotation in the range from 175 to 185 degrees, this change is due to the fact that when the shaft rotates, the cam mechanism reduces angular velocity, this method introduces uneven rotation of the shaft. 4. Conclusion Computer-Aided Technologies in Applied Mathematics Journal of Physics: Conference Series 1680 (2020) 012009 IOP Publishing doi:10.1088/1742-6596/1680/1/012009 By measuring the non-uniformity of the rotation of the shaft of piston machines it is possible to collect a volume of information about the dependence of the change in the angular velocity of the machine shaft at various angles of rotation from emerging defects" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0002521_ei250167.2020.9346917-Figure2-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0002521_ei250167.2020.9346917-Figure2-1.png", + "caption": "Fig. 2 Basic structure of the high speed permanent magnet synchronous machine (HS-PMSM)", + "texts": [ + " When energy storage is required, the high speed permanent magnet synchronous machine (HS-PMSM) works in the motor mode, and drives the flywheel to accelerate, thus the electrical energy is converted into kinetic energy. When energy needs to be released, the HS-PMSM works in the generator mode, and the flywheel decelerates, thus the kinetic energy is converted into electrical energy. Therefore, HS-PMSM is the hub of energy conversion of the magnetic levitation flywheel energy storage system. 2734 Authorized licensed use limited to: Miami University Libraries. Downloaded on June 15,2021 at 08:41:38 UTC from IEEE Xplore. Restrictions apply. Figure 2 shows the structure of the proposed high speed permanent magnet synchronous machine (HS-PMSM). From the point of view of structure, the HS-PMSM is composed of a stator and a rotor. While in terms of function, the HS-PMSM consists of several components, i.e., permanent magnet, sleeve, windings, and core. The permanent magnet, together with the sleeve, construct the rotor of HS-PMSM. While the stationary parts including windings and core construct the stator of HSPMSM. For the purpose of reducing current frequency in thewindings and magnetic field frequency in the machine, the number of poles of high speed machine rotor is normally designed to be two or four" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0000254_ecce.2019.8911883-Figure1-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0000254_ecce.2019.8911883-Figure1-1.png", + "caption": "Fig. 1. Cross section. (a) Conventional structure. (b) Proposed structure.", + "texts": [ + " As mentioned in [20], the partitioned-stator FSPM machine can improve the torque density and decrease the copper loss. Based on tooth-slot structure, the FSPM machine can produce 50% higher torque than that of regular PM machine [9]. However, for high pole ratio FSPM machine, the flux barrier effect was noticed and has been proved to weaken the flux modulation effect and torque capability [21]. In order to overcome this drawback, a FSPM machine with alternative flux bridges is proposed in this paper, as shown in Fig. 1. (b). The flux bridges are put on the top of stator yoke as \u201cC\u201d \u2013 shaped, which can provide flux path for the working harmonic field and weaken the flux barrier effect. The organization of this paper is as follow. In Section II, the origination and working principle of the proposed FSPM machine will be introduced and the influence of pole ratio on flux barrier effect will be analyzed. Then, in Section III, three proposed FSPM machines with different rotor tooth numbers will be investigated in torque characteristics considering the effects of split ratio and flux bridge width factor", + " From the viewpoint of the flux path, the reversely excited PMs are flux barriers, so this phenomenon is called as \u201cflux barrier effect\u201d [21]. One can find that the higher pole ratio, the more reversely excited PMs, and thus the severer \u201cflux barrier effect\u201d and lower torque density. Therefore, this paper will aim to solve the \u201cflux barrier effect\u201d of FSPMs and increase the torque density at high pole ratios. In order to solve the problem in high pole ratio FSPM machine, a novel FSPM machine with alternative flux bridges is proposed. As shown in Fig. 1 (b), the proposed FSPM machine has the same combination of stator and rotor slots, winding pole pair and reluctance rotor with the conventional FSPM machine. The major difference is the stator core configuration. In the conventional FSPM machine, the magnets are inserted into the \u201cU\u201d-shaped stator iron and the adjacent spoke-array PMs have different polarities. In the proposed FSPM machine, the magnets are circumferential magnetized and the flux bridges are placed at the top of coupled magnets as \u201cC\u201d-shaped in the stator" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0003051_icelmach.2018.8507049-Figure7-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0003051_icelmach.2018.8507049-Figure7-1.png", + "caption": "Fig. 7. Concentrated windings for 2pole 24-slot 6PIM: (a) single-layer CW 10-slot pole pitch, (b) single-layer CW 6-slot pole pitch, (c) single-layer CW 2-slot pole pitch, (d) two-layer selected FSCW with q=2/13.", + "texts": [ + " The fundamental slip value is defined as the function of synchronous speed (nsyn) and rotation speed (nrot): 1 syn rot syn n n S n \u2212 = (9) The slip value for the forward-rotating space harmonics with h=6k+1 is computed as in (10), while the slip-value for the backward-rotating space harmonics with h=6k-1 is computed as in (11) [20]. 1(1 ) syn rot syn h rot h synsyn h n nn n hS h hS nn h \u2212 \u2212 \u2212\u2212 = = = \u2212 + (10) 1(1 ) syn rot syn h rot h synsyn h n nn n hS h hS nn h \u2212 \u2212 ++ = = = + \u2212 (11) This part discusses about the results of using winding function to study the concentrated windings on the baseline induction motor. 1) Single-layer non-overlapping CW layouts In this analysis three single-layer non-overlapping concentrated winding layouts with different pole-pitch values have been considered; they are as shown in Fig. 7(a), (b) and (c), respectively. For these three layouts, the harmonic spectrums of the air gap MMF and the winding factors are calculated analytically by using the WFA. The normalized results (with respect to the fundamental harmonic of the 2- pole DW solution) are presented in Table III. By considering these results, among the three alternatives, the single-layer CW with 6-slot pole pitch features the higher fundamental harmonic amplitude and the smoother harmonic spectrum of Bg, and higher winding factor. Therefore, the single-layer CW with 6-slot pole pitch (Fig. 7 (b)) is the selected singlelayer non-overlapping concentrated winding to compare with other CWs. This winding layout is analyzed by using the WFA and FEA. Fig. 8 and Fig. 9 show the analytical results for the air gap MMF distribution and its harmonic spectrum, respectively. The winding function results, supported by FEA, show that the selected single-layer concentrated winding has a fundamental harmonic value of the air gap flux density that is far below than that of the 2-pole distributed winding", + " In order to have the comparison between these FSCW layouts, the fundamental component of the air gap flux density and the winding factor are calculated and reported in Table IV for the 2-layer layouts. Please note that all the results of this section are normalized with respect to the fundamental harmonic of the 2-pole DW solution. Only layouts that generates a 2-pole in their air gap MMF (PIM=2), are candidate to be selected for the comparison. As it can be seen in Table IV, the two-layer configuration with q=2/13 has the higher first harmonic value of both air gap flux density and winding factor comparing to the other layouts with PIM=2. Hence, this concentrated winding (Fig. 7 (d)) is chosen as the selected 2-pole layout among all the considered two-layer non-overlapping concentrated windings. The normalized results of the magnetic analysis for the selected winding are illustrated in Fig. 10 and Fig. 11. Nevertheless, comparing the harmonic spectra of Fig. 4 with that of Fig. 11, it is evident that the CW layout features a much lower MMF fundamental harmonic compared to the baseline 2-pole distributed winding configuration. For this reason, it is expected that also the torque-speed characteristic of the machine equipped with this 2-layer concentrated winding cannot compete at all with that produced by the baseline machine with DW" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0003736_icems.2014.7013521-Figure3-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0003736_icems.2014.7013521-Figure3-1.png", + "caption": "Figure 3 Models of permanent magnet shifting for a 48-slot 16-pole and 40- slot 16-pole motor", + "texts": [ + "75 degrees that relative to the P1 and other odd poles; The second offset is based on the first offset, each consisting of a set of four permanent magnets , P1 and P2 remain immobile while P3 and P4 are offset 3.75 degrees that relat ive to the P1 and P2. Considering the limited space of the pole shift in the third offset, select P16 ~ P7 as be mobile group and P8 ~ P15 be not moving group, and offset P16 ~ P7 3.75 degrees that relative to P8 ~ P15. The offset method of 16 pole 40 slot is similar to 16 pole 48 slot offset method, the offset angle is shown in Table . After the pole shift fin ite element models by applications Ansoft / Maxwell simulat ion software are shown in Fig.3. Analysis of cogging torque by Ansoft/Maxwell is shown in Fig.4. Before the pole shift, the cogging torque maximum minimum reduction value is 2.945N.m, and it is reduced to 1.498N.m after the first migrat ion, which reduced by 49.13%; Then it has been reduced to 0.73N.m after the second shift, which reduced by 75.21%; finally it is reduced to 0.572 N.m after the third times, which reduced by 80.72%. W ith the increase of the times of offset, reducing the amplitude leveled off. It also shows that a larger proportion of the motor cogging torque is the fundamental share" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0000603_iccas47443.2019.8971558-Figure6-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0000603_iccas47443.2019.8971558-Figure6-1.png", + "caption": "Figure 6. TakoBot 2 kinematic structure", + "texts": [], + "surrounding_texts": [ + "219\nresults. The new prototype supposes to improve robot dexterity and reachability capabilities. In chapter II will explain the concept design, III chapter will explain kinematic kinetic formulation, IV chapter will present experiments and simulation results.\nII. CONCEPT DESIGN\nA. Application Analysis The intended application of the robot is shown in Fig.1. The manipulator is designed to be used in the agriculture field for harvesting, weeding and inspection operations if necessary. The intended working environment is highly constrained and by the time workspace will be different because of growing plants.\nTo perform necessary tasks the robot should have the following functions:\n(1) Flexible dexterity: the robot should work in the confined workspace\n(2) Obstacle avoidance capability: the robot should avoid contact with solid surfaces and not collide.\n(3) Reachability: the robot should get the required position in spite of narrow space and obstacles.\n(4) Safety: the robot should be safe enough to avoid breaking any sticks of the plant.\n(5) Portability: the robot should be compact to be used as a tool for farming CNC platform (Fig.1).\nB. System Design Based on application analysis, the robot should have a slender structure with redundant DOFs. After consideration of previous TakoBot 1 prototype capabilities, we made changes in the design and actuation system as well. After numerous experiments with TakoBot 1, demonstrated some limitations related to the design. For example, a robot could not perform torsional motion during the work which led to the accumulation of strain energy inside of manipulator. Secondly, in case of bending it could not perform a pure bending shape. In the new prototype, we added a passive sliding mechanism (Fig.4). Proposed sliding mechanism demonstrated the following benefits:\n1) Smart bending stress distribution: This ability helped to the manipulator to bend purely and it distributed spring potential energy to the whole segment of the manipulator. In comparison with previous, it required less torque and less load to the cables. 2) Torsional motion: passive sliding disc mechanism provided better dexterity ability in case bending helical motion. Helical motion increases torsional stress of the structure unless a disc does not generate yawing around the z-axis. In spite of the heavier weight of the TakoBot 2 was able to perform better and more accurate.\nPassive sliding disc mechanism works for all discs attached to the backbone except base and end-effector discs, which means total length always remains constant. Fig.4. demonstrates a sliding mechanism structure. According to the design, travelling distance of the disc is 10mm, while the disc diameter is 50mm, an average distance of the single section (between discs) about 35 mm. 3D printed discs connected by coil compression springs, such design provides stiffness to the manipulator. The spring constant could be variable depends on motor torque, in this prototype we used spring with constant 0.63 N/mm. Each section consists of four segments, end segments discs connected by four springs, while mid-section segments connected by 8 springs between discs.\nAuthorized licensed use limited to: Auckland University of Technology. Downloaded on October 25,2020 at 08:37:21 UTC from IEEE Xplore. Restrictions apply.", + "220\nC. Actuating mechanism Manipulator's tendons actuate by a linear lead shaft\nconnected to the stepper motor by a coupler. Tendons attached to both sides of specially designed inlets on screw housing and screw with housing travels along the shaft. 1mm steel cable utilized as a tendon.\nIII. KINEMATICS AND KINETIC FORMULATION\nA. Forward kinematic formulation Coordinate systems are set at every universal joint. The homogeneous coordinate transform matrices:\n\u03a30\u2192\u03a31 , H0,1= ( R u0,1 0 0 0 1 ), u0,1=( x0 y0 l0 )\n(1)\n\u03a3i-1\u2192\u03a3i, Hi-1, i= ( R ui-1,i 0 0 0 1 ) , ui-1,i=( 0 0 L ),\n(i=2, \u22ef, n)\n(2)\nR= Rz(\u03b8zi)Rx(\u03b8xi)Ry(\u03b8yi) (3)\nwhere, \ud835\udc65\ud835\udc650 and \ud835\udc66\ud835\udc660 are an initial position of the base. \ud835\udc45\ud835\udc45\ud835\udc65\ud835\udc65(\ud835\udf03\ud835\udf03\ud835\udc65\ud835\udc65\ud835\udc65\ud835\udc65) and \ud835\udc45\ud835\udc45\ud835\udc66\ud835\udc66(\ud835\udf03\ud835\udf03\ud835\udc66\ud835\udc66\ud835\udc65\ud835\udc65) are rotation matrices of ith universal joint that has two rotation angles \ud835\udf03\ud835\udf03\ud835\udc65\ud835\udc65\ud835\udc65\ud835\udc65 and \ud835\udf03\ud835\udf03\ud835\udc66\ud835\udc66\ud835\udc65\ud835\udc65, \ud835\udc45\ud835\udc45\ud835\udc67\ud835\udc67(\ud835\udf03\ud835\udf03\ud835\udc67\ud835\udc671) is a rotation matrix of the ith disk with a rotation angle \ud835\udf03\ud835\udf03\ud835\udc67\ud835\udc671 along the axial axis and L is a length between neighbouring universal joints. Three rotation matrices have:\nMultiplying the H-matrices successively, we get unit\nvectors and the position vector of the ith coordinate system; H0,i=H0,1H1,2\u22efHi-1, i= (\nii ji ki ui 0 0 0 1 ) (4)\nwhere, \ud835\udc62\ud835\udc62\ud835\udc65\ud835\udc65 is the position of the ith universal joint Ui (\ud835\udc56\ud835\udc56 =\n1,\u22ef , \ud835\udc5b\ud835\udc5b \u2212 1). The position vector \ud835\udc5d\ud835\udc5d\ud835\udc65\ud835\udc65 of the end-point P\ud835\udc5b\ud835\udc5b and position of sliding plates Pi (i=1, \u22ef, n-1) of the manipulator are obtained by, (pi\n1)=H0,i(0 0 li 1)T, (i=1, \u22ef, n) (5)\nwhere, \ud835\udc59\ud835\udc59\ud835\udc5b\ud835\udc5b is a fixed length between the nth universal joint and the most distal plate. Position vectors of 8 hole A0, B0, C0, D0 , A\u03020, B\u03020, C\u03020, D\u03020 at the base plate are determined as,\na0=( ax ay 0 ) , b0=( bx by 0 ) , c0=( cx cy 0 ) , d0=( dx dy 0 ) ,\na\u03020=( a\u0302x a\u0302y 0 ) , b\u03020=( b\u0302x b\u0302y 0 ) , c\u03020=( c\u0302x c\u0302y 0 ) , d\u03020=( d\u0302x d\u0302y 0 ) ,\n(6)\nPosition vectors of 4 hole A\ud835\udc65\ud835\udc65, A\u0302\ud835\udc65\ud835\udc65, C\ud835\udc65\ud835\udc65, C\u0302\ud835\udc65\ud835\udc65 at the ith plate\n(\ud835\udc56\ud835\udc56 = 1,\u22ef , \ud835\udc5b\ud835\udc5b) are obtained as\n(ai 1)=H0,i( ax ay li 1 ) , (a\u0302i 1)=H0,i( a\u0302x a\u0302y li 1 ) ,\n(ci 1)=H0,i( cx cy li 1 ) , (c\u0302i 1)=H0,i( c\u0302x c\u0302y li 1 ) , (i=1, \u22ef, n)\n(7)\nwhere, \ud835\udc59\ud835\udc59\ud835\udc65\ud835\udc65 is an axial length between the ith universal joint and the ith plate, which varies as the plate slides along rods, except \ud835\udc59\ud835\udc59\ud835\udc5b\ud835\udc5b.\nIn the same way, position vectors of 4 hole B, B\u0302\ud835\udc65\ud835\udc65, D\ud835\udc65\ud835\udc65, D\u0302\ud835\udc65\ud835\udc65 at the ith plate (\ud835\udc56\ud835\udc56 = 1,\u22ef ,\ud835\udc5a\ud835\udc5a) are obtained as,\n(bi 1)=H0,i( bx by li 1 ) , (b\u0302i 1 )=H0,i ( b\u0302x b\u0302y li 1) ,\n(di 1)=H0,i( dx dy li 1 ) , (d\u0302i 1 )=H0,i ( d\u0302x d\u0302y li 1) , (i=1, \u22ef, m)\n(8)\nB. Kinetic formulation Our continuum manipulator is divided by two segments.\nAuthorized licensed use limited to: Auckland University of Technology. Downloaded on October 25,2020 at 08:37:21 UTC from IEEE Xplore. Restrictions apply.", + "221\nThe first segment is located by the distal part and the second segment is located by the proximal part. The first segment is operated by 2 pairs of 2 wires; total 4 wires. One pair of 2 wires is controlled by one motor that pulls one wire and push the other wire in the same length by using the pulley. While, the second segment is operated by 4 pair of 4 wires, total 8 wires. Therefore the second segment is controlled by 4 motors The second segment has m units and the first segment has n-m units. 4 pairs of wires are labeled by \ud835\udc4e\ud835\udc4e and ?\u0302?\ud835\udc4e, \ud835\udc4f\ud835\udc4f and ?\u0302?\ud835\udc4f, \ud835\udc50\ud835\udc50 and ?\u0302?\ud835\udc50, d and ?\u0302?\ud835\udc51,\nEquilibrium in moments at Un belonging to the first segment is (Sa,n-fa)(an-an-1)\u00d7(an-un)+(Sa\u0302,n-fa\u0302)(a\u0302n-a\u0302n-1)\u00d7(a\u0302n-un)+\n(Sc,n-fc)(cn-cn-1)\u00d7(cn-un )+(Sc\u0302,n-fc\u0302)(c\u0302n-c\u0302n-1)\u00d7(c\u0302n-un)++\nmw(pn-un)\u00d7g= ( 0 0 0 )\n( (9)\nwhere, an-an-1= an-an-1\n|an-an-1| , etc. m\ud835\udc64\ud835\udc64 is a payload applying at the\nend-point and g is the gravity acceleration vector. Equilibrium in moments at Ui , (i=m+1, \u22ef, n-1), belonging to the first segment is (Sa,n-fa)(an-an-1)\u00d7(an-un)+\n(Sa\u0302,n-fa\u0302)(a\u0302n-a\u0302n-1)\u00d7(a\u0302n-un)+ (Sc,n-fc)(cn-cn-1)\u00d7(cn-un )+ (Sc\u0302,n-fc\u0302)(c\u0302n-c\u0302n-1)\u00d7(c\u0302n-un)+\nmp \u2211 (pk-ui) n\nk=i+1\n\u00d7g= ( 0 0 0 )\n(10)\n(10)\nwhere. fa, fa\u0302, fc, fc\u0302 are wire tensions, Sa,i, Sa\u0302,i, Sc,i, Sc\u0302,i , (i=m+1, \u22ef, n) are spring tensions of the ith unit. \u201c\u00d7\u201d means a cross product and \u201c|*|\u201d, means the modulus of a vector \u2217. \ud835\udc5a\ud835\udc5a\ud835\udc5d\ud835\udc5d is the mass of one unit including the plate, the rod and the universal joint.\nThe spring tensions are obtained as, Sa,i\uff1dk(L-|ai-ai-1|), Sa\u0302,i\uff1dk(L-|a\u0302i-a\u0302i-1|),\nSc,i=k(L-|ci-ci-1|), Sc\u0302,i=k(L-|c\u0302i-c\u0302i-1|),)\n(11)\nwith spring coefficient k. Equations (9) and (10) contain 3(nm) equations including 4(n-m) -1variables of the n-m universal joints angles \u03b8xi, \u03b8yi, \u03b8zi, (i=m+1, \u22ef, n) and slide length of plates \ud835\udc59\ud835\udc59\ud835\udc56\ud835\udc56 (i=m+1, \u22ef, n-1). Equilibrium in force at ith plate (i=m+1, \u22ef, n-1) is,\n[-Sa,i+1(ai+1-ai) +Sa,i(ai-ai-1)+-Sa\u0302,i+1(a\u0302i+1-a\u0302i)+ Sa\u0302,i(a\u0302i-a\u0302i-1)-Sc,i+1(ci+1-ci)+Sc,i(ci-ci-1)-Sc\u0302,i+1(c\u0302i+1-c\u0302i))\nSc\u0302,i(c\u0302i-c\u0302i-1)+(n-i)mpg ]\u2219 (pi-ui)=0\n(12)\n(12) provide n-m-1equations. Combined it with (9) and (10), we obtain 4(n-m)-1 equations, which suffices in number to solve for 4(n-m)-1 variables; \u03b8x,i, \u03b8y,i , \u03b8z,i (i=m+1,\u22ef,n) and li (i=m+1, \u22ef, n-1) for a given set of wire tensions fa, fa\u0302, fc, fc\u0302.\nEquilibrium in moments at Um , the universal joint located at the most distal position belonging to the second segment is\n-Sa, m+1(am+1-am)\u00d7(am-um)+(Sb,m-fb)(bm-bm-1)\u00d7(bm-um)\n-Sa\u0302,m+1(a\u0302m+1-a\u0302m)\u00d7(a\u0302m-um)+(Sb\u0302,m-fb\u0302) (b\u0302m-b\u0302m-1) \u00d7(b\u0302m-um) -Sc, m+1(cm+1-cm)\u00d7(cm-um)+(Sd,m-fd)(dm-dm-1)\u00d7(dm-um)\n-Sc\u0302,m+1(c\u0302m+1-c\u0302m)\u00d7(c\u0302m-um)+(Sd\u0302,m-fd\u0302) (d\u0302m-d\u0302m-1) \u00d7(d\u0302m-um)\n+ (mw(pn-um)+mp \u2211 (pk-um) n-1\nk=m+1\n) \u00d7g= ( 0 0 0 )\n(13)\nFor the second segment, we can derive similar equations as (10), (11) and (12) by replacing {ai, a\u0302i, ci, c\u0302i} with {bi, b\u0302i, di, d\u0302i}, {Sa,i, Sa\u0302,i, Sc,i, Sc\u0302,i} with {Sb,i, Sb\u0302,i, Sd,i, Sd\u0302,i} for i=1, \u22ef, m-1 in (10) and for i=1, \u22ef, m in (11) and (12).\nAs a result, we obtain 4m equations included by (13), which suffices in number to solve for 4m variables; \u03b8x,i, \u03b8y,i, \u03b8z,i and li (i=1, \u22ef,m) for a given set of wire tensions \ud835\udc53\ud835\udc53\ud835\udc4f\ud835\udc4f, \ud835\udc53\ud835\udc53?\u0302?\ud835\udc4f, \ud835\udc53\ud835\udc53\ud835\udc51\ud835\udc51, \ud835\udc53\ud835\udc53?\u0302?\ud835\udc51.\nWire tensions fa, fa\u0302, fc, fc\u0302, fb, fb\u0302, fd, fd\u0302. are determined according to 4 motors\u2019 angles \u03d5a, \u03d5b,\u03d5c,\u03d5d As\nfa=kp ( \u03bb (\u03d5p+\u03d5a) 2\u03c0 -nL+ \u2211|ai-ai-1| n\ni=1\n) ,\nfa\u0302=kp ( \u03bb (\u03d5p-\u03d5a) 2\u03c0 -nL+ \u2211|a\u0302i-a\u0302i-1| n\ni=1\n) ,\nfc=kp ( \u03bb (\u03d5p+\u03d5c) 2\u03c0 -nL+ \u2211|ci-ci-1| n\ni=1\n) ,\nfc\u0302=kp ( \u03bb (\u03d5p-\u03d5c) 2\u03c0 -nL+ \u2211|c\u0302i-c\u0302i-1| n\ni=1\n) ,\nfb=kp ( \u03bb (\u03d5p+\u03d5b) 2\u03c0 -nL+ \u2211|bi-bi-1| m\ni=1\n) ,\nfb\u0302=kp ( \u03bb (\u03d5p-\u03d5b) 2\u03c0 -nL+ \u2211|b\u0302i-b\u0302i-1| m\ni=1\n) ,\nfd=kp ( \u03bb (\u03d5p+\u03d5d) 2\u03c0 -nL+ \u2211|di-di-1| m\ni=1\n) ,\nfd\u0302=kp ( \u03bb (\u03d5p-\u03d5d) 2\u03c0 -nL+ \u2211|d\u0302i-d\u0302i-1| n\ni=1\n) ,\n(14)\nwhere, \ud835\udf19\ud835\udf19\ud835\udc5d\ud835\udc5d is a motor rotation angle to generate a pretension, \ud835\udf06\ud835\udf06 is a lead of the screw rod and \ud835\udc58\ud835\udc58\ud835\udc5d\ud835\udc5d is the spring constant of the pretension spring.\nAuthorized licensed use limited to: Auckland University of Technology. Downloaded on October 25,2020 at 08:37:21 UTC from IEEE Xplore. Restrictions apply." + ] + }, + { + "image_filename": "designv11_71_0000915_012029-Figure1-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0000915_012029-Figure1-1.png", + "caption": "Figure 1. Structure of the V-ribbed belt.", + "texts": [ + " In response to the country's promotion of new energy vehicles and further reduction of production costs, EPDM is increasingly being applied to power systems such as engine cushions and transmission belts [1, 2]. Generally, automotive transmission belts are divided into V-belts, flat belts, and V-ribbed belts. The V-ribbed belt combines the advantages of flat belt and V-belt. In the overall structure, the main base is flat belt, and a plurality of V-shaped wedges is closely arranged inside the contact with a gear such as a gear. Therefore, it has the stability of the flat belt, the firmness and the transmission performance of the V belt. Its structural composition is shown in figure 1, which mainly includes four parts: top cloth, cushion rubber, tensile body and wedge rubber. EPDM is a terpolymer prepared by solution polymerization with ethylene- propylene as main monomers and addition of unsaturated monomers (non-conjugated diolefins). It is a non-polar saturated carbon chain rubber which is chemically stable and has excellent resistance to ozone and weathering, but it has low strength and poor self-adhesiveness [3]. In order to prepare a compound with superior comprehensive properties, it is often necessary to modify it" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0001960_ro-man47096.2020.9223579-Figure6-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0001960_ro-man47096.2020.9223579-Figure6-1.png", + "caption": "Fig. 6: The paths of evolution of proposed formulation compared to the classical DMP, in the case of spatial scaling.", + "texts": [ + "7445 0]Tm, having a rotation of 180o around zaxis with respect to the inertial frame. The motion consists of a simple arc-like pattern on the part\u2019s surface, which is kinesthetically demonstrated by the user as shown in Fig. 5. Based on this motion demonstration, the following DMPs are trained utilizing 60 kernels and LS for solving (4): the proposed model, a classical DMP in the Cartesian space, and a b-DMP in the Cartesian space. To compare the methods with respect to their spatial generalization properties, a different target pose is given for the execution of the task. Fig. 6 and 7 depicts the evolution of the classical DMP formulation and the b-DMP respectively, compared to the proposed formulation. Notice that the constraint is only respected by the evolution of the proposed implementation. In Fig. 8, the tool axis during the evolution of the proposed DMP and b-DMP is depicted. Notice that the tool axis during the execution of the b-DMP does not satisfy the constraint, which is to remain orthogonal to the surface. In this work, a dynamical system based learning by demonstration approach is proposed for constrained tasks" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0000338_iecon.2019.8926794-Figure8-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0000338_iecon.2019.8926794-Figure8-1.png", + "caption": "Fig. 8. Measurement of turning angle", + "texts": [ + " We experimented whether an Ackerman link implemented in the car functioned well. The angle at which the tire was turned to maximum was defined as the maximum turning angle. In that state, we shot an image from directly above the tire. From this image, with the pixel at the center of the tire width as coordinates, normal vector was determined from Pt1 and Pt2 of the two points at both ends of the tire. In addition, the normal vector, which is the vertical reference of the inclination, was also obtained from Pb1 and Pb2 of the two vertical parts of the car (Figure 8). In order to obtain tire movement plots, we created the field as shown in Figure 9. The radius of the circumference circle drawn on the experimental field was 590 mm, which is equal to the minimum turning radius of the car. That is, this circle is the ideal trajectory of the outer tire of the car at the time of maximum turning. In addition, scales were placed on this field as shown in Figure 10 at equal intervals. The scale used in the experiments graduated in millimeters from zero to seven. Finally, a smartphone camera was placed just above the outer tire of the car as shown in Figure 11" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0001178_isef45929.2019.9097079-Figure1-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0001178_isef45929.2019.9097079-Figure1-1.png", + "caption": "Fig. 1. Proposed RBMLS topology with discretized semi-rings forming approximated helical structures [9].", + "texts": [ + " Nevertheless, for long-stroke applications, these two topologies are not economically suitable, since they contain PMs on both movable parts. The reluctance-based magnetic lead screw (RBMLS) topology has PMs only in one of its movable parts, making them attractive for long-stroke applications. Since the experimental data for this topology is absent in literature to date, this paper proposes to experimentally test a realized prototype according to the topology proposed in [9]. The proposed RBMLS is displayed in Fig. 1 and its main dimension are listed in Table I. This topology comprehends two main structures moving in synchronism: a translator moving back and forth along the z-axis and a rotor rotating about the same axis. While the rotor is formed by a doublestart thread lead screw made of ferromagnetic material, the translator is composed by a set of cells arranged in such a way to form approximated helices. Each cell is composed of an axially magnetized PM semi-ring embedded in two other semi-ring pieces made of soft ferromagnetic material" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0000967_0954406220916504-Figure4-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0000967_0954406220916504-Figure4-1.png", + "caption": "Figure 4. Wrenches w\u0302C and w\u0302Y acting on the vehicle body.", + "texts": [ + " The steady state16 implies constant path radius of curvature, constant translation speed, constant angular speed and constant magnitude of lateral acceleration, although the velocity and acceleration are not constant in direction relative to the ground. Then, both the speed v and the angular velocity ! of the car are constant. Hence, the longitudinal acceleration along the X-axis and angular acceleration about the Z-axis are negligible. For the vehicle traveling forward with a constant speed v, the sum of side forces must be equal to the centrifugal force, i.e. FY \u00bc FC. Considering a moment equilibrium of the vehicle about the Z-axis, it becomes clear that RY \u00bc 0. Therefore, the sum of the tyre side forces w\u0302Y, as shown in Figure 4, can be rewritten as w\u0302Y \u00bc FC 0 1 0 0 0 0 T . Consequently, the sum of the wrenches w\u0302C and w\u0302Y becomes w\u0302 \u00bc w\u0302C \u00fe w\u0302Y \u00bc FC 0 0 0 Zcg 0 0 T \u00f012\u00de and acts on the vehicle body during cornering. The wheel of the vehicle maintains contact with the road surface at the contact point P, as shown in Figure 5. Since the wheel can slide and rotate about P, the contact may be modelled as a kinematic pair consisting of a plane pair (E-joint) and a spherical pair (S-joint). E-joint can be represented by two intersecting P-joints with two twists S\u0302Ex \u00bc 1 0 0 0 0 0 T and S\u0302Ey \u00bc 0 1 0 0 0 0 T ", + "9 When the wheel is steered, only the position of the tie-rod in the double-wishbone suspension is changed. Here, in order to investigate the effects of the initial positions of the spatial suspension mechanisms the steering angle is neglected. If the steering angle is considered in the analysis, it only needs to determine the position of the tie-rod at the beginning. Analysis procedure of full vehicle model This example begins with the computation of the roll motion of the vehicle body at the initial state as follows. Referring to Figure 4, the centrifugal force acting through the mass centre of the vehicle body during cornering is given by w\u0302c \u00bc 3000 0 1 0 900 0 0 T and the sum of the tyre side forces FYi can be expressed by w\u0302Y \u00bc 3000 0 1 0 0 0 0 T The resultant wrench caused by the cornering motion acting on the vehicle body is obtained from equation (12) w\u0302 \u00bc w\u0302c \u00fe w\u0302Y \u00bc 3000 0 0 0 900 0 0 T We first find the screw-based Jacobian Jp of the vehicle model given by equation (29). The four line vectors r\u0302Hi (i \u00bc 1, . . . , 4) in equation (24) are perpendicular to the ground and pass through the contact points between the wheels and the ground" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0003634_icems.2014.7013745-Figure1-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0003634_icems.2014.7013745-Figure1-1.png", + "caption": "Fig. 1. The structure of PMRM", + "texts": [ + "78-1-4799-5162-8/14/$31.00 \u00a92014 IEEE I. INTRODUCTION Many types of double-rotor motors in recent years are studied such as double-rotor permanent magnet (PM) machine[1], double-rotor switched reluctance motor[2], and double-rotor induction motor[3], to take place of the planetary gear in electrical variable transmission. In this paper, a doublerotor permanent magnet reluctance motor (PMRM) is proposed. The structure of PMRM is shown in Fig. 1. As Fig. 1 shows, the PMRM has two motors named outer motor and inner motor. Outer motor is composed of stator and outer rotor which is similar to a doubly salient permanent magnet motor. And as well, inner motor is composed of outer rotor and inner rotor which is similar to a permanent magnet synchronous motor. When PMRM works, inner motor and outer motor convey energy though inner air-gap and outer airgap. Compared with other double-rotor motors, PMRM has a better heat dissipation structure because of outer doubly salient structure, as well as, higher power density and higher efficiency because of permanent magnets" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0002542_iros45743.2020.9340900-Figure2-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0002542_iros45743.2020.9340900-Figure2-1.png", + "caption": "Fig. 2. Model of a single joint of the robot: a flexible element links a body of mass m to the rotor shaft.", + "texts": [ + " Given that the controller 3439 Authorized licensed use limited to: Univ of Calif Santa Barbara. Downloaded on June 15,2021 at 22:48:49 UTC from IEEE Xplore. Restrictions apply. structure (10) does not include any coupling between joints, we simplify the dynamics by considering that all other joints are fixed. Thus, the model reduces to a simple planar serieselastic actuator: motor i is fixed, and is linked to a single body B, aggregating all links downstream of joint i, through a spring damper. This single joint model is represented in Figure 2. Let \u03b1 be the flexibility angle (i.e. the ith component of \u03b1), and q the angle between the vertical axis and the center of mass of B, that is, the link angle qi plus a constant angle. The dynamics of the system writes J\u03b1\u0308 =\u2212 u+ \u03bdm(q\u0307 \u2212 \u03b1\u0307)\u2212 k(1 + J I )\u03b1 \u2212 \u03bd(1 + J I )\u03b1\u0307+ J I mlg sin (q) Iq\u0308 =\u2212 k\u03b1\u2212 \u03bd\u03b1\u0307+mlg sin(q) (11) where J is the rotor inertia, m and I the mass and inertia of body B, k the spring stiffness, and \u03bd, \u03bdm viscous friction coefficients for respectively the flexibility and the motor joint" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0000421_icee44586.2018.8938001-Figure13-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0000421_icee44586.2018.8938001-Figure13-1.png", + "caption": "FIGURE 13. Antenna beam width in the far field.", + "texts": [ + " To determine the ability of the ground station antenna to achieve these requirements, we must first consider the antenna as an array of radiating elements. This is depicted in Fig. 12. Note that there are M columns of line arrays each with N elements for a total of MxN elements. Thus, for an array for M = 20 and N = 20, there will be 400 elements. Also notice that there are M elements spaced apart in the x-direction by dX and N elements spaced apart in the y-direction by dY. Given the configuration in the figure, the antenna will create a beam in the far field with a beam width as defined in Fig. 13. In the direction broadside to the antenna, the antenna 3dB beam width in degrees can be approximated using BW\u03b8 = 50.76\u25e6 \u03bb NdY (2) and BW\u03c6 = 50.76\u25e6 \u03bb MdX . (3) Where, \u03bb= the wavelength at the operating frequency. Using (2) and (3), the magnitude of the directivity can be approximate by D(\u03b8, \u03c6) = 16 sin(BW\u03b8 ) sin(BW\u03c6) . (4) If the efficiency of the antenna is given by \u03b5eff, then the gain of the antenna in dB can be calculated using G(dB) = 10 log10(\u03b5effD). (5) Assuming the center of the band to be 11" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0000623_012064-Figure1-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0000623_012064-Figure1-1.png", + "caption": "Figure 1. Constructive scheme of the coupling roller bearing and its rings: (a) - constructive diagram of a typical axle box of a freight car: 1 - box housing; 2 - labyrinth ring; 3 - rear bearing; 4 - front bearing; 5 - cover fastening; 6 - tag; 7 - viewing cover; 8 - disc washer; 9 - lock washer; 10 - M20 bolt of the disc washer; 11 - M12 cover bolt; 12 - sealing ring; (b) - axial section of a fragment of the rear bearing of an axle box corresponding to one roller", + "texts": [ + " One of the effective ways to improve them is to search for rational design options based on the use of mathematical models of their stress-strain state (SSS) and experimentally verified criteria for assessing their durability [10\u201312]. Under the conditions of cyclic operation of high-loaded bearing elements of structures, such a criterion, as a rule, is the intensity of stresses [13\u201322]. Let us consider for definiteness a constructive scheme for coupling a bearing roller and its rings by the example of one of the typical axlebox nodes (figure 1) [23]. The durability of the elements of the axle assembly in many cases is determined by the moment the contact fatigue shells [24] appear on the rolling contact surfaces of the bearings and their inner and outer rings (figure 2). The determining factor in the occurrence of fatigue damage is an increased level of SibTrans-2019 IOP Conf. Series: Materials Science and Engineering 760 (2020) 012064 IOP Publishing doi:10.1088/1757-899X/760/1/012064 stress intensity in a possible focus of damage. To calculate this factor, a mathematical model of elastic deformation of the elements of the axle box can be used" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0000451_hsi47298.2019.8942634-Figure1-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0000451_hsi47298.2019.8942634-Figure1-1.png", + "caption": "Fig. 1. Our prototype", + "texts": [ + " Using two laser range scanners, our robot measures the right and left foot gait of its user and estimates its user\u2019s body balance by calculating his/her position of COG. Using estimated body balance, our proposed walker controls assistance force vector for maintaining body balance stably. Proposed system does not require sensors which are installed on its user\u2019s body for practical usage. The effectiveness of our proposed idea is verified by experiments with our prototype. II. SYSTEM CONFIGURATION A. System Overview Fig. 1 shows our proposed walker. Our walker has two actuated wheels (rear wheel) with in-wheel motor on each side. Its body size is enough small for practical usage in typical Japanese house. Its width is 550[mm] and its length is 740[mm]. Its user can change the height of the walker from 735 to 860[mm] according to its user\u2019s height. (In Fig. 1, its This work is supported in part by a Grant-in-Aid for Scientific Research C (16K01580) from the Japan Society for the Promotion of Science and the Matching Planner Program (VP29117940231) from Japan Science and Technology Agency, JST. 978-1-7281-3980-7/19/$31.00 \u00a92019 IEEE 231 Authorized licensed use limited to: University of Exeter. Downloaded on June 29,2020 at 10:30:30 UTC from IEEE Xplore. Restrictions apply. height is 860[mm].) The radius of its rear wheel (actuated wheel) is 150[mm] and its front wheel is 240[mm] for easy to small step-climbing" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0003488_msec2015-9243-Figure11-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0003488_msec2015-9243-Figure11-1.png", + "caption": "Figure 11: Calculation of a point of the middle line (Mj).", + "texts": [ + "org/about-asme/terms-of-use For generating the tool paths, a middle line is calculated (see Figure 10, top) as starting line for the equidistant tracks (see Figure 10, bottom) to reduce the amount of single tracks. This also reduces the amount of turning points in the tool path and avoids defects in these areas. Furthermore, for thin boundary lines, there is only a single track, represented by the middle line, needed for a best-fit with the geometry. This middle line is generated by calculating a middle point (Rj, see Figure 11) for each point (Pj) of the boundary. This point is lying on the line perpendicular to the tangent at this point and is located where the smallest possible concentric circle (in the 3D case it is a sphere, the 2D case will be described here to simplify the observation) around the middle point touches the trajectory itself. After calculating a middle point for every point of the boundary, a point cloud is generated. This point could is filtered and afterwards composed to a middle line. The equidistant lines themselves are generated directly onto the surface of the substrate starting from the middle line and thereby close-contoured tracks are generated" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0001770_s00170-020-05701-3-Figure7-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0001770_s00170-020-05701-3-Figure7-1.png", + "caption": "Fig. 7 Angle thinning turning operation", + "texts": [ + " The CNC code is generated and stored in (FileCNC.txt). In order to demonstrate the application of the parameterized programming module proposedmethodology, machining routines\u2019 creation examples with the hybrid generic parameterized language for specific programming cycles are presented. Although the requirements of different machining strategies are frequent, this requisite cannot be always satisfy, since the cycle possibilities offered by commercial CNC controls are limited [17]. In this context, Fig. 7 shows an unconventional machining strategy\u2014angle thinning turning operation, not found in most of the commercial CNC commands. Machining cycles normally used for this procedure correspond to longitudinal and transversal thinning. In Fig. 8, the result of the longitudinal and transversal thinning enlargement is illustrated, indicating the \u201crungs\u201d forming on the machined surface. Through Fig. 8, it is also possible to identify the Variables Definitions xinicio Diameter greater part contact area variation among tool and part, by adopting the angle thinning cycle" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0001057_0021998320920920-Figure14-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0001057_0021998320920920-Figure14-1.png", + "caption": "Figure 14. Contours of 0yz in the warps for unit-cell 2 with binders and wefts shown for context.", + "texts": [ + " Throughout all of the warps, 0xx remained the largest stress component, and 0xx was the largest contributor to failure index corresponding to LT failure, HLT, which was the maximum index for all locations except the region surrounding and including the concentrations shown in Figure 13. Where the concentrations of Hmax (refer to equation (5)) occurred, the largest index corresponded to transverse tension, HTT, which was dominated by shear stresses. The largest contributor of the shear stresses was 0yz. Figure 14 shows contours for 0yz within the warps, and the locations of 0yz concentrations match the locations of the Hmax concentrations that were shown in Figure 13. As mentioned before, a previous study considered a plain orthogonally woven textile composite with larger binders and a smaller textile thickness.25,26 However, the warps in the twill orthogonal weave studied in this paper experienced very different stress states compared to the plain orthogonal weave. In the plain orthogonal weave,25,26 the warps were shown to experience severe longitudinal shear stresses, but the twill weave considered herein did not exhibit nearly as severe shear stresses" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0002586_edpc51184.2020.9388196-Figure12-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0002586_edpc51184.2020.9388196-Figure12-1.png", + "caption": "Fig. 12. Temperature distribution out of CFD simulation.", + "texts": [ + " The amount of this power results from the load-related tooth power loss, which is already determined in section IV-A at the maximum possible load of the reference drive. Also, an increase by the safety factor S = 1.3 is applied. Besides the heat input an angular velocity of \u03c9 = 166 rad/s is considered. The simulation reveals that the pressure between inlet and outlet decreases by about \u2206p = 1.8 bar. This has to be considered perspectively if the cooling is additionally connected to the rotor cooling of an EV. The temperature of the cooling fluid heats up by \u2206T = 9.5 K in this special load case. Fig. 12 shows the temperature distribution of the critical cross-section Authorized licensed use limited to: San Francisco State Univ. Downloaded on June 20,2021 at 01:51:06 UTC from IEEE Xplore. Restrictions apply. at the driving side of the shaft. The tooth flank reaches a maximum temperature of \u03b8 = 150 \u00b0C. The mass temperature of the shaft is about 90 \u00b0C. This prevents damage to the coating in the tooth mesh due to excessive temperatures. It should also be noted that the resulting temperature distribution is similar to deep toothing with splash lubrication, which is typical for electric drives" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0001117_kem.841.327-Figure2-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0001117_kem.841.327-Figure2-1.png", + "caption": "Fig. 2 Loading Conditions.", + "texts": [ + " A mesh element size was 2 mm was chosen for the layers. Face mesh was applied in order to produce less warping and better aspect ratio of the elements. The FEM Model was solved for Equivalent Stresses, Equivalent Elastic strain in the specimen and equivalent stress in the GFRP and sugarcane fibres. Non-linear analysis was conducted in 1000 steps of computation. Also, the model was solved considering large deflection in the specimen. Loading Conditions. A CAD model of metal clamps was imported into the simulation as shown in Fig. 2, to function as the holders of the specimen when applying tensile loads. One of the holders was constrained as fixed support and the other holder was made to displace 0.35 mm as shown in Fig. 2, only in X-direction as per the experimental results. The other translational and rotational degrees of freedom were made constrained. The contact between the specimen and the holders was fixed type connection. Tensile Test. In the process of Tensile test, loads are gradually applied either mechanically with screw driver or hydraulically with pressurized oil in Universal Testing Machines. The advantage of using a Universal Tensile test machine is that there is gradual deformation with respect to time in response to the load applied" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0003721_1.4879277-Figure4-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0003721_1.4879277-Figure4-1.png", + "caption": "Fig. 4. Diagram showing the effective forces applied to the two objects in the rotational motion experiment.", + "texts": [ + " The rotation rate of the platform is controllable with a power supply and the angular velocity is measured by means of a slit disk and photogate. The containers are acrylic cylinders with their axes oriented horizontally, and a string is attached to the top and bottom of these containers to hold the lead and polystyrene masses, respectively. To determine the tilt angle during circular motion, a video camera is fixed to the rotating platform in front of the two containers. For uniform circular motion (see Fig. 4), the objects undergo a constant centripetal acceleration ac \u00bc x2R, where x is the angular velocity and R is the radius of the circular motion. Using Eq. (2), the title angle of both masses is thus given by tan h\u00f0 \u00de \u00bc ac g \u00bc x2R g : (3) Figure 5 shows a photograph of the experimental apparatus and a close-up view of the containers while in rotational motion shown as an inset. It is evident that the tilt angles have the same magnitude, as expected from our earlier discussion. In addition, using our experimental values of x\u00bc 3" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0001362_j.mechatronics.2020.102385-Figure11-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0001362_j.mechatronics.2020.102385-Figure11-1.png", + "caption": "Fig. 11. Kitchen waste processor and the plugging test simulator", + "texts": [ + " Startup and steady-state performance test of the proposed scheme From the previous comparative analysis, we can see that the system roposed in this paper has many advantages. In this paper, a kitchen aste processor system is taken as the research object, and the mechan- cal and electrical drive system is designed with the proposed scheme. n order to further test and verify the driving performance of the elec- t c d l t i w a 1 k t t w a 1 w s o 3 t 2 f p t m b r i v w i t a w a 1 n i c c 4 D i t C W v Y romechanical system, a prototype is developed for testing. The designed ombined cutter head is shown in Fig. 11 (a), the designed motor and river are shown in Fig. 11 (b), and the developed plugging test simuation device is shown in Fig. 11 (c). The two-phase current and speed waveforms of the motor are obained, as shown in Fig. 12 (a) and (b) respectively. When the motor s tested at a given speed of 2000 r/min with no load. The observed aveforms show that the motor runs from standstill to 2000r/min after bout 8.4ms, and the steady-state error of the motor speed is -12rpm ~ 3rpm. The maximum two-phase current of the motor is 0.5A when the itchen waste processor works at the no-load operation state. The two-phase current and speed waveforms of the motor are obained as well, shown in Fig" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0002338_1369433220971728-Figure2-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0002338_1369433220971728-Figure2-1.png", + "caption": "Figure 2. Determination of contact state of deployable structure with clearance.", + "texts": [ + " According to the size of the clearance circle, the motion of the shaft in the bearing can be divided into three different modes: the free flight, the impact, and the contact mode, as shown in the lower left of Figure 1. The clearance exists in the real joint to ensure the relative movement between the connecting components, which is also the source of the impact forces resulting the friction and wear, uncertainty in motion position, and degradation of the system performance. A fundamental requirement of dynamic analysis of a deployable structure with clearance is to accurately determine whether the bearing boundaries and journal are in contact with each other. In Figure 2, points Pb and Pj represent the centers of bearing and journal, respectively, and ebj denotes the eccentricity vector connecting the centers of bearing and journal. In Figure 2, rPj and rPb respectively represent the position vectors of the geometric center of journal and bearing in the global coordinate system, which can be defined in equation (1): ebj = rPj rPb \u00f01\u00de The unit normal and tangential vectors at the point of collision is defined by n=cosa i+sina j t= sina i+cosa j \u00f02\u00de where a is the angle between ebj and the x-axis. The relative penetration between two contact bodies can be expressed as: d= ebj c , \u00f03\u00de where ebj is the modulus of vector ebj. The relative indentation can be used to determine whether there is contact between the bearing and journal (Flores and Ambro\u0301sio, 2010), and the change between different contact states can be calculated by equation (4)" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0002816_thc-2010-0566-Figure9-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0002816_thc-2010-0566-Figure9-1.png", + "caption": "Fig. 9. Shear strain.", + "texts": [ + " In the human body most of the bone structure and teeth are good at bearing compressive stresses whereas ligaments are specialized structures meant to bear tensile stresses. We rarely feel the presence of yet another type of stress, which is called shear stress. This occurs when the force is active parallel to the cross-sectional area (Fig. 8A) and tries to slide one part of the body in relation to the other part (Fig. 8B). Shear stress is denoted by \u03c4 (tau) and also equal to the force acting per unit cross-sectional area with the unit Newton (N) Eq. (4). \u03c4 = F A (4) The material reacts to this shear stress as shown in Fig. 9A and the shear strain, denoted by \u03b3 (gamma), is calculated using Eq. (5) (Fig. 9B) and again it is a dimension-less number. \u03b3 = d L (5) A material experiencing shear stress also demonstrates a shear stress-strain curve very similar to a normal stress strain curve, with an elastic region where it follows Hooke\u2019s law and shear strain is linearly proportional to shear stress and having a modulus of rigidity, G, Eq. (6). Above the elastic limit the A B material deforms permanently. \u03c4 = G\u03b3 (6) Normal stresses (both tensile and compressive) and shear stress are the only two types of stresses which can exist inside the material and hence the material manifests only normal and shear strain" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0002917_b136378_7-Figure7.13-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0002917_b136378_7-Figure7.13-1.png", + "caption": "Fig. 7.13 \u201cDual\u201d glucose sensor based on two oxygen electrodes, one (a) covered with active glucose oxidase and other (b) with deactivated glucose oxidase", + "texts": [ + " As of the writing of this text, an Internet entry \u201cglucose sensor\u201d netted over 33,000 citations of which more than 5,000 were electrochemical glucose sensors! The Clark electrode-based glucose sensor is a good example of a cosubstrate sensor, one substrate being oxygen and the other one glucose. The earliest design of the glucose electrode (which is generally applicable to any oxidase and uses oxygen as the ultimate electron acceptor) is based on the differential measurement of oxygen deficiency at the oxygen electrode, caused by the enzymatic reaction. The principle of operation of this sensor is shown in Fig. 7.13. It consists of two identical oxygen electrodes: one (A) covered with the \u201cactive\u201d glucose oxidase layer and the other one (B), containing the \u201cdeactivated\u201d enzyme layer. The inactivation can be done either chemically, by ionizing radiation, or thermally. In the absence of the enzymatic reaction, the flux of oxygen to these electrodes, and therefore the diffusion limiting currents, are approximately equal \u0394i = (iL1\u223c iL2). In practice, there is always some difference given by the unequal transport properties of the two layers, however, it is easy to account for this difference" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0003392_mmar.2015.7283971-Figure1-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0003392_mmar.2015.7283971-Figure1-1.png", + "caption": "Fig. 1. Model of a tri-wheel mobile robot with castor wheel", + "texts": [ + " Hence, this paper analyses the influence of a castor wheel for the robot trajectory along a straight line. The paper is organized as follows. Section II discusses an extended model of a tri-wheel mobile robot, which accounts for the castor wheel impact [5, 6]. Section III presents the castor wheel properties according to the test measurements. Section IV shows the results of motion control analysis of the mobile robot, and Section V concludes the paper. II. EXTENDED MODEL OF A TRI-WHEEL MOBILE ROBOT Fig. 1 shows the notation that refers to a mobile robot with a castor wheel, where the following coordinate systems correspond to: (x1, y1) \u2013 the center of mass of the first driving wheel, (x2, y2) \u2013 the center of mass of the second driving wheel, (x4, y4) \u2013 the attachment point of the castor wheel, (x3, y3) \u2013 the center of mass of the castor wheel, (x5, y5) \u2013 the center of mass of the robot frame. Moreover, we assumed the following notation [5, 6]: S \u2013 the center of the mass of the robot frame, A \u2013 the characteristic point (xA, yA), l1 \u2013 the distance between the driving wheel and the characteristic point A, l2 \u2013 the distance between the characteristic point A and point S, l \u2013 the distance between the characteristic point A and the attachment point of the castor wheel, l3 \u2013 the frame of the castor wheel, \u2013 the torsion angle of the castor wheel, \u2013 the angle of rotation of the robot frame, 1, 2, 3 \u2013 the angles of rotation of the driving wheels ( 1, 2) and the castor wheel ( 3)" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0000526_icra40945.2020.9197293-Figure4-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0000526_icra40945.2020.9197293-Figure4-1.png", + "caption": "Fig. 4: Left: An example where the 6DFC algorithm finds a robust grasp when the Point algorithm does not. Right: An example where the Point algorithm finds a robust grasp when the 6DFC algorithm does not. In the first case, the thin part of the object results in a low-quality prediction for the point contact algorithm, whereas in the second case, the large contact area produces a false positive prediction from the 6DFC algorithm.", + "texts": [ + " To test this hypothesis, we place each of the 12 objects in their 3 most probable stable poses [14] and attempt the top 3 grasps that each algorithm labels as the highest reliability grasps for that stable pose. We remove nearby grasps so that the algorithm cannot choose 3 similar grasps, simulating other objects blocking the grasp from being executed. If multiple grasps are labeled with the same probability, one of them is chosen at random. We evaluated 875 unique grasps in total for 32 stable poses of the 12 objects. Both algorithms found successful grasps in at least one stable pose that the other could not (Figure 4 shows two examples), but overall the point grasps surprisingly succeeded more often (339 successes compared to 272 for the 6DFC algorithm). This result may be due to an inflation in grasp quality in the 6DFC algorithm; we 7895 Authorized licensed use limited to: Carleton University. Downloaded on October 04,2020 at 15:25:44 UTC from IEEE Xplore. Restrictions apply. found that many grasps that had large contact areas were rated highly but did not succeed, as shown in the bottom row of Figure 4. We will investigate this discrepancy further in future work. As each of the algorithms contains several parameters such as friction and elasticity coefficients, robustness parameters, we include an analysis of each algorithm\u2019s sensitivity to a subset of these parameters. 1) Effect of Robustness Parameters: All of the algorithms benefit from \u201crobustness\u201d, or sampling grasp poses around the nominal pose and averaging the predicted grasp qualities. We sample a random 3D translation from a zeromean Gaussian with variance \u03c32 r and a uniformly random 3D rotation with angle \u03b8 proportional to \u03c32 r and apply them to the nominal grasp" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0003435_icinfa.2015.7279707-Figure1-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0003435_icinfa.2015.7279707-Figure1-1.png", + "caption": "Fig. 1. Frame system with a body fixed frame B and the inertial frame E.", + "texts": [ + " Reference [10], the first step before control development is an adequate dynamic system modelling. The dynamic model of quadrotor ideally includes the gyroscopic effects resulting from both the rigid body rotation in space, and the four propeller\u2019s rotation. These have been often neglected in previous works, as in [15]. However, the main effects acting on a helicopter are described briefly in table1. Let us consider an inertial coordinate frame E fixed to ground and a body fixed coordinate frame B (see fig.1). Using Euler angles parametrization, the airframe orientation in space is 978-1-4673-9104-7/15/$31.00 \u00a92015 IEEE given by a rotation R from B to E, where R\u2208 SO3 is the rotation matrix. The dynamics of quadrotor under external forces expressed in the body fixed frame in Newton-Euler formalism: ( mI3\u00d73 0 0 I )( V\u0307 \u03c9\u0307 ) + ( \u03c9 \u00d7mV \u03c9 \u00d7 I\u03c9 ) = ( F \u03c4 ) (1) Where I\u2208 R3\u00d73 is the inertia matrix, V the body linear speed vector and \u03c9 the body angular speed. Reference [3], we can get mx\u0308 = (sin \u03b8 cos\u03d5 cos\u03c8 + sin\u03d5 sin\u03c8) 4\u2211 i=1 Ti, my\u0308 = (sin \u03b8 cos\u03d5 sin\u03c8 + sin\u03d5 cos\u03c8) 4\u2211 i=1 Ti, mz\u0308 = cos\u03d5 cos \u03b8 4\u2211 i=1 Ti \u2212mg, Ix\u03d5\u0308 = \u03b8\u0307\u03c8\u0307(Iy \u2212 Iz)\u2212 Jr \u03b8\u0307\u03a9 + l(T4 \u2212 T2), Iy \u03b8\u0308 = \u03d5\u0307\u03c8\u0307(Iz \u2212 Ix)\u2212 Jr\u03d5\u0307\u03a9 + l(T3 \u2212 T1), Iz\u03c8\u0308 = \u03d5\u0307\u03b8\u0307(Ix \u2212 Iy) + d(\u03a92 2 +\u03a92 4 \u2212 \u03a92 1 \u2212 \u03a92 3)" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0000161_012109-Figure3-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0000161_012109-Figure3-1.png", + "caption": "Figure 3. Functional diagram of working out the inclination of the body of the grinding head relative to the treated surface", + "texts": [ + " An important task in grinding large-sized surfaces is the determination of the position of the flap wheel relative to the workpiece [17]. This condition ensures uniform removal of the stock from the surface to be treated. To solve this problem the scientists developed an adaptive control system for the position of the grinding head. It works using a pair of ultrasonic sensors installed in a seal at the ends of the working window. (fig. 2). In the figure 2, numerals denote the following elements: 1 \u2013 work surface; 2 \u2013 ultrasonic sensors; 3 \u2013 resilient seal; 4 \u2013 flap wheel. This system functions as follows (fig. 3). ICI2AE 2019 IOP Conf. Series: Materials Science and Engineering 632 (2019) 012109 IOP Publishing doi:10.1088/1757-899X/632/1/012109 Before processing, we fix the part (panel or fuselage shell) in a vertical position in the fixing device. Next, we set the required width of the abrasive flap wheel and control at the working position with the help of the corresponding positioning unit. After, at some distance, there is a coordination alignment of the center of the width of the flap wheel with the middle line of the processing band. Then, the grinding head is moving towards the surface to be treated, while using the sensor set 2 by the adaptive control system, the distances L1 and L2 to the surface 1 are continuously monitored (fig. 2). During processing, the sensors maintain constancy of distances, and, if necessary, correct it by tilting the housing from or to the surface at an angle \u03b1 (fig. 3\u0430). If necessary, there is the correction of the movements along the axis \u0394Y and \u0394Z (fig. 3b). The angle of inclination and corrective movements is calculated by the following formulas: H SS arctg 21 (1) 2 2 COS SinSinR Z (2) 2 2 COS COSSinR Y (3) where S1 \u0438 S2 \u2013 readings of the upper and lower ultrasonic sensors, H \u2013 distance between sensors, R, \u03b2 \u2013 polar coordinates of the center of seal. Depending on the width of the circle used, the corresponding pairs of signals from the sensors are processed, and the angle is determined, by which the grinding head should be turned" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0002989_978-3-642-16214-5_9-Figure9.21-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0002989_978-3-642-16214-5_9-Figure9.21-1.png", + "caption": "Fig. 9.21. Powertrain with coaxial 5-speed countershaft transmission (\u201cin-line\u201d)", + "texts": [ + " The acceptable stress values are derived from experience. Reduction of Moments of Inertia As a result of the steps between ratios, the masses involved in the synchronizing process are subject to different angular accelerations. In order to be able to use only one angular velocity for all the masses involved in the calculation, the masses are related to one axis. This is normally the rotation axis of the idler gear to be shifted. In general: \u2211 = += i 1k 2 k kiired, 1 i JJJ . (9.9) Example: When shifting the transmission shown in Figure 9.21 from second to first gear, the masses are reduced onto the rotation axis of the idler gear 7. In this case: . )( )( 2 8 7 2 13 14 13 2 11 10 11 2 9 10 9 2 5 6 5 2 3 4 3 2 8 7 14108642CS 2 1 2 2 8 7 1ISC77red, \u239f\u239f \u23a0 \u239e \u239c\u239c \u239d \u239b \u23a5 \u23a5 \u23a6 \u23a4 \u23a2 \u23a2 \u23a3 \u23a1 \u239f\u239f \u23a0 \u239e \u239c\u239c \u239d \u239b +\u239f\u239f \u23a0 \u239e \u239c\u239c \u239d \u239b +\u239f\u239f \u23a0 \u239e \u239c\u239c \u239d \u239b +\u239f\u239f \u23a0 \u239e \u239c\u239c \u239d \u239b +\u239f\u239f \u23a0 \u239e \u239c\u239c \u239d \u239b +\u239f\u239f \u23a0 \u239e \u239c\u239c \u239d \u239b ++++++ +\u239f\u239f \u23a0 \u239e \u239c\u239c \u239d \u239b \u239f\u239f \u23a0 \u239e \u239c\u239c \u239d \u239b +++= z z z z J z z J z z J z z J z z J z z JJJJJJJ z z z z JJJJJ (9.10) On the assumption that the output shaft OS and the components connected to it are not subject to any change of angular velocity during synchronization, their moments of inertia may be ignored" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0000959_0954406220917995-Figure2-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0000959_0954406220917995-Figure2-1.png", + "caption": "Figure 2. The view of (a) the computational domain, boundary condition, and (b) the mesh around the wing.", + "texts": [ + "26 The coupling of the pressure and velocity is handled using the coupled algorithm. The spatial discretization is carried out using a least-square cell-based scheme. The pressure and momentum equations are solved with a second-order scheme. All scaled residuals leveled off and reached 10 5. In the present study, we used sliding mesh to model the motion of the flapping wing,24 a hybrid grid method is utilized to generate the grid. The domain size of the 3D wing model for flapping flight is essential for CFD simulation. The computational domain is 20c, illustrated in Figure 2(a). The domain size setting is the same as Srinidhi and Vengadesan.27 The cell size was increased by keeping a factor of 1.05 away from the wing and 1.1 for the outer boundary. Figure 2(b) displays the view of the mesh around the wing. The outer boundary is set as a pressure outlet. The wing edge was set with the condition of no-slip wall. The aims of this section are to find out the effects of different wing corrugation patterns and corrugation inclinations on steady and unsteady flapping flight performances. First, the CFD code is validated using experimental aerodynamic forces and the time step and grid independence are briefly discussed in the next section. Next, different corrugated patterns and inclination angles during steady flight were studied" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0001701_0954405420949757-Figure1-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0001701_0954405420949757-Figure1-1.png", + "caption": "Figure 1. Machine tools coordinate system of pinion processing.", + "texts": [ + " Ball end milling cutter is widely used in NC machining due to its low preparation cost and high flexibility,16,17 which is applicable for processing complex curved shape.18 According to the analysis above, a tooth crest chamfering is proposed based on four-axis CNC machine tools and ball end milling cutter, which can meet requirements of high processing efficiency and low processing cost. Mathematical model of tooth surface Machine tools coordinate system Tooth crest chamfering is researched by taking Gleason left-hand modified roll pinion19,20 as an example. Machine tools coordinate system of pinion processing is shown in Figure 1. Intersection point of top plane of cutter and axis of generating gear is defined as origin Om. The direction from cutter holder to cutter tip is defined as Zm, the straight up direction is defined as Ym, the direction of Xm is determined by the right hand rule. O0 is the intersection point of top plane and axis of cutter, the length of O0Om is called radial setting S1, the angle between O0Om and Xm is called cradle angle u1. The distance between pinion axis p1 and plane XmOmZm is called vertical offset E01" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0003970_rev.2014.6784196-Figure3-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0003970_rev.2014.6784196-Figure3-1.png", + "caption": "Fig. 3. 3D model parts of electromagnetic emulator (1 \u2013 cantilever, 2 \u2013 electromagnets, 3 \u2013 rotor disc)", + "texts": [], + "surrounding_texts": [ + "This paper is a result of activities within the project 543667-TEMPUS-1-2013-1-RS-TEMPUS-JPHES \u201cBuilding Network of Remote Labs for strengthening university-secondary vocational schools collaboration\u201d supported by The Education, Audiovisual and Culture Executive Agency (EACEA)." + ] + }, + { + "image_filename": "designv11_71_0002483_cac51589.2020.9326899-Figure1-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0002483_cac51589.2020.9326899-Figure1-1.png", + "caption": "Fig. 1. Simplified exoskeleton model", + "texts": [ + " Thus, this paper will use GRNN-DMP model to learn gait trajectory, which is considered to enhance the comfort and adaptability in the process of movement of patients with the help of exoskeleton. The rest of this paper is structured as follows: In section II, the structure and kinematic model of lower limb rehabilitation exoskeleton are introduced. In section III, the strategy of gait trajectory generation based on GRNN-DMP is derived. In section IV, the simulation results are given to verify the performance of the proposed method. Finally, the conclusion is given in section V. II. EXOSKELETON DESCRIPTION AND KINEMATIC MODELLING As shown in Fig. 1, the lower limb rehabilitation exoskeleton is simplified into a multi-link model. Taking a single leg as example, there are three degrees of freedom(DOFs) in the sagittal plane, one DOF in the coronal plane and one DOF in the transverse section. The x-axis is the forward direction, the y-axis is the horizontal direction, and the z-axis is the vertical direction. The link length parameters are shown in Table I. Only the sagittal DOF is considered here, and other DOFs are set to 0. The whole walking process can be divided into This work is partially supported by the National Natural Science Foundation of China (61773212), the Natural Science Foundation of Jiangsu Province (BK20170094), and the International Science & Technology Cooperation Program of China (2015DFA01710)," + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0001371_8823102-Figure1-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0001371_8823102-Figure1-1.png", + "caption": "Figure 1: Compound fault theory model of internal ring spalling and rolling element spalling.", + "texts": [ + " In this case, the vibration signal of the rolling bearing compound fault collected by the acceleration sensor is more complicated, and the corresponding envelope signal and its spectrum are also more complicated. In order to make a more realistic interpretation of the envelope signal and the envelope spectrum and improve the accuracy of the diagnosis, a model for analyzing the envelope signal and its frequency spectrum of the compound bearing compound fault is proposed in the following. In this chapter, the theoretical model is studied by taking the compound fault of the inner ring spalling and the rolling element spalling as an example. Figure 1 shows the theoretical model of compound fault of inner ring spalling and rolling element spalling. When t\ufffd 0, the position of the spalling point of the inner ring is exactly at the peak of the load distribution density curve, that is, the lowest position in the axis direction, and it just collides with the rolling element. Since the fit between the inner ring and the shaft is an interference fit during assembly, the inner ring will rotate as the shaft rotates. -e spalling point on the inner ring also rotates with the rotation of the bearing, so that the position of the spalling point does not always fall within the load distribution area as shown in Figure 1; most of the time it will fall outside the load distribution area. Due to the different load, the resulting pulse force is different. When the spalling point of the inner ring falls within the load distribution area as shown in the figure, the magnitude and direction of the vibration pulse force are related to the position of the spalling point and the angle of the axis. When the spalling point is outside the load area shown in Figure 1, it can be known from the force analysis that if the rotational inertia force generated when the rolling bearing rotates is ignored at this time, no vibration pulse force is generated in this area. In summary, when the internal ring has a spalling fault, the vector value of the generated vibration pulse force is not only affected by the load distribution but also related to the location of the spalling fault point. From this analysis, it can be seen that the magnitude and direction of the impulse force caused by the damage point of the inner ring and rolling element are affected by the load distribution and the position of the damage point" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0002393_j.matpr.2020.11.109-Figure1-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0002393_j.matpr.2020.11.109-Figure1-1.png", + "caption": "Fig. 1. Solid model of SUV car design with special tapered edges at rear end.", + "texts": [ + " Each and every surface in this car models can be adjusted using flexible modeling tool. In analysis Autodesk Flow Design software is used. This is an interactive application called Virtual wind tunnel. Using this we can analyze the drag force and turbulence without any complexity. Reason for using Auto Desk Flow Simulation software is to make analysis sim- ple. It is system built-in-properties equipped software that uses standard air properties for the analysis purpose. So we can obtain actual results as we can expect from the real wind tunnel. Fig. 1 represents the solid model of SUV car design with special tapered edges at rear end and Fig. 2 represents the solid model of Sedan car design with pressure diffuser fixed at the bottom end. B. Meshing of car models with Creo The solid model which is taken for drag analysis is meshed with different type of meshing based on the accuracy at particular area of the body. In this analysis process, the front area is meshed with triangular mesh and back fillet portion are meshed with triangular mesh. Finally, the side portion of the model is meshed with tetrahedral mesh due to simple operation and less time taken for simulating the results" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0001752_022053-Figure4-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0001752_022053-Figure4-1.png", + "caption": "Figure 4. FEM model.", + "texts": [ + "\" The dynamic loads acting on the tooth have been calculated using Buckingham dynamic load relations as follows [15]. (1) (2) For 20O Pressure angle (3) (4) (5) (6) (7) (8) In this study, the three-dimensional model of spur gear with involute profile is employed in both the healthy and faulty spur gear models. For this analysis, the driver and driven gear tooth were assumed as rigid body and have same material properties which are listed in the table.2. The perfect and defective gear is designed using SOLID185 element in ANSYS as shown in figure.4. Table.2 Material properties Material Steel-Steel Young\u2019s Modulus (N/mm2) 2.1 \u00d7 105 Poisson\u2019s ratio 0.3 In this analysis, dynamic load is calculated using Buckingham\u2019s equations (1), (2), (3), (4), (5) and (6) which is listed in the table.3. The dynamic load is applied at more than several points of contact when the gear pair rotates from the point of engagement to disengagement of the single gear tooth pair [18]. 3rd International Conference on Advances in Mechanical Engineering (ICAME 2020) IOP Conf" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0000245_ecce.2019.8912784-Figure5-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0000245_ecce.2019.8912784-Figure5-1.png", + "caption": "Fig. 5: The magnetic flux density in the conductors of an example AFPM coreless machine calculated with 3D FEA. For such machines, the winding is directly exposed to the airgap field, the magnitude and multi-dimensional variation of which can be substantial, resulting in significant eddy current losses.", + "texts": [ + " For an example AFPM machine, the magnitude of the flux density in each of the winding turns placed in the slot and surrounding a tooth was estimated with one sample per conductor using 2D and 3D models, respectively (Fig. 6). The 3D samples are taken for each of the 21 turns placed at equally distanced radial locations and then averaged. The 2D model was set-up for the mean diameter. The 2D model overestimation of the flux density, particularly for the turns that are closer to the top of the slot and closer to the teeth, is noticeable. The coreless AFPM machine example from Fig. 5 is illustrative of the typical very large 3D variation of the flux density in the stator windings. In order to account for this, one solution would be to study multiple 2D models representative of slices at different radial coordinates and combine their contributions [12]. An increased number of 2D slices would increase, in principle, the accuracy of the simulation at the expense of increased computational time, but won\u2019t still account for the end field, which makes the use of 3D models worthwhile even more so", + "0 49.4 Approximate FEA computational time 37 hours 35 minutes 2 minutes 12 seconds Equation (6), which comprehensively represents the discrete implementation of the proposed new hybrid 3D FEA method, was employed in the following case studies. Two example case studies were conducted: a 20 pole 24 slot AFPM machine with open slots and precise-wound concentrated windings with coils placed around the core teeth, as shown in Fig. 1, and a coreless AFPM machine with 12 coils and 16 poles, as shown in Fig. 5. The winding conductors were connected in series. It should be noted that in the case of parallel connections between turns, conductor transposition may reduce the circulating currents and additional losses. The calculations were performed for open-circuit and load, at a fundamental frequency of 1.6 kHz for the open-slot machine and at 480 Hz for the coreless design, respectively. A coreless machine operating at higher frequencies would typically require the use of Litz wire, which is beyond the scope of the current paper" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0001279_piicon49524.2020.9112890-Figure4-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0001279_piicon49524.2020.9112890-Figure4-1.png", + "caption": "Fig. 4 25% Inter-turn fault in stator winding", + "texts": [], + "surrounding_texts": [ + "A 3HP, 415V, 50 Hz, 1440rpm , 4 pole three phase Induction motor is designed using the basic equations of motor and it has been modelled in RMxprt in Ansys Maxwell. All the parameters such as stator diameter, length, number of stator slots and specifications, Type of winding, dimensions, Type of steel, Number of conductors/slot, No of Parallel branches, rotor length, diameter, number of slots are given from the calculated values. The performance characteristics is analysed for both healthy and inter-turn short circuit fault of three phase Induction motor. The inter-turn fault is modelled for the same motor by short circuiting 16% and 25% of the turns in one phase. The Transient analysis for healthy as well as faulty induction motor is done in Ansys Maxwell 2D. Table I Induction motor Specification" + ] + }, + { + "image_filename": "designv11_71_0000175_012073-Figure2-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0000175_012073-Figure2-1.png", + "caption": "Figure 2. The rotor revolving in a horizontal plane with two chains moving in three planes.", + "texts": [ + "k k k k k k kR x x y y z z (3) Add the coefficient to the system (2) that allows scaling the coordinates of points; it will be equal to the ratio of the constant length of the flexible working body rk to the length obtained after the increment of the angles of rotation of its points \u2013 Rk. 1 1 2 2 1 1 _ sin( ) cos( ) / ; cos( ) cos( ) / ; ( ) ( ) sin( ) / . k k k k k k k k k k k k k k k k k k k k b k k x r t t t r R y r t t t r R z x x y y t r R (4) This mathematical model (1-4) enables to calculate the trajectory of a flexible working body, for known angular velocities of its constituent links [7]. We developed a simulation model of a rotor operating in a horizontal position Using the Motion SolidWorks module (figure 2). Inertial forces, gravity and acceleration characteristics are taken into account in the study of the rotor kinematics. The general formulation of the problem is as follows: [ ] ( ) [ ] ( ) [ ] ( ) ( ),M s t C s t K s t f t (5) where [K] \u2013 stiffness matrix; [C] \u2013 damping matrix; [M] \u2013 mass matrix; s(t) \u2013 displacement vector; ( )s t \u2013 speed vector; ( )s t \u2013 acceleration vector Forestry 2019 IOP Conf. Series: Earth and Environmental Science 392 (2019) 012073 IOP Publishing doi:10.1088/1755-1315/392/1/012073 Modeling settings in Motion SolidWorks: material \u2013 steel (dry); dynamic friction velocity k= 10" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0002338_1369433220971728-Figure3-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0002338_1369433220971728-Figure3-1.png", + "caption": "Figure 3. Contact points detection in the general clearance joint.", + "texts": [ + " The relative penetration between two contact bodies can be expressed as: d= ebj c , \u00f03\u00de where ebj is the modulus of vector ebj. The relative indentation can be used to determine whether there is contact between the bearing and journal (Flores and Ambro\u0301sio, 2010), and the change between different contact states can be calculated by equation (4). d qt\u00f0 \u00ded qt 1\u00f0 \u00de.0, d qt\u00f0 \u00de\\0 free flight d qt\u00f0 \u00ded qt 1\u00f0 \u00de.0, d qt\u00f0 \u00de.0 continuous contact d qt\u00f0 \u00ded qt 1\u00f0 \u00de\u0142 0, d qt\u00f0 \u00de\u0142 0 from contact to seperation d qt\u00f0 \u00ded qt 1\u00f0 \u00de\u0142 0, d qt\u00f0 \u00de\u00f8 0 from seperation to contact 8>< >: \u00f04\u00de Figure 3 shows the position relationship when the journal and bearing in the clearance joint are about to contact. Qb and Qj are the candidate contact points when the contact bodies are about to impact. In the dynamic analysis of clearance mechanism, the potential contact points detection is the basis for calculating contact deformation and relative contact velocity, which determines the accuracy and precision of the calculation. For the general clearance joint, the positional relationship between potential contact points should meet the following geometric constraint conditions: d= r Q b r Q j , \u00f05\u00de nj 3 nb = 0 d 3 nb = 0 , \u00f06\u00de where r Q b and r Q j are the position vectors of potential contact points in the global coordinate systems, nj and nb are the normal vectors of the possible contact points (Li et al" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0003788_detc2014-34213-Figure15-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0003788_detc2014-34213-Figure15-1.png", + "caption": "Fig. 15 Eight-bar linkages with input joints A0, A and B", + "texts": [ + " For Type-112, the inputs are given through the joint combinations A0AB0, A0FB0 and BFB0. Type-222 is the input condition in which the inputs are given through the joints A, B and B0 or the joints A0, F and E. Since the eight-bar linkage has a symmetric structure, the inputs given in the other six-bar loop are the same. For Type-123 input condition, taking A0, A and B for example, there are two decomposition ways. One is by holding the input joint B, the three-DOF planar eight-bar linkage degenerates a two-DOF seven-bar planar linkage, as shown in Fig. 15(a). The other is by holding the input joints A0 and B, then, the three-DOF planar eight-bar linkage degenerates a Stephenson six-bar planar linkage, as shown in Fig. 15(b). The singularities of the two-DOF seven-bar linkage and the singleDOF Stephenson linkage are known in the above discussion. Thus, according to the proposed criteria, if the Stephenson sixbar linkages or the two-DOF seven-bar planar linkage is at the singular positions, the three-DOF planar eight-bar linkage becomes singular. The singularity of the six-bar linkages or the two-DOF seven-bar planar linkage happens when the joints E, F and B0 are in the common line or the joints D, C and C0 are collinear, which is also the singular condition for the three-DOF eight-bar linkage" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0000885_s42947-020-0303-x-Figure1-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0000885_s42947-020-0303-x-Figure1-1.png", + "caption": "Fig. 1. Sample trimming process and the prepared sample (a) sample dimension and (b) prepared sample.", + "texts": [ + " All of them are Superior Performance (SP) III mix, which has a Nominal Maximum Aggregate Size (NMAS) of 19.0 mm. Both warm and hot mix are included. The RAP content among the mixes varies from 0 to 35%. Note: HMA=Hot Mix Asphalt; WMA= Warm Mix Asphalt; PG=Performance Grade; RAP=Reclaimed Asphalt Pavement. HWTD testing. About 12 replicate samples were prepared for each mix using the Superpave gyratory compactor following the AASHTO T 312 [20] test standard. The samples were then cut into specific dimensions using a laboratory cutting saw, as shown in Fig. 1(a). Cylindrical samples are prepared to a height of 62\u00b11 mm (Fig. 1(b)). Samples\u2019 volumetric properties, theoretical maximum specific gravity (Gmm), air voids, and bulk specific gravity (Gmb) were determined at the laboratory. The Gmm, Gmb, and air void were determined according to the AASHTO T 209 [21], AASHTO T 166 [22], and AASHTO T 269 [23] test protocols respectively. The samples' air voids content was kept within 6\u00b11% to avoid possible deviation in the test results due to air voids variation. The Hamburg wheel tracking test was conducted in accordance with the AASHTO T 324-11 [24] test standard" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0002119_icaccm50413.2020.9212874-Figure12-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0002119_icaccm50413.2020.9212874-Figure12-1.png", + "caption": "Fig. 12. Applied rotational velocity (case2).", + "texts": [ + " 60 Authorized licensed use limited to: Carleton University. Downloaded on May 30,2021 at 15:29:13 UTC from IEEE Xplore. Restrictions apply. A Quadcopter\u2019s propeller is subjected to horizontal bending due to collision of the propeller with obstacles or a wall. To calculate the bending deformation under the applied rotational moment about the rotation axis, the propeller has been fixed by both the tip as shown in Figure 10.In addition, it is fixed at the point where it is connected to the motor shaft (Figure 11).Figure 12 shows the applied rotational velocity of 897 rad/s [8].Figure 13, 14 and 15, 16 show the total deformation and equivalent stress for CFRP and GFRP materials, respectively. The comparison of the obtained results is presented in Table III. 61 Authorized licensed use limited to: Carleton University. Downloaded on May 30,2021 at 15:29:13 UTC from IEEE Xplore. Restrictions apply. Fig. 14. Total deformation of GFRP propeller (case2). Fig. 15. Equivalent stress for CFRP propeller (case2). Table III. Results due to rotational velocity CFRP GFRP Max" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0002861_978-3-030-55889-5-Figure3.3-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0002861_978-3-030-55889-5-Figure3.3-1.png", + "caption": "Fig. 3.3 Plate, a system, b deflection under gravity load, c influence function for the deflection w at a node x", + "texts": [ + "5 = 2.75 = w(x) , (3.16) but the result is then of course only an approximate value. \u201cBut doesn\u2019t an FE-program calculate the nodal values by solving the system Kw = f and the values in between by interpolating between the nodes?\u201d Correct, but they are as large as if they had been calculated with the approximate influence functions. This is the key point. 154 3 Finite Elements 3.2 Why the Nodal Values of the Rope Are Exact 155 In 2-D and 3-D this is true as well. The nodal values of the plate in Fig. 3.3 are the result of the superposition of the nodal Green\u2019s functions with the load wh(xi ) = \u222b \u03a9 Gh( y, xi ) p( y) d\u03a9 y . (3.17) Of course, the nodal values wi in the output are the solution of the system Kw = f , but they are just as large as if the FE-program had integrated over the whole plate, the results are the same1 wh(xi ) = wi = \u2211 j k(\u22121) i j f j = \u222b \u03a9 Gh( y, xi ) p( y) d\u03a9 y . (3.18) This is the secret, little-known law behind finite elements. The quality of the influence functions determines the quality of the results" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0001285_icieam48468.2020.9112086-Figure5-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0001285_icieam48468.2020.9112086-Figure5-1.png", + "caption": "Fig. 5. The electrical diagram of the CD", + "texts": [], + "surrounding_texts": [ + "The operation of the control system is based on the obtaining and processing of measurement information. The relationship between the m-dimensional vector of the measured parameters z and the n-dimensional vector x of the estimated parameters is given by the linear algebraic relation ( ) ( ) ( ) ( );t t t t z H x ( ) ,k k k k z h x where, z , are the m-dimensional vectors of measurement and measurement noise. III. THE SYNTHESIS OF THE CONTROLLED DEVICE The controlled device (CD) is a basic element of the complex system of automation of the gate operation. The problem of synthesis of the controlled device is solved using a fuzzy logic controller. At its realization the CO is considered to be set. The measurement system is represented by a set of primary sensors and an optimal nonlinear filter for evaluating the state vector connected with a fuzzy logic controller [11]. At the synthesis of the controlled device a five-level method of functional decomposition was used. The global CD integration layer provides data exchange with other information systems (Ethernet network with TCP/IP exchange protocol and using Linux network operating systems). At this level the server/work stations communication is provided with mandatory protection against unauthorized access. At the next level the information interaction between the information collection and processing systems and subsystems for monitoring parameters of the technological process is carried out. At this level the central processor combines all subsystems and ensures their interaction. The third system level via communication channels or through peripheral device integration interfaces provides interaction of local or peripheral information acquisition and processing systems. In this case a combination of vertical (between the central and subsystem controllers using the RS232 interface) and horizontal integration (between homogeneous controllers in each of the subsystems using the RS-485 interface) is possible. It allows to build a network with good noise immunity and sufficient data exchange rate. The fourth (modular) level implements the interaction between the peripheral system of information collection and processing and controlling and the subsystem of the testing system. The modular architecture is quite simple to configure from the standard set of the blocks for almost any controlled object and allows to expand the functions. The lower (fifth) level of the CD architecture is necessary for communication of technical means used in additional security systems (for example, Closed Circuit Video Equipment) through the local plumes or integration interfaces. One of the possible variants to implementate the measuring unit of the controlled device with using the software package ZETLab is shown in Fig. 3. In Fig. 4 we show the dependences of the torque and braking torque explaining the control law implemented at the full operating cycle of the valves. This cycle includs the modes \"movement-closing / closing end / closed state / movementopening / opening end / open state\". The control device allows to realize the controlling of valve drive with help of a frequency converter. The gate drive control device is implemented in the form of a hardware-software module providing matching the interface of the microcomputer with the frequency converter (FC) Authorized licensed use limited to: Auckland University of Technology. Downloaded on July 26,2020 at 17:03:56 UTC from IEEE Xplore. Restrictions apply. necessary for remote control of the electric drive, the implementation of the specified operating modes of the drive and motor protection against overload and short circuit [13, 14]. The CD device is built on the basis of a microcontroller. It has the digital control inputs (DI) of the computer \"Master\", the analog output (AO) 0...+10 V and the digital outputs switching the voltage +24 V at the current of not more than 4 mA. The device additionally contains a USB-USART converter with the ability to receive upper-level control signals from the computer \"Master\" via the USB interface. The microcomputer type Raspberry Pi 4 with an activated terminal program can acts as the computer \"Master\" [14]. At the end positions of the gate the signals from the measuring sensors are received to the control. The device is powered from the pulse power supply DR15-12 (+12 V, 1.25 A). The device power input is protected from polarity reversal by a rectifier diode. The control device contains an optoelectronic pair providing the galvanic isolation between the computer \"Master\" and the frequency converter. We realize the commutation of the frequency converter to the control device in accordance with the Table 1. frequency converter is shown in Fig.6. The gate is controlled is made with the digital and analog inputs of the converter at the drive start and using the standard parameter values. By default, the device is configured for discrete control. The device will be completely controlled from the computer when it connected to the computer after receiving the appropriate command from it. There are two ways to connect to the external computer \"Master\" issuing commands to open and close the gate: - with using the discrete digital signals; - with using the USB-USART serial interface signals. The algorithm of device operation in the discrete mode of the input of commands from the computer \"Master\" is given in Table 2: Authorized licensed use limited to: Auckland University of Technology. Downloaded on July 26,2020 at 17:03:56 UTC from IEEE Xplore. Restrictions apply. 2020 International Conference on Industrial Engineering, Applications and Manufacturing (ICIEAM) here X is an arbitrary value \"1\" or \"0\", Digital Input 1 \u2013 \"1\" (the function \"START FC\", Digital Input 2 \u2013 \"0\" (the function \"STOP FC\", CVX2 \u2013 \"0\" (the function \"Closing\"), CVX2 \u2013 \"1\" (the function \"Opening\"). The forced \"STOP\" is carried out by feeding the discrete signal \"1\" from the computer \"Master\" to the input \"STOP\". To provide a mode of command input via the USB-USART interface the CD is connected by a USB cable to a computer with the installed software to identify the FTDI232 chip, then any terminal program is started in the computer \"Master\". The message 0\u00d732 0\u00d755 (the symbols \"2U\") is generated from the terminal program to the control device. It corresponds to the beginning of the CD operation on the serial interface. To return to the discrete mode we must to send the message: 0\u00d730 0\u00d744 (the symbols \"0D\") or reset the settings by the \"RESET\" button. The control commands in the symbolic form are \"0U\" \u2013 to open on a signal from the sensor; \"1U\" \u2013 to close on the signal from the sensor; \"2U\" \u2013 stop; \"3U\" \u2013 to close (forcibly); \"4U\" \u2013 to open (forcibly). An example of entering a command in a terminal program is shown in Fig. 7. The hardware implementation of the control system is shown in Fig. 8. The control device 1 is made of impactresistant heat-resistant ABS plastic UL-94V0 in the D3MG housing. It has the dimensions of 53\u00d790\u00d757 mm and together with the frequency converter 2 is fixed on a DIN rail. The electromechanical drive for the linear movement of the main pipeline valve is shown in Fig. 9." + ] + }, + { + "image_filename": "designv11_71_0002361_ffe.13405-Figure4-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0002361_ffe.13405-Figure4-1.png", + "caption": "FIGURE 4 Schematic of the bearing fatigue test: (A) test rigs for cylindrical roller bearings, (B) the housing and test bearing (magnified view of arrow X), (C) the fitting-gap and rolling-element load, (D) bending stress on the raceway", + "texts": [ + " The specimen surface was polished with emery papers and buffed with a diamond paste, then the surface layer was electro-polished at 40 \u03bcm per diameter, using a liquid consisting of 2 L of phosphoric acid, 40 g of oxalic acid and 40 g of gelatin. A semicircular notch was later introduced by EDM onto the specimen surface, as shown in Figure 3B. Regarding the materials used in the tensioncompression and four-point bending tests, it must be noted that their heat-treatment conditions were specifically determined so that their hardness values corresponded to those of the bearing test samples. Hence, Rockwell hardness values of all materials fell within the range of HRC 60\u201362. The bearing test is schematically illustrated in Figure 4. VG68 viscosity-grade machine oil was continuously supplied via a forced circulating style (cf. Figure 4A). The bearing load was applied to the outer ring in a radial direction through the housing (cf. Figure 4B). The outer ring was positioned to ensure that the defect was located in the load-zone centre. In all tests, the bearing load was 8.0 kN and the rotation speed was 500 min\u22121. The bearing load generates a rolling-element load, caused by each rolling element in action on the raceway (cf. Figure 4C), variable depending on their circumferential position. The utmost, rolling-element load was calculated as 2.8 kN, matching the maximum contact stress of 1.7 GPa, with a half-contact width of 0.11 mm.36\u201338 The fitting-gap and outer-ring load conditions are shown in Figure 4C,D. The fitting gap has been defined as the difference between the inner housing radius and the exterior outer-ring radius. In this study, the fitting gap was modified by altering the interior housing diameter so as to vary bending stress. In actual rolling bearings, the fracture process can be accelerated by an inadequately loose fitting. Specifically, a loose fitting is caused occasionally by severe housing wear and/or by housing stiffness-decay due to oil-circulation grooves/holes. Therefore, to simulate these effects, bearing tests were conducted under a few loosefitting conditions, whereby the fitting gap for bearings with Type A defects was 0" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0003632_0954406214549786-Figure6-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0003632_0954406214549786-Figure6-1.png", + "caption": "Figure 6. Schematic diagram of the experimental setup.", + "texts": [ + " The torsional stiffness of the coupling, the worm shaft, and the worm wheel shaft is 0.004 kN m/mrad, 0.1 kN m/mrad, and 1.5 kN m/mrad, respectively. The worm is supported by angular contact ball bearings (model 7207AC) with the preload 380N. The total axial stiffness of the worm supporting bearings kwb s\u00f0 \u00de is determined by equations (15) and (16) with the torque at different tilting angles. The impact testing for the natural frequency of the tilting table is performed by LMS Test. Lab as shown in Figure 6. The hammer was excited at one point and the response was measured at multi-point. Accelerometers were glued at six different positions by ceresin wax on the tilting table, two (Nos. 1\u20132) of them on the supporting part and the others (Nos. 3\u20136) on the top surface of tilting base. When these six measure points rotate around x-axis simultaneously, the associated natural frequency was chosen as the one in the tilting direction. The tests were repeated every 10 from 0 to 90 . Comparison between experimental and theoretical results Figure 7 plots the torque applied on the transmission system at different angles by equations (8) to (10), and it can be clearly seen that the torque increases greatly with the tilting angle" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0003102_978-1-4419-1126-1_16-Figure16.4-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0003102_978-1-4419-1126-1_16-Figure16.4-1.png", + "caption": "Fig. 16.4 A three section swimming tail with actuators that apply: (a) torquesMi or (b) forces FLi", + "texts": [ + " In the case of a piezoelectric actuator the elastic moment, MEi(t), is MEi\u00f0t\u00de \u00bc Xn j\u00bc1 ZijYijAijdij Eij\u00f0t\u00de 8i \u00bc 1; 2; . . . ; n: (16.5) dij is the effective piezoelectric coefficient of the jth layer and Eij\u00f0t\u00de \u00bc Vij\u00f0t\u00de=tij is the electric field on the j th layer. In the case of a magnetic actuator the Lorenz force is FLi\u00f0t\u00de \u00bc NibiIi\u00f0t\u00deB0 8i \u00bc 1; 2; . . . ; n (16.6) B0 is a constant magnetic field in the direction of the beam\u2019s longitudinal axis. Ni is the number of turns in the ith coil, bi is the width of the coil, Ii(t) is the current in the coil. Figure 16.4 illustrates how the different actuators drive the elastic tail. One should notice that applying a torque or force with the same phase and amplitude nulls out the actuators in the sub-domain boundaries and the actuator behaves as a single beam actuator with the boundary conditions (16.3). The PDE set (16.2)\u2013(16.4) is solved analytically by the method of separation of variables. A more detailed solution of a piezoelectric and magnetic three sectioned swimming tail is provided in [20] and [31] accordingly" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0001334_s40435-020-00661-8-Figure3-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0001334_s40435-020-00661-8-Figure3-1.png", + "caption": "Fig. 3 Rotor forces and moments in non-rotating hubfixed coordinate system", + "texts": [ + " The equations of motion of the fuselage are those of a rigid body, formulated (15) CMy = B2 a 2 ( 1 4 0 y \u2212 1 4 y z \u2212 1 8 B2 1c + 1 3 B y 0 + 1 4 B2 y 1 ) (16) CMz = B2 a 2 ( B2 Cd 0 4a \u2212 1 2 2 0 \u2212 1 2 2 z + Cd 2 2a 2 0 + Cd 0 4a 2 x + Cd 0 4a 2 y + Cd 2 2a 2 z + 1 4 Cd 1 a 0 ( 2 y + 2 x ) \u2212 1 8 B2( 2 1c + 2 1s ) + B2 Cd 2 8a ( 2 1c + 2 1s ) + Cd 2 4a B2 2 0 + Cd 2 6a B4 2 1 + Cd 2 4a 2 0 ( 2 x + 2 y ) + Cd 1 3a B ( 0 \u2212 z ) + 0 z a ( 1 \u2212 Cd 2 ) + Cd 1 4a B2 0 + Cd 1 5a B3 1 + 1 8a B2Cd 2 ( 2 x + 2 y ) 2 1 \u2212 1 4 B2 1 ( 0 \u2212 z) + Cd 2 2a B2 0 1 + 1 6a BCd 1 1 ( 2 x + 2 y ) \u2212 Cd 2 2a B2 z 1 + 2Cd 2 5a B3 0 1 + Cd 1 6a B x ( 1s \u2212 y ) \u2212 1 3 B 0 ( 0 \u2212 z ) + 2Cd 2 3a B 0 ( 0 \u2212 z) \u2212 1 6 B 0 ( 1s x + 1c y) + Cd 2 3a B 0 ( 1s x 0 \u2212 1c y ) \u2212 1 8 B2 1 ( 1s x + 1s x + 1c y + 1c y ) + Cd 2 4a B2 1 ( 1s x \u2212 1c y ) + Cd 2 3a B 0 1 ( 2 x + 2 y )) (17) ( rb, ) = 0 + rb R ( 1c cos b + 1s sin b ) (18) 0 = x tan TPP + CT 2 \u221a 2 x + 2 0 in the body-fixed coordinate system are given by Eq.\u00a0(19). Based on Fig.\u00a01, the longitudinal equations of motion of fuselage are obtained as: and the lateral-directional equation of motions are: where the fuselage forces and moments are given by (Fig.\u00a03): (19) XF + 4\u2211 i=1 XMi = m(u\u0307 + qw \u2212 vr + gsin\ud835\udf03) ZF + 4\u2211 i=1 ZMi = m(w\u0307 + pv \u2212 qu \u2212 g cos\ud835\udf19 cos \ud835\udf03) MyF + 4\u2211 i=1 Myi \u2212 ZM1zM1 + ZM2lM2 + ZM3lM3 + ZM4lM4 \u2212 4\u2211 i=1 XMi hM \u2212 XFhF + ZFlF = Iyyq\u0307 + ( Ixx \u2212 Izz ) pr (20) MxF + 4\u2211 i=1 Mxi + 4\u2211 i=1 YMihM \u2212 ZM4yM4 + ( ZM1 + ZM3 ) yM1 + ZM2yM2 + YFhF + ZFyF = Ixxp\u0307 + ( Izz \u2212 Iyy ) qr YF + 4\u2211 i=1 YMi = m(v\u0307 + ru \u2212 wp \u2212 g sin\ud835\udf19 cos \ud835\udf03) MzF + 4\u2211 i=1 Mzi + XM1yM1 \u2212 XM2yM2 + XM3yM3 + XM4yM4 + YM1lM1 \u2212 YM2lM2 \u2212 YM3lM3 \u2212 YM4lM4 \u2212 XFyF \u2212 YFlF = Izzr\u0307 + ( Iyy \u2212 Ixx ) pq and the fuselage moments are approximated as MxF = MyF = MzF = 0 . In the proceeding equations, ( , , ) is the Euler angles, is sideslip angle, and FP presents the flight path. The subscript F denoted the fuselage and M refers to main rotors. Therefore, ( DF, SF, LF ) are fuselage drag, side force, and lift, respectively, as shown in Fig.\u00a03. Of the different examinations conducted, three rotor inflow models (i.e., uniform, Drees, and Pitt-Peters) are chosen to generate the induce velocity distribution over the rotor disc in hover and forward flight [26]. The uniform inflow model is used to show the significant reduction of the computation burden in real time simulation environment. However, to provide a more realistic representation of the inflow, in this work, first harmonic non-uniform inflow of Drees and Pitt-Peters dynamic inflow model are employed" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0001877_icarm49381.2020.9195392-Figure6-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0001877_icarm49381.2020.9195392-Figure6-1.png", + "caption": "Fig. 6. Flexural stiffness test setup comprising of an electronic microbalance and a three-axis micromanipulator testing the VT-1 forewing prototype.", + "texts": [ + " A total of six samples are tested; (a-c) CF-PETG-1,2,3 being 3D printed forewings with different scaling factor and membrane, (d) CRT, cyanoacrylate reinforced thermoformed 50\u03bcm polyvinyl chloride wing, and (ef) VT-1,2 are 150 and 300\u03bcm vacuum thermoformed PVC wing. The relation between wing aerodynamic loading and deformation is called the flexural stiffness. As a primary mechanical property we measure this relation to validate our prototypes in terms of robustness, and further compare it to the available literature. Generally in a mechanical analysis non-destructive tests (NDTs) are prioritised due to their recycling nature. As shown in Figure 6, a test platform is designed specifically to analyse the mechanical properties of our wing prototypes. The setup is standardised according to the literature [39], [40]. It consists of an elec- tronic microbalance (Sartorius BSA323S-CW, China), and a three-axis manual micromanipulator (Shengling, LD60RM, China) fitted with 10\u03bcm precision micrometer-actuated linear slides. The prototyped wing root is clamped onto the 978-1-7281-6479-3/20/$31.00 \u00a92020 IEEE 605 Authorized licensed use limited to: University of Brighton", + " left edge of the micromanipulator platform with its chord parallel to the XY plane. Here, the aerodynamic forces acting on the wing are assumed to be a concentrated point force exerted with a pin fixed vertically on the microbalance at almost 70% of the wingspan and chord for the estimation of flexural stiffness spanwise and chordwise, respectively. Upon counterclockwise rotation of the Z-axis micrometer, the pin comes in contact with the wing placed at a distance (\u03b4) from wing clamp and leading edge for spanwise and chordwise measurement, respectively. Figure 6 shows the exact force points marked on the wing prototypes. The reading on Z-axis micrometer from the count microbalance begins measurement is denoted as wing displacement (\u0394). Hence, flexural stiffness [19], [39] can be estimated as, EI = f\u03b43 3\u0394 (1) where f is the applied static force recorded on the microbalance, with the wing displacement (\u0394) limited to 5% of \u03b4, as EI reflects only minor deformation rate. A plot of flexural stiffness for different wing prototypes is illustrated in Figure 7. The graph consists of different plot types (circle, triangle, etc" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0000893_msf.982.75-Figure5-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0000893_msf.982.75-Figure5-1.png", + "caption": "Fig. 5. Comparison between the experimental and numerical deformed shapes when a high blank-holder force applied - Stress distribution of different layers along the normal direction (S11)", + "texts": [ + " Delamination: In this investigation, observation measure was used to compare the severity of delamination of the formed cup. Effect of Blank-holder Force. Simulations and experimental results provided us a comprehensive overview of the effect of blank-holder force on the of the FML system formability. Fig. 4 shows cups formed at low blank-holder force of 6 kN, all specimen in simulation and experiment\u2019s shown the presence of wrinkling, demonstrating that a low blank-holder force wrinkling as the primary failure mode. Fig. 5 shows formed cups at the high blank-holder force of 6 kN; all formed cups showed very small or non-presence of wrinkling. Some Simple failed due to tearing or fracture, signifying that a high blank holder force increases the possibility of failure due to tearing and fracture and decreases the severity of wrinkling. Effect of Cavity Pressure. Lower cavity pressure than the optimal values lead to in serious wrinkling, because of the easy flow of material inside the die due to inadequate supporting force" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0002075_s42417-020-00261-y-Figure4-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0002075_s42417-020-00261-y-Figure4-1.png", + "caption": "Fig. 4 Finite-element model of the marine twin-layer gearbox case, a the integral structure of the gearbox case, b the carrier of the encased stage mounted in the inner shell. c The output end of the gearbox case. d The input end of the gearbox case", + "texts": [ + " PL represents the equivalent force of the output torque. The substructure method that can effectively reduce the degree of freedom of the gearbox model is used to calculate the mass, stiffness and damping matrices for the marine twin-layer gearbox case. This gearbox case is made of alloy steel with a density of 7900\u00a0kg/m3, an elastic modulus of 2.05 \u00d7 105\u00a0MPa, and Poisson\u2019s ratio of 0.3. The finite element model of the twin-layer gearbox case includes 828,873 nodes and 585,756 elements, as shown in Fig.\u00a04. The spring element with stiffness and damping is used to simulate the cylindrical vibration isolator placed between the inner and outer shells. The stiffness and damping of the spring element are 7.3 \u00d7 106\u00a0N/m and 6.6 \u00d7 105\u00a0N\u00a0s/m, respectively. The master node is built in the center of the bearing hole and is rigidly coupled to the slave node on the surface of the bearing hole (see Fig.\u00a04b\u2013d). The bolt holes at the bottom of the gearbox case are given a fixed constraint. When solving, the mass, stiffness, and damping matrices of the slave nodes are all condensed to the master node. Therefore, the dynamic parameters of the gearbox case can be characterized by the mass matrix, stiffness matrix and damping matrix on the master node. After condensing, the dynamic equation of twin-layer gearbox case is expressed as where Mg, Kg, and Cg are mass, stiffness, and damping matrices of the master nodes on the twin-layer gearbox case, respectively" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0001103_s1064230720010037-Figure3-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0001103_s1064230720010037-Figure3-1.png", + "caption": "Fig. 3. Settlement scheme of the movement of the robot.", + "texts": [ + " The task of determining the optimal foot movement according to some criterion while raising and lowering it onto the supporting surface in the presence of obstacles has also not been considered. The translational motion of a robot with orthogonal walking motors, such as the Ortonog walking robot, is considered [15] (Fig. 2). The distance profile S from the foot entering the transfer phase up to an obstacle in the form of a protrusion of height H and distance L to its location at height h from the initial level (Fig. 3) is considered to be well known, for example, according to the information and measurement system of the robot. At each moment in time, one of the movers of mass m is in transport. To solve the problem, differential equations of the motion of the mass m of the foot are compiled in horizontal and vertical directions and the equation describing the uniform progressive movement of the robot\u2019s body: (1.1) where x and y are, respectively, the horizontal and vertical coordinates of the foot of the mover in absolute motion; g is the acceleration of gravity, Q is the resistance force to the movement of the robot, due, for example, to the hook load, etc", + " By defining this function for a particular robot moving in specific operating conditions, as a result of solving the problem on the minimum, it is possible to realize the optimal mode of movement. The translational movement of the Ortonog walking robot with the rectilinear movement of its center of mass along the horizontal axis at a constant speed is considered. The stride length L = 0.91 m and the mass of the transferred walking mechanism m = 70 kg. The robot moves along a horizontal deformable surface (h = 0), characterized by the dimensionless parameter \u03be, which defines various resistance forces. The profile of the supporting surface (Fig. 3) is described by the dimensionless parameters \u03b8 = 0.3, \u03bc = 0.2, and \u03c3 = 0. In real conditions, these parameters are not regulated. Thus, the goal of solving the model problem is to test the method for determining the dimensionless parameters of the control actions \u03b5 and \u03b3, which ensure the optimal motion of the robot according to the criterion of minimum energy consumption. The graphs (Figs. 5\u20137) show the corresponding dependences of heat loss per unit path A at \u03b1 = \u03b2 = 0.0001 (this is possible when using the same motors in the hoist and ( ) ( ) ( ) ( ) ( ) ( ) \u2212 = + + + + + + + + + 1 1 \u03c4 2 2 2 1 1 1 0 \u03c4 \u03c4 2 2 2 2 2 2 0 {\u03b1[ ] \u03b2 }d {\u03b1[ ] \u03b2 }d " + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0002861_978-3-030-55889-5-Figure1.32-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0002861_978-3-030-55889-5-Figure1.32-1.png", + "caption": "Fig. 1.32 Test of a transmission belt", + "texts": [ + "21 Test Functions 49 Unfortunately, however, the concept of virtual displacements is historically so burdened that one is tempted to replace itwith amore harmless term like test function. After all, testing, wiggling, deflecting from the equilibrium position is an everyday maneuver. To determine the weight of a suitcase, we lift it up. Since force = mass \u00d7 acceleration we conclude from the acceleration a and the force in the arm, the mass m of the suitcase. To determine the tension in a transmission-belt or the pressure in a football, we press our thumb against it, see Fig. 1.32. Likewise, the draftsman projects the structure onto three orthogonal planes, and the drafts are the scalar product between the structure and the two unit vectors ei , e j of each of the three planes. From the more sides we look at something, the more we learn about the object. In a figurative sense it means, the more often we wiggle an object, the more tests \u03d5i we run, the clearer the picture becomes. No, virtual displacements need not be small, see Fig. 1.33. The equation \u03b4We = \u222b l 0 p \u03b4w dx = \u222b l 0 M \u03b4M E I dx = \u03b4Wi , (1" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0002156_icee50131.2020.9260646-Figure5-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0002156_icee50131.2020.9260646-Figure5-1.png", + "caption": "Fig. 5. The magnetic flux density distribution of partial and fully HTSISM. a) Partial HTS-ISM. b) Fully HTS-ISM.", + "texts": [ + " The intensity of magnetic flux line at fully HTS-ISM is lower than partial HTS-ISM, the reason behind this difference is the angle of radiation of magnetic flux into the surface of HTS tapes, since the geometry of HTS coil are flat, the HTS coil is exposed to perpendicular magnetic flux which decreases the critical current density significantly. In another words, critical current density of the HTS tape is dependent to flux density magnitude and angle of flux radiation, so that by variation of each of these components, critical current density will be changed. Due to the reduction of critical current density, the magnetic flux intensity at fully HTS-ISM is lower than partial HTS-ISM. Fig. 5 also shows the magnetic flux density distribution of both machines. As shown in Fig. 5, none of the structures is saturated. The B-H curve of the silicon-steel sheet used in the simulation is shown in Fig.6. Torque ripple is a phenomenon considered in the design of electric motors and is related to the increase and decrease of the periodic output torque of the motor. Its value is obtained by examining the difference between the maximum and minimum electromagnetic torque over a full period and is expressed as a percentage. Torque ripple is given in relation 5: Torque ripple = T - Tmax min \u00d7100 Tavg (5) Where Tmax is the maximum torque, and Tmin is the minimum torque and Tave is the mean torque" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0002115_ab78b1-Figure3-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0002115_ab78b1-Figure3-1.png", + "caption": "Figure 3: The schematic diagram of experimental setup. The experimental setup consists of LED spotlights, LED arrays, a digital camera, and a screen. The digital camera is used to record the falling of the Yang-na seed. The position-time data is then extracted from the recorded video using Tracker.", + "texts": [], + "surrounding_texts": [ + "The experimental setup, as shown in Fig.(3), consists of a light panel, a high speed shutter camera: Model(EX-FH100), and a screen. The light panel consists of 4 LED spotlights at each corner and 10 arrays of LED. The LED light source is supplied by a 12V battery to generate stationary light at 400-800 Lux. The height of the camera and distance between camera and the screen are 1.5 and 2.5 m respectively. At the beginning of the experiment, a Yang-na seed is released vertically at 3.0 m height above the floor in a close room without air flow. The digital camera is set to record in the audio-video-interleave (.avi) format at 240 frames per second (fps). The physical resolution of our video is about 9.78 \u00d7 10\u22124 meters per pixel. The recorded videos 1 2 3 4 5 6 7 8 9 10 11 12 13 14 15 16 17 18 19 20 21 22 23 24 25 26 27 28 29 30 31 32 33 34 35 36 37 38 39 40 41 42 43 44 45 46 47 48 49 50 51 52 53 54 55 56 57 58 59 60 A cc ep te d M an us cr ip t 9 are calibrated within Tracker for the displacement of the falling seed. The vertical falling of three Yang-na seeds were independently tested. The vertical positions and time of each falling seed were acquired using Tracker analysis. An example of Tracker analysis is shown in the Fig.(4). The vertical positions and time of falling were further analyzed in the next section. 1 2 3 4 5 6 7 8 9 10 11 12 13 14 15 16 17 18 19 20 21 22 23 24 25 26 27 28 29 30 31 32 33 34 35 36 37 38 39 40 41 42 43 44 45 46 47 48 49 50 51 52 53 54 55 56 57 58 59 60 A cc ep te d M an 10" + ] + }, + { + "image_filename": "designv11_71_0001375_i2mtc43012.2020.9129496-Figure3-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0001375_i2mtc43012.2020.9129496-Figure3-1.png", + "caption": "Fig. 3. Bearing fault conditions. (a) ORWD. (b) ORWC. (c) SRD.", + "texts": [ + " The analog to digital conversion was performed by an Analog to Digital Converter (ADC), 4-Channel SingleEnded ADS7841 (from Burr-Brown), and configured to convert at 6.4 ksps to 12 bits. The acquisition was performed by a Data Acquisition System (DAS) implemented on an FPGA to manage and send the data to a Personal Computer (PC) for storage and labeling for future use. An ordinary alternator was band-couple to simulate the mechanical load on the IM. The setup is shown in Fig. 2. Four conditions are studied in this paper: healthy, ORW with a drill of 1.6mm \u2205 (ORWD Fig. 3.a), ORW with a transversal cut of 1.2mm of width (ORWC, Fig. 3.b), and a bearing without lubrication and roughness in its ORW and IRW caused by oxidation (SRD, Fig. 3.a). On the one hand, the drill and the transversal cut were accomplished to simulate real, common, and localized defects on bearing. On the other hand, the SRD bearing fault was obtained from a real operative and environmental condition to study the distributed defects. Bearings 6204 were used in this paper. 3 Authorized licensed use limited to: University College London. Downloaded on July 05,2020 at 20:18:31 UTC from IEEE Xplore. Restrictions apply. The application of the sparse representation for fault detection is based on the reconstruction of the signal (or spectrum)" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0002769_ijhvs.2019.102683-Figure1-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0002769_ijhvs.2019.102683-Figure1-1.png", + "caption": "Figure 1 (a) Torsio-elastic suspension design; (b) schematic representation; and (c) illustration of suspension kinematics leading to reduced sprung mass roll under a 15 cm high bump", + "texts": [ + " The model validity is also demonstrated by comparing the rear-axle suspended vehicle responses with the reported measured data. Optimal design parameters of the three suspension configurations are obtained using the genetic algorithm. The relative ride performance potentials are analysed in terms of frequency-weighted rms accelerations and 8-h equivalent exposure values near the driver\u2019s seat location. The torsio-elastic suspension comprises a lateral link coupling the sprung chassis to the solid axle via two torsion bars oriented along the longitudinal axis, as shown in Figure 1(a) (Pazooki et al., 2011). The suspension system is schematically shown in Figure 1(b). Each torsion bar consists of a steel rod encased in a tubular elastomer, and is tightly positioned in clevis like mountings fixed to the chassis and the solid axle. Elastomer of elliptical cross-section is preferred to achieve unequal stiffness along the vertical and lateral axes. The lateral link couples the two torsion bars with certain torsion preload. The length and diameter of the torsion rod are determined in accordance with the suspension\u2019s load carrying capacity, which also relate to the radial and torsional flexibility of the suspension. The static equilibrium orientation of the lateral link and static deformations of the torsion bar also depend on the suspension load. Relative motion between the wheel and chassis yields restoring force and roll moment due to elastic deformations of the torsion bar along the radial and torsion directions, respectively. The elastomer also provides some light damping in the radial and torsional directions. Additional dampers may be introduced for damping in multiple directions, as seen in Figure 1(b). The major advantages of the torsio-elastic suspension include: \u2022 substantially smaller vertical motion of the sprung mass compared to the unsprung mass, which is due to kinematics of the lateral link that primarily undergoes roll motion \u2022 energy absorption due to torsional deflection of the elastic shaft \u2022 high effective roll stiffness attributed to restoring moment of the elastic torsional shafts \u2022 compensation of sprung mass roll by differential orientations of the left- and rightside suspension links, as shown in Figure 1(c) \u2022 low relative vertical motion between the sprung and unsprung masses, and therefore lower sensitivity to load variations \u2022 rugged design requiring minimal maintenance. A three-dimensional ride dynamic model of an off-road vehicle is formulated with frontand rear-axle torsio-elastic suspensions, so as to investigated relative ride performance potentials of the front-, rear- and two-axle suspension configurations. Owing to light damping due to torsion bars, two additional dampers are introduced in the front- and rearaxles with inclinations \u03b8f and \u03b8r in the roll plane, respectively, with respect to the vertical axis" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0002816_thc-2010-0566-Figure12-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0002816_thc-2010-0566-Figure12-1.png", + "caption": "Fig. 12. Shear stress distribution.", + "texts": [ + " If we know the magnitude of the externally applied torsional moment, also called torque, on a cylinder of radius r then we can calculate the induced maximum shear stress at the cylinder surface using Eq. (6). \u03c4max = 2T \u03c0r3 (7) If we look at the cross section of the cylinder then we will see that the induced shear stresses vary linearly from zero in the center to a maximum value at the surface, whereas in a hollow tube the stress is a finite value at the inside diameter with maximum at the outer diameter (Fig. 12). The maximum shear stress at the hollow tube surface with outside radius, ro, and inside radius, ri, is calculated using Eq. (7). For the same applied torque and outside tube diameter, as the inside tube diameter increases the maximum induces shear stress at the outside surface increases. \u03c4max = 2Tro \u03c0(r4o \u2212 r4i ) (8) If we know the material properties then we can calculate the shear strain at each point, this shear strain cumulatively results in twisting of the cylinder. The angle of twist, \u03b8 (theta), (Fig" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0001138_0954406220925836-Figure3-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0001138_0954406220925836-Figure3-1.png", + "caption": "Figure 3. Articulated moving platforms with three directions: (a) AMP1[ AMP2[ AMP3; (b) AMP1[ AMP2[ AMP2; (c) AMP2[ AMP2[ AMP3; (d) AMP2[ AMP2[ AMP2.", + "texts": [], + "surrounding_texts": [ + "Parallel manipulator, articulated moving platform, rotational capability, type synthesis, constraint synthesis Date received: 6 March 2020; accepted: 19 April 2020" + ] + }, + { + "image_filename": "designv11_71_0000930_s1068798x20020173-Figure2-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0000930_s1068798x20020173-Figure2-1.png", + "caption": "Fig. 2. Structure of the two-mass elastic system.", + "texts": [ + " and angular velocity at the drive\u2019s motor shaft, while \u041cfm and \u03c9fm are the torque and angular velocity at the shaft of the functional mechanism. If we neglect dissipative losses in the transmissions and take into account that the rigidity of the mechanical shafts connecting the half-clutches to the other components is much higher than that of the magnetic clutch, we may combine the rotating masses of the motor and driving half-clutch and likewise combine the masses of the driven half-clutch and the functional mechanism to obtain a two-mass system (Fig. 2). In Fig. 2, the total reduced moment of inertia of the components rigidly connected to the motor is J1 = Jd + J1cl, while the total reduced moment of inertia of the components rigidly connected to the functional mechanism is J2 = Jfm + J2cl. The elastic coupling between these masses is due to the forces exerted by the magnetic field and is characterized by the equivalent rigidity \u0441m. The two-mass electromechanical system is the most expedient model of the magnetic clutch in study- RUSSIAN ing the influence of elastic couplings" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0000136_012076-Figure2-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0000136_012076-Figure2-1.png", + "caption": "Figure 2. Material cutting areas by plates in the form of Reuleaux Triangle Profile", + "texts": [ + " In this way various methods of material destruction work together: compression, stretching, abrasion, impingement attack, constantly changing the forces directions operating on the material, change of numerical values and force vectors [8, 9]. HIRM-2019 Journal of Physics: Conference Series 1353 (2019) 012076 IOP Publishing doi:10.1088/1742-6596/1353/1/012076 The grinder design with combine rollers is implemented as follows: plates in a form of Reuleaux Triangle Profile are located on the shaft with an 120\u00b0 offset relative to each other. Figure 2 shows the material cutting areas. HIRM-2019 Journal of Physics: Conference Series 1353 (2019) 012076 IOP Publishing doi:10.1088/1742-6596/1353/1/012076 During cutting, the guillotine working principle is realized (with knives rotation) with non-linear distribution load [10]. Figure 3 shows directions of effort in the cutting zone along the radius of roll Reuleaux Triangle Profile. During roll rotation, the point of forces application shifts tangentially to the curve formed by the opposite edges of Reuleaux Triangle Profile" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0003825_j.expthermflusci.2014.10.004-Figure2-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0003825_j.expthermflusci.2014.10.004-Figure2-1.png", + "caption": "Fig. 2. Photograph of the experimental apparatus.", + "texts": [ + "0 319 sensors, a connected pipeline, a PARM, an encoder, a linear guide, a payload, and an AD/DA converter. The temperatures at the inlet and end parts of the PARM are measured by temperature sensors, and data are sent to the PC through a DA converter. In the present study, two temperature sensors are used in order to confirm the temperature change at different locations. These two temperature sensors (T type) are attached at the outer surface of the PARM. One sensor is attached at the inlet part, and the other sensor is attached at the end part of the PARM, as shown in Fig. 2. Here, the temperature measured at the inlet part of the PARM is denoted as hIN, and the temperature measured at the end part of the PARM is denoted as hEND. A linear guide is used to fix the payload to the PARM. Two payloads are used in the present study. Payload 1 is a mass of 1.18 kg, and payload 2 is a mass of 2.0 kg. The supply pressure is 600 kPa (gauge). Two types of PARMs are considered herein. One type (MXAM-10-AA) has an inner diameter of 10.0 mm and natural lengths of 213 mm and 317 mm" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0002247_tcyb.2020.3036289-Figure1-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0002247_tcyb.2020.3036289-Figure1-1.png", + "caption": "Fig. 1. (a). Example of curvature-bounded homotopy that bridges the pairs x and y. The corresponding Dubins paths and adjacent circles are illustrated as well. Note that some trajectories are distorted for better visual effect. (b). Onion-like multilayer trajectory homotopy structure.", + "texts": [ + " The definitions of the other three adjacent circles, cXR, cYL, and cYR, are analogous. OXL, OXR, OYL, and OYR denote the respective centers of these circles. Definition 4 (Curvature-Bounded Homotopy) [35]: A curvature-bounded homotopy H is a homotopy function satisfying that H(u) is a curvature-bounded trajectory for all u \u2208 U. Within a curvature-bounded homotopy, the trajectory states, lengths, and midway lengths continuously vary with respect to the homotopy parameter u. An example of curvature-bounded homotopy that bridges the pair x and y is shown in Fig. 1(a), and the corresponding Dubins paths and adjacent circles are shown as well. For convenience, we will refer to curvaturebounded homotopy simply as homotopy. We introduce two models for multiple vehicle coordination problems. 1) Isolated State Consensus Model: An isolated state is a free point in Euclidean space that can be represented by a vector. The objective of isolated state consensus is the convergence of the isolated states of vehicles to a common value. Much of the literature [9], [36], [37] focuses on isolated state consensus models and algorithms for their application to vehicles", + " The following conditions must be satisfied with the design of the protocol fC: 1) the controls of state variables are independent and 2) the transformation from homotopic description to FK description, that is, the function S , is always valid. Considering the above conditions, an algorithm needs to be built to decouple the control of each trajectory state. Decoupling is based on the algebraic structure of the homotopic trajectory description. The algebraic structure of trajectory homotopy is designed with an onion-like multilayer structure, and the trajectory states can be independently controlled on each layer, as illustrated in Fig. 1(b). For a given initial trajectory, this structure can be built as follows. Select M \u2212 1 nodes y2, . . . , yM along the initial trajectory to divide it into segments. Establish trajectory homotopies to connect the endpoint pairs y1/y, y2/y,. . . , yM/y, where y1 = x. The parameters of the M homotopy functions are denoted as u = [u1 \u00b7 \u00b7 \u00b7 uM]. To independently control M midway lengths of a trajectory, we consider the following constraints on the structure: r1 \u2265 \u2016y2 \u2212 y\u2016 > r2 ... rM\u22121 \u2265 \u2016yM \u2212 y\u2016 > rM" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0003990_2014-01-0856-Figure2-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0003990_2014-01-0856-Figure2-1.png", + "caption": "Figure 2. Detailed view of MBD model", + "texts": [ + " Based on road testing and the MBD model, the loading-time histories in the suspensions are extracted for a subsequent cab durability analysis. The MBD model used for VI and loading extraction consists of a flexible frame and a rigid cab that are coupled by bushings characterized by stiffness and damping. The flexible frame is imported from Nastran finite element model, and the cab is simplified to be a mass center with the real characteristic information including location, quality, and rotational inertias obtained from the simple measurement. Figure 2 shows the view of the MBD model. In the road testing, the main relevant signals that have to be gathered are vertical, lateral and longitudinal accelerations in the frame and the cab, some for iteration and some for reference. The number of signals chosen as objective signals for iteration equals 9. The whole objective signals are provided in table 1 below. The original signals need pre-processing and identification as noted above, and after that they can be accepted as VI objective signals. For example, taking the vertical acceleration signals of the left-front suspension, a comparison of the signals in the time domain before and after these procedures and the frequency domain analysis results are depicted in figure 3: For the choice of 7-channel drive signals, the commonly used force signals are substituted by displacement signals because the former ones may lead to uncontrolled deformations in the MBD model" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0001138_0954406220925836-Figure10-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0001138_0954406220925836-Figure10-1.png", + "caption": "Figure 10. 5-DOF GPMs: (a) 3T2R-AMP1AMP2; (b) 2T3R-AMP2AMP2AMP2.", + "texts": [ + " According to the constraint synthesis method, only two combinations (C and D and C and C) are available to synthesize the 3T2R GPMs. When the R3RMPaR3R 0 3 H2C-limb and the R3RMRMR03 R03 H3C-limb serve as the connected limbs, the directions of revolute joints R3 and R03 in the H2C-limb is perpendicular to the constraint couple of H3C-limb. To guarantee this requirement is satisfied during continuous motion, the axes of joints R3 are parallel to or coaxial, and two groups of parallel revolute joints are arranged on the end-effector. As shown in Figure 10(a), the 3T2RAMP1AMP2 GPM is acquired after installing the revolute joints equipped with actuators on one line. As for the 2T3R mechanisms, the constraint-force of the wrench system can be generated by F and D, F and F. The R3RMPaR3R 0 3R D-H2limb and the R1RMRMR2R2 F-H3limb are employed to build the 2T3R GPMs. To conduct rotational motions about three directions, three groups of parallel revolute joints are assembled to the end-effector. Then the R3RMPaR3R 0 3R D-H2limb is transformed into the R3RMPaR3PR 0 3R2 D-H2limb, and the 2T3RAMP2AMP2AMP2 GPM is derived, as depicted in Figure 10(b). 6-DOF GPMs with articulated moving platforms When constructing 6-DOF mechanisms, only two hybrid limbs with three active joints can be used to fully control the GPMs, and the topological structure is drawn in Figure 7(e). With no constraint exerting on the end-effector, only unconstrainted limbs are feasible to design mechanisms. When employing two R3RMRMR03R 0 3R D-H3limbs to connect the two sides of the end-effector, three groups of parallel revolute joints are included in the limb. To make sure the hybrid limbs can perform the necessary three-dimensional translation, the R3 revolute joins in the two limbs are not parallel" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0003618_ijvsmt.2015.067521-Figure4-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0003618_ijvsmt.2015.067521-Figure4-1.png", + "caption": "Figure 4 Symbology for position analyses and location of the Ai, Bi, Ci points, (a) front view (b) top view (see online version for colours)", + "texts": [ + " Figure 3 shows the fixed frame Ox0y0z0 and the moving frame Ox1y1z1, which is attached to the car body. Owing to the fact that there are no irregularities on the flat road, the wheel vertical motion with respect to the moving frame is only a mathematical function of the body roll angle, which is represented by the angle .\u03c6 The actuators 1 and 2, and the bars 1 and 2 are in the Oy0z0 plane. The goal of the kinematic position analysis is to obtain the mathematical transformation between the end effector location, defined by the vector [\u0398,\u0393,\u03a8]T\u03c7 = [Figure 4(a) and Figure 5(a)] and the displacements provided by the actuators, defined by the vector 1 2 3[ , , ]TS s s s= [Figure 4(a) and Figure 4(b)]. Both vectors are defined in the moving frame. Regarding to the limb 1 [Figure 4(a)], the centres of the spherical joint (point D1) and the universal joint (point E1) that connect the actuator 1 are given by the equation (3): 11 1 1 1[0, cos\u0398, sen\u0398] 0, , TT ED E z= \u2212 \u2212 = \u2212\u23a1 \u23a4\u23a3 \u23a6 (3) where 1 1| 0 | .D= \u2212 In Figure 4(a), if the angle is null, the distance between points D1 and E1 is 1 .Ez However, when theta is not zero, the following equation can be obtained: ( )1 2 2 1 1 1ED E z s\u2212 = + (4) Using the equations (3) and (4), the relation between the displacement of the actuator 1, s1, and the angle \u0398 can be obtained by equation (5). 1 1 1 2 2 2 1 1 1 12 2 cos\u0398 2 sin\u0398 2 0E S Es z z+ + \u2212 \u2212 = (5) The centre of the joint that connects the limb 1 to the end-effector is defined by equation (6) [Figure 4(a)]. 1 [0, cos\u0398, sin\u0398]TA = (6) With respect to the limb 2, the origin of the displacement of the actuator 2, C2, is defined by equation (7) [Figure 4(a)]: 2 [0,0, ]TC h= (7) The centre of the revolute joint that connects the bar 2 to the actuator 2 (point B2) and the centre of the spherical joint that connects the bar 2 to the end-effector (point A2) are obtained by equation (8) [Figure 4(a) and Figure 4(b)]: [ ] [ ]2 2 20, , 0, cos\u0398 cos\u0393, sin\u0398 sin\u0393 TTB s h A h h= = + + (8) Regarding to the limb 3, the origin of the displacement provided by the actuator 3, C3, is defined by equation (9) [Figure 4(b)]: 3 [ ,0, ]TC d h= (9) The point representing the centre of the universal joint that connects the bar 3 to the actuator 3 (point B3) and the point that represents the centre of the spherical joint that connects the bar 3 to the end-effector (point A3) are obtained by equation (10) [Figure 4(b) and Figure 5(a): [ ]3 3 3 2, , TB d s h A A u\u2032= = + (10) The vector u\u2032 is calculated by applying the equation (11) which comes from the quaternion algebra (Yang and Freudenstein, 1964). ( ) ( )2\u02c6 \u02c6 \u02c6sin(\u03a8) 2sin (\u03a8 / 2)u u p u p p u\u2032 \u23a1 \u23a4= + \u00d7 + \u00d7 \u00d7\u23a3 \u23a6 (11) where the vector u is defined by equation (12), \u03a8 is the rotation angle of the vector u around the king-pin [ 1 2A A in the Figure 5(a)] and d is the distance between points A3 and A2. 1u dl= (12) The unit vector \u02c6 ,p defined by equation (13), has the direction of the king-pin" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0003929_scis-isis.2014.7044649-Figure5-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0003929_scis-isis.2014.7044649-Figure5-1.png", + "caption": "Figure 5. Angle direction in absolute direction control", + "texts": [ + " Direction control based on the compass sensor controls is used to match both orientation angles measured by the compass sensor with respect to the magnetic North Pole of the smartphone and riding robot. The rider standing on the riding robot changes the direction of the smartphone to the desired direction to make the riding robot to move in the desired direction by pressing a start button on the screen of the 978-1-4799-5955-6/14/$31.00 \u00a92014 IEEE 19 smartphone. Assuming that the rider moves the smartphone to the desired angle by rotating the smartphone to the left or right as shown in Fig. 5(a), the orientation angle of the smartphone s\u03b8 is measured by the compass sensor with respect to the magnetic North Pole. s\u03b8 is contained in the command packet that is transmitted to the riding robot through Bluetooth wireless communication in 100 ms intervals. The riding robot parses the s\u03b8 from the command packet transmitted from the smartphone. The riding robot drives a motor mounted on both sides for matching the and its own heading angle c\u03b8 . In order to perform direction control based on the compass sensor, the velocities of both sides of the motor lv and rv are given by ( ) ( ) l m p s c r m p s c v s k v s k \u03b8 \u03b8 \u03b8 \u03b8 = + \u2212 = + + (5) where ms is the forward velocity of the riding robot, and ( )p s ck \u03b8 \u03b8\u00b1 is a proportion controller depending on the direction angle of the riding robot" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0001209_icemi46757.2019.9101798-Figure2-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0001209_icemi46757.2019.9101798-Figure2-1.png", + "caption": "Fig. 2 Simplified single axle dynamic model", + "texts": [ + " The running process of the locomotive is a highly nonlinear and complicated problem, and the accurate dynamic model of the running process is hard to obtain. 1201 Authorized licensed use limited to: University of New South Wales. Downloaded on July 26,2020 at 04:06:41 UTC from IEEE Xplore. Restrictions apply. To facilitate the research in adhesion control and reduce the complexity of modeling, the simplified locomotive dynamics model is generally used[15]. The single-axis simplified dynamic model is shown in Fig.2. The dynamic mathematic model of the locomotive is as follows: t s d dvM F F dt (5) = ( )sF vs W g (6) 2( ) ( )d t t tF v a bv cv Mg (7) where, M is the mass of the train; tv is the locomotive velocity; ( )d tF v is the running resistant when operating; , ,a b c is the basic resistant coefficient. Supposing that the gear box m g d wR w , according to the characteristic of the traction system, the mathematic equation is described as follows: m m m L dwJ T T dt (8) ( )L g rT vs W g R (9) 2 d equ m g JJ J R (10) where, sF is the adhesion force between the wheel and rail; mJ is the motor moment of inertia; dJ is the wheelset moment of inertia; equJ is the average wheel inertia; mT is the motor torque; LT is the equivalent load torque on the motor of the adhesion force; W g is the axle weight" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0002368_icicm50929.2020.9292221-Figure4-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0002368_icicm50929.2020.9292221-Figure4-1.png", + "caption": "Figure 4. Ultrasound detection data.", + "texts": [], + "surrounding_texts": [ + "When a partial discharge occurs in the GIS, impact vibration and sound will be generated, so the partial discharge signal can be measured by installing an ultrasonic sensor on the outer wall of the cavity, as shown in Figure 2. The advantage of this method is that there is no connection between the sensor and the electrical circuit of the GIS equipment, and it is not subject to electrical interference. The basic method of the ultrasonic positioning method is to measure the time delay or time difference of the ultrasonic signal generated by the partial discharge to multiple ultrasonic sensors at different positions. The method calculates the positioning of the local discharge source and realizes the positioning of insulation defects." + ] + }, + { + "image_filename": "designv11_71_0003851_e2014-02130-2-Figure4-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0003851_e2014-02130-2-Figure4-1.png", + "caption": "Fig. 4. Beam, which position is described by 3 angles.", + "texts": [ + " As the consequence, numerical integration is ambigious, since it is unknown which solution has to be chosen, similarly as in Sect. 2.2.1 2.2.3 Three angles\u2019 desription In this section, motion of the beam is describe using three independent variables. Cartesian coordinates of first beam\u2019s end A can be obtained by angles: \u03b1 and \u03b2, like in spherical pendulum. Third variable is is given by the following geometric construction: draw a line from point O2 to point A, the space between beam and strings is divided to 2 triangles: BO2A and O1O2A (Figure 4). The angle \u03b3, between those triangles, describes unequivocally position of point B. To fully describe the motion of the beam, Cartesian coordinates of points: A, B and C have to be known. Point A is described by angles \u03b1 and \u03b2. Coordinates of point C are the average of the coordinates of point A and B. To obtain point B one has to add auxiliary variables. Let us mark line segment O2A as d. Basing on cosine theorem, one obtains: d 2 = l2 + b2 \u2212 2lb cos(\u03b1 + \u03c0 2 ) = l 2 + b2 + 2lb sin\u03b1. Marking angle \u2220O1O2A as \u03d5, sin\u03d5 = l cos\u03b1 d , cos\u03d5 = b+l sin\u03b1 d and \u2220O1AO2 as \u03b8, then using sine theorem for O1O2A one receives: l sin\u03d5 = b sin \u03b8 , sin \u03b8 = b sin\u03d5 l = b cos\u03b1 d . From cosine theorem cos \u03b8 can be obtained: cos \u03b8 = b2\u2212l2\u2212d2 \u22122ld = l+b sin\u03b1 d . Due to the similarity between triangles O1O2A and ABO2 we have the following relations between angles: \u2220O1O2A is equal to \u2220O2AB, \u2220O1AO2 is equal to \u2220AO2B. The height l1 of triangle ABO2 is given by l1 = l sin \u03b8 = bl cos\u03b1 d and it crosses O2A at point D. Denoting O2D as l2, we have l2 = l cos \u03b8 = l2+bl sin\u03b1 d . Several transformations of local B coordinate system have to be performed as in Fig. 4, in order to obtain Cartesian coordinates of point B corresponding to main coordinate system. Let us consider B coordinate system as follow: x axis is parallel to d, y axis is perpendicular to d and center of this system is in point B. First, the translation of the system from point B to point D by l1 about z6 axis is performed. In the next step, rotation of system by \u03b8 about x5 axis takes place. Then, translation from point D to point O2 by l2 along x4 axis occurs. Furthermore, one rotates system by \u03d5 about y3 axis and by \u03b2 about x2 axis" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0003909_mmar.2014.6957361-Figure1-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0003909_mmar.2014.6957361-Figure1-1.png", + "caption": "Fig. 1. Planar model of a crane, where M, m, l, u and \u03b1 are, respectively, masses of a crane and payload, rope length, controlling signal corresponding to the control force acting on a crane, and sway angle of a payload.", + "texts": [ + " Section II describes the assumptions for a planar crane model and TS fuzzy control scheme. Section III presents the iterative procedure and EA employed to design a fuzzy controller. Numerical examples and conclusions are delivered in section IV and V, respectively. II. FUZZY CONTROL SCHEME FOR CRANE MODEL The planar crane system is considered as a point-mass payload suspended at the end of a rope, which is simplified to the massless and rigid link connecting a payload with a cart which motion is affected by the control signal corresponding to the control force acting on a cart (Fig. 1). Linearizing the motion equations (1) of a cart-pendulum system (through assuming: \u03b1\u03b1 \u2245sin , 1cos \u2245\u03b1 , and 0sin2 \u2245\u03b1\u03b1 ) derived from the Lagrange's second law type equation, the model can be presented in the form of two transfer functions (2-3) representing the linear second-order oscillatory (2) and astatic (3) systems with varying parameters l and m. 258978-1-4799-5081-2/14/$31.00 \u00a92014 IEEE 0sincos sincos)( 2 =++ =\u2212++ \u03b1\u03b1\u03b1 \u03b1\u03b1\u03b1\u03b1 mgmlxm umlmlxmM (1) 22)( )( nssU s \u03c9 \u03b1 + \u03a9\u2212= (2) 2 2 )( )( s gls s sX \u2212\u2212= \u03b1 (3) where: Ml 1=\u03a9 , l g M m n \u239f \u23a0 \u239e \u239c \u239d \u239b += 1\u03c9 - the natural not dumped pulsation of a system, g = 9", + ", n, and determine the controllers parameters Kk according to (6); 4: complete the rules base according to the assumption 2; 5: test the condition (9) for operating points (17) through interpolating controllers parameters according to the function (15); 6: if the condition (9) is satisfied 7: remove the fuzzy set and rules created in steps 2 and 4; 8: s := s + 1, r := r - 1; 10: assume br-1 := ls-1 and repeat the step 3 to recalculate the parameters of rules conclusions created in step 4; 11: end if; 12: end while; The EA-based method is utilized to design a TS fuzzy controller without initial conditions i) and ii) assumed in the previous subsection. A single proposition of a fuzzy controller can be presented by a real-valued chromosome consisting of the vectors of membership functions parameters (Fig. 2): [ ] [ ]110110 ,,,,,,,,, ++ == rrnn bbbbaaaa \u2026\u2026 ba (18) The objective of the evolutionary algorithm is to determine the minimum set of membership functions parameters, including the boundary parameters a0, an-1, b0 and br+1, to satisfy the condition (9) in the presence of varying parameters l and m of a system (Fig. 1). The fitness of an individual is determined through testing the condition (9) for most hazardous operating points corresponding to the all possible combinations of the crossover points of triangular membership functions (Fig. 2), their midpoints, and the bounds of system parameters intervals [l -, l +] and [m -, m +]. The objective function is formulated as a sum of the normalized distances between closed-loop system characteristic equation coefficients at the most hazardous operating points and the closest bounds of desired polynomials coefficients intervals associating with the fuzzy controller's rules which has been activated with the firing strength factor wk > 0: \u2211\u2211\u2211 = = = \u2212+ + \u2212+ \u2212 \u239f\u239f \u239f \u23a0 \u239e \u239c\u239c \u239c \u239d \u239b \u2212 \u2212 \u2212 \u2212 \u22c5\u22c5= H h N k z kzz kzz kzz kzz pp pp pp pp f 1 1 3 0 21 ,min\u03b2\u03b2 (19) where: \u23a9 \u23a8 \u23a7 > = = 0,1 0,0 1 k k w w \u03b2 , [ ] [ ]\u23a9 \u23a8 \u23a7 \u2209\u22c5> \u2208\u22c5= k k mlmlif mlmlif PKS PKS ),(),(,1 ),(),(,1 2 2 \u03b2 \u03b2 , pz - the closed-loop characteristic polynomial coefficients, H - the number of operating points at which the condition (9) is tested, \u03b22 > 1 - the factor used to penalize an individual for violating the condition (9)", + " The arithmetical-based crossover and non-uniform mutation results in adding to the current population (\u03bb1+\u03bb2) new individuals \u03bb3 and \u03bb4, respectively, hence the final population size in a single generation equals to \u03bb1+\u03bb2+\u03bb3+\u03bb4, and from this group of individuals, the new population \u03bb1 is selected using tournament method to be the parents of the next generation. The genes deletion/insertion mutation ensures diversity of genome size, whiles the recombination and non-uniform mutation results in fine exploration of the best promising regions of solution space by tuning the membership functions parameters. IV. NUMERICAL EXPERIMENTS The iterative procedure and EA described in previous section were used to design a fuzzy controller for a cartpendulum model (Fig. 1) for following assumptions: i) the constant mass of a crane equals to M = 200kg, ii) desired stable poles intervals specified at each operating point corresponding to the midpoints of membership functions (ai, bj) is formulated as follows: [ ] [ ] ( )[ ]\u03be\u03c9 \u2212== +\u2212 1/2, lgsss nkk \u2213 (20) where \u03b6 is a parameter determining the width of a desired stable poles interval. The numerical experiments were conducted to design a fuzzy controller for the value of a parameter \u03be and intervals of scheduling variables, which were assumed in the first experiment as \u03be = 0" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0002144_icem49940.2020.9270848-Figure7-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0002144_icem49940.2020.9270848-Figure7-1.png", + "caption": "Fig. 7. Left: CAD drawing of the prototype; Right: Picture of the manufactured prototype.", + "texts": [ + " As predicted the reduction effect on the steady component of the torque was only 0.3%. The topology of the manufactured prototype is chosen as described in Sec. IV. The machine is not especially optimized and just a preliminary design for a proof of concept. The rotor disks are made of uniformed structural steel 9S20K. Quarter disk shaped magnets are glued to the rotor disk. The air cored stator is based on a 3D printed housing made from polylactide (PLA). Computer-aided design (CAD) drawings of the stator and rotor are shown in Fig. 7. In the housing, coil holders are integrated, which simplify the manufacturing process of the winding. The windings are connected in star topology. Small ventilation slots are provided to ensure a convectional air cooling. More material and construction parameters are listed in Tab. III, IV. All measurements are conducted on the test bench shown in Fig. 8. The rotary speed and mechanical position is mea- TABLE III MATERIAL DATA Stator symbol value sured with an Incremental Encoder HEDS from Broadcom with 1024 steps per turn" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0000356_iecon.2019.8927121-Figure8-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0000356_iecon.2019.8927121-Figure8-1.png", + "caption": "Fig. 8. A modular integrated PM RLM [23]", + "texts": [ + " 7, was already widely investigated in the literature in the case of planar (surface) motors [17]. A first set of its coils were placed on the axial direction, to ensure the rotation, while the other one was distributed circumferentially to provide the axial force [19], [20]. Several such RLMs were developed by P. Bolognesi and his collaborators [21], [22], [23]. These all were of modular construction, an efficient and relatively new approach in electrical machines [24]. The simplest, three-phase variant with two stator modules is given in Fig. 8. The stator modules have salient poles with concentrated coils, which can be connected in different ways to obtain a poly-phase symmetrical winding. The modules are very similar with the stators of the brushless dc machines. The mover is an empty ferromagnetic cylinder, having an even number of surface mounted PMs, alternatively magnetized on the radial direction. The linear stroke of the RLM depends on the distance between the stator modules and the axial length of the PM array. This LRM can be constructed in diverse variants with different number of stator modules (which in some cases can also be shifted symmetrically on the radial direction)", + " Comparing the structure of the integrated variants with other PM RLMs, it can be stated that it is simple, and has short axial length. It can be manufactured via widespread techniques and by using common materials. Also, it's very good decoupled control of the axial force (F) and torque (T) should be obligatory mentioned. This enables the independent control of the rotary and linear motions to be performed by the RLM. This feature is well-demonstrated by the simplified mathematical model of the machine shown in Fig. 8 [23]: ))(( )( 0 qq dd BA BA iyliyp ii lT F (1) where y is the axial position of the mover, p the number of pole pairs, l the linear stroke and 0 the transformed magnetic flux at no-current condition. In the case of the currents, the A and B subscripts are related to the two stator modules (module A and B), and d and q to the two axes in quadrature. By analyzing equation (1) it can be easily observed, that the axial force is linearly proportional to the difference between the two d-axis currents components, while the torque generated by the RLM is equal to a linear combination of q-axis currents with coefficients that vary in complementary proportional way with the linear position" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0002355_j.egyr.2020.11.179-Figure14-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0002355_j.egyr.2020.11.179-Figure14-1.png", + "caption": "Fig. 14. BEESM1 with skewed rotor. (a) Structure. (b) 2-D section view.", + "texts": [ + "84 W, respectively. Since the currents of armature winding and field winding of BEESM1 are the same as those of the BEESM2, the field and armature winding copper losses of them are the same, which are 4.13 and 6.18 W, respectively. Therefore, the efficiencies of BEESM1 and BEESM2 are 91.1% and 89.6%, respectively. The losses and efficiencies of both machines are summarized in Table 2. To reduce the torque ripple, the rotor of BEESM is skewed by \u03c0 electrical degree. The BEESM1 with skewed otor is shown in Fig. 14(a). The BEESM1 with skewed rotor can be regarded as a combination of left and right arts, and the rotor tooth axes of both parts are differed by \u03c0 electrical degree. Meanwhile, the field windings of eft and right parts of BEESM are cross connected. The 2-D cross section and magnetizing flux of both parts of EESM1 are given in Fig. 14(b). The output torques of BEESM1 and BEESM2 with skewed rotor are given in ig. 15. The torque ripples of BEESM1 and BEESM2 are 30.56% and 9.28%, respectively. Comparing with the utput torque in Fig. 12, it can be found that the torque ripples of both machines can be effectively suppressed by kewing the rotor and the BEESM2 has the lower torque ripple than the BEESM1. . Conclusion This paper has proposed a novel brushless electrically excited synchronous machine (BEESM) with arc-shaped otor structure" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0003666_s1068371214050101-Figure2-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0003666_s1068371214050101-Figure2-1.png", + "caption": "Fig. 2. Coordinate transformation of five phase induction motor.", + "texts": [ + " 5 2014 ROTOR FIELD ORIENTED CONTROL STRATEGY 337 composed of many pairs of d\u2013q components. That is, the first and the second rows constitute d1\u2013q1 subspace while the third and the fourth rows constitute d2\u2013q2 subspace, and so on. (9) Multiple phase induction motor contains many orthogonal d\u2013q planes, which means the induction motor has a large amount of control degrees of free dom. For the five phase system, it possesses d1\u2013q1 and d2\u2013q2 plane, having two control degrees of freedom. The diagram of five phase induction motor coordi nate transformation is shown in Fig. 2. According to the formula (9) and Fig. 2, the coor dinate transformation matrix of five phase system is as formula (10). (10) Cs 2 n \u03d51( )cos \u03d51 \u03be\u2013( )cos \u03d51 2\u03be\u2013( )cos \u22c5\u22c5\u22c5 \u03d51 n 1\u2013( )\u03be\u2013( )cos \u03d51( )sin\u2013 \u03d51 \u03be\u2013( )sin\u2013 \u03d51 2\u03be\u2013( )sin\u2013 \u22c5\u22c5\u22c5 \u03d51 n 1\u2013( )\u03be\u2013( )sin\u2013 \u03d52( )cos \u03d52 2\u03be\u2013( )cos \u03d52 4\u03be\u2013( )cos \u22c5\u22c5\u22c5 \u03d52 2 n 1\u2013( )\u03be\u2013( )cos \u03d52( )sin\u2013 \u03d52 2\u03be\u2013( )sin\u2013 \u03d52 4\u03be\u2013( )sin\u2013 \u22c5\u22c5\u22c5 \u03d52 2 n 1\u2013( )\u03be\u2013( )sin\u2013 \u03d53( )cos \u03d53 3\u03be\u2013( )cos \u03d53 6\u03be\u2013( )cos \u22c5\u22c5\u22c5 \u03d53 3 n 1\u2013( )\u03be\u2013( )cos \u03d53( )sin\u2013 \u03d53 3\u03be\u2013( )sin\u2013 \u03d53 6\u03be\u2013( )sin\u2013 \u22c5\u22c5\u22c5 \u03d53 3 n 1\u2013( )\u03be\u2013( )sin\u2013 \u22c5\u22c5\u22c5 \u03d5 v ( )cos \u03d5 v v\u03be\u2013( )cos \u03d5 v 2v\u03be\u2013( )cos \u22c5\u22c5\u22c5 \u03d5 v v n 1\u2013( )\u03be\u2013( )cos \u03d5 v ( )sin\u2013 \u03d5 v v\u03be\u2013( )sin\u2013 \u03d5 v v\u03be\u2013( )sin\u2013 \u22c5\u22c5\u22c5 \u03d5 v v n 1\u2013( )\u03be\u2013( )sin\u2013 1/ 2 1/ 2 1/ 2 \u22c5\u22c5\u22c5 1/ 2 1/ 2 1\u2013 / 2 1/ 2 \u22c5\u22c5\u22c5 1\u2013 / 2 " + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0000058_icmae.2019.8881011-Figure6-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0000058_icmae.2019.8881011-Figure6-1.png", + "caption": "Figure 6. Initial position of the platform", + "texts": [ + " (15) The control law must consider the dynamics of the error between the desired signal and the current signal, with the purpose of feedback the robot variables and make corrections if it is necessary. This dynamic is considered in equation (16) (13). (16) III. ANALYSIS OF RESULTS AND DISCUSSION For the verification of the results, an algorithm with the calculations described above was performed and validated using MatLab\u00ae SimMechanics tool, considering the robot dimensions described in Table III, the simulation in figure 5 and the controller constant K of 10. (6) (7) (8) (9) The initial position of the robot is specified as illustrated in figure 6, with angle at the joints equal to zero. Different trajectories are plotted in order to verify the calculations raised in the inverse kinematics algorithm, starting with the values shown in Table IV, and the results obtained in figure 7. For the verification of the direct kinematics algorithm, the values from Table V and the results obtained in figure 8 are stipulated. Finally, a helical trajectory is traced to verify the behavior of the platform in the face of variations in time. From what is obtained the behavior of figure 9 in angular positions, figure 10 for positions in X, Y and Z, and figure 11 for positions in" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0001259_j.matpr.2020.05.353-Figure3-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0001259_j.matpr.2020.05.353-Figure3-1.png", + "caption": "Fig. 3. Solidworks design.", + "texts": [ + " Pneumatic cylinder punches the bag with a V shaped Die to make the handles Solid works Design: The solid works design describes the modelling of semi-automatic cloth bag making machine. First the solid works model are designed and checked for the stability of the design by using the features available in the software and the 3D model is validated to ensure the smooth working. This proposed method is fully based on Arduino programming and some of mechanical components are fabricated using mild steel. As shown in Fig. 3. The solid works model is designed with provisions where the two stitching machines are placed exactly at the two sides of the table. Pneumatic piston is used for the purpose of thread cutting and making handle for the bag. By the way of design, it is ensured that the Semi-automatic cloth bag making machine can fabricate cloth bag with the high efficiency [7,8]. Operation Feasibility: The Mechanical setup is designed and modelled using solid works software and are analyzed in the ANSYS. The shaft is tested for its load bearing capacity" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0000896_0954406220912005-Figure7-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0000896_0954406220912005-Figure7-1.png", + "caption": "Figure 7. Kinematic scheme of the ith leg of an RSRR wheel in the xli yli plane.", + "texts": [ + " To perform this analysis, equations (4) and (7) are considered in all mechanisms. These equations represent a system of four nonlinear equations. The unknowns are the pair of passive joint variables, and variables qr and qp. Following is the analysis of the RSRR wheel. In this analysis, the system of equations is reduced to a pair of quadratic equations, which can be obtained for the RSRP and RSPR wheels. Subsequently, the solution for the RSPP wheel is presented. To analyze this mechanism, consider the scheme of the ith leg shown in Figure 7. The orientation of the traction link is indicated by the auxiliary variable 2 \u00bc 2 qr l, which defines the angle of bi with respect to the axis xli where l \u00bc \u00fe i =2. Similarly, the orientation of the proximal link is defined by 3 \u00bc 3 qr l. The leg extension depends on the value of Sp. Therefore, the magnitude of the vector ti depends only on the value of the variable qp. The system of equations (4) and (7) can be reduced to two quadratic equations, whose unknowns are the length Sp and the orientation of the traction link given by 2 a1S 2 pu 2 2 \u00fe a2Spu 2 2 \u00fe a3Spu 2 \u00fe a4S 2 p \u00fe a5u 2 2 \u00fe a6Sp \u00fe a7u 2 \u00fe a8 \u00bc 0 \u00f047\u00de b1S 2 pu 2 2 \u00fe b2Spu 2 2 \u00fe b3Spu 2 \u00fe b4S 2 p \u00fe b5u 2 2 \u00fe b6Sp \u00fe b7 \u00bc 0 \u00f048\u00de where the coefficients a1 to a8 are given in Table 2, and the coefficients of the second polynomial are b1 \u00bc 1, b2 \u00bc 2lT sin#, b3 \u00bc 4lT cos#, b4 \u00bc 1, b5 \u00bc l2T t2, b6 \u00bc 2lT sin#, and b7 \u00bc l2T t2" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0003636_978-81-322-1768-8_8-Figure2-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0003636_978-81-322-1768-8_8-Figure2-1.png", + "caption": "Fig. 2 Classification of faults", + "texts": [ + " If condition monitoring has been carried out in this case complete destruction of the motor could have been avoided and hence a huge amount of cost might have been saved. Thus CM is desirable for such cases. This motivates the need for developing new reliable and efficient CM methods to avoid such damages. The most prevalent faults in Induction Motor are briefly categorized as \u2022 Rotor faults \u2022 Bearing faults \u2022 Eccentricity faults \u2022 Stator faults. The surveys indicate that in general, failures in electrical machines are dominated by bearing and stator faults with rotor winding problems being less frequent. Figure 2 shows the statistical spread in the various dominant mechanisms. There are many papers regarding rotor condition monitoring, and rotor fault detection. Some of these works have mainly used the frequency spectrum of the stator current for rotor condition monitoring. The rotor magnetic field orientation pendulous oscillation, due to broken bars, was recently presented as an index for rotor fault diagnostic purposes. There are other techniques which are based on the artificial intelligence and data mining methods as well as some investigations based on parameter estimation or parameter identification techniques" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0000316_icems.2019.8921951-Figure11-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0000316_icems.2019.8921951-Figure11-1.png", + "caption": "Fig. 11 Fourth-order natural frequencies of rotor system", + "texts": [ + "7 times the secondorder critical speed [7]. Modal analysis and calculation of natural frequency are the key steps to determine the critical speed of the rotor system. The calculation of the frequency and the analysis of the vibration mode can make the motor avoid the resonance phenomenon. The finite element method is used to analyze the modality of the rotor system, and the first four natural frequencies and vibration modes of the rotor system are obtained as shown in Fig. 8 to Fig. 9. From Fig. 8 to Fig. 11, the natural frequencies of the various stages of the rotor are known, and the critical speeds corresponding to the various circles are calculated as shown in Table \u2163. It can be seen from Table 4 that the first-order critical speed of the rotor of the motor is much larger than the rated working speed of the rotor, and the motor does not resonate during normal operation, which proves that the rotor design is reasonable and can ensure safe and reliable operation. In the design of high-speed motors, the rotor's operating speed needs to avoid the first-order critical speed" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0001034_ec-06-2019-0272-Figure1-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0001034_ec-06-2019-0272-Figure1-1.png", + "caption": "Figure 1. Conceptual model and structure of GTS", + "texts": [ + " Dynamic reliability assessment methodology To solve the above issues, this paper takes a 2.0MW class GTS of WT to study the dynamic reliability modeling method with considering time-varying mesh stiffness, timevarying damping, input varying load, bearings support stiffness, bearings support damping, random errors of different part with its phase angles, random errors based failure dependency and the effects of non-normal random load. 2.1 Equivalent dynamics model 2.0 MW WT gearbox is composed of two planetary and one parallel stage, its schematic diagram is shown in Figure 1: where sj, pj, rj and cj denote the sun gear, planet gear, ring gear and planet carrier of the j-th planet stage, respectively; g1 and g2 denote driven and driving gear of the parallel stage. There are three helical planet gears in every planetary stage. In the same planetary stage: the three planet gears distributed evenly along the circumference direction and their rotate speeds are same at any time, namely, v j pi \u00bc v j p; the external working pressure angles between each of the planet gear and sun gear are same, namely, aj spi \u00bc aj w; the internal working pressure angle between each planet gear and ring gear are equal, namely, aj rpi \u00bc aj n" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0002144_icem49940.2020.9270848-Figure8-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0002144_icem49940.2020.9270848-Figure8-1.png", + "caption": "Fig. 8. Overview picture of the test bench.", + "texts": [ + " The air cored stator is based on a 3D printed housing made from polylactide (PLA). Computer-aided design (CAD) drawings of the stator and rotor are shown in Fig. 7. In the housing, coil holders are integrated, which simplify the manufacturing process of the winding. The windings are connected in star topology. Small ventilation slots are provided to ensure a convectional air cooling. More material and construction parameters are listed in Tab. III, IV. All measurements are conducted on the test bench shown in Fig. 8. The rotary speed and mechanical position is mea- TABLE III MATERIAL DATA Stator symbol value sured with an Incremental Encoder HEDS from Broadcom with 1024 steps per turn. This leads to an angular resolution of 0.09 rad. For torque measurement a dual range torque sensor of the type 4503A by Kistler Instrumente GmbH is used. The sensor has a nominal range of \u00b11Nm. The measurement technology is connected to the load machine and prototype via backlash free couplings. The load machine is a speed controlled Induction Machine (IM) by AMK with a speed rating of nnom = 7000min\u22121 and nominal torque of Tnom = 2Nm" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0001892_012007-Figure2-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0001892_012007-Figure2-1.png", + "caption": "Figure 2. Display (a) 3D dual stator AFPM engine, (b) 3D sketch pf a double stator AFPM machine [6]", + "texts": [ + " AFPM is more compact than the RFPM engine shown in Figure 1; it is recognized as having the ability to get more power density than an RF engine [4]. AFPM machines are recognized that can get better power density than RFPM machines. Therefore AFPM can be applied because it is efficient and effective, and its capabilities are superior to RFPM [5] To assess the potential of a multi-stator engine, a double stator AFPM engine is considered. This machine has two disks (stator) that support entanglement (Figure 2). On the rotor, parts are connected three magnetic disks which are connected to each other in order to act as a driving shaft in the drive motor system which will later be connected to the wheels. In this paper, the axial flux motor is designed with a multi-sided model that has two winding disks, which become the stator while three interconnected magnetic disks which become the rotor Figure 2. ICECME 2019 IOP Conf. Series: Materials Science and Engineering 931 (2020) 012007 IOP Publishing doi:10.1088/1757-899X/931/1/012007 To develop an axial flux motor requires an initial step by determining the parameters as targets to be achieved in Table 1. The multi-sided flux axial motor design in this paper is designed with a coreless type (figure 3), where there is no metal element on the winding disk. Still, here the windings are coated with resin material. So there is no slot winding (slot less)" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0003251_icuas.2015.7152410-Figure1-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0003251_icuas.2015.7152410-Figure1-1.png", + "caption": "Figure 1 Flow and aerodynamic of gyroplane and helicopter", + "texts": [ + " Bin Xu is with the Vehicle Research Center, Beijing Institute of Technology, Beijing, 100081 China. (e-mail: xubin@bit.edu.cn). Han Han is with School of Mechanical Engineering, Beijing Institute of Technology, Beijing, 100081 China. (e-mail: hanhan2@126.com) Liyuan Liu is with Vehicle Research Center, Beijing Institute of Techno logy, Beijing, 100081 China. (e-mail:liuliyuanbit@gmail.com) autorotative working state, the power of spinning the rotor comes from a relative flow that is directed upward through the rotor disk (fig. 1). In terms of power, gyroplane rotor absorbs the wind energy from a relative flow while helicopter rotor consumes the inputting engine power. If unfortunately helicopter engines shut off, the abnormal emergency working state is same as normal one of straight-and-level forward flight. 978-1-4799-6009-5/15/$31.00 \u00a92015 IEEE 1178 In the same way, it is easy to see wind turbine absorbs wind energy just like the gyroplane. The relative flow makes gyroplane rotor work in autorotative state without power export but wind turbine outputs the power from wind to gain green electric energy (fig" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0001360_apec39645.2020.9124433-Figure1-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0001360_apec39645.2020.9124433-Figure1-1.png", + "caption": "Fig. 1 Blade bearing test rig (a) front view (b) back view.", + "texts": [ + ",M +B is [2]: where the notation F and F\u22121 indicate discrete Fourier transform (DFT) and inverse DFT. u(t) is defined as: u(t) = 1, t = M + 1,M + B 2 + 1 2, t = M + 2,M + 3, ...,M + B 2 0, t = M + B 2 + 2, ...,M +B (21) The discrete Hilbert envelope hd(t) can be derived as: hd(t) = \u221a Re[zd(t)]2 + Im[zd(t)]2 (22) where the notation Re and Im are the real part and imaginary part of hd(t). Finally, the frequency spectrum is analyzed with the aim of detecting the presence of bearing FCFs. The industrial-scale and slow-speed wind turbine blade bearing (see Fig.1) is used in our study. The geometric parameters of the bearing are listed in Table II. According to Eq.1 to Eq.3. its theoretical FCFs are calculated in Table III. The test blade bearing has been in service for 15 years in the wind farm, so the possible faults inside the bearing are naturally formed. Unlike artificial defect bearings, the test blade bearing provides the opportunities to demonstrate real-world blade bearing AE characteristics. In order to diagnose this bearing, we built a blade bearing test-rig to imitate the wind turbine pitch system. As presented in Fig.1(a), the test bearing is able to be rotated via the chain drive system; the three-phase induction motor coupled with the gearbox can produce power to drive the bearing within 5 rpm and the motor inverter can control the bearing rotation speeds. In the experiment, to provide a more realistic simulation of blade bearing working environment, a jack is utilized to provide a radial load to the bearing (see Fig.1(b)). Below the jack, we use a load sensor to measure the provided forces. When the bearing rotates, the bearing acoustic emission signals can be acquired through the PK15I type AE sensor which is mounted on the bearing outer ring surface; and a high-speed data acquisition device is used to collect data at the sampling rate of 100 kS/s. Fig.2(a) shows the blade bearing raw AE signal where the rotation speed is at 1.86 rpm. For this case study, the load is removed from the test rig. As can be seen, the raw signal is masked by heavy noise and the 148 Authorized licensed use limited to: University of Exeter" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0003753_amr.891-892.81-Figure1-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0003753_amr.891-892.81-Figure1-1.png", + "caption": "Figure 1. Photographs of the brake disc, which had a through-crack. The area boxed in (a) is shown at higher magnification in (b).", + "texts": [ + " Details of the forensic investigations relating to each case study are provided and the ensuing remedial actions discussed. Fatigue cracking is by far the most common cause of failure for aircraft components, accounting for more than half of all failures. Identifying the causes of fatigue cracking is vital in ensuring aircraft safety and preventing recurrence of the failure. This paper outlines examples of typical fatigue crack investigations conducted by the Aircraft Forensic Engineering group at DSTO. During servicing of an ADF rotary wing aircraft, the main rotor brake disc was found to have cracked through, Figure 1. The brake disc had reportedly been visually inspected 113 airframe hours prior discovery of the crack. 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-13/07/15,19:16:29) Examination of the brake disc revealed that, in addition to the through-crack, the surface of the disc exhibited severe craze-cracking as well as radial cracks up to 13 mm in length, Figure 2" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0001295_tmech.2020.3002396-Figure6-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0001295_tmech.2020.3002396-Figure6-1.png", + "caption": "Fig. 6. Device for measuring the foot position/orientation and ZMP of the human. It consists of two data acquisition devices for the joints and two force/torque sensors for the feet. Each data acquisition device for the joints has 6 DOFs for measuring motion in the Cartesian space, and each force/torque sensor has 6 DOFs for measuring the ground reaction force. The measurement device includes an onboard battery and PC to provide mobility.", + "texts": [ + " 1083-4435 (c) 2020 IEEE. Personal use is permitted, but republication/redistribution requires IEEE permission. See http://www.ieee.org/publications_standards/publications/rights/index.html for more information. To express the motion of the lower body, the position and orientation of both feet and the ZMP of the human should be measured. We measured the position and orientation of each foot with a data acquisition device, and the ZMP of the human was measured with a force/torque sensor, as shown in Fig. 6. In this study, we used the HUBO2+ humanoid platform developed by the Korea Advanced Institute of Science and Technology (KAIST) Humanoid Research Center, as shown in Fig. 7 [34], [35]. HUBO2+ has approximate weight of 40 kg and height of 120 cm. It has 6 DOFs in each lower limb, and hence, 12 DOFs in the actuating joints and 6 DOFs in the floating base joints should be controlled to imitate human walking. We used WiFi for communication between the robot and data acquisition device, and controller area networks for internal communications for the robot and data acquisition device" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0001869_elan.202060313-Figure3-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0001869_elan.202060313-Figure3-1.png", + "caption": "Fig. 3. Schematic illustrations of (A) the individual electrified cell culture well with electrolyte and (B) the 24-well electroculture ware of ECSARA showing a cell culture insert in well A1.", + "texts": [ + " The charge transfer resistance, (RCT), is seen to be ordered according to graphitic carbon showing the least charge transfer resistance followed by glassy carbon, platinized stainless steel, platinum, gold, and titanium. Table 2 shows the effective area, Aeff, and \u03b3 (area correction factor) extracted from the R S plots, RCT (\u03a9) extracted from equivalent circuit analysis of the Nyquist plots (EIS Data) and the normalized RCT /\u03b3 extracted for the various electrode materials. The normalization of the charge transfer resistance to the area correction factor, (RCT /\u03b3 (\u03a9)), gives a corrected evaluation of the relative facility of charge transfer kinetics at the true electrode interface [20]. Figure 3A shows an exploded schematic of a single well outfitted with opposing electrodes sitting within the test electrolyte and Figure 3B shows a schematic of the 24- well electroculture ware of ECSARA. Multiple-scan rate cyclic voltammetry of the GrC and Ti electrodes served to compare both these substrate electrodes within the 24- well electroculture ware system. Electrodes were studied in 0.1 M PBS (pH 7.2) as reference and in 25 mM Fe(II)/ 25 mM Fe(III) in 0.1 M PBS (pH 7.2). The multiplexed 24-well test environment was used to investigate the wellto-well variability when using different candidate electrodes. 3.2.1 Variability Among Graphitic Carbon and Titanium Electrodes from MSRCV MSRCV in 25 mM Fe(II)/25 mM Fe(III) in 0" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0002276_s12541-020-00431-8-Figure1-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0002276_s12541-020-00431-8-Figure1-1.png", + "caption": "Fig. 1 Regions of torsion beam", + "texts": [ + " Beam elements are used to model the torsion beam, which is a part of the CTBA experiencing a large amount of deformation. Rigid bodies are used to model parts where little deformation occurs, such as the trailing arm, knuckle, and spring mount. This method allows models to be created using hardpoints and torsion beam properties even when the shape of the CTBA has not been determined. The torsion beam used in CTBA is divided into a constant region with a fixed cross section shape and transition regions in which the cross-section shape gradually changes, as shown in Fig.\u00a01. To depict this, the constant region is configured as a single beam element. To reflect changes in the shape of the cross section, 5 beam elements are used in each of the left and right transition regions (Fig.\u00a02). For beam element modeling, the Timoshenko beam 6 DOF stiffness matrix in Eqs. (1)\u2013(3) is used in this study. Figure\u00a03 shows the coordinate system of the Timoshenko beam element. In the Timoshenko beam element used in this paper, the warping effect is ignored. (1) \u23a1\u23a2\u23a2\u23a2\u23a2\u23a2\u23a2\u23a3 Fx Fy Fz Tx Ty Tz \u23a4\u23a5\u23a5\u23a5\u23a5\u23a5\u23a5\u23a6 = \u2212 \u23a1 \u23a2\u23a2\u23a2\u23a2\u23a2\u23a2\u23a2\u23a2\u23a2\u23a3 EA L 0 0 0 0 0 0 12EIzz L3(1+Py) 0 0 0 \u22126EIzz L2(1+Py) 0 0 12EIyy L3(1+Pz) 0 6EIyy L2(1+Pz) 0 0 0 0 GJ L 0 0 0 0 6EIyy L2(1+Pz) 0 4EIyy L(1+Pz) 0 0 \u22126EIzz L2(1+Py) 0 0 0 4EIzz L \ufffd 1 + Py \ufffd \u23a4\u23a5\u23a5\u23a5\u23a5\u23a5\u23a5\u23a5\u23a5\u23a5\u23a6 \u23a1\u23a2\u23a2\u23a2\u23a2\u23a2\u23a2\u23a3 dx \u2212 L dy dz x y z \u23a4\u23a5\u23a5\u23a5\u23a5\u23a5\u23a5\u23a6 \u2212 C \u23a1\u23a2\u23a2\u23a2\u23a2\u23a2\u23a2\u23a3 Vx Vy Vz x y z \u23a4\u23a5\u23a5\u23a5\u23a5\u23a5\u23a5\u23a6 Lumped mass elements are placed between the beam elements to create the torsion beam\u2019s mass properties" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0002728_978-3-642-03016-1_3-Figure3.9-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0002728_978-3-642-03016-1_3-Figure3.9-1.png", + "caption": "Fig. 3.9 Ballistic model of a single step with three phases when the model can be approximated by a system with one degree of freedom of predominant mobility in: a. -ankle (A), b. -metatarsal (M), and c. -joints in both legs (MAHA2)", + "texts": [ + " For example, the terminal link may follow a sinusoidal trajectory, while the COM follows a straight-line trajectory. However, the gait analysis confirms that in human gait, the COM follows a trajectory that can be approximated by a superposition of the sinusoids in the sagittal and frontal planes. 3.3 Model of a One-Step Cycle: Passive Phase A standing inverted pendulum of length l with its mass m able to rotate relative to a pin joint, which is an analog of the ankle joint, will simulate the body (Fig. 3.9). Let the supporting foot be bound to the ground until the angle of rotation of the pendulum relative to the vertical line reaches its value j. We define the beginning of a gait cycle as the moment when the COM takes its highest position (a), and the pendulum\u2019s rod is vertical. We will synthesize one step for the model, and will define the step\u2019s completion as the moment when the second leg will also be in the vertical position after advancing the first leg (c). When the second leg reaches its vertical position, the configuration of the model will be identical to those shown in Fig", + " These limitations reflect our initial interest in evaluating the role of the ankle joint in the mechanics of gait, and in developing specific recommendations to the design of the prosthetic ankle unit. Any model, by virtue of being a model, is a substitute for the targeted real object or process by a simpler object or process. The adequacy of the model is its ability to describe the cause\u2013effect relationships of the critical features of the target. The usefulness of the model is validated by the outcomes in solving practical problems. Having said that, how can we assess the model (Fig. 3.9)? In Chap. 4, we will show how the model was applied to the synthesis of a mechanism of the prosthetic ankle joint. In Chap. 5, we will present a biomechanical validation of the use of the actual prostheses, which was developed and manufactured following the current modeling of the gait cycle. We divide the step into three phases: A, M, and MAHA2 depending on the mobility in joints such that during each phase, the model will be a system with one degree of freedom. Mobility in both ankle joints A, and A2 are shown in Fig. 3.9 a, c. Mobility in metatarsal joint of one leg M, is shown in Fig. 3.9 b, and in the hip joint H, in Fig. 3.9c. By definition, that means that during each of these phases, there will be only one variable needed to calculate coordinates of any point of the model. When the model\u2019s configuration changes, as it passes from one phase to another, there will still be one variable to follow and one equation to generate. We will use this approach to extract the most critical characteristics of biped human gait, which can be utilized in lower limb prosthetics design. 3.3.1 The first phase is the pendulum\u2019s rotation relative to the ankle joint A of the first leg (phase A)", + " Once the ankle becomes fixed, the model\u2019s motion continues with a rotation around the metatarsal joint as shown in the Working Model movie in c. For the rotation relative to the metatarsal joint, the software calculates a new trajectory 2 of the COM, which is positioned higher when compared with trajectory 1, which the model would follow if the ankle joint were not fixed by the stop 4 (see. a, b). The advantage of the trajectory of the COM going higher due to transfer of rotation from ankle to the metatarsal joint has been demonstrated earlier with (3.9). The equation describes the motion of a more simplistic model (Fig. 3.9) with the weightless pendulum rod and foot. The Working Model calculates the decreasing velocity of the current model\u2019s COM (c) and shows the moment when the velocity becomes zero, and changes its direction during reverse rotation (d). Here is an illustration of the importance of the transfer of rotation in the ankle to the metatarsal joint for maintaining dynamic balance. We also see an illustration of the important role of the stop 4 that provides for a passive switch of rotation from the ankle to the metatarsal joint", + " The resistance increases nonlinearly and rapidly until the end of the dorsiflexion period, resulting in deceleration of articulation in the ankle and lifting of the heel when the articulation in the metatarsal joints begins. The maximum magnitude for the moment around the talocrural joint (articulation of the ankle in the sagittal plane) averages from 80 to 120 Nm. A similar concave nonlinear pattern is seen in the subtalar (talocalcaneal) joint (frontal articulation) with the maximal moment of resistance of 23\u201325 Nm (Scott and Winter 1991). The period of articulation in the ankle without a resistive moment from the calf muscles has been simulated by phase A of the model (Fig. 3.9), and can be viewed also in the Working Model simulation (Fig. 3.10). During phase M of the model when rotation occurs in the metatarsal joint due to limiting of the mobility in the ankle, the heel lifts by inertia. That suggests that a prosthetic foot and ankle unit can also produce the heel lift without external power if the \u201cfree-limited mobility\u201d sequence is included in the prosthesis\u2019 design (Pitkin 1994). We assumed the limiting of the mobility in the model\u2019s ankle to be energy-free, since we wanted to learn about the possibility of replicating that mechanism with a passive stop in the prosthetic joint", + " This activity (eccentric contraction of triceps surae, gastrocnemius, and the intrinsic foot flexors) takes place at the end of the dorsiflexion period, when the heel has to be raised. The dorsiflexion period is followed by rapid plantar flexion (concentric contraction), which occurs during late stance until \u201ctoe-off.\u201d It is well documented that during plantar flexion, foot flexor muscles generate much less power than at the end of dorsiflexion period (Perry 1992). Following their activity that limits the mobility in ankle during the dorsiflexion period of gait (phases A and M of the model (Fig. 3.9)), plantar flexor muscles generate a rotational moment in the ankle, plantarflexing the foot when the second leg is advancing and is positioning on the ground. As we saw in Chap. 2, the value of the moment for plantarflexion is significantly smaller than the value required for limiting dorsiflexion and locking the ankle prior to heel lift. To see why there is no need for a large plantarflexion moment to complete the step, we will continue to consider the ballistic model (Fig. 3.9) in its active phase c. 3.4 A Model of a One-Step Cycle 3.4.1 Active Phase \u201cMAHA 2 \u201d We are returning to configuration c of the model (Fig. 3.9), which corresponds to the period when the calf muscles and foot intrinsic flexors generate the rotational moment for the foot plantarflexion. This is a simulation of the second plantarflexion that occurs at the end of the stance phase. The ankle moment has positive sign for a viewer of the picture (Fig. 3.9). It is in contrast with the negative moment generated by the m tibialis anterior and for a first plantarflexion occurring immediately after the heel-on at the beginning of stance. We will calculate the minimal value of the propulsive moment in the ankle of the back leg, which will complete the step. For that, we will assume that the trailing leg is not flexing the knee, as in the experiment of intentional propulsion described in Sect. 3.1.2.2 (Fig. 3.3). A hypothesis, which we will be validating, is that the moment for propulsion decreases with increased velocity of the model\u2019s COM. A transfer from phase M to the final phase called MAHA2 is associated with the third change of the model\u2019s configuration. During phase MAHA2 , the model will be a four-bar linkage with rotation permitted relative to the metatarsal joint M and ankle joint A of the first leg, the hip joint H and ankle joint A2 of the second (advanced) leg (Fig. 3.9). One of the limitations of the model is that we will not consider energy costs for changing the model\u2019s configuration. The previous two phases were ballistic, i.e., there was no active control over the model\u2019s motion. Now, we request inclusion of an actuator to be placed at the ankle joint A. The actuator\u2019s task is to generate the minimal rotational moment needed to bring the advancing second leg to a vertical position, with the velocity of the COM equal to zero, which completes the step cycle", + " One may characterize the decrease of the needed work as an advantage due to inertia of the body during walking. Falling down, which begins at stage A, is the origin of the velocity n2, as well as velocity n1. The work A(m) cannot, however, be equal to zero if we want a nonzero step of the advancing leg. Indeed, if we assume A(m) = 0, 2 m(n2 cos l)2 mgD3 = , and since D3 = D1 \u2212 D2, we will get cosg = 1. That indicates that the angle g will be equal to zero, or that the vector n2 will be parallel to the directional vector d r (Fig. 3.9). The latter means that virtual rotation of the model comes to joint M instead of A2, which might be possible only if the length of the advancing step is zero. At the same time, cos g \u2260 0, since in that case the vector n2 would be perpendicular to the directional vector d r bringing the step length to infinity. We see here that height l of the model, foot dimensions, and step length of the model in phase MAHA2 have to meet conditions of geometrical compliance to assure adequacy of the model to human biped gait (Pitkin 1997b)" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0003275_aim.2015.7222530-Figure11-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0003275_aim.2015.7222530-Figure11-1.png", + "caption": "Figure 11. Configuration4: 4 segments placed in the region 00 < y < 3600 in the best possible configuration.", + "texts": [], + "surrounding_texts": [ + "A 3-channel gauss meter (Lakeshore-Model 460) was used to measure the magnetic flux density. A torque gauge (HTG2-40 made by IMADA) with its respective torque sensor held the IPM at the centre of the system. The IPM was connected to the torque sensor via a green plastic connector that was prototyped using a 3D printer. The torque sensor and the probe tip of the gauss meter were mounted on plastic holders which were also fabricated with a 3D printer. Both the torque sensor and the probe tip of the gauss meter can be moved along the X and Z axes and the arrays of magnets can only be moved along the Y axis. These displacements are controlled by a micromanipulation system constructed of XYZ stages as shown in Fig. 13." + ] + }, + { + "image_filename": "designv11_71_0002656_jahs.64.042008-Figure1-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0002656_jahs.64.042008-Figure1-1.png", + "caption": "Fig. 1. Helicopter (H), load (L), and earth (E) fixed coordinate systems.", + "texts": [ + " Finally, the conclusions of this work are summarized. The external load model used is that of an 8 ft \u00d7 6 ft \u00d7 6 ft CONEX cargo container with two rear-mounted stabilization fins. The fins prevent load rotation but do not guarantee load stability throughout the helicopter flight envelope. This model was selected due to its extensive use in the studies mentioned previously and to extend its stability from 60 kt up to 130 kt. The two fins are inclined at 33 deg relative to the box side faces, trailing edge out (Fig. 1). The model also assumes a total weight of 2489 lb, which represents an empty container with the four sling cables. For this study, the load center of gravity was set to be 0.3 ft aft of the CONEX geometric center. This makes the load unstable at an airspeed of 100 kt, the airspeed selected as the target airspeed for load stabilization in the current research. The dynamic model described above has been thoroughly validated using dedicated wind tunnel tests and flight tests. The aerodynamic model of the fin-stabilized load uses static aerodynamic forces and moment coefficients measured in wind tunnel testing for the complete load, fins included", + " Coupled helicopter external load system For the studied configuration, it was assumed that the load was connected to the helicopter cargo hook by a swivel, which enabled free yaw rotations of the load with a zero resisting friction moment. The new coupled system created in this way was studied by using its state-space representation, obtained by combining the load and the helicopter model (including the flight control system): \u23a1 \u23a3x\u0307 H x\u0307L x\u0307C \u23a4 \u23a6 = \u23a1 \u23a3F H (x H , xL, xc) FL (x H , xL) FC (x H , xC , u) \u23a4 \u23a6 (9) In the equations above, the functions F H , FL , and FC are the corresponding state vector functions that describe the helicopter, the load dynamics, and the DI controller, respectively. Figure 1 presents the coupled system and the coordinate systems used. The load relative cable angles were calculated by obtaining the distance from the cargo hook to the load center of mass, pL , using the earth-to-helicopter coordinate transform matrix TEH, respectively: pL = TEH(r L \u2212 r H ) \u2212 rCH = [xrp, yrp, zrp]T (10) The values of the relative cable angles were obtained assuming the transformation from (H) to (L), TEH, follows the order pitch (\u03b8C), roll (\u03c6C), and yaw (\u03c8C). Using this order, the transformation matrix is derived" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0002203_icem49940.2020.9271015-Figure1-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0002203_icem49940.2020.9271015-Figure1-1.png", + "caption": "Fig. 1. Axial cross section of one pole pitch of the 2D FEA Prius 2004 model with a total of 48 slots and 8 poles, at 1000 rpm and 250 A rms.", + "texts": [ + " thermal transient mode, not only a transient thermal LPN model is needed, but also the machine losses in different machine parts as functions of motor torque and speed. To account for its spatial distribution, the losses are determined in the following parts separately: \u2022 Lamination core: stator yoke, stator teeth, rotor yoke \u2022 Winding: active part, end-winding part \u2022 Magnets \u2022 Bearings All losses except the bearing losses, have been calculated with Finite Element Analysis (FEA) using a 2D Maxwell model, shown in Fig. 1, fed with sinusoidal current excitation, based on [10] and [6]. Magnet losses are calculated from the 978-1-7281-9945-0/20/$31.00 \u00a92020 IEEE 860 Authorized licensed use limited to: University of Prince Edward Island. Downloaded on May 17,2021 at 03:11:04 UTC from IEEE Xplore. Restrictions apply. TABLE I MAGNET DATA USED IN 2D FEA 20\u25e6C 70\u25e6C 120\u25e6C Coercivity, Hc 920 kA/m 759 kA/m 598 kA/m Rel. permeability, \u03bcr 1.03 1.246 1.463 Remanenet flux, Br 1.191 T 1.189 T 1.099 T El. conductivity, \u03c3PM 667 000 S/m 660 396 S/m 653 922 S/m induced eddy currents by defining the magnets as non-excited coils" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0002989_978-3-642-16214-5_9-Figure9.44-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0002989_978-3-642-16214-5_9-Figure9.44-1.png", + "caption": "Fig. 9.44. Conventional arrangement: double-sided lined plates (DSP = Double-Sided Plates). 1 Inner plate, double-sided lining; 2 outer plate", + "texts": [ + " Design measures to compensate the rotational pressure include: \u2022 stronger return spring: inexpensive, but reduces the effective shifting pressure, \u2022 ball valve for clutch discharge: low effort, but unsuitable for controlled clutches and overlapping shiftings, since the function is influenced by speed and oil temperature, \u2022 spring-controlled clutch discharge valve: relatively high effort, unsuitable for controlled clutches and overlapping shiftings, since the function is influenced by speed and oil temperature, \u2022 shift pressure-controlled discharge valve: functionally reliable, but very high effort, also unsuitable for controlled clutches and overlapping gearshifting, \u2022 pressure compensation chamber (Figure 9.41, rotational pressure compensation B 2 and E 11): requires a lot of installation space, yet always functions and can also be used for controlled clutches and overlapping shiftings. As opposed to conventional lined plates with double-sided friction lining (DSP = Double-Sided Plates, Figure 9.44), Single-Sided Plates (SSP) have only one friction lining alternating between the inner and outer plates (Figure 9.45). By increasing the available volume of steel for thermal storage, the load capacity of the clutch can be increased at equal installation space. However, SSPs are more expensive than DSPs, they increase the danger of assembly faults and tend to wobble at high rotational speeds due to their low weight. In the following, a few questions of detail concerning multi-plate clutches will be addressed" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0001304_j.matpr.2020.05.658-Figure1-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0001304_j.matpr.2020.05.658-Figure1-1.png", + "caption": "Fig. 1. Parts of the rocke", + "texts": [ + "658 nism is preferred over gear mechanism because with the help of motor mechanism, the overall power provided to the different wheels would be equal and no wheel in the six wheel mechanism would lag in the power transmission. Moreover, maneuvering it over different terrains require different powers. 3. Methodology 3.1. Design requirements and modelling The wheel chair using the rocker bogie mechanismmajorly consists of 4 different parts over which the motion of the system is dependent. These are the longer arm, shorter arm, main arm and the wheels and it is shown in Fig. 1. 3.1.1. Longer arm The longer arm is the bigger arm of the system. It is placed in the bogie part of the system. r bogie mechanism. uctural analysis of wheel chair using a rocker bogie mechanism, Materials The main arm is one which connects the two opposite sets of rocker and bogie and combines to form a four arm arrangement. The dimensions of the longer arm, shorter arm and main arm are taken according to a scale model that is designed. The dimension of longer arm is 400 units 282 units with an inclination of 90 degrees" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0003788_detc2014-34213-Figure19-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0003788_detc2014-34213-Figure19-1.png", + "caption": "Fig. 19 Eight-bar linkages with input joints A0, A and C0", + "texts": [ + " 18(a) or a two-DOF seven-bar linkage with the input joint B0 held as shown in Fig. 18(b). Thus, for the Stephenson six-bar linkages or the seven-bar linkage, the singularities happen when the links AE, B0B and CD intersect at a common point. Therefore, the three-DOF eight-bar linkage is at the singular positions. If the input joints do not contain one of the joints B, F and B0, such as the input joints given through A0, A and C0, the eight-bar linkage can be decomposed into a single-DOF six-bar linkage with the input joints A0 and C0 held as shown in Fig. 19(a) or a two-DOF seven-bar linkage with the input joint A0 held shown in Fig. 19(b). Thus, for the single six-bar linkages or the seven-bar linkage, the singularities happen when links A0E, B0B and CD intersect at a common point. Therefore, the three-DOF eight-bar linkage is at the singular positions. The proposed method can be also applied to three-DOF planar linkages with prismatic joints. Based on the criteria, a three-DOF planar linkage with prismatic joints can also be decomposed into a corresponding two-DOF planar linkage and an additional input joint or a corresponding single-DOF planar linkage and two additional input joints" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0001691_j.oceaneng.2020.107700-Figure2-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0001691_j.oceaneng.2020.107700-Figure2-1.png", + "caption": "Fig. 2. The deployment of the thrusters (propellers) on the kth module.", + "texts": [ + " Then the connector force vector fh that acts on the mass centers of modules for the floating structure is expressed as fh = \u2212 DT fh (8) The motion equation of the modularized floating structure Eq. (1) changes to a differential algebraic equation as follow Mx\u0308 + (R + S)x = fw \u2212 DT fh + fc D\u22c5x = 0 (9) To implement the control process, the thrusters (propellers) are mounted on the bottom of each module identically. Vector fc in Eq. (1) is the control force vector acting on the mass centers of modules of the floating structure. For the kth module, the deployment of the thrusters (propellers) is shown in Fig. 2. To control the heave, roll and pitch motions of the floating structure, thrusters (propellers that produce vertical forces) are mounted in the bottom of each module. The more propellers are, the less thrust required from each propeller, the better reliability obtained for control. On the other hand, larger number of propellers may have more costs. Here, we Case 1 Case 2 Case 3 dominant frequency (rad/s) 0.618 1.257 0.3\u20131.5 wave angle (degree) 0\u2218 45\u2218 0\u2218 S. Xia et al. Ocean Engineering 216 (2020) 107700 do not discuss the optimal option on the number of propellers, which really depends on particular engineering applications. In this paper, we choose 12 propellers for control. Three propellers as a sub-group are placed in each corner of a module, as shown in Fig. 2. The thrusts of the propellers on the kth module are expressed as uk = [uk,1,\u22ef, uk,12] T that generates a control force vector on the center of the kth module of the floating structure. When the floating structure is static, the control thrust of the jth propeller acts on the kth module at the mounting point of ck,j,u0 in the local frame. The mass center of the kth module is at the point ck,c0 in the global frame. When the system moves, the motion of the kth module isxk = [xk, yk, zk, \u03b1k, \u03b2k, \u03b3k] T " + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0003715_cdc.2014.7039544-Figure2-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0003715_cdc.2014.7039544-Figure2-1.png", + "caption": "Fig. 2. AMB operating principle.", + "texts": [ + " Furthermore, in order to reproduce the transmission delays observed in remote turbomachinery applications, a time delay can be added to the input of the AMB system. Active magnetic bearings (AMBs) use electromagnetic forces to levitate the rotor within the bearing clearance. This allows the rotating parts of the machine to move freely in the direction of rotation with minimum parasitic losses. This characteristic makes AMBs attractive for high-speed rotating machines, such as compressors, gas turbines and artificial heart pumps [16]. The basic AMB operating principle in one control axis is illustrated in Fig. 2. The actuators of the AMB system is a pair of electromagnets acting on opposite sides of the target rotor. Each electromagnet operates on a fixed bias current Ib, perturbed by a control current i. A rotor levitation controller calculates a control command to the actuators based on the feedback signal from the sensors measuring the rotor displacements. The command signal from the levitation controller is forwarded to the current amplifiers powering the coils of the AMB actuators, thus closing the control loop" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0002313_10402004.2020.1856991-Figure2-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0002313_10402004.2020.1856991-Figure2-1.png", + "caption": "Fig. 2. Schematic of four kinds of representative film gap cases", + "texts": [ + " Under normal operating conditions, saucer-shaped warping deformation plays a major role in almost all the forms of disk deformation due to temperature variation and constraints [15]. In addition, misalignment between the driving disk and the driven disk or surface irregularity of the disk may lead to the difference in oil pressure or temperature. By studying all the possible shape of the disks generated by the above reasons, four kinds of representative film gap cases, i. e. divergent, convergent, convergent-divergent and divergent-convergent, are added to the soft-start dynamic model, as shown in Fig. 2. In general, the friction surfaces along the radial direction are assumed to be of a linear change regardless of actual cambered surfaces. This is due to the fact that the gradient change of the film thickness is much larger than that of the radial distance. Fig. 2(a) means the divergent film gap case in which d is the divergent angle and the minimum initial film thickness minh is located at the inner radius of the disk. Similarly, as shown in Fig. 2(b), c is the convergent angle and the minimum initial film thickness minh is located at the outer radius of the disk. Fig. 2(c) means that the film thickness decreases first and then increases along r direction. minh is located at the middle part of the area through which the oil film flows, i. e. 0r r . 1cd and 2cd are the convergent angle and the divergent angle, respectively. The divergent-convergent film gap case that is similar to the convergent-divergent film gap case is shown in Fig. 2(d), which is not being covered here. Acc ep te d M an us cr ipt 6 2.1 Film lubrication model In the HVD lubrication analysis, the film thickness across the film (in z direction) has a much smaller dimension than either the disk circumference (in \u03b8 direction) or the radial length (in r direction). As compared to axial velocity zv , radial velocity rv , and tangential velocity, v are predominant velocities since fluid flows through the disks\u2019 surfaces. According to J. Z. Cui et al. [18], for iso-viscous lubricants with the above assumptions, by neglecting the inertia and body force in the very thin lubricant film, the modified averaged Reynolds equation in cylindrical coordinates can be expressed as follows t h r r p rhr r th ir 12)]([ 3 (1) where is the dynamic viscosity of the oil film" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0000175_012073-Figure1-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0000175_012073-Figure1-1.png", + "caption": "Figure 1. Rotor with flexible working bodies: (a) projections of the material points of the working body in the XYY plane; (b) projections of material points of the working body in the XOZ plane.", + "texts": [ + " The rotor was developed at Voronezh State Forestry University and is a flywheel weighing 10 kg with a diagonal of 0.3 m to which chain links with a weight of 0.35 kg each are attached. The rotor through the flange is attached to the shaft of the hydraulic motor. Brand hydraulic motor MG2.28 / 32 manufacturer Shakhty plant GIDROPRIVOD. The working volume is 28 cm3, the rotational speed is 25-4800 rpm, the torque is 83 Nm, the nominal pressure of the working fluid is 20 MPa. In the process of movement, the flexible working body simultaneously performs rotational and portable translational motion (figure 1). Forestry 2019 IOP Conf. Series: Earth and Environmental Science 392 (2019) 012073 IOP Publishing doi:10.1088/1755-1315/392/1/012073 Since the angular velocity is an order of magnitude greater than the translational velocity, the latter one can be neglected. Given the possibility of moving the material points of the working body in both the horizontal and vertical planes, we obtain: sin( ); cos( ); sin( ), k k k k k k k k kb x r t y r t z r t (1) where xk, yk, zk \u2013 the coordinates of the k-th point of the flexible working body, m; rk \u2013 the length of the working body from the point of attachment to its k-th point, m; k \u2013 the angular velocity of the k-th point of the flexible working body in the horizontal plane, s-1; k_b \u2013 the angular velocity of the deviation of the k-th point of the flexible working body in the vertical plane of the axis, s-1. We adopt a flexible working body, as a system of material points that can move relative to each other (figure 1). Since we consider the work of a flexible working body in three-dimensional space, the coordinates that determine its position are found as projections of points on the corresponding axes: 1 1 2 2 1 1 _ sin( ) cos( ); cos( ) cos( ); ( ) ( ) sin( ). k k k k k k k k k k k k k k k k b x r t t t y r t t t z x x y y t (2) Add links that allow each point to rotate relative to each other, but do not allow to change the length of the working body. To do this, we calculate the length of the working body, on the basis of the system (2), 2 2 2 1 1 1( ) ( ) ( ) " + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0001775_0278364920955242-Figure5-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0001775_0278364920955242-Figure5-1.png", + "caption": "Fig. 5. The operational space theory postulates a projection of the articulated-body dynamics to dynamics at the task point (shown as a yellow dot). This projection yields the effective mass at the task point. (a) Belted ellipsoid of effective mass of the body at the task point traced by varying direction u\u0302 and plotting the effective mass in that direction. (b) Belted ellipsoid of effective mass for two different configurations of the manipulator. Note that the principle directions as well as the lengths of the minor and major axes change drastically with the configuration.", + "texts": [ + " This is illustrated for the KUKA IIWA manipulator in Figure 4 where the task motion at the end-effector has six DOFs and the manipulator has seven joints. The null-space motion describes additional motion of the manipulator keeping the task motion unchanged. For a particular state of the system given by q, _q\u00f0 \u00de, this null-space motion is unique, and the task-space forces do not create a motion in the null space. Inspecting the task-space inertia matrix reveals important characteristics of the dynamics at the end-effector (Khatib, 1995). One particular property of interest is the effective mass at the end-effector. As shown in Figure 5, the effective mass of the manipulator at the end-effector in direction u\u0302 is defined to be the force required to be applied in this direction such that the resulting acceleration at the end-effector has a unit component in the u\u0302 direction. The effect of other forces such as gravity and Coriolis/centrifugal forces is not considered, effectively treating the manipulator as static in a gravity-free environment. Formally, the effective mass at the end-effector in u\u0302 direction is derived to be mu = 1 u\u0302TL 1 v u\u0302 \u00f04\u00de where L 1 v is the task-space inverse-inertia matrix for an end-effector positioning task. The task velocity for a positioning task is given by y = Jv _q where Jv is the linearvelocity Jacobian for the manipulator at its end-effector. The belted ellipsoids shown in Figure 5(b) represent the contour traced out by the effective mass value as u\u0302 is varied while keeping both the chosen point on the end-effector, and the configuration of the manipulator constant. In the same lines as (4), an effective inertia in u\u0302 direction can be defined with respect to an orientating task at the endeffector. In the following subsection, we use the previous results to derive the governing equations for a system of articulated bodies in a multi-point contact scenario. Consider an articulated body B in contact with a fixed surface at time t as shown in Figure 6", + " The operational space theory was developed by Khatib (1987) as a framework for the unified force and motion control of redundant robot manipulators. It introduced the principle of projected task space dynamics that is at the cornerstone of whole-body control for many-DOF robots today such as humanoids (Khatib et al., 2004). In this framework for robots, or articulated bodies in general, a clear physical interpretation is available for the inertial dynamics in the task space. The effective mass and rotational inertia can even be visualized at the chosen task point, as illustrated in Figure 5. For systems of free rigid bodies, such a physical intuition is readily available in reference to the dynamics for each body at its center of mass. However, the multi-body dynamics theory provides no such physical intuition for articulated-rigid-body systems, where the canonical approach is to resolve it as free rigid bodies with joint reaction forces. By applying the operational space theory to articulated-body systems in contact, we are able to gain the same level of physical intuition. As shown in the following, we can regard collisions between articulated rigid bodies fundamentally as collisions between equivalent free rigid bodies, given the projected dynamics in the contact space" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0000327_j.mechmachtheory.2019.103729-Figure14-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0000327_j.mechmachtheory.2019.103729-Figure14-1.png", + "caption": "Fig. 14. Rubik\u2019s Cube unit after equivalent the generalized kinematic pair.", + "texts": [ + " Therefore, when the three three-loop mechanisms are closed, each of them can be equivalent to a generalized serial chain with the same number and properties of the DOF. The DOF of the generalized kinematic pair of the generalized serial chain is that of the 3-rank component relative to the fixed platform, that is, the DOF of the highest rank of component in the three-loop mechanism. Obviously, the equivalent mechanism is a 3-PPPRRR parallel mechanism with a 3-rank component which works as the moving platform, as shown in Fig. 14 . The mechanisms of L(1) + L (2) + L (4) and L(1) + L (3) + L (5) are compared with L(2) + L (3) + L (6) which is the last section analysis have the same mechanisms, so it has six DOFs. The equivalent of the three branches has no common constraint, and its order is d = 6 . No redundant constraint when the link is closed, so v = 0 . In this case, the mechanical DOF is M = d(n \u2212 g \u2212 1) + g \u2211 i =1 f i + v = 6 \u00d7 (17 \u2212 18 \u2212 1) + 18 + 0 = 6 (17) At this point, it can be said that the DOF of the 3-rank component as the moving platform in the unit mechanism is 6" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0001419_jemmr.20.00119-Figure4-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0001419_jemmr.20.00119-Figure4-1.png", + "caption": "Figure 4. (a) Computer image of the data obtained using a Br\u00fcel & Kj\u00e6r data collection device; (b) test set-up; (c) specimens", + "texts": [ + " The RT Photon Pro software16 was used to interpret rials Research 2020.9:805-811. the data obtained from the tensile tests, and a Br\u00fcel & Kj\u00e6r Photon+ dynamic signal analyzer device17 was used for the collection of data. The experimental set-up involved hanging test specimens from a cotton thread to allow free movement, the damping of the oscillation by the specimen after a hammer strike and the placement of the cables of the acceleration sensors in a position such that they would not affect the results during oscillation (Figure 4). The test specimens were tested in three different ways. For each test specimen, the acceleration sensor was first placed at the farthest end of the specimen at a location not overlapping with the knotting point (Figure 5). Force was applied by way of a hammer strike at a point on the opposite side of the sensor, close to the central point of the specimen but not overlapping with the knotting point, and data were collected. In the second test, the acceleration sensor was fixed to a point one-quarter along the total length of the test specimen" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0000825_s41315-019-00114-2-Figure5-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0000825_s41315-019-00114-2-Figure5-1.png", + "caption": "Fig. 5 2WD Smart Robot Car Chassis Kit Fig. 6 HC-SR04 ultrasonic sensor", + "texts": [ + " We use ultrasonic sensor to measure the free distance from the robot to avoid obstacle and collision, and override motor control if the distance is below a configured threshold value. The ultrasonic sensor hardware provides the lowest and safety layer of the robot control. The mobile phone is played as a main controller of the robot. It is responsible for sending image and video signals to host computer via Wi-Fi connections, the phone can also connect to micro-controller through Bluetooth communications. The main body of the Arduino robot is a 2WD Smart Robot Car Chassis Kit, as shown with Fig.\u00a05. The mobile platform uses two DC motor gear box (1:48) to drive. The two wheels on the chassis are driven separately, so the robot can turn around and go in any direction. The wheels have silicone tires and measure 60\u00a0mm (2.36\u2033) in diameter. The Arduino robot can avoid obstacles with a distance measuring sensor: HC-SR04 ultrasonic sensor. This low-cost sensor provides less than 400\u00a0cm of remote measurement ability with accuracy up to 3\u00a0mm. Every HC-SR04 module includes a control circuit, an ultrasonic transmitter and a receiver" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0000058_icmae.2019.8881011-Figure7-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0000058_icmae.2019.8881011-Figure7-1.png", + "caption": "Figure 7. Position No 1", + "texts": [ + " ANALYSIS OF RESULTS AND DISCUSSION For the verification of the results, an algorithm with the calculations described above was performed and validated using MatLab\u00ae SimMechanics tool, considering the robot dimensions described in Table III, the simulation in figure 5 and the controller constant K of 10. (6) (7) (8) (9) The initial position of the robot is specified as illustrated in figure 6, with angle at the joints equal to zero. Different trajectories are plotted in order to verify the calculations raised in the inverse kinematics algorithm, starting with the values shown in Table IV, and the results obtained in figure 7. For the verification of the direct kinematics algorithm, the values from Table V and the results obtained in figure 8 are stipulated. Finally, a helical trajectory is traced to verify the behavior of the platform in the face of variations in time. From what is obtained the behavior of figure 9 in angular positions, figure 10 for positions in X, Y and Z, and figure 11 for positions in. . Figure 11 shows a 5-degree oscillation due to the kinematics propagation error present in the platform, therefore a more robust control is necessary, which implies a higher computational cost" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0001059_s42417-020-00207-4-Figure1-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0001059_s42417-020-00207-4-Figure1-1.png", + "caption": "Fig. 1 Air-ring bearing", + "texts": [ + "in 1 Department of\u00a0Mechanical Engineering, Coimbatore Institute of\u00a0Technology, Coimbatore\u00a0641\u00a0014, India 1 3 F Absolute dynamic load-carrying capacity (N) f Incremental dynamic load-carrying capacity (N) g Acceleration due to gravity (m\u00a0s\u22122) h Axial length (m) K Stiffness coefficient (N\u00a0m\u22121) k Thermal conductivity (W\u00a0m\u22121\u00a0K\u22121) L Bearing length (m) m Mass (kg) N Number of time steps; rotational speed of the rotor (RPM) n Number of feed-holes per row (see Table\u00a01); an index P Static load (N) p Static pressure (Pa) R Gas constant (J\u00a0kg\u22121\u00a0K\u22121) r Radial coordinate (m); number of rows of feedholes (see Table\u00a01) S Absolute displacement (m) S\u0307 Absolute velocity (m\u00a0s\u22121) s Incremental displacement (m) s\u0307 Incremental velocity (m\u00a0s\u22121) T Temperature (K); time period of vibration (s) t Time (s) U A target quantity of interest u Fluid velocity component (m\u00a0s\u22121) x A coordinate (m) y A coordinate (m) z A coordinate (m) \u03b1 Attitude angle (rad) \u03b2 Angular position of baffle (rad) (see Fig.\u00a01) \u03b3 Angular width of baffle (rad) (see Fig.\u00a01) \u03b4 Radial distance (across film thickness) measured from bushing surface (m) (= e/c) Eccentricity ratio \u03b8 Angle coordinate (rad) \u03bb Rate of reduction of truncation error \u03bc Dynamic viscosity (kg\u00a0m\u22121\u00a0s\u22121) \u03bd Angular frequency (rad\u00a0s\u22121) \u03c1 Density (kg\u00a0m\u22123) \u03c9 Angular velocity (rad\u00a0s\u22121) AB Air-bearing AR Air-ring B Bushing dm Dynamic mesh i Grid-level J Journal max Maximum v At constant volume r Radial rel Relative s Supply z Axial \u03b8 Circumferential Superscripts i An index j An index x x-Component y y-Component F Absolute dynamic load-carrying capacity (N) f Incremental dynamic load-carrying capacity (N) r Position vector (m) S Absolute displacement (m) \u0307 Absolute velocity (m\u00a0s\u22121) s Incremental displacement (m) \u0307 Incremental velocity (m\u00a0s\u22121) \u0308 Acceleration (m\u00a0s\u22122) u Velocity (m\u00a0s\u22121); orthonormal vector v Orthogonal vector w A vector \u03c4 Viscous stress tensor (N\u00a0m\u22122) \u03c9 Angular velocity vector (rad\u00a0s\u22121) Air bearings are suitable for supporting high-speed, lightly loaded rotors for efficient transmission of energy", + " [15] described a test bench setup for investigating SE whirl of a high-speed rotor with air bearings floating on O-rings. However, the selection of an elastomeric material having suitable dynamic coefficients (DC), i.e., stiffness coefficients and damping coefficients, requires a trial-and-error approach. As an alternative to elastomeric support for the bushing, several novel designs of aerostatic bearings such as Sixsmith, dual gas film or externally-pressurized journal air-bearing (AB) with an air-ring (AR), or air-ring bearing (ARB) (see Fig.\u00a01), and tangential orifices were proposed [2]. Brzeski and Kazimierski [16] explained the operating principle of ARB. Kazimierski and Trojnarski [17] presented methods of computing mass flow rates of air through the plain feedholes (FH) and the pocketed FH in ARB. Czo\u0142czy\u0144ski et\u00a0al. [18] analyzed the dynamic response of ARB to a step-change in force. A procedure for determining the DC of ARB using the Gram-Schmidt orthogonalization process is available [3]. The range of values of DC for an ARB that would enable dynamic stability of a symmetrically loaded rigid rotor was estimated in Ref", + " [22, 23] presented the analysis of an air bearing system. References [12\u201315] explained that SE vibration could be avoided by introducing an elastic supporting structure between the bearing bushing and the casing, characterized by properly selected DC. An elastic supporting structure could be realized in practice via an externally-pressurized AR [2, 3, 16\u201323]. Based on the previous studies [3, 16\u201323], the following observations are made. (1) Analysis for p-distribution in the AB and AR regions (see Fig.\u00a01) were made by solving RE separately in each region, and thereby their DC were determined separately. (2) The geometric center of the journal (CGJ) was prescribed to follow a simple harmonic motion (SHM) to determine the DC of AB, by fixing the position of the bushing center (CGB). Similarly, by securing the position of CGJ, CGB was allowed to follow an SHM to determine the DC of AR. This procedure is referred to in this paper as a single-degree of freedom (1-DOF) approach. (3) Air mass flow through FH into AR and AB regions were computed using zero-dimensional (algebraic) equations", + " Air at supply pressure ps enters the FH provided in C before it flows through the AR and into the FH in B. It then flows in the AB and is exhausted to the atmosphere through the bearing ends. Figure\u00a02c shows the geometry of FH in B and C. 1 3 For determining the DC, the RARBS shown in Fig.\u00a02a can be replaced by an idealized physical system as shown in Fig.\u00a03 [24]. The following quantities are identified to describe the behavior of the idealized system: (1) absolute displacements of CGJ and CGB [measured relative to the fixed casing center (O) in Fig.\u00a01]: (j) J and (j) B (the superscript j is used to indicate motion of CGJ and CGB along direction j; j can be either x or y as in Fig.\u00a01); (2) absolute velocities: \u0307 (j) J and \u0307(j) B ; and (3) absolute dynamic load-carrying capaci- ties: (ij) J and (ij) B (the superscripts i and j in (ij) J are used to indicate the bearing forces in direction i when CGJ moves along direction j; both i and j can be x or y). When i is x and j is y, (ij) J can be interpreted as a vector of instantaneous values of F(xy) J . Quantities (1), (2), and (3) are obtained via TSS. In addition, system constants mJ and mB, and the stiffness and damping coefficients of AB and AR, i.e., KAB, KAR, CAB, CAR, respectively, are used to characterize the idealized RARBS. The rotor load PJ = mJg\u22152 and the joint-bushing load PB = mBg\u22152 (see Fig.\u00a03). The SSS provides the following: (a) coordinates of the SEP identified at t = 0, i.e., ( S (x) J (0), S (y) J (0) ) of CGJ, and ( S (x) B (0), S (y) B (0) ) of CGB which can be defined in terms of eccentricity (e) and attitude-angle (\u03b1) as ( e0 J , 0 J ) and ( e0 B , 0 B ) , respectively (see Fig.\u00a01); and similarly (b) components of static load-carrying capacity ( FJ (x)(0),FJ (y)(0) ) and ( FB (x)(0),FB (y)(0) ) . At time t = 0, the TSS starts from the SEP of CGJ (OJ) and CGB (OB) as shown in Fig.\u00a04. The incremental displacements of CGJ at any t is s (j) J (t) = S (j) J (t) \u2212 S (j) J (0) . Similar expressions can be set up for CGB measured about SEP. Together, these form the elements of (j) J and (j) B . The TSS consists of two stages: In stage-1, CGJ is simulated to follow an SHM i.e., s(x) J (t) = Asin( t) with s(y) J (t) = 0 ", + " During TSS, B1\u2013B3 are stationary wall boundaries and B4\u2013B7 are moving boundaries, whereas B8\u2013B12 are deforming boundaries. Both SSS and TSS are performed by the algorithm described in Fig.\u00a0 6, using ANSYS FLUENT. Computations of thermo-hydraulic variables are performed with double-precision accuracy. Overall mass and energy conservation through the entire fluid flow domain are verified during the simulations. During SSS, trial values of the positions ( e0 J , 0 J ) and ( e0 B , 0 B ) are assumed (see Fig.\u00a01). At the journal boundary B7 (Fig.\u00a05a), \u03c9 and ( S (x) J (0), S (y) J (0) ) are specified. Using Eqs. (11), (19) and (20) (excluding the transient terms), the p-distributions around journal and bushing are computed by FLUENT and integrated to find ( FJ (x)(0),FJ (y)(0) ) and ( FB (x)(0),FB (y)(0) ) . Since the positions of CGJ and CGB are arbitrary to start with, the fluid forces do not balance PJ and PB. Several trials with different positions of CGJ and CGB are needed to arrive at the correct candidate positions ( e0 J , 0 J ) and ( e0 B , 0 B ) at which FJ (y)(0) \u2245 PJ and FB (y)(0) \u2245 PB . In terms of force components F(R) along, and F(T) perpendicular to the l ine of centers, 0 J \u2245 tan\u22121 |||F (T) J \u2215F (R) J ||| and 0 B \u2245 tan\u22121 |||F (T) B \u2215F (R) B ||| (see Fig.\u00a01). Finding SEP is the most time-consuming part of the 2-DOF solution procedure. Computations of force components are performed to an accuracy of \u00b1 5% to PJ and PB. The SSS provides ( e0 J , 0 J ) , ( e0 B , 0 B ) , ( FJ (x)(0),FJ (y)(0) ) and ( FB (x)(0),FB (y)(0) ) at various values of \u03c9 of the journal. User-defined functions (UDF) are codes used to customize the FLUENT solver: (a) to compute \u03bc using Eq.\u00a0(21); (b) to define boundary conditions on the journal and the bushing during TSS; and (c) to write out the time series of (j) J , (j) B , \u0307 (j) J and \u0307(j) B ", + " The motion of both interior and boundary nodes in a deforming zone is defined using a diffusion-based smoothing process that enables them to absorb the movement of journal and bushing. In this process, udm of the deforming boundary is calculated in FLUENT using \u2207\u2022(\u03b7\u2207udm) = 0 where \u03b7 is a diffusion coefficient ( = \u22120.5 , where \u03b4 is the normalized distance between interior and moving boundary nodes). Fig. 7 a Computation of moving/deforming mesh velocity. b Velocity of moving nodes of journal boundary during stage-1 of TSS During TSS, the elastic properties of baffles (see Fig.\u00a01) and seals (see Fig.\u00a02b) are not considered. It is assumed that these parts would deform due to the radial movement of CGB, without any resistance. A rigid rotor condition is also assumed. Hence the study in this paper is limited to fluid dynamic analysis. The discretization error (or mesh sensitivity) Ei in a quantity U during SSS and TSS at a given computational grid-level i (spatial or temporal) is [32]: where bi is either a reference cell size or \u0394t, C is a constant, \u03bb is the rate of reduction of truncation error, and Uexact is the exact value of U", + " The following inferences are made: (1) K(ij) and C(ij) are measures of vibrational energy that can be stored and dissipated in an ARB. Since Kxx AR > Kxx AB and Kyy AR > K yy AB (see Fig.\u00a0 20), and Cxx AR > Cxx AB and C yy AR > C yy AB (see Fig.\u00a021), the ability of an ARB to store and dissipate vibrational energy improves by the incorporation of an AR with the plain AB. (2) For all (\u03c9, \u03bd), K(xx) AB > 0 and K(yy) AB > 0 (see Fig.\u00a020a) implying a statically stable operating condition, i.e., the CGJ would return to its SEP (OJ, see Fig.\u00a01) when displaced by any disturbance. Also, during dynamic conditions, fluid forces f (xx) AB (t) and f (yy) AB (t) due to Kxx AB and Kyy AB will act on the CGJ in a direction opposite to its displacement from its SEP. The resultant radial force shall pull CGJ towards OJ, insuring stable operation. This is true during both forward whirl and backward whirl of the journal. (3) For all sets of (\u03c9, \u03bd), Kyx AB < 0 and Kxy AB > 0 (see Fig.\u00a020b). K(xy) AB > 0 implies a static stability, whereas K (yx) AB < 0 implies a static instability where any disturbance would displace CGJ away from OJ" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0002569_isas49493.2020.9378867-Figure1-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0002569_isas49493.2020.9378867-Figure1-1.png", + "caption": "Fig. 1. Illustrative details on the two-section soft manipulator with three actuators in each section.", + "texts": [ + " 3) Comparative studies with both conventional modelbased and cutting-edge model-less methods are pre- 978-0-7381-1262-6/20/$31.00 \u00a92020 IEEE 217 20 20 In te rn at io na l S ym po siu m o n Au to no m ou s S ys te m s ( IS AS ) | 9 78 -1 -6 65 4- 18 82 -9 /2 0/ $3 1. 00 \u00a9 20 20 IE EE | D O I: 10 .1 10 9/ IS AS 49 49 3. 20 20 .9 37 88 67 Authorized licensed use limited to: Carleton University. Downloaded on June 17,2021 at 16:15:36 UTC from IEEE Xplore. Restrictions apply. sented. In this paper, we select a kind of soft manipulator as the testing object shown in Fig.1. Based on the work of Webster III et al. [5], we can get the two mappings as shown in Fig.2. We can get the position of the end-effector of the manipulator by making use of the mappings. Variables related with the mappings are depicted in (1) (2) (3) (4). More details about the definition of the mappings can be found in [5]. Section 2 Section 1 End Effector (a) Vertical cross section of the manipulator containing two sections in which section-2 is on the top of section-1. Actuator 3 \u03d5i = arctan( \u221a 3(qi2 + qi3 \u2212 2 \u2217 qi1) 3 \u2217 (qi2 \u2212 qi3) )" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0003291_chicc.2015.7260446-Figure1-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0003291_chicc.2015.7260446-Figure1-1.png", + "caption": "Fig. 1: Three-dimensional relative motion of missile-target", + "texts": [ + " Section 3 describes the design method of the finite-time ISS guidance law. In Section 4, simulation results are provided to verify the effectiveness of the design. Section 5 makes some conclusions. In this section, we first give the dynamics of three-dimensional relative motion of missile-target and then, introduce some basic concepts of finite-time ISS theory. 2.1 Three-dimensional Relative Motion of Missile -target The relative motion of missile-target under the three-dimensional spherical coordinate system , ,r is shown in Fig.1, where , ,re e e are the unit vectors whose direction is the same as , ,r , M and T are separately the centroids of missile and target, r represents the relative distance of missile-target, and , respectively denote the azimuth angle and pitching angle. Suppose that the components of the relative accelerations of missile and target are respectively , ,M M Mr a a a and , ,T T Tr a a a . The relative motion can be expressed as the following differential equations [10]: 2 2 2= cos T Mr r r 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) Remark 1: It can be seen that there is singularity at 0r in the above model" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0002793_9781119288190.ch18-Figure18.3-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0002793_9781119288190.ch18-Figure18.3-1.png", + "caption": "Figure 18.3 Constructional details of an electrodeless discharge lamp.", + "texts": [ + " These lamps are called multi\u2010element lamps. The instruments are provided with a switching system, which puts each lamp into its approximate current range. Multi\u2010element lamps emit very complex spectra. Therefore, only dual or triple element lamps, especially for related elements such as calcium\u2013magnesium, sodium\u2013potassium, and copper\u2013nickel\u2013cobalt, are generally constructed. Electrodeless Discharge Lamp The electrodeless discharge lamp (EDL) is a useful source of atomic line spectra. Its design is shown in Figure\u00a018.3. A small amount of the element or salt of the element for which the source is to be used is sealed inside a quartz bulb containing a few torr of an inert gas such as argon. The lamp contains no electrodes but w9781119288121c018.indd 109 10-12-2019 19:49:14 is energized through a coiled antenna from a radio frequency (RF) or microwave generator operating at 2450 MHz and 200 W power output. When an RF field of sufficient power is applied, the coupled energy will vaporize and excite the atoms inside the bulb into emitting their characteristic spectrum" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0001775_0278364920955242-Figure12-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0001775_0278364920955242-Figure12-1.png", + "caption": "Fig. 12. Fabricated experimental setups consisting of aluminum links and high-density plastic collision surfaces. For the planar setups, acrylic disks mounted on sleeve bearings act as the collision surfaces. For the spatial setup, the colliding spheres are made out of 3D-printed polymer. Contact occurs at a single point between each pair of colliding surfaces.", + "texts": [ + " Instead, if one only considered the effective mass of the rigid body RF (Figure 10(c)), the inertia of the colliding body would be underestimated. Lastly, we note that although the previous inferences were drawn for the multi-point collision scenario, identical inferences can be drawn for the multi-point steady contact scenario regarding the generalized accelerations and change of generalized momenta. To test the accuracy of the CSR model, we devised several setups consisting of pendulums with serial kinematic chain structures as shown schematically in Figure 11. The fabricated setups are shown in Figure 12. In each experiment, a pair of pendulums hangs from a rigid and grounded frame, close enough such that they collide often when swinging. Figure 13 shows a few frames from a video of the pendulums colliding on one of the devised setups. The left pendulum is dropped in frame 1, and collision occurs in frame 3. To ensure that the pendulums are rigid and joint friction is negligible, the links are constructed out of aluminum, and steel precision ball bearings are used at each joint. Unlike real robots, the joints are not actuated", + " Our goal through this experimental study was two-fold: to test model accuracy in predicting the outcome of multi-point collisions in articulated bodies; and to test the validity of the model\u2019s underlying assumption that an instantaneous projection of the system dynamics into the contact space is physically admissible. This requires an experimental setup where the DOFs of the articulated-body system are larger than the dimension of the contact space. We performed experiments on three different test setups of colliding pendulums that satisfy this criterion. Figure 12 shows the fabricated articulated-body test systems as described in the following. Setup TP1. As shown in Figures 11(a) and 12(a), this setup consists of a 1-DOF and a 2-DOF pendulum freely hanging from the grounded support. The pendulums are planar, and collisions occur at a single point between the disks mounted on the pendulums. The stiff disks are lasercut out of acrylic, and mounted on bushings to allow them to rotate freely. This is done so as to allow the disks to roll against each other upon contact and minimize the frictional impulse" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0000463_s12541-019-00282-y-Figure6-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0000463_s12541-019-00282-y-Figure6-1.png", + "caption": "Fig. 6 Top view of the pad\u2019s contact area", + "texts": [ + " Notice that F1 and F2 are the spring and damping forces, and x1 and x2 are the deformed displacements of the spring and damping elements in series. Equation\u00a0(1) can be rewritten as From Eq.\u00a0(2), it is noted that the force generated by pressing the sponge is similar to that of an ideal spring. Figure\u00a05 shows the front view of the mopping pads. The forces can be divided into two components: one is generated between the floor and pad by the weight, as shown in Fig.\u00a05a, and the other generated at the tilting angle \u03b8, as shown in Fig.\u00a05b. The friction force of the pad tilted at angle at point p(r, ) , is illustrated in Fig.\u00a06, where r is the radius of center of the pad and is angle of the pad. From Figs.\u00a05 and 6, the force F generated by the sponge can be determined from zd and zp , which can be expressed as follows: (1)F = F1 + F2, (2)F = k1 ( x1 + x2 ) + c2x\u03072 \u2245 k1 ( x1 + x2 ) . 1 3 where M is the weight, g is the gravitational acceleration, S is the contact area between the pad and the floor, and k is the spring coefficient. The force generated by the sponge can be rewritten as The force model of the floor and rotating mopping pad is derived by limiting the rotating angle to a small value (1\u00b0\u20133\u00b0). Thus, Eqs.\u00a0(3)\u2013(5) can be combined and resulted as Furthermore, the friction force of the pad tilted at angle at point p(r, ) , as shown in Fig.\u00a06, can be expressed as follows: where is the friction force constant between the mopping pad and the floor. The force generated by friction for the total area can be calculated by dividing it into x and y axes as follows: (3)zp = r cos tan , (4)zd = Mg kS , (5)Fn = k ( zd + zp ) . (6)Fn = k ( zd + r cos ) . (7)f = Fn = k ( zd + r cos ) , (8) Fx = r \u222b 0 2 \u222b 0 (fsin )rd dr = r \u222b 0 2 \u222b 0 [ k ( zd + r cos ) sin ] rd dr = 0, As shown in Eqs.\u00a0(8) and (9), when the mopping pads rotate, the force in the y-axis direction is determined by the rotating angle and the friction constant " + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0003328_amm.761.329-Figure2-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0003328_amm.761.329-Figure2-1.png", + "caption": "Fig. 2 Assembly view of FDM nozzle. Fig. 1 Exploded view of FDM model.", + "texts": [], + "surrounding_texts": [ + "The dimensions of the FDM nozzle were measured and a 3D CAD model was reverse engineered. Solidworks CAD software was used. Fig. 1 shows the exploded view of the FDM nozzle and Fiq. 2 shows the assembly view of FDM Nozzle. Table 1 shows the details of each parts of the FDM Nozzle. ANSYS is a CAE software that was used to run the simulation and analysis. The Solidworks file was saved in an IGES format to allow ANSYS to read the file. Static structural analysis was carried out to analyze the nozzle reaction of the force that was experienced. Total deformation of the FDM nozzle, virtual stress tests and the factor of safety (FoS) were obtained in the analysis. Modal analysis was used to analyze the reaction of the nozzle when subjected to the vibration from the ultrasonic transducer. The total deformation of the FDM nozzle and an equivalent stress with the frequency of 20 kHz - 30 kHz and 30 kHz - 40 kHz were studied in this analysis. Table 1 shows the detail of each part of FDM nozzle including the part\u2019s name, quantity, volume mm , the type of material and its density. Fig. 3 shows the equivalent stress of FDM nozzle for frequency 20 kHz to 30 kHz. The highest value is 11.721 MPa and the lowest value is 21.324 MPa. The factor of safety (FoS) for this model was obtained by using this result. The calculations showed that the factor of safety of the model was 20.56, referring to the fact that the nozzle can withstand a frequency range of between 20 kHz to 30 kHz that will be transmitted from the ultrasound transducer. Even though the FoS is high, we observed that bending still took place on the nozzle due to the fact that the part has a thin thickness profile. Our observations also showed that no loose screws were found on the part after being subjected to vibrations. Fig. 4 shows the equivalent stress of FDM nozzle for the range of frequency between 30 kHz to 40 kHz. The highest value is 12.753 MPa and the lowest value is 6.464 MPa. The FoS for this model was found to be 18.8975, referring to the fact that the nozzle can withstand frequencies between 30 kHz to 40 kHz from the ultrasonic transducer. The FoS is higher than 20 kHz to 30 kHz because the ultimate tensile strength is also high. The ultimate tensile strength becomes higher because of the frequency applied is higher. The supposition is that having a higher frequency will result in a lower FoS for FDM nozzle." + ] + }, + { + "image_filename": "designv11_71_0001019_j.precisioneng.2020.04.003-Figure6-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0001019_j.precisioneng.2020.04.003-Figure6-1.png", + "caption": "Fig. 6. CAD model of the 3-V kinematic coupling tested for this paper.", + "texts": [ + " The 3-V kinematic coupling designed for this experiment has adjustable inclination angles which cover a broad practical range from 25\ufffd to 80\ufffd as Fig. 4 shows. Each vee is formed by two cylinders secured with threaded fasteners to a plastic plate, which allows an electrical conductivity test to confirm contact (although it was not used nor very reliable using DC). Five triangles having different angles were machined from plastic sheet and used to set the inclination angle \u03b1 of each vee in 5\ufffd increments, as Fig. 5 demonstrates. Referring to Fig. 6, cylinder pairs that form vees are arranged with their axes in three vertical planes that are tangent to and equally spaced on a 203.2-mm [8-in] circle. If all six inclination angles are set the same, then the coupling is considered geometrically symmetric. The coupling is considered symmetrically loaded if the load is the same in all three planes. The load can be made nonsymmetrical either by shifting a weight off center along the linear pattern of holes or by tilting the base plate with respect to gravity. The former method allows the proportion of weight \u03c9 carried by the front vee in Fig. 6 to range nearly a factor of two from 0.224 to 0.442. The latter method was not used. Note, nonsymmetrical configurations greatly expand the number and flexibility of tests that can be run. Another way to make the coupling nonsymmetrical is by having different coefficients of friction among the vees. Antifriction vees may be formed using, for example, rolling-element, aerostatic or flexure bearings. This approach has been used to achieve highly repeatable support of critical structures such as EUV optics [12]", + " Hale Precision Engineering 64 (2020) 200\u2013209 Supplementary data to this article can be found online at https://doi.org/10.1016/j.precisioneng.2020.04.003. Appendix The equilibrium equation (1) describes a 3-V kinematic coupling with five of six constraints engaged under the following restrictions and assumptions: 1) Three constraint planes, each containing a pair of constraint lines, are tangent to and equally spaced on a base cylinder of radius R, which for this experiment had a vertical axis with respect to gravity, see Fig. 6. 2) The base plane defined by the instantaneous centers for three vee constraints is perpendicular to the base cylinder. 3) Either constraint 1 or 2 as identified in Fig. 7 is engaged and has a CoF \u03bc1-2 and inclination angle \u03b11-2. 4) The remaining four constraints 3\u20136 are engaged and have the same CoF \u03bc3-6 and inclination angle \u03b13-6. 5) The external load including the nesting force is parallel to the base cylinder and distributed such that constraint 1 or 2 carries proportion \u03c9 and constraints 3\u20136 carry 1 \u2013 \u03c9" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0001346_tee.23184-Figure1-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0001346_tee.23184-Figure1-1.png", + "caption": "Fig. 1. A 6/4 PM-SRM with two magnets in stator yoke", + "texts": [ + " The analysis has been carried out under variable speed and torque conditions. A comparison between the PM-SRM and SRM results has been conducted on various aspects. The results give a good opportunity to evaluate the capability of the proposed drive system. It has been demonstrated that the PM-SRM have potential to be used in near future motor drive technology. 2. Operating Principles of PM-SRM Drive 2.1. The PM-SRM The PM-SRM has similar structure with a conventional SRM except permanent magnets (PMs) buried in stator yoke (Fig. 1). The phase voltage can be expresses as vk = Rk ik + d\u03bbk (\u03b8 , ik ) dt (1) where Rk is the winding resistance, i k is the current and \u03bbk is the flux linkage of k th phase; \u03b8 is the rotor position. Phase torque developed in the motor can be found by partial derivation of coenergy with respect to the rotor position: Tk = \u2202W \u2032 k (ik , \u03b8) \u2202\u03b8 ik =const (2) where W \u2019k is the coenergy of phase k th and can be obtained from the partial integration of flux linkage with respect to the phase current: W \u2032 k = \u222b \u03bbk (ik , \u03b8)ik \u03b8=const (3) Assuming magnetic linearity, phase torque can be simplified as follows: Tk = 1 2 i 2 k dLk d\u03b8 \u2212 1 2 \u03d52 m dRgk d\u03b8 + N ik d\u03d5mk d\u03b8 (4) In this equation, \u03c6m is the flux going out from the magnet, \u03c6mk represents the magnet flux coupled with the k th winding; Lk and Rk are inductance of the phase winding and the reluctance seen from the kth winding, respectively; N is the turn number of a phase winding" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0001845_icra40945.2020.9196746-Figure4-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0001845_icra40945.2020.9196746-Figure4-1.png", + "caption": "Fig. 4 Perspective view of displacement-magnification mechanism.", + "texts": [], + "surrounding_texts": [ + "The structure of the developed parallel gripper is shown in Figs. 3\u20136, and the specifications are shown in Table II. Because the total mass of the developed parallel gripper is 1.36 kg and the mass of the commercially available parallel gripper is 0.4 kg, the mass of the added mechanisms is 0.96 kg. Each mechanism is outlined below. Here, the movable portion of the commercially available parallel gripper is referred to as the drive portion. Because each finger mechanism has the same structure, only one finger mechanism will be described. 9989 Authorized licensed use limited to: Auckland University of Technology. Downloaded on October 06,2020 at 06:28:00 UTC from IEEE Xplore. Restrictions apply. (a) Displacement-magnification mechanism The displacement-magnification mechanism consists of two sets of rack-and-pinion systems. The rack-and-pinion system of the first finger mechanism consists of a first rack, a third rack, and Gear A. The rack-and-pinion system of the second finger mechanism consists of a second rack, a fourth rack, and Gear B. The first and second racks are connected to the drive portion. Gear A and Gear B rotate on a rotating shaft provided on the base portion. Both the first and second guide portions have a structure in which two sets of guiderails overlap. Gear A comprises two gears, A1 and A2, with different diameters on the same axis. Gear A1 and Gear A2 are integrally connected to form Gear A. Gear A1 meshes with the first rack and Gear A2 meshes with the third rack. The diameter of Gear A1 is smaller than that of Gear A2. The movement of the first and third racks engaged with Gear A is as follows: The pitch circle diameter of Gear A1 in Gear A is Da, and the pitch circle diameter of Gear A2 is Db. When the drive portion is driven to move the first rack in the horizontal direction, Gear A1 meshes with the rotating first rack and Gear A2 also rotates at the same rotational speed as the rotating shaft. Then, as Gear A2 rotates, the third rack moves in the horizontal direction. At this time, the ratio between the moving distance l1 of the first rack and the moving distance l3 of the third rack can be derived in advance. When Gears A1 and A2 make one rotation, the moving distance l1 of the first rack becomes \u03c0Da and the moving distance l3 of the third rack is \u03c0Db. In other words, the value obtained by multiplying the diameter ratio of gears A1 and A2 by the movement amount l1 of the first rack is the movement amount l3 of the third rack, and thus the following equation is established: l3 = l1 (Db / Da). (1) The pitch circle diameter of gear A1 is Da = 12, the pitch circle diameter of gear A2 is Db = 24, and the movement distance ratio is 2. If loss is ignored, the distance traveled is double the output relative to the input. Regarding the opening/closing speed, which is the time derivative of the movement amount, the output is twice the input. Furthermore, the relationship between the moving force F1 of the first rack and the moving force F3 of the third rack is expressed by the following equation. Here, the moving force F1 of the first rack corresponds to the gripping force of a commercially available parallel gripper. The movement force F3 of the third rack corresponds to the gripping force of the developed parallel gripper. F3 = F1 (Da / Db). (2) From the above formula, if loss is ignored, the gripping force of the developed parallel gripper is half that of the commercially available parallel gripper. (b) Finger mechanism The first finger mechanism is connected to the third rack and the first guide. The second finger mechanism is connected to the fourth rack and the second guide. Each finger mechanism has two nails that extend and contract. Each nail is connected to a displacement sensor (LP15-014-R5) for measuring the amount of expansion/contraction and to a compression coil spring via a curved link. The compression spring coil is for maintaining the positioning of the nail. The displacement sensor inserts a metal rod inside the metal pipe, and the output voltage (1 to 5 V) changes when the metal rod moves in the measurement area (14 mm). The greater the amount of metal rod inserted into the metal pipe, the lower the output voltage. The configuration is such that the output voltage of the displacement sensor rises when the nail extends, and the output voltage of the displacement sensor decreases when the nail shortens. The shape of the curved link is devised so that the curved link does not come into contact with other members when each nail extends and contracts. The movement range of the nail is set by the curved link coming into contact with a mechanical stopper. The developed parallel gripper has a total of four displacement sensors. 9990 Authorized licensed use limited to: Auckland University of Technology. Downloaded on October 06,2020 at 06:28:00 UTC from IEEE Xplore. Restrictions apply. (c) Drive control system Figure 7 shows the configuration of the drive control system. The control unit is a single microcomputer (Raspberry Pi 3 Model B) and it controls the opening/closing operation of the parallel gripper. The amount of expansion/contraction of the finger mechanism is detected by a displacement sensor (LP15-014-R5), which is input into the control unit via the displacement sensor amplifier (LP15-014-R5-A) and the A/D converter on the developed board. The displacement sensor is connected to the developed electronic circuit board via the GPIO terminal of the microcomputer. The external host system controller and the control unit of the parallel gripper can communicate with each other wirelessly or by a wired connection. For the above configuration, the operation of the parallel gripper changing from the closed state to the open state will be described here. By driving the drive portion, the first rack and the second rack are moved in opposite directions from each other along the horizontal direction. Gear A1, which meshes with and is driven by the first rack, rotates counterclockwise. Furthermore, Gear B1, which meshes with and is driven by the second rack, rotates clockwise. As a result, the third rack that meshes with Gear A2 and the fourth rack that meshes with Gear B2 move in opposite directions from each other in the horizontal direction. Therefore, the first finger mechanism and the second finger mechanism also move away from each other." + ] + }, + { + "image_filename": "designv11_71_0001140_1.i010790-Figure4-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0001140_1.i010790-Figure4-1.png", + "caption": "Fig. 4 Schematic of 5 UAVs A1; : : : ;A5 pursuing an intruder UAV swarm.", + "texts": [ + " The reason for this could be, for example, that the intruders seek to perform a coordinated attack on a protected area or because the intruder UAVs are flying in a leader\u2013follower configuration and the followers need to be close to the leader. Because of their need to stay close together, the intruder UAVs aremoving as a flock.We furthermore assume that this flock of intruder UAVs lies within a circle of radiusR. A scenario comprising five UAVs performing a cooperative pursuit of an intruder swarm bounded within a circle is illustrated in Fig. 4. Later, in Sec. V, we allow for the radiusR of the circle to be time varying; in Sec. VII, we remove the circular assumption altogether and consider that the intruder swarm can be of any arbitrary shape. Let r1; : : : ; rn represent the position vectors of the pursuing UAVs, and V1; : : : ;Vn represent their respective velocity vectors. Let VB represent the velocity vector of the circle encompassing the intruder UAV swarm. Consider a virtual point X that lies in the convex hull of A1; : : : ; An defined by X Xn i 1 \u03bbiri; Xn i 1 \u03bbi 1; \u03bbi > 0; i 1; : : : ; n (7) An algorithm to compute \u03bb1; : : : ; \u03bbn is discussed in Sec" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0001525_acc45564.2020.9147798-Figure2-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0001525_acc45564.2020.9147798-Figure2-1.png", + "caption": "Fig. 2: Planar PAA underactuated manipulator in the absence of gravity. The second and third joints are active, while the first joint is passive.", + "texts": [ + " However, we have to admit that our strategy cannot achieve global stabilization due to the existence of singularities. This paper is structured as follows: Section II presents the dynamic model of the planar PAA underactuated manipulator; Section III introduces the proposed control strategy, including partial integrability, chained form transformation, and the null space avoidance control framework; Section IV shows the simulation results and gives some discussions. The parameters of the planar three-link underactuated manipulator are illustrated in Fig. 2. Specifically, qi is the angle of the i-th link; mi is the mass of the i-th link; Ii is the moment of inertia of the i-th link around its center of mass; `i is the length of the i-th link; `ci is the length between the i-th joint and the center of mass of the i-th link; and \u03c4i is the applied torque of the i-th joint. For simplicity, s2 \u201c sinpq2q, c2 \u201c cospq2q, s3 \u201c sinpq3q, c3 \u201c cospq3q, s23 \u201c sinpq2 ` q3q, c23 \u201c cospq2 ` q3q. Due to the absence of gravity, the dynamics can be formulated as [12] Mpqq:q `Cpq, 9qq \u201c \u03c4 , (1) where q \u201c rq1, q2, q3s T and \u03c4 \u201c r0, \u03c42, \u03c43sT " + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0001031_icimia48430.2020.9074843-Figure9-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0001031_icimia48430.2020.9074843-Figure9-1.png", + "caption": "Fig. 9 Rotary-Linear Induction Motor: 1-primary stator, 2- secondary stator, 3-housing, 4- rotor [12]", + "texts": [ + " A resultant two-dimensional magnetic field is generated across the air gap between the stator and rotor. The tubular rotor is designed for interacting with the resultant magnetic field. The interaction of two fields produces both rotational torque and linear force simultaneously. Here both rotational and linear movements are not independent of each other. Design 2: Design 2 is simpler as compared to design 1, which employs two stators with ring and distributed windings respectively is shown in Fig.9. The primary stator is placed horizontally and has distributed winding generates the rotating magnetic field. It leads to rotatory motion. The secondary stator is placed vertically have the ring winding generates the traveling magnetic field. It leads to linear motion. Both the primary and secondary stators are placed mutually perpendicular to each other. The length of the rotor is long as compared to design 1. When the primary stator is energized, then the rotating magnetic field is produced. The resultant magnetic field across the airgap links with the rotor and develops a torque" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0000427_j.ifacol.2019.12.578-Figure2-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0000427_j.ifacol.2019.12.578-Figure2-1.png", + "caption": "Fig. 2. Laboratory training workplace.", + "texts": [ + " The simulator contains 16 digital inputs, 16 digital outputs, 1 analog input, and 1 analog output for general purposes. For testing hardware and software counters there are provided two pulse outputs with fixed frequencies of 4 kHz and 40 kHz and amplitude 24V. Also there are intended BCD switches possible to set numbers in the range from \u2013999 to 999. To display BCD numbers in the same range four 7- segment indicators are provided. The analog input and output range is from -10Vdc to +10Vdc and both voltages are measured and displayed. The laboratory training workplace depicted in Fig. 2 consists of the described above simulator \u2013 1 [P. Yakimov et al. (2019)], PLC Simatic S7-1200 \u2013 2, PC - 3, smart AC variable speed drive model Commander SKA1200037 \u2013 4 [Control Techniques], AC induction motor \u2013 5 and reducer - 6 [Motovario], and incremental rotary encoder \u2013 7 [Sensata Technologies]. The configuration of the working place allows PLC software development and testing using real industrial devices. There are highlighted the electrical and mechanical connections in the laboratory set-up for practicing loop control of AC induction motor" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0003714_esda2014-20232-Figure10-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0003714_esda2014-20232-Figure10-1.png", + "caption": "Figure 10:The five first frequencies of the pedestal (1st:538.2, 2nd:1172.7, 3rd:1951.5, 4th:2945, 5th:3063.1)", + "texts": [], + "surrounding_texts": [ + "The water brake specification is presented in table 1. In the first step, an initial estimation of loading about the rotor is made. Considering the mechanism of power absorption which is based on passing a water flow with an approximate pressure of 3.5 bar through the mechanical structure which produces friction between moving and stationary part, the initial loading is estimated. By measuring the amount of the torque, the turbine torque can be calculated precisely. In this regard, a torque load is considered on the load cell arm. Bearings components should be chosen carefully to handle the loading. Consequently, angular contact bearings are used for their ability to deal with high rotational speed and their widely use in precise applications. All connections are bolted and lock nuts are used to create preload on the bearings. Twelve anchor bolts are used tohold the \u201cCabAssy\u201don the housing and maintain bearings preload on their position. To transmit torque from the engine to the dynamometer, three options can be considered including: flexible coupling, diaphragm coupling and floating quill shaft (splined). Here, floating quill shaft is used. Anti-seize lubrication is applied to the splines. Water brake shaft splines are carefully designed and machined to be mounted on the shaft and engine spline. To prevent premature wear of the splines, the shaft is hardened. Of course, if the mechanical analysis is performed on all parts, more accurate results will be achieved. However, the analysis has been carried out on some important parts according to the need and importance which is more time and cost effective. The effect of the other parts are only considered in boundary conditions of the important components during the analysis. In some cases, transient analysis is needed to assess critical situations. The static load analysis in is done in 1 second period of time which is equivalent to a time step. Mode shape analysis is also performed for the structural components for predicting critical damage locations." + ] + }, + { + "image_filename": "designv11_71_0002989_978-3-642-16214-5_9-Figure9.31-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0002989_978-3-642-16214-5_9-Figure9.31-1.png", + "caption": "Fig. 9.31. Multi-plate clutch and belt brake unit. 1 Rectangular ring (rotating); 2 grooved ring; 3 O-ring (static); 4 outer plate carrier and brake drum; 5 housing; 6 shaft; 7 piston; 8 pressure oil supply; 9 brake belt; 10 steel plate; 11 end plate; 12 snap ring; 13 lined", + "texts": [ + " In the case of automatic transmissions of planetary design, brakes serve to support components of the planetary gear in the housing as needed. Multi-plate brakes and belt brakes are established designs. In belt brakes, a metal belt is looped around a brake drum. The important advantage of the belt brake is: \u2022 small radial installation space requirement. On the other hand, its limitations include: \u2022 non-harmonic, poorly reproducible torque built-up, \u2022 uneven load share with wear on the belt ends, \u2022 radial forces and \u2022 high sensitivity to adjustment tolerances. Figure 9.31 shows a multi-plate clutch and belt brake unit from the MercedesBenz W5A 030 5-speed automatic transmission (Figure 12.22). plate; 14 inner plate carrier; 15 piston return spring In modern transmission designs, belt brakes have been to the greatest extent replaced by multi-plate brakes. Belt brakes will not be discussed in the following, whereas the layout and design of multi-plate clutches and brakes will be explored further. Wet multi-plate clutches and brakes (i.e. those that are through-flowed with oil) are used in many transmissions" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0000967_0954406220916504-Figure7-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0000967_0954406220916504-Figure7-1.png", + "caption": "Figure 7. Multi-link suspension mechanism: (a) the schematic diagram of the suspension mechanism and (b) the kinematic model.", + "texts": [ + " The infinitesimal displacement D\u0302H of the vehicle body with respect to the wheel can be expressed as D\u0302H qHS\u0302H, where qH represents the magnitude of the angular displacement of the vehicle body about the axis S\u0302H. If a small change ls of the length of the strut spring along N2 is given at any instant, the infinitesimal displacement D\u0302H of the vehicle body can be found from equations (8) and (9) as D\u0302H \u00bc qHS\u0302H \u00f014\u00de where qH \u00bc ls r\u0302 T s S\u0302H \u00f015\u00de Multi-link suspension. The schematic diagram of a multilink suspension mechanism is shown in Figure 7(a). In this figure, each joint is numbered by Ni (for i\u00bc 1, 2, . . ., 13). The strut spring mounted on one of five links is aligned with the prismatic joint N2 (P-joint). The other joints Ni are spherical ones. Figure 7(b) shows the kinematic model of the multi-link suspension mechanism. The vehicle body can be considered to be the moving platform of the spatial parallel mechanism connected to the wheel by one SPS-serial (N1N2N3) kinematic chain of a shock absorber and five SSserial (N4N5, . . . ,N12N13) kinematic chains. Referring to Figure 7(b), there are five constraint wrenches s\u0302i (i\u00bc 1, . . . , 5) that pass through two S-joints in each SS chain. Due to the strut spring force along P-joint N2, there exists a wrench which can act on the vehicle body along r\u0302s of the reciprocal Jacobian of the SPS chain. Since it has five constraint wrenches acting on the vehicle body, the multi-link suspension mechanism is a 1-DOF parallel mechanism. The reciprocal Jacobian Jrp of the suspension mechanism has only one unit column vector S\u0302H which is reciprocal to s\u0302i (i\u00bc 1, " + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0001059_s42417-020-00207-4-Figure3-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0001059_s42417-020-00207-4-Figure3-1.png", + "caption": "Fig. 3 Idealized rotor air-ring bearing system", + "texts": [ + " The first row of FH is provided in C and B at distances of L/6 and L/4, respectively, where L is the length of journal bearing (see Fig.\u00a02b). The second row of FH is provided in C and B at 5L/6 and 3L/4, respectively. Air at supply pressure ps enters the FH provided in C before it flows through the AR and into the FH in B. It then flows in the AB and is exhausted to the atmosphere through the bearing ends. Figure\u00a02c shows the geometry of FH in B and C. 1 3 For determining the DC, the RARBS shown in Fig.\u00a02a can be replaced by an idealized physical system as shown in Fig.\u00a03 [24]. The following quantities are identified to describe the behavior of the idealized system: (1) absolute displacements of CGJ and CGB [measured relative to the fixed casing center (O) in Fig.\u00a01]: (j) J and (j) B (the superscript j is used to indicate motion of CGJ and CGB along direction j; j can be either x or y as in Fig.\u00a01); (2) absolute velocities: \u0307 (j) J and \u0307(j) B ; and (3) absolute dynamic load-carrying capaci- ties: (ij) J and (ij) B (the superscripts i and j in (ij) J are used to indicate the bearing forces in direction i when CGJ moves along direction j; both i and j can be x or y). When i is x and j is y, (ij) J can be interpreted as a vector of instantaneous values of F(xy) J . Quantities (1), (2), and (3) are obtained via TSS. In addition, system constants mJ and mB, and the stiffness and damping coefficients of AB and AR, i.e., KAB, KAR, CAB, CAR, respectively, are used to characterize the idealized RARBS. The rotor load PJ = mJg\u22152 and the joint-bushing load PB = mBg\u22152 (see Fig.\u00a03). The SSS provides the following: (a) coordinates of the SEP identified at t = 0, i.e., ( S (x) J (0), S (y) J (0) ) of CGJ, and ( S (x) B (0), S (y) B (0) ) of CGB which can be defined in terms of eccentricity (e) and attitude-angle (\u03b1) as ( e0 J , 0 J ) and ( e0 B , 0 B ) , respectively (see Fig.\u00a01); and similarly (b) components of static load-carrying capacity ( FJ (x)(0),FJ (y)(0) ) and ( FB (x)(0),FB (y)(0) ) . At time t = 0, the TSS starts from the SEP of CGJ (OJ) and CGB (OB) as shown in Fig", + " Components of incremental dynamic load-carrying capacity (xj) J and (yj) J are: f (xj) J (t) = F (xj) J (t) \u2212 FJ (x)(0) and f (yj) J (t) = F (yj) J (t) \u2212 FJ (y)(0) , where j can be x or y. Similar sets of expressions can be written for (xj) B and (yj) B . The equations of motion of journal and bushing: (1\u22152)mJ\u0308 (j) J = (jx) J + (jy) J and (1\u22152)mB\u0308 (j) B = (jx) B + (jy) B , are equations of free vibration since the problem considered here investigates SE vibration [25]. Referring to Fig.\u00a04, and allowing i to be x or y, and j to be x or y: and For example, when i is x and j is y, K(ij) AR = K (xy) AR , C (ij) AB = C (xy) AB etc. (see Fig.\u00a03). In Eqs. (1, 2) vectors such as (j), \u0307(j) are linearly independent. A set of orthonormal vectors can be formed from them using the Gram-Schmidt orthogonalization process [3, 21, (1) (ij) J = \u2212K (ij) AB ( (j) J \u2212 (j) B ) \u2212 C (ij) AB ( \u0307 (j) J \u2212 \u0307 (j) B ) (2) (ij) B = \u2212K (ij) AR (j) B + K (ij) AB ( (j) J \u2212 (j) B ) \u2212 C (ij) AR \u0307 (j) B + C (ij) AB ( \u0307 (j) J \u2212 \u0307 (j) B ) 26]. Letting, (j) = (j) J \u2212 (j) B , and \u0307(j) = \u0307 (j) J \u2212 \u0307 (j) B , and allow- ing i to be x or y, and j to be x or y, Eq.\u00a0(1) is re-written as: Again letting, (j) = (j) B , and \u0307(j) = \u0307 (j) B , Eq" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0003878_s11434-014-0376-5-Figure5-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0003878_s11434-014-0376-5-Figure5-1.png", + "caption": "Fig. 5 The unsymmetrical situation of the horse standing posture", + "texts": [ + " Furthermore, especially during the locomotion, three-point supporting and two-point supporting situations might occur. When standing, horses adjust leg postures to make itself comfortable. Actually, the adjustment is to find a posture for mitigating the leg muscles burden and reducing energy consumptions. As a result of the adjustment, there are two final standing postures: one is that two forelimbs and two hind limbs are respectively symmetric about the longitudinal plane of the horse (Fig. 4) while the other is not symmetrical (Fig. 5). For the symmetrical situation, showing in Fig. 4, A, B, C and D stand for foothold of the right forelimb, the left forelimb, the right hind limb, the left hind limb respectively. G stands for the center of gravity of the horse. W refers to the whole weight of the horse. V1, V2, V3 and V4 represent the vertical ground reaction force of the left forelimb, the right forelimb, the left hind limb, the right hind limb respectively. x12 represents the distance between two forelimb footholds and the center of gravity in the longitudinal plane while x34 refers to the distance between two hind limb footholds and the center of gravity", + " For symmetrical about the longitudinal plane, there is: y1 \u00bc y2; y3 \u00bc y4: \u00f01\u00de All external forces should meet the force equilibrium and moment equilibrium. For the force equilibrium, it yields: V * 1 \u00fe V * 2 \u00fe V * 3 \u00fe V * 4 \u00bc W * : \u00f02\u00de For the moment equilibrium, the following relationship can be given: V * 1 x*12 \u00fe V * 2 x*12 \u00fe V * 3 x*34 \u00fe V * 4 x*34 \u00bc 0; V * 1 y*1 \u00fe V * 2 y*2 \u00fe V * 3 y*3 \u00fe V * 4 y*4 \u00bc 0: \u00f03\u00de An angle is between two forelimbs as same as two hind limbs in the unsymmetrical case (Fig. 5) and the ground reaction forces have the horizontal components, which are denoted by H1, H2, H3, and H4. Therefore, the ground reaction force denoted by F1, F2, F3 and F4 respectively is the resultant force of the vertical component and the horizontal component. In the unsymmetrical situation, all external forces must meet the force equilibrium and the moment equilibrium. For the force equilibrium, it is easy to get the following relationships: V * 1 \u00fe V * 2 \u00fe V * 3 \u00fe V * 4 \u00bc W * ; H * 1 \u00fe H * 2 \u00fe H * 3 \u00fe H * 4 \u00bc 0: \u00f04\u00de For the moment equilibrium, we can find that: V * 1 x*1 \u00fe V * 2 x*2 \u00fe V * 3 x*3 \u00fe V * 4 x*4 \u00bc 0; V * 1 y*1 \u00fe V * 2 y*2 \u00fe V * 3 y*3 \u00fe V * 4 y*4 \u00bc 0; H * 1 y*1 \u00fe H * 2 y*2 \u00fe H * 3 y*3 \u00fe H * 4 y*4 \u00bc 0; H * 1 h * \u00fe H * 2 h * \u00fe H * 3 h * \u00fe H * 4 h * \u00bc 0: \u00f05\u00de where h is the height of the center of gravity of the horse" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0002119_icaccm50413.2020.9212874-Figure10-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0002119_icaccm50413.2020.9212874-Figure10-1.png", + "caption": "Fig. 10. Propeller fixed support at the tips (case2).", + "texts": [ + " Figure 6, 7 and 8, 9 show the total deformation and stress for CFRP and GFRP materials, respectively. Table II presents the comparison of results for CFRP and GFRP. 60 Authorized licensed use limited to: Carleton University. Downloaded on May 30,2021 at 15:29:13 UTC from IEEE Xplore. Restrictions apply. A Quadcopter\u2019s propeller is subjected to horizontal bending due to collision of the propeller with obstacles or a wall. To calculate the bending deformation under the applied rotational moment about the rotation axis, the propeller has been fixed by both the tip as shown in Figure 10.In addition, it is fixed at the point where it is connected to the motor shaft (Figure 11).Figure 12 shows the applied rotational velocity of 897 rad/s [8].Figure 13, 14 and 15, 16 show the total deformation and equivalent stress for CFRP and GFRP materials, respectively. The comparison of the obtained results is presented in Table III. 61 Authorized licensed use limited to: Carleton University. Downloaded on May 30,2021 at 15:29:13 UTC from IEEE Xplore. Restrictions apply. Fig. 14. Total deformation of GFRP propeller (case2)" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0003632_0954406214549786-Figure3-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0003632_0954406214549786-Figure3-1.png", + "caption": "Figure 3. Constrained surfaces in the finite element analysis.", + "texts": [ + " All nodes on the end surface of the worm connected to the coupling are fixed, and the nodes on the surfaces of the worm in contact with the supporting bearings can only rotate around the axis of the worm. Step 3: Define the contact between worm and worm wheel with contact elements. The elements on the potential contact area are refined. Step 4: Define a master node to simulate rotation with the rigid shaft. All nodes on the surface of the worm wheel in contact with its shaft are coupled with a master node MC as shown in Figure 3. These nodes can rotate around MC. Step 5: Perform a static finite element analysis to determine the rotation angle r of the master at UNIV OF CONNECTICUT on May 22, 2015pic.sagepub.comDownloaded from node around the x-axis when the torque T s\u00f0 \u00de is applied on MC. After that, the equivalent tangential meshing stiffness of worm and worm wheel can be calculated by equation (11) kc s\u00f0 \u00de \u00bc T s\u00f0 \u00de r2wg r \u00f011\u00de Step 6: Repeat the five steps above to get the meshing stiffness of the worm gear pair at different tilting angles" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0001892_012007-Figure3-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0001892_012007-Figure3-1.png", + "caption": "Figure 3. Display (a) three magnetic disks that function as rotors (b) two winding discs that function as stators", + "texts": [ + " In this paper, the axial flux motor is designed with a multi-sided model that has two winding disks, which become the stator while three interconnected magnetic disks which become the rotor Figure 2. ICECME 2019 IOP Conf. Series: Materials Science and Engineering 931 (2020) 012007 IOP Publishing doi:10.1088/1757-899X/931/1/012007 To develop an axial flux motor requires an initial step by determining the parameters as targets to be achieved in Table 1. The multi-sided flux axial motor design in this paper is designed with a coreless type (figure 3), where there is no metal element on the winding disk. Still, here the windings are coated with resin material. So there is no slot winding (slot less). The coreless type will not produce cogging torque so that the efficient level is better and very useful for use in electric vehicles ICECME 2019 IOP Conf. Series: Materials Science and Engineering 931 (2020) 012007 IOP Publishing doi:10.1088/1757-899X/931/1/012007 To assume the maximum power density with no load, the thickness of the rotor seat can be calculated by the following equation: (1) Copper loss is calculated using the following equation: (2) On the stator stand the angle changes \u03b2 depending on the angle \u03b8 by following the following equation: (3) The simulation used in this paper is Software Finite Element Method Magnetics version 4" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0002651_3439231.3439268-Figure2-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0002651_3439231.3439268-Figure2-1.png", + "caption": "Figure 2: EI-Edurobot in design", + "texts": [ + " This decision was based on the realization that training for empathy towards children with deficiencies (mostly ASD) who participate in inclusive classroom is under-researched, especially utilizing ER. Also, the cost of robots in similar cases is rather high or the utilized robots are specifically designed for a specific study with specific settings. Thus, an open design for a robot which could be used for empathy training, but also other goals was the main aim of this attempt. The technical aspects and the intended research are presented in the next sections, along with pre-decision considerations (see discussion section) which affected the robot\u2019s design overall. The EI-Edurobot (Figure 2) is an open source training robot that can initially targeted children aged 4 to 9 y.o. but can be also described as multi-purpose teaching robot which can be used in numerous occasions. It incorporates, several sensors (Figure 4) and a user-friendly programming language and management platform (still under development - Figure 5). The sensors are mainly divided into two categories: a) safety-related sensors, and b) the education-related sensors. The EI-Edurobot is addressing children and special attention has been paid to reassuring both their and its safety" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0002442_icarcv50220.2020.9305376-Figure2-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0002442_icarcv50220.2020.9305376-Figure2-1.png", + "caption": "Fig. 2: Simulated printing the start of a single trajectory.", + "texts": [ + " This algorithm has been developed [24]. Fig. 1e shows way-points in the generated trajectories from this algorithm and Fig.1f shows an overlapped view of the way-points and the 3D model of a GSS. To simulate the printing process, a simulation environment has been developed in Matlab according to the actual dimension of the printer and robots (ABB IRB 120) and the robot is modeled using Denavit\u2013Hartenberg (DH) parameters. The robot simulation has been done using the robotics toolbox developed by Peter Corke [25]. Fig. 2 shows the simulation environment in Matlab and the printing of the first trajectory. During the printing process, the robot arm affixed with a printhead, starts by depositing the first layer of melted material on the rotating column. It then progressively adds 855 Authorized licensed use limited to: Cornell University Library. Downloaded on May 24,2021 at 13:12:17 UTC from IEEE Xplore. Restrictions apply. 1.5 layer-upon-layer and expands the print radially outwards. Each layer has multiple trajectories and printing starts from the bottom and ends at the top" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0001381_s1068798x20060210-Figure5-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0001381_s1068798x20060210-Figure5-1.png", + "caption": "Fig. 5. Dobroborskii\u2019s ERD-3 test bench.", + "texts": [ + " In view of the polytropic relation between the pressure and volume of the compressed air in the working chamber, the following formula was proposed where is the piston displacement at impact; is the initial height of the working chamber; is the work of impact; is the polytropic index ( = 1.25\u20131.35); is the initial compressed-air pressure in the working chamber; and is the cross-sectional area of the piston. On the basis of this design, numerous organizations created test benches. Dobroborskii himself produced the first, known as the ERD-3 system [6]. In this test bench, as we see in Fig. 5, a powerful pneumatic cylinder 3 applies force to handle 4 of the jackhammer. The cylinder is attached to column 2 on base 1. Cylinder 5 with piston 6 is also mounted on base 1. The piston is constantly pressed against cap 7 by compressed air supplied through pipe 8 and valve 9 with manometer 10. Upward piston motion is limited by cap 7, to which clamp 11 is attached. Impacts of the hammer face are transmitted through mandrel 12. At the piston, we see pin 13 with a projection which touches contact 14 as the piston is moved by the hammer blows" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0002510_yac51587.2020.9337612-Figure1-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0002510_yac51587.2020.9337612-Figure1-1.png", + "caption": "Fig. 1 Construction of quadrotor", + "texts": [ + "1 10 9/ YA Authorized licensed use limited to: UNIVERSITY OF CONNECTICUT. Downloaded on May 17,2021 at 04:09:41 UTC from IEEE Xplore. Restrictions apply. 145 .Dynamic Model of Quadrotor UAV A. Dynamic Modeling of Quadrotor UAV in Normal Operation. Quadrotor UAVs are underactuated system with four rotors and six degrees of freedom. The altitude motion is achieved by changing the power of four motors at the same time. The attitude angle of UAVs can be adjusted by adjusting the speed of motor by control quantity. Fig. 1 is the construction of traditional quadrotor UAV. The nonlinear equation of quadrotor is shown in (1) with following assumptions. 1. The fuselage of quadrotor UAV is rigid with symmetrical structure and centroid at geometric center. 2. Consider the small motion range of quadrotor UAV, the influence of centripetal force of earth revolution on UAV motion is ignored. 3. It is assumed that the air resistance of quadrotor UAV is proportional to its flight speed. 4. Neglecting the influence of physical and electrical characteristics of the quadrotor UAV on its body parameters" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0002310_asemd49065.2020.9276162-Figure4-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0002310_asemd49065.2020.9276162-Figure4-1.png", + "caption": "Figure 4. simulation model and experimental prototype.", + "texts": [ + " It can be seen that under the combined action of signal wave and carrier wave, the amplitude of its voltage frequency is the maximum at 50Hz of the fundamental wave, and there will also be large amplitude near the switching frequency, and the frequencies are fc 2f1and fc 4f1. The two-dimensional Fourier analysis results of the air gap electromagnetic wave are shown in Fig.2. Combined with theoretical analysis, the previous radial electromagnetic force and the electromagnetic vibration results are calculated and simulated according to the frequency. Then the electromagnetic vibration response under the main frequency is calculated by full method. The calculation results are shown in Fig.3. The simulation model and experimental prototype are shown in Fig. 4. After the stator vibration deformation at 1400Hz, it drives the foot to vibrate together, and further magnifies on the foot. The vibration mode of the foot is mainly the overall swing and the end bending deformation and the result is that the vibration amplitude of the foot end is the largest. The vibration acceleration amplitude of the main vibration frequency is shown in Fig.5. The vibration acceleration measured by the experiment is in good agreement with the simulation results. IV. CONCLUSIONS When the motor is fed by frequency converter, higher harmonic is introduced into machine winding, and the air-gap flux density can produce new frequency" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0003270_icpe.2015.7167862-Figure5-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0003270_icpe.2015.7167862-Figure5-1.png", + "caption": "Fig. 5. Concept diagram of weighted demagnetization", + "texts": [ + " When demagnetization in permanent magnet of motors is occurred, each permanent magnet has a little difference on demagnetization ratio. Therefore, we determined irreversible demagnetization pattern having different demagnetization ratio between N-pole and S-pole. Fig. 3 shows a concept diagram of inequality demagnetization. In addition, Fig. 4 shows magnetization shape diagram according to demagnetization ratio. Ratio of irreversible demagnetization patterns is N-pole 50%, S-pole 60% and N-pole 50%, S-pole 70%. Fig. 5 shows a weighted demagnetization pattern. One of N-pole and S-pole is demagnetized. Demagnetization pattern of this model cannot be occurred in real conditions. However, it has a special meaning for BEMF harmonic characteristics analysis. Because the BEMF waveform is directly influenced by flux linkage variance of each phase, we selected weighted demagnetization model. Namely, flux linkage is occurred to difference on increment and decrement of flux linkage. In addition, it affects distortion of BEMF waveform" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0001211_012001-Figure3-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0001211_012001-Figure3-1.png", + "caption": "Figure 3. Calculation model of a two-unit road train.", + "texts": [ + " The mathematical model of road train planar motion described in detail in [15] may be used to study the curvilinear motion of multi-unit vehicles with active trailer units. Longitudinal and transverse angles of relative motion of moving road train units along the horizontal base are not large. Therefore, the movement of each unit as a solid body is considered in the horizontal plane taking into account the lead angle in the direction of movement on flat non-yielding surface and consists of translational and rotational motion of the center of mass (Figure 3). IASF-2019 IOP Conf. Series: Materials Science and Engineering 819 (2020) 012001 IOP Publishing doi:10.1088/1757-899X/819/1/012001 The following coordinate systems are introduced to describe this road train motion case: fixed coordinates X'OY', the origin (point \u041e) coincides with the start point of simulated route; moving coordinates X1C1Y1 and X2C2Y2 referenced to centers of mass (CM) of the tractor and semitrailer, respectively; coordinates XjiOjiYji referenced to the i-th wheel of the j-th unit" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0001285_icieam48468.2020.9112086-Figure9-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0001285_icieam48468.2020.9112086-Figure9-1.png", + "caption": "Fig. 9. The electromechanical drive for the linear movement", + "texts": [ + " The control commands in the symbolic form are \"0U\" \u2013 to open on a signal from the sensor; \"1U\" \u2013 to close on the signal from the sensor; \"2U\" \u2013 stop; \"3U\" \u2013 to close (forcibly); \"4U\" \u2013 to open (forcibly). An example of entering a command in a terminal program is shown in Fig. 7. The hardware implementation of the control system is shown in Fig. 8. The control device 1 is made of impactresistant heat-resistant ABS plastic UL-94V0 in the D3MG housing. It has the dimensions of 53\u00d790\u00d757 mm and together with the frequency converter 2 is fixed on a DIN rail. The electromechanical drive for the linear movement of the main pipeline valve is shown in Fig. 9. [1] GOST 33852-2016, Pipe fittings. Gate valves for oil trunk pipelines. General specifications. [2] TU 372100-002-81484267-2016, Wedge gate valves. [3] A. Garganeev, A. Karakulov, and S. Landgraf, Electric drive of shut-off valves. Tomsk: TPU, 2013. [4] D. White and G. Woodson, Electromechanical energy conversion. Moscow: Energia, 1964. [5] K. Kim and S. Ivanov, \u201cSimulation of a combined electric drive,\u201d Russian Electromechanics, vol. 62, no. 3(197), pp. 44\u201351, 2019. DOI: 10.17213/0136-3360-2019-3-44-50" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0000315_icems.2019.8921590-Figure4-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0000315_icems.2019.8921590-Figure4-1.png", + "caption": "FIGURE 4. Exploded view of the mechanical design of the OMT.", + "texts": [ + " The concept of our OMT design is based on the following rules: 1) the matching element should not be split into four identical blocks intersecting on the axis of the input circular waveguide (as described in [16]); 2) the matching element should not be a mobile part to be glued, screwed, soldered or just VOLUME 1, 2013 481 inserted in the center of the turnstile junction (as described in [16] and [17]); and 3) the total number of blocks to assemble the OMT should be minimized. Respecting these rules is possible with the final structure of the OMT presented in Fig. 3(a) and 3(b). For each polarization, an E-plane 180\u02da out-of-phase power combiner is used in order to recombine the RF signals split by the turnstile junction through opposite waveguide outputs. From a practical point of view, the final structure of the OMT presented in Fig. 3(a) and 3(b) is realized by superimposing three aluminum blocks (see Fig. 4). The upper block consists of a circular waveguide input, a circular waveguide transition, two E-plane power combiners and four H-plane 90\u02da bends (as shown in Fig. 5(a)). The lower block consists of the base of the turnstile junction with two superimposed cylinders as a matching stub, four H-plane 90\u02da bends and four E-plane 90\u02da bends (as depicted in Fig. 5(b)). The middle aluminum block acts as an interface between the lower and upper blocks, connecting the outputs of the turnstile junction to the inputs of the power combiners", + " 9, the return loss at the input of the power combiner is below \u221240 dB over our band of interest. In order to keep a finite thickness during the manufacturing process, a cusp at the intersection (Fig. 9) of the two-miter bends is truncated at a width of 0.1 mm. In order to be able to machine the 4-section transformer and the 3-miter bends on a cut length of 6.33 mm (represents the full width of the band 1 waveguide), a standard 2 mm-diameter end-mill was used. III. EXPERIMENTAL VERIFICATION OF THE OMT The structure of the OMT presented in Fig. 4 is realized by superimposing three aluminum blocks made of Aluminum 6061 with a T6 temper (see Fig. 10(a) and 10(b)). The three parts are machined using a CNC milling machine and two dowel pins are used for alignment. The machining required for the manufacturing of this OMT is standard for a CNC machine with tolerance within reach of a production-type workshop. The dowel pins and the features of the turnstile and bends have +/\u22120.01 mm tolerances which is the tightest tolerance in the OMT. The rest of the features have +/\u22120" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0001004_icit45562.2020.9067177-Figure1-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0001004_icit45562.2020.9067177-Figure1-1.png", + "caption": "Fig. 1: 7-phase SPM machine with toothconcentrated winding under consideration", + "texts": [ + "5 slot per pole per phase) is evaluated in [6]. As third harmonic winding factor is higher than the fundamental one, the rated machine current control aims at generating first and third harmonic components of the same order. Thus a particular magnet layer segmentation is determined to achieve acceptable torque density and low torque ripple. To some extent, the resulting machine could be qualified as bi-harmonic or double-polarity one [7]. The resulting magnetic circuit with the magnet segmentation and the winding layout are represented in Fig.1a. From this study, a prototype has been manufactured and the real winding distribution is depicted in Fig.1b. The present paper specifically addresses the fault-tolerant control of this machine. Introduced in [8], the generalized Maximum Torque Per Ampere (MTPA) control strategy is applied with, for the present study, the particularity of looking for an easy implementation (classical PI current controllers and intersective modulation). Furthermore, referring to other effective post-fault control strategies for multi-phase machine [9], [10], the proposed control structure is the same whatever the machine status is (healthy or open-circuit faults)" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0003878_s11434-014-0376-5-Figure10-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0003878_s11434-014-0376-5-Figure10-1.png", + "caption": "Fig. 10 (Color online) Changing the direction of ground reaction force hF", + "texts": [ + " The force analysis of the forelimb is divided into two parts: one is to keep the forelimb posture and change hF, which is the direction of the ground reaction force, this part will explore the joint moment variations to explain how the ground reaction force affects the joint forces and moments. The other one is to keep the ground reaction force unchanged and change the joint angles h1 and h2, which indicates the leg postures. This section will explore that how the leg postures affect the joint moments. 4.1 Influences on the joint forces and moments by changing hF For a given leg posture (Fig. 10), obviously, it is not at the dead-point supporting state. Therefore, no matter how the direction of the ground reaction force hF is varied, all joint moments could not be zero. However, there must be a line called the best supporting line that comes close to the deadpoint line. If the ground reaction force F is collinear with the best supporting line, the overall evaluation M of the leg will reach the minimum. Now, changing the direction of the ground reaction force (Fig. 10), the joint moments and the overall evaluation are calculated. The line determined by the hF when the overall evaluation M is minimum is the best supporting line of the given leg posture. Estimating the actual situation, the range of hF is [20 ,160 ]. The variations of the joint moments M1, M2, M3, M4 and the overall evaluation M are showed in Fig. 11. In Fig. 11, the minimum values of four joints are zero, it shows that the single joint moment could get to the minimum when the ground reaction force along the corresponding directions of the ground reaction force hF" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0003275_aim.2015.7222530-Figure1-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0003275_aim.2015.7222530-Figure1-1.png", + "caption": "Figure 1. The proposed drug delivery system for capsule endoscopy.", + "texts": [ + " The remainder of this paper is organized as follows. Section II provides the details of the magnetomechanical system under study. Section III presents the theoretical methods for the optimization of the flux density with multiple arc-shaped magnets. Section IV provides the verification of one of our optimized structures with experimental results for flux density, magnetic torque and piston force generated to release drug. Finally, discussions of the results and future work are presented in Section V. Fig. 1 shows a ring-shaped external magnet that produces a rotating magnetic field around the patient bed when it is physically rotated. A 3.17 S mm cubic IPM with magnetization grade of 1.4 T (i.e., NSO) is placed in the prototype of a capsule. The IPM is rotated by the external rotational magnetic field. This rotational movement is then converted into a translational movement by a slider-crank mechanism. In this way, a piston will push drug out of a reservoir when the external magnet is rotated around the patient" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0000115_012066-Figure1-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0000115_012066-Figure1-1.png", + "caption": "Figure 1. Using a tool cone to install the laser head in the machining center", + "texts": [ + " Enterprises with these types of production do not use robotic lines or sections in which industrial robots would automatically change workpieces or additionally process the surfaces of workpieces with a tool mounted on the robot's arm. That is, under these conditions, in order to realize the possibility of using a laser head on a CNC machine, it is necessary to partially or completely modernize the laser head frame and adapt it for manual installation in the machine spindle by an operator or a preset person. Fig. 1 shows an example on how to use a tool cone to install the laser head in a CNC machine. A diagram of the laser head integration into the machining center, developed with using a circuit diagram of a fiber laser setup, presented in [8], is shown in Fig. 2. IPDME 2019 IOP Conf. Series: Earth and Environmental Science 378 (2019) 012066 IOP Publishing doi:10.1088/1755-1315/378/1/012066 When solving the problem of integrating the laser head into the machining center, it is necessary to determine the installation location of the main system unit" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0002509_978-981-15-6868-8_1-Figure17.1-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0002509_978-981-15-6868-8_1-Figure17.1-1.png", + "caption": "Fig. 17.1 Dual chamber microbial fuel", + "texts": [ + " Such MFCs may prove valuable to generate electricity in remote sensing regions. Dual chamber MFC is one of the simplest designs in MFCs (Niessen et al. 2004; Phung et al. 2004). Dual chamber MFC consists of two chambers; they are anode and cathode separated either using salt bridge or proton exchange membrane (PEM). The substrate is used in anode chamber and catholyte water or other catholytes are used in cathode chamber, salt bridge or proton exchange membrane separates the anode and cathode chamber, which helps in the proton transport between the two chambers (Fig. 17.1). The appellation of MFC is also defined by the catholyte or cathode chamber configuration. If air (oxygen) is the electron acceptor for the cathode in the cathode chamber, MFC will be termed as air cathode MFC (Ringeisen et al. 2006; Shantaram et al. 2005). This design although applied for basic research generally produces low power output due to the intricate design, high internal resistance and electrode based losses 380 M. Umesh and H. M. C. Hamza (Du et al. 2007; Logan and Regan 2006a, b). Hamza et al" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0001892_012007-Figure4-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0001892_012007-Figure4-1.png", + "caption": "Figure 4. Design of multi-sided flux axial motors", + "texts": [], + "surrounding_texts": [ + "The simulation used in this paper is Software Finite Element Method Magnetics version 4.2 (FEMM 4.2). Figure 5. shows the topology of the axial flux motor in 2D, there are two sides of back iron with a thickness of 10 mm, on each side there are 4 magnets, the type of magnet used is NdFeb 40, with a water gap of 1 mm. ICECME 2019 IOP Conf. Series: Materials Science and Engineering 931 (2020) 012007 IOP Publishing doi:10.1088/1757-899X/931/1/012007" + ] + }, + { + "image_filename": "designv11_71_0001137_012001-Figure2-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0001137_012001-Figure2-1.png", + "caption": "Figure 2. Location of accelerometer on the experimental jig", + "texts": [ + " From the equation, it is shown that the contact pressure is contributed by material properties: - there are Young\u2019s modulus and Poisson\u2019s ratio, load and also specimen geometry which is the radius of the rolling elements that were in contact. In this work, the roller geometry and the load are similar for all specimens and only different materials were used. Therefore, the different contact pressure obtained in all 4 tests are contributed by the material properties. The vibration measurement is carried out using acceleration parameter where the accelerometer is used. The accelerometer is placed on the 3 different locations on the jig as shown in Figure 2. The first accelerometer C4, is placed on the upper roller mounting. The second accelerometer C2, is placed on the lower roller mounting. The third accelerometer C3, is placed on the base of the experimental rig. The acceleration level is recorded after the motor is running stably which is after a few minutes switched on the motor. ICVSSD 2019 IOP Conf. Series: Materials Science and Engineering 815 (2020) 012001 IOP Publishing doi:10.1088/1757-899X/815/1/012001 Surface scanning is carried out using Infinite Focus Microscope (IFM)" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0002307_syscon47679.2020.9275901-Figure1-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0002307_syscon47679.2020.9275901-Figure1-1.png", + "caption": "Fig. 1. Quadrotor inside the construction block.", + "texts": [ + " In this paper, we look at the autonomous assembly problem as a formation problem, and therefore, positioning multiple blocks in a structure is modelled as arranging robots in a formation. Also, given that pinning controllers can be applied to change the state of a group of agents, in our approach pinning controllers drive the robots to the mounting area. To best of our knowledge, this modelling method constitutes a novel approach to solve the autonomous construction problem with actuated parts. In this work, we study the problem of autonomous assembling of structures using actuated blocks. The blocks have a cubic frame with quadrotors attached (Fig. 1). Also, the three dimensions (height, width, length) of each block are equal to d. Let us describe three coordinate frames needed for our mathematical formulation hereafter. First, consider an inertial frame fixed in the environment, SI . The axes XI and YI describe the horizontal plane, and ZI points upwards. Second, we assume that each block has a coordinate frame Sb located on its geometric center, with axes Xb, Yb, and Zb. Lastly, we can describe Sq as a reference frame fixed at the center of gravity of the quadrotor. The axes Xq and Yq are coplanar with the plane formed by the four motors. Xq points toward the front of the quadrotor between the two front motors. The Yq axis is directed 90o to the left of Xq . Zq points upwards (towards the origin of Sb), perpendicular to the plane of the motors. Both Sb and Sq are depicted in Fig. 1a. The relationship between the inertial frame (SI ) and the block frame (Sb) is given by the angles of pitch (rotation about Yb), roll (rotation about Xb), and yaw (rotation about Zb). We can assume that there is no rotation between Sq and Sb, but Sb is translated with respect to Sq along Zq = Zb axis (see Fig. 1a). Thus, the angles of (roll=\u03c6, pitch=\u03b8, yaw=\u03c8) describe the rotation between quadrotor frame (Sq) and inertial (SI ), as well as between block frame (Sb) and inertial (SI ). The rotation matrix from Sq to SI is given by RI q (1). RI q = c\u03b8c\u03c8 s\u03c6s\u03b8c\u03c8 \u2212 c\u03c6s\u03c8 c\u03c6s\u03b8c\u03c8 + s\u03c6s\u03c8 c\u03b8s\u03c8 s\u03c6s\u03b8s\u03c8 + c\u03c6c\u03c8 c\u03c6s\u03b8s\u03c8 \u2212 s\u03c6c\u03c8 \u2212s\u03b8 s\u03c6c\u03b8 c\u03c6c\u03b8 (1) Authorized licensed use limited to: Carleton University. Downloaded on May 26,2021 at 21:19:08 UTC from IEEE Xplore. Restrictions apply. where c(\u00b7) and s (\u00b7) denote cosine (.) and sine (" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0003670_2014-01-1727-Figure1-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0003670_2014-01-1727-Figure1-1.png", + "caption": "Figure 1.", + "texts": [ + " Thus, while both flywheels have equal but opposite angular velocity the output of the differential is zero, but when the inertia's of both flywheels are inversely varied, an output angular velocity is caused by the difference of the velocities of the two flywheels. A mechanical rotational differential can be used to form a split-path power transmission, wherein an input rotational torque source can be delivered to two inputs of a differential in a manner such that one input is reversed, so that regardless of the angular velocity of the input source, the output of the differential will be zero [10, 11]. Figure 1 illustrates a differential with one output (P3 -the body of the differential) and two counter-rotating inputs. P2 is a direct-coupled input and P1 is coupled through a variablespeed belt/pulley combination. From the drive input, the P1 input can be varied from 1:1 to plus or minus about 2:1 determined by the variable belt range. The standard equation for a differential is: (1) Thus, if P1 and P2 are equal, put opposite direction, P3 will always be zero. P1 and P2 can be rotating at exceptionally high angular velocity and P3 can be accelerated from zero by varying the angular velocity of P1 or P2 or both" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0000642_s12206-020-0134-3-Figure1-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0000642_s12206-020-0134-3-Figure1-1.png", + "caption": "Fig. 1. The planar four bar RRPR linkage.", + "texts": [ + " (f) States clear instructions in order to detect successful designs of planar RRPR linkages, which are free of branch and order defects. It is expected that the foregoing features will have a positive influence on the development of an improved methodology for the kinematic design of planar RRPR linkages. Two detailed case studies are provided to validate the applicability of the proposed synthesis approach. The linkage under study is a planar four bar linkage consisting of two types of dyads, namely, the revolute-revolute dyad (RR), and the revolute-prismatic dyad (RP). These two dyads are connected as shown in Fig. 1. In Fig. 1, two fixed revolute joints are located at points A and ,D whereas a mobile revolute joint is located at point .B Moreover, a prismatic joint allows a translation of body 3 with respect to link 2 along center line CQ of the corresponding slot. Finally, it should be noted that the guided body, namely, body 4, is assumed to be rigidly attached to and move with the coupler of the linkage, namely, link 2. The input data of the problem under study is a given set of finitely separated poses of a rigid body on a plane", + " Summarizing, the synthesis process of the RP dyad with five given poses may be geometrically described in terms of the intersection of the circle ( ),, u vg Eq. (30), with the line ( )1 , ,h u v Eq. (31), being the given poses subject to satisfy constraint Eq. (40). Each obtained solution , ,u v must be substituted into Eq. (33). This yields a system of four lines in the D Dx y- plane. A solution will be valid only if there is a common intersection of these four lines. The coupler link is basically a plate with a pinhole and a slot, see Fig. 1. Point B is located at the center of the pinhole, whereas points Q and E are along the line center of the slot. The previous synthesis process showed that design parameters BQL and e can be arbitrarily defined. Moreover, without losing generality, and because of simplicity reasons, we decided that the design parameter BQL will be perpendicular to the slot of the coupler link. The foregoing discussion suggests that the final geometry of the coupler link is that shown in Fig. 5. Thus, the designer is free to choose a numerical value for design parameter ,BQL whereas the design parameter e may be computed as follows: ( )0 0 BQe L= \u00b1 - \u00d7 +d b n (41) where the double sign \u201c\u00b1\u201d allows to the designer more freedom to deal with space limitations related to the place where the designed linkage will operate" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0001314_1464419320933382-Figure5-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0001314_1464419320933382-Figure5-1.png", + "caption": "Figure 5. PTL maintenance robot: (a) three-dimensional model and (b) schematic diagram.", + "texts": [ + " Through walking and different joint actions using camera vision positioning, the robot can achieve entering or exiting operation space, nut (or bolt head) alignment and bolt tightening due to the fourarm collaboration aiming at the target object and operational suitability for the space constraints on PTL. The description of the kinematic and dynamic representation of the PTL maintenance robot belonging to a multibody system is discussed in the framework of ST with Cartesian coordinates and Euler angles. The Euler\u2013Rodrigues formula is introduced for the kinematic model of robot, and the robotic dynamics is derived by using the Newton\u2013Euler dynamic equations combined with the D\u2019Alembert\u2013 Lagrange principle of virtual work. Figure 5 shows the three-dimensional (3D) model and schematic diagram of the mechanism and coordinate frames for the PTL maintenance robot used in this investigation, respectively. The robot model consists of 13 rigid bodies, a root joint, 7 revolute joints and 5 Prismatic joints whose data are provided in Table 1. Equations (5) to (8) express the forward kinematics of the robot by introducing Euler\u2013Rodrigues formula Er(v,t) and the quaternion pj of the robot which are derived by the location vector representing the Cartesian coordinates and Roll-Pitch-Yaw (RPY) angles in the body frame Bf g" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0000408_s00707-019-02584-8-Figure2-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0000408_s00707-019-02584-8-Figure2-1.png", + "caption": "Fig. 2 Visualization of the gear pair. The bore is highlighted with bold lines", + "texts": [ + " Yet, no hex mesh could be created even after over eight hours of manual work by an expert using the standard partitioning tools of Abaqus. It is concluded that automatic or semi-automatic hex meshing is not an appropriate way for the complex gearbox housing geometry which has also been noted by many industrial users. 3.2 Validation Several simulations were performed in order to validate the simulation methods and contact algorithms. An impact of a rotating bull gear onto a pinion is simulated, as illustrated in Fig. 2. In this model, the gear motion is not constrained to a bearing, the bull gear has an initial rotational velocity of\u03c9b 0 = 10 rad/s, while the pinion is stationary. In Abaqus, the HHT time-integration scheme is used. Numerical damping can be disabled by setting the \u2018alpha\u2019 parameter to zero, see [29,30]. Additionally, in order to activate the strictest tolerance settings, the option \u2018transient fidelity\u2019 is used, see [31]. In GTM, the explicit Runge\u2013Kutta time-integration method \u2018ode45\u2019 is used with the default tolerance settings, see [32,33]" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0002869_ijhm.2018.094879-Figure1-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0002869_ijhm.2018.094879-Figure1-1.png", + "caption": "Figure 1 Schematic and coordinates of the slipper and disk", + "texts": [ + " a recess radius ratio, = R1 / R2 B pad thickness Bi parameter, /TBh \u03bb 0pc specific heat at constant pressure, = cp / cp0 Ec parameter, = (\u03c90R2)2 / (cp0T0) Eu parameter, = 6\u03bc0 / (\u03c10\u03c90H2) 1 2 3 4, , ,F F F F functions in equation (4) H representative film thickness h film thickness, = h / H hc central clearance hT heat transfer coefficient M moment load-carrying capacity, = M / (6\u03bc0\u03c90R2 5 / H2) mr, m\u03b8 parameters in equation (5) N disk rotational speed Nu parameter, = HhT / \u03bb0 p fluid pressure, = p / (6\u03bc0\u03c90R2 2 / H2) ap ambient pressure, = pa / (6\u03bc0\u03c90R2 2 / H2) Pe parameter, = PrRe rP recess pressure, = pr / (6\u03bc0\u03c90R2 2 / H2) Pr parameter, = \u03bc0cp0 / \u03bb0 ps supply pressure mQ mass flow rate, = Qm / (\u03c10\u03c90R2 3) , ,r \u03b8 z coordinate, = r / R2, \u03b8, z / h Re parameter, = \u03c10\u03c90H2 / \u03bc0 R2 pad radius T temperature, = T / T0 aT ambient temperature, = Ta / T0 rT recess temperature, = Tr / T0 wT disk temperature, = Tw / T0 tc oil temperature in chamber u velocity along r-axis, = u / (\u03c90R2) v velocity along \u03b8-axis, = v / (\u03c90R2) W load-carrying capacity, = W / (6\u03bc0\u03c90R2 4 / H2) w velocity along z-axis, = w / (\u03c90H) inclination angle of slipper, = H / R2 r restrictor parameter, = 3rc 4 / (4lcH3) \u0394t pad temperature rise \u0394tm pad mean temperature rise \u03b4t difference in temperature \u0398 circumferential coordinate \u03bb thermal conductivity, = \u03bb / \u03bb0 \u03bb thermal conductivity of pad \u03bc viscosity, = \u03bc / \u03bc0 \u03c1 density, = \u03c1 / \u03c10 dissipation function, = \u03a6 / (\u03c90R2 / H)2 azimuth angle of slipper Subscripts 0 reference; standard Figure 1 shows a hybrid thrust pad bearing modelling a slipper of swashplate type axial piston pumps and motors, in which the bearing and counter surface correspond to the test slipper and rotating disk of the experiment, respectively. The central equations including the generalised Reynolds equation, three-dimensional energy equation, and the heat conduction equation are given in the following non-dimensional forms (Kazama, 2009): 33 3 2 22 0 4 0 4 0 1 1 sin sin 1 p p FF h r F h \u03c9 h r r r r \u03b8 \u03b8 \u03b8 F F h R F hR \u03b8 \u03b8 r r \u03b8 (1) 0 2 1 p r \u03b8 T T v T T w T\u03c1 c u m m r z r \u03b8 z zh p v p T EcEcEu T T u \u03bb u r r \u03b8 Pe z Reh z (2) 22 2 2 2 2 2 1 1 0T T R Tr r r r r \u03b8 B z (3) Therein, Ec, Eu, ,Pe and Re are comparable to the Eckert, Euler, Peclet, and Reynolds numbers, respectively" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0003310_romoco.2015.7219745-Figure1-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0003310_romoco.2015.7219745-Figure1-1.png", + "caption": "Fig. 1. Nonholonomic mobile manipulator with holonomic constraint", + "texts": [ + " In Section III, dynamics reduced to the surface of holonomic geometric constraint have been introduced. Section IV contains formulation of the control problem. Control algorithm, including position and force control loop, has been proposed in Section V. Simulation results illustrating behaviour of new control law have been presented in Section VI. Conclusions and directions of future works are formulated in Section VII. Let\u2019s consider an p-DOF manipulator mounted on a nonholonomic mobile platform, see Fig. 1. A model of nonholonomic system can be obtained from d\u2019Alembert principle [7] as below Q(q)q\u0308 + C(q, q\u0307)q\u0307 +D(q) = B(q)u+ fn + fh, (1) where q \u2208 Rn+p denotes the generalized coordinates, Q(q) is (n+p)\u00d7(n+p) symmetric positive definite inertia matrix, C(q, q\u0307) denotes matrix of centripetal and Coriolis forces, D(q) \u2208 Rn+p is a vector of gravity, B(q) is an input matrix, u is a vector of control signals and fn and fh are external nonholonomic and holonomic forces acting on the system. Let the vector of generalized coordinates of the mobile manipulator be denoted as q = (qTm, q T r )T , where qm \u2208 Rn defines a vector of generalized mobile platform coordinates 978-1-4799-7043-8/15/$31" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0002059_ecce44975.2020.9236256-Figure5-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0002059_ecce44975.2020.9236256-Figure5-1.png", + "caption": "Fig. 5. No-load flux distributions of HE4 (E-core stator) at different rotor positions. (a) \u03b8e=156o. (b) \u03b8e=336o.", + "texts": [ + " VPR= EIV VDC \u00d7100% (1) In order to explain the cause of no-load induced voltage pulsation in the DC field winding of the HESFMs, the HESFM with E-core stator (HE4) is taken as an instance. HE4 has 6 field coils in the DC winding as illustrated in Fig. 1(d). As shown in Fig. 4, the flux-linkages of 6 DC field coils are variable versus the electric position (\u03b8e) of rotor and they contain both DC components and other harmonics. By way of example, the minimum value of the flux-linkage of DC field coil f1 (\u03a61) occurs at \u03b8e=156o, while the maximum value of \u03a61 occurs at \u03b8e=336o. By referring to the flux distributions illustrated in Fig. 5(a), it can be observed that the maximum value of the reluctance in the magnetic circuit of DC field coil f1 (Rm1) occurs at \u03b8e=156o, since the middle tooth of the E-core stator that accommodates f1 does not align with a rotor tooth but align with the middle of a rotor slot and hence very few fluxes pass through it. At \u03b8e=336o, the middle tooth of the Ecore stator that with f1 on it aligns with a rotor tooth and thus the majority of fluxes can flow through it, since Rm1 is minimum at this rotor position as depicted in Fig. 5(b). The variation in flux-linkage exists in other DC field coils at the same time but with different phase angles for different waveforms as depicted in Fig. 4(a). As a matter of fact, 6 DC field coils possess the same flux-linkage spectrum as shown in Fig. 4(b) and the phase displacement is 120o between two adjoining DC field coils. As a result, pulsating voltages are produced in different DC field coils as shown in Fig. 6, and likewise, they possess the same voltage spectrum and the phase displacement between two adjoining pulsating voltage waveforms is 120o" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0003805_02678292.2015.1006153-Figure4-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0003805_02678292.2015.1006153-Figure4-1.png", + "caption": "Figure 4. Definition of the coordinate system of the conical capillary used in this study. R0 is the capillary radius measured at the centre of the loop and \u03b1 is the half cone angle. The left (BP1) and right (BP2) branch points move axially at different speeds.", + "texts": [ + "[28,29] As the energy of the PR texture is higher than the PP one, the BP region moves slowly to the left and right but with different velocities. For conical capillaries, the convection of the branching region changes the disclination loop shapes, but for small cone angles one can expect that small shape changes are faster than the slow convection of the branching region. A key issue in this paper is the study of the effect of R0 and cone angle \u03b1 on the bending energy, the branching region thicknesses (XB1 and XB2), on the BP velocities (\u03c9b1 and \u03c9b2), and ovoidal disclination loop shapes. Figure 4 shows the coordinate system used in this study and a typical ovoidal disclination loop. The disclination shape is the result of the elasticity of the disclination as well as the capillary confinement. [30\u201334] The disclination is described by a planar space curve, r = r(s), where r is the position vector. The unit tangent t(s) and the unit normal N(s) to the filament are given by [35] t\u00f0s\u00de \u00bc \u2202r\u00f0s\u00de \u2202s ; \u2202t\u00f0s\u00de \u2202s ; \u22022r\u00f0s\u00de \u2202s2 \u00bc \u03ba N\u00f0s\u00de (1a; b) where \u03ba \u00bc N \u2202t=\u2202s is the curvature, quantifying the deviation from linearity" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0001949_j.mechmachtheory.2020.104139-Figure4-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0001949_j.mechmachtheory.2020.104139-Figure4-1.png", + "caption": "Fig. 4. Reference frames: (a) Conformal ball CVT, and (b) Non-conformal ball CVT.", + "texts": [ + " The reason why the spin ratio of the non-conformal ball CVT is smaller than the spin ratio of the conformal ball CVT is that r i, C is larger than r i, NC . Fig. 3 (c) shows spin ratios according to B . When B is larger, the spin ratios converge to a small constant value, while spin ratios rapidly increase as B becomes smaller. According to B , the spin ratios of non-conformal ball CVTs are smaller than the ratios of conformal ball CVTs because r i, C is larger than r i, NC . For analysis, contact between ball and disks in two ball CVTs is considered in the reference frames shown in Fig. 4 . It is set so that the force and moment direction that the ball exerts on the disk match the reference frames. The direction in which the ball inhibits the motion of the input disk and induces motion of the output disk is the direction of the force reference frames, and the direction of the moment is set based on \u03c9 ib and \u03c9 ob , indicated in red in Fig. 2 ; this is the direction of the moment that the ball exerts on the disk. Fig. 5 is a free body diagram for the input disk, output disk and one ball for each ball CVT", + " 201 \u00d7 ( \u02dc \u03c1e q x i )0 . 134 ( 1 \u2212 0 . 61 e \u22120 . 73 \u03b5 i ) (74) H o = ( h \u03c1e q x ) o = 2 . 81 [ ( 2 B \u2212 1 ) ( 1 \u2212 0 . 5 C r o ) \u0303 \u03c9 o ] 0 . 67 G 0 . 53 0 . 201 \u00d7 ( \u02dc \u03c1e q x o )0 . 134 ( 1 \u2212 0 . 61 e \u22120 . 73 \u03b5 o ) (75) The contact model lets us obtain traction coefficients and spin momentum coefficients by integrating the shear stresses over the contact area. \u03bci = \u02dc ax i \u0303 ay i 1 \u222b 0 d R 2 \u03c0\u222b 0 \u02dc \u03c4i b x Rd \u03c8 (76) \u03bco = \u02dc ax o \u0303 ay o 1 \u222b 0 d R 2 \u03c0\u222b 0 \u02dc \u03c4o b x Rd \u03c8 (77) Because the z direction of the reference frame in Fig. 4 and the direction in which the spin momentum is defined in the free body diagram in Fig. 5 are different according to each CVT and contact point, the spin momentums are defined as follows. \u03c7i,NC = \u2212\u03c7i,C = \u02dc ax i \u0303 ay i \u0303 rCV T 1 \u222b 0 d R 2 \u03c0\u222b 0 ( \u02dc ax i \u0303 \u03c4i b y cos\u03c8 \u2212 \u02dc ay i \u0303 \u03c4i b x sin\u03c8 ) R 2 d \u03c8 (78) \u03c7o,C = \u2212\u03c7o,NC = \u02dc ax o \u0303 ay o \u0303 rCV T 1 \u222b 0 d R 2 \u03c0\u222b 0 ( \u02dc ax o \u0303 \u03c4o b y cos\u03c8 \u2212 \u02dc ay o \u0303 \u03c4o b x sin\u03c8 ) R 2 d \u03c8 (79) where the following coordinate transformation rule is used { X = Rcos\u03c8 Y = Rsin\u03c8 , 0 \u2264 R \u2264 1 , 0 \u2264 \u03c8 \u2264 2 \u03c0 (80) The traction coefficients and spin momentum coefficients between ball and cavity can be obtained by a similar process" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0002728_978-3-642-03016-1_3-Figure3.8-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0002728_978-3-642-03016-1_3-Figure3.8-1.png", + "caption": "Fig. 3.8 Model of ballistic extension of the knee by providing the vertical position of unstable equilibrium. O ankle joint; A hip joint; B locked knee joint", + "texts": [ + " In an attempt to more fully explain the effectiveness of the muscles\u2019 work, we are going to now simulate a ballistic knee extension with a mathematical model. Knee stance flexion\u2013extension occurs during the stance period of gait, when the foot is in full contact with the ground, and the contralateral leg is in swing. In terms of events, we will begin simulation at the moment when knee flexion is switching to knee extension. The model is a two-link inverted pendulum with joints in the ankle and hip zones (Fig. 3.8). For that short period of time, the knee can be assumed locked by the extensor muscles, and can therefore be represented by a solid member. We will examine the conditions under which the knee can be unlocked and can start its extension. We will find the mathematical conditions of stability of the model and then interpret these conditions in biomechanical terms. Then, we are going to apply the conditions of equilibrium of the inverted pendulum with the requirements of the elastic properties of the prosthetic knee (Pitkin 1984)", + "2): c11 = [-(m1a1 + m2 (a1 + b)) g cos j1 + c1 + c2]0 = (m1a1 + m2 (a1 + b))g + c1 + c2, c22 = [-m2ga2 cos j2 + c2 ]0 = -m2ga2 + c2 , c12 = -c2 and after substituting them to (3.3), we get ( -(m1a1 + m2 (a1 + b )) g + c1 + c2 ) (-m2 ga2 + c2 ) - (- c2 )2 > 0 or m2 ga2c2 c2 > m2 ga2 . m2 ga2 + c2 c1 > (m1a1 + m2 (a1 + b))g + , (3.4) The inequalities (3.4) constitute the boundaries for the stiffness of the upper and bottom torsion springs, which guarantee the stability of dynamic equilibrium. As we discussed earlier, these stability conditions would also be the conditions for ballistic extension of the knee due to unlocking of the joint B of the model (Fig. 3.8). Let us estimate numerically the boundaries of the stiffness c1 of the lower spring. To be in a safe domain, we assume the stiffness c2 of the upper spring to be 2 times higher than that in (3.4) c2 = 2m2ga2 , and assume that a1 = a2 = b = 0.5 m; m2 = 2m1 = 400 N. That gives us c1 > 14 Nm/deg. (3.5) That estimate is almost 10 times greater than those found in experiments on standing subjects (Cavagna 1970; Fitzpatrick et al. 1992), and in experiments on passive stretching of the ankle (Sinkjaer et al. 1988). An explanation for why the higher stiffness is required in the inverted pendulum to maintain its stability when compared with the standing subject was provided in our earlier work (Pitkin 1984) and was confirmed in a recent study (Loram and Lakie 2002). The conditions (3.4) for stability of equilibrium of the pendulum shown in Fig. 3.8 were obtained without considering the movement of the foot. The attachment to the joint O from a mathematical point of view is the attachment to the ground. In the anatomical ankle, with its angulation relative to the foot, a smaller stiffness of intrinsic and triceps surae muscles maintain balance in a ballistic-like manner, as explained in Chap. 2. The estimate (3.5) should also be decreased when we are not limited by standing, but consider inertial effects of gait. As we will demonstrate later, the greater the velocity of the model, the smaller the work that needs to be performed by the moment in ankle joint to complete the step" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0001875_icra40945.2020.9197391-Figure1-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0001875_icra40945.2020.9197391-Figure1-1.png", + "caption": "Fig. 1: (a) Bernoulli-Ball system. Air is forced into an expansion chamber and through a flow straightening nozzle. A ping-pong ball hovers on the vertical air jet. (b) Schematic of dynamics modelling assumptions.", + "texts": [ + " In the case of the Bernoulli ball, the self-stabilizing properties of the system can significantly simplify the mathematical description for the purpose of control, despite the complex three dimensional body-fluid interaction. In the following 978-1-7281-7395-5/20/$31.00 \u00a92020 IEEE 3974 Authorized licensed use limited to: University of Canberra. Downloaded on October 04,2020 at 01:17:58 UTC from IEEE Xplore. Restrictions apply. sections we will outline how to model, analyze, and synthesise a controller for a self-stable system exemplified by the Bernoulli ball. The Bernoulli-Ball system\u2014as shown in Figure 1a\u2014 exhibits two forms of self-stability. Firstly, it is self-stable in the horizontal plane; when perturbed horizontally the ball returns to it\u2019s original position. This is commonly explained using Bernoulli\u2019s principle and the Coanda effect. As the speed of a fluid increases, the fluid pressure decreases. Hence, the pressure within an airflow is lower than the surrounding environment, creating a self-stabilising force about the jet centerline [18]. Secondly, it is self-stable in the vertical direction; when released into the airflow, the ball will eventually settle around a nominal height", + " We start by considering the force balance on the ball in the vertical direction, mz\u0308 = Fd(z, z\u0307) + Fb \u2212mg (1) where z is the vertical distance between the nozzle outlet and ball centre of mass, Fd(z, z\u0307) is the drag force on the ball at a height z and velocity z\u0307, Fb is the buoyant force due to the weight of displaced air and mg is the ball weight. The buoyant force is minimal, so we have Fb \u2248 0. We assume that the air-jet flow expands conically from the nozzle with a cone angle \u03b8 and that the velocity profile Vf (z) is approximately parabolic in the radial direction, as shown in Figure 1b. We define V\u0304f (z) as the mean flow velocity at a height z, with the mean velocity at the nozzle outlet being denoted by V0 = V\u0304f (0). Applying flow continuity over z yields an expression for V\u0304f (z), V\u0304f (z) = V0r 2 n (rn + z tan \u03b8)2 (2) where rn is the nozzle radius. The mean nozzle outlet velocity changes in response to a change in fan power, which is controlled using a PWM signal denoted by u. The dynamics of the fan, expansion chamber and airflow, and delays in the control PC and microcontroller introduce latency into the system, which manifests as a delay between a change in control input and a change in the ball force balance", + " Control was achieved using an electronic speed control (HobbyKing 20A ESC 3A UBEC) connected to an Arduino UNO interfaced with MATLAB. A Pulse-width modulation (PWM) signal was used to vary motor speed by changing the available current. The ball height was measured using a high frame rate webcam (Logitech BRIO) capable of delivering 90- 120 frames per second. In each frame the ball was located using a simple thresholding algorithm, after which the height was determined using the predetermined camera parameters. Figure 1a shows the setup. The controller was used to calculate the switching times zinitial = 0 m and ztarget = 0.1, 0.15, 0.2 m. The calculated policies were applied in the real system, as were the corresponding open-loop control policies for the same target heights. Figure 4 shows snapshot images of these controlled responses, while Figure 5 shows the state space controlled and open-loop responses for ztarget = 0.1, 0.15 and 0.2 m. In all cases, the open-loop response is characterized by a spiralling trajectory ending in steady-state behaviour, which manifests as an oscillation around the target height" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0001701_0954405420949757-Figure6-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0001701_0954405420949757-Figure6-1.png", + "caption": "Figure 6. Intersection calculating of auxiliary circle and tooth surface of convex side.", + "texts": [ + " The problem of solving (x0Fi, y 0 Fi, z 0 Fi) can be transformed into a numerical optimization problem as follow min f= j ffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi x0Fi 2 + y0Fi 2 q ffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi x0Fi 2 + y0Fi 2 q j+ jz0Ei z0Fij s:t: p \\ t\\ p, p \\ u1 \\ p, s1 . 0 ( \u00f019\u00de This problem is solved by particle swarm algorithm, algorithm parameters are as same as above. After coordinates of point F0i is obtained, let d 0 EF be the distance between E0i and F0i. The value range of f0i is set to [p/2, 3p/2]. As f0i changes, d 0 EF changes accordingly, when d0EF= 0, points E0i, F0i and S0i coincide, which are shown in Figure 6. The coordinates of point S0i are denoted by (x0Si, y 0 Si, z 0 Si), the problem of solving (x0Si, y 0 Si, z 0 Si) can be transformed into a single variable optimization problem as follow min f= d0EF s:t:p=2\\ f0i \\ 3p=2 \u00f020\u00de This problem is solved by adopting gold-segmentation algorithm, solving convergence condition is f\\ 10 4mm. The coordinates of point S00i are denoted by (x00Si, y 00 Si, z 00 Si), whose solving method is as same as point S0i. Intersection calculating of auxiliary circle and face cone The coordinates of point Of and A0i are denoted by (0, 0, zf) and (x0Ai, y 0 Ai, z 0 Ai)" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0000603_iccas47443.2019.8971558-Figure10-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0000603_iccas47443.2019.8971558-Figure10-1.png", + "caption": "Figure 10. End-effector orientation vector", + "texts": [], + "surrounding_texts": [ + "222\nC. Inverse kinematic solution\nAccording to given set of variables \ud835\udf03\ud835\udf03\ud835\udc65\ud835\udc65,\ud835\udc56\ud835\udc56, \ud835\udf03\ud835\udf03\ud835\udc66\ud835\udc66,\ud835\udc56\ud835\udc56 , \ud835\udf03\ud835\udf03\ud835\udc67\ud835\udc67,\ud835\udc56\ud835\udc56 (\ud835\udc56\ud835\udc56 = 1, \u22ef , \ud835\udc5b\ud835\udc5b) and \ud835\udc59\ud835\udc59\ud835\udc56\ud835\udc56 (\ud835\udc56\ud835\udc56 = 1, \u22ef , \ud835\udc5b\ud835\udc5b \u2212 1), we calculate the end-point position by Eq. (6),\n(pn 1 ) =H0,n ( 0 0 ln 1 ) (in jn kn rn 0 0 0 1 ) ( 0 0 ln 1 ) = (knln+rn 1 )\n(15)\nTaking a total differentiation of pn=knln+rn with respect to \u03b8x,i, \u03b8y,i , \u03b8z,i (i= 1, \u22ef, n) and li (i=1, \u22ef, n-1) and also motor angles \u03d5a, \u03d5b, \u03d5c, \u03d5d,\n\u2206pn= \u2202pn \u2202v \u2206v+ \u2202pn \u2202\u03d5 \u2206\u03d5 (16)\nWhere, v=(\u03b8x1,\u03b8x2, \u22ef,\u03b8xn, \u03b8y1,\u03b8y2, \u22ef,\u03b8yn ,\u03b8z1,\u03b8z2, \u22ef,\u03b8zn, l1 , l2 , \u22ef,ln-1 )\n\u2208R4n-1 and \u03d5=(\u03d5a, \u03d5b, \u03d5c, \u03d5d ). \u2202pn \u2202v \u2208R3\u00d74n-1 and \u2202pn \u2202\u03d5 \u2208R3\u00d74. Whereas, let w=(w1, w2, \u22efw4n-1)T=04n-1 represents the 4n-1 equations provided by Eqs. (9)(10)(12)(13), which also includes \u03b8x,i, \u03b8y,i , \u03b8z,i (i=1, \u22ef, n) , li (i=1, \u22ef, n-1) and also motor angles \u03d5a, \u03d5b, \u03d5c, \u03d5d. Taking a total differentiation for w=04n-1 as well, we have,\n\u2206w= \u2202w \u2202v \u2206v+ \u2202w \u2202\u03d5 \u2206\u03d5=04n-1 (17)\nwhere, \u2202w \u2202v \u2208R(4n-1)\u00d7(4n-1) and \u2202w \u2202\u03d5 \u2208R(4n-1)\u00d74 . Since \u2202w \u2202v is a square matrix, we can solve (19) with respect to the vector \u2206\ud835\udc63\ud835\udc63 as,\n\u2206v=- (\u2202w\n\u2202v ) -1 \u2202w \u2202\u03d5 \u2206\u03d5 (18)\nSubstituting (18) into (16), we have\n\u2206pn=\u2202pn \u2202v (\u2202w \u2202v ) -1 \u2202w \u2202\u03d5 \u2206\u03d5+ \u2202pn \u2202\u03d5 \u2206\u03d5=\n( \u2202pn \u2202\u03d5 - \u2202pn \u2202v (\u2202w \u2202v ) -1 \u2202w \u2202\u03d5 )\u2206\u03d5=J\u2206\u03d5\n, which can be solved for \u2206\ud835\udf19\ud835\udf19, by using a generalized inverse of the Jacobian J\u2208R3\u00d74\n\u2206\u03d5=J\u2020\u2206pn+P\u22a5(J)\u03a8 (19)\nwhere, J\u2020\u2208R4\u00d73 is a generalized inverse of \ud835\udc71\ud835\udc71 and P\u22a5(J)\u2208R4\u00d74 is a null projection operator of J, and \u2206\u03d5N\u2208R4 is a correction of \ud835\udf53\ud835\udf53 so as to minimize a positive scalar potential \ud835\udf11\ud835\udf11 by making use of a redundant actuation.\nWe use J\u2020=JT(J JT)-1 and P\u22a5(J)=I-J\u2020J. Eq.(19) provides a variation of motor angles \u2206\ud835\udf53\ud835\udf53 for a given position and direction variation \u2206pn . Applying the Euler method, we have the following variational equation, \u03c6+(\u2202\u03c6/\u2202\u03d5)\u2206\u03d5N=0\n,which is solved by\n\u2206\u03d5N=\u03c6 (\u2202\u03c6/\u2202\u03d5)(\u2202\u03c6/\u2202\u03d5)T (\u2202\u03c6 \u2202\u03d5)\nT\n( 20) As a candidate of \u03c6 , we take \u03c6=knz\n2 , where, k\ud835\udc5b\ud835\udc5b\ud835\udc67\ud835\udc67 is the z component of kn: the unit vector of the end-point orienting an axial direction. It means that the axial direction the endpoint takes on a horizontal plain as far as possible while keeping a designated position.\nAuthorized licensed use limited to: Auckland University of Technology. Downloaded on October 25,2020 at 08:37:21 UTC from IEEE Xplore. Restrictions apply.", + "223\nIV. EXPERIMENTS AND SIMULATION TakoBot 2 kinematics computed by Wolfram Mathematica software. As a robot controller, Arduino UNO was utilized, and for stepping motors, we used SilentStick TMC2208 motor drivers to provide sufficient power to the motors. During the experiments, manipulator demonstrated basic motions and helical motion in both directions. According to the conducted experiments, passive sliding mechanism increased robot bending stress tolerance and acted smoother than the previous prototype. Likewise, the mechanism discovered new horizons for continuum robot applications and imagination about continuum manipulators.\nV. REFERENCES\n[1] A.Yeshmukhametov, K. Koganezawa, Y. Yamamoto, \u201cDesign and kinematics of cable driven continuum robot arm with universal joint backbone\u201d,IEEE ROBIO 2018 conference proceeding, Kuala \u2013Lumpur, Malaysia. 2018. [2] V.C Anderson, R.C. Horn: Tensor Arm Manipulator Design, Mech. Eng 8998), 54-65. 1967 [3] R. Buckingham, A. Graham, \u201cNuclear snake-arm robots\u201d, Industrial Robot: An International Journal, Vol.39 Issue: 1,pp 6-11, https://doi,org/10.1108/01439911211192448, 2012. [4] M.W. Hannan and I.D.Walker, \u201cKinematics and the\nImplementation of an Elephant\u2019s Trunk Manipulator and other Continuum Style Robots\u201d, Journal of Field Robotics, pp 45-63,February 2003. https://doi.org/10.1002/rob.10070\n[5] Graham, R., Bostelman, R., \u201cDevelopment of EMMA Prototype suspended from the 6 m RoboCrane Prototype.\u201d Proc. ANS Seventh Topical Meeting on Robotics and Remote Systems,\u201d Augusta, GA, April 27-May 1, 1997. [6] I.A. Gravagne ; C.D. Rahn ; I.D. Walker, \u201cLarge deflection dynamics and control for planar continuum robots\u201d, IEEE/ASME Transactions on Mechatronics, Volume: 8 , Issue: 2 , June 2003. [7] Camarillo, D., Milne, C., Carlson, C., Zinn, M., Salisbury, J.: Mechanics modeling of tendon driven continuum manipulators. IEEE Transactions on Robotics (accepted for publication) (2008) [8] S. Neppalli, B. Jones, W. McMahan, V. Chitrakaran, I. Walker M. Pritts, M. Csencsits, C. Rahn, M. Grissom, \u201cOctArm - A Soft Robotic Manipulator\u201d, Proceedings of the 2007 IEEE/RSJ International Conference on Intelligent Robots and Systems San Diego, CA, USA, Oct 29 - Nov 2, 2007 [9] Jones, B., Walker, I.: Kinematics for multisection continuum robots. IEEE Transactions on Robotics 22(1), 43\u201355 (2006) [10] Haitham E., Usman M., Zain M.,Sang-Goo J., Muhammad U., Elliot W.H., Alisson M.O.,Jee-Hwan R., \u201cDevelopment and Evaluation of an Intuitive Flexible Interface for Teleoperating Soft Growing robots\u201d, 2018 IEEE/RSJ IROS, Madrid,Spain, October 1-5,2018. [11] A. Benouhiba, K. Rabenorosoa, P. Rougeot, M. Ouisse, N. Andreff, \u201c A multisection Electro-active Polymer based milli-Continuum Soft Robots. 2018 IEEE/RSJ IROS, Madrid,Spain, October 1-5,2018. [12] B. Ouyang, H. Mo, H. Chen, Y. Liu, D. Sun, \u201cRobust Model-Predictive Deformation Control of a Soft Object by Using a Flexible Continuum Robot, 2018 IEEE/RSJ IROS, Madrid,Spain, October 1-5,2018. [13] R. Kang, D.T. Branson, T. Zheng, E. Guglielmino, D.G. Caldwell, \u201cDesign, modelling and control of a pneumatically actuated manipulator inspired by biological continuum structures\u201d, Bioinspiration&Biomimetics 8 (2013) 14pp. [14] Bryan A. Jones, Ian D. Walker, \u201cA New Approach to Jacobian Formulation for a Class of Multi-Section Continuum Robots\u201d, Proceedings of the 2005 IEEE, ICRA, Barcelona, Spain, April 2005. [15] Han Yuan and Zheng Li, \u201cWorkspace analysis of cabledriven continuum manipulators based on static model\u201d, Robotics and Computer \u2013 Integrated Manufacturing. 49 (2018) 240-252. [16] X. Dong, M. Raffles, S. Gobos-Gusman, D. Axiente, J. Kell, \u201c A Novel Continuum Robot Using Twin-Pivot Compliant Joints: Design, Modelling and Validation, Journal of mechanism and robotics, February 16, 2015.\n[17] T. Liu, Z. Mu, H. Wang, W. Xu, Y. Li, \u201c A Cable-Driven Redundant Spatial Manipulator with Improved Stiffness and Load Capacity\u201d, 2018 IEEE/RSJ IROS, Madrid, Spain, October 1-5,2018\n[18] J. Starke, E. Amanov, M.Taha Chihaoui, J. Burgner-Kahrs, \u201cOn the Merits of Helical Tendon Routing in Continuum Robots\u201d, 2018 IEEE/RSJ IROS, Vancouver, BC, Canada, September 24-28, 2017.\nAuthorized licensed use limited to: Auckland University of Technology. Downloaded on October 25,2020 at 08:37:21 UTC from IEEE Xplore. Restrictions apply." + ] + }, + { + "image_filename": "designv11_71_0000191_24748668.2019.1697580-Figure2-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0000191_24748668.2019.1697580-Figure2-1.png", + "caption": "Figure 2. Definitions of angles and hip height.", + "texts": [ + " The segments were head, trunk, upper arms, forearms, hand, hip, thighs, shanks, and foot. The right shoulder, right elbow, hip, right knee, and left knee joint angles were measured. The joint angles and javelin projection angles were measured at the time point of javelin initial release. The release velocity of the wrist was calculated as the velocity from the time point of the javelin\u2019s initial release to the release (two frames of throwing image). The definitions of joint angles and javelin projection angles are shown in Figure 2, and all these angles were observed within the sagittal plane. The shoulder joint angle was defined as the angle from the trunk to the upper arm. The elbow joint angle was defined as the angle from the forearm to the upper arm. The trunk was defined as a vector from the hip to shoulder, while the trunk inclination angle was the vector relative to the perpendicular line. The knee joint angle was defined as the angle from the thigh to the shank. The javelin angle was defined as the projection angle relative to the horizontal" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0000987_0954406220915499-Figure4-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0000987_0954406220915499-Figure4-1.png", + "caption": "Figure 4. The dual-axle steering mechanism shimmy system. 1\u2014steering wheel, 2\u2014steering gear, 3\u2014first rocker, 4\u2014anterior longitudinal rod, 5\u2014steering knuckle arm, 6\u2014first transition rod, 7\u2014intermediate rocker, 8\u2014tie rod, 9\u2014second transition rod, 10\u2014second rocker, 11\u2014steering power cylinder, 12\u2014posterior longitudinal rod, 13\u2014tire, 14\u2014left trapezoid arm, and 15\u2014right trapezoid arm.", + "texts": [ + " The caster angle can ensure a stable self-righting performance under high-speed driving conditions; The steering axis inclination angle can ensure a stable self-righting performance under low speed and heavy load conditions. The function of the camber angle is to prevent the wheel from tilting inward under heavy loads. The reason for designing the toe-in angle is to reduce the abnormal uneven wear of the tire due to the camber angle. The mechanics model of dual-axle steering mechanism shimmy system is shown in Figure 4. In the mechanics model, the elastic of the rod is considered; it is equivalent to a spring\u2013damper unit. In addition, the split road adhesion coefficients and the dry friction between the suspension and steering system are also considered when modeling. In line with the law of the right-hand coordinate system, the center of the mass of the vehicle is deemed as the coordinate origin; the vehicle forward direction is for X-axis, the vehicle left direction is for Y-axis, and perpendicular to the ground up direction is for Z-axis. According to Figure 4, the following assumptions are made when establishing the mechanical model of the dual-axle steering mechanism: 1. The steering wheel is immobile; 2. The impact of the force of air is ignored; 3. The various parts associated with the vibration and their couplings are simplified according to the moments of inertia, springs, and dampers; 4. The angle between the steering trapezoid plane and XY plane and the angle between the steering linkage and the XZ plane are ignored; 5. The direction and speed of the vehicle is constant; 6" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0003780_acc.2014.6859036-Figure1-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0003780_acc.2014.6859036-Figure1-1.png", + "caption": "Fig. 1: Geometry of the guidance problem", + "texts": [ + " The desired trajectory is defined as an arbitrary sequence of waypoints without any geometric requirement on the path between two successive waypoints. A waypoint position is given in 3D Cartesian-Space coordinates (Xw, Yw, Hw) in a Earth fixed reference frame, and the desired heading angle (\u03c7w) is used to obtain a preferred approach direction. Each waypoint defines a coordinate 978-1-4799-3274-0/$31.00 \u00a92014 AACC 5254 frame centered in the waypoint position and rotated by the angle \u03c7w + (\u03c0/2) around the Z-axis with respect to the earth fixed frame, as is shown in Fig. 1. We assume, as common in these cases, that a waypoint manager exists, and it checks the aircraft position and updates the desired waypoint when the previous one has been reached within a given tolerance. For the purposes of this paper we will concentrate on horizontal plane guidance, thus an autopilot capable of tracking a desired altitude reference (i.e.Hw) is also assumed available. The desired approach trajectory to the waypoint is achieved via the heading fuzzy controller f\u03c7, which generates the desired heading angle \u03c7d using the position errors along the X and Y axes of the current waypoint reference frame [ewX , e w Y ]: \u03c7d = f\u03c7(e w X , e w Y , S) with [ewX , e w Y ] T = Rz(\u03c7w + (\u03c0/2)) \u00b7 [eX , eY ]T (1) Where eX = Xw \u2212 X and eY = Yw \u2212 Y , with (X,Y ) the position of the vehicle in inertial coordinates in a local level frame" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0000886_012117-Figure2-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0000886_012117-Figure2-1.png", + "caption": "Figure 2. Deviation of the position of the spatial three-link manipulation system", + "texts": [ + " When analyzing the convergence process, the average value of fitness function by population (Mean fitness) is often used. As a fitness function for solving inverse kinematics problem of manipulating robots, we will use a function of calculating the square of position deviation of the robot\u2019s output link characteristic point from selected point on a motion path. We use the genetic algorithm to search for values of hinge coordinates minimizing fitness function using example of a flat two-link (figure 1) and three-link spatial manipulation systems (figure 2). For a two-link manipulation robot, the chromosome will be a vector X=[x1 x2]T which elements are the hinged coordinates, x1 \u2013 is rotational, \u0430 x2 \u2013 is translational. For these coordinates, we introduce restrictions , . We make up a fitness function where Y=[y1 y2]T \u2013 vector of Cartesian coordinates of point selected on the trajectory. We set coordinates of selected point y1= 1.0000 m \u2013 rectangular coordinate along x axis, y2= 0.0000 m \u2013 rectangular coordinate along y axis. We set the population size at 10 individuals" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0002816_thc-2010-0566-Figure13-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0002816_thc-2010-0566-Figure13-1.png", + "caption": "Fig. 13. Angle of twist due to torsion.", + "texts": [ + " The maximum shear stress at the hollow tube surface with outside radius, ro, and inside radius, ri, is calculated using Eq. (7). For the same applied torque and outside tube diameter, as the inside tube diameter increases the maximum induces shear stress at the outside surface increases. \u03c4max = 2Tro \u03c0(r4o \u2212 r4i ) (8) If we know the material properties then we can calculate the shear strain at each point, this shear strain cumulatively results in twisting of the cylinder. The angle of twist, \u03b8 (theta), (Fig. 13) for a cylinder made of material with modulus of rigidity G and length L experiencing an external torque of, T , can be calculated using Eq. (8). The longer and thinner a cylinder greater is the angle of twist. \u03b8 = 2TL \u03c0Gr4 (9) If a square cross-section beam is fixed at one end and a force is applied at the other end perpendicular to its axis then there is a induced bending moment M in the bar. A bending moment is the applied force multiplied by the distance of the force from the fixed end (M = F" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0000798_icaee47123.2019.9015101-Figure8-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0000798_icaee47123.2019.9015101-Figure8-1.png", + "caption": "Fig. 8. Simulated spectral evolutions for (a) Gaussian and (b) CS pulses through PCF3.", + "texts": [ + " Nevertheless as we will show in the discussion section, the calculated SC spectral flatness with the CS pulse is the highest among the other three input pulses. In this regard, the highly-structured spectrum in the AD region formed with the CDS input pulse seems to be detrimental in terms of spectral flatness. Thus now we focus on the CS pulse as the most eligible amongst the four pulse types for forming balanced SC spectra through PCF3. In addition, to explicitly analyze the advantage of the spectral shape of the CS pulse over the cases of the Gaussian or the soliton pulses, we plot the spectral evolutions for the Gaussian and CS pulses in Fig. 8. In the case of the Gaussian pulse, the power spectral density distribution has a wide spectral dip with a regular fringe pattern across the double ZDWs, having transferred most of the pulse energy to the side spectra centered at 900 and 1200 nm. In comparison, such effects are relatively diminished in the case of the CS pulse, having a decent level of new spectral components arising in the middle of the split spectra as well as producing a spectral BW as large as that of the Gaussian pulse. As discussed just before, these new spectral components are created from the SPM-induced spectral broadening of the RSPs of the CS pulse" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0003981_s0005117915020058-Figure2-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0003981_s0005117915020058-Figure2-1.png", + "caption": "Fig. 2. A single-link robot manipulator.", + "texts": [ + " Under assumption (4.5) we have upper bounds for the functional (4.13): J0(u) tr (\u03a3X(0)) \u03bc0 \u03bbmax(X(0)), \u03a3 = \u222b S0 \u03bc(x0)x0x T 0 dx0, \u03bc0 = \u222b S0 \u03bc(x0)\u2016x0\u20162 dx0. Therefore in the formulated optimization problems (1)\u2013(3) instead of the first condition of (4.4) we can use inequalities tr (\u03a3X(0)) \u03c9 or \u03bc0 \u03bbmax(X(0)) \u03c9. Example. Consider a control system for a single-link robot manipulator whose link\u2019s circular motion from one end to another is done with a flexible connection of the link and the executive mechanism (Fig. 2). AUTOMATION AND REMOTE CONTROL Vol. 76 No. 2 2015 A linear torsion spring is located between the executive mechanism and the end of a link. This system is defined with two nonlinear differential equations of order two that follow from the mechanical balance of the executive mechanism (motor shaft) and the manipulator link discarding the friction and external disturbances, or, in vector-matrix form [12], x\u0307 = A(x)x+Bu, (4.14) where A(x) = \u23a1 \u23a2 \u23a2 \u23a2 \u23a2 \u23a3 0 1 0 0 \u2212(\u03bcgh\u03d5(\u03b81) + k)/J1 0 k/J1 0 0 0 0 1 k/J2 0 \u2212k/J2 \u2212d/J2 \u23a4 \u23a5 \u23a5 \u23a5 \u23a5 \u23a6 , B = \u23a1 \u23a2 \u23a2 \u23a2 \u23a2 \u23a3 0 0 0 1/J2 \u23a4 \u23a5 \u23a5 \u23a5 \u23a5 \u23a6 , x = \u23a1 \u23a2 \u23a2 \u23a2 \u23a2\u23a2 \u23a3 \u03b81 \u03b8\u03071 \u03b82 \u03b8\u03072 \u23a4 \u23a5 \u23a5 \u23a5 \u23a5\u23a5 \u23a6 , where \u03b81 and \u03b82 are angular coordinates of the manipulator link and motor shaft respectively, u is the controlling moment produced by the electric drive, J1 and J2 are moments of inertia respectively for the manipulator link and the electric drive, k is the rigidity of the transmission gear, d is the damping coefficient, \u03bc is the manipulator link\u2019s mass, h is the manipulator link\u2019s length, g is the gravitational acceleration, and \u03d5(\u03b8) = (sin \u03b8)/\u03b8 is a continuous function" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0002604_j.ifacol.2020.12.1397-Figure1-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0002604_j.ifacol.2020.12.1397-Figure1-1.png", + "caption": "Fig. 1. The Body Frame and Inertial Frame", + "texts": [ + " Our contributions in this paper are that we have derived the quadcopter nonlinear model with least small signal approximations and assumptions in section (2). Feedback linearization based control of the derived model has been designed in section (3). Finally the simulation results of the proposed control scheme and future work has been discussed in section (??) and (7) respectively. 2. QUADCOPTER MODELLING For quadcopter modelling we have considered three coordinate frames that are inertial, vehicle and body frame. which have been explained in the Fig. 1 and Fig. 2 The state vector is defined as X =\u2206 [ x y z \u03c6 \u03b8 \u03c8 x\u0307 y\u0307 z\u0307 \u03c6\u0307 \u03b8\u0307 \u03c8\u0307 ] Where x, y and z are the components of position vector of Ob in inertial frame along i\u0302i, j\u0302i and k\u0302i respectively as Control System Analysis and Design of Quadcopter in the Presence of Unmodelled Dynamics and Disturbances Muhammad Z. Ali \u2217 Aftab Ahmed \u2217\u2217 Hamad K. Afridi \u2217\u2217\u2217 \u2217 Pakistan Institute of Engineering and Applied Sciences, Islamabad, Pakistan (e-mail: msse1714@pieas.edu.pk). \u2217\u2217 Pakistan Institute of Engineering and Applied Sciences, Isla abad, Pakistan(e-mail: aftab@pieas", + " Our contributions in this paper are that we have derived the quadcopter nonlinear model with least small signal approximations and assumptions in section (2). Feedback linearization based control of the derived model has been designed in section (3). Finally the simulation results of the proposed control scheme and future work has been discussed in section (??) and (7) respectively. 2. QUADCOPTER MODELLING For quadcopter modelling we have considered three coordinate frames that are inertial, vehicle and body frame. which have been explained in the Fig. 1 and Fig. 2 The state vector is defined as X =\u2206 [ x y z \u03c6 \u03b8 \u03c8 x\u0307 y\u0307 z\u0307 \u03c6\u0307 \u03b8\u0307 \u03c8\u0307 ] Where x, y and z are the components of position vector of Ob in inertial frame along i\u0302i, j\u0302i and k\u0302i respectively as Copyright \u00a9 2020 The Authors. This is an open access article under the CC BY-NC-ND license (http://creativecommons.org/licenses/by-nc-nd/4.0) shown in the Fig. 1. And \u03c6, \u03b8 and \u03c8 are the Euler angles that define the transformation from vehicle frame to body frame as expressed in Fig. 2. Oi is the origin of inertial frame located at the ground control station where Ob is the origin of vehicle frame and body frame located at the center of mass of the quadcopter. It should be noted that inertial frame and vehicle frame have parallel coordinate axis only the origin is located elsewhere. We have derived the expression for the transformation from vehicle to body frame" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0000058_icmae.2019.8881011-Figure1-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0000058_icmae.2019.8881011-Figure1-1.png", + "caption": "Figure 1. Scheme of Hunt type platform", + "texts": [ + " Section 2 presents the direct kinematics and Differential Kinematic Control of the same robot, obtained from Newton Raphson's method of successive approximations. Finally, in section 3, conclude on the models made, their configurations and the results obtained. II. GEOMETRIC ANALYSIS OF HUNT PLATFORM In order to obtain the kinematic model of a parallel robot type RUS, it is necessary to analyze the structure and geometry to know the mobility of the robot and the spatial relationships of the elements that compose it. The diagram of the parallel robot analyzed is shown in Figure 1 and consists of a base defined by an irregular hexagon and a similar platform that are connected through six arms of two links: inlet and coupling, with lengths of and , respectively. The geometry of hexagons is defined by two magnitudes: the dimensions of the major edge and the minor edge. Defining a reference system located in the geometric center 978-1-7281-5535-7/19/$31.00 \u00a92019 IEEE of the hexagon, the six vertices are located as shown in Table I, considering that the lengths of the major and minor edges are denoted as and , respectively; and thus, forming a vector of positions denoted as whose in the 6 points is at zero" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0002119_icaccm50413.2020.9212874-Figure3-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0002119_icaccm50413.2020.9212874-Figure3-1.png", + "caption": "Fig. 3. Mesh model of the propeller.", + "texts": [], + "surrounding_texts": [ + "The propellers are subject to bending in the vertical direction due to applied thrust force F. To analyze the propeller under thrust the propeller is fixed at the point where it connected to the motor shaft as shows in Figure 4. 1 2 3 4F mg F F F F (1) 2 i t iF K (2) Where m is the total mass (Quadcopter and payload) and g is acceleration due to gravity. Kt is the coefficients, depends upon propellers geometry and air density. If the value of m is 20 kg then the value of the total thrust force is calculated as follows. Total Thrust force (F)= mg= 20*9.81=196.2 N Thrust force required per propeller= F/4 =196.2/4 =49.05 N The calculated thrust force is approximately 50 N for a single propeller. Thus, the propeller is subjected to 50 N of thrust force to find out the deformation and stress due to the bending effect as shown in Figure 5. Figure 6, 7 and 8, 9 show the total deformation and stress for CFRP and GFRP materials, respectively. Table II presents the comparison of results for CFRP and GFRP. 60 Authorized licensed use limited to: Carleton University. Downloaded on May 30,2021 at 15:29:13 UTC from IEEE Xplore. Restrictions apply." + ] + }, + { + "image_filename": "designv11_71_0000090_s38313-019-0145-6-Figure6-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0000090_s38313-019-0145-6-Figure6-1.png", + "caption": "FIGURE 6 Safety factors of the hybrid cylinder crankcase (prototype) (\u00a9 Volkswagen)", + "texts": [ + " The concept shows a high number of areas with low safety factors. Especially the second main bearing, the inner and outer walls of the ventilation window and partly the top deck are affected, FIGURE 5. An additional calculation of the contact pressure indicates that displacements of the plastic within the fusion area are likely. Particularly the sealing zone in the lower part of the water jacket is affected. The reinforcements of the prototype significantly reduce the number of critical areas, FIGURE 6. Nevertheless there is still optimization potential regarding the ventilation window of the first and second main bearing and some spots in the top deck. The newly designed form-locking structures for the metal/plastic fusion on the main bearings also need to be improved. After the calculation of the acoustics, the temperature and the strength, ecological aspects have to be considered as a further important criterion. Below, the global warming potential of the hybrid cylinder crankcase is estimated" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0002083_niles50944.2020.9257973-Figure6-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0002083_niles50944.2020.9257973-Figure6-1.png", + "caption": "Fig. 6. Inertial and Body Frames", + "texts": [ + " Here the repulsion force represented by equation 3 is used twice, first for calculating the repulsion field force for the environment obstacles (defined here as the walls of the building) the value of \u03b7 is 850 and second for calculating the force estimated for each agent when assumed to be an obstacle the value of \u03b7 is 18. The attraction force represented by equation 2 is used to calculate the attraction field force for the goals of the quadrotors, where each goal force affects its agent, the value of \u03be is 1/2100. To illustrate the 6DOF motion of the quadrotor two frames are defined, the earth inertial frame (E-frame) and the bodyfixed frame (B-frame) shown in Figure 6. Assuming the state\u2019s vector in the following way: 499 Authorized licensed use limited to: UNIVERSITY OF CONNECTICUT. Downloaded on May 17,2021 at 18:41:15 UTC from IEEE Xplore. Restrictions apply. x = [ \u03c6 \u03b8 \u03c8 p q r u v w x y z ]T \u2208 R12 (5) Eventually, the dynamic model of the quadrotor in the inertial frame is: x\u0308 = \u2212 ftm [sin(\u03c6)sin(\u03c8) + cos(\u03c6)cos(\u03c8)sin(\u03b8)] y\u0308 = \u2212 ftm [cos(\u03c6)sin(\u03c8)sin(\u03b8)\u2212 cos(\u03c8)sin(\u03c6)] z\u0308 = g \u2212 ft m [cos(\u03c6)cos(\u03b8)] \u03c6\u0308 = Iy\u2212Iz Ix \u03b8\u0307\u03c8\u0307 + \u03c4x Ix \u03b8\u0308 = Iz\u2212Ix Iy \u03c6\u0307\u03c8\u0307 + \u03c4y Iy \u03c8\u0308 = Ix\u2212Iy Iz \u03c6\u0307\u03b8\u0307 + \u03c4z Iz (6) The parameters\u2019 values of the quadrotor used in this paper are in equation 7" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0002989_978-3-642-16214-5_9-Figure9.25-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0002989_978-3-642-16214-5_9-Figure9.25-1.png", + "caption": "Fig. 9.25. Mercedes-Benz external-cone synchronizer system. 1 Idler gear with dog gearing; 2 annular spring; 3 synchronizer ring; 4 synchronizer body; 5 gearshift sleeve; 6 locking lug; 7 groove in idler gear", + "texts": [ + " The counter-cone ring 4 is rotationally fixed to the synchronizer body 5. The gearshift effort is reduced and the torque capacity, i.e. performance, is increased because of the increased number of friction surfaces and the larger friction surface area of the double-cone synchronizer. The parallel multi-cone design requires closer manufacturing tolerances and therefore entails higher production costs. Multi-cone synchronizers are therefore used only in the lower gears. In the Mercedes-Benz external-cone synchronizer system (Figure 9.25) the synchronizer unit 3 is fixed by an annular spring 2 to the idler gear 1. The synchronizer ring has three inward-facing locking lugs 6, which engage in corresponding grooves 7 in the idler gear. It can turn relative to the wheel circumferentially a certain distance and axially once the annular spring has been overcome. When shifting, the gearshift sleeve is pressed against the synchronizer ring. The friction torque turns the synchronizer ring as far as it will go. Its locking lugs 6 are then so placed before a bevel in the idler gear that the gearshift sleeve 5 and synchronizer ring 3 can move no further as long as the friction torque is not equal to zero" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0000576_icpads47876.2019.00090-Figure6-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0000576_icpads47876.2019.00090-Figure6-1.png", + "caption": "Fig. 6. (a) Substrate without ridged quantum well. (b) Substrate with ridged quantum well. \u03c6e is the emitter work function. EF is the emitter Fermi energy level [11].", + "texts": [ + " All of these studies demonstrate potential TEC electrode materials. Surface nanostructuring of the thin metal ridged quantum-well layer coating on the emitter is one of the methods proven in work function reduction [11]. Additional boundary conditions imposed on the electron wave function at these periodic ridges prevent free electrons from filling into it. As a result, these electrons have to occupy high energy states, increasing the Fermi energy and reducing the work function of the emitter, as depicted in Fig. 6. A semiconductor substrate with a wide bandgap, e.g., silicon, is preferred as the emitter in this method, as it allows for greater electron confinement and a significant work function reduction. Alkali metals are well known for their low work function characteristics, and hence, their unique behavior has been intensively studied. According to Langmuir [12], alkali atom adsorption on metal or semiconductor surfaces can reduce the work function. When alkali metal atoms with very low electronegativity values fall onto the substrate surface, their valence electrons tend to transfer to the substrate surface" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0000846_eiconrus49466.2020.9039424-Figure1-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0000846_eiconrus49466.2020.9039424-Figure1-1.png", + "caption": "Fig. 1 OWPMSM motor equivalent circuit", + "texts": [ + " Topology with voltage source and floating bridge capacitor achieves less speed increasing level but more useful in regeneration mode [11-12]. This paper represents modelling of OWPMSM motor topology with two inverters and two independent voltage sources (traction batteries) mounted on electric vehicle powertrain in order to get presented topology\u2019s dynamic performance. Motor parameters fluctuations, unsymmetrical states and faults conditions are out of presented thesis scope [13]. Equivalent circuit of OWPMSM showed on Fig.1. Generally, there is no difference with PMSM with stator star secondary winding connection. PMSM motor equations Voltage equations are given by: + \u22c5 +\u2212 + = f fr d q dsqr drqs d q i i LRL LLR V V \u03c1\u03bb \u03bb\u03c9 \u03c1\u03c9 \u03c9\u03c1 (1) where dV \u2013 d-axis voltage (V); qV \u2013 q-axis voltage(V); sR \u2013 stator resistance (Ohm); dL \u2013 d-axis self-inductance (H); qL \u2013 q-axis self-inductance (H); r\u03c9 \u2013 electrical speed (RPM); f\u03bb \u2013 field flux linkage (Wb); \u03c1 \u2013 derivative operator; di \u2013 d-axis current (A); qi \u2013 q-axis current (A)" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0003102_978-1-4419-1126-1_16-Figure16.2-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0003102_978-1-4419-1126-1_16-Figure16.2-1.png", + "caption": "Fig. 16.2 Conceptual assembly drawing of a swimming micro robot for neurosurgery. The components of the robot are: (1) Three swimming tails, (2) Power source (here 3 Renata ZA10 batteries in series), (3) Packaged IC for command control and communication, (4) Antenna, (5) Endoscopic Camera, (6) LEDs, (7) Force sensors, (8) Tool for intervention, (9) Localization sensor (here an Aurora magnetic tracker receiving coil), (10) Casing (Transparent in the drawing to show the internal components) (Copyright 2008 IEEE)", + "texts": [ + " The robot has a capsular form and when it is released from the cannula three swimming tail AP actuators pop out to enable maneuvering in five DOF. This swimming robot has a diameter of \u00d86.6 mm (determined mostly by the diameter of the batteries) and a length of 31 mm which results in a net volume of 1.1 cm3. The propulsive units\u2019 angle with the axis of the body is 30 and the outer diameter of the robot with open swimming tails is \u00d826 mm. The different components of this robot match the categories of Sect. 16.2.1\u201316.2.6 in this fashion (See Fig. 16.2): 1. AP \u2013 Three flagellar piezoelectric swimming tails, (for detailed description of their operation method see Sect. 16.3). 2. PS \u2013 Power source made of batteries or magnetic coils for RF induction. 3. CU \u2013 Custom designed integrated circuits (IC) for command and control. The main tasks of the CU here is process the data of the SU, convert it to a proper communication protocol, generate the actuation signals for the AP according to external command. 4. CT \u2013 IC for communication and an antenna that is able to transmit the signal out of the body" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0003788_detc2014-34213-Figure20-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0003788_detc2014-34213-Figure20-1.png", + "caption": "Fig. 20 Eight-bar linkages with three prismatic joints", + "texts": [ + " Based on the criteria, a three-DOF planar linkage with prismatic joints can also be decomposed into a corresponding two-DOF planar linkage and an additional input joint or a corresponding single-DOF planar linkage and two additional input joints. For example, the 6 Copyright \u00a9 2014 by ASME Downloaded From: http://proceedings.asmedigitalcollection.asme.org/pdfaccess.ashx?url=/data/conferences/asmep/82119/ on 04/05/2017 Terms of Use: http://www.asme.org/about-asme/terms-of-use three-DOF eight-bar linkage with the prismatic joints in Fig. 20 can be decomposed into a two-DOF seven-bar linkage with the same prismatic joints and an additional input joint or a corresponding a single-DOF six-bar linkage and two input joints. The singularity of the two-DOF planar linkage with prismatic joints is discussed in Section 3. For example, when the joints A0, B0 and C0 in Fig. 14(a) are replaced with three prismatic joints, such as the prismatic joints A, K and C in the Fig.20(a), the three-DOF eight-bar planar linkage can be decomposed into a single-DOF six-bar linkage with the prismatic joints A, K, C and the additional joints E and B, as shown in Fig. 20(b). Hence, the singular condition for the three-DOF eight-bar linkage is the same as the single-DOF six-bar linkage with the prismatic joints in Fig. 2. As discussed above, whether the N-DOF(N\u22652) planar linkage contains the prismatic joints, the N-DOF planar linkage can be decomposed into a corresponding (N-1)-DOF planar linkage and an additional joint or (N-2)-DOF planar linkage and two additional joints. Similarly, an N-DOF planar linkage can be decomposed into (N-M)-DOF planar linkage and M additional joints, the singularity of N-DOF planar linkage happens when the singularity of (N-M)-DOF planar linkage happens" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0001717_icmae50897.2020.9178892-Figure4-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0001717_icmae50897.2020.9178892-Figure4-1.png", + "caption": "Figure 4. 3D models of the tested samples with their basic dimensions.", + "texts": [ + " A simple Body-Centred Cubic (BCC) unit cell sizes 5x5x5 mm with strengthened linear edges and diameter of a strut 1 mm was selected for this research. The basic cell has been regularly patterned inside the cylindrical shape. The second type of the samples was the same shape with the same outside sizes; however, the lattice structure in the central part of the sample was replaced by a shell with wall thickness 2 mm. The basic BCC cell; the virtual models of both porous, as well as shell-shaped samples with the basic dimensions are presented in Fig. 4. FDM (Fused Deposition modelling) or FFF (Fused Filament Fabrication) 3D printers is primarily used for ABS material processing. Plastics come in the form of a long filament wound around a spool. Employing a print head, a molten layer of plastic is deposited on the print bed, which then adheres. Once the first layer has been drawn, the print bed drops and a new layer is built on the previous layer. This is repeated several times, ultimately resulting in a 3D printed model. The principle of FDM technique is shown in Fig" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0002204_icem49940.2020.9270855-Figure1-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0002204_icem49940.2020.9270855-Figure1-1.png", + "caption": "Fig. 1. Topology comparison of 12 slots/7 pole pairs FSCW PM machines. (a) Regular SPM machine. (b) Regular CPM machine. (c) Proposed DPM machine.", + "texts": [ + " In Section II, the topology of the proposed machine is compared with SPM and CPM machines, and the operation principle, especially the modulated air gap flux density harmonics and the working harmonics, are illustrated based on the flux modulation. After that, in Section III, the influence of stator PM on no-load electromagnetic performance, including the flux-linkage, back-EMF, cogging torque is analyzed with 12-slot, 7-pole pair machine by FEA. Then, in Section IV, the torque capability is researched. Finally, conclusions are given. P Authorized licensed use limited to: Rutgers University. Downloaded on May 15,2021 at 05:01:08 UTC from IEEE Xplore. Restrictions apply. II. TOPOLOGY COMPARISON AND OPERATION PRINCIPLE Fig. 1 compares the topologies of the proposed DPM machine (Fig. 1 (c)) with regular SPM (Fig. 1 (a)) and CPM (Fig. 1 (b)). Compared with SPM, nearly half PM is saved, and the material cost is reduced. However, the electromagnetic performance is sacrificed. Therefore, keeping the same rotor structure of CPM and considering the large space of stator slot opening, PMs are arranged in each stator slot opening. In order to increase the magnetic field density, Halbach magnetization type is used, each pole contains three magnet blocks, and the central block is magnetized in radial direction while the two side blocks are magnetized towards the adjacent teeth" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0002077_icma49215.2020.9233867-Figure5-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0002077_icma49215.2020.9233867-Figure5-1.png", + "caption": "Fig. 5 Description of ship coordinates and the key points", + "texts": [ + " (14) Finally, the estimated harmonics of different frequencies are superposed as the signals after the prediction time preT ( ) ( ) ( ) ( ) , , , 1 sin 2 1,..., N pre Obs i Obs i Obs i i pre w t A f t t i N T t T \u03c0 \u03d5 \u03c1 = = + + = \u2264 \u2264 (15) compensation value ( )t\u03c1 cannot be predicted, the constant at time t can be calculated as: ( ) ( )2 1\u02c6 N pret x T T t T\u03c1 += \u2264 \u2264 (16) Build the coordinate system B as shown in the Fig 8, the origin is the center of gravity of the ship; the z-axis is vertical upward; the y-axis points to the front of the ship, and the x-axis is determined by the right-hand rule. In the Fig. 5, the point M is the expected target point of the aircraft recovery lanyard. By adjusting the joint displacement VL VR , the position of the point M relative to the aircraft in the recovery process is ensured to be unchanged. MR ML, respectively, represent the control points of the recovery mechanism rope. Then at T0 time, the target point M0 is determined, and the coordinate of M0 in coordinate system B is 0 0 0 B B B M M MX Y Z and when the ship shakes due to the waves, the coordinate M0 in geodetic coordinate system D is 0 0 0 D D D M M MX Y Z " + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0002108_s40799-020-00405-5-Figure4-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0002108_s40799-020-00405-5-Figure4-1.png", + "caption": "Fig. 4 Geometry of the load head", + "texts": [ + " However, accommodating such joints would have made the load head bigger and heavier. We recall that load heads are meant to be as light as possible so as to not affect the dynamic behavior of the very light electrical equipment to which they will be connected. Different spherical joint ends were tested to create the connection, and the one finally chosen to achieve an optimal trade-off between backlash and friction. Each load head weighs \u223c35 lb. The geometry of the load head is illustrated in Fig. 4. The inner (solid) hexagon is the horizontal projection of the top and bottom plates. The coordinate system in which measured force and moment components are represented is shown at the center of this hexagon. The geometric details are illustrated with reference to load cell 3. Therefore, the elevation shows only the side on which this load cell is. Load cell 3 is shown in thick solid lines in plan and elevation. The solid dots at the ends represent the centers of its pin connections that standoff from the side of the hexagon by p", + " The vectors of forces applied by the six load cells on the top plate are labeled with circled numbers, and are shown at the points of intersection of their lines of action. The force vector, f3, corresponding to load cell 3 is shown explicitly in plan and elevation. The projections of each load cell on the horizontal and vertical planes are w and h respectively, so that the center-to-center length of a load cell is l = \u221a w2 + h2. Distances s and t to the centers of the pin connections are as shown in the Fig. 4. From the geometry, the side of the outer hexagon, and hence the distance from the center to a corner, is w\u0304 = w + 2 ( s + p\u221a 3 ) (1) The vertical distance, t\u0304 , from the point of intersection of forces to the pin center is t\u0304 = t s ( s + p\u221a 3 ) (2) The unit vectors in the directions of the load-cell forces are v\u03021 = \u2212w 2l \u0131\u0302 \u2212 \u221a 3w 2l j\u0302 \u2212 h l k\u0302 v\u03022 = \u2212w 2l \u0131\u0302 + \u221a 3w 2l j\u0302 \u2212 h l k\u0302 v\u03023 = w l \u0131\u0302 \u2212 h l k\u0302 v\u03024 = \u2212w 2l \u0131\u0302 \u2212 \u221a 3w 2l j\u0302 \u2212 h l k\u0302 v\u03025 = \u2212w 2l \u0131\u0302 + \u221a 3w 2l j\u0302 \u2212 h l k\u0302 v\u03026 = w l \u0131\u0302 \u2212 h l k\u0302 (3) so that force vectors are f1 = f1v\u03021, . . . , f6 = f6v\u03026. \u0131\u0302, j\u0302 , and k\u0302 are the unit vectors in the direction of the X, Y, and Z axes, respectively. The position vectors to the force intersection points, denoted r1, r2 and r3 in the Fig. 4, are given by r1 = w\u0304\u0131\u0302 + (t\u0304 \u2212 t)k\u0302 r2 = \u2212 w\u0304 2 \u0131\u0302 + \u221a 3w\u0304 2 j\u0302 + (t\u0304 \u2212 t)k\u0302 r3 = \u2212 w\u0304 2 \u0131\u0302 \u2212 \u221a 3w\u0304 2 j\u0302 + (t\u0304 \u2212 t)k\u0302 (4) r2 is shown in both plan and elevation for illustration. Denoting the measured force and moment components, R = (FX, FY , FZ, MX, MY , MZ) (5) and collecting the load-cell forces (tensions) in the form, f = (f1, f2, f3, f4, f5, f6) (6) the equilibrium of the top plate gives R = Cf (7a) where C is the 6\u00d76 matrix, C = [\u2212v\u03021 \u2212v\u03022 \u2212v\u03023 \u2212v\u03024 \u2212v\u03025 \u2212v\u03026\u2212r1\u00d7v\u03021 \u2212r1\u00d7v\u03022 \u2212r2\u00d7v\u03023 \u2212r2\u00d7v\u03024 \u2212r3\u00d7v\u03025 \u2212r3\u00d7v\u03026 ] (7b) If k denotes the stiffness of an individual load cell, it is easy to show that the stiffness matrix of the load head is given by K = kCC (8) assuming that the plates and pins are relatively stiff" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0000947_9783527342587.ch34-Figure34.4-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0000947_9783527342587.ch34-Figure34.4-1.png", + "caption": "Figure 34.4 Schematic of nodule\u2013fibril model representing microstructure of typical auxetic microporous polymer. (a) The polymer at rest and (b) the polymer at the tensile load. Source: Adapted from Liu and Hu 2010 [17].", + "texts": [ + " When the structure is subjected to an extension in a direction, the structure also expands in a direction that is perpendicular to the extension direction due to the free rotation of the rectangle units as shown in Figure 34.3b resulting in the negative PR effect. The auxetic effect depends on the strain and dimensions of the rectangles [16]. Auxetic materials based on nodule and fibrils achieve auxetic behavior due to the transition of nodes connected by fibrils under tension. One such example is illustrated in Figure 34.4. In the stress-free state, the nodules overlap with the fibrils wound around them as shown in Figure 34.4a. When the material is stretched in the direction of the fibrils, they get straight, and the nodules rotate and snap into a rigid grid-like arrangement as shown in Figure 34.4b [17]. Chiral structures are another kind of structures that have been developed for auxetic honeycombs. As shown in Figure 34.5, in this kind of structures, basic chiral units (highlighted in bold) are firstly formed by connecting straight ligaments (ribs) to central nodes, which may be circles or rectangles or other geometrical forms. The whole chiral structures are then formed by joining the chiral units. The auxetic effects are achieved through wrapping or unwrapping of the ligaments around the nodes in response to an applied force [17]" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0000802_gncc42960.2018.9018937-Figure1-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0000802_gncc42960.2018.9018937-Figure1-1.png", + "caption": "Figure 1. The different configurations of the morphing aircraft", + "texts": [ + " In section 3, the robust adaptive back-stepping controller based on RBFNN is designed, and the robust term is designed to eliminate approximation error between the real value and evaluated value approximated by RBFNN. Subsequently, the Lyapunov synthesis based on stability analysis is used to prove that all the signals in the closed systems are uniformly ultimately bounded with tracking error is convergent. The simulation results are shown and analyzed in Section 4. And Conclusions are given in Section 5. II. PROBLEM DESCRIPTION The backward-sweeping wings morphing aircraft which researched in this paper is shown in Figure 1. The control-oriented model of the longitudinal dynamics of a morphing aircraft considered in this paper is shown in Authorized licensed use limited to: Western Sydney University. Downloaded on August 15,2020 at 21:56:59 UTC from IEEE Xplore. Restrictions apply. equations (1)-(5). The thrust direction of this aircraft is along the axis of body frame. 1 ( cos sin )wV D T mg m \u03b1 \u03b3= \u2212 + \u2212 (1) sinh V \u03b3= (2) ( )1 sin coswL T mg mV \u03b3 \u03b1 \u03b3= + \u2212 (3) q\u03b1 \u03b3= \u2212 (4) 1 by y q M I = (5) where \u03b1 is angle of attack, \u03b3 is the flight path angle, q is the pitch rate, V is the velocity, h is the altitude" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0002347_j.compstruct.2020.113512-Figure1-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0002347_j.compstruct.2020.113512-Figure1-1.png", + "caption": "Fig. 1. An unsymmetrical cross-ply single-patch laminate with stable configurations at room temperature.", + "texts": [ + " In the past, a great deal of research has been carried out on morphing structures, but studies on morphing skins are predominately on flexible material research [1,2], which was summarized by a review paper on methods and applications [3]. Multi\u2010stable plates are considered as alternative morphing skins that can meet both shape changing and stiffness requirement. However, it is generally difficult to achieve continuous multi\u2010stable composite materials. The multiple stabilities of a composite laminate are usually accomplished by combining several bi\u2010stable patches. Fig. 1 exhibits a typical bi\u2010stable composite that has two distinct stable configurations created by a combination of geometric nonlinearity and residual pre\u2010stress during curing. The switch between two stable geometries is called snap\u2010through/snap\u2010back, excited either by a bending moment or a normal force to change from the primary stable shape to the secondary one. Around the world, a considerable amount of research has been carried out on bi\u2010stable composites. In the early 1980s, Hyer studied the cured shape of a thin unsymmetrical laminate [4\u20136]" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0001642_0309364620950850-Figure1-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0001642_0309364620950850-Figure1-1.png", + "caption": "Figure 1. Diagram of low-cost socket fabrication method. The protective sleeve is placed on the residual limb, and the hard plastic shell placed over the limb, to create a mold for the foam. The protective sleeve is folded over the proximal end of the shell to seal the mold. After the foam socket is cast, a friction wrist is installed in a wrist mounting cup positioned in the shell using adhesive or fasteners.", + "texts": [ + " However, like the International Committee of the Red Cross (ICRC) polypropylene socket,10,11 the CIR approach requires an oven, vacuum pump, electricity, and trained clinicians. We developed a rapid socket fitting protocol (Table 1) using a mass-producible outer socket shell and expandable rigid foam. Four individuals with unilateral transradial amputations were fit by a staff engineer with no formal prosthetic training. This technical note describes our method and presents results from initial functional evaluation of these sockets. The socket shell, which forms the outer structure of the socket (including the forearm) (Figure 1), was designed in SolidWorks (Dassault Syst\u00e8mes SolidWorks Corporation, Waltham, MA, USA), using 25th to 95th percentile male arm circumference values for the wrist, forearm, and elbow and the length12 to create a composite 3D arm envelope that can be cut to fit a wide range of people. Prototype shells were 3D-printed (Stratasys uPrint SE Plus, Eden Prairie, MN, USA) in acrylonitrile butadiene styrene (ABS; 0.254 mm), which allowed for iterative design and rapid fabrication. The socket was formed directly on the residual limb, using expandable polyurethane rigid foam (Foam IT 8, Smooth On Inc", + " The shell was cut to length with a customary U-shaped trimline at the cubital fold. The gel sheath was donned and the outer shell positioned over the limb. The proximal edge of the sheath was then folded back over the shell (Figures 1 and 2(a)) and sealed using self-adhesive wrap (Figure 2(d)). The shell could touch the residual limb but not fit so tightly that the flow of foam around the residual limb would be obstructed. The foam resins were mixed, then poured through the opening (gate) at the distal end of the shell (Figure 1). The shell was held in place on the arm as the foam expanded to capture the shape of the limb, forming a smooth interface, and bonded with the outer shell, which was incorporated into the final socket. The gate functioned as an escape valve to contain the flow of expanding foam and to create backpressure to ensure that foam filled all space between the limb and outer shell. Terminal devices with standard \u00bd\u201d-20 threads were attached to the socket using a one-piece 3D-printed prototype friction wrist (Figure 2(c))" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0003783_0954406215583522-Figure6-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0003783_0954406215583522-Figure6-1.png", + "caption": "Figure 6. Normalized acceptance surface for four contact point slewing bearings (s\u00bc 0.943, \u00bc 45 ).6", + "texts": [ + " As it has been mentioned, in previous works the authors developed and validated via FEA an analytical method that studies the interference field caused by the acting loads and provides all of the Fa\u2013Fr\u2013M load combinations that cause that contact pressure value, and therefore the static failure of the bearing.4\u20137 Figure 5 shows both the ball\u2013raceway interference field of the analytical method, and the FE model. The results of the analytical method can be represented in the three-dimensional \u2018\u2018acceptance surface\u2019\u2019 shown in Figure 6, where all of the Fa\u2013Fr\u2013M load combinations that lead to the static failure of the slewing bearing are represented. The axes are normalized with respect to the axial capacity of the bearing C0a, defined as the pure axial load Fa that causes the static failure. According to ISO 76 C0a \u00bc f0 z d 2 w sin \u00f01\u00de where dw is the ball diameter, z is the number of balls, is the initial ball-raceway contact angle, and f0 is a tabulated factor which can be approached by the following formula f0 \u00bc 61:6 80 dw=d cos \u00f02\u00de where d is the mean diameter of the bearing", + " The authors want to remark that each manufacturer should provide its own experimentally correlated expression for the axial load capacity C0a; the methodology developed in this work is valid regardless of the expression of C0a. Manufacturers usually use the value of C0a from expression (1), divided by a safety factor (in addition to the aforementioned acting load magnification factor); for example, IRAUNDI uses a factor of 2 for POS214 series. The overturning moment axis is also normalized with respect to the mean diameter of the bearing d, whereas the radial force axis is also normalized with respect to the tangent of the initial ball-raceway contact angle . As pointed out in Figure 6, the maximum values for the surface in the normalized axes are found to be Fa/C0a\u00bc 1, Fr/C0a tan \u00bc 0.4577 and M/d/C0a\u00bc 0.2288.1,4 Taking advantage of this normalization, the acceptance surface allows to calculate all of the admissible load combinations for four contact point slewing bearings taking as input data the mean diameter d, the axial load capacity C0a and the initial contact angle ; in other words, the acceptance surface in Figure 6 is a at UQ Library on June 4, 2016pic.sagepub.comDownloaded from unified surface, valid for any four contact point slewing bearing. It must be noted that the form of the surface slightly varies with the conformity of the ball-raceway contact s, and the initial contact angle (in Figure 6, s\u00bc 0.943 and \u00bc 45 ).6 The three-dimensional acceptance surface in Figure 6 can be represented in a bidimensional graph with normalized axes Fa/C0a (abscissa axis) and M/d/C0a (ordinate axis), plotting Fr/C0a tan as isolines (see Figure 7). As can be observed in Figure 7, the direction of the acting loads affects the static load-carrying capacity of the bearing. In order to obtain conservative and generalist curves, valid for any load direction, a superposition of the curves of the four quadrants is used, giving rise to the load capacity curves illustrated in Figure 8. Study of the bolted joint failure" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0000905_012027-Figure1-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0000905_012027-Figure1-1.png", + "caption": "Figure 1. Schematic diagram of the blocking door working.", + "texts": [ + " These methods are used to create dry ground construction conditions to repair metal structures. The blocking gate method refers to the method of using a floating gate to block the bottom hole of the water inlet directly under water to form dry ground construction conditions inside the bottom hole. This is especially advantageous when a cofferdam or a reservoir cannot be emptied Significantly, it has broad application prospects and practical value. The working principle of the blocking gate is shown in figure 1. This solution has been successfully applied to the structural restoration and management projects of reservoir dams such as Zhelin Hydropower Station and Gongzui Hydropower Station [7-8]. The 6th International Conference on Environmental Science and Civil Engineering IOP Conf. Series: Earth and Environmental Science 455 (2020) 012027 IOP Publishing doi:10.1088/1755-1315/455/1/012027 For the spillway of a concrete dam and the water inlet of a power station, accidental gates are often provided with a smooth and smooth chest wall" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0001099_j.promfg.2020.02.102-Figure4-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0001099_j.promfg.2020.02.102-Figure4-1.png", + "caption": "Fig. 4. Evolution of the gripper design: (a) benchmark version; (b) result of applying lightweight construction principles; (c) result of a simulation based topology optimization; (d) manufactured gripper (design from (c) with flexure hinge; (e) generative design variant with lowest weight.", + "texts": [ + " Non-positive gripping and holding of objects, corrosion resistance as well as a high number of gripping cycles had to be combined with a low weight for high dynamics. Based on existing CAD gripper data, several lightweight variants were designed using the conventional principles of lightweight construction such as frameworks. Additional to the weight, the effort for removal of support structures was an evaluation criterion. A weight reduction of 38% was achieved compared to the conventional benchmark gripper with a weight of 312 g (Fig. 4). The weight of the gripper was further reduced by applying a simulation based topology optimization with the target of maximum component stiffness and with specified mass savings of 60% and 80%, resp. (see Fig. 3 c). Exploiting a generative design method implemented in a CAD software, reduced the weight of the gripper to only 50 g. Compared to the benchmark, this is a reduction of 84%. At present, a gripper variant with a weight of 132 g (Fig. 3 d), equipped with an additively manufactured flexure hinge, is under experimental evaluation", + " Non-positive gripping and holding of objects, corrosion resistance as well as a high number of gripping cycles had to be combined with a low weight for high dynamics. Based on existing CAD gripper data, several lightweight variants were designed using the conventional principles of lightweight construction such as frameworks. Additional to the weight, the effort for removal of support structures was an evaluation criterion. A weight reduction of 38% was achieved compared to the conventional benchmark gripper with a weight of 312 g (Fig. 4). Fig. 4. Evolution of the gripper design: (a) benchmark version; (b) result of applying lightweight construction principles; (c) result of a simulation based topology optimization; (d) manufactured gripper (design from (c) with flexure hinge; (e) generative design variant with lowest weight. The weight of the gripper was further reduced by applying a simulation based topology optimization with the target of maximum component stiffness and with specified mass savings of 60% and 80%, resp. (see Fig. 3 c)" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0001846_eit48999.2020.9208283-Figure6-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0001846_eit48999.2020.9208283-Figure6-1.png", + "caption": "Fig. 6 Chain drive (left) from motor to rear", + "texts": [ + " It was specially designed and fabricated from the box tubing to accommodate the wheels and the steering mechanism. The steering mechanism was mounted using the clamps as shown. The steering assembly was similarly fabricated using stock metal and universal couplings that mated with the rack-andpinion steering mechanism. It was mounted using Unistrut C-channel as shown in Fig. 5. Due to space restrictions and a mismatch between the differential and motor splines, they were connected using a chain drive rather than directly, as shown in Fig. 6. The chain used was an ANSI #40 chain. Special spline couplers were fabricated to mate with the output of the motor Authorized licensed use limited to: University of Glasgow. Downloaded on November 01,2020 at 13:18:06 UTC from IEEE Xplore. Restrictions apply. and the differential. They were fitted with chain sprockets to mount the chain, as shown in Fig. 7. differential (right) The accelerator and brake pedals were fabricated and assembled, after a couple of prototypes were tried, as shown in Fig" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0000126_j.ifacol.2019.10.069-Figure15-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0000126_j.ifacol.2019.10.069-Figure15-1.png", + "caption": "Fig. 15. Collision of the nozzle to the table when making section 2.", + "texts": [ + "2 Round partitioning of the thin wall dome Based on the orientation showed in Fig. 10 (c), it is needed to be partitioned properly to be able to be built. Here, a 3 + 2 machine configuration is assumed. In order to prevent collisions, the part is split radially, as shown in Fig. 13. The building sequence is shown by numbers in the illustration. The nozzle should be tangent to the surface during production. Top edge of each sub-section is the substrate for starting the next section. Fig. 14. Splitting the surface dome into subsections. Fig. 15. Collision of the nozzle to the table when making section 2. The solution to prevent collisions is to build the part at the end of a plate. The plate is located vertically in the machine table. Therefore, when making the first layer of the first section, the trunnion table rotates 90\u00b0 to put the plate in a horizontal orientation. As the deposition of the layers progress, the plate turns vertically, and the dome gradually forms (Fig. 16). The benefit of the radial partitioning over the rotary tool path is that by increasing the number of partitions in Fig" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0003814_12.2040049-Figure3-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0003814_12.2040049-Figure3-1.png", + "caption": "Figure 3. Laser scanning approaches for the post-processing of LAM-produced parts. Left: Ablation of a single material layer at a constant z-offset to the design shape. Right: Processing of multiple layers with varying step \u0394z.", + "texts": [ + "org/terms by a Lumera Super Rapid system was used (pulse duration <15 ps, wavelength 1064 nm, repetition rate range frep = 0 - 500 kHz, maximum average power 10 W). Laser irradiations were performed using a 5\u00d7/NA = 0.1 microscope objective, and a CCD-based imaging system was utilized to control the beam position. The samples were translated by a 3D linear positioning stage Aerotech ALS130 at a constant speed v = 10 mm/s. Different scanning approaches were applied to maintain the design consistency, improve the resulting surface finish and processing efficiency (Figure 3). Curved surfaces were processed in multiple overlapping tracks resulting in layers that followed the final design shape at a constant vertical z-offset. This procedure was repeated for different offset values with varying step \u0394z = 10 \u2013 50 \u00b5m until the final design shape corresponding to a position z = 0 was revealed. The lateral separation between the tracks was \u0394x = 25 \u00b5m in all irradiation experiments. The resulting surface profile of the processed areas was investigated by optical microscopy (Olympus BX-51) and whitelight interferometry (Zygo NewView 6300)" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0003929_scis-isis.2014.7044649-Figure4-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0003929_scis-isis.2014.7044649-Figure4-1.png", + "caption": "Figure 4. Heading angles of a smartphone and riding robot", + "texts": [ + " The tangent velocity v and angular velocity \u03c9 of the robot are given by cos sinc c c cx y v\u03b8 \u03b8\u0394 \u2212 \u0394 = and c\u03b8 \u03c9\u0394 = (2) The equations of motion can be summarized as follows: cos 0 sin 0 0 1 c c c c c x v y \u03b8 \u03b8 \u03c9 \u03b8 \u0394 \u0394 = \u0394 (3) v and \u03c9 are given by , 2 2 r l c r l vv vv v d \u03c9 \u2212+ = = (4) where lv and rv are the pulse width modulation (PWM) of the left and right motors, and 2 cd is the horizontal distance between the wheels on both sides. The riding robot controlled by the smartphone controls a motor on both sides by a command contained in the communication packet transmitted from the smartphone. (a) Angle of smartphone Fig. 4 shows an orientation angle of the smartphone and riding robot. The orientation angles measured on the smartphone and riding robot are the variation between the Yaxis shown in Fig. 4 and the Earth\u2019s magnetic North Pole, and a rotation angle with respect to a Z-axis named yaw. The yaw of the smartphone is measured on the compass sensor for changing angles and in 30 ms intervals. Direction control based on the compass sensor controls is used to match both orientation angles measured by the compass sensor with respect to the magnetic North Pole of the smartphone and riding robot. The rider standing on the riding robot changes the direction of the smartphone to the desired direction to make the riding robot to move in the desired direction by pressing a start button on the screen of the 978-1-4799-5955-6/14/$31" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0000812_cdc40024.2019.9029757-Figure14-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0000812_cdc40024.2019.9029757-Figure14-1.png", + "caption": "FIGURE 14. Hydraulic principle of EHHT test system.", + "texts": [ + "13 shows the step signal response at amplitude 8MPa. Given that the load port pressure is about 0.15s delay at both control strategies, and considering the friction torque of EHHT, load pressure begins to increase when the EHHT\u2019s control angle increases and the driven torque is greater than friction torque. The simulation results show that the Fuzzy-PID controller can increase the EHHT\u2019s control performance. V. EXPERIMENTS We built a special laboratory bench for experiment of EHHT, the hydraulic system principle of EHHT is shown in Fig.14. The experiment bench is composed of following three parts. VOLUME 4, 2016 8613 (1) Main power of hydraulic circuit. This section provides a constant source of pressure for the electro-hydraulic servo plate-inclined plunger hydraulic transformer. It consists of a main motor, constant pressure variable pumps, hydraulic accumulators, safety valves, etc.. (2) Auxiliary oil source. This section provides controlled oil source for the laboratory bench. It mainly consists of an auxiliary motor, gear pumps, relief valves, hydraulic accumulators, hydraulic servo valve and filter" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0003788_detc2014-34213-Figure9-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0003788_detc2014-34213-Figure9-1.png", + "caption": "Fig. 9 Seven-bar linkage with input joints A and B0", + "texts": [ + " Thus, for the Watt six-bar linkages, the singularity happens when the joint E, A and A0 become collinear or the joints D, C and C0 become collinear. Therefore, the two-DOF seven-bar linkage is at singular positions. This conclusion is also obvious in that the inputs are given through the two common joints B and B0 of the two-five bar loops of the linkage. If the inputs are given through the non-neighboring joints in the same five-bar loop, such as joints A and A0, the sevenbar linkage can be decomposed into a Watt six-bar linkage shown in Fig. 9(a) with the input joint B0 held or a Stephenson six-bar linkage shown in Fig. 9(b) with the input joint A held. Thus, for the Watt six-bar linkage and the Stephenson linkages, the singularity happens when the joints E, B and A0 become 3 Copyright \u00a9 2014 by ASME Downloaded From: http://proceedings.asmedigitalcollection.asme.org/pdfaccess.ashx?url=/data/conferences/asmep/82119/ on 04/05/2017 Terms of Use: http://www.asme.org/about-asme/terms-of-use collinear or the joints D, C and C0 become collinear. Therefore, the two-DOF seven-bar linkage is at singular positions. Input joints not in the same five-bar loop In this group, the inputs are given through the joints not in the same five-bar loop of the seven-bar linkage, such as joints A0 and C0, the seven-bar linkage can be decomposed into a Stephenson six-bar linkage shown in Fig" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0000974_icase48783.2019.9059229-Figure1-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0000974_icase48783.2019.9059229-Figure1-1.png", + "caption": "Fig. 1. Sketch of aerodynamic forces acting on a golf ball. Lift force is generated in the direction from the side moving against to the side moving along the direction of the flow, commonly known as Magnus effect.", + "texts": [], + "surrounding_texts": [ + "Keywords\u2014Aerodynamic Efficiency; Dimple Characteristics; Dimples Depth; Wind Tunnel; CFD; Drag Reduction. Nomenclature d = Golf Ball diameter \u03b5 = Parameter for relative roughness k = Golf Ball dimple depth" + ] + }, + { + "image_filename": "designv11_71_0001304_j.matpr.2020.05.658-Figure3-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0001304_j.matpr.2020.05.658-Figure3-1.png", + "caption": "Fig. 3. Assembly of wheel chair using the rocker bogie mechanism.", + "texts": [ + " Since, the mechanism has retractable wheels, it is designed such as all the wheels can distribute the load angle automatically and motion of the wheel should not remain disrupted. All the joints are designed as mechanically hinged joints and stub axle joint or spring joint is not considered. A hinged joint is one, where, motion of one part takes place with respect to that axis of the other part. With the help of a hinged joint, it can be guaranteed that the system would remain stiff and rigid, even when it is placed at uneven surfaces. The assembly of the wheel chair using the rocker bogie mechanism is represented in Fig. 3. The design and simulation of the wheel chair using a rocker bogie mechanism is done using SOLIDWORKS 2014 and ANSYS R20 & R14. 14. The models of the rocker bogie mechanism is created using SOLIDWORKS. Meshing can be defined as the grid representation of any structure. Using meshing, it is easier to divide the forces applied upon the system into various domains. Structured mesh is done for the wheel chair with rocker bogie mechanism using ANSYS. Structural mesh can be represented in two different forms - either triangular form or rectangular form" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0002505_cac51589.2020.9327372-Figure1-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0002505_cac51589.2020.9327372-Figure1-1.png", + "caption": "Fig. 1. Model of the 2-DOF robot", + "texts": [ + " ROBOT MODELING Ignoring factors such as friction and torque disturbance, the dynamic model of a multi-joint robot can be expressed as: ( ) ( , ) ( )+ + =M q C q q q G q (1) where M is the inertia matrix. C is the centrifugal force and Coriolis force matrix. G is the gravity matrix. q , q , and q are the angular displacement, angular velocity, and angular acceleration of the robot joints, respectively. is the driving torque of the robot joint motor. This paper uses a 2-DOF joints robot as an example for research. Its physical model is shown in Fig. 1. The kinematic equation of the robot end position and joint displacement can be expressed as: 1 12 1 1 12 2 c c s s l l =x (2) where T[ ]x y=x is the position of the robot end in the Cartesian coordinate system. 1l and 2l are the lengths of two links of the robot; 1 1c cos( )q= , 1 1s sin( )q= , Assuming that the mass center of each link is at the end of it. According to the Newton-Euler equation of each link, the dynamic matrix parameters of the system under the joint coordinate system can be written as[17]: 2 2 2 1 2 1 2 2 2 1 2 2 2 2 2 1 2 2 2 2 2 2 2 1 2 2 2 2 ( ) 2 c c c m m l m l m l l m l m l l m l m l l m l + + + + = + M (3) 2 1 2 2 2 2 1 2 2 2 2 1 2 1 2 2 s s s 0 m l l q m l l q m l l q \u2212 \u2212 =C (4) 1 2 1 1 2 2 12 2 2 12 ( )g c g c g c m m l m l m l + + =G (5) where 1m and 2m are the mass of the links, and g is the gravitational acceleration" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0001138_0954406220925836-Figure14-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0001138_0954406220925836-Figure14-1.png", + "caption": "Figure 14. Kinematic diagram of the 3T3R-AMP2AMP2AMP3 GPM.", + "texts": [ + " As a result, the actuation wrenches of the hybrid limbs together with the wrench system of the middle limb constitute the basis wrenches.44 Therefore, the actuation scheme of the 3T3RAMP2AMP2AMP3 GPM is effective. Through similar analysis, the actuators\u2019 validity of other proposed GPMs in this paper can also be identified to be reasonable. The inverse kinematics is investigating the input variables of active joints after giving the position and orientation parameters of the end-effector with respect to the fixed base. To establish the geometric model of the proposed 3T3R-AMP2AMP2AMP3 GPM, the kinematic diagram is sketched in Figure 14. The global coordinate frame represented as {O-XYZ} is established at the middle point O of the fixed base. While the moving reference frame signified by {P-uvw} is attached to the intermediate moving platform D1D2. For the 3T3R-AMP2AMP2AMP3 GPM, the end-effector is the clamp assembled on the moving point P, and the motion of the endeffector can be described as the three-dimensional translation and three rotations around u-axis, v-axis, and w-axis. To solve the inverse kinematics, the geometric parameters of the links and platforms are defined in the following equation AiBik k \u00bc EiHik k \u00bc l1, BiCik k \u00bc l2, A3B3k k \u00bc l3, AiEik k \u00bc c, OAik k \u00bc a, EiFik k \u00bc b, CiDik k \u00bc h, PDik k \u00bc e, OA3k k \u00bc PC3k k \u00bc k, ffC2B2Ci 0 \u00bc \u2019, i \u00bc 1, 2 \u00f011\u00de The moving point P concerning the global coordinate frame can be represented by the vector P \u00bc x y z ", + " As a result, the transformation from the initial position to the moving reference frame can be described as the coordinate system {P-X0Y0Z0} rotates about X0-axis by the angle , then rotates b around v0-axis in {P-u0v0w0} after the first rotation, finally, rotates around the derived w-axis following previous rotations. It should be noted that the rotational angle about the w-axis is decoupled from other output parameters and can be calculated directly by the electric rotating machinery arranged at the middle limb. Therefore, the angle can be set as a constant for convenience. Then the output parameters degenerate into the system consisting of five variables, and the procedure of the transformation of the reference frames is depicted in Figure 14. The orientation matrix O P1R of the end-effector concerning the initial frame is carried out as \u00bdOPR \u00bc \u00bdRx\u00f0 \u00de \u00bdRy\u00f0 \u00de \u00bc 1 0 0 0 cos sin 0 sin cos 2 64 3 75 cos 0 sin 0 1 0 sin 0 cos 2 64 3 75 \u00bc cos 0 sin sin sin cos sin cos cos sin sin cos cos 2 64 3 75 \u00f012\u00de Let q1, q2, q3, q4, and q5 indicate the input angles of active revolute joints, and q6 signifies the distance of the prismatic joint in the middle limb. Afterward, the analysis of the inverse kinematics can be described as giving the position (x, y, z) and orientation ( , b, ) parameters of the end-effector, compute the angles and the displacement (q1, q2, q3, q4, q5, q6) of the active joints" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0000492_tmag.2019.2954900-Figure6-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0000492_tmag.2019.2954900-Figure6-1.png", + "caption": "Fig. 6. Fabricated motor for verification (a) Rotor assembly; (b) Stator assembly", + "texts": [ + " In 2D FEA analysis with the correction factor, the 3D FEA analysis and the analysis error were mostly within 0.5%, and the analysis time was only about 1/30 to 1/100 of 3D FEA. Fig. 5 shows the analysis result waveforms for each model when a stacking length of 20 mm was considered. The waveforms based on 2D FEA with the proposed correction coefficient approach closely matches the result of the 3D FEA approach. IV. VERIFICATION BY TESTING THE FABRICATED MOTOR The fabricated motor for the test is shown in Fig. 6, and its specifications are shown in Table \u2162. Fig. 6(a) and Fig. 6(b) show the rotor assembly and stator assembly, respectively. In Fig. 7, the motor and the test equipment for performance evaluation are shown. 0018-9464 (c) 2019 IEEE. Personal use is permitted, but republication/redistribution requires IEEE permission. See http://www.ieee.org/publications_standards/publications/rights/index.html for more information. CMP-539 5 Fig. 8 shows the results of no-load back EMF measurement. The results of the proposed 2D FEA and conventional 3D FEA was consistent with the experimental results" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0001341_s11227-020-03370-3-Figure12-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0001341_s11227-020-03370-3-Figure12-1.png", + "caption": "Fig. 12 The Environment of SIMULINK during real-time data acquisition, and SOLIDWORKS during mold design", + "texts": [ + "201, that is the result of the weighted decisions of the system. This output value can be considered as the final response of the constructed ANFIS model to the arbitrary input value. The following table represents a qualitative comparison between the investigated method and other modeling approaches. The superiority of the investigated method in generality and low calculation, in comparison with modeling as an uncertain system [38] and exact mathematical modeling, is the main reason that this paper has studied the ANFIS modeling approach. Figure\u00a012 is a brief view on modeling environment (Table\u00a02). In this paper, we investigated the ANFIS modeling of a soft pneumatic actuator. This actuator is built using a 3D printed mold and is capable of bending in the x\u2013y plane. This kind of finger-like actuator can have medical applications like its use in rehabilitation gloves. One challenge in controlling these kinds of actuators is 1 3 the modeling of their movement. In this paper, bending motion is considered for the actuator, and in order to gather its curvature data, a commercial bending sensor is installed on the outer side of this soft module" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0002253_lcsys.2020.3045669-Figure1-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0002253_lcsys.2020.3045669-Figure1-1.png", + "caption": "Fig. 1. (a) Front and (b) isometric views illustrating all joints in the human subsystem. (c) Isometric view showing joints in the robotic subsystem. (d) CAD illustration of the robot. (e) Assembled powered prosthesis.", + "texts": [ + " (2) We study the properties of the Jacobian of the Poincar\u00e9 map of the centralized and distributed controllers and show that there exists an upper bound on the norm of the difference of Jacobian matrices that significantly simplifies the search for optimal local control parameters. (3) We numerically validate the analytical foundation by designing distributed controllers for multi-domain and collaborative human-robot locomotion with a powered knee-ankle prosthetic leg and nontrivial feet, having a total of 19 DOFs (see Fig. 1). We show that the proposed approach can effectively synthesize local controllers for this comprehensive model of locomotion. In particular, we reduce the number of local controller parameters to be tuned and optimized to almost 5.2% of that in [22]. In this section, we consider single-domain hybrid models of locomotion to simplify the development of distributed control algorithms. The result can, however, be extended to multi-domain hybrid models of locomotion. We assume that the generalized coordinates of the mechanical system can be given by q \u2208 Q \u2282 R nq ", + " The proposed approach, however, significantly reduces the number of parameters to be optimized for the synthesis of distributed controllers. More specifically, we do not search for Cii and only focus on Ei. The objective of this section is to numerically verify the theoretical results for human-robot locomotion. We consider a lower-extremity amputee as the human subsystem and assume that a 3D tree-like structure with square-profile foot represents the human body (see Fig. 2(a), Fig. 2(c)). The robot subsystem on the other hand comprises a knee-ankle powered prosthesis where the foot is modeled with a square profile (see Fig. 1). While we model the powered prosthesis after [30], we extend it with an additional \u201cpassive\u201d DOF, attributed to the compliance in the ankle roll for simulation purposes. Under this consideration, we assume the entire structure to consist of 19 DOFs. The first 6 DOFs are attributed to the absolute position and orientation of the human body. The remaining 13 DOFs are the internal shaping variables that are distributed between both the legs as follows: 3 DOFs are associated with the hip to model the roll, pitch, yaw angles, one DOF captures the pitch Authorized licensed use limited to: Carleton University" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0003714_esda2014-20232-Figure1-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0003714_esda2014-20232-Figure1-1.png", + "caption": "Figure 1: A schematic cross section of the objective water brake", + "texts": [], + "surrounding_texts": [ + "Water brake is a type of Hydro Dynamometer. The mechanism of power absorption is based on making turbulence to achieve enough friction between rotor/stator and water. Rotor itself has to be perforated to amplify the friction and achieve this goal. Figure (2) shows this mechanism schematically. R o to r R o to r S ta to r S ta to r S ta to r S ta to r Gas Turbine Shaft Coupling W.B. Shaft Turbulence Turbulence Figure 2: Water brake mechanism Making turbulence is the major concern in water brake design that is attained by trapping water in the space between perforated rotor and stator. Although, there are several ways to measure the absorbed power, it is inferred that the best way is testing the mechanism. The operating envelope of the dynamometer is a region enclosed by several curves in its boundaries. It is common to show this envelope in power-rotational speed or torquerotational speed diagram. Since power is the main purpose of the measurements in dynamometer utilization purposes (studies), it is better to use power-rotational speed for design and analysis purposes [12]. As show in figure (3), the restrictions include: (a) maximum power absorption curve when the dynamometer is in fully filled mode, (b) the maximum torque curve that is related to mechanical shaft design constraint, (c) the maximum power absorption capacity with the maximum permissible temperature rise in water flow, (d) the maximum rotational speed associated with the system mechanical design limitations, (e) minimum allowable fill percent which is related to controllability and initial cooling. In addition to the above limitations, it is required to consider constraints such as water outlet temperature during the transient process in controller design process. The aforementioned restrictions demand a well-designed active control strategy to cover the limitations and provide safe operation of the water brake dynamometer. Downloaded From: http://proceedings.asmedigitalcollection.asme.org/ on 01/26/2016 Terms of Use: http://www.asme.org/about-asme/terms-of-use 3 Copyright \u00a9 2014 by ASME" + ] + }, + { + "image_filename": "designv11_71_0002861_978-3-030-55889-5-Figure1.50-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0002861_978-3-030-55889-5-Figure1.50-1.png", + "caption": "Fig. 1.50 Application of the Reduction principle, a and b moment distribution and Mohr\u2019s approach, c different paths have the same result, d Reduction principle", + "texts": [ + "277) But an engineer does it differently: The engineer installs a hinge and rotates both sides of the hinge so that tan \u03d5l + tan \u03d5r = 1. The bending energy is then simply the energy necessary to rotate the beam ends, and this energy is finite. In the case of the statically determinate beam in Fig. 1.49b it is even zero, since w\u2032\u2032 = 0. Remark 1.5 For further remarks aboutSobolev\u2019s embedding theorem, seeSect. 9.23. 74 1 Foundation The Reduction principle is a special version of Mohr\u2019s equation. Mohr would calculate the horizontal displacement ui of the frame in Fig. 1.50 by applying a unit point load X1 = 1 in the direction of ui , and then evaluate the integral 1 \u00b7 ui = \u2211 e \u222b le 0 ( M M1 E I + N N1 E A ) dx (1.278) by integrating over all frame elements. But according to the Reduction principle it suffices to apply the point load X1 = 1 to a substructure of the original structure, where substructure means any statically determinate system contained in the original system, as for example the single post in Fig. 1.50d. This surprising result is better understood when we realize that the node, which carries the load, can be reached from different starting points, see Fig. 1.50c, and that the sum of the horizontal displacement increments along 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. 1.50d 1 \u00b7 ui = \u222b l 0 ( M M1 E I + N (N1 = 0) E A ) dx . (1.279) The Reduction principle essentially states that the Dirac delta, the X1, need not be applied to the original system, but it can be any substructure, which is \u201ccontained\u201d in the original system, see Fig. 1.50d. Contained means, that in the transition to the substructure nodes of the original structure can be released, but no nodes may additionally be fixed. In the language of mathematics this means the Dirichlet boundary may shrink, but it may not grow [1], p. 149. The admissible substructures are typically the structures, which we 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_71_0002653_3449301.3449330-Figure4-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0002653_3449301.3449330-Figure4-1.png", + "caption": "Figure 4: Work Layout for Simulation Scenarios", + "texts": [ + " The control system performs by programming a Viper robot employing its step-by-step guide to programming robots. Then, three scenarios start with Pick from a fixed position to a fixed position, Pick from a fixed position to a pallet type. Finally, a Band Pick to a Pallet Place. This last scenario is based on the Honduras industry, where they have been implementing robots (harnesses, food, and education industry). A conveyor belt is used from which the Pick is originated and a pallet to make the Place. Most offline programs allow determining the working area of the robot as shown in Figure 4. Robot duty cycles are compared using a conventional tool, and it is necessary to be able to insert the design of the clamping system to perform these operations. The handling system is based on the Viper 650 robot, whose characteristics are shown in Table 1. It is a robot with industrial characteristics where it stands out its six degrees of freedom, allowing excellent mobility to develop industrial applications of any kind. It is limited by its reach of 653 mm and a maximum load of 5 kg, ideal for Pick & Place in the food industry, and its repeatability of 0" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0000986_cyber46603.2019.9066539-Figure1-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0000986_cyber46603.2019.9066539-Figure1-1.png", + "caption": "Fig. 1. Schematic representation of the 3D-LIPM with point foot. The model contains a point foot at position rankle, a point mass at position r with mass m and a massless leg linking the point mass and the point", + "texts": [ + " However, the ZMP method is unable to recover from large unexpected disturbance due to its key drawback. When the robot becomes unbalanced and the robot begins to rotate about the edge of the foot, the measured actual ZMP can\u2019t provide useful information, so the stepping strategy is applied and it will be introduced in section in detail. To use the ZMP control method on the robot, we should simplify the robot model. The well-known simplified model is Linear Inverted Pendulum Model (LIPM) [7] and depicted in Fig.1, containing a point mass body and a massless leg. The support polygon of the pendulum is a point, with position rankle, and the position of the point mass is located at r, the height of the point mass is kept at Zc by the force f of the simplified leg. The polygon of the dynamic system is a simple point, that means the only way to maintain balance is to keep the ZMP at rankle, so we need to control the CoM\u2019s position of the system following the equations of motion as m m= +r f g (1) where T[ ]x y zr r r=r means the position of the CoM corresponding to the world coordinate, f is the vector force acting on the point mass, g is the gravitational acceleration of the value [0 0 \u2013g]T" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0000865_sii46433.2020.9026171-Figure6-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0000865_sii46433.2020.9026171-Figure6-1.png", + "caption": "Fig. 6. AR.Drone 2.0", + "texts": [ + "01 s observation range Robs of agents dj radius 25 m Communication range Rcom of agents dj radius 50 m Control method of dj coverage control Control method of ci P control c1 c2 c3 Target position for agent ci [x, y] = [380, 120] [x, y] = [325, 170] [x, y] = [240, 250] Each weight W i W1 = 83/111 W2 = 23/111 W3 = 5/111 Fig. 8. This figure is an enlargement of initial position of all agetns. Blue dots are agents dj controlled by coverage control and red and green and cyan dots are agents ci controlled by P control. Green circles are observation range of dj . And Red and Magenta lines mean network graph. In this study, the controlled object handled in the simulation is AR.Drone 2.0 made by Parrot Corporation. As shown in Fig. 6, we define the appearance of A.R.Drone 2.0 and drone coordinate system x, y, z, attitude angle \u03d5, \u03b8, \u03c8. For other detailed information, please refer to [13]. We defined the equation of state model of AR.Drone2.0 as follows: x\u0307i = Axi +Bui (9) xi = [ xi vxi yi vyi ]T ui = [ uxi uyi ]T A = 0.0 1.0 0.0 0.0 0.0 \u22120.4430 0.0 0.0 0.0 0.0 0.0 1.0 0.0 0.0 0.0 \u22120.4644 B = 0.0 0.0 6.565 0.0 0.0 0.0 0.0 5.953 , Where xi, vxi , yi, vyi are the state of the i-th agent, and those are x coordinate, x axial velocity, y coordinate, y axial velocity" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0003921_coase.2014.6899477-Figure15-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0003921_coase.2014.6899477-Figure15-1.png", + "caption": "Figure 15. Reason why the contact force suddenly decreased.", + "texts": [ + " Table I shows the average value and standard deviation of the maximum contact force values. From Fig. 14 and Table I, we consider that the safety device switched off the two motors of the robot and stopped the robot after the contact force approximately reached the detection contact force level. The reason why the contact force suddenly decreased is because the robot was moved in the backward direction by the spring forces of Linear Springs Y after the two motors of the robot were switched off and the robot was stopped, as shown in Fig. 15 (see (v) in Section II). The differences between the detection contact force level and the experimental values are attributed to the output force errors of Linear Springs Y, friction in the safety device, among others. B. Velocity-based Detection Mechanism Next, for safety reasons, we mounted the robot on a spacer and attached some markers on Gear A (i.e. Shaft A), Claw A, and Plate A of the safety device. Then, we measured the velocity of Shaft A, the motion of Claw A, and the motion of Plate A by using a motion capture system (HAS-500, DITECT Corporation) while increasing the velocity of Shaft A by the motor (Motor 1) (see Fig" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0000451_hsi47298.2019.8942634-Figure8-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0000451_hsi47298.2019.8942634-Figure8-1.png", + "caption": "Fig. 8. Kinematical relashinship of waist, thigh and lower thigh.", + "texts": [ + " Its user always grasps the hand grip during walking ( lefthandP and righthandP are same position of hand grips). Its user\u2019s left arm and right arm are same posture. Using these assumptions, our walker can estimate COG position of its user. In this model, our walker knows a left and right foot position ( ,leftfoot rightfootP P ) and a waist position ( waistP ) by proposed scheme on a previous subsection. As the same, our walker derives the position of each joint. From these results, we can estimate the COG position of each linkage. For example, COG position of a left thigh link is shown in Fig. 8(b) and (12). We use human parameters as TABLE III. HUMAN PARAMETERS IV. PROPOSED WAKLING ASSISTANCE Our proposed walker assists its user\u2019s walking as follows; Usually, our walker generates the brake traction for negating the unbalance of its user\u2019s body caused by walking. It its user\u2019s posture exceeds the safety area, for example, its user falls down, our walker generates the maximum brake traction for safety reason. No Name M [%] C.G [%] K [%] Length [m] * 1 Lower arm 1.6 41.5 27.9 0.35 2 Upper arm 2" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0001138_0954406220925836-Figure7-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0001138_0954406220925836-Figure7-1.png", + "caption": "Figure 7. Available topological arrangements with AMP1 and AMP2.", + "texts": [ + " According to the above analysis, the 2F2C-H2limbs with R1RM and PRM arrangements serve as basic limbs to deduce all hybrid limbs with full constraints, as shown in Tables 6 and 7. Taking the R1RM limb as an example, the following flowchart to demonstrate the procedure of constructing F-H2limbs and C-H2limbs is given in Figure 6. To achieve high rotational capability, only two limbs are connected to the end moving platforms in the following proposed mechanisms in this section. Then the available topological arrangements are derived, as depicted in Figure 7. Afterwards, the design philosophy of GPMs with articulated moving platforms can be generalized as follows: (a) design articulated moving platforms and serial kinematic chains according to the expected motion. (b) Assemble the articulated moving platforms on the serial kinematic chains according to the arrangement of kinematic joints obtained in the previous section. By integrating the serial kinematic chains with the parallel revolute joints together with the necessary translation of the articulated moving platforms, the serial limbs are derived", + " Then the parallelogram mechanism Pa that plays the role of magnifying the translational range is approved to substitute the passive prismatic joint P. 3-DOF GPMs with articulated moving platforms After integrating the articulated moving platforms with serial or hybrid limbs, a novel class of GPMs with articulated moving platforms is derived. With different arrangements of kinematic joints, connected limbs provide two constraint systems to the endeffector. For the 3-DOF mechanisms, three active joints installed on or next to the fixed base are required. Hence, the practicable topological arrangement can only be the structure of Figure 7(a). As a result, one serial limb and one hybrid limb with two active joints are arranged on the two sides of the end-effector. For the 2T1R mechanisms, the wrench system is composed of one constraint-force and two constraint-couples. According to the constraint synthesis method,37 the constraint-couple can be generated by two parallel constraint-forces, and the geometrical requirements of assembling two limbs are demanded to be satisfied.42,43 Then all the groups of limbs that work together to realize the above wrench system are obtained as follows: 1F2C and D, 1F2C and F, 1F2C and C, 1F2C and 2C, 1F2C and 1F1C, 1F2C and 1F2C, 1F1C and C, 1F1C and 1F1C, 1F1C and 2C, F and 2C", + " As a result, the R3RMPaR3R 0 3 C-H2limb is transformed into the R1RMPaR1R3 C-H2limb. Where R3 is parallel to the direction of the R2 joint in the 2F1C-limb. Then the 1T2R-AMP2AMP2 GPM that can conduct high rotational performance about two directions is derived, as drawn in Figure 8(b). 4-DOF GPMs with articulated moving platforms When synthesizing 4-DOF mechanisms, four active joints are necessary to fully control the GPMs. Then the available topological arrangement of the structures can be seen in Figure 7(b) and (c). Hence, one serial limb and one hybrid limb with three active joints, or two hybrid limbs with two active joints are organized on the two sides of the end-effector. For 3T1R mechanisms, the wrench system is constituted by two constraint-couples. The combination of two limbs to provide the wrench system can be derived as follows: 2C and D, 2C and C, 2C and 2C, C and C. When employing the topological arrangement depicted in Figure 7(b), the symmetrical GPMs with identical limbs are desired. Then two R3RMPaR3 2CH2limbs are used to construct the 3T1R GPMs with AMP1 or AMP2. When assembling the R3 revolute joints of the limbs parallelly, the 3T1R GPMs with AMP2 is obtained; if the revolute joints equipped with actuators are coaxial, the 3T1R GPMs with AMP1 is acquired. Afterward, the 3T1R GPMs with AMP1 or AMP2 are obtained, and the 3T1R-AMP2 GPM is drawn in Figure 9(a). The wrench system of 2T2R mechanisms is composed of one constraint-force and one constraintcouple, which can be realized by 1F1C and D, 1F1C and C, 1F1C and F, F and F, F and C", + " To make sure the constraint-couples are parallel, the axis of R3 is parallel to or coaxial with the direction of R1, and R03 is parallel to R2. Besides, two groups of parallel revolute joints are assembled in turn to perform rotational motions. Hence, the R1RM R1R2 1F1C-H2limb evolves into the R3RMR3R 0 3 1F1C-H2limb, and the obtained 2T2R-AMP1AMP2 GPM is depicted in Figure 9(b). Similarly, the wrench system of 1T3R mechanisms is formed by 2F and D, 2F and F, 2F and 2F, F and F. The R1R1R2R2 2F-limb and R3RMRMR03R 0 3R D-H3limb are selected to design the GPMs with the topological structure shown in Figure 7(c). Where the axis of R joint in the R3RMRMR03R 0 3R D-H3limb is different from the R3 and R03 joints. Then, the 3T1RAMP2AMP2AMP2 GPM is obtained, as shown in Figure 9(c). 5-DOF GPMs with articulated moving platforms For the 5-DOF mechanisms, the topological arrangement is sketched in Figure 7(d). Since five active joints are required, two hybrid limbs connecting to the endeffector possess two and three active joints respectively. The wrench system of the 3T2R mechanisms is equivalent to a constraint-couple. According to the constraint synthesis method, only two combinations (C and D and C and C) are available to synthesize the 3T2R GPMs. When the R3RMPaR3R 0 3 H2C-limb and the R3RMRMR03 R03 H3C-limb serve as the connected limbs, the directions of revolute joints R3 and R03 in the H2C-limb is perpendicular to the constraint couple of H3C-limb", + " The R3RMPaR3R 0 3R D-H2limb and the R1RMRMR2R2 F-H3limb are employed to build the 2T3R GPMs. To conduct rotational motions about three directions, three groups of parallel revolute joints are assembled to the end-effector. Then the R3RMPaR3R 0 3R D-H2limb is transformed into the R3RMPaR3PR 0 3R2 D-H2limb, and the 2T3RAMP2AMP2AMP2 GPM is derived, as depicted in Figure 10(b). 6-DOF GPMs with articulated moving platforms When constructing 6-DOF mechanisms, only two hybrid limbs with three active joints can be used to fully control the GPMs, and the topological structure is drawn in Figure 7(e). With no constraint exerting on the end-effector, only unconstrainted limbs are feasible to design mechanisms. When employing two R3RMRMR03R 0 3R D-H3limbs to connect the two sides of the end-effector, three groups of parallel revolute joints are included in the limb. To make sure the hybrid limbs can perform the necessary three-dimensional translation, the R3 revolute joins in the two limbs are not parallel. Then the 3T3R-AMP2AMP2AMP2 GPM with rotations about three directions is obtained, as shown in Figure 11" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0002342_rcar49640.2020.9303289-Figure2-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0002342_rcar49640.2020.9303289-Figure2-1.png", + "caption": "Figure 2. Robot system overall structure", + "texts": [ + " Kefan Xing, Yinghan Wang, Diansheng*Chen Min Wang and Sitong Lu are with Institute of Robotics, Beihang University, Beijing, 100191, China. chends@163.com. 978-1-7281-7293-4/20/$31.00 \u00a9 2020 IEEE 464 Authorized licensed use limited to: Univ of Calif Santa Barbara. Downloaded on May 19,2021 at 02:45:08 UTC from IEEE Xplore. Restrictions apply. The flexible ankle-assist robot prototype mainly includes a rope drive mechanical module, a spring recovery module, a foot wearing module and a calf wearing module. As shown in figure 2, the rope drive mechanical module provides the driving force. And dorsiflexion is completed by pulling the feet with the rope. Spring recovery module can provide elastic potential energy which can make up for the disadvantage that the rope drive can\u2019t be driven in the reverse direction, and provides spring tension for ankle joint plantarflexion. The foot wearing module is mainly integrated with pressure sensors, inclination sensors, and integrated insole. The calf wearing module mainly binds the rope drive module and the spring module to the calf" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0003559_iccpct.2014.7054910-Figure1-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0003559_iccpct.2014.7054910-Figure1-1.png", + "caption": "Fig. 1: A 6/4 SRM", + "texts": [ + " Though the machine is a type of synchronous machine, it has certain novel features that make the machine more advantageous with respect to the other machines used in this field. The stator windings are wound field which is similar to that of the DC motor and it has no windings or magnets on its rotor. The machine is a doubly salient one as it has salient poles on both rotor and stator. A 6/4 pole Switched Reluctance Motor is used in this paper for the simulation of the sensor less scheme. The basic structure of a 6/4 SRM is shown in Fig.1. The rotor comes to an aligned position whenever the diametrically opposite stator poles are energized. The rotor moves to the minimum reluctance position in accordance with the magnetic circuit theory. Here, two rotor poles are in alignment with the two stator poles and two other rotor poles are out of alignment with the current stator poles. These rotor poles then come to its 250978-1-4799-2397-7/14/$31.00 \u00a92014 IEEE minimum reluctance position when the next set of rotor poles is energized. Hence, this system involves the switching of currents into the stator windings when there is a variation of reluctance" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0001148_1077546320927598-Figure2-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0001148_1077546320927598-Figure2-1.png", + "caption": "Figure 2. Simplified model of robotic manipulator with base vibration.", + "texts": [ + " It can produce vibration in three directions: (1) in vertical direction (linear vibration along axis y), (2) in pitching direction (rotational vibration relevant to axis z), and (3) in swing direction (rotational vibration relevant to axis x). xoy denotes the local coordinates, which is attached to the profiled frame structure. Intuitively, the transfer arm rotates relative to the motor output shaft in the xoy plane. Hence, in its moving process, it will be mainly influenced by the vibration in vertical and pitching direction. The simplified model of the robotic manipulator system with compound vertical and pitching vibration is shown in Figure 2. In Figure 2, XOY denotes the Cartesian coordinate system; xoy is the local coordinate system, which is attached to the installation base; G0, G1, and G2 represent the mass centers of installation base, transfer arm, the load mounted at the end of transfer arm, respectively; O2 represents the rotation center of the transfer arm; yb and \u03b8b are the vibration displacements in vertical and pitching direction, respectively; \u03b8 is the angular displacement of the transfer arm relative to the installation base; \u03b1 represents the angle between the central line of the transfer arm and the line linked by O2 and G2; L0, L1, L2, and H are geometric sizes as shown in Figure 2. By virtue of the second kind Lagrange equation, the system dynamic equation can be derived as Je\u20ac\u03b8 \u00bc U \u00feMG MR Mf \u00fe Sd (1) where Je \u00bc m1L21 \u00fe m2L22 \u00fe I1 \u00fe I2 is the equivalent moment of inertia of the robotic manipulator system.m1 andm2 are the masses of the transfer arm and load. Correspondingly, I1 and I2 are the moment of inertia of the transfer arm and load. U \u00bc Mdi\u03b7 \u00bc KTI i\u03b7 is the general control torque. In which, Md \u00bc KTI is the driving torque from the motor, KT is torque constant of the motor, and I is the current value" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0001964_biorob49111.2020.9224403-Figure1-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0001964_biorob49111.2020.9224403-Figure1-1.png", + "caption": "Fig. 1: (a) Image of the actuator with lower leg attachement. Source: smarcos.eu, (b) Kinematic model of the actuator.", + "texts": [ + " Section III describes the control strategy which compensates for stiffness faults. Section IV shows biomechanical data obtained from a knee flexing experiment, which work as the reference trajectory for the control strategy evaluation discussed in Section V. Finally, the conclusions of the paper are presented in Section VI. 978-1-7281-5907-2/20/$31.00 \u00a92020 IEEE 660 Authorized licensed use limited to: Middlesex University. Downloaded on November 03,2020 at 09:12:06 UTC from IEEE Xplore. Restrictions apply. Fig. 1a shows the SMARCOS actuator, a MACEPPAbased smart variable stiffness actuator, which is used to demonstrate the feasibility of the control method proposed in this paper. The MACCEPA is a concept conceptualized by the Vrije Universiteit Brussel (VUB) [15], which generates a quasi linear torque function with respect to the deflection of two rotating bodies connected through a linear spring. The kinematic diagram is shown in Fig. 1b. The system consists of an input link (orange colored), an output link (blue colored), and a lever arm (gray colored) rotating around point R. The output link and the lever arm are connected at points C and D through a spring with stiffness k. The output stiffness of the actuator is adjusted by modifying the length of pretension P of the spring. The angular position of the arm \u03d5a and output link \u03d5l are measured with respect to the input link, with the neutral position attaining a 120\u25e6 angle as shown in Fig. 1b. The arm position \u03d5a can be modified by adjusting the length between points A and B with linear motion performed by a servomotor coupled with a screw drive. The low-level control is performed on the integrated electronics containing a BLDC (brushless) servo control module configured to operate in current control mode. The actuator measures the force and estimates the output torque. Furthermore, the motion range of the arm is mechanically limited to \u03d5a \u2208 [\u2212\u03c0/4,\u03c0/3]. Fig. 2 shows the configuration of the lever arm section with length LC, and output link with length LD" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0001877_icarm49381.2020.9195392-Figure2-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0001877_icarm49381.2020.9195392-Figure2-1.png", + "caption": "Fig. 2. Locust-inspired 3D printed forewing, from left to right clockwise; unprocessed 3D printer-adapted forewing frame, wing thickness, PLA venation mould, and PVC reinforced forewing prototype.", + "texts": [ + " Although most of the presented work exploit experimental facilities such as Objet EDEN260V, we made an attempt to obtain similar if not improved results with a consumer-grade printer to prove concealed potentials of readily available equipment. This is achieved by further experimenting with the numerous printer parameters (extruder & bed temperature, infill, layer height, extrusion & travel speed, etc.) and environment to create an almost ideal condition for a precision rapid prototyping of complex and delicate CAD models. As shown in Figure 2, the major venations enclosed in the outer wing frame of a locust-inspired forewing is fabricated with the planform parallel to the XZ plane of the printer to compensate for the nominal (150\u03bcm) layer-resolution achieved on a LulzBot Taz5 machine. Initially, biodegradable and odourless polylactic acid (PLA) is used to print at 150\u03bcm layer thickness. This material is favourably characterised with low melting point and minimum warping that produces precise corrugated venation patterns. The CAD files for both versions of the wings (original and optimised) are sliced and processed for printing. Furthermore, the same samples were printed using carbon fibre infused PETG to further extend the scope of this investigation. Post-processing involved surface finishing and removal of residues with the help of alcohol wipes and fine grit (>1000) sandpapers. Finally, the CF-PETG wing skeletal structures shown in Figure 2, are placed in a PLA printed venation mould for reinforcing the wings with a 50\u03bcm thick Polyvinyl Chloride (PVC) membrane also called vinyl wrap coated with a thin layer of thermoplastic rubber and plasticizer solution (35% VM & P Naphtha, 30% Resins, 15% Hexane, 15% Toluene, 5% Methyl ethyl ketone) sprayed evenly to form a membrane masking venations, respectively. Salient advantages are high flexural stiffness when compared to a chitosan (linear polysaccharide 1,4 -linked 2aminodeoxy-b-D-glucan) membrane that is much similar to a real insect wing membrane introduced in the literature [33], [36]" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0001566_012084-Figure7-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0001566_012084-Figure7-1.png", + "caption": "Figure 7. Dynamic finite element model.", + "texts": [ + " The density of the virtual material layer can be obtained by 1 1 2 2 1 2 1 2 2 V V + = V V (12) where \u03c1 is the density of the virtual material layer, \u03c11, V1 are the density and volume of structure 1, respectively. \u03c12, V2 are the density and volume of structure 2, respectively. The material property and structural parameters of the virtual material layer are obtained as shown in Table 3. The material property parameters in Table 3 are in-built in ANSYS. The virtual material layer and the component surfaces are operated by GLUE. The dynamic model finite element is established, the local finite element model is shown in Figure 7, and modal analysis is carried out. Depending on the plate size of the lapped beam structure in the previous section, two plates which are consistent with the simulation parameters are machined for the modal test. In this study, the hammering modal test of the overlap beam structure with bolt joint is carried out using the LMS Test. Lab 14A vibration test and analysis system. The experimental device is composed of a hammer, threeway acceleration sensor, LMS Test Lab vibration test and analysis software, data acquisition system, structural specimen of overlapped beam, signal line, elastic rope, suspension, and the PC" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0000253_2019045-Figure4-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0000253_2019045-Figure4-1.png", + "caption": "Fig. 4. Helix diagram of left tooth surface.", + "texts": [], + "surrounding_texts": [ + "The tooth surfaces of double helical gear is involute helicoid. Figures 4 and 5 are its left and right tooth surface helix diagram, respectively. The left tooth surface equation is expressed as xk1 \u00bc rb1 sinmk1 rb1mk1 cosmk1 \u00fe rk1 sin uk1 yk1 \u00bc rb1 cosmk1 \u00fe rb1mk1 sinmk1 \u00fe rk1 cos uk1 rb1 zk1 \u00bc rk1uk1 tan gk1 8< : \u00f04\u00de The right tooth surface equation is expressed as xk2 \u00bc rb2 sinmk2 rb2mk2 cosmk2 \u00fe rk2 sin uk2 yk2 \u00bc rb2 cosmk2 \u00fe rb2mk2 sinmk2 \u00fe rk2 cos uk2 rb2 zk2 \u00bc rk2uk2 tan gk2 : 8< : \u00f05\u00de Equations (4) and (5) are all the equation of involute helicoid. Where, rbi (i=1, 2) is the radius of base circle, mki (i=1, 2) is the roll angle of point k in the involute, rki (i=1,2) is the radiusofpoint k in the involute, uki (i=1,2) is theexpansionangleofpoint k in the involute,gki (i=1,2) is the helix angle of ascent of point k in the involute, bki (i=1, 2) is the helix angle of point k in the involute." + ] + }, + { + "image_filename": "designv11_71_0000254_ecce.2019.8911883-Figure12-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0000254_ecce.2019.8911883-Figure12-1.png", + "caption": "Fig. 12. Comparison of on-load flux lines: (a) Conventional FSPM machine. (b) Proposed FSPM machine.", + "texts": [ + " 0 60 120 180 240 300 360 -150 -100 -50 0 50 100 150 200 P ha se b ac k -E M F Rotor position (deg) Conventional FSPM machine Proposed FSPM machine (a) 0 2 4 6 8 10 0 25 50 75 100 125 150 P h as e ba ck -E M F ( V ) Order Conventional FSPM machine Proposed FSPM machine (b) Fig.10. Comparison of phase back-EMFs: (a) Waveforms. (b) FFT spectrums. 0 60 120 180 240 300 360 -200 0 200 C og gi n g to rq u e (m N m ) Rotor position (deg) Conventional FSPM machine Proposed FSPM machine Fig.11. Comparison of cogging torques The on-load flux lines and flux contour plots of the compared two FSPM machines are presented in Fig. 12 and Fig. 13. It can be found that the flux density in the stator yoke of the proposed FSPM machine is larger than that of the conventional one, because the flux barrier effect of the proposed topology becomes lower. Moreover, the rated torque waveforms of two FSPM machines are compared in Fig.14 (a). It reveals that the proposed FSPM machine has 17.6% higher average torque than the conventional FSPM machine. Moreover, the PM usage of the proposed FSPM is also 35.5% smaller. Therefore, the torque per PM usage of the proposed FSPM machine is about 1" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0002213_icem49940.2020.9271042-Figure2-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0002213_icem49940.2020.9271042-Figure2-1.png", + "caption": "Fig. 2. Field solution for a 2D magnetostatic problem, visualizing the lines of constant z in black, the absolute value of the flux density in the ferromagnetic parts and the airgap as well as z in the conducting parts.", + "texts": [ + " This section provides an overview of the model workflow (III.-A.), presents the necessary simplifications (III.-B.) and the model equations for the electromagnetic machine submodel (III.-C.). In the sampling process, the state variables (currents and the rotor position ) are varied in appropriate ranges and resolutions. If the number of state variables is , and the number of sampling points for these state variables are then the number of samplings , is (1) For every state, a magnetostatic problem is solved (see Fig. 2). The corresponding internal parameters (linkage fluxes ) and output parameters (forces and torque ) are evaluated from the field solution and stored in look-up tables. In a transient simulation, at every time step the controller/ inverter model based on pulse-width modulation (PWM) provides a voltage vector . In case of the SCIM, this is the vector of stator voltages (see Fig. 1). Note that this vector, like all other vectors and scalars exchanged by the submodels depicted in Fig. 1, contains instantaneous values for a single time step" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0002974_ijmpt.2019.104568-Figure2-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0002974_ijmpt.2019.104568-Figure2-1.png", + "caption": "Figure 2 Dimension SST 1200es FDM machine crest ultrasonic generator (see online version for colours)", + "texts": [ + " Most of the previous works focused on correcting dimensional accuracy in the FDM process by obtaining optimum parameter settings using control factors of the process and finding optimal shrinkage compensation factors. So far, to the best of the authors\u2019 knowledge, no work has been reported on building a predictive model for the volumetric error in the FDM process. In the present work, an ANFIS model was built to predict the volumetric shrinkage in the FDM process. The effects of the FDM process parameters on the circularity and straightness of small hollows was also analysed. The geometrical structures designed were manufactured using the Stratasys \u2018Dimension SST 1200es\u2019 FDM machine shown in Figure 2(a). This machine is easy to operate, and it also consists of a large build chamber with dimensions of 406 mm \u00d7 355 mm \u00d7 406 mm, as such it could be used to build both medium and small sized prototypes. Hence, it is very suitable for RP manufacturing purposes. Moreover, the P430 ABSplus Ivory model was used as the build material and ABS material was utilised as support material. In addition, the crest ultrasonic generator shown in Figure 2(b) was used to detach the support material from the built model as a post-processing process. In this work, five geometric models were designed as the control parts using CATIA-V5 3D software. Among the models, one model is kept as a full solid, while the other four models contain different geometries of hollow shapes. Table 1 shows the theoretical (CAD) volumetric dimensions of the designed samples, which are similar to the specifications in ASTM D638 for standard tests. The total theoretical volume VCAD in each sample as obtained using equation (1) are shown in Table 1" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0003183_icra.2015.7139672-Figure9-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0003183_icra.2015.7139672-Figure9-1.png", + "caption": "Fig. 9. The path consists of 3 singular (red) and 2 regular primitives (blue)", + "texts": [ + " To analyze different scenarios, we consider adding a singular primitive S1 between Si and Sg. Here we present the analysis regarding \u03b2 at the switching point from Si to S1. When the trajectory of the grain cart transitions to S1, it is evident that the grain cart has \u03c9 > 0 and \u03c9\u0307 < 0. Since the path merges into the line segment in S1, we can conclude that \u03b2\u0307 =\u2212vsin\u03b2 +\u03c9 < 0, which implies \u2212vsin\u03b2 < 0. Since v = 1, we have \u03b2 < 0 at the switching point to S1. Similarly, we can derive that \u03b2 > 0 at the switching point to Sg. Figure 9 depicts the path that consists of three singular primitives. However, since the constraints (11) and (13) hold for all singular primitives, we can be conclude that the direction of the line segment in S1 is constrained in the interval [\u2212\u03c0 3 , \u03c0 3 ]. Thus spatial restrictions on the destination still exist. To resolve this issue, we additionally propose a path containing 2 regular and 3 singular primitives, as shown in Fig. 8. The grain cart proceeds on a regular primitive R1 when it is safe to turn" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0003240_ascc.2015.7244503-Figure1-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0003240_ascc.2015.7244503-Figure1-1.png", + "caption": "Fig. 1. TRMS system.", + "texts": [ + " The switching surface is defined using Quadratic minimization as in [6]. The switching control is added with equivalent control obtained from feedback linearization [7] to get highly robust controller. The remaining portion is arranged as follows. In Section II, an overview of the system parameters and decoupling are provided. In Section III surface design is given. In Section IV, sliding mode controller design has been described. In Section V the sliding mode performance is presented. Finally concluding remarks are made in Section VI. . As shown in Fig. 1 , the TRMS is a mechanical unit consisting of two rotors, one main rotor and another tail rotor placed on a beam together with a counter balance [2]. The main rotor produces a lifting force allowing the beam to rise vertically while the tail rotor is used to make the beam to turn left and right around the horizontal axis. The following momentum equation can be derived for the vertical movement as it appears in [1] and [2]. The moment of inertia of main motor is given by \u03c8\u0308I1 =M1 \u2212MFG \u2212MB\u03c8 \u2212MG (1) where \u03c8\u0308 is the pitch angle acceleration, M1 is the nonlinear static characteristic of main rotor,MFG is the gravity momentum, MB\u03c8 is frictional force momentum and MG is gyroscopic momentum" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0000603_iccas47443.2019.8971558-Figure7-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0000603_iccas47443.2019.8971558-Figure7-1.png", + "caption": "Figure 7.Cable eyelet arrangement for end and mid -section discs", + "texts": [], + "surrounding_texts": [ + "220\nC. Actuating mechanism Manipulator's tendons actuate by a linear lead shaft\nconnected to the stepper motor by a coupler. Tendons attached to both sides of specially designed inlets on screw housing and screw with housing travels along the shaft. 1mm steel cable utilized as a tendon.\nIII. KINEMATICS AND KINETIC FORMULATION\nA. Forward kinematic formulation Coordinate systems are set at every universal joint. The homogeneous coordinate transform matrices:\n\u03a30\u2192\u03a31 , H0,1= ( R u0,1 0 0 0 1 ), u0,1=( x0 y0 l0 )\n(1)\n\u03a3i-1\u2192\u03a3i, Hi-1, i= ( R ui-1,i 0 0 0 1 ) , ui-1,i=( 0 0 L ),\n(i=2, \u22ef, n)\n(2)\nR= Rz(\u03b8zi)Rx(\u03b8xi)Ry(\u03b8yi) (3)\nwhere, \ud835\udc65\ud835\udc650 and \ud835\udc66\ud835\udc660 are an initial position of the base. \ud835\udc45\ud835\udc45\ud835\udc65\ud835\udc65(\ud835\udf03\ud835\udf03\ud835\udc65\ud835\udc65\ud835\udc65\ud835\udc65) and \ud835\udc45\ud835\udc45\ud835\udc66\ud835\udc66(\ud835\udf03\ud835\udf03\ud835\udc66\ud835\udc66\ud835\udc65\ud835\udc65) are rotation matrices of ith universal joint that has two rotation angles \ud835\udf03\ud835\udf03\ud835\udc65\ud835\udc65\ud835\udc65\ud835\udc65 and \ud835\udf03\ud835\udf03\ud835\udc66\ud835\udc66\ud835\udc65\ud835\udc65, \ud835\udc45\ud835\udc45\ud835\udc67\ud835\udc67(\ud835\udf03\ud835\udf03\ud835\udc67\ud835\udc671) is a rotation matrix of the ith disk with a rotation angle \ud835\udf03\ud835\udf03\ud835\udc67\ud835\udc671 along the axial axis and L is a length between neighbouring universal joints. Three rotation matrices have:\nMultiplying the H-matrices successively, we get unit\nvectors and the position vector of the ith coordinate system; H0,i=H0,1H1,2\u22efHi-1, i= (\nii ji ki ui 0 0 0 1 ) (4)\nwhere, \ud835\udc62\ud835\udc62\ud835\udc65\ud835\udc65 is the position of the ith universal joint Ui (\ud835\udc56\ud835\udc56 =\n1,\u22ef , \ud835\udc5b\ud835\udc5b \u2212 1). The position vector \ud835\udc5d\ud835\udc5d\ud835\udc65\ud835\udc65 of the end-point P\ud835\udc5b\ud835\udc5b and position of sliding plates Pi (i=1, \u22ef, n-1) of the manipulator are obtained by, (pi\n1)=H0,i(0 0 li 1)T, (i=1, \u22ef, n) (5)\nwhere, \ud835\udc59\ud835\udc59\ud835\udc5b\ud835\udc5b is a fixed length between the nth universal joint and the most distal plate. Position vectors of 8 hole A0, B0, C0, D0 , A\u03020, B\u03020, C\u03020, D\u03020 at the base plate are determined as,\na0=( ax ay 0 ) , b0=( bx by 0 ) , c0=( cx cy 0 ) , d0=( dx dy 0 ) ,\na\u03020=( a\u0302x a\u0302y 0 ) , b\u03020=( b\u0302x b\u0302y 0 ) , c\u03020=( c\u0302x c\u0302y 0 ) , d\u03020=( d\u0302x d\u0302y 0 ) ,\n(6)\nPosition vectors of 4 hole A\ud835\udc65\ud835\udc65, A\u0302\ud835\udc65\ud835\udc65, C\ud835\udc65\ud835\udc65, C\u0302\ud835\udc65\ud835\udc65 at the ith plate\n(\ud835\udc56\ud835\udc56 = 1,\u22ef , \ud835\udc5b\ud835\udc5b) are obtained as\n(ai 1)=H0,i( ax ay li 1 ) , (a\u0302i 1)=H0,i( a\u0302x a\u0302y li 1 ) ,\n(ci 1)=H0,i( cx cy li 1 ) , (c\u0302i 1)=H0,i( c\u0302x c\u0302y li 1 ) , (i=1, \u22ef, n)\n(7)\nwhere, \ud835\udc59\ud835\udc59\ud835\udc65\ud835\udc65 is an axial length between the ith universal joint and the ith plate, which varies as the plate slides along rods, except \ud835\udc59\ud835\udc59\ud835\udc5b\ud835\udc5b.\nIn the same way, position vectors of 4 hole B, B\u0302\ud835\udc65\ud835\udc65, D\ud835\udc65\ud835\udc65, D\u0302\ud835\udc65\ud835\udc65 at the ith plate (\ud835\udc56\ud835\udc56 = 1,\u22ef ,\ud835\udc5a\ud835\udc5a) are obtained as,\n(bi 1)=H0,i( bx by li 1 ) , (b\u0302i 1 )=H0,i ( b\u0302x b\u0302y li 1) ,\n(di 1)=H0,i( dx dy li 1 ) , (d\u0302i 1 )=H0,i ( d\u0302x d\u0302y li 1) , (i=1, \u22ef, m)\n(8)\nB. Kinetic formulation Our continuum manipulator is divided by two segments.\nAuthorized licensed use limited to: Auckland University of Technology. Downloaded on October 25,2020 at 08:37:21 UTC from IEEE Xplore. Restrictions apply.", + "221\nThe first segment is located by the distal part and the second segment is located by the proximal part. The first segment is operated by 2 pairs of 2 wires; total 4 wires. One pair of 2 wires is controlled by one motor that pulls one wire and push the other wire in the same length by using the pulley. While, the second segment is operated by 4 pair of 4 wires, total 8 wires. Therefore the second segment is controlled by 4 motors The second segment has m units and the first segment has n-m units. 4 pairs of wires are labeled by \ud835\udc4e\ud835\udc4e and ?\u0302?\ud835\udc4e, \ud835\udc4f\ud835\udc4f and ?\u0302?\ud835\udc4f, \ud835\udc50\ud835\udc50 and ?\u0302?\ud835\udc50, d and ?\u0302?\ud835\udc51,\nEquilibrium in moments at Un belonging to the first segment is (Sa,n-fa)(an-an-1)\u00d7(an-un)+(Sa\u0302,n-fa\u0302)(a\u0302n-a\u0302n-1)\u00d7(a\u0302n-un)+\n(Sc,n-fc)(cn-cn-1)\u00d7(cn-un )+(Sc\u0302,n-fc\u0302)(c\u0302n-c\u0302n-1)\u00d7(c\u0302n-un)++\nmw(pn-un)\u00d7g= ( 0 0 0 )\n( (9)\nwhere, an-an-1= an-an-1\n|an-an-1| , etc. m\ud835\udc64\ud835\udc64 is a payload applying at the\nend-point and g is the gravity acceleration vector. Equilibrium in moments at Ui , (i=m+1, \u22ef, n-1), belonging to the first segment is (Sa,n-fa)(an-an-1)\u00d7(an-un)+\n(Sa\u0302,n-fa\u0302)(a\u0302n-a\u0302n-1)\u00d7(a\u0302n-un)+ (Sc,n-fc)(cn-cn-1)\u00d7(cn-un )+ (Sc\u0302,n-fc\u0302)(c\u0302n-c\u0302n-1)\u00d7(c\u0302n-un)+\nmp \u2211 (pk-ui) n\nk=i+1\n\u00d7g= ( 0 0 0 )\n(10)\n(10)\nwhere. fa, fa\u0302, fc, fc\u0302 are wire tensions, Sa,i, Sa\u0302,i, Sc,i, Sc\u0302,i , (i=m+1, \u22ef, n) are spring tensions of the ith unit. \u201c\u00d7\u201d means a cross product and \u201c|*|\u201d, means the modulus of a vector \u2217. \ud835\udc5a\ud835\udc5a\ud835\udc5d\ud835\udc5d is the mass of one unit including the plate, the rod and the universal joint.\nThe spring tensions are obtained as, Sa,i\uff1dk(L-|ai-ai-1|), Sa\u0302,i\uff1dk(L-|a\u0302i-a\u0302i-1|),\nSc,i=k(L-|ci-ci-1|), Sc\u0302,i=k(L-|c\u0302i-c\u0302i-1|),)\n(11)\nwith spring coefficient k. Equations (9) and (10) contain 3(nm) equations including 4(n-m) -1variables of the n-m universal joints angles \u03b8xi, \u03b8yi, \u03b8zi, (i=m+1, \u22ef, n) and slide length of plates \ud835\udc59\ud835\udc59\ud835\udc56\ud835\udc56 (i=m+1, \u22ef, n-1). Equilibrium in force at ith plate (i=m+1, \u22ef, n-1) is,\n[-Sa,i+1(ai+1-ai) +Sa,i(ai-ai-1)+-Sa\u0302,i+1(a\u0302i+1-a\u0302i)+ Sa\u0302,i(a\u0302i-a\u0302i-1)-Sc,i+1(ci+1-ci)+Sc,i(ci-ci-1)-Sc\u0302,i+1(c\u0302i+1-c\u0302i))\nSc\u0302,i(c\u0302i-c\u0302i-1)+(n-i)mpg ]\u2219 (pi-ui)=0\n(12)\n(12) provide n-m-1equations. Combined it with (9) and (10), we obtain 4(n-m)-1 equations, which suffices in number to solve for 4(n-m)-1 variables; \u03b8x,i, \u03b8y,i , \u03b8z,i (i=m+1,\u22ef,n) and li (i=m+1, \u22ef, n-1) for a given set of wire tensions fa, fa\u0302, fc, fc\u0302.\nEquilibrium in moments at Um , the universal joint located at the most distal position belonging to the second segment is\n-Sa, m+1(am+1-am)\u00d7(am-um)+(Sb,m-fb)(bm-bm-1)\u00d7(bm-um)\n-Sa\u0302,m+1(a\u0302m+1-a\u0302m)\u00d7(a\u0302m-um)+(Sb\u0302,m-fb\u0302) (b\u0302m-b\u0302m-1) \u00d7(b\u0302m-um) -Sc, m+1(cm+1-cm)\u00d7(cm-um)+(Sd,m-fd)(dm-dm-1)\u00d7(dm-um)\n-Sc\u0302,m+1(c\u0302m+1-c\u0302m)\u00d7(c\u0302m-um)+(Sd\u0302,m-fd\u0302) (d\u0302m-d\u0302m-1) \u00d7(d\u0302m-um)\n+ (mw(pn-um)+mp \u2211 (pk-um) n-1\nk=m+1\n) \u00d7g= ( 0 0 0 )\n(13)\nFor the second segment, we can derive similar equations as (10), (11) and (12) by replacing {ai, a\u0302i, ci, c\u0302i} with {bi, b\u0302i, di, d\u0302i}, {Sa,i, Sa\u0302,i, Sc,i, Sc\u0302,i} with {Sb,i, Sb\u0302,i, Sd,i, Sd\u0302,i} for i=1, \u22ef, m-1 in (10) and for i=1, \u22ef, m in (11) and (12).\nAs a result, we obtain 4m equations included by (13), which suffices in number to solve for 4m variables; \u03b8x,i, \u03b8y,i, \u03b8z,i and li (i=1, \u22ef,m) for a given set of wire tensions \ud835\udc53\ud835\udc53\ud835\udc4f\ud835\udc4f, \ud835\udc53\ud835\udc53?\u0302?\ud835\udc4f, \ud835\udc53\ud835\udc53\ud835\udc51\ud835\udc51, \ud835\udc53\ud835\udc53?\u0302?\ud835\udc51.\nWire tensions fa, fa\u0302, fc, fc\u0302, fb, fb\u0302, fd, fd\u0302. are determined according to 4 motors\u2019 angles \u03d5a, \u03d5b,\u03d5c,\u03d5d As\nfa=kp ( \u03bb (\u03d5p+\u03d5a) 2\u03c0 -nL+ \u2211|ai-ai-1| n\ni=1\n) ,\nfa\u0302=kp ( \u03bb (\u03d5p-\u03d5a) 2\u03c0 -nL+ \u2211|a\u0302i-a\u0302i-1| n\ni=1\n) ,\nfc=kp ( \u03bb (\u03d5p+\u03d5c) 2\u03c0 -nL+ \u2211|ci-ci-1| n\ni=1\n) ,\nfc\u0302=kp ( \u03bb (\u03d5p-\u03d5c) 2\u03c0 -nL+ \u2211|c\u0302i-c\u0302i-1| n\ni=1\n) ,\nfb=kp ( \u03bb (\u03d5p+\u03d5b) 2\u03c0 -nL+ \u2211|bi-bi-1| m\ni=1\n) ,\nfb\u0302=kp ( \u03bb (\u03d5p-\u03d5b) 2\u03c0 -nL+ \u2211|b\u0302i-b\u0302i-1| m\ni=1\n) ,\nfd=kp ( \u03bb (\u03d5p+\u03d5d) 2\u03c0 -nL+ \u2211|di-di-1| m\ni=1\n) ,\nfd\u0302=kp ( \u03bb (\u03d5p-\u03d5d) 2\u03c0 -nL+ \u2211|d\u0302i-d\u0302i-1| n\ni=1\n) ,\n(14)\nwhere, \ud835\udf19\ud835\udf19\ud835\udc5d\ud835\udc5d is a motor rotation angle to generate a pretension, \ud835\udf06\ud835\udf06 is a lead of the screw rod and \ud835\udc58\ud835\udc58\ud835\udc5d\ud835\udc5d is the spring constant of the pretension spring.\nAuthorized licensed use limited to: Auckland University of Technology. Downloaded on October 25,2020 at 08:37:21 UTC from IEEE Xplore. Restrictions apply.", + "222\nC. Inverse kinematic solution\nAccording to given set of variables \ud835\udf03\ud835\udf03\ud835\udc65\ud835\udc65,\ud835\udc56\ud835\udc56, \ud835\udf03\ud835\udf03\ud835\udc66\ud835\udc66,\ud835\udc56\ud835\udc56 , \ud835\udf03\ud835\udf03\ud835\udc67\ud835\udc67,\ud835\udc56\ud835\udc56 (\ud835\udc56\ud835\udc56 = 1, \u22ef , \ud835\udc5b\ud835\udc5b) and \ud835\udc59\ud835\udc59\ud835\udc56\ud835\udc56 (\ud835\udc56\ud835\udc56 = 1, \u22ef , \ud835\udc5b\ud835\udc5b \u2212 1), we calculate the end-point position by Eq. (6),\n(pn 1 ) =H0,n ( 0 0 ln 1 ) (in jn kn rn 0 0 0 1 ) ( 0 0 ln 1 ) = (knln+rn 1 )\n(15)\nTaking a total differentiation of pn=knln+rn with respect to \u03b8x,i, \u03b8y,i , \u03b8z,i (i= 1, \u22ef, n) and li (i=1, \u22ef, n-1) and also motor angles \u03d5a, \u03d5b, \u03d5c, \u03d5d,\n\u2206pn= \u2202pn \u2202v \u2206v+ \u2202pn \u2202\u03d5 \u2206\u03d5 (16)\nWhere, v=(\u03b8x1,\u03b8x2, \u22ef,\u03b8xn, \u03b8y1,\u03b8y2, \u22ef,\u03b8yn ,\u03b8z1,\u03b8z2, \u22ef,\u03b8zn, l1 , l2 , \u22ef,ln-1 )\n\u2208R4n-1 and \u03d5=(\u03d5a, \u03d5b, \u03d5c, \u03d5d ). \u2202pn \u2202v \u2208R3\u00d74n-1 and \u2202pn \u2202\u03d5 \u2208R3\u00d74. Whereas, let w=(w1, w2, \u22efw4n-1)T=04n-1 represents the 4n-1 equations provided by Eqs. (9)(10)(12)(13), which also includes \u03b8x,i, \u03b8y,i , \u03b8z,i (i=1, \u22ef, n) , li (i=1, \u22ef, n-1) and also motor angles \u03d5a, \u03d5b, \u03d5c, \u03d5d. Taking a total differentiation for w=04n-1 as well, we have,\n\u2206w= \u2202w \u2202v \u2206v+ \u2202w \u2202\u03d5 \u2206\u03d5=04n-1 (17)\nwhere, \u2202w \u2202v \u2208R(4n-1)\u00d7(4n-1) and \u2202w \u2202\u03d5 \u2208R(4n-1)\u00d74 . Since \u2202w \u2202v is a square matrix, we can solve (19) with respect to the vector \u2206\ud835\udc63\ud835\udc63 as,\n\u2206v=- (\u2202w\n\u2202v ) -1 \u2202w \u2202\u03d5 \u2206\u03d5 (18)\nSubstituting (18) into (16), we have\n\u2206pn=\u2202pn \u2202v (\u2202w \u2202v ) -1 \u2202w \u2202\u03d5 \u2206\u03d5+ \u2202pn \u2202\u03d5 \u2206\u03d5=\n( \u2202pn \u2202\u03d5 - \u2202pn \u2202v (\u2202w \u2202v ) -1 \u2202w \u2202\u03d5 )\u2206\u03d5=J\u2206\u03d5\n, which can be solved for \u2206\ud835\udf19\ud835\udf19, by using a generalized inverse of the Jacobian J\u2208R3\u00d74\n\u2206\u03d5=J\u2020\u2206pn+P\u22a5(J)\u03a8 (19)\nwhere, J\u2020\u2208R4\u00d73 is a generalized inverse of \ud835\udc71\ud835\udc71 and P\u22a5(J)\u2208R4\u00d74 is a null projection operator of J, and \u2206\u03d5N\u2208R4 is a correction of \ud835\udf53\ud835\udf53 so as to minimize a positive scalar potential \ud835\udf11\ud835\udf11 by making use of a redundant actuation.\nWe use J\u2020=JT(J JT)-1 and P\u22a5(J)=I-J\u2020J. Eq.(19) provides a variation of motor angles \u2206\ud835\udf53\ud835\udf53 for a given position and direction variation \u2206pn . Applying the Euler method, we have the following variational equation, \u03c6+(\u2202\u03c6/\u2202\u03d5)\u2206\u03d5N=0\n,which is solved by\n\u2206\u03d5N=\u03c6 (\u2202\u03c6/\u2202\u03d5)(\u2202\u03c6/\u2202\u03d5)T (\u2202\u03c6 \u2202\u03d5)\nT\n( 20) As a candidate of \u03c6 , we take \u03c6=knz\n2 , where, k\ud835\udc5b\ud835\udc5b\ud835\udc67\ud835\udc67 is the z component of kn: the unit vector of the end-point orienting an axial direction. It means that the axial direction the endpoint takes on a horizontal plain as far as possible while keeping a designated position.\nAuthorized licensed use limited to: Auckland University of Technology. Downloaded on October 25,2020 at 08:37:21 UTC from IEEE Xplore. Restrictions apply." + ] + }, + { + "image_filename": "designv11_71_0000884_s12650-020-00640-3-Figure2-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0000884_s12650-020-00640-3-Figure2-1.png", + "caption": "Fig. 2 a Projectile and b nozzle geometry", + "texts": [ + " The pressure relief section had a length of 38.5 cm, which was designed to diminish the blast wave in front of the projectile. The pressure relief section had four slots; each slot is 4 mm wide and 345 mm long. The test chamber was a square tank of 350 9 350 mm in width and 590 mm in length with two polymethyl methacrylate (PMMA) windows on two sides for visualization. The projectile was made of polymethyl methacrylate (PMMA) and is a cylindrical shape with a diameter of 15 mm, length of 8 mm and a weight of 0.92 g as shown in Fig. 2a. The HSSPG was employed to generate the high-speed liquid jet velocity of 550 to 2290 m/s in each gunpowder weights. The nozzle with exit diameter of 0.7 mm (Fig. 2b) that was connected to the pressure relief section was made of mid-steel. Gunpowder of 5 g was used in this study, which could launch the projectile to a speed of about 952 \u00b1 4% m/s. Please note that at the smaller size of the nozzle, its friction loss could affect the jet formation or shock phenomena. However, this exit diameter of the nozzle is used with reference from past research (Pianthong 2002; Milton and Pianthong 2005b; Pianthong et al. 2002), where the maximum jet velocity can be reached" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0000761_s00202-020-00962-3-Figure1-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0000761_s00202-020-00962-3-Figure1-1.png", + "caption": "Fig. 1 Dual-airgap spoke-type PMVM model", + "texts": [ + " The performance parameters such as average torque, torque ripple, torque density and efficiency in the dual-airgap PMVMs are analyzed using different rotor and stator pole combinations, while the model reaches the maximum torque and efficiency. The outer and inner dimensions of all models for different pole ratios are kept constant. 2D-FEM analysis is used to analyze the effect of pole ratio on electromagnetic torque performance of dual-airgap spoketype PMVM. The PMVM has the stator toothed-pole structure [21] to realize the rotor magnetic field modulation effect, and the windings are of conventional overlap type. The PMVM used in this paper is radial-type dual-airgap PMVM with spoke-type magnets in the rotor. Figure 1 shows the basic dual-airgap dual-stator winding (DDSW) model of PMVM with pole ratio 5 (12 stator slots/20 rotor poles). The basic dimensions of DDSW model are presented in Table 1. The DDSWmodel has ferrite spoke-type permanent magnets, and the magnets are magnetized circumferentially. The relationship between the number of rotor poles, stator poles and stator slots in a Vernier machine as presented in [21] is: \u00b1 Ps/2 Pr/2 \u2212 Ss (1) wherePs,Pr andSs are stator poles, rotor poles and stator slot, respectively" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0003499_muh-1404-6-Figure2-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0003499_muh-1404-6-Figure2-1.png", + "caption": "Figure 2. 3D model of a 3-axis open category milling machine tool with main dimensions of 1920 mm \u00d7 1790 mm \u00d7 1200 mm and weight of 4555 kg.", + "texts": [ + " The time-varied deformation in x and y directions generated from the analysis can be used later to evaluate the characteristics of the work pieces surface quality. The mechanical structure of a machine tool center can be considered to have the major contribution of its rigidity. It mainly consists of the column, bed, table, saddle, slider, and spindle head. When the cutting process is established on hard materials, the dynamic characteristics of the mechanical structure of a machine tool center become crucial. The CAD model such as that shown in Figure 2 is created using any of the commercial CAD software programs and imported to ANSYS, which is the analysis tool package used in this work. Before creating the FE model of the machine tool structure, small holes, chamfers, fillets, and other tiny details are ignored to simplify the model so that a high-quality mesh can be obtained with minimum computation time. Contacts are accurately defined to truly describe the actual relation in practice from the aspect of load transfer. For the mechanical structure, the contact between any bolted subsystems such as the column and bed or between any sliding subsystems such as the column and spindle head is defined as frictional with a user-defined coefficient of friction" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0002607_j.ifacol.2020.12.1281-Figure2-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0002607_j.ifacol.2020.12.1281-Figure2-1.png", + "caption": "Fig. 2. Left : Crazyflie 2.0. The spherical reflective markers used by the Motion Capture system (attached to the frame) and the UWB chip (with the marking \u201c1\u201d) can be seen. Right : Frame Coordinate System.", + "texts": [ + " Each drone also has a unique set of reflective markers so we can track their ground truth position with a motion capture system. The position data from ultra-wideband is fused with the data from on-board sensors with an Extended Kalman Filter (EKF). The filter is based on the papers Mueller et al. (2015) and Mueller et al. (2016), and was mostly implemented by the designers of the Crazyflie. The estimator code was modified to improve performance and to report all the system states to the controller Fig. 2. Left : Crazyflie 2.0. The spherical reflective markers used by the Motion Capture system (attached to the frame) and the UWB chip (with the marking \u201c1\u201d) can be seen. Right : Frame Coordinate System. 3.2 Dynamical Model The quadcopter dynamics has been widely studied before. For our controller synthesis we will consider the NewtonEuler equations such as the ones presented in Luukkonen (2011). Let p[m] denote the position of the Crazyflie in a Global Cartesian coordinate system, and \u03b7[rad] be its Euler angles" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0002923_9781119633365-Figure1.20-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0002923_9781119633365-Figure1.20-1.png", + "caption": "Figure 1.20 (a) Switched-capacitor and (b) switched-inductor cells.", + "texts": [ + " Although this approach can 1 Introduction20 derive new PWM converters and is straightforward, it is still tedious and lacks of mechanism to explain the converters with identical transfer ratio but with different configurations. Again, it is based on a cell level. Based on a cell level, another approach to developing new converters with higher step\u2010up and step\u2010down voltage ratios by introducing switched\u2010capacitor or switched\u2010inductor cells to the PWM converters shown in Figures\u00a01.7 and 1.8 was proposed. Typical switched\u2010capacitor and switched\u2010inductor cells are shown in Figure\u00a0 1.20, and their derived converters have been shown in Figure\u00a0 1.10. Applications of this approach are quite limited, and the chance of deriving new converters is highly depending on experience. Otherwise, it might need many trial and errors. When inserting a cell into a PWM converter, one has to use volt\u2010second balance principle to verify if the converter is valid. This approach still needs a lot of ground work to derive a valid converter. Another synthesis approach based on a converter cell concept, 1L and 2L1C, was proposed" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0000148_mrs.2019.8901063-Figure2-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0000148_mrs.2019.8901063-Figure2-1.png", + "caption": "Fig. 2: a) Motion enabled by module rotation and then translation to the adjacent connection point. b) Motion enabled by simultaneous module rotation and translation to the adjacent connection point. c) DONUt module configurations and their difficulties when populated with EPMs. d) Top view of a DONUt module with the measured IR transceiver field-ofview.", + "texts": [ + " To switch \u2019off\u2019 an SEP, we charge the bank for a shorter time, dumping less energy into the coil such that only about half of the dipoles are switched in a particular direction within the low coercivity magnet. The ability to control both the polarity and the on/off-state of the SEP allows a greater number of possible module connection topologies compared to those feasible with Electro-Permanent Magnets (EPMs) as used in previous work [7], [2] which can only switch between an on/off-state. This concept is illustrated in Fig. 2c. While testing the motion of the modules we found that each module is capable of three separate types of movement. The first entails a module rotating and translating across a surface to an adjacent connection point (Fig. 2a). We conducted reliability tests and found that a module can perform this type of motion with a success rate of 48 out of 50 times. To enable this motion, a module i conducts the following sequence of polarity switches in four adjoining SEPs: 1) [O-N-S-O], 2) [O-N-N-O], 3) [O-S-N-O], while the opposing module j remains fixed at [O-S-S-O]. O, N, and S correspond to off, north, and south, respectively. It is important to note that this requires only one module to switch its SEPs, and therefore does not necessitate precise timing coordination between the modules. The second type of movement involves a module separating from another module, rotating in the spot, and reconnecting to the module at the same location. This type of motion would allow a collective to keep a configuration while rotating a module to, e.g., improve its field of view. The third type of motion involves simultaneous rotation and translation, such that a module rolls about the perimeter of another module from one connection point to another (Fig. 2b). The following sequence is executed to move module i in this fashion: 1) [S-O-N-N], 2) [S-O-O-N], 3) [S-N-O-N], 4) [N-S-S-N], and 5) [S-N-S-N]; while module j keeps a fixed polarity [NS-S-N]. To roll about the perimeter of another module, the moving module must keep a constant distance between its center of mass and the point of contact with the adjacent module. This is only possible if the module simultaneously exhibits the kinetic energy needed to roll and translate to the new location. Due to the limited strength of the SEPs (caused by the design of the battery bank), this type of motion is currently only possible with a sparsely populated module, or with two full modules aided by an external force pressing them slightly together" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0003319_icuas.2015.7152288-Figure1-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0003319_icuas.2015.7152288-Figure1-1.png", + "caption": "Fig. 1. Formation geometry", + "texts": [ + " In [9], robustness of SDRE method to erroneous measurements was also demonstrated. In the following section leader-follower formation equations are derived. Then, guidance of the follower using fuzzy logic control approach is given. In section III, flight control design with SDRE method is presented. The simulations results and concluding remarks are given in sections IV and V respectively. 978-1-4799-6009-5/15/$31.00 \u00a92015 IEEE 167 In this part, formation control method is explained. First, geometric relations of formation flight are derived, as shown in Fig. 1. The inertial frame and leader fixed frame shown in the figure are separated by the heading angle, \u03c8L, only. Thus, the z-axes of the Inertial Frame and Leader Fixed Frame are always parallel to each other. The vector from the follower \u201dF\u201d to the reference point \u201dR\u201d to be followed, \u2212\u2192 FR, given in leader fixed frame, may be expressed as follow: \u2212\u2192 FR|L = \u2212\u2192 R |L \u2212 \u2212\u2192 F |L (1) \u2212\u2192 R |L = \u2212\u2192 L |L + [\u2212a b c]T |L (2) In equation (2), c, is a distance from leader to reference point \u201dR\u201d towards z axis. Normally all positions may be expressed in the inertial frame, \u2212\u2192 L |I = [XL YL \u2212HL]T (3) \u2212\u2192 F |I = [XF YF \u2212HF ]T (4) and transformed to the leader frame using the leader heading angle, to obtain, \u2212\u2192 FR|L" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0002321_012009-Figure5-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0002321_012009-Figure5-1.png", + "caption": "Figure 5. The process of sampling the gaps at the time of passage of TDC and BDC, respectively.", + "texts": [ + " Computer-Aided Technologies in Applied Mathematics Journal of Physics: Conference Series 1680 (2020) 012009 IOP Publishing doi:10.1088/1742-6596/1680/1/012009 Figure 4a. Vibration spectrum of the piston compressor in time interval 0 \u2013 100 mc. Figure 4b. Vibration spectrum of the piston compressor in time interval 101 \u2013 200 mc. Figure 4c. Vibration spectrum of the piston compressor in time interval 201 \u2013 300 mc. One possible way to solve this problem is to identify defects by monitoring the uneven rotation of the shaft of the piston machine. So in a piston compressor at the moment of reaching the top dead center (Figure 5), when the piston body is pushed out, the angular velocity of the shaft at the moment of sampling the gaps in the joints of the crank mechanism increases, and at the time of the occurrence of an impact load, the speed sharply decreases. At the bottom dead center, at the time of the return of the piston one can also observe an abrupt change in the angular velocity of the shaft. If a defect occurs in articulated moving parts, for example, a crankshaft with a connecting rod or a connecting rod with a piston, the change in angular velocity will be greater [3]" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0002861_978-3-030-55889-5-Figure9.5-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0002861_978-3-030-55889-5-Figure9.5-1.png", + "caption": "Fig. 9.5 The market woman checks the equilibrium with the principle of virtual displacements", + "texts": [ + " In practice, of course, the finite elements never reach the limit h \u2192 0 and only a finite number of tests is performed, since there are only a finite number of shape functions \u03d5i on a mesh. The finite element method can be seen as a method of replacing a functional J (\u03b4u) = (p, \u03b4u) with a functional Jh(\u03b4u) = (ph, \u03b4u), or, if you have infinite patience, h \u2192 0, with a sequence of functionals Jh(\u03b4u) = (ph, \u03b4u). The practical significance of the concept of a weak solution becomes apparent when we watch a market woman, see Fig. 9.5, since she also draws conclusions indirectly. She has to solve the equation Pl \u00b7 hl = Pr \u00b7 hr , (9.76) which means, as she knows, that at every turn \u03b4\u03d5 of the balance beam, the work on the left and right side are the same Pl \u00b7 hl = Pr \u00b7 hr \u21d2 Pl \u00b7 hl \u00b7 tan tan \u03b4\u03d5 = Pr \u00b7 hr \u00b7 tan \u03b4\u03d5 , (9.77) and so she concludes, by wiggling the balance, indirectly, arguing \u201cbackwards\u201d Pl \u00b7 hl = Pr \u00b7 hr \u21d0 Pl \u00b7 hl \u00b7 tan \u03b4\u03d5 = Pr \u00b7 hr \u00b7 tan \u03b4\u03d5 . (9.78) The same is done by the toolmaker, who rolls a cylinder back and forth with his fingers, see Fig" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0001625_s11694-020-00601-2-Figure1-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0001625_s11694-020-00601-2-Figure1-1.png", + "caption": "Fig. 1 a Experimental structure of the sensor. b Sensor block diagram", + "texts": [ + " The total flavonoids (TF) were determined using the method previously described by Patil et\u00a0al. [15] with some modifications. Two grams of samples were homogenized in 25\u00a0mL of the extractor solution (isopropanol: acetic acid: water. 8:1:1,v:v:v) and the absorbance was measured at 370\u00a0nm in a spectrophotometer (Kasuaki UV/VIS, model IL-592-BI). Total flavonoids were calculated by using a standard curve of quercetin at concentrations from 0 to 160\u00a0\u00b5g\u00a0mL\u22121, and the results were expressed as mg of quercetin 100\u00a0g\u22121 FW. The structure of the experimental optical sensor is shown in Fig.\u00a01a: (1) LED; (2) LED holder; (3) Optical filter holder; (4) Optical filter; (5) Cuvette holder; (6) 3.5\u00a0mL Cuvette; (7) Photodiode. Figure\u00a01b shows a block diagram of the prototype sensor and the data acquisition system which includes an electronic circuit composed of instrumentation amplifiers with low noise and low offset. A specific LED and Optical filter set was used for each parameter as shown in Table\u00a02. To validate the sensor the spectrophotometer Kazuaki UV/VIS, model IL-592-BI. was used, following the same methodology as described above. After analyzing the standards and the samples in the spectrophotometer using the specific wavelength for each bioactive, the standards and the samples were placed in the cuvette and excited with a specific LED wavelength (Table\u00a02) for each parameter" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0002346_cce50788.2020.9299140-Figure1-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0002346_cce50788.2020.9299140-Figure1-1.png", + "caption": "Fig. 1: quadrotor with pay load", + "texts": [ + " The design is different from other techniques based on the differential flatness concept and leads to a simpler controller which achieves trajectory tracking in the presence of disturbances with a minimum swing od the payload. This article is organized as follows. Section II presents the model of a quadrotor with a hanged payload and its linear approximation together with the parametrized dynamics based on differential flatness. Section III describes the control strategy using sliding mode techniques. In Section IV some simulation results are given. Finally, conclusions are drawn in Section V. As shown in Fig. 1 a quadrotor slung load system is considered, which consists of a quadrotor UAV, a rigid cable and a suspended payload. The following assumptions are made in order to obtain the dynamic model of the system: 978-1-7281-8987-1/20/$31.00 \u00a92020 IEEE Authorized licensed use limited to: University of Newcastle. Downloaded on June 16,2021 at 07:00:10 UTC from IEEE Xplore. Restrictions apply. 1) The quadrotor is symmetrical. 2) The cable attachment point is at the UAV center of mass. 3) The cable is mass-less, inelastic and rigid" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0002338_1369433220971728-Figure5-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0002338_1369433220971728-Figure5-1.png", + "caption": "Figure 5. Normal and tangential contact force components.", + "texts": [ + " Additionally, mk and ms are the kinetic friction coefficient and static friction coefficient, respectively. The specific values of the above parameters can refer to Table 1. After some further mathematical calculations, the magnitude of contact force and its orientation are obtained as: QC = Kdn\u00bd1+ 8(1 cr) 5cr _d _d ( ) ; g =a+ e e=tan 1 msign vt\u00f0 \u00de\u00f0 \u00de ( , \u00f016\u00de where e is the angle between the contact force vector QC and the normal direction n, a is the angle between the clearance vector and the axis Mx, while the relevant forces, directions and angles are shown in Figure 5. K is calculated by adding the resulting tangential element from equation (13) to the normal component of the contact force from equation (10), which can be expressed as: K = ffiffiffiffiffiffiffiffiffiffiffiffiffiffi 1+m2 p Kg \u00f017\u00de Dynamic model and control in planar deployable structure Dynamic model of planar scissor deployable structure including clearance The examined scissor deployable structure consists of four bars and a slider, in which bar 1 is the active bar and a constant angle velocity is applied" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0003880_scis-isis.2014.7044812-Figure14-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0003880_scis-isis.2014.7044812-Figure14-1.png", + "caption": "Fig. 14. Force representation for symmetrical obstacle distribution", + "texts": [ + " MATHEMATICAL EXPLANATION OF SOLVING DEAD- The proposed APF algorithm does not consist of local minimum which causes dead-lock problem for the mobile robot. Simply, local minimum point exists when the robot, obstacle and the goal points are collinear or for a symmetrical distribution of the obstacles around the line which connects the robot and the goal. In a single obstacle case, the velocity of the robot should be collinear according to the set of equations of APF. The motion direction of the robot should be the same as shown in Fig.14. If there are several obstacles in the environment and they are symmetrical within the detectable range of LRF, this can be happened, but it highly depends on the previously detected obstacle distribution and the velocity direction. Dead-lock can be happened in the situation described above and for the most possible orientation describes in Fig. 14 and the governing equation of the modified APF is given in (11). of the equation represents the distance between the robot and the obstacle, and is the measurable range of the LRF. For the situation descried in Fig. 14, even the attractive force is cancelled out by the repulsive force component , there is a new repulsive force component in existence perpendicular to the original repulsive force. Therefore the total potential force never becomes zero in the proposed APF. As the is existing, this force is used to bring the robot away from its current path in order to avoid the collision or dead-locking. (11) VI. CONCLUSION One of the problems associated with the APF based path planning of mobile robot which is known as dead-lock, a result of local minimum has been discussed in this paper" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0000256_icems.2019.8921568-Figure1-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0000256_icems.2019.8921568-Figure1-1.png", + "caption": "Fig. 1. Complete structure of the motor", + "texts": [ + " Then the characteristics of demagnetization of AlNiCo are introduced, and the simplified mathematical model of AlNiCo is established to explain the principle of field regulation. In order to obtain the largest magnetic field linear control range, the combination scheme of the two permanent magnet thicknesses is analyzed by finite element method. Finally, the rule of influence of permanent magnet thickness on permanent magnet working point, magnetic field regulation ability, gap flux density linear control range, no-load back EMF and rated load torque is summarized. II. STRUCTURE OF THE FVPSCPM The structure of FVPSCPM is shown in Fig. 1 and Fig. 2. Memory permanent magnet material AlNiCo and high performance permanent magnet material NdFeB are used in This work was supported by Major Project of Application Technology Research and Development Plan of Heilongjiang Province of China under Grant GA15A401. 978-1-7281-3398-0/19/$31.00 \u00a92019 IEEE series as excitation part to ensure that the motor has good regulation ability of gap flux density as well as high energy density. The rotor adopts claw pole shape. Since the presence of magnetic regulation coil, in order to get rid of brush and slip ring, a partitioned stator structure is adopted, and the excitation structure is separated from claw pole rotor as partitioned inner stator" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0001892_012007-Figure1-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0001892_012007-Figure1-1.png", + "caption": "Figure 1. (a) RFPM engine, (b) AFPM engine [4]", + "texts": [ + " This software is used and focused on analysing the magnetic flux density. To analyse the magnetic field working in the motor [3]. ICECME 2019 IOP Conf. Series: Materials Science and Engineering 931 (2020) 012007 IOP Publishing doi:10.1088/1757-899X/931/1/012007 Nowadays, many areas need more things like power density, low weight, low material, low cost, excellent cooling, and ventilation. Axial flux motors or often abbreviated as AFPM are motors that have good criteria to be used in the future. AFPM is more compact than the RFPM engine shown in Figure 1; it is recognized as having the ability to get more power density than an RF engine [4]. AFPM machines are recognized that can get better power density than RFPM machines. Therefore AFPM can be applied because it is efficient and effective, and its capabilities are superior to RFPM [5] To assess the potential of a multi-stator engine, a double stator AFPM engine is considered. This machine has two disks (stator) that support entanglement (Figure 2). On the rotor, parts are connected three magnetic disks which are connected to each other in order to act as a driving shaft in the drive motor system which will later be connected to the wheels" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0001940_rpj-01-2020-0010-Figure5-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0001940_rpj-01-2020-0010-Figure5-1.png", + "caption": "Figure 5 (a) Additively manufactured maraging steel cooling test artifact with allocated temperature sensors. (b) Transparent view of the cooling test artifact showcasing the internal channel", + "texts": [ + " The second case was an example of a conformal cooling channel design. This design was chosen to demonstrate the actual applicability of the methodology in an industrial setting. Both channel designs had a nominal channel diameter of 2mm. To validate the application of the proposed methodology a cooling test artifact, containing an internal helical channel with a nominal diameter of 3mm, was additively manufactured in Maraging Steel 300 using a Concept Laser M2 system. The component and internal channel design can be seen in Figure 5. The test artifact is placed within an experimental setup, wherein a heating element is placed just below the artifact while water is pumped in through the channel. The experimental setup allows a controlled mass flow rate of water through the channel, while the water is allowed to drain into an open container at ambient pressure (i.e. 1 atm). The temperature of the water at the inlet and outlet, as well as the temperature at six different locations (as seen in Figure 5) on the test artifact aremeasured during the experiment. The validation experiment is carried out in two stages. In the first stage, the heating element is turned ON for 6,600 s and then turned OFF for 6,000 s. The temperature results from this experimental stage allow calibration of the heat transfer coefficients between the test setup and the ambient and provide a reference thermal response of the test setup. In the second stage, water is pumped into the channel at a flow rate of 0.493L/min and subsequently, the heating element is turned ON for 2,700 s and turnedOFF for 2,700 s" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0001019_j.precisioneng.2020.04.003-Figure14-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0001019_j.precisioneng.2020.04.003-Figure14-1.png", + "caption": "Fig. 14. Schematic of constraint plane 1\u20132 showing: (a) components of the sliding direction; and (b) forces at the engaged constraint.", + "texts": [ + " 7) All the geometry defined above applies for the coupling virtually close to full engagement so that small angle approximations may be used and the shapes of contact surfaces do not make a substantial difference. The derivation of (1) begins by identifying the one-degree-of-freedom motion of the coupling and the sliding directions at the five engaged constraints. Two axes of rotation admissible for constraints 3\u20136 are apparent in Fig. 13: axis A is out of the page and identical to the instantaneous center demonstrated in Fig. 7; and axis B passes through the instantaneous centers for constraints 3\u20134 and 5\u20136. A fifth constraint at 1 or 2 ties rotation about axes A and B to a fixed ratio. Fig. 14-a shows the fifth constraint undergoing a unit displacement along its sliding direction and its components about axes A and B. Combining this with lever arms from Fig. 13, Equations (2) and (3) give corresponding rotations about axes A and B. \u03b8A\u00bc cos \u03b11 2 3 R (2) \u03b8B\u00bc 2 sin \u03b11 2 3 R (3) L.C. Hale Precision Engineering 64 (2020) 200\u2013209 The sliding directions at constraints 3\u20136 are less apparent, involving components in the plane of the vee and out. Upon careful study, constraints 3\u20136 have slightly different sliding directions and resultant forces", + " \u03b4A\u00bc ffiffiffiffi 3 p R \u03b8A \u00bc cos \u03b11 2 ffiffiffi 3 p (4) \u03b4B\u00bc ffiffiffi 3 p 2 r \u03b8B \u00bc \u03c1 sin \u03b11 2 ffiffiffi 3 p (5) \u03b4\u00bc cos \u03b11 2 ffiffiffi 3 p ffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi 1\u00fe \u00f0\u03c1 tan \u03b11 2\u00de 2 q (6) Note as in the manner of \u03c9, all the forces are dimensionless having been normalized by the total load on the moving body. In Equation (7), the friction force along the sliding direction depends on the CoF, the weight proportion that constraints 3\u20136 carry and a wedge action due to the inclination angle. Referring to Fig. 14-b for a moment, the forces f A and f B at constraint 1 or 2 arise from friction at constraints 3\u20136 due to rotations about axes A and B, respectively, and are found by summing moments about these axes. Equation (8) expresses the out-of-plane component of friction force through a lever ratio about axis A, and similarly Equation (9) expresses the in-plane component about axis B. f \u00bc \u03bc3 6 \u00f01 \u03c9\u00de cos \u03b13 6 (7) fA\u00bc ffiffiffi 3 p R 3 R f \u03b4A \u03b4 \u00bc \u03bc3 6 \u00f01 \u03c9\u00de ffiffiffi 3 p cos \u03b13 6 1 ffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi 1\u00fe \u00f0\u03c1 tan \u03b11 2\u00de 2 q (8) fB\u00bc ffiffiffi 3 p =2 r 3 =2 R f \u03b4B \u03b4 \u00bc \u03bc3 6 \u00f01 \u03c9\u00de ffiffiffi 3 p cos \u03b13 6 \u03c12 tan \u03b11 2 ffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi 1\u00fe \u00f0\u03c1 tan \u03b11 2\u00de 2 q (9) Returning to Fig. 14-b, the balance of forces in this plane may be expressed in two orthogonal directions such as horizontal and vertical, or along and perpendicular to the sliding direction. Using the former yields Equations (10) and (11) by inspection. Dividing (11) by (12) eliminates the normal force giving Equation (12). Now it is a simple matter to substitute (8) and (9) into (12), and after some manipulation, Equation (13) expresses a slightly more general form of (1) containing \u03bc1-2 and \u03bc3-6. fA\u00bcN\u00f0sin \u03b11 2 \u03bc1 2 cos \u03b11 2\u00de (10) \u03c9 fB\u00bcN\u00f0cos \u03b11 2\u00fe \u03bc1 2 sin \u03b11 2\u00de (11) fA \u03c9 fB \u00bc sin \u03b11 2 \u03bc1 2 cos \u03b11 2 cos \u03b11 2 \u00fe \u03bc1 2 sin \u03b11 2 (12) sin \u03b11 2 \u03bc1 2 cos \u03b11 2 cos \u03b11 2 \u00fe \u03bc1 2 sin \u03b11 2 \u00bc \u03bc3 6ffiffi 3 p cos \u03b13 6 \u03c9 ffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi 1\u00fe\u00f0\u03c1 tan \u03b11 2\u00de 2 p 1 \u03c9 \u03bc3 6 \u03c12 tan \u03b11 2ffiffi 3 p cos \u03b13 6 (13) For the test with antifriction constraints 3\u20136, the right side of (13) is effectively zero and the limiting CoF \u03bc1-2 is easily found to be the tangent of the L" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0003778_1.4030163-Figure5-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0003778_1.4030163-Figure5-1.png", + "caption": "Fig. 5 Possible kinematic states at initial configuration with interval of 60:01m. Example additional patterns due to the interval are shown in (b) and (c). (a) Superimpose of 25 patterns, (b) four taut-wire pattern, and (c) three taut-wire pattern.", + "texts": [ + " 4(d)\u20134(f) are actually two taut-wire patterns and only the wires 2 and 5 take the tension. In a practical case that there exists some asymmetry, however, such patterns shown in Figs. 4(b), 4(d), and 4(e) can actually be possible with nonzero tension values for all the wires expressed in solid line. In the case that the uncertainty is taken into account in the form of the interval of 60.01 m for the wire lengths and the platform mass center position, the number of patterns of possible tension states becomes 25. Figure 5(a) shows the superimpose of all of the 25 patterns. The result is obtained by means of the incremental forward kinematics calculation with the interval, based on Dq \u00bc 0 and Dc\u0302 \u00bc 0, from the configuration shown in Fig. 4. Figures 5(b) and 5(c) show some of the additional patterns of tension states due to the introduction of the interval. These patterns could not be possible without interval; for example, the configuration shown in Fig. 5(c) is corresponding to an asymmetry condition where the wires 1, 3, and 5 are slightly longer and the wires 2, 4, and 6 are slightly shorter. Their representative wire lengths and platform mass center position referring to the platform-fixed coordinate are the same as the case without interval; however, their position and orientation of the platform do not coincide with the patterns shown in Fig. 4. This is because the placement of taut wires in these patterns is actually asymmetric. Figure 6 shows two example results of incremental forward kinematic calculation from the initial configuration shown in Fig. 5. Displacement of the platform mass center with no change in wire lengths is conducted and the results are shown in Fig. 6(a). By displacing the mass center 0.4 m right from the geometrical center of the platform, the configuration shown in Fig. 6(a1) is obtained; this is a superimpose of four possible patterns. Displacing the 041021-8 / Vol. 7, NOVEMBER 2015 Transactions of the ASME Downloaded From: http://mechanismsrobotics.asmedigitalcollection.asme.org/ on 02/01/2016 Terms of Use: http://www", + "org/about-asme/terms-of-use mass center rightward another 0.1 m, we obtain the configuration shown in Fig. 6(a2); this is a superimpose of two possible patterns. As indicated in these results, the number of possible configurations decreases in accordance with the asymmetry of the system in general. By displacing the mass center in another direction, that is, 0.5 m left from the geometrical center, the configuration shown in Fig. 6(a3) is obtained as a matter of course through the configuration same as shown in Fig. 5(a). Table 3 gives the wire tension values of the two possible kinematic states shown in Fig. 6(a2). Although they have the same pattern of taut and slack wires, the wire tension values are significantly different. This is a typical example of the case that two states are essentially in different modes of equilibrium conditions under the same tension pattern. Shortening the central two wires, 2 and 5, by 0.3 m from the configuration shown in Fig. 5, we obtain the configuration shown in Fig. 6(b1); this is a superimpose of five patterns. It typically illustrates the possible kinematic states obtained by means of the approach based on the many-worlds interpretation. Displacing the mass center 0.5 m rightward, the configuration shown in Fig. 6(b2) is obtained; this is a superimpose of four right-leaning patterns. The left-leaning and flat patterns shown in Fig. 6(b1) no longer exist, because they cannot attain any succeeding possible kinematic states at a certain incremental step due to the termination explained in Sec", + " The introduced unification parameters aUx and aUh represent a kind of degree of resolution of the position and orientation of the platform, based on which we deal with the distributed kinematic states. Figure 8 shows typical examples of intermediate kinematic states based on unification parameters aUx and aUh of different values; the possible kinematic states at 10th, 20th, and 30th steps are shown in the case of attaining the configuration shown in Fig. 6(b1) from the initial configuration shown in Fig. 5(a) in 100 incremental steps, corresponding to aUx \u00bc aUh \u00bc 0:4 and aUx \u00bc aUh \u00bc 0:1. The influence of these parameters on the number of possible kinematic states is significant as shown in Figs. 8(a1) and 8(b1); however, the possible kinematic states in both cases immediately converge to the same patterns in accordance with the remoteness from the bifurcation point as shown in Figs. 8(a3) and 8(b3). At the end of the 100 incremental steps, the obtained result is the configuration shown in Fig. 6(b1) in the both cases accordingly", + " 8(a), we assign a value on the large side to the number of possibilities, that is, PS\u00bc 30. Now we can estimate, quite roughly, the computational cost in a single incremental step, CS \u00bc 960CI , where CI is the computational cost of the inverse calculation of matrix of 12 12 at the maximum. The estimated computational cost is considerable, but still fairly practical. As a matter of fact, the kinematics calculation to obtain the configuration shown in Fig. 6(b1) from the initial configuration shown in Fig. 5 in 100 incremental steps takes 2.6 s by means of a typical desktop computer; a realtime calculation can be possible in the case of a quasi-static motion. We have discussed the incremental forward kinematics of wiresuspended parallel mechanical system. The approach has been developed based on the two basic equations as well as their incremental relations, which are the geometrical and equilibrium conditions in a general form. On the basis of the generality, the developed approach can deal with the kinematics of this type of mechanical system of any number and any placement of the wires" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0002989_978-3-642-16214-5_9-Figure9.27-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0002989_978-3-642-16214-5_9-Figure9.27-1.png", + "caption": "Fig. 9.27. Eaton LF (Low Force) synchronizer of the Eaton S Series. 1 Synchronizer body;", + "texts": [ + " The LF Synchronizer (LF = Low Force) is used in multi-range transmissions with two countershafts of the Eaton S Series for simultaneous shifting of idler gears (see Figure 12.38). All synchronizers of this three-range transmission are arranged on the central shaft. An essential aim of this LF synchronizer is reducing the shifting force and increasing the speed of gear change. This is achieved by converting the rotation force resulting from the synchronizing process into an axial force. This axial force strengthens contact between the synchronizer ring and the synchronizer body of the gearwheel to be shifted. Figure 9.27 shows the structure of this synchronizer. At the beginning of the shifting process, the driver moves the gearshift lever in the direction of the new gear position. In this way, the synchronizer plate 3 brings the synchronizer ring 2 into frictional contact with the synchronizer body 1 of the new gearwheel 6 by means of the sliding sleeve 8 and the axially shiftable pin 10. 2 synchronizer ring; 3 synchronizer plate; 4 preload mechanism; 5 ramps in the spline shaft profile; 6 gearwheel; 7 main shaft; 8 sliding sleeve; 9 selector teeth; 10 pin This motion preloads the system by means of the spring-loaded preloadmechanism 4" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0000451_hsi47298.2019.8942634-Figure4-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0000451_hsi47298.2019.8942634-Figure4-1.png", + "caption": "Fig. 4. Coordination", + "texts": [ + " With a full-charged battery, the user can use the proposed walker for four hours. The total amount of energy in the battery is 41 Wh, which is less than 160 Wh, thereby complying with the airplane safety standard set by International Civil Aviation Organization (ICAO). Thus, our device can be brought onto airplanes as well. III. COG POSITION ESTIMATION A. Body Shape Detection For estimating the position of its user\u2019s COG, our proposed walker measures a foot position and a waist position. Our walker has two laser range scanners as Fig.4(a). LRS1 is attached on the height which can measure its user\u2019s upper body, thigh and lower thigh. Its height is y=810[mm]. LRS2 is attached on the height of the ankle and y=150[mm]. (In this paper, we use ankle as the narrow section of the leg joining the foot.) 1) Estimating a foot position For estimating a foot\u2019s position, our walker uses LRS2 as Fig. 4(b). We uses UTM30LX laser scanner (HOKUYO AUTOMATIC CO., LTD). It has 270[deg] measuring range with 4[point/deg] and our walker uses 74[deg] measuring range as Fig.4(b). Thus, we uses about 300 measuring point for foot detection. 232 Authorized licensed use limited to: University of Exeter. Downloaded on June 29,2020 at 10:30:30 UTC from IEEE Xplore. Restrictions apply. For easy calculation for a built-in small computer, our walker detects a foot position using simple algorithm as follows; If the measuring position 2, ,i i lrs ix y zP by ith laser line fulfills (1), our walker detects iP and 1iP shows same obstacle. thresholddist is 0.1[mm] experimentally and 2lrsy is the position of LRS2 (y=150[mm])", + " If our system detects two objects, our system evaluates first object (j=1) as a left foot and second object (j=2) as a right foot. If our system detects more than two objects, our system selects two objects which has larger value in (2), and evaluates each object as feet. Table 1 shows the measuring error comparing with the result by motion capture system. (a) Kicking a left foot (b) Kicking a right foot Fig. 5. Mesuring position of each foot by LRS2. TABLE I. FOOT POSITION MESURING ERROR 2) Estimating a waist position For estimating a waist position, our walker uses LRS1 as Fig. 4(b). Our walker uses 50[deg] range for detecting its user\u2019s upper body as Fig. 6(a). Our proposed walker evaluates the intersection point between an upper body straight line and a thigh straight line as a waist position. Our walker detects a waist position as follows; LRS1 measures the position information as 1, ,i lrs i ix y zP by ith laser line. 1lrsx is the position of LRS1 (x=0[mm]). If we assume , 0n n m P is a waist position, an upper body straight line and a thigh straight line as (5) and (6)" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0000823_032122-Figure3-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0000823_032122-Figure3-1.png", + "caption": "Figure 3. The principle diagram of bisector plane positioning", + "texts": [], + "surrounding_texts": [ + "The UHF, transient ground voltage and ultrasonic comprehensive diagnostic technology are used to analyze the partial discharge type of the switchgear, and the UHF time difference method and the ESMA 2019 IOP Conf. Series: Earth and Environmental Science 440 (2020) 032122 IOP Publishing doi:10.1088/1755-1315/440/3/032122 bisector plane method are used to location the partial discharge source. Further explain the pattern diagnosis and fault location method of multi-source partial discharge." + ] + }, + { + "image_filename": "designv11_71_0000756_gcce46687.2019.9015395-Figure1-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0000756_gcce46687.2019.9015395-Figure1-1.png", + "caption": "Fig. 1. Shear force measuring device", + "texts": [ + " Therefore, in this study, a shear deformation and stress sensor system with piezoelectric polymer films embedded in soft material was developed. The designing principle and process will be firstly introduced, and the actual measurement of the shear force distribution in different underneath depth will be followed. It is also expected to be applied to the organic robotics, such as artificial skin of humanoid robot or prosthesis, or simulated organ for surgical practice and pathology. The basic structure of shear stress measuring system used in this study is shown in Fig. 1 (a). The sensing elements of PVDF are fixed perpendicularly to the surface and embedded in the soft substrate material. PVDF, which was manufactured in the form of micro film, was applied as sensing element. The PVDF films embedded in the soft material is deformed like a beam along with the deformation of the substrate material. The electric charges qi induced by piezoelectric property vary along with different local stress [6]. Based on the idea that the sensor films are arranged inside of soft substrate with different embedded lengths as illustrated in of Fig. 1 (b), the three-dimensional distribution of stress can be obtained. PVDF is usually attached to flexible plate such as PET sheet, whose rigidity decides the PVDF\u2019s deformation behavior. Thus, the performances of 3 kinds of sensor elements with 50, 25, and 16 \u00b5m thickness of PET sheets were compared. As a result, the thinnest one, whose flexibility was highest, was selected because of the highest sensitivity. In this paper, a kind of artificial skin (HITOHADA GEL, EXSEAL Co., Ltd., Japan) with ASKER-C hardness of 0 was utilized as the substrate material. Using this substrate and PVDF film, a testbed inlaying clamps matching films\u2019 thickness was manufactured as shown in Fig. 1 (a). All the sensors are supported vertically without being clamped the sensing area, and then electric wiring from the pins for signal transmission is done below the testbed. 2019 IEEE 8th Global Conference on Consumer Electronics (GCCE) 978-1-7281-3575-5/19/$31.00 \u00a92019 IEEE656 To test whether this system can observe the deformation and stress distribution of substrate, an experiments were conducted. At the same time, the effect of the distance of adjacent PVDF elements in horizontal direction was also evaluated" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0001870_icra40945.2020.9197245-Figure1-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0001870_icra40945.2020.9197245-Figure1-1.png", + "caption": "Fig. 1. Model of a variable stiffness spring. In this model, z denotes the length of the spring, Fz denotes the force by the spring, while k denotes the stiffness of the spring changed by the \u201cmotor\u201d m.", + "texts": [ + " Downloaded on September 21,2020 at 05:02:58 UTC from IEEE Xplore. Restrictions apply. II. VARIABLE STIFFNESS SPRINGS Variable stiffness actuators use two motors; one to control the deformation, and another to set the stiffness of the spring [1], [2], [17]. In this work, we consider variable stiffness spring actuators [18], sometimes referred to as stiffness modulators [7], [8], which only provide stiffness modulation without changing the equilibrium position of the actuator. The conceptual model of a variable stiffness spring actuator is shown in Fig. 1. A. Variable stiffness helical springs The potential energy function of a helical extension spring, with changeable active length and active length dependent stiffness, is given by: V (z,z0) = 1 2 k(z0)(z\u2212 z0) 2, (1) where z\u2212 z0 defines the deformation of the spring, while k(z0) is the stiffness of the spring. The latter relation may be justified by noting that the stiffness of a helical spring is the function of the number of active coils which, in turn, defines the active length z0 of the spring [12]", + " This assertion leads to the following condition: \u2202Vmax \u2202 z0 = \u2202 \u2202 z0 max z\u2208[zmin,zmax] V (z,z0)> 0. (7) Assuming that the maximum potential energy is attained at z = zV \u2208 {zmin,zmax}, the inequality can be reduced to the following set of basis functions: B3 = {(z\u2212 z0) mzn 0}, where (m,n) \u2208 {N\u00d7Z : (8) sgn[Cm,n]sgn[zV \u2212 z0] m\u22121 ( m+n\u2212 zV z0 n ) < 0}. Similar to the condition derived in the previous section, (8) is sufficient but not necessary to achieve (7). For variable stiffness springs that do not require energy to hold stiffness, the stiffness modulating actuator, e.g., the motor in Fig. 1, should experience zero force when the spring is not deformed, i.e., when it does not store energy [17]. This imposes the following condition: \u2200z0 : Fz0(z0,z0) =\u2212 \u2202V \u2202 z0 (z0,z0) = 0. (9) This equality relation implies the following set of basis functions: B4 = {(z\u2212 z0) mzn 0}, where (m,n) \u2208 {N\u00d7Z : m\u2265 2}\u222a (0,0). (10) According to (10), the exponent of the deformation should be greater than one or both exponents must be zero. Potential energy functions that describe the class of variable stiffness springs, ideal for energy storage applications, can be obtained by a linear combination of basis functions that satisfy the following conditions (3), (5), (7), and (9)" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0002850_978-94-007-6046-2_41-Figure1-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0002850_978-94-007-6046-2_41-Figure1-1.png", + "caption": "Fig. 1 A typical walking step of the simplest walking model. Just after footstrike, the swing leg (heavy line) swings forward past the stance leg (thin line) until the swing leg hits the ground and a new step begins. is the angle between the stance leg and the slope normal, is the angle between the two legs, l is the leg length, M is the hip mass, m is the foot mass, g is the gravitational acceleration, and is the slope angle (Courtesy of Martijn Wisse [1])", + "texts": [ + " Stability analysis of fixed points can be determined by analyzing a linearization of fcl around x , fcl x x A x x ; (7) where A is a constant matrix. The eigenvalues of A will provide the local stability about the fixed point x . If the real parts of the eigenvalues of A are all less than 0, x will be a stable fixed point as a small deviation from x will converge back to x , given sufficient time. This analysis can be useful for analyzing standing balance in a legged system. When a legged system, such as the one depicted in Fig. 1, is walking, the state does not remain at a fixed point. Instead, it makes a cyclic path through state space as shown in Fig. 2. This path in state space is called a limit cycle [2]. This limit cycle is considered stable when the system converges back to the nominal path after having deviated from it over a certain number of steps without necessarily requiring that the state is locally stable at every instant in time. Limit cycle walking is often used to refer to the latter [1]. Suppose the limit cycle repeats with a period of T seconds, x(t)D x(tC T)" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0001057_0021998320920920-Figure2-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0001057_0021998320920920-Figure2-1.png", + "caption": "Figure 2. Unit-cells within clipped region. (a) All four unit-cells in the clipped region. (b) Unit-cell selected for results in later section (unit-cell 2).", + "texts": [ + " Figure 1(b) shows the tow architecture more clearly by removing the matrix and highlights the variation of cross-section shape within the tows. Because VTMS relies on contact to compact the textile geometry, the outermost tows experience more deformation than the tows nearest the midplane, which is shown in Figure 1(b). The extreme deformation of the outermost tows is reduced by compacting one layer at a time, but the outermost tows will still experience more deformation than the interior tows. The tow VF is important to consider for a textile model. As shown in Figure 2(a), there are four unitcells within the clipped region for postprocessing, and each unit-cell is expected to have some variation due to how VTMS simulates the compaction and relaxation of the digital chains. Table 1 shows the VF of each type of tow in each unit-cell shown in Figure 2, with the difference in the tow VF between unit-cells illustrating how much the tow architecture varies throughout the model. The largest relative change of tow VF between unitcells occurred between unit-cells 2 and 3 where there was a 6% difference for the binders. Overall, the tow VFs remained quite similar across unit-cells. The total tow VF is lower than typical specimens found in the literature. The average tow VF across the unit-cells was 42.5%, which would translate to a fiber VF of about 30%, while the fiber VF in typical textile composites is between 40% and 55%", + " As mentioned before, the boundary conditions are applied to the full analysis region, and the analysis region is clipped for post-processing to reduce any artifacts near the boundaries due to non-periodic boundary conditions. Figure 4 shows the full analysis region and the subregion used for post-processing. The clipped region used for post-processing is only approximately a full RVE, since the textile geometry is non-periodic. Within the clipped region, there are approximately four unit-cells, as illustrated in Figure 2(a). It should be noted that the unit-cells shift along the x-direction by a quarter wavelength of the binder\u2019s path. Presumably, the different unit-cells should experience similar stress distributions, after accounting for the shift of binders between the unit-cells. Consequently, in later sections, stress concentrations are investigated for the unit-cell shown in Figure 2(b), unless otherwise stated. Quantitative metric for stress severity This paper does not aim to predict the progression of damage in a 3D textile, but a failure initiation criterion based on the stress state as predicted by a linear elastic analysis can offer insight into which geometrical features within an orthogonally woven textile cause severe stresses for a given set of boundary conditions. The textile community has yet to develop a failure criterion specialized for tows that is widely accepted", + " The failure indices are used for relative comparison, and if the expected point at which initial damage occurs is needed, the stresses can be linearly scaled with the applied load, since the elastic response is considered. The results herein describe the stresses as predicted by a linear elastic analysis with the aim to understand which features of the tow architecture have a strong influence on creating stress concentrations that can lead to damage. Figure 5 shows contours of the maximum index of Hashin\u2019s criteria, Hmax, for the binder in unit-cell 2 (refer to Figure 2 for the unit-cell enumeration). The maximum failure index provides a metric of severity of the stress state. The regions of severe stresses tended to form near two types of features of the tow architecture. First, the most severe stresses in the binder developed where the binder transitions between travelling along the x-axis and travelling through the thickness, such as label A in Figure 5. In these regions where the tow path of the binder transitions and the most severe stresses occur, transverse tension is the mode corresponding to the maximum failure index, as shown in Figure 6", + " However, since there is only one wavelength of a single binder within a unit-cell, Figure 10 shows contours of the maximum index of Hashin\u2019s criteria for the binder in unitcell 4, which can be compared to Figure 5 that shows the same results for unit-cell 2. The locations of high Hmax matched well between the two unit-cells. However, the highest concentrations were more severe within the volume of the binder in unit-cell 2 compared to unit-cell 4, and the cross-sectionally averaged Hmax in the regions of concentrations in the binders differed by up to 16% across all four unit-cells shown in Figure 2(a). Figure 11 shows contours of the maximum index of Hashin\u2019s criteria, Hmax, for the wefts in unit-cell 2. Within the wefts, local maxima of Hmax exclusively occurred in locations near a binder, with the largest regions with a high Hmax occurring in the wefts at the end of the z-crowns, such as at labels A in Figure 11. Unlike the concentrations of Hmax near the ends of the z-crowns, concentrations of Hmax that occurred in locations near a z-aligned binder remained highly localized, such as at labels B in Figure 11" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0002045_ecce44975.2020.9236116-Figure4-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0002045_ecce44975.2020.9236116-Figure4-1.png", + "caption": "Fig. 4. FEM analysis result of magnetic field in Litz wire.", + "texts": [ + " This paper analyzed the static magnetic field in the rectangular and circular solid wires with the uniform current distribution. Then, the square of the magnetic field was calculated and averaged over the cross-section. Finally, the ratio of the averaged square magnetic field of the rectangular wire to that of the circular wire was calculated, to determine M. The 2-dimensional FEM static magnetic field analysis was performed using JMAG-Designer 18.1 (JSOL Corp.) As an example, the simulation result of the static magnetic field for the rectangle with the aspect ratio of 3:1 is presented in Fig. 4. Consequently, the factor M was determined as a function of the aspect ratio as shown in Fig. 5. IV. EXPERIMENT An experiment was carried out to verify the proposed AC resistance analysis method. For this purpose, the AC resistance of the Litz wires with the circular and rectangular cross-section was experimentally measured and compared with the analytical prediction of the previous method and the proposed method, respectively. These two Litz wires have the same length, the same number of strands, the same strand diameter, the same cross-section area" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0000407_6.2020-1184-Figure5-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0000407_6.2020-1184-Figure5-1.png", + "caption": "Fig. 5 Deformed tape spring geometry in opposite sense bending.", + "texts": [ + " The natural fold radius for a partially flat anisotropic tape spring can be derived by constructing the flexural energy using superposition, in a similar manner to [17]. The strain energy due to bending, Ub , of a thin plate/shell element is given by [30]: Ub = 1 2 \u222b \u2206\u03baMdA = 1 2 \u222b \u2206\u03baD\u2206\u03badA (13) where D is the bending stiffness matrix for a laminate and A is the area. The central flat region of the tape spring undergoes a change in curvature, \u2206\u03ba f lat : \u2206\u03ba f lat = [ \u00b11/rx 0 ] (14) where rx is the longitudinal bend radius, described in Figure 5, for opposite and equal sense bending respectively. Substituting Equation 14 into 13 and integrating over the area of the fold, w f latrx\u03b2, gives the energy of the flat section, Ub, f lat : Ub, f lat = D11w f lat \u03b2 rx . (15) The arc regions of the tape spring undergo changes in curvature, \u2206\u03baarc , given by: \u2206\u03baarc = [ \u00b11/rx 1/Rarc ] (16) Substituting Equation 16 into 13 and integrating over the area of the fold, 2R\u03b1rx\u03b2, gives the energy of the two arc sections, Ub,arc , as: Ub,arc = 2\u03b2\u03b1 ( D11Rarc rx \u00b1 2D12 + D22rx Rarc ) (17) Adding Equation 15 and Equation 17 and minimising with respect to rx yields the natural fold radius, r\u2217x : r\u2217x = \u221a RarcD11w f lat + 2\u03b1R2 arcD11 2\u03b1D22 (18) for both opposite and equal sense bending" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0002119_icaccm50413.2020.9212874-Figure7-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0002119_icaccm50413.2020.9212874-Figure7-1.png", + "caption": "Fig. 7. Total deformation of GFRP propeller(case1).", + "texts": [], + "surrounding_texts": [ + "A Quadcopter\u2019s propeller is subjected to horizontal bending due to collision of the propeller with obstacles or a wall. To calculate the bending deformation under the applied rotational moment about the rotation axis, the propeller has been fixed by both the tip as shown in Figure 10.In addition, it is fixed at the point where it is connected to the motor shaft (Figure 11).Figure 12 shows the applied rotational velocity of 897 rad/s [8].Figure 13, 14 and 15, 16 show the total deformation and equivalent stress for CFRP and GFRP materials, respectively. The comparison of the obtained results is presented in Table III. 61 Authorized licensed use limited to: Carleton University. Downloaded on May 30,2021 at 15:29:13 UTC from IEEE Xplore. Restrictions apply. Fig. 14. Total deformation of GFRP propeller (case2). Fig. 15. Equivalent stress for CFRP propeller (case2). Table III. Results due to rotational velocity CFRP GFRP Max. Deformation 0.2012 mm 0.268 mm Max. stress 807.54 MPa 1127 MPa C. Vibration analysis The propeller has been tested for vibrational effect to find out the resonance frequencies for both the materials. The modal analysis has been performed in Ansys, Figure 17 and 18 show the frequency variation with first six mode shapes for CFRP and GFRP material, respectively. The color contour describes the range of deformation from minimum to maximum. Table 4 and figure 19 present the frequency comparison for CFRP and GFRP materials and figure 19 also present the comparison with previously publish work. Fig.17. CFRP based propeller\u2019s vibration frequencies and mode shapes (case3). 62 Authorized licensed use limited to: Carleton University. Downloaded on May 30,2021 at 15:29:13 UTC from IEEE Xplore. Restrictions apply. IV. CONCLUSION A Quadcopter\u2019s propeller has been analyzed for the structural and vibrational aspects. The propeller has been analyzed under the thrust (case1) and rotational loading (case2) conditions to find out the stress and deformation produced in two materials. It is found that for case1 both the materials show almost same results, while for case2 GFRP shows the high values of maximum deformation and stress compared to CFRP. Moreover, from the results of the modal analysis, it is found that CFRP is having high values of frequencies compared to GFRP. From these observations, it can be concluded that CFRP perform better than GFRP material under static and vibration loading condition. Furthermore the natural frequency of propeller is validate with publish literature [10] in figure 19. REFERENCES [1] Ahmad, F., Kumar, P., Bhandari, A., & Patil, P. P. (2020). Simulation of the Quadcopter Dynamics with LQR based Control. Materials Today: Proceedings, 24, 326-332. [2] Ahmad, F., Kumar, P., & Patil, P. P. (2018). Modeling and simulation of a Quadcopter with altitude and attitude control. Nonlinear Studies, 25(2). [3] Wei, P., Yang, Z. and Wang, Q. (2015). The Design of Quadcopter Frame Based On Finite Element Analysis,In 3rd International Conference on Mechatronics, Robotics and Automation. Atlantis Press. [4] Jaouad, H., Vikram, P., Balasubramanian, E., & Surendar, G. (2020). Computational Fluid Dynamic Analysis of 63 Authorized licensed use limited to: Carleton University. Downloaded on May 30,2021 at 15:29:13 UTC from IEEE Xplore. Restrictions apply. Amphibious Vehicle. In Advances in Engineering Design and Simulation (pp. 303-313). Springer, Singapore. [5] Kumar, S., & Mishra, P. C. (2016). Finite element modeling for structural strength of quadcoptor type multi modevehicle. Aerospace Science and Technology, 53, 252- 266. [6] Ahmed, M. F., Zafar, M. N., & Mohanta, J. C. (2020, February). Modeling and Analysis of Quadcopter F450 Frame. In 2020 International Conference on Contemporary Computing and Applications (IC3A) (pp. 196-201). IEEE. [7] Martinetti, A., Margaryan, M., & van Dongen, L. (2018). Simulating mechanical stress on a micro Unmanned Aerial Vehicle (UAV) body frame for selecting maintenance actions. Procedia manufacturing, 16, 61-66. [8] Singh, R., Kumar, R., Mishra, A., & Agarwal, A. (2020). Structural Analysis of Quadcopter Frame. Materials Today: Proceedings, 22, 3320-3329. [9] Xiu, H., Xu, T., Jones, A. H., Wei, G., & Ren, L. (2017, December). A reconfigurable quadcopter with foldable rotor arms and a deployable carrier. In 2017 IEEE ROBIO (pp. 1412-1417) [10] Ahmad, F., Bhandari, A., Kumar, P., & Patil, P. P. (2019, November). Modeling and Mechanical Vibration characteristics analysis of a Quadcopter Propeller using FEA. In IOP Conference Series: Materials Science and Engineering (Vol. 577, No. 1, p. 012022). IOP Publishing. 64 Authorized licensed use limited to: Carleton University. Downloaded on May 30,2021 at 15:29:13 UTC from IEEE Xplore. Restrictions apply." + ] + }, + { + "image_filename": "designv11_71_0003778_1.4030163-Figure10-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0003778_1.4030163-Figure10-1.png", + "caption": "Fig. 10 Conceptual illustration of wire-suspended platform system", + "texts": [ + " The author would like to thank Dr. Jean-Pierre Merlet for his valuable comment on this paper. We formulate the two basic equations in general form, that is, the geometrical relation (1) and the force and moment acting on the platform (2) due to the wire tension and the gravity. Their Jacobian matrices are also given. Symbols are listed in Table 4. A conceptual illustration expressing a fixed connection point, a platform connection point and the corresponding kinematical and geometrical wire lengths is shown in Fig. 10. A.1 Geometrical Relation A.1.1 Formulation of l\u00f0X\u00de. Equation (1) is formulated based on the following geometrical relation: l \u00bc l\u00f0x1;\u2026; xN\u00de \u00bc \u00bdl1\u00f0x1\u00de;\u2026; lN\u00f0xN\u00de T; li\u00f0xi\u00de \u00bc \u00f0xi xi\u00de T\u00f0xi xi\u00de h i\u00f01=2\u00de (A1) and the coordinate transformation of the position of wireconnection point on the platform xi \u00bc xi\u00f0X\u00de \u00bc R\u00f0h\u00dex\u0302i \u00fe x (A2) where R is an appropriate rotation matrix of h. A.1.2 Formulation of @l=@X. The Jacobian matrix of Eq. (1) is obtained based on the chain rule as Table 4 Symbols used in formulation of basic equations Symbol Comment N Number of wires q \u00bc \u00bdq1; ; qN T Wire lengths vector (kinematical lengths) l \u00bc \u00bdl1; ; lN T Distances between the wire connection points (geometrical lengths)bxi \u00bc \u00bdx\u0302i; y\u0302i; z\u0302i T Position of the connection point of ith wire at platform, referring to the platform-fixed coordinate (i \u00bc 1; ;N) xi \u00bc \u00bdxi; yi; zi T Position of the connection point of ith wire at platform, referring to the global coordinate (i \u00bc 1; ;N) xi \u00bc \u00bdxi; yi ; zi T Position of the fixed connection point of ith wire at the wall, at the ceiling or at the floor, referring to the global coordinate (i \u00bc 1; ;N)bc \u00bc \u00bdx\u0302c; y\u0302c; z\u0302c T Mass center position of the platform, referring to the platform-fixed coordinate c \u00bc \u00bdxc; yc; zc T Mass center position of the platform, referring to the global coordinate X \u00bc \u00bdxT ; hT T Platform position (x \u00bc \u00bdx; y; z T) and orientation (h \u00bc \u00bdw; h;/ T ) m Platform mass s \u00bc \u00bds1; ; sN T Wire tension vector F \u00bc \u00bdf T ; gT T Force f \u00bc \u00bdfx; fy; fz T and moment g \u00bc \u00bdgx; gy; gz T acting on the platform, due to the wire tension and the gravity g \u00bc \u00bdgx; gy; gz T Gravitational acceleration vector 041021-12 / Vol" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0002276_s12541-020-00431-8-Figure5-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0002276_s12541-020-00431-8-Figure5-1.png", + "caption": "Fig. 5 Static test for torsion beam model (a bending test, b torsion test)", + "texts": [ + " Figure\u00a04 shows the parameters of length properties for torsion beam modeling. (2)ASY = A I2 yy \u222b A ( Qy lz )2 dA ASZ = A I2 zz \u222b A ( Qz ly )2 dA (3)Py = 12EIzzASY GAL2 Pz = 12EIyyASZ GAL2 1 3 The developed torsion beam model is compared with the finite element (FE) torsion beam model through static bending tests and static torsion tests. The beam parameter values used in the torsion beam modeling are shown in Table\u00a02. In the bending and torsion tests, one side of the torsion beam is fixed, as shown in Fig.\u00a05 On the other side, 200\u00a0N of force is applied in a downward direction while 200\u00a0Nm of torque is applied in an axial direction. Table\u00a03 shows that the test results are accurate with an overall error within 5%. These errors are considered to occur because the Timoshenko beam elements cannot express the warping effect. 1 3 The torsion beam model presented in Sect.\u00a02.1 is combined with the remaining parts to create the CTBA model. First, parts such as the trailing arm, knuckle, and spring lower mount are configured as a rigid body model, such that the CTBA\u2019s hardpoints can be directly changed" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0002586_edpc51184.2020.9388196-Figure9-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0002586_edpc51184.2020.9388196-Figure9-1.png", + "caption": "Fig. 9. Design of contour cooling.", + "texts": [ + " The function of friction reduction is provided by the tooth flank coating, while the heat is removed by the contour cooling channels inside the LBM-printed structure. A glycol-water mixture will be used as a coolant. Although the planned demonstrator will be installed and tested in a test bench setup, in perspective the entire system is intended to drive an EV. Therefore the circuit could be integrated into the cooling system of the electric machine. The final internal structure of the pinion shaft is shown in Fig. 9. It is recognizable that the inlet (blue) and outlet (green) of the cooling medium takes place on one side of the shaft. The Reason for this is that the opposite shaft side is used for torque input in the test bench of the demonstrator as well as in the perspective e-gearbox integration. Finally, a rotary union allows the coolant to be transported between pinion shaft and cooling system. This can be seen at Fig. 1 on the left. Via the inflow line (blue) the coolant reaches the distribution channel (turquoise) via four cross-like arranged branches" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0001406_s11431-019-1525-0-Figure3-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0001406_s11431-019-1525-0-Figure3-1.png", + "caption": "Figure 3 (Color online) Capability zonotope at the given pose. (a) 3D view; (b) XY view; (b) ZX view; (d) YZ view.", + "texts": [ + " In the simulations, the snake robot keeps tracking of a square trajectory, and the operational force capability is evaluated by the developed method at each pose, which is visualized in MATLAB as shown in Figure 2(b). All the parameters of the dynamics model are calculated and updated with the CASPR [10]. At a given pose (q=[0.1, 0.1, 0.1, \u2026, 0.1, 0.1]T), the simulated snake arm and its end effector force zonotope are visualized in MATLAB. Detailed views such as 3D, XY, ZX, YZ views of the zonotope are shown in Figure 3. From the figures we can see there are many redundant vertices and faces which contribute little to the volume of capability zonotope. The number of cables added in each step (Sc) of the iterative is an important parameter, which effects the computation efficiency of the method. To study the time efficiency related to Sc, we generate the end effector force zonotope when the snake robot keeps tracking of a square with different Sc and record the time consuming at each pose. The initial cable number is 6, and can be chosen randomly or not randomly", + " If the cables are chosen randomly, if rand is true, otherwise is false. The step number of cables are 1, 2, 3, 5, 6, and 10. The time consuming of each pose of the robot is recorded and shown in Figure 4. Detailed average time consuming of zonotope generation with different step cable numbers Sc are listed in Table 1. From Figure 4 and Table 1, we can see that the optimal step number of cables is 5. Another problem for the zonotope representation is there are too many redundant vertices and faces as shown in Figure 3, which has little contribution to the volume of the zonotope. We may mention that if a vertex has little contribution to the volume of the zonotope, it can be regarded as a redundant one. Thus, we can omit this vertex in the calculation for simplicity, which will be verified in the following simulation results. However, such redundant faces corresponding to the constraints of the robot make the post-process like manipulability optimization and motion planning difficult and inefficiency. Thus, we put forward a threshold to diminish the number of vertices faces" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0001320_j.mechmachtheory.2020.103953-Figure1-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0001320_j.mechmachtheory.2020.103953-Figure1-1.png", + "caption": "Fig. 1. A 2-DOF mechanism with mobile frame: scheme and notations.", + "texts": [ + " 4 As consequence, it can be concluded that, in general, the IMT is \u201cefficient enough\u201d to be implemented in a general multibody-dynamics software and that its real efficiency can be greatly improved by combining it with suitable notations that depend on the set of mechanism types to analyze. The identification of these notations is out of the scope of this paper. 4. Case studies In this section, the IMT is applied to three case studies to better illustrate its effectiveness. Even though the technique is general, the three cases have been intentionally chosen simple enough for the reader to be easily able to follow and check all the steps. The first case ( Fig. 1 ) refers to a mechanism with mobile frame (i.e., with time-dependent constraints). The second case ( Fig. 6 ) refers to a mechanism with non-holonomic constraints. The third case ( Fig. 11 ) refers to a spatial mechanism usable as a pointing system, for instance, to orientate antennas [ 35 , 36 ]. 4.1. A 2-DOF mechanism with mobile frame Fig. 1 shows a two-DOF planar mechanism whose frame (link 1) is mobile with respect to an observer fixed to the Cartesian reference Ox 0 y 0 . This situation occurs when a mechanism is mounted on a ground or air vehicle. The mechanism of Fig. 1 controls the position of point P, which is fixed to link 5, with respect to link 1. With reference to Fig. 1 , the angles \u03b82 , \u03b83 , and \u03b85 define the orientations of links 2, 3, and 5, respectively, with respect to link 1; whereas, the angle \u03b81 and the coordinates (x A , y A ) of point A in Ox 0 y 0 define the pose of link 1 in Ox 0 y 0 , and depend on time according to a known motion law. The lengths of the segments AD, AB, BC, DC, EP, and BE are denoted a 1 , a 2 , a 3 , a 4 , a 5 , and a 6 , respectively. The angles \u03b82 and \u03b85 are chosen as generalized coordinates; whereas, the angle \u03b83 and the coordinates (x P , y P ) of point P in Ox 0 y 0 are the secondary variables. The geometric relationships (see Fig. 1 ) 5 ( P \u2212 O ) = ( P \u2212 E ) + ( E \u2212 B ) + ( B \u2212 A ) + ( A \u2212 O ) (21a) 4 In this comparison taking into account the jerk analysis contribution somehow roughly compensates the fact that, in [27] , the torques\u2019 computation is included. 5 Here, the square of a vector denotes the dot product of the vector by itself. ( C \u2212 D ) 2 = [ ( C \u2212 B ) + ( B \u2212 A ) \u2212 ( D \u2212 A ) ] 2 , (21b) after the introduction of the above-defined notations and some rearrangements, become the following constraint equations x P = x A + [ a 2 c ( \u03b82 + \u03b81 ) + a 5 c ( \u03b85 + \u03b81 ) + a 6 c (\u03b3 + \u03b83 + \u03b81 ) ] (22a) y P = y A + [ a 2 s ( \u03b82 + \u03b81 ) + a 5 s ( \u03b85 + \u03b81 ) + a 6 s (\u03b3 + \u03b83 + \u03b81 ) ] (22b) a 2 4 = a 2 1 +a 2 2 +a 2 3 + 2 [ a 2 a 3 c ( \u03b82 \u2212\u03b83 ) \u2212 a 1 ( a 2 c \u03b82 + a 3 c \u03b83 ) ] (22c) where c x and s x stand for cosx and sinx", + " Jerk analysis and higher-order analyses After the velocity and acceleration analyses, the jerk analysis follows the same scheme as the acceleration analysis, and it requires only the preliminary computation of the drift terms e 1 and e 2 , but the matrices appearing in expression ( 18a ) of e 1 have been already computed and only the ( m + r )-tuples H x ( d ) appearing in expression ( 18b ) of e 2 must be computed. These computations are omitted since they are similar to the previous ones and yield big analytic expressions. The same holds for higher-order analyses with order greater than 3. 4.1.3. Numerical example The above-deduced formulas have been applied to a mechanism geometry defined as follows (see Fig. 1 ; the linear sizes are measured in a generic length unit (l.u.)): a 1 = 1 l.u., a 2 = 0.5 l.u., a 3 = 0.9 l.u., a 4 = 0.8 l.u., a 5 = 0.5 l.u., a 6 = 0.5 l.u., \u03b3 = ( \u03c0 /3) rad. The motion of the frame has been assigned as follows: x A = 0 l.u., y A = 0.05sin( \u03c9t) l.u., \u03b81 = ( \u03c0 /18)sin( \u03c9t) rad where t is the time, and \u03c9= (2 \u03c0 /T) rad/s with T = 10s; whereas, the motion imposed by the input variables has been assigned as follows: \u03b82 = \u03c9t rad, \u03b85 = ( \u03c0 /3)[1 + 2sin(2 \u03c9t)] rad. The motion simulation has been implemented in Matlab R2019b and the results are reported in the diagrams of Figs" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0002902_978-94-007-6046-2_43-Figure4-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0002902_978-94-007-6046-2_43-Figure4-1.png", + "caption": "Fig. 4 Early platforms based on the SLIP model. (a) Raibert\u2019s monopod (Photo courtesy of B. Brown) (b) The ARL monopod (Photo from [1]). (c) The Bow-Leg monopod (Photo courtesy of G. Zeglin)", + "texts": [ + " In either case, the agility, stability, and robustness of these machines are tied to the agility, stability, and robustness of the underlying model, propelling active research in dynamics and control of legged locomotion based on this template. Legged robot platforms inspired by the SLIP model range from mechanical designs that directly embody the model to higher degree of freedom platforms that seek to embed SLIP behavior by active control. The Raibert hopper was the first legged robot that directly embodied the SLIP [45, Fig. 4a]. The platform used a prismatic piston leg with an air chamber providing elastic rebound. This pneumatic actuation was combined with hydraulic actuation to supply energy in stance and to place the leg in flight. The hopping robot was the first machine to demonstrate running behaviors in the sagittal plane. Subsequent generalization of this monopedal design led to bipedal and quadrupedal robots running in three dimensions [45]. The gait stability and robustness of these running machines were not well understood, triggering a chain of theoretical research on legged dynamics and control that used and further shaped the SLIP model. Following the success of Raibert\u2019s running robots, the ARL family of monopod robots investigated more effective mechanisms for prismatic leg actuation to regulate energy and for angular control of the leg during both flight and stance [20, Fig. 4b]. These platforms were capable of faster and more efficient running with their passive dynamic mechanisms incorporating series elastic actuation both in the radial and angular degrees of freedom [2]. The Bow-Leg platform (Fig. 4c) is an example of a Raibert-style hopper that more directly approximated the energy efficiency of the SLIP model, which does neither inject nor dissipate energy. Even though such an energetically conservative behavior is impossible to reproduce on a physical platform, the Bow-Leg platform achieved high energy efficiency by using a very lightweight leg design to minimize collision losses and a novel radial actuation mechanism that pre-compressed the leg spring during flight for discrete energy injection [69]", + " A third approach was introduced by [43], who assume an ideal actuator in series with the leg spring and implement partial feedback linearization to cancel the nonlinearities in (2). The resulting dynamics lead to approximate solutions that provide a basis for gait controllers with an actuated spring leg. Control strategies of the planar SLIP model have evolved over time from intuitive to optimal control methods. The early control strategies go back to the 1980s. They arose from the work on legged machines that roboticists were interested in. Raibert and colleagues [45] devised an intuitive controller for their planar hopping machine (Fig. 4a), concentrating on the attitude of the center body, hopping height, and forward speed (Fig. 8). A proportional-derivative feedback on the global body angle regulated the attitude of the center body during stance. The control of the hopping height relied on thrusting the hydraulic piston of the hopper\u2019s leg, which injected energy at maximum leg compression by shifting the rest length of the air spring that was connected in series with the piston. In effect, the duration of the thrust regulated hopping height" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0003808_s00542-015-2524-5-Figure3-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0003808_s00542-015-2524-5-Figure3-1.png", + "caption": "Fig. 3 a Finite element model and b pressure distribution of FDBs with stationary herringbone grooves", + "texts": [ + "5\u2033 hard disk drive design depicted in Fig. 1. The mass of the rotor, including a 2.5\u2033 disk, is 20.988 g. FDBs consist of coupled journal and thrust bearings. They only allow an axial rotational motion by constraining remaining five degrees of freedom, so the dynamic characteristics of coupled journal and thrust bearings have to be analyzed in five degrees of freedom. In this chapter, the dynamic characteristics and the stability are investigated with the consideration of five degrees of freedom. Figure 3 shows a finite element model of the stationary grooved FDBs, and the pressure distribution of the (46)\u03beij = Vij/m11\u03c7 2 (47)\u03c1ij = uij/m11\u03c7 (48)\u03c4ij = Uij/m11\u03c7 (49)|[P =0] + [I]| = 0 (50)det[P] < 0. 1 3 stationary grooved FDBs with the whirl radius of 0.185 \u00b5m at the rotational speed of 7200 rpm. This finite element model consists of two herringbone-grooved journal bearings, four plain journal bearings, two spiral-grooved thrust bearings and one plain thrust bearing. The fluid film was discretized into 6120 isoparametric bilinear elements with four nodes" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0003634_icems.2014.7013745-Figure3-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0003634_icems.2014.7013745-Figure3-1.png", + "caption": "Fig. 3. No-load magnetic field distribution (one position)", + "texts": [ + " Therefore, it is necessary to analyze by FEM in order to figure out detailed electromagnetic coupling of PMRM. III. ANALYSIS BY FINITE ELEMENT METHOD In order to analyze electromagnetic coupling of PMRM quantitatively, finite element method (FEM) is used. In the PMRM, because of the doubly salient structure of outer motor, magnetic circuit of outer motor will be different when outer rotor rotates to a different position. Magnetic field distribution of PMRM when permanent magnets act alone is shown in Fig. 3 and Fig. 4. As Fig. 3 shows, when outer rotor rotates to this position, magnetic flux will go through two aligned teethes of stator and outer rotor. When there are no aligned teethes, magnetic flux will go through adjacent teethes as Fig. 4 shows. Magnetic field distributions of PMRM when inner rotor or stator armature windings energized alone are shown in Fig. 5 and Fig. 6. motor will increase or decrease when +D axis or \u2013D axis currents of inner motor act alone. By comparing Fig. 5(c) with Fig. 5(d), it can be seen that when +Q axis currents act alone, magnetic fluxes of outer motor little change" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0001782_s40430-020-02586-x-Figure2-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0001782_s40430-020-02586-x-Figure2-1.png", + "caption": "Fig. 2 Mechanical model of the system. The manipulator will be have a dynamical behavior that coincides with the model of the human arm", + "texts": [ + " The squared joint torque is also an interesting choice because it is a measure of the energy consumption, whose minimization is a principle found in biological motion, and it is also associated with the size, power and cost of actuators required to control the system. Using functions 16 to 21, we can achieve pointwise optimality to the problem of minimizing the objective function. This is the quadratic-programming formulation, and it is presented in the simulations. To compare the performance of both approaches with a practical application, a redundant manipulator behaving as a simplified human arm model [7] was simulated under different conditions. The robot consisted of a planar four-DOF serial link manipulator (see Fig.\u00a02). The desired behavior of the robot end-effector is inspired on human arm motion. It consists of a virtual two-DOF serial manipulator representing the shoulder and elbow joints (see Fig.\u00a02c), with asymmetrical joint stiffness and damping, that varies according to external disturbances applied on the limbs. The wrist and hand are rigidly attached to the forearm, and they are considered the virtual end-effector. The virtual elbow et and shoulder st equilibrium positions move cyclically. The positions of each virtual joint equilibrium position are shown in Fig.\u00a03a. A cyclic force is applied on the robot end-effector, with a profile according 1 3 to Fig.\u00a03b. The robot end-effector translates the forces to the virtual end-effector, interfering with the arm motion. The resulting nonlinear impedance of the virtual human arm resists the interference. The robot end-effector task is to behave exactly like the virtual human arm, as it moves under the external forces (Fig.\u00a02c). The optimization criterion is to minimize the sum of squared joint torques and the sum of the squared error between the desired and the actual behavior. Additionally, an inertia-weighted torque optimization is also considered for the QP approach, as it is known to yield better results than pure torque optimization. The task space for both the robot end-effector and virtual end-effector is an horizontal plane. The task coordinates are the horizontal position and end-effector orientation, defined by [ xq yq q ]T for the robot end-effector and [ x y ]T for the virtual end-effector. The mapping p(q) \u2236 q \u21a6 [ xq yq q ]T between joint- and task-space coordinates of the robot end-effector is where Lk and qk denote, respectively, link lengths and joint coordinates for k = 1, 2, 3, 4 (Fig.\u00a0 2b). The mapping r( ) \u2236 \u21a6 [ x y ]T , between the virtual joint (92) p(q) = \u23a1 \u23a2\u23a2\u23a3 L1 cos(q1) + L2 cos(q2) + L3 cos(q3) + L4 cos(q4) L1 sin(q1) + L2 sin(q2) + L3 sin(q3) + L4 sin(q4) q4 \u23a4 \u23a5\u23a5\u23a6 , coordinates and virtual task-space coordinates of the virtual end-effector, is expressed as where Ls and Le represent, respectively, upper and lower arm lengths, with s and e denoting, respectively, shoulder and elbow angles (Fig.\u00a02c). Note that in the ideal case, both (92) and (93) map their arguments to the same point. In state-space representation, the virtual joint angles and velocities are denoted as The dynamical equation of the virtual system is (93)r( ) = \u23a1\u23a2\u23a2\u23a3 Ls cos( s) + Le cos( e) Ls sin( s) + Le sin( e) e \u23a4\u23a5\u23a5\u23a6 , (94) [ \ud835\udf03s \ud835\udf03e ?\u0307?s ?\u0307?e ]T = [ z1 z2 z3 z4 ]T . (95) z\u0307 = \u23a1 \u23a2\u23a2\u23a2\u23a3 z3 z4 M\u22121 v \ufffd \u2212Kv \ufffd z1 \u2212 \ud835\udf03st z2 \u2212 \ud835\udf03et \ufffd \u2212 Cv \ufffd z3 \u2212 ?\u0307?st z4 \u2212 ?\u0307?et \ufffd\ufffd \u23a4\u23a5\u23a5\u23a5\u23a6 + \u23a1\u23a2\u23a2\u23a2\u23a3 0 0 +Gv + JT r + \ufffd fm?\u0308?st ?\u0308?et \ufffd \u23a4\u23a5\u23a5\u23a5\u23a6 , Journal of the Brazilian Society of Mechanical Sciences and Engineering (2020) 42:513 1 3 513 Page 12 of 20 The physical system is a four-DOF ideal serial manipulator with diagonal moment of inertia tensor and negligible joint stiffness" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0003008_978-3-319-76138-1_1-Figure1.1-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0003008_978-3-319-76138-1_1-Figure1.1-1.png", + "caption": "Fig. 1.1 Conventional robot architectures: Industrial robots, Stewart\u2013Gough platform, and delta robot", + "texts": [ + " Serial robots are inspired by the structure and function of the human arm which consists of a series of joints and bones connecting the torso with the hand. Using the muscles as actuators, the hand can be freely moved in space to manipulate objects in the environment. For a robot as technical system, muscles and bones are replaced by \u00a9 Springer International Publishing AG, part of Springer Nature 2018 A. Pott, Cable-Driven Parallel Robots, Springer Tracts in Advanced Robotics 120, https://doi.org/10.1007/978-3-319-76138-1_1 1 links and actuated joints in order to mimic the human\u2019s motion capacities (Fig. 1.1a). Theoverall kinematic chain is called serial,meaning that there is oneunique sequence of links and actuated joints that create the desired motion. Serial robots are often referred to as robotic arms or articulated robots due to their kinematic similarity with the human arm. The difference of a parallel robot with respect to a serial robot is to add more kinematic chains to connect the end-effector with the base (see Fig. 1.1b, c). In order to allow for independentmobility in such a structure,most of the joints are left passive without motors. Although there are many possible choices to distribute the motors on the parallel robot, a beneficial solution is to put one motor on each kinematic chain to distribute the load amongst all kinematic chains. Again, this concept of parallel actuation can be found in the human body. Although the overall kinematic structure of human arms and legs is serial, the actuation of the joints has a parallel topology" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0003270_icpe.2015.7167862-Figure1-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0003270_icpe.2015.7167862-Figure1-1.png", + "caption": "Fig. 1. Concept diagram of equality demagnetization", + "texts": [ + " In addition, when high starting torque is demanded, armature reaction is main cause of demagnetization of PM [7]. Commonly, the irreversible demagnetization of PM is occurred to unspecific magnetization shape. Therefore, we classified irreversible demagnetization of PM into equality, inequality, and weighted demagnetization patterns. In addition, we performed characteristic analysis according to various irreversible demagnetization of PM. Finally, we propose irreversible demagnetization detection technique. Fig. 1 shows equality demagnetization pattern, which means that demagnetization ratio of all PMs, is equal. Demagnetization ratio is composed of 50% and 70%, and magnetization shape is appeared at lopsided step shape. In general, a demagnetization of an IPM does not occur evenly at whole face of PM. To be more specific, the portion of demagnetization is determined according to direction of rotate. This is because, when a motor is operating, the magneto-motive force from the supplied power affects the q-axis, not the d-axis [6]" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0002816_thc-2010-0566-Figure19-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0002816_thc-2010-0566-Figure19-1.png", + "caption": "Fig. 19. Mohr\u2019s circle for uniaxial tension.", + "texts": [ + " 17, the position of the center of the circle c is the average of the two normal stresses (\u03c3x and \u03c3y and reference point a can be plotted with \u03c4 xy on the y-axis and \u03c3x on the x-axis. Now we can join points c and a to get the radius and then the circle can be drawn with a radius given by Eq. (12). From this Mohr\u2019s circle for a given situation we can find out the maximum normal stress, point b, and shear stress, point e, active in the material. R = \u221a( \u03c3x \u2212 \u03c3y 2 )2 + \u03c42 xy (13) If we take a simple situation where a tensile stress of 100 Pa is active on a square cross section beam then the state of stress at a very small volume element will be as shown in Fig. 19A. Surprisingly, the Mohr\u2019s circle Fig. 19B) of this volume element shows that there exists an orientation (45\u25e6 clockwise, Fig. 19C) where maximum shear stress of 50 Pa is active. Therefore for supporting 100 Pa of stress we have to choose a material for which the \u03c3allowed is at least 100 Pa and \u03c4allowed is at least 50 Pa, all the material with \u03c4allowed < 50 Pa cannot be used even when they can support the normal stress of 100 Pa and will deform (Fig. 19D). Looking at a more complex combined loading of compressive and torsional nature, Fig. 20 shows a cylindrical beam loaded with 100 Pa compressive stress and a torque giving rise to a shear stress of \u221250 Pa on its surface. From point d in Mohr\u2019s circle we can see that the maximum normal stress induced is compressive and 120 Pa and point e tells us that the maximum shear stress is about 71 Pa. Therefore a material for this application should be capable of bearing a compressive stress> 120 Pa and a shear stress > 71 Pa" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0003439_transfun.e98.a.1973-Figure1-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0003439_transfun.e98.a.1973-Figure1-1.png", + "caption": "Fig. 1 DC motor system for experiments.", + "texts": [ + " Let \u03b1(F\u0304) and \u03b1(F\u0304) be the maximum and the minimum real parts of all eigenvalues of F\u0304, respectively. Under Assumption 1 there exist L\u03041 and L\u03042 such that \u03b1(F\u03042) < \u03b1(F\u03041) and \u03b1(F\u03041) < 0 since the pair (C\u0304, A\u0304) is observable. Then, there exists \u03b4 > 0 such that K\u03041 and K\u03042 are well defined and det(In\u0304 \u2212 eF\u03042\u03b4e\u2212F\u03041\u03b4) 0, \u2200t \u2208 (0, \u03b4] [7], [8]. Performance of the proposed observer is tested through computer simulations and laboratory experiments on velocity estimation of the one-link manipulator shown in Fig. 1. The system is represented by x\u03071 = x2, x\u03072 = \u2212a1x2 + a2u \u2212 a3\u03c4l (14) where x = [\u03b8 \u03c9]T represents the angular position and the velocity, respectively. The corrupted measurement output y = x1+w. In the experiments a sinusoidal signal w has been added to the measured output for testing the robustness. The input u is the armature current and \u03c4l is the load torque disturbance. The parameters a1 = Bm/Jm, a2 = Kt/Jm, a3 = 1/Jm; Kt is the torque constant; Jm and Bm denote the rotor inertia and the friction coefficient, respectively" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0000315_icems.2019.8921590-Figure7-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0000315_icems.2019.8921590-Figure7-1.png", + "caption": "FIGURE 7. (a) 3-step E-plane bend when the milling is done parallel to the narrow walls of the waveguide. (b) 3-miter E-plane bend when the milling is done perpendicular to the narrow walls of the waveguide.", + "texts": [ + "4 mm, LH2 = 5.37 mm and LH3 = 2.43 mm. For the second bend (with perpendicularmilling), two different standard diameters of endmills were used because larger endmills can cut deeper (see Fig. 6(b)). For shallow cuts a 1 mm diameter end mill was used. For deeper cuts, corresponding to the total width of the broad walls of the waveguide, a 1.5 mm end mill was chosen. Physical dimensions for this bend are the following; H1 = 3.98 mm, H2 = 1.64 mm, H3 = 5.52 mm, H4 = 3.96 \u00b5m, H5 = 1.6 mm and H6 = 6.33 mm. Fig. 7(a) shows a 3-step E-plane bend when the milling is performed parallel to the narrow walls of the waveguide. Physical dimensions for this bend (using a standard 1 mm diameter end mill) are the following; E1 = 2.5 mm, E2 = 2 mm, E3 = 0.67 mm, E4 = 2.01 mm, E5 = 0.66 mm, E6 = 2.5 mm and E7 = 3.25 mm. Fig. 7(b) presents a 3-miter E-plane bend. In this last case the milling is done perpendicular to the narrow wall of the waveguide. Physical dimensions are the following (using a standard 2 mm diameter end mill); P1 = 1.88 mm, P2 = 2.26 mm, P3 = 1.6 mm, P4 = 2.87 mm and21 = 15.1\u25e6. Simulated return losses as a function of frequency for the four bends are shown in Fig. 8. In all cases, the reflection coefficient is maintained below \u221244 dB over the 31\u201345 GHz band. These excellent results of compact devices can all be reached using standard end-mill diameters" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0002119_icaccm50413.2020.9212874-Figure2-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0002119_icaccm50413.2020.9212874-Figure2-1.png", + "caption": "Fig. 2. CAD model of the propeller.", + "texts": [ + " The remaining paper is arranged in the following sections: Section II presents the design of propeller and the materials used, Section III presents the static and vibration analyses of the propeller model based on the FEA, finally the paper is concluded in Section IV. 978-1-7281-9785-2/20/$31.00 \u00a92020 IEEE 59 Authorized licensed use limited to: Carleton University. Downloaded on May 30,2021 at 15:29:13 UTC from IEEE Xplore. Restrictions apply. II. PROPELLER DESIGN AND MATERIAL PROPERTIES The propeller has been designed in Creo 2.0, which is a CAD software for 3D modeling. The developed CAD model has been imported in Ansys for further structural analysis. Figure 2 and 3 show the solid CAD model and mesh model of the Quadcopter\u2019s propeller, respectively. The meshing has been done in Ansys by applying the body sizing mesh tool in which default mesh size is 2 mm. Table I shows the material properties of carbon fiber reinforced polymer(CFRP) and glass fiber reinforced polymer (GFRP). FEA is the most wildly used approach for various applications in various domains such as thermal, fluid, structural, harmonic, vibration, and medical. Ansys is based on the theory of FEA, which provides the user friendly interface to check the product performance under various boundary conditions" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0003364_ceit.2015.7233057-Figure1-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0003364_ceit.2015.7233057-Figure1-1.png", + "caption": "Fig. 1. Schematic Diagram of the WECS", + "texts": [ + "eywords\u2014Self excited induction generator (SEIG), excitation capacitance, wind turbine, renewable energy, wind energy conversion system (WECS) I. INTRODUCTION Wind energy is, for many reasons, one of the most promising renewable energy sources. The wind energy can be harnessed using a Wind Energy Conversion System (WECS), as shown in Fig. 1, comprising a wind turbine, an electric generator, a power electronic converter and the corresponding control system [1]. Self-excited alternating-current generators have obtained increased attention in recent years. This could be due to the suitability of these generators for various applications such as wind [2]. The suitability of the self-excited induction generator has made it widely used for different applications such as wind, tidal, and small hydroelectric renewable energy conversion. Due to isolated and decentralized nature of such resources, it is possible to develop the unit\u2019s off-grid and decentralized to feed local needs", + " Permanent magnet generators can also be used for wind energy applications but they suffer from uncontrollable magnetic field, which decays over a period due to weakening of the magnets, and the generated voltage tends to fall steeply with load. [3] - [4]. The SEIG has a self-protection mechanism because the voltage collapses when there is a short circuit at its terminals. Further, it has more advantages such as cost, reduced maintenance, rugged with simple construction, have a brushless rotor (squirrel cage), etc. Wind is a result of the movement of atmospheric air. The prime mover for the induction generator is the wind turbine connected via a gear box as shown in Fig.1. II. MATHEMATICAL MODEL OF A WECS In this section, the mathematical model of a wind energy conversion system comprising wind turbine, multiplier of speed, and an induction generator is developed. The aerodynamic model must produce an aerodynamic torque starting from the speed of the wind and the number of revolutions of the turbine. This speed corresponds to the speed of rotation three low-speed ( ) [3]. To obtain the aerodynamics model, we can use, initially, the expression of the mechanical power produced by a wind mill", + " It is given by the following relation: (1) The expression of the mechanical power can be modified to represent the torque extracted from the power contained in the wind: (2) Tip-speed ratio ( ) is a function of linear velocity of tip of the blade and the wind velocity and is given by the following equation: (3) The coefficient of power is generally modeled by the following analytical expression: (4) (5) With The multiplier adjusts the speed (slow) of the turbine at the speed of the generator as in Fig. 3. This multiplier is modeled mathematically by the following equations: Fig. 3 shows the axes equivalent circuit of a SelfExcited Induction Generator supplying an inductive load. Equation (10) and Figure 1 is a classical matrix formulation using D-Q axes modeling. It is used to represent the dynamics of conventional induction machine operating as a generator. The representation includes the self and mutual inductances as coefficients widely used in machine theory. Using such a matrix representation, one can obtain the instantaneous voltages and currents during the self-excitation process, as well as during load variation. The dynamic model of the three-phase squirrel cage induction generator is developed by using stationary D-Q axes references frame and the relevant volt-ampere equations are as [6]: (8) From which, the current derivative can be expressed as: (9) (10) Where , , , and are defined below: ; , And The SEIG operates in the saturation region and its magnetizing characteristics are non-linear in nature", + " Excitation with and without saturation When SEIG is excited with capacitance value C=270 F and value of rotor speed =1500 rpm the generated voltage and current attain their steady state values of 380 Volts and 19 amp in 0.8 sec, respectively as shown in Fig. 6 C. Variation of speed 0.2 0.3 0.4 0.5 0.6 0.7 0.8 0 200 400 Time (s) S ta to r V ol ta ge V s (V ) Vs w ith saturation Vs w ithout saturation D. Variations of speed after Full-Excitation E. Variation of excitation capacitance F. Variations of load and excitation after Full-Excitation 1) RLC balanced load and balanced excitation: In this case, the rotor of the induction machine is driven at 1500 r/min while , , and for each phase (see Fig. 1). Resolving (10), (11), (15), (17) and (19), the behavior of the SEIG was studied when it is connected to a balanced capacitive bank and a balanced RLC load. Different cases were investigated. Fig. 10 gives a set of results at steady-state. 2) RLC Unbalanced load and balanced excitation: In the second case with balanced excitation and unbalanced load, the SEIG is loaded with the following component values: , , , , , and , , , 0 0.01 0.02 0.03 0.04 -400 -200 0 200 Time (s) V ol ta ge V ab c (V ) Phase A Phase B Phase C 0 0", + "04 -20 0 20 Time (s) C ur re nt I ab c (A ) Phase a Phase b Phase c Therefore, using equations (10), (11), (15) and (18), the operation of the SEIG is simulated and the waveforms of the electrical variables are drawn in Fig. 11. However, within this range, it is possible to use the SEIG to feed an isolated unbalanced load without the use of static converters and complex controls. 3) RLC balanced load and unbalanced excitation: In the case of balanced RLC load and unbalanced excitation, all currents, stator voltage, load currents are balanced but stator currents and capacitor currents are unbalanced as shown in Fig. 12. The rotor of the induction machine is driven at 1500 r/min with , , , and for each phase (see Fig. 1). IV. CONCLUSION This paper presents some encouraging results that have been determined for the SEIG under balanced and unbalanced load conditions using balanced and unbalanced excitation conditions. The simulations have been obtained running MATLAB\u00ae/SIMULINK models developed in the previous chapters. Based on the results obtained for all the cases, a diphase analytical model of the induction generator, taking the saturation effects into account by means of a variable magnetizing inductance is obtained and presented in this paper" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0000245_ecce.2019.8912784-Figure3-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0000245_ecce.2019.8912784-Figure3-1.png", + "caption": "Fig. 3: Eddy current 3D distribution in the conductors (top) and a typical eddy current path considered in a 2D analysis, having a go and return path along the conductor Leff and not including an end coil section.", + "texts": [ + " For such methods, a combination of analytical equations and FEA is employed. Typically, flux density values are sampled from the simplified coil cross section representation of a 2D FE model, e.g. shown in Fig. 1c. The flux density is then used with analytical equations in order to calculate the eddy current losses. Although hybrid methods may lead to more accurate estimations than pure analytical methods, they have the risk of inaccuracy due to the fact that they disregard 3D effects. Neglecting end paths, as shown in Fig. 3 may result into un- derestimation of resistance and hence overestimation of eddy current losses. The end turn paths become a larger contributor to the resistance of the eddy current paths for shorter coils, i.e., lower stack length in case of radial flux machines or larger split ratio in case of axial flux machines. Therefore, it is necessary to consider the entire path of the eddy current loop. One typical approach is to employ correction factors, such as: Ks = 1 \u2212 tanh ( \u03c0Leff dc ) \u03c0Leff dc (1) where Leff is the effective length of the coil and dc is the conductor diameter", + " For example, if the real conductor dimensions are larger than the skin depth and the variations of flux density inside each conductor may be considerable, multiple flux samples in each plane are needed. The eddy current losses for a circular conductor can be calculated based on a general analytical formulation: Peddy = 1 R ( d\u03a6 dt )2 ; dR = (Leff + 2r)\u03c1\u221a d2c 4 \u2212 r2 dr (3) where R is the resistance; \u03a6, the magnetic flux through the conductors; and \u03c1 the conductor resistivity. Each conductor cross section can be divided into M segments as shown in Fig. 3. Assuming constant flux density for each segment, one field sample per segment may suffice. Therefore, the eddy current loss for one coil side with N\u03c6\u00d7 Nz turns can be estimated as: Peddy = 4L2 eff \u03c1 N\u03c6\u2211 n\u03c6=1 Nz\u2211 nz=1 \u221e\u2211 h=1 M\u2211 m=1 ... ( d dt ( 1 Leff \u222b OD ID Bm,h ( D 2 , n\u03c6, nz ) dD ))2 \u222b rm2 rm1 r2 \u221a d2c 4 \u2212 r2 (Leff + 2r) dr = Leff 2\u03c1 Ks N\u03c6\u2211 n\u03c6=1 Nz\u2211 nz=1 \u221e\u2211 h=1 M\u2211 m=1 ... [( d dt ( 1 Leff \u222b OD ID Bm,h( D 2 , n\u03c6, nz) dD ))2 (f(rm2)\u2212 f(rm1)) ] ; f(r) = r \u221a d2c 4 \u2212 r2 (2r2 \u2212 d2c 4 ) + ( d2c 4 )2 arctan r\u221a d2c 4 \u2212 r2 , (4) where Bm,h is the hth harmonic of the flux density in the mth section for the conductor associated with n\u03c6 and nz at the diameter D" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0001575_ls.1519-Figure1-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0001575_ls.1519-Figure1-1.png", + "caption": "FIGURE 1 Geometry of the aerostatic journal bearing", + "texts": [ + " In the combined method, the basic idea of analytic modelling in the traditional ESA method16 is referred to, and the CFD simulation is introduced for the detailed calculation. With the proposed approach, the load carrying capacity, frequency-dependent stiffness and damping can all be obtained. Based on the specified journal bearing, the analysis process is illustrated and the influences of bearing parameters on stiffness and damping are particularly discussed. The aerostatic journal bearing with orifice-type restrictor is shown in Figure 1. The externally pressurised gas passes through the air supply passage and the orifice, then feeds into the bearing clearance to form the air film, and finally escapes into the atmosphere from the outside edge of the bearing, that is, sides A and B. There are two rows of orifices along the bearing axis, which are named as row N1 and N2, respectively. Each row contains 12 orifices. The bearing diameter \u03a6 equals 98 mm, and the length L is 120 mm. The distance l between the orifice row and the same side bearing outside edge is 30 mm, and the orifice diameter d equals 0", + " In the flow field analysis, part of the air supply passage is considered upstream of the orifice. The pressure inlet condition is used at the inlet position and the pressure equals the air supply pressure. The pressure outlet condition is used at sides A and B, and the pressure equals the atmospheric pressure. Because of the symmetrical property of the bearing geometry, the air film is also geometrically symmetry about its middle cross section, that is, the position of the dash line as shown in Figure 1A. As thus, only half of the air film is considered in the analysis for simplification. The load carrying capacity, stiffness and damping of the whole model are twice of those of the half model. In analysis of the bearing dynamic performances, the small displacement perturbation method is commonly adopted. Firstly, the small displacement perturbation in the lateral direction of the spindle is imposed, and each physical quantity can then be regarded as summation of a steady term and a time-varying term,6,19,20 for example, the pressure p in the air film can be expressed as p= p+ p0, \u00f01\u00de where p is the steady pressure and p 0 is the time-varying pressure", + " However, as mentioned above, because of the complex air film geometry, these methods are not easy to be implemented in calculating the journal bearing mechanical performances. Therefore, the ESA-CFD combined method is proposed. In this study, the basic idea of analytic derivation of the ESAmethod is referred to, and with the application of CFD simulation, the ESA-CFD combined method is established. To expound the ESA-CFD combined method, the ESA method is firstly introduced,16 which includes three steps in calculation. In step 1, the air film as shown in Figure 1 is unfolded into a plane and it is divided uniformly into n parts along the circumferential direction as shown in Figure 3, where n is the orifice number per row. This simplification is due to the fact that the bearing diameter is far larger than the air film thickness, and the change of air film thickness in each divided part is small when the orifice number per row is enough, which are always satisfied in most engineering cases. In step 2, the hypotheses that \u201cgas flows only in each divided part\u201d and \u201conly 1D flow exists in each divided part\u201d are introduced", + " It should be noted that in fitting model of FLi xd\u00f0 \u00de, e0j and \u03c60 j are no longer considered in xd, because FLi is a time-independent quantity as discussed in Equation (6). The solution process of the ESA-CFD combined method is shown in Figure 6. 5 | APPLICATION OF THE ESACFD COMBINED METHOD AND ANALYSIS OF THE BEARING MECHANICAL PERFORMANCES In this work, the experiments are firstly conducted to verify the calculation accuracy of the combined method, and then, for the specified journal bearing configuration, the performances are calculated and the influences of bearing parameters on performances are discussed. The aerostatic journal bearing as shown in Figure 1 is considered for analysis. The given bearing parameters include: \u03a6 = 98 mm, L = 120 mm, d = 0.15 mm, l = 30 mm, n = 12, Ps = 0.6 MPa, atmosphere pressure Pa = 0.1 MPa. Firstly, the load carrying capacity is calculated by using the combined method, and the analysis results are compared with the experimentally achieved ones. Afterwards, the stiffness and damping are calculated, and the influences of bearing parameters h0, \u03b5 and \u03c6 on those dynamic performances are discussed in detail. The considered range of h0 is 5-15 \u03bcm, and the range of \u03b5 is 0-0" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0003961_s1068798x14070119-Figure1-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0003961_s1068798x14070119-Figure1-1.png", + "caption": "Fig. 1. Machining gear teeth in a worm gear by a hob with d0 > d1: (a) with the center of the contact spot (point K) in the midplane of the gear system (YK = 0); (b) with YK \u2260 0; (1) input; (2) output.", + "texts": [ + " This means that the rated values of the machine tool\u2019s interaxial distance aw0 and the splitting diameter d0 of the hob must be increased so that the curvature of the helical surface of the generating worm forming the lateral surface of the gear tooth is less than the curva ture of the working worm\u2019s turn, with the worst com bination of manufacturing errors of the gear system\u2019s housing and installation errors of the machine tool\u2019s interaxial distance. In machining gear teeth by a hob with an increased diameter d0, the rated value of the machine tool\u2019s interaxial distance aw0 must be aw0 = aw1 + 0.5(d0 \u2013 d1). Specifying aw0 \u2013 aw1 = 1.75fa, we obtain the required increment \u0394d0a = d0 \u2013 d1 in the hob diameter d0 relative to the worm diameter d1: \u0394d0a = 3.5fa. OF INCREASED DIAMETER FOR LONGITUDINAL MODIFICATION OF GEAR TEETH We now consider the machining of teeth in a simple gear (Fig. 1a) and the engagement of the correspond ing orthogonal worm gear. The diameter d0 of the hob\u2019s splitting cylinder is greater than the diameter d1 of the worm\u2019s splitting cylinder. However, the genera trices of these cylinders intersect at the engagement pole: point F. We will consider the case where the rated center of the contact spot (point K) is in the midplane of the gear system. In Fig. 1a (on the left side), we show the cross sec tion of the worm gear and engagement in the plane perpendicular to the worm axis and passing through the gear axis. The cross section A\u2013A is tangential to the worm\u2019s splitting cylinder (and that of the hob) and passes through a straight line parallel to the gear axis at a distance equal to the radius r2 of the gear\u2019s splitting cylinder (r2 = 0.5mxz2). This straight line is the engagement axis [1]. Its properties are as follows: 1. In all cases (in linear or localized contact), all the contact normals of the gear system (normals common to the active surfaces at the contact points) intersect this axis", + " In a gear system with longitudinally localized contact and a constant gear ratio, points of instanta neous contact of the active surfaces in immobile space form the engagement line, and this line must intersect the engagement axis. In view of these properties, it is expedient, in the synthesis of a gear system with longitudinal localiza tion, to select the point K at the center of the constant spot so that it lies on the engagement axis, either in the midplane of the gear system at point F (YK = 0) or at some distance toward the disengagement of the worm (YK > 0; Fig. 1b). For the synthesis of such a gear system, it is neces sary and sufficient to align the hob axis in machine tool engagement so that its generating surface touches the worm surface at this point\u2014in other words, so that the normals to these two surfaces meet at point K. The normal to any helical surface intersects its axis at a distance equal to the radius rb of the basic cylinder in the equivalent invo lute worm . In Fig. 1a (on the left side), the worm axis O1\u2013O1 is perpendicular to the plane of the page and corre sponds to point O1, while the basic cylinder is shown by a circle of radius rb. Therefore, the straight line that passes through point K and is tangential to a cylinder of radius rb is the projection of the common normal at the calculation point. At the same time, this tangent shows the direction of motion of the center of the con tact spot at the surface of the tooth in worm rotation. In Fig. 1a (on the right side), the worm axis O1\u2013O1 runs horizontally, while the hob axis O0\u2013O0 is inclined at an angle \u0394\u03b3. The cross section of the worm turn in plane A\u2013A tangential to the worm\u2019s splitting cylinder (identified by cross hatching) is in contact with the turn cross section of the hob\u2019s generating surface at point K in the midplane of the gear system. Straight line KC0 is the common normal to the lon gitudinal profiles of the turns at point K. It is inclined to the worm axis O1\u2013O1 at an angle \u03b31, equal to the splitting inclination of the worm turn; it is inclined to the projection of the hob at an angle \u03b30, equal to the splitting inclination of a turn of its generating surface", + " The distances C1K = R1 and C0K = R0 are the radii of curvature of these cross sections at the calculation point (1) Here \u03b1x1, \u03b1x0 are, respectively, the axial profile angles of the worm turn and the hob\u2019s generating surface. rb 0.5mx1z1/ \u03b1 2 x1tan \u03b3 2 1tan+( ) 1/2 = R1 0.5d1/ \u03b1x1tan \u03b3 3 1cos( );= R0 0.5d0/ \u03b1x0tan \u03b3 3 0cos( ).= If the calculation point K is to be on the worm\u2019s splitting cylinder in the midplane of the gear system (YK = 0), the inclination of the hob axis must be increased by \u0394\u03b3, where \u0394\u03b3 = \u03b31 \u2013 \u03b30. When the worm turns in the direction indicated by arrow \u03c9, the side of the profile that passes through point K is active, and the gear\u2019s entry and departure from engagement may be seen in Fig. 1a (on the right side). The measure \u03b41 of surface convexity (curvature) of the worm turn at the length of the gear tooth in the given cross section may be calculated from a parabolic formula: \u03b41 = 0.125 /(R1cos2\u03b31). The decrease \u0394\u03b4 in the convexity in that turn cross section of the hob\u2019s generating surface on account of increase \u0394R in the radius of curvature and correspond ingly in the splitting diameter is determined from the familiar equation [2] \u0394\u03b4 = \u20130.125 \u0394R/ The minus sign shows that \u0394\u03b4 and \u0394R have opposite signs", + " Thus, to obtain longitudinally localized contact in the worm gear by increasing the hob diameter, the increase \u0394d0m in the hob\u2019s splitting diameter relative to that of the working worm must be calculated so as to obtain the minimum required size of the contact spot with a gap at the spot boundary no greater than 0.005 mm. The hob\u2019s splitting diameter is d0 = d1 + \u0394d0m. When using hobs of enlarged diameter, it is possible to regulate the position of the center of the contact spot, as shown in [4]. By an increase in the inclination \u0394\u03b3 of the hob\u2019s axis, the contact spot may be shifted toward the point of worm disengagement. That improves the formation of a hydrodynamic lubricant layer between the surfaces of the worm turn and the gear tooth. In Fig. 1b, we show the hobbing process when \u0394\u03b3 > \u03b31 \u2013 \u03b30. At the left, we see that the line corresponding to cross section A\u2013A passes through the engagement pole. The point K must lie precisely on this line (the engagement axis) and is displaced along it by an amount YK in the direction of worm disengagement. The other notation in Fig. 1b is the same as in Fig. 1a. In the vicinity of the engagement pole F, the dis placement YK of the expected center K of the contact spot (depending on the inclination \u0394\u03b3 of the hob axis) is determined in terms of the radii of curvature R1 and R0 of the turn cross sections of the worm and hob (5)YK R0R1 \u0394\u03b3 \u03b30+( )sin \u03b31sin\u2013[ ]/ R0 R1\u2013( ).= Table 1 Precision class 6 7 8 9 10 11 12 |fa|, mm([fa]) 0.053 0.085 0.13 0.21 (0.195) 0.34 (0.31) 0.53 0.85 \u0394d0a, mm 0.186 0.298 0.455 0.735 1.19 1.855 2.975 |f\u03a3|, mm 0.012 0.017 0.022 0.028 0" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0000737_j.matpr.2020.01.583-Figure1-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0000737_j.matpr.2020.01.583-Figure1-1.png", + "caption": "Fig. 1. Schematic illustration of incorporation of polar functional groups (carboxyl and am adhesion, spreading and proliferation.", + "texts": [ + " The physicochemical properties of PAni:Ch nanocomposites before and after surface functionalization have been studied with the help of various characterization tools viz. SEM, I-V measurement, FTIR, and XPS. The effect of altered surface properties of glycine NHS ester functionalized PAni:Ch nanocomposites has been investigated to modulate the blood compatibility by Hemolysis assay and cytotoxicity by MTS proliferation assay with 3 T3 fibroblasts followed by cell adhesion, spreading and proliferation by live/dead assay and SEM. The schematic representation of the work carried out has been shown in Fig. 1. Aniline (p.a. Merck Germany, 99.5%) was distilled under reduced pressure before use. Ammonium peroxydisulfate (APS) (p.a. Merck Germany, 98%) and Hydrochloric acid (HCl) (p.a. Sigma Aldrich, 37% AR grade) were used without further purification. Deionised water (12 MX cm) used for the synthesis was obtained from a Milli-Q system. Chitosan extrapure (p.a. SRL India, 99%) and BOC glycine N-Hydroxysuccinimide (NHS) ester (p.a. Acros Organics, 99%) were used as obtained. Acetic acid was purchased from (p" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0002541_iros45743.2020.9340742-Figure6-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0002541_iros45743.2020.9340742-Figure6-1.png", + "caption": "Fig. 6: Approach 3: Initial pose of phone as the shared frame of reference. With the phone mounted on the robot, when the AR application starts, the global (initial pose) frame of phone links the virtual object location to its actual real-world coordinate on the map.", + "texts": [ + " As in Approach 2, the robot is localized on the mapped area, and the pose of the phone on the map is computed based on the robot\u2019s pose. The phone is initially (when the AR application starts) mounted on the robot with a fixed frame of reference in the robot\u2019s coordinate system. The global axis in the AR scene is the real-world location of the phone where the AR application starts. In other words, the location of the virtual object in global axis (globalPobject) is same as that in the phone\u2019s initial coordinate frame (piPobject), i.e., globalPobject = piPobject (see Fig. 6). Once the AR application starts, the user can take off the mounted phone from the robot and walk around the area to mark any locations by creating virtual objects. The global pose of the virtual object is transformed into the map\u2019s coordinate frame through a set of transformations given below. Step 1: Transform the global pose of virtual object in the AR system (piPobject)AR to the initial pose of phone in the robot system (piPobject)robot as in (2). Step 2: Compute the virtual object location in the map\u2019s frame mapPobject = (mapTpi)( piPobject)robot" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0002496_3443467.3443832-Figure1-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0002496_3443467.3443832-Figure1-1.png", + "caption": "Figure 1: Structure diagram of a four-rotor UAV", + "texts": [ + " Therefore, on the basis of ADRC, this article replaces the non-singular error feedback rate link with nonsingular terminal sliding mode control to obtain the SM-ADRC, and at the same time, the non-smooth fal function in the composite controller is improved. The improved composite controller is applied to the attitude control of the aircraft, and combined with simulation analysis, the effectiveness of the method is verified. In order to facilitate the modeling of the quadrotor UAV, the aircraft is simplified as shown in Figure 1. Permission to make digital or hard copies of all or part 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. Copyrights for components of this work owned by others than ACM must be honored. Abstracting with credit is permitted. To copy otherwise, or republish, to post on servers or to redistribute to lists, requires prior specific permission and/or a fee. Request permissions from Permissions@acm.org. EITCE 2020, November 6\u20138, 2020, Xiamen, China \u00a9 2020 Association for Computing Machinery. ACM ISBN 978-1-4503-8781-1/20/11\u2026$15.00 https://doi.org/10.1145/3443467.3443832 Establish two basic coordinate systems as shown in Figure 1: the ground coordinate system E E E EO X Y Z and the body coordinate system B B B BO X Y Z .Among them, the roll angle , pitch angle , and yaw angle are the angles of the airframe coordinate system around the X, Y, and Z axes and the ground coordinate system. Therefore, the body coordinate system can coincide with the ground coordinate system after three axis rotations, and the conversion relationship is shown in (1) [10-11]: + + Z Y X \u03a8 \u0398 \u0398 \u03a6 \u03a8 \u03a8 \u03a6 \u03a6 \u03a8 \u0398 \u03a6 \u03a8 \u0398 \u03a8 \u03a6 \u03a8 \u0398 \u03a6 \u03a8 \u03a6 \u03a8 \u0398 \u03a6 \u03a8 \u0398 \u03a6 \u0398 \u03a6 \u0398 T T T T C C S S C S C C C S S S C S S S S C S C S S S C S S C C C (1) In the formula, T is the conversion matrix, and S and C are respectively expressed as sin and cos functions" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0003183_icra.2015.7139672-Figure2-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0003183_icra.2015.7139672-Figure2-1.png", + "caption": "Fig. 2. Grain cart", + "texts": [ + " For most of the combines, the auger is mounted on the left side, as shown in Fig. 1. At this point, a vehicle has to be on the left side of the combine to empty the tank. Here we denote the unloading rate of the tank using auger as ru. Therefore, a combine involved in harvesting and unloading operation simultaneously unloads at a rate ru\u2212 r f (ru > r f ). A grain cart, also known as chaser bin, is a trailer towed by a tractor. In this paper, we use term grain cart to represent the system including both the tractor and the trailer. Fig. 2 (a) shows the appearance of a grain cart. Because of the larger capacity, one can use it to collect grains from multiple combines so that the combines could work without interruption. Figure 2 (b) shows a grain cart. We model the grain cart as a trailer attached to a car-like robot. The robot is hitched by the trailer at the center. The equations of motion for the grain cart are as follows: q\u0307 = x\u0307 y\u0307 \u03b8\u0307 \u03b2\u0307 = vcos\u03b8 vsin\u03b8 \u03c9 \u2212vsin\u03b2 +\u03c9 , where q = (x,y,\u03b8 ,\u03b2 ) \u2208 R2\u00d7S1\u00d7S1 is the configuration, u = (v,\u03c9) \u2208 U = [\u22121,1]2 is the control [15] [16]. In the configuration q, (x,y) is the coordinate of robot\u2019s center, \u03b8 is the robot\u2019s orientation, \u03b2 is the angle between trailer\u2019s orientation, and the robot\u2019s orientation" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0002923_9781119633365-Figure7.16-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0002923_9781119633365-Figure7.16-1.png", + "caption": "Figure 7.16 Code configurations of the high step-down transfer code D/(2 \u2212\u00a0D).", + "texts": [ + " In the following, derivation of the several examples is presented. 7.3.1.1 High Step-Down Converter with\u00a0Transfer Code D/(2 \u2212\u00a0D) A high step\u2010down switched\u2010inductor converter with the transfer code D/(2 \u2212 D) is\u00a0 first decoded and synthesized as follows. Through cross multiplication, Vo/Vi\u00a0= D/(2 \u2212 D) can be expressed as V D V Do i2 . (7.14) or V D V D Vi o o1 . (7.15) 7 Converter Derivation with\u00a0the\u00a0Fundamentals168 Equation\u00a0(7.15) can be put into a feedback and feedforward form as V V D D D Vi o o 1 . (7.16) Equation\u00a0(7.16) can be configured as shown in Figure\u00a07.16a. One more possibility to decode transfer code D/(2 \u2212 D) is shown as follows: V V D D D D o i 2 1 2 . (7.17) Define V V Do i/ /1 2 and we have V V D V V o i o i . (7.18) Through cross multiplication, V Vo i/ can be expressed as V V D o i 1 2 (7.19) V Vo i2 D . (7.20) or V V D Vi o o1 . (7.21) Combining the feedback expression of (7.21) with a forward term D yields the code configuration shown in Figure\u00a07.16b. To synthesize the code configuration shown in Figure\u00a07.16a, we use a buck converter as a forward path and use I\u2010buck\u2013boost as a feedback path, as shown in Figure\u00a0 7.17a. To establish a common node for switches S1 and S2, switch S1 is moved to the return path, as shown in Figure\u00a0 7.17b. It can be identified that switches S1 and S2 have a common D\u2013S node and they can be replaced with a\u00a0 \u03a0\u2010type grafted switch, as shown in Figure\u00a0 7.17c. Next step is to find the relationship between the currents through switches S1 and S2. When both switches are turned on simultaneously, the voltage across inductor L1 is Vi(1 \u2212 D)/(2 \u2212 D)", + " To satisfy with ampere\u2010second balance, there is no average current through capacitor C2 in the steady state. Therefore, the currents through switches S1 and S2 are identical, diodes DF1 and DF2 are no longer needed, and capacitor C2 can be removed from the converter, too. The converter becomes the one shown in Figure\u00a07.17d, in which there is only switch S12 left. Rearranging the components in a special form will yield the high step\u2010down converter as shown in Figure\u00a07.17e. For the code configuration shown in Figure\u00a07.16b, although there is an I\u2010boost converter that can serve as the feedback path, there is no converter to synthesize the unity when its input is combined with the negative feedback path. Thus, even though we have a correct code configuration, it is not necessary to have converters to synthesize the code configuration because there is no isolation transformer in the synthesis. Code configuration is not unique for a given transfer code. Except to the configuration shown in Figure\u00a07.16, transfer code D/(2 \u2212 D) can be expressed alternately as follows, and they can be configured into different code configurations: V V D D o i 2 . (7.22) With cross multiplication, we have V D V Do i2 , (7.23) or V D V D V V V D Do i o i o1 . (7.24) or V V V D D Do i o 1 . (7.25) A code configuration to realize (7.25) is shown in Figure\u00a07.18a, in which Vo has a feedback gain \u22121/D, combined with input Vi, which is fed to a forward gain D/(1 \u2212 D). With similar processes shown in (7.17)\u2013(7.21), we can modify the code configuration shown in Figure\u00a07", + " Therefore, we have the same current through switches S1 and S2 under their on states, and diodes DF1 and DF2 conduct no current, so does capacitor C2, and they can be removed from the converter, as shown in Figure\u00a07.19d. After rearranging the component positions, we have the final high step\u2010down switched\u2010inductor converter, as shown in Figure\u00a07.19e. For the code configuration shown in Figure\u00a07.18b, since the forward path has a transfer code of 1/(1 \u2212 D), it will be synthesized with a boost converter. However, a boost converter has no negative feedback. Thus, there is no converter to synthesize the code configuration shown in Figure\u00a07.18b. Similar to that of Figure\u00a07.16b, not all of the valid code configurations can be synthesized with converters. For instance, other code configurations for achieving the transfer code D/(2 \u2212\u00a0D) are shown in Figure\u00a07.20. However, there is no converter to synthesize the transfer code D/2 shown in Figure\u00a07.20a, and there is no negative feedback for an I\u2010buck converter with the transfer code 1/D, either, as shown in the inner loop of Figure\u00a07.20b. Comparing the code configurations shown in Figures\u00a0 7.16b, 7.18b, and 7.20 with those shown in Figures\u00a07" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0001917_icuas48674.2020.9214023-Figure4-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0001917_icuas48674.2020.9214023-Figure4-1.png", + "caption": "Fig. 4. Plots of the LSL and LSR paths to the straight line segment AB", + "texts": [ + " 1) First derivative of the length of the LSL path: The coordinates of the position corresponding to \u03bb on AB is given as (ax +\u03bbvx,ay +\u03bbvy), where vx = bx\u2212ax and vy = by\u2212ay. The length of the line segment AB, lAB = \u221a v2 x + v2 y . Let \u03b3 be the orientation of the line segment AB with respect to the x\u2212axis. Let \u03c6 LSL 1 and \u03c6 LSL 2 be the angles subtended by the first and second arc of the LSL path; and let LLSL s be the length of the second segment (straight line) of the LSL path. A sample LSL path to AB is shown in Fig. 4(a). The sum of the angles \u03c6 LSL 1 and \u03c6 LSL 2 are equal to \u03b3 + 2n\u03c0 , n = 0,1,2. Therefore, the derivative of \u03c6 LSL 1 +\u03c6 LSL 2 with respect to \u03bb is zero. Using elementary geometry, the length, LLSL s is derived and its derivative are shown below. LLSL s = \u221a (ax +\u03bbvx\u2212\u03c1 sin\u03b3)2 +(ay +\u03bbvy +\u03c1 cos\u03b3\u2212\u03c1)2. (1) d d\u03bb LLSL s = 1 LLSL s (vx(ax +\u03bbvx\u2212\u03c1 sin\u03b3) + vy(ay +\u03bbvy +\u03c1 cos\u03b3\u2212\u03c1)), = vx cos\u03c6 LSL 1 + vy sin\u03c6 LSL 1 , = lAB cos(\u03b3\u2212\u03c6 LSL 1 ). (2) Let \u2206T LSL(\u03bb ) be the difference in time of arrival by the target and the pursuer following an LSL path, \u2206T LSL(\u03bb ) = ta + \u03bb lAB vt \u2212 LLSL(\u03bb ) vp ", + " Proof: Equating the first derivative to zero gives the following, d d\u03bb \u2206T LSL = lAB vt \u2212 1 vp lAB cos(\u03b3\u2212\u03c6 LSL 1 ) = 0 The length, lAB cannot be zero, and the above equality gives cos(\u03b3\u2212\u03c6 LSL 1 ) = vp vt . If vt < vp, then it results in an infeasible value for cos(\u03b3\u2212 \u03c6 LSL 1 ). This implicates that d d\u03bb \u2206T LSL never equals zero, and no extremum exists. 2) First derivative of the length of the LSR path: Let \u03c6 LSR 1 and \u03c6 LSR 2 be the angles subtended by the first and second arc of the LSR path, and let LLSR s be the length of the second segment (straight line) of the LSR path. A sample LSR path is shown in Fig. 4(b). Using elementary geometry, the length LLSL s and the angles \u03c6 LSR 1 and \u03c6 LSR 2 are derived and their derivative are shown below. Let us define lx and ly as lx := ax +\u03bbvx +\u03c1 sin\u03b3 and ly := ay +\u03bbvy\u2212\u03c1 cos\u03b3\u2212\u03c1 . 802 Authorized licensed use limited to: University of New South Wales. Downloaded on October 18,2020 at 17:32:00 UTC from IEEE Xplore. Restrictions apply. LLSR s = \u221a L2 cc\u22124\u03c12, (4) \u03c6 LSR 1 = \u03c81 +\u03c82, (5) \u03c6 LSR 2 = \u03c6 LSR 1 \u2212 \u03b3, (6) where Lcc,\u03c81 and \u03c82 are given as below, Lcc = \u221a l2 x + l2 y , \u03c81 = arctan ly lx and \u03c82 = arctan( 2\u03c1 LLSR s )" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0000058_icmae.2019.8881011-Figure3-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0000058_icmae.2019.8881011-Figure3-1.png", + "caption": "Figure 3. Coordinated System of Kinematic Chains", + "texts": [ + " (4) (5) where are the points of the mobile platform obtained from Table II. B. Inverse Kinematics In the previous section it found expressions for the location of the fixed platform in , and for the mobile platform in in terms of the dimensions of the edges that make up the hexagon, as well as the desired position and orientation. Proceed with the mathematical analysis of the extremities to find the solution to the inverse kinematic problem. To do this, each arm is placed in a coordinated system as shown in figure 3. From the above, it obtain vectorially the expression that defines from the triangle OPL that, being a non-rectangle triangle, its sides are related by the cosine law together with the angle and mathematically described in the equation (3), (1) (1) (3) and given that the value of is unknown, we apply the definition of the product point as it is shown in the equation (4), to finally obtain the equation (5) that mathematically expresses . Where, y . In addition, it noted that the coordinate system for each arm rotates with angles described in (6) and defined in figure 2" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0001381_s1068798x20060210-Figure6-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0001381_s1068798x20060210-Figure6-1.png", + "caption": "Fig. 6. Vil\u2019ner\u2019s LIOT test bench.", + "texts": [ + " In 1966, for the first time, the SN626\u201366 standard laid out health requirements and rules for the operation of tools and devices transmitting vibrations to their operators\u2019 hands. From that point on, manufacturing enterprises developed test benches for determining the vibrational parameters of such equipment. The first theoretical work on this topic was undertaken at the Leningrad Institute of Worker Health and Safety by G.S. Vil\u2019ner [7]. He believed that the hammer vibrations in the test bench must be identical to those when used by an operator in order to obtain objective results. On that basis, he developed the LIOT test bench (Fig. 6). Guides 2 and 3 on either side of column 1 permit the attachment of displacement sensors. At the top of the column, arm 4 has two bushes 5 that serve as guides for rod 6. The upper end of the rod is threaded. In rotation of f lywheel 7, the rod moves vertically in bush 5. Guide bush 9 connected to clamp 10 is attached to the lower end of the rod by pin 8. Cylindrical spring 11, which moves in bush 9, supports load 12 (mass 20 kg). Ball bearings are placed within bush 9 and load 12. The jackhammer 13 is mounted on tool 14, which has a spherical working part" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0002216_icem49940.2020.9270977-Figure1-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0002216_icem49940.2020.9270977-Figure1-1.png", + "caption": "Fig. 1: IM cross section.", + "texts": [ + " The paper is organized as follow: 1) section II reports relevant data of the induction motor considered in this study; 2) section III reports the test bench description and the measurement procedures; 3) section IV describes the equations adopted for flux orientation in the measurement post-process; 4) section V briefly presents the simulation strategy based on Rotor Field Oriented Analysis (RFOA); 5) section VI compares the experiment flux measurement with simulation results; 6) section VII shows the estimated quantities throughout the flux measurement. II. IM UNDER CONSIDERATION In Figure 1 the IM geometry is reported, whereas in Table I the geometric, winding and nominal electrical data are shown. The IM here considered has a skewed rotor with closed slots. III. MEASUREMENT SETUP AND PROCEDURES The idea is to supply the machine with a given stator current and frequency. On the other hand, the rotor speed will be varied controlling the master motor speed. In this way, from the synchronous speed, the rotor is gradually slowed down, to the standstill condition. Inside the motor, the stator current will be differently divided into d- and q-axis components, at the different imposed rotor frequencies" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0001511_jsyst.2020.3006990-Figure10-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0001511_jsyst.2020.3006990-Figure10-1.png", + "caption": "Fig. 10. Typical pulley with a belt tensioned around it.", + "texts": [ + " Using the upper bounds obtained for different segments of the optimal path under the 3PA, an upper bound can be obtained for the overall optimal path under the 3PA as summarized in Table II. The main step and distinguishing feature of each algorithm described in the previous section is the derivation of the orientation angle at each viewpoint. It is desired now to obtain this angle based on a natural physical system, namely pulley. Given some viewpoints, the arriving/departing angles at/from each one can be considered as those in a pulley depicted in Fig. 10 with a belt tensioned around it. A new algorithm is introduced accordingly, called the PA, as described in the following two steps. 1) Place a circle of radius rmin at each viewpoint such that it intersects the arriving and departing paths in two equal chords with the viewpoint in the middle of the arc that starts with the first intersection and ends with the second one. 2) Connect every pair of consecutive circles with a line tangent to both. The PA can be illustrated by a simple 3-viewpoints example depicted in Fig" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0002146_icem49940.2020.9270956-Figure1-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0002146_icem49940.2020.9270956-Figure1-1.png", + "caption": "Fig. 1. Sketch of the rotor center with respect to the stator center. (a) healthy, (b) static eccentricity, and (c) dynamic eccentricity", + "texts": [ + " Gerada1 are with the Power Electronics, Machines and Control group, the University of Nottingham, Nottingham, UK. (e-mail: michele.degano@nottingham.ac.uk) H. Mahmoud2 is with the Electric Power and Machines Department, Cairo, University, Cairo, Egypt. (e-mail: eng.hanafy4@yahoo.com) M. Degano3, Xiaochen Zhang3 and C. Gerada3 are with the Power Electronics, Machines and Control group, the University of Nottingham, Ningbo, China. (e-mail: michele.degano@nottingham.ac.uk) vibrational analysis is conducted by FE simulations based on different conditions of the rotor as shown in Fig. 1. Their vibrational results are compared for figuring out the effects of eccentric rotor. The cases shown in Fig. 1 are considered for understanding the difference between the healthy (concentric rotor), static eccentric and dynamic eccentric rotor position. As shown in Fig. 1 (b), which is a static eccentricity, the rotor is shifted from the center point to the stator for 0.15 mm, but still rotates around its own center. As shown in Fig. 1 (c), the rotor is similarly shifted from the center point to the stator for 0.15 mm and rotates around the stator center. Both two cases are having an uneven airgap, that there is a smaller airgap on the positive x-axis direction. The healthy rotor is also considered in the case study as a comparison, which is shown in Fig. 1 (a) [7]. A 4 pole/36 slots PMaSynRel machine with three layers of flux barriers per pole is adopted here as shown in Fig. 2, and its main dimensions are list in Tab. I [8]. TABLE I MAIN DIMENSIONS OF STUDIED MACHINE Parameters Value Parameters Value External stator diameter 200 mm Number of slots 36 Internal stator diameter 125 mm Number of poles 4 Airgap 0.35 mm Rated current 5.29 A Stack length 40 mm Rated torque 12.5 Nm Impact on Vibration of Eccentric Permanent Magnet Assisted Synchronous Reluctance Machine Jiaqi Li1, Hanafy Mahmoud2, Michele Degano1,3, Anuvav Bardalai1, Xiaochen Zhang3, Chris Gerada1,3 P Authorized licensed use limited to: Carleton University. Downloaded on May 31,2021 at 17:56:31 UTC from IEEE Xplore. Restrictions apply. Fig. 2 shows a PMaSynRel machine in healthy condition, which means it has a concentric rotor and the machine is rotating around the center. Then, the rotor is moved 0.15 mm along the d- axis to simulate the static and dynamic eccentric conditions. In addition, the three cases shown in Fig. 1 are analyzed and compared in the following sections for their electromagnetic forces, airgap flux densities, and finally their vibrational level. For the radial flux permanent magnet machines, the vibration is mainly caused by radial force. The radial force density can be computed through [9] 2 2 0 1 ( ) 2 r tB B\u03c3 \u03bc = \u2212 (1) from where represents the radial force density, 0 represents the free space permeability, Br and Bt represent the airgap flux density in radial and tangential direction, respectively" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0002989_978-3-642-16214-5_9-Figure9.24-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0002989_978-3-642-16214-5_9-Figure9.24-1.png", + "caption": "Fig. 9.24. Double-cone synchronizer (ZF-D). 1 Idler gear;", + "texts": [ + " Multi-cone synchronizers are therefore increasingly being used in the low gears (first and second). The number of friction cones and the friction materials used will depend on the intended use. 2 synchronizer hub with dog gearing; 3 double-cone ring; 4 counter-cone ring; 5 synchronizer body For example, the use of a triple-cone synchronizer may make it possible to use low-cost special brass synchronizer rings [9.5]. Two cone friction surfaces are required to achieve synchronization in the case of double-cone synchronizers (Figure 9.24). The link between the double-cone ring 3 and the synchronizer hub 2 is rotationally fixed using several dogs, but axially flexible. The counter-cone ring 4 is rotationally fixed to the synchronizer body 5. The gearshift effort is reduced and the torque capacity, i.e. performance, is increased because of the increased number of friction surfaces and the larger friction surface area of the double-cone synchronizer. The parallel multi-cone design requires closer manufacturing tolerances and therefore entails higher production costs" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0000331_icaaid.2019.8934965-Figure1-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0000331_icaaid.2019.8934965-Figure1-1.png", + "caption": "Figure 1. Internal and Body Frame of Quadcopter", + "texts": [ + " The grounding effect and gyroscopic effect may present other terms of nonlinearities that if a designer set nonlinear controllers these nonlinearities are for sure going to be compensated and dealt with for granted. Unfortunately for simplicity reasons this model pays attention to the moment and acceleration equations that we will tend to control in the next chapter as our main focus will be to control the altitude of the quadcopter while pitch, roll and yaw angles will be null and allow our model to hover and attempt trajectory tracking along one axis. The quadcopter structure is presented in Figure 1 including the corresponding angular velocities, torques and forces created by the four rotors (numbered from 1 to 4) Authorized licensed use limited to: UNIVERSITY OF BIRMINGHAM. Downloaded on June 13,2020 at 19:29:11 UTC from IEEE Xplore. Restrictions apply. The absolute linear position of the quadcopter is defined in the inertial frame x,y, and z axes with \u03be. The attitude (the angular position) is defined in the inertial frame with three Euler angles \u03b7. Pitch angle \u03b8 determines the rotation of the quadcopter around the y-axis" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0002084_physrevfluids.2.033901-Figure2-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0002084_physrevfluids.2.033901-Figure2-1.png", + "caption": "FIG. 2. Left frame (a) shows a sketch of an inverted pendulum, where the vertical position is unstable. Frame (b) shows an analogous fluid problem, consisting of a cylinder and splitter plate exposed to incoming flow U\u221e. Instead of the gravitational force, pressure force arising due to the reverse flow UR in the wake of the cylinder act as destabilizing force, turning the plate out of the backflow region, either to the right or to the left. In frame (c), the quantities defined within the IPL model are shown. The model of backflow region (MBFR) is shown in gray. Figure adapted from Ref. [5].", + "texts": [ + " We also observe that for very short appendages, the turn angle approaches some finite value, which is close to the wake attachment angle \u03b8a \u2248 55\u25e6 for the cylinder alone. The drift angle, on the other hand, has a maximum value for intermediate plate lengths, and approaches zero drift for very small appendages. This is expected, since the cylinder alone does not exhibit any transverse motion. The observations above can be explained by a simple\u2014yet quantitative\u2014model based on an analogy to an inverted pendulum confined between two vertical walls. The sketch in Fig. 2(a) shows a pendulum consisting of a circular cylinder and a plate. The body is free to rotate around the center of the cylinder, which is some distance below the center of mass of the pendulum. The straight vertical )b()a( (c) 033901-3 )b()a( position of the pendulum is thus an unstable equilibrium and any small disturbance will make it fall either to left or to the right. The fluid-structure-interaction mechanism of the freely falling body is similar; the only difference is that instead of gravitational forces the pressure forces are acting on the plate to destabilize it, as illustrated in Fig. 2(b). In other words, if the splitter plate is sufficiently short, in the presence of a small disturbance the pressure forces will turn the splitter plate out of the backflow region (i.e., the region with a significant reversed flow behind the cylinder). However, as the plate is pushed out of the backflow region, it is exposed to forward flow, which provides a stabilizing force and acts in a similar way as the wall acts for the inverted pendulum [Fig. 2(a)]. In order to obtain more quantitative predictions, a model of the backflow region can be defined as shown in Fig. 2(c). Here, B(\u03b8 ) is the distance from the cylinder surface to the point on the plate where the normal force on the plate changes sign. Figure 3(a) shows these points (with black dots) on the plate identified from a series of simulations of the flow around cylinder with splitter plate at various equilibrium turn angles. It is, however, not always possible\u2014for example, in many experiments\u2014to directly evaluate the force distribution on the appendage. An estimate of the backflow region can be made from simulations or experimental measurements of the wake of a body without the appendage", + " Both calibration coefficients CA and k can be estimated with reasonable accuracy from measurements of the wake without a splitter plate [5]. Using the model of normal forces, we can obtain equilibria angles by finding zero torque around the center of the cylinder. The total torque due to the splitter plate around the center of cylinder is T (\u03b8 ) = F+ n [ D 2 + B(\u03b8 ) 2 ] + F\u2212 n [ D 2 + B(\u03b8 ) 2 + L 2 ] . (3) To construct the torque expression, it has been assumed that the force on the plate is located at the center of each segment of the plate, as illustrated in Fig. 2(c). Inserting the expressions for normal forces into (3), we obtain T (\u03b8 ) = sin(\u03b8 ){(1 + k)[B\u03022(\u03b8 ) + B\u0302(\u03b8 )] \u2212 L\u03022 \u2212 L\u0302}CA\u03c1fU 2 \u221eD2, (4) where B\u0302 = B/D and L\u0302 = L/D. By choosing an appropriate value for k based on appendage-free wake measurement, one can find equilibrium torque angles \u03b80 (i.e., angles for which T (\u03b80) = 0) for each splitter plate length. The results from the IPL model are compared to the direct numerical simulation (DNS) results of freely falling cylinder with splitter plate at Re = 45 in Fig" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0003251_icuas.2015.7152410-Figure8-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0003251_icuas.2015.7152410-Figure8-1.png", + "caption": "Figure 8 Grid of a couple of interface surface of two domains", + "texts": [ + " The grid densification is enforced not only at the boundary layer development locations, but also in the rotor wake and on the path of the tip and root vortices (Fig. 7.b). a Sectional surface mesh of the stationary domain at spanwise direction is exhibited in Fig. 7.c. c c) Structured O-H grid for stationary domain, cut at r/R=0.7 projected on y=const plane The rotary and stationary domain have three pairs of interface surfaces, the overlapping grids at a couple of interface surfaces are shown (Fig. 8). Grid data of one interface surface of is delivered to the other one through numerical interpolation. (4) Boundary Condition The two computational domains are surrounded by boundary conditions of six types: viscous and inviscid walls, velocity inlet, pressure outlet, symmetry (Fig. 9). The surfaces of the blade are modelled as a moving, no-slip condition with a zero heat flux wall-black. To save computational resources, a shaft surface, the port and starboard surface, are modelled as an inviscid wall-green" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0000623_012064-Figure6-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0000623_012064-Figure6-1.png", + "caption": "Figure 6. The pattern of stress intensity distribution that occurred when the rollers interact with the bearing rings when a load is applied.", + "texts": [ + " Series: Materials Science and Engineering 760 (2020) 012064 IOP Publishing doi:10.1088/1757-899X/760/1/012064 SibTrans-2019 IOP Conf. Series: Materials Science and Engineering 760 (2020) 012064 IOP Publishing doi:10.1088/1757-899X/760/1/012064 In the process of modeling, it was assumed that a vertical uniformly distributed calculated load of 107H acts on the thrust surfaces of the housing of the axle box. The general picture of the stress intensity distribution arising in the roller of the axle box under this load is shown in figure 6. From the figure it can be seen that in the rollers of the rear bearing the stress level is higher than in the rollers of the front one. At the same time, the third roller (with respect to the vertical axial section of the axle box), the rear bearing roller, experiences the greatest load. When considering the features of contact and rings, a preliminary breakdown of the rollers and bearing rings into fragments was used to achieve the required modeling accuracy (Figure 7a). In this case, each of the fragments was split on the FE independently with increasing concentration as it approaches the contact surface and uniformly within the fragment (Figure 7b)" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0002776_978-94-007-6046-2_29-Figure3-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0002776_978-94-007-6046-2_29-Figure3-1.png", + "caption": "Fig. 3 An example of classical \u201cshape\u201d for the transition function, usually empirically defined. The maximum distance between the feet is limited by the length of the legs and kinematic constraints. The transition function should also prevent self-collisions from occurring, so El and Er are defined in a way that keeps the feet at a safe distance from each other. A constraint also bounds the difference of orientation of the feet, as very different orientations can typically result in collisions between the knees or violations of kinematic constraints at the hips", + "texts": [ + " There are different ways to define the goal region, and we choose to define it as a subset of F FS , reached as soon as one of the footsteps belongs to it. We can now formally specify the input and output of footstep planning in 2D: Data: \u2013 S W C FS ! P .C FS /, the transition function \u2013 F FS , the free space (usually known implicitly via the description of the obstacles) \u2013 .s0; s1/ 2 F 2 FS : the initial state, verifying s1 2 S.s0/ \u2013 G F FS : the goal region Result: A sequence of footsteps s2, s3, . . . , sn 2 F FS such that 8i 2 f1; : : : ; n 1g, siC1 2 S.si /, and sn 2 G. Figure 2 illustrates footstep planning in 2D, and Fig. 3 shows a classical example of transition function model for a humanoid robot. An ideal transition function should have a relatively simple structure and yet model accurately the stepping capabilities of the robot. For a given robot with a given walking algorithm, a good Fig. 4 Flea motion planning: a variant of 2D footstep planning with state-space R 2 choice for the function S can be progressively obtained after thorough tests and experiments. Conservativeness of the transition function is always preferred over allowing infeasible steps" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0001494_ur49135.2020.9144981-Figure3-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0001494_ur49135.2020.9144981-Figure3-1.png", + "caption": "Fig. 3: The prototype of sensing device", + "texts": [ + " In this section, we explain the latest sensing device design: device concept, hardware composition and control system. Our basic development characteristics are low power consumption, omni-directional motion design, low cost and portability. These concepts realize long-term robust monitoring and high usability water environment monitoring system. \u2022 Low power consumption: Since it is difficult to supply electric power to USV 495 Authorized licensed use limited to: Macquarie University. Downloaded on July 25,2020 at 19:00:28 UTC from IEEE Xplore. Restrictions apply. Fig. 3 shows the whole image of the device. In autonomous surface navigation, it is not easy to supply electric power to the devices when they are deployed on water. Therefore, it is required to suppress electric power consumption. Fig. 5: The sensing device in fixed coordinate Fig. 6: The relationship between the world coordinate and the sensing device in fixed coordinate To achieve this requirements, the sensing device is consists of low power consumption components. The device is compact to be deployed easily" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0001701_0954405420949757-Figure10-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0001701_0954405420949757-Figure10-1.png", + "caption": "Figure 10. Both sides chamfering tool path.", + "texts": [ + " Distance between discrete point (x0Ak, y 0 Ak, z 0 Ak) and Z1 axis is denoted by rj, which is calculated as rj = ffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi x0Ak 2 + y0Ak 2 q \u00f028\u00de When cutter surface contacts with upper edges of convex side and concave side, the cutter position is constrained by d0A = rc and d00A = rc, coordinates solution of cutter center is transformed into an optimization problem as follow min f= d0A rcj j+ d00A rcj j s:t:rj \\ r \\ rj +2rc, 0\\ u \\ 2p \u00f029\u00de This optimization problem is solved by particle swarm algorithm, number of particles is set to 50, number of iterations is set to 10,000, solving convergence condition is f\\ 10 4mm. When P j is given, a convergence solution with r and u is obtained, then coordinates of cutter center are calculated by r and u. When subscript j varies from 1 to n, different coordinates of cutter center are obtained. Both sides chamfering tool path is obtained by linking all cutter centers, as shown in Figure 10. Due to extensions at inner and outer ends along tooth width direction, on both sides chamfering tool path in Figure 10, segment PAPB only performs concave side chamfering in practice, segment PCPD only performs convex side chamfering in practice, segment PBPC performs both sides chamfering. Chamfering dimension error of j-th point of both sides chamfering on lower edges of concave and convex sides are calculated as ej = d0S rcj j+ d00S rcj j \u00f030\u00de Maximum value among ej is called maximum chamfering dimension error and denoted by emax, which is expressed as emax= max (e1, e2, :::ej, :::en), j 2 (1, 2, :::n) \u00f031\u00de Suppose that cutter surface and upper edge of convex side contact at points of P0B, P 0 C, P 0 D in Figure 11 when cutter center moves to points of PB, PC, PD in Figure 10 in turn. Similarly, suppose that cutter surface and upper edge of concave side contact at points of P00A, P 00 B, P00C in Figure 11 when cutter center moves to points of PA, PB, PC in Figure 10 respectively. Points of P0B, P 0 C, P0D, P 00 A, P 00 B, P 00 C divide chamfering process into three phases including concave side phase, both sides phase and convex side chamfering phase, as shown in Figure 11. Chamfering tool path for concave side chamfering phase Segment P00AP 00 B in Figure 11 is processed in concave side chamfering phase. Cutter surface contacts with upper and lower edges of concave side, the position relationship between cutter and pinion is constrained by d00S = rc and d00A = rc, coordinates solution of cutter center is transformed into an optimization problem as follow min f= d00S rcj j+ d00A rcj j s:t:rj \\ r \\ rj +2rc, 0\\ u \\ 2p \u00f032\u00de This optimization problem is solved by particle swarm algorithm with the same parameters mentioned above", + " Coordinates of point P00B in Figure 11 is denoted by (x00B, y 00 B, z 00 B), if the distance between cutter center and point P00B is larger than cutter radius, interference between cutter and tooth crest of convex side can be avoided, therefore a constraint is added into tool path solving process of concave side chamfering, which is written as ffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi (r cos u x00B) 2 + (r sin u y00B) 2 + (zj z00B) 2 q . rc \u00f033\u00de Cutter can not approach the region near point P00B after the constraint above is added, a part of segment PAPB near point PB in Figure 10 is chosen to process the region near point P00B in Figure 11. Chamfering tool path for convex side chamfering phase Segment P0CP 0 D in Figure 11 is processed in convex side chamfering phase. Cutter surface contacts with upper and lower edges of convex side, the position relationship between cutter and pinion is constrained by d0S = rc and d0A = rc, coordinates solution of cutter center is transformed into an optimization problem as follow min f= d0S rcj j+ d0A rcj j s:t:rj \\ r \\ rj +2rc, 0\\ u \\ 2p \u00f034\u00de This optimization problem is solved by particle swarm algorithm with the same parameters mentioned above", + " In order to avoid interference, a constraint is added into tool path solving process of convex side chamfering, which is written as ffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi (r cos u x0C) 2 + (r sin u y0C) 2 + (zj z0C) 2 q . rc \u00f035\u00de Cutter can not approach the region near point P0C after the constraint above is added, a part of segment PCPD near point PC in Figure 10 is chosen to process the region near point P0C in Figure 11. When solving processes for concave side, both sides and convex side chamfering are completed, integral chamfering tool path for concave side, both sides and convex side is obtained, as shown in Figure 12. There are two discontinuities on borders of three phases, cutter moves up and then moves down in those positions to ensure processing correctly. Cutter axis direction calculating Cutter axis is set to be perpendicular to pinion axis Z1, therefore chamfering can be realized on four-axis CNC machine tools" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0000090_s38313-019-0145-6-Figure5-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0000090_s38313-019-0145-6-Figure5-1.png", + "caption": "FIGURE 5 Safety factors of the hybrid cylinder crankcase (concept) (\u00a9 Volkswagen)", + "texts": [ + " As a following step, the High Cycle Fatigue (HCF) endurance strength and the contact pressure within the metal/plastic fusion zone are calculated. The applied bearing forces, piston side forces and ignition forces correspond to the baseline engine at full load operation. All components are built up by a linear elastic material model with isotropic material properties. The concept shows a high number of areas with low safety factors. Especially the second main bearing, the inner and outer walls of the ventilation window and partly the top deck are affected, FIGURE 5. An additional calculation of the contact pressure indicates that displacements of the plastic within the fusion area are likely. Particularly the sealing zone in the lower part of the water jacket is affected. The reinforcements of the prototype significantly reduce the number of critical areas, FIGURE 6. Nevertheless there is still optimization potential regarding the ventilation window of the first and second main bearing and some spots in the top deck. The newly designed form-locking structures for the metal/plastic fusion on the main bearings also need to be improved" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0000328_14484846.2019.1699720-Figure1-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0000328_14484846.2019.1699720-Figure1-1.png", + "caption": "Figure 1. Contact between cylindrical roller and outer ring of roller element bearing.", + "texts": [ + " Finally, the conclusion of the proposed work is presented in Section 7. The thermal EHL line contact problem models the lubricant flow between two rotating ball and cylindrical rollers under an applied load. The governing physical problem is expressed by coupling of the Reynolds equation for the lubricant flow and elastic deformation equation of ball and rollers. The compressible laminar flow under steady-state lubricant flow and contacting surfaces, viz., outer ring and cylindrical roller of roller element bearing under an applied load as shown in Figure 1, is considered. The one-dimensional Reynolds equation for steady-state condition can be written as (Khan et al. 2009) @ @x \u03c1h3 \u03b7 dp dx 12um @ @x \u00f0\u03c1h\u00de \u00bc 0; where um \u00bc ua \u00fe ub 2 : (1) The relevant boundary and cavitation conditions for the proposed problem are p\u00f0xin\u00de \u00bc 0 ; and p \u00bc @p @x \u00bc 0 at x \u00bc xout: (2) The film thickness equation for line contact which incorporates roughness term in the form is given as h\u00f0x\u00de \u00bc h0 \u00fe x2 2Rx \u00fe vd\u00f0x\u00de \u00fe r\u00f0x\u00de; (3) where h0 is the central film thickness, x2 2Rx denote undeformed contact shape of geometry, vd\u00f0x\u00de denote elastic deformation of the contacting surfaces, given by vd\u00f0x\u00de \u00bc 2 \u03c0E \u00f0xout xin p\u00f0x0\u00de ln \u00f0x x0\u00de2dx0: and r\u00f0x\u00de is the surface roughness term given by r x\u00f0 \u00de \u00bc A sin 2\u03c0x The load balance equation is given by \u00f0 p\u00f0x\u00dedx \u00bc w: (4) For the line contact problems neglecting heat convection across and heat conduction along fluid film, the energy equation is given Ghosh and Hamrock by (1985) k @2t @z2 \u00bc \u03c1cpu @t @x t \u03c1 u @p @x \u03b7 @u @z 2 ; (5) where u \u00bc ua \u00fe @p @x \u00f0z 0 z \u03b7dz z 0 z \u03b7dz h 0 1 \u03b7dz \u00f0z 0 z \u03b7dz 0 @ 1 A\u00fe \u00f0ub ua\u00de h 0 1 \u03b7dz \u00f0z 0 z \u03b7dz: The surface temperature increase in the thermal EHL of line contact is given by ta\u00f0x\u00de \u00bc kfffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi \u03c0\u03c11c1k1ua p \u00f0x 1 @t @z z\u00bc0 dsffiffiffiffiffiffiffiffiffiffi x s p ; (6) tb\u00f0x\u00de \u00bc kfffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi \u03c0\u03c12c2k2ub p \u00f0x 1 @t @z z\u00bch dsffiffiffiffiffiffiffiffiffiffi x s p : (7) Equations (1\u20133) are non-dimensionalised using the following dimensionless variables: X \u00bc x b ;P \u00bc p ph ;H \u00bc hRx b ; \u03b7 \u00bc \u03b7 \u03b70 ; \u03c1 \u00bc \u03c1 \u03c10 ; H0 \u00bc h0Rx b2 ;U \u00bc \u03b70um ERx ;W \u00bc w ERx ;G \u00bc \u03b1E (8) The dimensionless Reynolds equation can be written as @ @X \u03c1H3 \u03b7 dP dX \u03bb @ @x \u00f0\u03c1H\u00de \u00bc 0; where \u03bb \u00bc 12\u03b70umRx 2 b3ph : (9) and boundary conditions become P\u00f0Xin\u00de \u00bc 0 ; and P\u00bc @P @X \u00bc 0 at X \u00bc Xc: (10) The dimensionless film thickness at any point Xi is given by (Houpert and Hamrock 1986) H\u00f0Xi\u00de \u00bc H0 \u00fe Xi 2 2 1 2\u03c0 \u00f0Xout Xin P\u00f0X0\u00de ln \u00f0X X0\u00de2dX0 1 4 ln\u00f0R2 8W \u03c0 \u00de \u00fe A sin\u00f02\u03c0Xi \u00de: (11) The dimensionless load carrying capacity is expressed as \u00f0Xout Xin PdX \u00bc \u03c0 2 : (12) The non-dimensional viscosity\u2013pressure\u2013temperature relation proposed by Roelands (1966) is (Sadeghi and Sui 1990) \u03b7 \u00bc exp ln\u00f0\u03b70\u00de \u00fe 9:67 1\u00fe 1\u00fe 5:1 10 9PHP\u00f0 \u00de z \u00fe \u03b3T0\u00f01 T\u00de (13) where \u03b70 is the absolute viscosity, PH \u00bc ph p0 is ambient pressure and zis a constant characteristic of the liquid (pressure\u2013viscosity index) as given (Cupu, Jamaluddin, and Osman 2013; Quinchia et al" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0002936_iciea.2019.8833691-Figure3-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0002936_iciea.2019.8833691-Figure3-1.png", + "caption": "Figure 3 \u2013 2-fingered gripper misalignment example.", + "texts": [ + " Although this binary criterion has been shown to be a useful measure to facilitate object gripping, it lacks any information related to how well the object has been handled \u2013 as evident by the fact that many methodologies [19, 20, 22, 27] are prone to a change in object orientation to some degree during the gripping process. The lack of descriptive power associated with this measurement \u2013 post-grip \u2013 is problematic for many industrial applications that require both controlled handling and placement of objects. In practice, a change in object orientation during the gripping process could be caused in part by a misalignment of gripper (figure 3) or gravity (figure 4). To better measure the affect the grasping process has on an object, a similarity metric is proposed in this paper to quantify the quality of candidate grasps. While the grasp only criterion is binary, a similarity score can be continuous. Such a score is calculated through a similarity test \u2013 in which an object is gripped and lifted vertically, moved away from the gripping location, moved back to its original location and placed at the gripping point. A new image is then captured and compared to its pre-gripped counterpart" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0001148_1077546320927598-Figure11-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0001148_1077546320927598-Figure11-1.png", + "caption": "Figure 11. Diagram of the control system.", + "texts": [ + " Angular velocity sensor, laser displacement sensor, and acceleration sensor are used here to measure the base vibration data. And Figures 8\u201310 give out the vibration curves tested by the three sensors. The amplitude of angular velocity of the base pitching vibration is 0:92\u00f0rad=s\u00de, the amplitude of displacement of the base vertical vibration is 32:6mm, and the amplitude of acceleration of the base vertical vibration is 4\u00f0m=s2\u00de. The robotic manipulator is driven by a DCmotor. Its type is Maxon RE50. Figure 11 shows the composition of the control system. In the control system, EPOS2 digital position controller is used to transmit control command between the personal computer and robotic manipulator system via USB. This controller is matched with the Maxon motor. Figure 12 shows the block diagram of designed compensated control in this article. The compensation of system friction and balance torque is based on the LuGre friction model and balance torque model identified in Section 2.2 and Section 2.3, respectively" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0000463_s12541-019-00282-y-Figure10-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0000463_s12541-019-00282-y-Figure10-1.png", + "caption": "Fig. 10 Generation of forward and backward motions Fig. 11 Generation of rightward and leftward motions", + "texts": [ + " In order to perform the forward and backward movements, the friction between the floor and the pad should be based on the y-axis, with x\u0308 = 0 and \u03a6\u0308 = 0. (15) Mtx\u0308 = \ud835\udefc[\u2212s1\ud835\udf031sin ( \ud835\udf131 ) + s2\ud835\udf032sin ( \ud835\udf132 ) + s3\ud835\udf033sin ( \ud835\udf133 ) \u2212 s4\ud835\udf034sin ( \ud835\udf134 ) ], (16)Mty\u0308 = \ud835\udefc[+s1\ud835\udefc\ud835\udf031cos ( \ud835\udf131 ) + s2\ud835\udefc\ud835\udf032cos ( \ud835\udf132 ) \u2212 s3\ud835\udefc\ud835\udf033cos ( \ud835\udf133 ) \u2212 s4\ud835\udefc\ud835\udf034cos ( \ud835\udf134 ) ], (17)IG\u03a6\u0308 = \u2212s1(T1 + d\ud835\udefc\ud835\udf031) \u2212 s2 ( T2 + d\ud835\udefc\ud835\udf032 ) \u2212 s3 ( T3 + d\ud835\udefc\ud835\udf033 ) \u2212 s4 ( T4 + d\ud835\udefc\ud835\udf034 ) . With the condition of x\u0308 = 0 and \u03a6\u0308 = 0 through Eqs.\u00a0(15)\u2013(17) and constraint 1, the force is generated in the y direction. The forward motion is generated when the pad velocity is given as follows: Figure\u00a010 shows the force generated when setting the direction of the pad and the moving direction of the robot. From Eqs.\u00a0(12)\u2013(14), the equations of the back and forth motions can be expressed as As shown in Eqs.\u00a0(17) and (18), the x-axis force and the moment of the robot cancel each other and become zero, and the force and acceleration are generated in the y direction. Thus, the robot can move forward. For the backward movement, the direction of the pad rotation is the reverse of that for forward rotation" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0001949_j.mechmachtheory.2020.104139-Figure1-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0001949_j.mechmachtheory.2020.104139-Figure1-1.png", + "caption": "Fig. 1. Features of ball CVTs: (a) Conformal ball CVT and (b) Non-conformal ball CVT; and conformity between disk cavity and ball.", + "texts": [ + " The main purpose of this study is to identify performance change for all ball CVT design factors based on the newly established ball CVT analysis model. The performance of ball CVTs was analyzed in terms of torque capability, efficiency, and torque density. In order to grasp trends of performance changes from the analysis results and increase utilization of the analysis results themselves, most design and performance factors are expressed as dimensionless values. It is an additional object of this study to compare performance between two ball CVTs of conformal type and non-conformal type. Fig. 1 shows the geometric characteristics of two ball CVTs, conformal and non-conformal, and the conformity between the disk cavity and the ball. In the conformal ball CVT shown in Fig. 1 (a), when clamping force is applied between the two disks for power transmission, the direction of the force received by the balls from the disks is the inner direction, which is the direction of the center of rotation of the disk. To support the force in this direction, the conformal ball CVT requires an internal carrier between the balls. On the other hand, in the non-conformal ball CVT shown in Fig. 1 (b), when the same force is applied, the balls are subject to a force directed outward from the center of rotation of the disk, and so the balls are supported by an external carrier installed on the outside the balls. In this study, the conformal ball CVT is represented by subscript C, and the non-conformal ball CVT is represented by subscript NC. Geometry of ball CVT is determined by the following four design factors: \u2022 Contact angle, \u03b1 \u2022 Conformity, F \u2022 Ratio occupied by balls within total diameter of ball CVT, B \u2022 Overall radius of ball CVT, r CVT The contact angle \u03b1 is an angle formed by a line connecting the rotation axis of the ball, the center of the ball, and the contact point between the disk and the ball when the rotation axis of the disk and that of the ball are parallel" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0003632_0954406214549786-Figure2-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0003632_0954406214549786-Figure2-1.png", + "caption": "Figure 2. Equivalent dynamic model of the tilting table driven by worm and worm wheel in the tilting direction.", + "texts": [ + " The position of the center of mass G of the titling table will change when the tilting base rotates an angle s around x-axis (right hand rule). Equivalent dynamic model in the tilting direction The dynamic characteristics of the tilting table driven by worm and worm wheel are mainly decided by the stiffness and damping of the worm wheel shaft, worm gear pair, supporting bearings, worm shaft, and the coupling. Its equivalent dynamic model in the tilting direction can be established as shown in Figure 2. The contact between the worm and worm wheel is represented by a spring-damper element in the tangential direction of the pitch circle of the worm wheel. sx is the angle displacement of the tilting table in the tilting direction. T s\u00f0 \u00de, the total gravity torque produced by tilting parts including the worm wheel shaft, the tilting base, and the rotary table, is applied on the transmission system at the tilting angle s. Jeq is the total moment of inertia of the tilting parts about x-axis. Kwgs and Cwgs are the torsional stiffness and damping of the worm wheel shaft" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0002118_14763141.2020.1834609-Figure5-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0002118_14763141.2020.1834609-Figure5-1.png", + "caption": "Figure 5. Anterior view stick figure diagram of the bat and upper body at ball impact (solid line: optimised performance, dashed line: measured performance).", + "texts": [], + "surrounding_texts": [ + "The RMSE between the measured data and simulation model was 0.19 m/s, 0.98\u00b0, 0.31\u00b0 for the bat-head speed, bat orientation angle and the joint angles of the upper body (17DOFs) respectively. The second and third optimisations, that distributed the force and moment of the individual hands, achieved RMSEs in the bat parameters of 0.51 m/s and 2.61\u00b0, and 0.46 m/s and 1.45\u00b0, respectively. The maximum bat-head speed achieved in the optimised performance simulation was 40.0 m/s, predominantly produced by an increase in the Y-axis component of its velocity around ball impact (Figure 4). The angle between the bat-head velocity vectors of the measured and optimised performance data was 1.81\u00b0, and the difference of the impact point of the bat at ball impact was 0.032 m, which are set to constrain the ball impact. The time histories of joint angles at the barrel- and knob-side upper limbs in the measured and optimised performance simulation data displayed differences in the second half, in particular around ball impact (Table 1, Figures 5, 6 and 7). The barrel-side shoulder abduction angle was smaller in the simulated than the measured performance (Figure 6 (b)) and wrist radial flexion angle was smaller in the simulated than the measured performance (Figure 6(g)). The knob-side elbow flexion/extension and wrist palmar flexion angles were smaller in the simulated than the measured performance (Figure 7(d,f)). In addition, the right lateral flexion angle of the torso joint was larger in the simulated than the measured performance during in the 0.15 s before ball impact (Figure 8(b)). Figure 4. Time histories of (a) bat-head speed and (b) velocity from the start of the swing to ball impact (solid line: optimised performance, dashed line: measured performance). Unit: degrees FE: flexion/extension; AA: adduction/abduction; IER: internal/external rotation; PS: pronation/supination; PDF: palmar/ dorsi flexion; RU: radial/ulnar flexion; ARF: retro/ante flexion; RL: right/left lateral flexion; CCR: counter-clockwise /clockwise rotation. Figure 7. Time histories of the joint angle of the knob-side upper limb, shoulder joint [a-c], elbow joint [d, e], and wrist joint [f, g], from the start of the swing to ball impact (solid line: optimised performance, dashed line: measured performance). Figure 8. Time histories of the torso joint angles from the start of the swing to ball impact (solid line: optimised performance, dashed line: measured performance). The RMSE of the joint torque time histories of the upper limbs between second and third optimisations was 5.8 \u00b1 5.4 Nm (Table 2). The largest discrepancies were observed in the knob-side shoulder adduction/abduction and elbow flexion/extension torques; 22.2 Nm and 11.2 Nm, respectively (Figure 9)." + ] + }, + { + "image_filename": "designv11_71_0003024_s1063454119040071-Figure1-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0003024_s1063454119040071-Figure1-1.png", + "caption": "Fig. 1. Illustration to Note 1.", + "texts": [ + " d d d dZ t z x t y x t x x t y t x t x dt d dt dt f x t y t t f x t x v = = \u2212v v v 0 0( ) ( ) ( ), (0) ,d t A t t y x dt VESTNIK ST. PETERSBURG UNIVERSITY, MATHEMATICS Vol. 52 No. 4 2019 where the matrix function A(\u00b7) in (17) may be calculated by the formula (18) and is independent of the homeomorphism \u03c4(t) used in determination of the residual z(\u00b7). To prove Statement 2 it is convenient to exploit the following fact. Note 1. Consider the two vectors \u2208 and \u2208 emerging from the same point O and the plane orthogonal to . If we decompose the vector = + , where is parallel to and is orthogonal to (see Fig. 1), then the vector is determined by (19) Indeed, if we denote the scalar product of the vectors and by \u03bb, that is, \u03bb := and, with no loss in generality, assume that | | = 1, then = \u03bb \u00b7 and We return to the proof of Statement 2. The linearization of the dynamics of system (1) in the neighborhood of the solution x0(t) = x(t, x0) permits approximately writing the perturbed solution x(t) = x(t, y0) as where the vector function w(\u00b7) is the solution to the linear system (20) We now take a look at Fig. 2 and use Note 1 to compute the approximation of the residual z(\u00b7)" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0003282_ccdc.2015.7161907-Figure1-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0003282_ccdc.2015.7161907-Figure1-1.png", + "caption": "Fig. 1 TORA system", + "texts": [ + " In this paper, the sliding mode control method is adopted to design the stability controller of TORA. Inspired by the vibration suppression of civil structures, we take the effect of earthquake wave on TORA system into consideration, and study the robustness of sliding mode controller when environmental disturbance exists. Simulation experiments are conducted to prove the validity and robustness of control algorithm proposed. 2 DYNAMICS OF TORA 2.1 Moedling of TORA without Disturbance As depicted in Fig. 1, TORA system consists of a rotating mass driven by a motor and a translational cart fixed by a spring with no damping. Cart mass is denoted by M. The rotating mass is m, and the stiffness coefficient of spring is k. The rotating mass rotates in the horizontal plane driven by the motor torque N, and the rotational radius is r. The angle is denoted by . And the Moment of inertia of rotating mass is J. The cart has one-dimensional movement in the horizontal plane, and its displacement is x. The modeling of TORA system can be done using the Lagrange equations", + " 2 2 2 2 1 2 1 1 1[ cos ( ) ] 2 2 2 \u03b8 \u03b8 \u03b8= + = + + + +K K K Mx mx mrx mr J (2) The total potential energy is 21 2 =P kx (3) So the Lagrangian of TORA system is 2 2 2 21 1 1( ) cos ( ) 2 2 2 \u03b8 \u03b8 \u03b8= \u2212 = + + + + \u2212L K P M m x mrx mr J kx (4) According to the Lagrange equations, the dynamics of TORA should be d ( ) 0 d d ( ) d \u03b8\u03b8 \u2202 \u2202\u2212 = \u2202 \u2202 \u2202 \u2202\u2212 = \u2202\u2202 L L t x x L L N t (5) By substituting the Lagrangian (4) into (5), the dynamics of TORA can be obtained ( ) cos sin 0\u03b8\u03b8 \u03b8+ + \u2212 + =M m x mr mr kx (6) 2cos ( )\u03b8 \u03b8+ + =mr x mr J N (7) We can rewrite the dynamics into the vectorial form ( ) ( , ) ( )+ + =M q q C q q G q U (8) where 2 cos ( ) cos \u03b8 \u03b8 + = + M m mr M q mr mr J , 0 sin( , ) 0 0 \u03b8 \u03b8\u2212 = mrC q q 0 , ( ) , 0 x kx q G q U N\u03b8 = = = Through observing the TORA system in Fig 1, and substituting 0, 0, 0, 0, 0\u03b8 \u03b8= = = = =x x N into the dynamics of TORA, the stable equilibrium points of TORA can be figured out 0; [0,2 ]\u03b8 \u03c0= \u2208x It means the spring will be its original length and the angle of rotating mass can be any value of interval [0, 2 ] when TORA system reaches the stable equilibrium point. In this paper, we try to design the torque input N to stabilize TORA system to the point ( , , , ) (0,0,0,0)\u03b8 \u03b8 =x x To stabilize two variables with only one control input shows the underactuated characteristics of TORA" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0002989_978-3-642-16214-5_9-Figure9.48-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0002989_978-3-642-16214-5_9-Figure9.48-1.png", + "caption": "Fig. 9.48. Assembly of the parking lock in the transmission ZF 6 HP 26", + "texts": [ + " The system has to be robust enough to function safely even in poor tolerance situations, under extreme temperatures and with the deterioration that sets in after extensive mileage. The surface quality of the parking lock wheels and pawls, frequently designed as a blanking part, must fulfil all requirements with respect to surface pressure and friction. F\u00f6rster provides suggestions for determining the essential geometric conditions of parking lock systems in [9.7]. The abovementioned criteria for parking lock systems of robustness and functional reliability must also be guaranteed in the case of electric activation. Figure 9.48 shows the assembly situation and Figure 9.49 the working principle of an electrically activated parking lock using the example of the automatic transmission ZF 6 HP 26. The detent plate in the transmission is omitted, being substituted with a parking plate as well as with a parking lock cylinder with a locking magnet. When leaving the park position, hydraulic pressure is introduced to the parking lock cylinder by means of a solenoid valve. The cylinder pushes back a piston, which extracts the locking cone under the pawl by means of the kinematics shown" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0003342_978-3-319-07398-9_9-Figure17-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0003342_978-3-319-07398-9_9-Figure17-1.png", + "caption": "Fig. 17 Proposed concept: hexapod", + "texts": [ + " The first concept, based on ball joints (Drutchas 1991), is similar to the bone structure of an arm with shoulder and elbow joints replaced with ball and pin joints, respectively. The serial mechanism achieves the required range of motion through a combination of rotations allowed by the joints. Figure 16 shows the example image of this concept. The second concept, a hexapod (Newport http://www.newport.com), supports the headrest by six links working in parallel, allowing triangulation of three points in space while deterministically defining a plane. Figure 17 shows the example image of this concept. The third concept, based on media jamming (Brown et al. 2010), uses granular media subject to a critical compressive stress to form a pseudo-solid. It is possible to rapidly switch a cleverly designed support from free-forming and freely adjustable to rigid, pseudo-solid by just applying an appropriate vacuum. Figure 18 shows the example image of this concept applying to robotic arm gripper. SmoothMotion team built three concept prototypes based on these initial concepts, and posted them to the Wiki for the UT team to review prior to an online design reviews" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0002586_edpc51184.2020.9388196-Figure10-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0002586_edpc51184.2020.9388196-Figure10-1.png", + "caption": "Fig. 10. Optimized cross-section of distribution channel (left) according to [22], elliptical cross-section at the tooth (center), rounded contour of single channel (right).", + "texts": [ + " The medium backflow (green) spreads over four single pipes, before these open into a ring-shaped cross-section. For the design of all channels, recommendations can be applied that are valid for contour cooling of LBMprinted injection molding or sheet metal forming tools [21], [22], [23]. The aim is to achieve a balance between sufficient cooling capacity, the strength of the structure, and the pressure losses in the channels. Since the course of the distribution channels is perpendicular to the additive build-up direction, the cross-section is designed drop-shaped instead of round (Fig. 10 left). In this way, a manufacturing-related loss of cross-section can be provided. Due to the compact and small shape of the teeth, elliptical shape is defined for the channel cross-section (Fig. 10 middle). In this way, the channels run just below the tooth tip without weakening its structure too much. At the same time, a smaller distance to the contour can be achieved at a larger channel surface. As the channels run at a small angle to the direction of construction, this cross-sectional shape can be realized without loss of surface area. The curve of a single channel can be seen in Fig. 10 on the right. Due to the helix angle of the gear, the shape of the cooling channels is also helical. Besides, large fillets are provided at the directional changes. This allows on the one hand a reduction of pressure Authorized licensed use limited to: San Francisco State Univ. Downloaded on June 20,2021 at 01:51:06 UTC from IEEE Xplore. Restrictions apply. losses and on the other hand a better de-powdering and flushing after additive manufacturing. It should be noted that the individual cooling channels do not run across the entire face width" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0002816_thc-2010-0566-Figure17-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0002816_thc-2010-0566-Figure17-1.png", + "caption": "Fig. 17. Mohr\u2019s circle.", + "texts": [ + " The tank is pressurized and if we look at a small square piece of material on the tank surface we can see that the square experiences tensile stresses on both sides as shown in Fig. 16. To handle these situations with combined loading we use Mohr\u2019s circle. Most of the materials have to be analysed 3-dimensionally but since they are geometric in nature, using the symmetry we can convert them to 2 dimensions like we did for the spherical storage tank (Fig. 16) Mohr\u2019s circle is a plot between the normal stress (\u03c3) on the horizontal or x-axis and the shear stress (\u03c4 ) on the vertical or y-axis (Fig. 17). Tensile normal stresses are positive and compressive are negative, for the shear stress we follow a sign convention shown in Fig. 18 where a very small volume element inside the bulk of the material is shown. Coming back to Fig. 17, the position of the center of the circle c is the average of the two normal stresses (\u03c3x and \u03c3y and reference point a can be plotted with \u03c4 xy on the y-axis and \u03c3x on the x-axis. Now we can join points c and a to get the radius and then the circle can be drawn with a radius given by Eq. (12). From this Mohr\u2019s circle for a given situation we can find out the maximum normal stress, point b, and shear stress, point e, active in the material. R = \u221a( \u03c3x \u2212 \u03c3y 2 )2 + \u03c42 xy (13) If we take a simple situation where a tensile stress of 100 Pa is active on a square cross section beam then the state of stress at a very small volume element will be as shown in Fig" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0003554_gtindia2014-8203-Figure3-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0003554_gtindia2014-8203-Figure3-1.png", + "caption": "Figure 3. The scheme of method for calculation compressor rotor blades forced oscillations.", + "texts": [ + " Based on this calculation, RW frequency diagram construction and IPC rotor speeds, at which the resonance may occur, are carried out. It should be noted that the coupled oscillations of the blade and disk were not considered in this paper. 4. Gas load is represented as a combination of backward traveling waves of load (harmonic waves) using the objectoriented programming language APDL, built in Ansys. Further dynamic calculation at the resonance mode in resonance with the most dangerous harmonic is performed. This method is presented schematically in Figure 3. The following areas were selected as measures to reduce the resonant stresses level: the usage of the blades with special Shvarov\u2019s profile at last rotor wheel; the usage of the guide vane in front of the middle support, with the blades set having different stagger angles and with circumferentially alternating blade pitch. Also, special attention was paid to the fact that the number of blades with stagger angles different from the standard should be as low as possible. Parameterization of the full circle compressor model was carried out for the introduction GV5 different stagger angles and circumferentially alternating blade pitch" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0001701_0954405420949757-Figure7-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0001701_0954405420949757-Figure7-1.png", + "caption": "Figure 7. Upper and lower edges.", + "texts": [ + " Since E 0 i is a point on auxiliary circle R0i, if E 0 i is also on face cone, E0i and A0i coincide, the following geometric relationship exists xi zi zf = ffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi x0Ei 2 + y0Ei 2 q z0Ei zf \u00f021\u00de The problem of solving (x0Ai, y 0 Ai, z 0 Ai) can be transformed into a single variable optimization problem as follow min f= j xi zi zf ffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi x0Ei 2 + y0Ei 2 p z0Ei zf j s:t:p=2\\ f0i \\ 3p=2 ( \u00f022\u00de This problem is solved by particle swarm algorithm, algorithm parameters are as same as above. The coordinates of point A00i are denoted by (x00Ai, y 00 Ai, z 00 Ai), whose solving method is as same as point A0i. Radius adjustment of auxiliary circle When subscript value i varies from 1 to m, points A0i, S 0 i, A00i and S00i are calculated. Upper edge of convex side is created by linking points A0i with different value i. By the same method, lower edge of convex side, upper edge of concave side and lower edge of concave side are created, which are shown in Figure 7. Let the distance between A0i and S0i be d 0 i, the distance between A00i and S00i be d 00 i . d 0 i and d00i increase as r 0 i and r00i increase. Design value of chamfering width is set to dc. In order to control chamfering width of convex side, d0i = dc should be satisfied, since there is no analytic solution to this equation, the problem of controlling chamfering width of convex side can be transformed into a single variable optimization problem as follow min f= d0i dcj j s:t:0\\ r0i \\ 2dc \u00f023\u00de Initial value of r0i and increment per time are set to 0" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0003321_oceans-genova.2015.7271659-Figure4-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0003321_oceans-genova.2015.7271659-Figure4-1.png", + "caption": "Fig. 4. Glider mission plan: a) way points and lane lines in the horizontal plane, and b) yo-yo trajectory in the vertical plane.", + "texts": [ + " According to [2], for linear models, the mth component of the control input vector (with m=1, 2, 3 in the 3D case) is given by an expression involving the state covariance matrix, the measurement matrix and its gradient with respect to the agent position as follows: 1 , , \u02c6 \u02c62 ( ) ( ) ( ) ( ) T T i m i i i i mA f SR dA p where ( ) ( )i i is the agent measurement matrix according to (3) and 2 i iR . The time step index has been dropped for the sake of clarity. The dynamic of the agents in (8) is adapted, as specified below, to model the typical behavior of an underwater glider [17], having a constant speed (in absence of sea current), a constrained vertical plane dynamic and a waypoint guidance system. The wave glider is similarly modeled in 2D. Generally, an underwater glider moves through a 3D space following a saw-tooth shape trajectory (see Fig. 4) in the vertical plane. The trajectory is composed of a certain number of dive/climb cycles in the interval between two surfacing phases of the glider (during which data transmission is possible). The data, collected during each dive or climb cycle, are stored and finally transmitted during the surfacing phase. The underwater glider dynamic model considered in this work assumes a constant velocity without water current disturbances, constrained to follow a yo-yo trajectory in the vertical plane with given climbing and diving target depths [18]" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0000254_ecce.2019.8911883-Figure6-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0000254_ecce.2019.8911883-Figure6-1.png", + "caption": "Fig. 6. Definition of kf", + "texts": [ + " In terms of [22], the ripple torque relates to the least common multiple (LCM) of the stator slot number and rotor tooth number. The larger LCM of Zs and Zr, the lower ripple torque. These values in 16-, 17- and 20-rotor-tooth proposed FSPM machines are 48, 204 and 60, respectively. Thus, the 12/17 proposed FSPM machine has smaller ripple torque than that of the 12/16 and 12/20 proposed FSPM machines. To take a comprehensive consideration, the optimal split ratios of 16-, 17- and 20-rotor-tooth proposed FSPM machines are 0.56, 0.54 and 0.58. B. Flux bridge width factor As shown in Fig. 6, the flux bridge width factor kf is defined as : IV. COMPARISON OF THE PROPOSED AND CONVENTIONAL FSPM MACHINES In this section, the electromagnetic performances, including flux density, back-EMF, torque, overload capability etc. of the proposed FSPM machine and the conventional FSPM machine are compared by using FEA method. Two FSPM machines have been optimized and some main parameters are listed in Table III. When at a high pole ratio, the function of flux bridges is obvious, hence the 12- stator-slot/ 17-rotor-tooth FSPM machines, whose pole ratio is 17, are chosen as an example" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0002276_s12541-020-00431-8-Figure7-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0002276_s12541-020-00431-8-Figure7-1.png", + "caption": "Fig. 7 Roll test rig model", + "texts": [ + " Modeling is performed using Dymola (Dassault Syst\u00e8mes, France), which is a development environment based on the Modelica language. To verify the accuracy and perform application cases for the beam element based CTBA model proposed in this study, two test modes were performed: roll test and VTL (virtual testing lab). This section describes the test rig models for performing these tests. In roll simulation, vertical displacement is applied in the opposite direction to the left and right wheels to create roll torque. As shown in Fig.\u00a07, the test rig is created by connecting actuators that can apply the vertical displacement to the wheel center of the CTBA model. The input profile used in this study is shown in Fig.\u00a08. A virtual testing lab (VTL) is an analysis method that uses a drive signal created based on the signal measured in vehicle durability tests. In this study, the wheel center force and moment signal measured by a wheel force transducer (WFT) during accelerated durability road (Belgian) tests are used as 1 3 the VTL drive signals" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0002917_b136378_7-Figure7.7-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0002917_b136378_7-Figure7.7-1.png", + "caption": "Fig. 7.7 Catheter Clark oxygen electrode", + "texts": [ + ", 1986) that the fastest speed of response for a hemicylindrical Clark electrode is obtained with an electrode radius of approximately 5\u201310\u03bcm. This is due to the fact that as the radius decreases, the effect of the layers that are closer to the electrode surface becomes relatively more important than those that are farther away. In fact, with the further decrease of the radius the time response becomes longer than for the corresponding planar electrodes. The greatest impact of the Clark oxygen electrodes has been in medicine and physiology. (A schematic diagram of a catheter-size Clark electrode is shown in Fig. 7.7.) On the other hand, a temperature- and pressure-compensated Clark electrode for oceanographic measurements up to 600 ft has also been developed (Fatt, 1976). The normal temperature coefficient of the Clark electrode is \u223c2%/\u25e6C and the linearity is usually better than 1%. The time response depends mainly on the thickness of the membrane. For a 5\u03bcm thick polypropylene membrane on a 7\u03bcm radius hemispherical electrode, the response time is below 1 s (Hurrell and Abruna, 1988). The idea of confining the electrical circuit behind a gas-permeable membrane is quite general" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0000294_icems.2019.8921576-Figure1-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0000294_icems.2019.8921576-Figure1-1.png", + "caption": "Fig. 1. Modular fault-tolerant motor system.", + "texts": [ + " The relationship between overload multiple of fault-tolerant current and safe operation time is also discussed. The faulttolerant strategy with temperature analysis can give a more reasonable fault-tolerant current and make the output capacity of the motor maximized on the premise of the safe temperature rise. II. MODULAR FAULT-TOLERANT MOTOR SYSTEM AND The motor is adopted 16/24 poles/slots, double-layers fractional slot concentrated winding (FSCW) and four modules structure. Each module is controlled by an independent drive with three-phase four-wire control circuit, as shown in Fig. 1. When the fault occurs, the currents of the remaining healthy windings are adjusted by topology reconstruction to compensate the MMF loss caused by the faulty windings or the faulty bridge arms. The generated zero-sequence current flows through the redundant bridge arm to ensure no effect on the normal operation of other healthy modules. The program combines the characteristics This work was supported by the National Natural Science Foundation of China under Grant 51677039. 978-1-7281-3398-0/19/$31" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0003550_asjc.1046-Figure4-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0003550_asjc.1046-Figure4-1.png", + "caption": "Fig. 4. The effect of the network is sampling, delay and quantization.", + "texts": [ + " Let the sampling interval be tk+1 \u2212 tk = \ud835\udf0fk. If \ud835\udf0fk is constant, the system is periodic[21]. Because of the random network scheduling and the stochastic perturbations in environment, \ud835\udf0fk is time varying and not necessarily less than the sampling period of the sensor. Under the event-driven assumption, the controller keeps using the information until the next one arrives (see Figure3). So, the control input is u(t) = Kq(x(tk\u22121)), t \u2208 [tk, tk+1). (5) Considering the effects of the network, system in Figure 1 is transformed into the one in Figure 4. The resultant feedback system is described by x\u0307(t) = Ax(t) + BKq(x(tk\u22121)), t \u2208 [tk, tk+1). (6) The state of the system is evolving as follows: x(tk+1) = eA\ud835\udf0fk x(tk) + \u222b \ud835\udf0fk 0 eAhdhBKq(x(tk\u22121)). (7) If all the transmission intervals are the same to h0, the coefficient matrices in (6) are constant and the system is periodic, which has been well studied in [24]. Actually, even a well-designed network may fail to perform an equal time interval transmission. In order to discuss the robustness against the variation of the network induced delay, we define \ud835\udf0fk = \ud835\udf0f0 + \ud835\udf03k, where \ud835\udf03k is the perturbation of the delay and belongs to [hl \u2212 \ud835\udf0f0, hu \u2212 \ud835\udf0f0]" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0001219_icemi46757.2019.9101620-Figure1-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0001219_icemi46757.2019.9101620-Figure1-1.png", + "caption": "Fig. 1 The diagram of the joint link", + "texts": [ + " The D-H parameter method is used to establish the kinematic equation of manipulator in this paper. For a manipulator consisting of arbitrary multiple links and joints, in order to establish kinematic equations, it is necessary to establish a three-dimensional coordinate system for each joint that satisfies the right-hand rule. The position and attitude relations between adjacent coordinate systems are obtained by D-H parameters, and then the position and attitude of the end in Cartesian space are described by the relationship between coordinate systems. Fig.1 shows two adjacent joints and one link. By translating and rotating, one coordinate system can be transformed to the next. According to the transformation order of i i i id a , the corresponding transformation matrix is expressed as (1). 1 ( ) Rot ( ) ( ) Rot ( cos sin cos sin sin cos sin cos cos cos sin cos = 0 sin cos 0 0 0 1 ) i i i i i i i i i i i i i i i z i z i x i x i i i ia a a T Trans d Tra d ns (1) Where, i-1 i T represents the transformation matrix between adjacent coordinate systems" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0001817_0959651820959164-Figure1-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0001817_0959651820959164-Figure1-1.png", + "caption": "Figure 1. Schematic diagram of the cable-driven manipulator.", + "texts": [ + " Section \u2018\u2018Implementation of algorithm\u2019\u2019 explains details about the implementation of the proposed algorithm and considers the desired trajectory as inputs of the FL system to improve the control performance. A detailed stability analysis is given to ensure the stability of the proposed algorithm. Section \u2018\u2018Experimental studies\u2019\u2019 gives comparative experiments to show the validity and effectiveness of the proposed algorithm. Section \u2018\u2018Conclusion\u2019\u2019 concludes the passage. A schematic diagram of the CDM is shown in Figure 1 where we can easily see kinematic coupling do exist. Suppose q represents rotation or translation positions of motors, and u represents rotation or translation positions of joints. The kinematic coupling of a CDM is given by Xu et al.36 q=C u \u00f01\u00de where C is a positive number for one joint, or a lower triangular matrix if the manipulator has more than 1 DOF. That lower triangular matrix consists of zero or positive numbers and all its diagonal elements are positive. Then considering the dynamic character of motors,37 we have tm = I\u20acq+Dm _q+ ts \u00f02\u00de where tm is electromagnetic torque; ts is output torque to drive the cables" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0003618_ijvsmt.2015.067521-Figure6-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0003618_ijvsmt.2015.067521-Figure6-1.png", + "caption": "Figure 6 Roll angle ( ),\u03c6 camber angle (\u03b3) and angle (\u0393) (see online version for colours)", + "texts": [ + " ( )2 1 2 1 \u02c6 [0,cos\u0393,sin\u0393]TA A p A A \u2212 = = \u2212 (13) In addition, the coupling between the location of the end-effector and the displacements of the actuators 2 and 3 is obtained by the equation (14): ( ) ( ) 2 ( 1, 2,3)T i i i iA B A B i\u2212 \u2212 = = (14) Developing the equation (14) for the bars 2 and 3, we obtain the equations (15) and (16), respectively: 2 2 2 2 22( cos\u0398 cos \u0393) 2 cos(\u0398 \u0393) sin\u0398 sin \u0393 0s h s h h h h\u23a1 \u23a4\u2212 + + + \u2212 \u2212 \u2212 =\u23a3 \u23a6 (15) 2 2 2 3 3 2 2 2 2 2( cos\u0398 cos \u0393 sin \u0393 sin\u03a8) (1 cos\u03a8) ( sin\u0398 sin \u0393 cos \u0393 sin\u03a8 ) ( cos\u0398 cos \u0393) ( sin \u0393 sin\u03a8) 2 sin \u0393sen\u03a8( cos\u0398 cos \u0393) 0 s h d s d h d h l h d d h \u2212 + + + \u2212 + + \u2212 \u2212 \u2212 + + + + = (16) The equations (5), (15) and (16) are second-degree polynomial, independent and decoupled in the variables s1, s2 and s3. These equations represent the mathematical transformation between the end-effector location and the displacements provided by the actuators. The angles \u03a8 and \u0393, which define the location of the end-effector in the moving frame can be related to the angles \u03b3 and \u03b5, which define the camber and toe of the wheel in the fixed frame, respectively. The relationship between the angles \u03b3 and \u0393 is given by the equation (17) (see Figure 6) while the relationship between the angles \u03b5 and \u03a8 is given by the equation 18 (see Figure 5). The angles in the both equations are defined applying the right-hand rule. \u0393 2 \u03b3 = \u2212 + \u03c0 \u03c6 (17) ( )arctan y x\u03b5 w w= (18) ( ) ( )0 0 1 0 1 0 \u02c6 \u02c6\u02c6 \u02c6 \u02c6x y uw Ru i w Ru j u d \u2032 \u2032 \u2032 \u2032= \u22c5 = \u22c5 = (19) where 0 1 R is the rotation matrix of the moving frame with respect to the fixed frame. Figure 7(a) shows the free-body diagram of the wheel-tyre assembled to wheel carrier. In order to simplify the force actuators analysis, the following hypotheses are assumed: \u2022 only the mass of the wheel-tyre set is taken into account, hence, the other parts masses are neglected \u2022 the product of inertia of the wheel-tyre set is neglected \u2022 the longitudinal acceleration is not considered \u2022 the wheel vertical motion is only a mathematical function of the body roll angle (it is assumed there are no irregularities on the flat road) \u2022 the vertical and angular accelerations of the wheel-tyre set are neglected \u2022 the rolling resistance at the tyre is neglected" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0001588_aim43001.2020.9158991-Figure3-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0001588_aim43001.2020.9158991-Figure3-1.png", + "caption": "Figure 3. TakoBot wire alignment structure", + "texts": [ + " The wires are aligned along with the wire eyelets on each disc (Fig. 2). Such a structure so-called sliding disc mechanism provides smart bending stress distribution along the slender part and improves manipulator bending features by decentralizing forces and bending torques [24]. The maximum traveling displacement of sliding discs is 10mm. The spacer disc diameter is 50mm, and the total length of the continuum part is 380 mm. The wire passes along the spacer discs via wire eyelets and passes through inside of the compressional springs (Fig. 3). In this prototype, we used a stainless-steel wire with a diameter of 1mm, and the eyelet hole size is about 1.5mm. Based on previous experiences, the hole size was determined to be close to the wire diameter. Otherwise, it might cause twist deformation in the continuum part. B. Pretension mechanism design The pretension mechanism (PtM) located in the middle part of the robot. Driving wires passes through the PtM device. In this robot, we drive eight wires, four wires for each section. A linear actuating unit drives wires; in this robot, one motor drives two paired wires with a strain and ease manner, Total four motors actuate eight wires (Fig 4)" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0002274_diped49797.2020.9273405-Figure1-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0002274_diped49797.2020.9273405-Figure1-1.png", + "caption": "Figure 1. Location of system elements", + "texts": [ + " Compared with the method of integral equations, the use of finite difference methods for modeling coupled electromagnetic and mechanical processes cannot be defined as rational, since they have such disadvantages as artificial limitation of the calculated range, sampling of the surrounding space, etc.). Therefore, when analyzing electromagnetic processes in linear systems, the method of integral equations for magnetic field sources is most suitable [4]. In this work, the calculation is based on solving a boundary value problem for scalar electric and vector Authorized licensed use limited to: San Francisco State Univ. Downloaded on June 20,2021 at 21:10:33 UTC from IEEE Xplore. Restrictions apply. 60 magnetic potentials, which is written based on Maxwell's equations. In Fig. 1, the arrangement of the elements is shown; the system consists of two ring inductors and a cylindrical ferromagnetic shell. We consider the known geometrical parameters of the system, as well as the electrical and magnetic properties of the materials: \u03b3 is the electrical conductivity of the shell material we take equal to 0; \u03bc is the absolute magnetic permeability of the shell material; J is the permanent magnet magnetization. The electromagnetic field is calculated using the Maxwell equation system" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0001489_ur49135.2020.9144980-Figure1-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0001489_ur49135.2020.9144980-Figure1-1.png", + "caption": "Fig. 1: Reference frames of a quadrotor in + configuration with 4 rotors and propellers.", + "texts": [ + " Let W : OW,xW,yW, zW denote a right-hand inertial reference frame with origin OW stationary with respect to the earth, with xW \u00b7 yW = 0, xW \u00d7 yW = zW and such that zW points out of the earth. Let p denote the position of the center of mass (CoM) of the vehicle relative to W , with components x, y, z such that p = xxW +yyW +zzW. Let B : OB,xB,yB, zB denote a right-hand frame whose origin OB is attached to the CoM of the vehicle, xB \u00b7yB = 0, xB\u00d7yB = zB and zB pointing in the thrust direction. Frame B\u2019s axes are aligned such that the inertia tensor J is a positive-definite diagonal matrix. Figure 1 shows the configuration of a typical quadrotor, reference frames and relevant quantities. The equations of motion for a multicopter with n rotors can be written as: p\u0307 = v (1) mv\u0307 = \u2212mgzW + TzB + Rfa (2) R\u0307 = R\u03c9\u0302 (3) J\u03c9\u0307 = \u2212\u03c9 \u00d7 J\u03c9 + q (4) where T = \u2211n i=1 Ti, Ti is the thrust generated by rotor i, q = \u2211n i=1 di \u00d7 TizB + Qi is the torque vector produced by the action of thrust TizB at a distance di from the CoM, Qi is moment produced by the rotation of propeller i. The state vector is X = [p,v,R,\u03c9]T where v is the velocity, R = RW B \u2208 SO(3) is the vehicle-to-world transformation that represents attitude, and \u03c9 the angular velocity in body frame and whose elements \u03c9x, \u03c9y, \u03c9z are such that \u03c9 = \u03c9xxB +\u03c9yyB +\u03c9zzB " + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0001066_hnicem48295.2019.9072698-Figure3-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0001066_hnicem48295.2019.9072698-Figure3-1.png", + "caption": "Fig. 3. Division of the Watt II six-bar linkage into a crank-rocker four-bar linkage (on the left) and a double rocker four-bar linkage (on the right).", + "texts": [ + " The selection of precision points from the generated solar trajectory was made through Chebyshev\u2019s spacing, which is defined in (2) as adopted from [5]. = 12 ( + ) \u2212 12 ( \u2212 ) cos (2 \u2212 1)2 (2) where: = th precision point; = upper limit of the input; = lower limit of the input; = precision point number = 1,2, \u2026 , ; = total number of precision points. For the case study, the number of precision points was set to = 3. The procedure for the kinematic synthesis of the Watt II six-bar linkage shown in Fig. 2 is described as follows. First, the linkage was divided into two (2) four-bar linkages shown in Fig. 3. Then, the double rocker linkage on the right was synthesized using Freudenstein\u2019s method, while the crankrocker linkage on the left was synthesized using the analytical position method. Freudenstein\u2019s method, as presented in [18], is described as follows. From the double rocker four-bar linkage shown on the right in Fig. 3, with link lengths , , , and , input angle , and output angle , the Freudenstein equation can be obtained, as shown in (3). \u2212 Z cos + Z = cos( \u2212 ) where: = , = , = (3) From the choice of the number of precision points, a system of equations can be formulated by varying and depending on the positions prescribed by the precision points. For = 3, a system of three equations can be formulated, as shown in (4) in matrix form, which can be solved using linear algebra. cos , \u2212 cos , 1cos , \u2212 cos , 1cos , \u2212 cos , 1 = cos( , \u2212 , )cos( , \u2212 , )cos( , \u2212 , ) (4) In order to determine the values of the precision points, and were related to input and output of the solar apparent motion trajectory model", + " From these design criteria, the precision points can be determined from Chebyshev\u2019s spacing. Then, the link ratios , , and can be solved from the system of equations shown in (4) using linear algebra. Finally, the length of can be determined based on the available space for installation and the lengths of , , and can be obtained from the solved link ratios. This ground link was set to = 4.000 in this study. The analytical position method, as presented in [19], is described as follows. From the crank-rocker linkage shown on the left in Fig. 3, the limiting positions of the crank , the minimum transmission angle between and , and the time ratio between the rotation of the crank and the oscillation of the rocker were determined as design criteria. In this study, the minimum transmission angle was set to =30\u00b0, which is the ideal design value according to the literature [18]. The time ratio was set to unity. The limiting positions can be computed as the values of when the rocker is at its extreme positions. Using these design criteria, the link ratios of the crank-rocker mechanism can be expressed by (7) to (9)" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0000356_iecon.2019.8927121-Figure7-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0000356_iecon.2019.8927121-Figure7-1.png", + "caption": "Fig. 7. Similarity between the tubular PM RLM and PM surface motor [18]", + "texts": [ + " A drawback of this RLM is its quite complicated construction (two stators with two sets of windings, two Halbach arrays, etc.). More compact PM RLM variants are the so-called integrated versions of these electrical machines, which usually have a single surface mounted PM pattern on their mover, and diverse sets of coils on their stator for generating the magnetic flux needed for the two DoF motion. The movers of the first such variants were with a checkerboard PM array. This way of placing the PMs, given in Fig. 7, was already widely investigated in the literature in the case of planar (surface) motors [17]. A first set of its coils were placed on the axial direction, to ensure the rotation, while the other one was distributed circumferentially to provide the axial force [19], [20]. Several such RLMs were developed by P. Bolognesi and his collaborators [21], [22], [23]. These all were of modular construction, an efficient and relatively new approach in electrical machines [24]. The simplest, three-phase variant with two stator modules is given in Fig" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0000942_ilt-01-2020-0030-Figure2-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0000942_ilt-01-2020-0030-Figure2-1.png", + "caption": "Figure 2 Force analysis", + "texts": [ + " b 1 is the pitch angle of the ball, b 2 is the yaw angle and c j is the position angle of the ball (for No. j ball). Under the load, an angle called contact angle naturally forms between the ball and inner or outer raceways. At a low speed, the contact angle of inner race ai and that of outer race ae is approximately equal. As speed increasing, the values of two contact angles are no longer equal due to centrifugal force, ai > ae, the black areas are contact areas between races and elements under loading, as shown in Figure 2. Because the ball is completely restrained by the cage, the angular velocity of the cage and the angular velocity of revolution of the ball are both vm according to Figure 2. Furthermore, the velocity of spin motion for inner and outer race is as follows: vSi \u00bc v isinai vRcosb 2cosb 1sinai 1vRcosb 2sinb 1cosai (1) vSe \u00bc v esinae 1vRcosb 2cosb 1sinae vRcosb 2sinb 1cosae (2) where v i and v e are the angular viscosity of the inner ring and outer ring relative to the cage, respectively. As the outer ring fixed, absolute angular velocity of the cage can be expressed by vm = \u2013v e, the relative angular velocity of inner and outer ring v i, v e, angular velocity of rotation of the ball vR can be expressed respectively: v i \u00bc v 11 rreRri cosb 2cosb 1cosae 1 sinb 1sinae\u00f0 \u00de rriRre cosb 2cosb 1cosai 1 sinb 1sinai\u00f0 \u00de (3) v e \u00bc v 11 rriRre cosb 2cosb 1cosai 1 sinb 1sinai\u00f0 \u00de rreRri cosb 2cosb 1cosae 1 sinb 1sinae\u00f0 \u00de (4) vR \u00bc v rre cosb 2cosb 1cosae 1 sinb 1sinae\u00f0 \u00de Rre 1 rri cosb 2cosb 1cosai 1 sinb 1sinai\u00f0 \u00de Rri (5) where v is the absolute angular velocity of inner ring, rri and rre are the radius of inner and outer race, and Rri and Rre are the distances between the rolling center of the inner and outer race and the central axis, respectively, as shown in Figure 2; Rri and Rre are calculated as follows: Rri \u00bc dm 2 rricosai (6) Rre \u00bc dm 2 1 rrecosae (7) Because there is friction between the ball and the inner and outer races, the angular velocity of ball because of gyroscopic moment is almost close to 0, i.e.b 2 = 0, according to equations (3)-(5). The ratio of angular velocity of rotation of ball and relative angular velocity of inner and outer ring v i and v e are calculated using equations (8) and (9). vR v i \u00bc Rri rricos b ai\u00f0 \u00de (8) vR v e \u00bc Rre rrecos b ae\u00f0 \u00de (9) Note that the spin to roll ratio of inner and outer races are calculated by equation (10) and (11), respectively: vSi v ri \u00bc g 0 sinai 1 Rri rri tan ai b 1\u00f0 \u00de (10) Fatigue life analysis Yue Liu Industrial Lubrication and Tribology vSe v re \u00bc g 0 sinae 1 Rre rre tan ae b 1\u00f0 \u00de (11) where v ri and v re are the angular velocity of ball relative to the inner and outer race, g 0 = D/dm, D is the diameter of the ball and dm is the pitch diameter of bearing", + " (2001), which yields the following equation: b 1\u00bc arctan Msi Mse 11 g 0 cosae 1 g 0 cosai 11 sinai 1 2sinae Msi Mse 11 g 0 cosae 1 g 0 cosai 1 1 cosai 1 2 cosae 1 g 0\u00f0 \u00de1A 0 B@ 1 CA (12) Note that A is the distance between raceway groove curvature centers of ball bearing having no clearance and having an initial contact angle. The ratio of the pin friction moment for inner and outer race Msi/Mse can be obtained by integrating the distributed friction in the contact area with the contact normal moment. Msi Mse \u00bc FniaiE x i\u00f0 \u00de FneaeE x e\u00f0 \u00de (13) Both Fni and Fne are the contact forces of inner and outer ring, respectively, and E(x i) and E(x e) are the complete elliptic integral of the second kind defined by ISO/TS 16281. The force analysis between ball and races is shown in Figure 2. The centrifugal force of the ball is given as follows: Fc \u00bc prD3v2 mdm 12 (14) where r is thematerial density. The gyroscopic couple of the ball can be computed as follows: Mg \u00bc prD5vRvmsinb 1 60 (15) Moreover, the equilibrium equations of mechanics are given as follows: Qci j\u00f0 \u00desinai j\u00f0 \u00de Qce j\u00f0 \u00desinae j\u00f0 \u00de Ffi j\u00f0 \u00decosai j\u00f0 \u00de 1Ffe j\u00f0 \u00decosae j\u00f0 \u00de \u00bc 0 (16) Fc j\u00f0 \u00de 1Qci j\u00f0 \u00decosai j\u00f0 \u00de Qce j\u00f0 \u00decosae j\u00f0 \u00de 1Ffi j\u00f0 \u00desinai j\u00f0 \u00de Ffe j\u00f0 \u00desinae j\u00f0 \u00de \u00bc 0 (17) Qci j\u00f0 \u00de 1Qce j\u00f0 \u00de D 2 \u00bc Mg j\u00f0 \u00de (18) where Qci(j) and Qce(j) are respectively the contact forces between the balls and races, Ffi(j) and Ffe(j) are respectively friction forces of inner and outer races, as shown in Figure 2. According to Harris and Barnsby (1998), the friction force Ff acting over the contact surface is obtained by integration: Ff \u00bc \u00f0 t f dA \u00bc ab \u00f01 1 1 \u00f0ffiffiffiffiffiffiffiffi1 q2 p ffiffiffiffiffiffiffiffi 1 q2 p Cv Ac A0 mas 1 1 Ac A0 t 1 fN 1 t 1 f lim -1 dtdq (19) Depending on the direction of sliding velocity, the coefficient of sliding Cv is 1 or 1. Ac is the area associated with asperity\u2013 asperity contact, A0 is the total contact area, t fN is the Newtonian portion of the friction shear stress, andt flim is the maximum shear stress that can be sustained at the contact pressure. Note that for full-film lubrication, t f \u00bc t 1 fN 1 t 1 f lim -1 . Contact pressure s at any point (x, y) is determined from the following equation (20): s\u00bc 3Qc 2pab 1 x a 2 y b 2 \" #1=2 (20) where a and b are semi-major and semi-minor axis of the projected contact ellipse, respectively, as shown in Figure 2, andQc is the contact force. The angle positions of elements in the plane are shown in Figure 3. When the outer ring is fixed, under the combined action of contact force, centrifugal force and friction force, the equilibrium equations of mechanics of inner ring are given by the following equations: Fa \u00bc XZ j\u00bc1 Qci j\u00f0 \u00desinai j\u00f0 \u00de Ffi j\u00f0 \u00decosai j\u00f0 \u00de (21) Fr \u00bc XZ j\u00bc1 Qci j\u00f0 \u00decosai j\u00f0 \u00de Ffi j\u00f0 \u00desinai j\u00f0 \u00de cosc j (22) M \u00bc XZ j\u00bc1 Qci j\u00f0 \u00desinai j\u00f0 \u00de Ffi j\u00f0 \u00decosai j\u00f0 \u00de Rri 1Qci j\u00f0 \u00de D 2 cosc j (23) Fatigue life analysis Yue Liu Industrial Lubrication and Tribology The curvature center of the inner groove, outer groove and the ball center areOii,Oei andOj, respectively, after deformation, as O 0 ij,O 0 ej,O 0 j " + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0001402_s12239-020-0083-y-Figure2-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0001402_s12239-020-0083-y-Figure2-1.png", + "caption": "Figure 2. Finite element model for contact analysis.", + "texts": [ + " To minimize the contact stresses of the housing and the roller, the track groove of the housing and the radius of the roller are used as design variables and the optimization analysis proceeds. The GAF of an optimized CV joint is compared with an existing CV joint through multibody dynamic simulation. In addition, to verify the reliability of the multibody dynamic simulation model, the GAF of the CV joint is measured using a rotating durability tester, and the experimental results are compared with the simulation results. 2.1. Finite Element Model Figure 2 shows finite element model used for the contact analysis between the groove track of the housing and the roller. The finite element model was created using HyperMesh (Altair Engineering, 2018a). As the roller and housing models are symmetrical, as shown in Figure 2, they are analyzed with a 1/4 symmetrical model. The roller material is fine machined 52100 steel and the housing is forged 4121-H steel. The dimensions and detailed properties of the model are presented in Table 1. As per the assumptions of contact theory in the Hertz theory, the contact surface was frictionless and smooth. The load used in the simulation was 14,492 N and this load was calculated by T = 900 Nm, that is the maximum engine torque during the starting conditions of a midsized vehicle (Schmelz et al" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0003766_0954406214543293-Figure1-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0003766_0954406214543293-Figure1-1.png", + "caption": "Figure 1. The simplified model for a trot and a transverse gallop. The arrow A indicates the direction of running. Each leg is characterized by a stiffness k. The uncompressed length of leg is l0. Two legs are attached to the shoulder joint, and the other two are attached to the hip joint. The angle of the leg is characterized by . The angle of the body is characterized by . The length of half beam is characterized by L.", + "texts": [ + " Sections \u2018\u2018Explanation of the PAVD property\u2019\u2019 and \u2018\u2018Conditions under which the PAVD emerges\u2019\u2019 explain the reason for and conditions of the emergence of the PAVD in transverse galloping pattern. Section \u2018\u2018Suitable ranges of speeds for the trot and the transverse gallop\u2019\u2019 discuss the suitable ranges of speeds for the trot and the transverse gallop. The last section presents the useful information for design of legged system. Models and searching methods for the trot and the transverse gallop Model and dynamic equations. A simplified planar model is employed to study the dynamic properties (Figure 1). The simplified model has been demonstrated to be helpful in capturing important qualitative properties of quadruped gaits, such as trot, bound, gallop.10,12\u201314,17 This simplification here can help us focus on the dynamic properties of different gaits. In this paper, the body is modeled as a rigid beam, the center of mass is at its geometrical center and the entire musculoskeletal spring system of a leg is represented as a massless linear spring.12,13,18,19 The leg can be treated as a spring whose stiffness does not change with varying speeds18,20 and the inertial effects of legs are negligible compared to the inertial effects of the body.10,19 Accordingly, it is reasonable to treat the leg of the model as a massless linear spring. Two legs are attached to the shoulder joint and the other two are attached to the hip joint. The model is presented in Figure 1, while the associated parameters are given in Table 1. The equations of motion are obtained by using the Lagrangian approach and can be stated in the following form \u20acx \u00bc k m \u00bdcrt\u00f0l0 lrt\u00de sin rt \u00fe crl\u00f0l0 lrl\u00de sin rl \u00fe cft\u00f0l0 lft\u00de sin ft \u00fe cfl\u00f0l0 lfl\u00de sin fl \u20acy \u00bc k m \u00bdcrt\u00f0l0 lrt\u00de cos rt \u00fe crl\u00f0l0 lrl\u00de cos rl \u00fe cft\u00f0l0 lft\u00de cos ft \u00fe cfl\u00f0l0 lfl\u00de cos fl g \u00f01\u00de at University of Birmingham on June 5, 2015pic.sagepub.comDownloaded from \u20ac \u00bc kL J \u00bd crt\u00f0l0 lrt\u00decos\u00f0 rt \u00de crl\u00f0l0 lrl\u00decos\u00f0 rl \u00de \u00fe cft\u00f0l0 lft\u00decos\u00f0 ft \u00de\u00fe cfl\u00f0l0 lfl\u00decos\u00f0 fl \u00de where lrl \u00bc ffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi \u00f0x L cos xtdrl \u00de 2 \u00fe \u00f0 y L sin \u00de2 q \u00f04\u00de lft \u00bc ffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi \u00f0x\u00fe L cos xtdft \u00de 2 \u00fe \u00f0 y L sin \u00de2 q \u00f05\u00de lfl \u00bc ffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi \u00f0x\u00fe L cos xtdfl \u00de 2 \u00fe \u00f0 y L sin \u00de2 q \u00f06\u00de rt \u00bc arctan\u00bd\u00f0xtdrt x\u00fe L cos \u00de=\u00f0 y L sin \u00de \u00f07\u00de rl \u00bc arctan\u00bd\u00f0xtdrl x\u00fe L cos \u00de=\u00f0 y L sin \u00de \u00f08\u00de ft \u00bc arctan\u00bd\u00f0xtdft x L cos \u00de=\u00f0 y\u00fe L sin \u00de \u00f09\u00de fl \u00bc arctan\u00bd\u00f0xtdfl x L cos \u00de=\u00f0 y\u00fe L sin \u00de \u00f010\u00de ci \u00bc 1 the ith leg is on the ground 0 the ith leg is not on the ground i \u00bc rt, rl, ft, fl \u00f011\u00de Searching methods for achieving the periodic motion of trot" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0003499_muh-1404-6-Figure3-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0003499_muh-1404-6-Figure3-1.png", + "caption": "Figure 3. FE model of a 3-axis open category milling machine tool.", + "texts": [ + " The definition of the contact type between any subsystems decides the mathematical relation between the in-contact nodes on both sides. There are three main formulation types: penalty function, Lagrange multiplier, and augmented Lagrange. The last type is preferred in this work. The FE model of the mechanical substructures is generated from their perspective CAD models using tetrahedron elements. The tetrahedron elements are proven to be more suitable than bricks elements to simulate machine tool structures. Figure 3 shows the FE model of the machine tool mechanical structure. The main characteristic of interest when evaluating a FE model is the mesh size. A fine mesh size leads to converged results, but it increases the degrees of freedom, which increases the computational cost as well. However, local refinements in mesh size can be specified at critical regions such as the TCP and work table using the sphere of influence. To obtain a mesh-independent model with the minimum computation time, constitutive iterations are carried out to properly select mesh characteristics in order to achieve convergence of the results to the exact value within an accepted residual error" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0000987_0954406220915499-Figure3-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0000987_0954406220915499-Figure3-1.png", + "caption": "Figure 3. Schematic of the wheel alignment parameters. (a) The caster angle, (b) the camber angle and the steering axis inclination angle, and (c) the toe-in angle (from top view).", + "texts": [ + " At the same time, the steering knuckle arms of the second axle move through the first transition rod, intermediate rocker, second transition rod, second rocker, and posterior longitudinal rod. Hence, the steering of the wheels of the second axle is in coordination with that of the first axle. A schematic of the intermediate steering transmission mechanism is shown in Figure 2. The wheel alignment parameters considered in this paper are the caster angle a, the steering axis inclination angle , the camber angle c, and the toe-in angle . The schematic of the wheel alignment parameters is shown in Figure 3. The steering axle not only needs to ensure the vehicle has a stable steering function, but also needs to make the steering vehicle wheel have a self-righting effect to ensure the vehicle can travel stably in a straight line. The caster angle can ensure a stable self-righting performance under high-speed driving conditions; The steering axis inclination angle can ensure a stable self-righting performance under low speed and heavy load conditions. The function of the camber angle is to prevent the wheel from tilting inward under heavy loads", + " The shimmy system has 9-DOF: h1 is the swing angle at which the right wheel of the first axle rotates around the kingpin; h2 is the swing angle at which the left wheel of the first axle rotates around the kingpin; 1 is the lateral swing angle of the first axle; h3 is the swing angle at which the right wheel of the second axle rotates around the kingpin; h4 is the swing angle at which the left wheel of the second axle rotates around the kingpin; 2 is the lateral swing angle of the second axle; d1 is the swing angle of the first rocker; d2 is the swing angle of the intermediate rocker; d3 is the swing angle of the second rocker. In this study, the mathematical model of the shimmy system is established by the Lagrange equation which can be expressed as d dt @T @ _qk @T @qk \u00fe @U @qk \u00fe @D @ _qk \u00bc Qk \u00f01\u00de where qk represents 9-DOF of the system, T represents the system\u2019s kinetic energy, U represents the system\u2019s potential energy, D represents the system\u2019s dissipated energy, and Qk represents the nine generalized forces to which the system is subjected. According to Figure 3, kinetic energy, potential energy, dissipated energy, and the nine generalized forces of the dual-axle steering shimmy system are obtained as follows. The kinetic energy of the shimmy system is T \u00bc 1 2 I1\u00f0 _ 1 2 \u00fe _ 2 2\u00de \u00fe 1 2 J1 _\u20191 2 \u00fe 1 2 I2\u00f0 _ 3 2 \u00fe _ 4 2\u00de \u00fe 1 2 J2 _\u20192 2 \u00fe 1 2 Ic1 _ 1 2 \u00fe 1 2 Ic2 _ 2 2 \u00fe 1 2 Ic3 _ 3 2 \u00f02\u00de where Ii (i\u00bc 1,2) is the moment of inertia of the wheel around the kingpin in the ith axle, Ici (i\u00bc 1,3) is the moment of inertia of the ith rocker, Ic2 is the moment of inertia of intermediate rocker, and Ji (i\u00bc 1,2) is the moment of inertia of the ith axle around its side off-axis" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0003788_detc2014-34213-Figure17-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0003788_detc2014-34213-Figure17-1.png", + "caption": "Fig. 17 Eight-bar linkages with input joints A, B and B0", + "texts": [ + " The other is the three-DOF eight-bar linage is decomposed into a two-DOF planar seven-bar linkage with the input joint B0 held, as shown in Fig. 16(b). Thus, for the Stephenson six-bar linkages or the two-DOF seven-bar planar linkage, the singularity happens when the joint E, B and F are collinear or the joints D, C and C0 are collinear. Therefore, the three-DOF eight-bar linkage is at the singular positions. For the Type-222 input condition, let the inputs be given through the joints A, B and B0 as an example, as shown in Fig. 17(a). There are also two decomposition ways. The three-DOF eight-bar linkage can be decomposed into a Watt planar sixbar linkage with the input joints B and B0 held, as shown in Fig. 17(a). The linkage can be also decomposed into a twoDOF planar seven-bar linkage with the inputs joint B0 held, as shown in Fig. 17(b). Thus, for the Watt planar six-bar linkages or the two-DOF seven-bar planar linkage being at singular positions, the joints A0, E and B are collinear or the joints D, C and C0 are collinear. Therefore, the three-DOF eight-bar linkage is at the singular positions. Input joints not in the same six-bar loop In this group, the inputs are given through the joints not in the same six-bar loop of the eight-bar linkages. Two cases can be found according to whether one of the joints B, F and B0 is the input joint" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0002861_978-3-030-55889-5-Figure6.5-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0002861_978-3-030-55889-5-Figure6.5-1.png", + "caption": "Fig. 6.5 Single force acting on a plate and on a slab. The stresses in the plate, as well as the shear force (vn), behave as 1/r , but since the plate deflection w is the threefold integral of vn , the deflection w = r2 ln r is finite at r = 0", + "texts": [ + " The dislocation \u00b10.5, which triggers the influence function, remains the same, but the flanks of the influence function become steeper and steeper when the distance shrinks, h \u2192 0, and so the influence function bulges out in the neighboring spans. A similar effect was observed in the plate in Fig. 6.4, where a slight offset of two supportingwalls produced a twistingmoment, which required large support reactions to stabilize the plate. Such situations are not so rare. If we place a single force P on a plate, see Fig. 6.5, and we draw circles with radius r around the source point, the horizontal stresses integrated over the circle must be equal to the point load, and this also in the limit r \u2192 0. What we integrate, though, are not the horizontal stresses but the horizontal tractions. If S is the stress tensor in the plate S = [ \u03c3xx \u03c3xy \u03c3yx \u03c3yy ] , (6.2) 364 6 Singularities - and we are on the circle, the tractions t in the direction of the normal vector n = {nx , ny}T are t = S n = [ \u03c3xx \u03c3xy \u03c3yx \u03c3yy ] [ cos \u03d5 sin \u03d5 ] on the circle , (6", + " When we place a force ei at a point x of an infinitely long plate, the stress tensor at a point y in the direction of a vector \u03bd = {\u03bd1, \u03bd2}T has the components [6], (4.7), Ti j ( y, x) = \u2212 1 4\u03c0 (1 \u2212 \u03bd) r [ \u2202r \u2202\u03bd ((1 \u2212 2 \u03bd) \u03b4i j + 2 r,i r, j ) \u2212 (1 \u2212 2 \u03bd)(r,i \u03bd j ( y) \u2212 r, j \u03bdi ( y))] (9.124) where r,i := \u2202r \u2202yi = yi \u2212 xi r . (9.125) If the point y lies on a circle with radius r and centered at x, \u03bdi and r,i are identical \u03bd1 = r,1 = cos \u03d5 \u03bd2 = r,2 = sin \u03d5 , (9.126) and so we have at any such point \u2202r \u2202\u03bd = \u2207r \u2022 \u03bd = [ cos \u03d5 sin \u03d5 ] \u2022 [ cos \u03d5 sin \u03d5 ] = cos2 \u03d5 + sin2 \u03d5 = 1 . (9.127) Since the force in Fig. 6.5 a points in x-direction, we let in (9.124) i = 1, and so the traction vector on the circle has the two components tx = T11 = \u2212 1 4 (1 \u2212 \u03bd) \u03c0 r \u00b7 [(1 \u2212 2 \u03bd) + 2 cos2 \u03d5] (9.128) ty = T12 = \u2212 1 4 (1 \u2212 \u03bd) \u03c0 r \u00b7 [2 cos\u03d5 sin \u03d5] (9.129) and the integrals are \u222b 2\u03c0 0 tx d\u03d5 = \u22121 r \u222b 2\u03c0 0 ty d\u03d5 = 0 . (9.130) Figure 9.11 and the following text is a supplement to page 129. The Taylor series of a curve w(x) w(x) = w(0) + w\u2032(0) x + 1 2 w\u2032\u2032(0) x2 + \u00b7 \u00b7 \u00b7 (9.131) leads to the representation w(x) = \u222b l 0 G0(y, 0) p(y) dy + \u222b l 0 G1(y, 0) p(y) dy \u00b7 x + 1 2 \u222b l 0 G2(y, 0) p(y) dy \u00b7 x2 + \u00b7 \u00b7 \u00b7 (9" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0002228_icem49940.2020.9270810-Figure6-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0002228_icem49940.2020.9270810-Figure6-1.png", + "caption": "Fig. 6. Experimental set up line diagram.", + "texts": [ + " It can be seen that the features can be extracted at any layer of scattering and the wavelet decomposition recovers all of the signal information at any level. Also, the filters in the scattering network are predesigned wavelet filters and do not require learning. However, the filter parameter and length are designed for a specific operating frequency and needs to be updated according to the operating frequency when a variable frequency drive is used. The experimental setup consists of an induction motor fed by a SiC-based inverter, built in the lab, mechanically coupled with a dc generator controlled by an ABB_DCS800 motor drive. Fig. 6 shows the line diagram of the set up. The dc generator was able to generate torque pulsations at different frequency levels by controlling the armature current with the ABB drive to emulate motor fault signatures. The bottom part of Fig. 7 shows the induction motor coupled with the dc generator. The drive circuit of both machines are shown in the upper part of Fig. 7 along with the data acquisition board. The SiC-based inverter is able to switch very fast, which can create computational and detection accuracy challenges in MCSA" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0002119_icaccm50413.2020.9212874-Figure11-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0002119_icaccm50413.2020.9212874-Figure11-1.png", + "caption": "Fig. 11. Cylindrical fixed support (case2).", + "texts": [ + " 60 Authorized licensed use limited to: Carleton University. Downloaded on May 30,2021 at 15:29:13 UTC from IEEE Xplore. Restrictions apply. A Quadcopter\u2019s propeller is subjected to horizontal bending due to collision of the propeller with obstacles or a wall. To calculate the bending deformation under the applied rotational moment about the rotation axis, the propeller has been fixed by both the tip as shown in Figure 10.In addition, it is fixed at the point where it is connected to the motor shaft (Figure 11).Figure 12 shows the applied rotational velocity of 897 rad/s [8].Figure 13, 14 and 15, 16 show the total deformation and equivalent stress for CFRP and GFRP materials, respectively. The comparison of the obtained results is presented in Table III. 61 Authorized licensed use limited to: Carleton University. Downloaded on May 30,2021 at 15:29:13 UTC from IEEE Xplore. Restrictions apply. Fig. 14. Total deformation of GFRP propeller (case2). Fig. 15. Equivalent stress for CFRP propeller (case2). Table III" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0002324_iceca49313.2020.9297379-Figure6-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0002324_iceca49313.2020.9297379-Figure6-1.png", + "caption": "Figure 6: Configuration of motors, their direction of rotation and orientations of axis for gliding mode", + "texts": [ + " Since there are no control surfaces to change the directions as in the conventional gliders, this drone uses variable thrust and angular moment principles to maneuver in gliding mode. Brushless DC motors are used in this drone whose angular velocit ies can be varied by varying frequency of the current supplied with the help of ESC (Electronic Speed Controller). The configuration of motors and their direct ion of rotation and orientations of axis for gliding mode are shown in the figure below. Let, x, y and z be three perpendicular axis attached to the fixed frame as shown in figure 6. Roll axis, Pitch axis and Yaw axis be three perpendicular axis attached to the frame of the drone as shown in figure 6. T1, T2, T3 and T4 be the thrusts developed by the motors 1, 2, 3 and 4 respectively. The length between motors 1 and 2 (and so between motors 3 and 4) be 2*L1 and the length between motors 1 and 3 (and so between motors 2 and 4) be 2*L2. ax, ay, and az be accelerat ions of the drone on x, y and z axis respectively. \u03d5, \u03b8 and \u03c8 be Pitch, Roll and Yaw angles respectively. \u03c91, \u03c92, \u03c93 and \u03c94 be angular velocit ies of motors 1, 2, 3 and 4 respectively. M1, M2, M3 and M4 be the moments generated by motors 1, 2, 3 and 4 respectively" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0001381_s1068798x20060210-Figure3-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0001381_s1068798x20060210-Figure3-1.png", + "caption": "Fig. 3. Rusetskii\u2019s test bench.", + "texts": [ + " (1) may be neglected, since the drag force cannot be measured. In the Sudnishnikov test bench, the operator is replaced by a simulator, whose role is to fix the jackhammer in the vertical position and to ensure that the hammer bit applies force to the object. The object employed is an actual sample, rather than a simulator. That complicates the research. This deficiency was taken into account in developing subsequent test benches. One such test bench was developed by F.I. Rusetskii to monitor the energy of jackhammer impact (Fig. 3) [6]. Plate 2 with measuring equipment is attached to column 1. Arm 3 moves along guides of the column and supports pneumatic clamp 4, while clip 5 holds the hammer handle. The hammer impacts are transmitted by pin 6, compressed by two cast-iron shoes 7. Pressure is transmitted to the shoes through rod 9, by means of piston 8 of the hydraulic press. Piston 10 raises the pin to its initial position. Oil is supplied to cavity 11 from a device outside the test bench. e f t m s F RUSSIAN ENGINEERING RESEARCH Vol. 40 No. 6 The test bench is used as follows. The jackhammer (not shown in Fig. 3) is mounted in a special device, the lower ring of which is at pin 6. Clamp 4 grips the hammer, whose handle is held in clip 5. Under the action of the hammer impacts, the pin, compressed by the hydraulic press, gradually descends. When the pin reaches its lower position, the hammer is automatically switched off. The initial and final pin positions are recorded. The hammer power is assessed from the time for the pin to travel this distance. The test bench is calibrated by repeatedly dropping a load from a fixed height, until the pin reaches its final position" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0002989_978-3-642-16214-5_9-Figure9.29-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0002989_978-3-642-16214-5_9-Figure9.29-1.png", + "caption": "Fig. 9.29. Porsche synchronizer. 1 Idler gear; 2 synchronizer hub with dog gear; 3 locking belt; 4 synchronizer ring; 5 circlip; 6 gearshift sleeve; 7 guide sleeve; 8 pad; 9 end stop", + "texts": [ + " Because of its large power transmission surface AR, it is suitable wherever there is a requirement for very high synchronizer performance. The cone angle \u03b1 of a multi-plate synchronizer is 90\u00b0. To be operated with the same gearshift effort as a single-cone synchronizer with \u03b1 = 6.4\u00b0, according to Equation 9.2 a multi-plate synchronizer of the same effective diameter must have j = 9 friction surfaces. The lengths of the two synchronizers are then roughly equal. Multi-plate synchronizers are complex and costly. The Porsche system locking synchronization (Figure 9.29) has a self-reinforcing locking effect, preventing premature gearshift action before speeds have been synchronized. The Porsche synchronizer requires relatively little gearshift effort, but its high manufacturing cost means it is no longer of practical significance. The quality of the synchronization process of the Porsche synchronizer is very much subject to variations in the friction coefficient. The synchronization process will be only briefly described. The slotted synchronizer ring 4 located in front of the selector teeth is crowned" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0003971_robio.2014.7090583-Figure1-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0003971_robio.2014.7090583-Figure1-1.png", + "caption": "Fig. 1. Quadrotor UAV.", + "texts": [ + " The notations addressed in this paper are listed as follows: \ud835\udc65(\ud835\udc61), \ud835\udc66(\ud835\udc61), \ud835\udc67(\ud835\udc61) Coordinate of the quadrotor at time \ud835\udc61 in Cartesian space \ud835\udf19(\ud835\udc61), \ud835\udf03(\ud835\udc61), \ud835\udf13(\ud835\udc61) Euler angles: roll, pitch, and yaw at time \ud835\udc61, respectively \u03a9\ud835\udc56 Angular speed at \ud835\udc56\ud835\udc61\u210e motor \ud835\udc53\ud835\udc56 Thrust at \ud835\udc56\ud835\udc61\u210e motor \ud835\udc3f Half distance between two motors on the same axis \ud835\udc40 Mass of the quadrotor \ud835\udc54 Gravitational acceleration constant \ud835\udc3c\ud835\udc65\ud835\udc65, \ud835\udc3c\ud835\udc66\ud835\udc66, \ud835\udc3c\ud835\udc67\ud835\udc67 Moment of inertia around x-axis, y-axis, and z-axis, respectively \ud835\udc4f Thrust coefficient \ud835\udc51 Drag coefficient \ud835\udc57\ud835\udc5f Propeller inertia coefficient The model of the quadrotor, shown in Fig. 1, has four rotors to generate the propeller forces, \ud835\udc531, \ud835\udc532, \ud835\udc533, \ud835\udc534. The vertical motion is achieved by adjusting the rotor speeds altogether with the same quantity. The two pairs of rotors (\ud835\udc451, \ud835\udc453) and (\ud835\udc452, \ud835\udc454) turn in opposite directions in order to balance the moments and produce yaw motion as desired. Meanwhile, forward or backward motion is achieved by pitching in the desired direction by increasing \ud835\udc533 or \ud835\udc531 and decreasing \ud835\udc531 or \ud835\udc533 to maintain the total thrust. Finally, a sideways motion is achieved by rolling in the desired direction by increasing \ud835\udc532 and decreasing \ud835\udc534 to maintain the total thrust" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0003634_icems.2014.7013745-Figure6-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0003634_icems.2014.7013745-Figure6-1.png", + "caption": "Fig. 6. Magnetic field distribution when outer motor is energized alone It can be seen that when +D axis currents of inner motor act alone, magnetic flux will go through outer air-gap to stator, and when \u2013D axis currents of inner motor act alone, little magnetic flux will go through outer motor, as shown in Fig. 5(a) and Fig. 5(b). It is obvious that magnetic fluxes of outer", + "texts": [ + " In the PMRM, because of the doubly salient structure of outer motor, magnetic circuit of outer motor will be different when outer rotor rotates to a different position. Magnetic field distribution of PMRM when permanent magnets act alone is shown in Fig. 3 and Fig. 4. As Fig. 3 shows, when outer rotor rotates to this position, magnetic flux will go through two aligned teethes of stator and outer rotor. When there are no aligned teethes, magnetic flux will go through adjacent teethes as Fig. 4 shows. Magnetic field distributions of PMRM when inner rotor or stator armature windings energized alone are shown in Fig. 5 and Fig. 6. motor will increase or decrease when +D axis or \u2013D axis currents of inner motor act alone. By comparing Fig. 5(c) with Fig. 5(d), it can be seen that when +Q axis currents act alone, magnetic fluxes of outer motor little change. From Fig. 5, it can be obtained that D axis currents of inner motor has much stronger effect on magnetic field distribution of outer motor than Q axis currents of inner motor. As Fig. 6(a), Fig. 6(b) and Fig. 6(c) show, stator teeth of phase C and outer rotor teeth align completely. When positive currents of phase C act alone, magnetic fluxes of inner motor decrease obviously. Meanwhile, when negative currents of phase C act alone, magnetic fluxes of inner motor are little changed. As Fig. 6(d), Fig. 6(e) and Fig. 6(f) show, stator teeth of phase C and outer rotor teeth align incompletely. When positive currents or negative currents of phase C act alone, magnetic fluxes of inner motor are little changed. In the actual operation process, when single phase of outer motor is energized, stator teeth of this phase and outer rotor teeth align incompletely. Based on the situation above, it can be obtained that the influence that outer motor acts on inner motor is much weaker than the influence that inner motor acts on outer motor from Fig. 5 and Fig. 6. Therefore, the rests of the paper focus on the analysis of influence that inner motor acts on outer motor. The numerical results of electromagnetic coupling effect on magnetic flux that inner motor acts on outer motor are shown in Fig. 7, Fig. 8, Fig. 9 and Fig. 10. Current shown in Fig. 7 and Fig. 8 is Q axis currents of inner motor, magnetic flux indicates that created by inner motor, and angle indicates mechanical angle that outer rotor rotates counterclockwise. Fully unaligned teeth position indicates 0 degree and fully aligned teeth position indicates 14 degree" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0000463_s12541-019-00282-y-Figure7-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0000463_s12541-019-00282-y-Figure7-1.png", + "caption": "Fig. 7 Force and moment", + "texts": [ + "\u00a0(8) and (9), when the mopping pads rotate, the force in the y-axis direction is determined by the rotating angle and the friction constant . Furthermore, the moment T at the center of the mopping pad can be expressed as follows: Equation\u00a0(10) indicates that the moment is changed by the friction angle and the friction constant . Furthermore, when the rotating direction of the motor changes, the direction of the friction force also changes. The direction of the force can be determined by the \u00b1 \u03c9 direction, as shown in Fig.\u00a07. (9) Fy = r \u222b 0 2 \u222b 0 (fcos )rd dr = r \u222b 0 2 \u222b 0 [ k ( zd + r cos ) cos ] rd dr = kr3 3 . (10) T = r \u222b 0 2 \u222b 0 (r \u22c5 f )rd dr = r \u222b 0 2 \u222b 0 [ k ( zd + r cos )] r2d dr = kr4 . 1 3 From Fig.\u00a07 and Eqs.\u00a0(8)\u2013(10), it can be observed that the force and direction can be controlled by the tilting angle and the pad\u2019s rotation direction. Figure\u00a08 shows the top view of the proposed robot. Each pad has three variables , , and \u03c9. is the angle of the pad, is the tile angle of the pad, and \u03c9 is the rotational velocity of the pad. The quadruple pads are arranged symmetrically with a distance of l along the x and y axes, and a total of 12 variables are controllable. Although controlling a robot with 12 degrees of freedom can enable various motions, it comes with the disadvantage of added weight and cost since the motors, controller, and battery are large and expensive" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0002540_iccr51572.2020.9344388-Figure6-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0002540_iccr51572.2020.9344388-Figure6-1.png", + "caption": "Figure 6. Simulation on stepping stones: (a) Stable walking for LIPM-like walking pattern; (b) Torso swinging for fixed-peroid walking pattern.", + "texts": [ + "t \u23a9\u23aa\u23aa \u23aa\u23a8 \u23aa\u23aa\u23aa \u23a7 \u0308 + \u0307 + = + ( ) \u0308 + (\u0307 ) \u0307 = 0\u2212 , < , < ,, < < ,\u0308 = ( \u2212 ) + \u0307 \u2212 \u0307 + \u0308 +\u0308 = ( \u2212 ) + ( \u0307 \u2212 \u0307) + \u0308 +{2,3,4} = 0< < (59) where is the constant weight of the tolerance of \u0308. 146 Authorized licensed use limited to: East Carolina University. Downloaded on June 21,2021 at 06:17:49 UTC from IEEE Xplore. Restrictions apply. V. SIMULATION AND RESULTS The simulation model of the robot is developed in Simscape Multibody and the control system is running at a frequency of 1 kHz. To verify the proposed approaches, a terrain with horizontal stepping stones of the same height is built in the simulation shown in Fig. 6. The preset terrain consists of 12 stepping stones spaced at different intervals (0.45, 0.29, 0.48, 0.37, 0.49, 0.32, 0.28, 0.46, 0.44, 0.35, 0.36 and 0.41 m), each of which is set up to 40 cm in length according to the size of the robot's foot. Both the LIPM-like walking approach and fixed period walking approach are tested in the same simulation conditions. More specifically, the mechanical parameters and control parameters for the simulations are presented in Table \u2160. Fig. 7 shows the velocity indicators, angular momentum, torso angle, and step time in the simulation of the LIPM-like walking approach (presented in Sec" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0000603_iccas47443.2019.8971558-Figure4-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0000603_iccas47443.2019.8971558-Figure4-1.png", + "caption": "Figure 4. Section design structure", + "texts": [ + " System Design Based on application analysis, the robot should have a slender structure with redundant DOFs. After consideration of previous TakoBot 1 prototype capabilities, we made changes in the design and actuation system as well. After numerous experiments with TakoBot 1, demonstrated some limitations related to the design. For example, a robot could not perform torsional motion during the work which led to the accumulation of strain energy inside of manipulator. Secondly, in case of bending it could not perform a pure bending shape. In the new prototype, we added a passive sliding mechanism (Fig.4). Proposed sliding mechanism demonstrated the following benefits: 1) Smart bending stress distribution: This ability helped to the manipulator to bend purely and it distributed spring potential energy to the whole segment of the manipulator. In comparison with previous, it required less torque and less load to the cables. 2) Torsional motion: passive sliding disc mechanism provided better dexterity ability in case bending helical motion. Helical motion increases torsional stress of the structure unless a disc does not generate yawing around the z-axis. In spite of the heavier weight of the TakoBot 2 was able to perform better and more accurate. Passive sliding disc mechanism works for all discs attached to the backbone except base and end-effector discs, which means total length always remains constant. Fig.4. demonstrates a sliding mechanism structure. According to the design, travelling distance of the disc is 10mm, while the disc diameter is 50mm, an average distance of the single section (between discs) about 35 mm. 3D printed discs connected by coil compression springs, such design provides stiffness to the manipulator. The spring constant could be variable depends on motor torque, in this prototype we used spring with constant 0.63 N/mm. Each section consists of four segments, end segments discs connected by four springs, while mid-section segments connected by 8 springs between discs" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0002989_978-3-642-16214-5_9-Figure9.7-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0002989_978-3-642-16214-5_9-Figure9.7-1.png", + "caption": "Fig. 9.7. Single-cone synchronizer (ZF-B), see also Figure 9.12. 1 Idler gear with needle roller bearings;", + "texts": [ + "16]: \u2022 synchronizing mechanism for each individual gear, \u2022 central synchronizer for the whole transmission (Section 9.2.6) and \u2022 speed synchronization by the prime mover (Section 9.2.6). It is technically possible to dispense with synchronizers when \u2022 there is a small gear step between the gears (\u03c6 < 1.15) or \u2022 the masses of the gears are small, for example in a motorcycle gearbox. Synchromesh is frequently omitted in commercial vehicles for reasons of economy and to improve transmission reliability. Transmissions without synchromesh are more robust. A mechanical synchromesh unit as shown in Figure 9.7 frictionally matches the different speeds of the transmission shaft 6 (and the gearshift sleeve 5 rotationally fixed to it) and of the idler gear to be shifted 1. When their speeds have been synchronized, the elements are positively engaged. This synchromesh unit incorporates a frictionally engaged clutch and a positive locking clutch (see also Figure 10.3 \u201cSystematic classification of moving-off elements\u201d). This section describes the gear changing process using as an example a notional vehicle with a 2-speed coaxial countershaft transmission (Figure 9" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0003565_ipec.2014.6869805-Figure5-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0003565_ipec.2014.6869805-Figure5-1.png", + "caption": "Fig. 5. Experimental system.", + "texts": [ + " Papp (21) By examining reactive power factor of the system, efficiency of power during free motion can be determined. Experiments are performed in order to examine two things: 1) Power factor difference due to structural difference of the system 2) Power factor difference between work space and joint space XY table (XY) and two link manipulator (2 link) are com pared to clarify power factor difference due to structure of the system. In addition, work space power factor and joint space power factor are compared in two link manipulator to show that power factor differs between work space and joint space. Figure 5 shows experimental systems. Fig. 5(a) is XY table composed of two sets of linear motors . .Toint space and work space of the system has common axis in XY table. Fig. 5(b) is two link manipulator composed of two sets of rotary motors . .Toint space and work space in two link manipulator requires coordinate transformation. Transformation function from joint space to work space is, To examine the power factor difference between different structured systems, the same motion was applied to XY table and two link manipulator. Circular motion given to the system and position response of each robots are shown in Fig. 6. Both systems are controlled by PD control with DOB" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0001523_iwcmc48107.2020.9148486-Figure4-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0001523_iwcmc48107.2020.9148486-Figure4-1.png", + "caption": "Fig. 4. Schematic calculation of high-altitude bright zone of UAV", + "texts": [ + " The parameters and configurations of 4G RRU product are shown in the following table 1: Gammar-436 TDD N78 adopts O-RAN architecture and supports the smooth evolution of future-oriented new technologies with a maximum 687 Authorized licensed use limited to: Univ of Calif Santa Barbara. Downloaded on August 01,2020 at 01:24:24 UTC from IEEE Xplore. Restrictions apply. output power of 4x20W. The 5G RRU product parameters and configuration are shown in the following table 2: \u2164. LINK BUDGET omnidirectional antenna at high altitude As shown in Figure 4, the propagation distance of the wireless line-of-sight of the Earth's curved surface can be expressed as: (km) 17 (m) 17 (m)t rd h h + (1) where: d is the effective receiving distance (km) between the base station antenna and the user; th is the effective height (m) of the base station antenna; rh is the effective height (m) of the user receiving antenna. At the same time, this paper adopts the free space propagation model for tethered UAV communication. The path loss model calculation formula can be expressed as: km MHz32" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0002119_icaccm50413.2020.9212874-Figure16-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0002119_icaccm50413.2020.9212874-Figure16-1.png", + "caption": "Fig. 16. Equivalent stress for GFRP propeller(case2).", + "texts": [], + "surrounding_texts": [ + "A Quadcopter\u2019s propeller is subjected to horizontal bending due to collision of the propeller with obstacles or a wall. To calculate the bending deformation under the applied rotational moment about the rotation axis, the propeller has been fixed by both the tip as shown in Figure 10.In addition, it is fixed at the point where it is connected to the motor shaft (Figure 11).Figure 12 shows the applied rotational velocity of 897 rad/s [8].Figure 13, 14 and 15, 16 show the total deformation and equivalent stress for CFRP and GFRP materials, respectively. The comparison of the obtained results is presented in Table III. 61 Authorized licensed use limited to: Carleton University. Downloaded on May 30,2021 at 15:29:13 UTC from IEEE Xplore. Restrictions apply. Fig. 14. Total deformation of GFRP propeller (case2). Fig. 15. Equivalent stress for CFRP propeller (case2). Table III. Results due to rotational velocity CFRP GFRP Max. Deformation 0.2012 mm 0.268 mm Max. stress 807.54 MPa 1127 MPa C. Vibration analysis The propeller has been tested for vibrational effect to find out the resonance frequencies for both the materials. The modal analysis has been performed in Ansys, Figure 17 and 18 show the frequency variation with first six mode shapes for CFRP and GFRP material, respectively. The color contour describes the range of deformation from minimum to maximum. Table 4 and figure 19 present the frequency comparison for CFRP and GFRP materials and figure 19 also present the comparison with previously publish work. Fig.17. CFRP based propeller\u2019s vibration frequencies and mode shapes (case3). 62 Authorized licensed use limited to: Carleton University. Downloaded on May 30,2021 at 15:29:13 UTC from IEEE Xplore. Restrictions apply. IV. CONCLUSION A Quadcopter\u2019s propeller has been analyzed for the structural and vibrational aspects. The propeller has been analyzed under the thrust (case1) and rotational loading (case2) conditions to find out the stress and deformation produced in two materials. It is found that for case1 both the materials show almost same results, while for case2 GFRP shows the high values of maximum deformation and stress compared to CFRP. Moreover, from the results of the modal analysis, it is found that CFRP is having high values of frequencies compared to GFRP. From these observations, it can be concluded that CFRP perform better than GFRP material under static and vibration loading condition. Furthermore the natural frequency of propeller is validate with publish literature [10] in figure 19. REFERENCES [1] Ahmad, F., Kumar, P., Bhandari, A., & Patil, P. P. (2020). Simulation of the Quadcopter Dynamics with LQR based Control. Materials Today: Proceedings, 24, 326-332. [2] Ahmad, F., Kumar, P., & Patil, P. P. (2018). Modeling and simulation of a Quadcopter with altitude and attitude control. Nonlinear Studies, 25(2). [3] Wei, P., Yang, Z. and Wang, Q. (2015). The Design of Quadcopter Frame Based On Finite Element Analysis,In 3rd International Conference on Mechatronics, Robotics and Automation. Atlantis Press. [4] Jaouad, H., Vikram, P., Balasubramanian, E., & Surendar, G. (2020). Computational Fluid Dynamic Analysis of 63 Authorized licensed use limited to: Carleton University. Downloaded on May 30,2021 at 15:29:13 UTC from IEEE Xplore. Restrictions apply. Amphibious Vehicle. In Advances in Engineering Design and Simulation (pp. 303-313). Springer, Singapore. [5] Kumar, S., & Mishra, P. C. (2016). Finite element modeling for structural strength of quadcoptor type multi modevehicle. Aerospace Science and Technology, 53, 252- 266. [6] Ahmed, M. F., Zafar, M. N., & Mohanta, J. C. (2020, February). Modeling and Analysis of Quadcopter F450 Frame. In 2020 International Conference on Contemporary Computing and Applications (IC3A) (pp. 196-201). IEEE. [7] Martinetti, A., Margaryan, M., & van Dongen, L. (2018). Simulating mechanical stress on a micro Unmanned Aerial Vehicle (UAV) body frame for selecting maintenance actions. Procedia manufacturing, 16, 61-66. [8] Singh, R., Kumar, R., Mishra, A., & Agarwal, A. (2020). Structural Analysis of Quadcopter Frame. Materials Today: Proceedings, 22, 3320-3329. [9] Xiu, H., Xu, T., Jones, A. H., Wei, G., & Ren, L. (2017, December). A reconfigurable quadcopter with foldable rotor arms and a deployable carrier. In 2017 IEEE ROBIO (pp. 1412-1417) [10] Ahmad, F., Bhandari, A., Kumar, P., & Patil, P. P. (2019, November). Modeling and Mechanical Vibration characteristics analysis of a Quadcopter Propeller using FEA. In IOP Conference Series: Materials Science and Engineering (Vol. 577, No. 1, p. 012022). IOP Publishing. 64 Authorized licensed use limited to: Carleton University. Downloaded on May 30,2021 at 15:29:13 UTC from IEEE Xplore. Restrictions apply." + ] + }, + { + "image_filename": "designv11_71_0002830_978-94-007-6046-2_42-Figure4-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0002830_978-94-007-6046-2_42-Figure4-1.png", + "caption": "Fig. 4 Potential energy conserving orbit [8]. (a) Ideal 2D biped model with massless legs. (b) Linear dynamics under gravity compensation", + "texts": [ + " As we observed, it is common to use approximations and linearization to design a suitable gait and controller, even if we have precise equations of the dynamics. Many researchers also presented the usefulness of various model simplifications to analyze global dynamics of biped locomotion [1, 15, 17]. In the next section, we explain the derivation of a linear model for biped walking which was inspired from these pioneering works. The first version of the linear inverted pendulum mode was derived by revisiting the 2D massless leg model of Gubina et al. Considering the gravity and moment compensation for all state space of Fig. 4a, the foot force F1 and the hip torque T2 are determined as F1 D mg.r1 C r2 cos. 1 2// r1 cos 1 C r2 cos 2 (14) T2 D mgr1r2 sin. 1 2/ r1 cos 1 C r2 cos 2 : (15) Thanks to the whole state-space compensation of gravity and the related moment, these inputs can yield the horizontal and upright body motion as shown in Fig. 4b. The trajectory of the center of mass .xg; yg/ was named potential energy conserving orbit [8], and its dynamics was described as follows: Rxg D .g=yg/xg (16) yg D yh .constant/ (17) Since this dynamics was derived without any approximation, it is valid in all range of xg. For example, we can use its analytical solution as the reference trajectory for walking with a big stride: xg.t/ D xg.0/ cosh.t=Tc/ C Tc Pxg.0/ sinh.t=Tc/ (18) Tc p yh=g Another useful equation is a conservation of the orbital energy E defined as 1 2 Px2 g g 2yh x2 g D constant E: (19) This can be obtained by multiplying (16) by Pxg and integrating it", + " By substituting them into (20), we obtain, following dynamics, the linear inverted pendulum mode (LIPM) [6]: Rx D g yc x C 1 myc u1 (23) Unlike the potential energy conserving orbit whose body height must be constant, LIPM can handle change of body height by using the slope k and body rotation with constant rate !c , whereas it holds the same linear dynamics. Examples of LIPM trajectories without use of ankle torque (u1 D 0) are shown in Fig. 6. In these simulations, a telescopic leg model of Fig. 4a was used to visualize the results. At first glance, the body movement of Fig. 6a, b looks different where the first one changes its height by the slope constraint and the second one moves horizontally while rotating. Nevertheless, if we pay attention to their horizontal motions, they have identical acceleration-deceleration patterns governed by (23). The LIPM can be controlled by the input torque u1. Although the magnitude of u1 is strictly limited by the foot size, it is useful to compensate model errors and disturbance which must be concerned in the real walking robot control" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0001701_0954405420949757-Figure13-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0001701_0954405420949757-Figure13-1.png", + "caption": "Figure 13. Contact positions on cutter surface.", + "texts": [ + " When solving processes for concave side, both sides and convex side chamfering are completed, integral chamfering tool path for concave side, both sides and convex side is obtained, as shown in Figure 12. There are two discontinuities on borders of three phases, cutter moves up and then moves down in those positions to ensure processing correctly. Cutter axis direction calculating Cutter axis is set to be perpendicular to pinion axis Z1, therefore chamfering can be realized on four-axis CNC machine tools. Cutter center is denoted by point E and cutter axis direction angle is denoted by uc, as shown in Figure 13. When the position of point E keeps unchanged, variation of uc has no effect to chamfering shape, but contact position on cutter surface changes with uc, which relates to tool wear.21,22 In both sides chamfering phase, suppose that upper edge of convex side and cutter surface contact at point F1 which projects to cutter axis at point G1, upper edge of concave side and cutter surface contact at point F2 which projects to cutter axis at point G2, as shown in Figure 13. In order to process correctly, point G1 and G2 should locate on segment EH. Variation range of uc satisfies two constraints as follows 04EG14rc 04EG24rc \u00f036\u00de In concave side chamfering phase, contact point F1 does not exist in Figure 13, there is only one constraint 04EG24rc left. Similarly, in convex side chamfering phase, contact point F2 does not exist in Figure 13, there is only one constraint 04EG14rc left. The uc of k-th discrete point on integral chamfering tool path under the constraints has maximum value ukmax and minimum value ukmin.ukmax and ukmin are calculated and taken as ordinate values, while sequence number of discrete point on integral chamfering tool path is taken as abscissa value, upper and lower borders are formed by linking all of ukmax and ukmin respectively. Feasible region of uc is determined by upper and lower borders, as shown in Figure 14" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0000431_dynamics47113.2019.8944673-Figure2-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0000431_dynamics47113.2019.8944673-Figure2-1.png", + "caption": "Fig. 2. Phasor diagram showing dq and xy axes", + "texts": [ + " DTC CONCEPT Direct torque control (DTC) is based on motor equations written in a coordinate system that is tied to the air gap flux, which includes the magnet flux in the d-axis and the armature reaction fluxes SdSdiL and SqSqiL in the d- and q-axis. Classical theory mainly uses scalar equations for components along the d and q axes, as in (7). DTC are most often expressed using space vectors. These vectors are physical quantities that not only move in time, but also rotate in space [8]. To determine these space vectors, we write (7) as follows qd j . (14) The magnetic flux-linkage is a complex number that is constant in the steady state. The value of leave over in the fixed position in Fig. 2. Rewrite (14) in x, y axes yx j , (15) where the coordinate origin xy coincides with the coordinate origin dq, while the x-axis is rotated relative to the d-axis by an angle . If is chosen as Fig. 2 (where the x-axis collocates along the space vector ), then x ; 0y . (16) Fig. 2 also shows that SdSdf SqSq iL iL arctan . (17) The component parts of space vector x and y can be obtained from (7) by the reexpression q d y x cossin sincos . (18) From (18) can be obtained y x q d cossin sincos . (19) Similar to (15), the current is qdyx jiijiii , and then the torque (8) can be rewritten as follows yxyyx i mp ii mp M 2 )( 2 , (20) where is the amplitude x . This equation is identical in form to (9). However now is the total flux linkage, which can be adjusted" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0002209_icca51439.2020.9264350-Figure3-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0002209_icca51439.2020.9264350-Figure3-1.png", + "caption": "Fig. 3. Generalized PI model operator", + "texts": [ + " Researchers have proposed modified models to characterize asymmetric hysteresis, including the GPI model which uses nonlinear play operators. The two straight lines of the play operator p(t) in the CPI model are replaced by strictly increasing and continuous envelope functions \u03b3r and \u03b3l. Example candidates for these envelope functions include second- and third-order polynomials, exponential and hyperbolic-tangent functions [8], [12], [13], [14]. The nonlinear play operator used to create an asymmetric hysteresis loop can have the characteristic as shown in Fig. 3(a) and so (1) becomes the following [8]: p(t)=Sr[v](t)=max{\u03b3r(v)\u2212r,min[\u03b3l(v)+r, p(t\u2212T )]} p(0)=Sr[v](0)=max{\u03b3r(v(0))\u2212r,min[\u03b3l(v(0))+r, 0]} . (4) Fig. 3(a) and (4) show that the zero outputs p(t) of the GPI play operator occur at thresholds \u03c21 and \u03c22 of the increasing and decreasing input v(t). The thresholds are expressed as: \u03c21 = \u03b3\u22121 r (r); v\u0307(t) > 0 \u03c22 = \u03b3\u22121 l (\u2212r); v\u0307(t) < 0. (5) Then, r = \u03b3r(\u03c21) and \u2212r = \u03b3l(\u03c22). Note that GPI reduces to CPI when \u03b3r(v) = \u03b3l(v) = v. As with the CPI model, the GPI model can be applied to the battery application using the generalized one-sided play operator as shown in Fig. 3(b), where the operator output is: p(t)=Sor[v](t)=max{\u03b3r(v)\u2212r,min[\u03b3l(v), p(t\u2212T )]} p(0)=Sor[v](0)=max{\u03b3r(v(0))\u2212r,min[\u03b3l(v(0)), 0]} . (6) To model asymmetric hysteresis nonlinearities, the GPI model output can be implemented using a finite number of play operators and the output y(t) can be created by a weighted summation of the play operators associated with the input v(t) as illustrated in Fig. 4 and can be expressed as: y(t) = n\u2211 i=1 w(ri)Sori [v](t). (7) 1579 Authorized licensed use limited to: London School of Economics & Political Science" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0003880_scis-isis.2014.7044812-Figure6-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0003880_scis-isis.2014.7044812-Figure6-1.png", + "caption": "Fig. 6. Representation of the new repulsive force and total potential force", + "texts": [ + " This will support robot to move away from the local minima, and for a smooth and dead-lock free motion. We modified the motion planning algorithm to guide the robot properly towards the goal when there is an obstacle closer to it. When the robot detects the distance between the robot and goal less than the distance between the robot and the obstacle within the obstacle detecting range , robot 978-1-4799-5955-6/14/$31.00 \u00a92014 IEEE 261 considers attractive force only for its motion. In this research, we utilize the function described in (9). (9) Fig. 6 explains how the new total potential force has been changed from its original total potential force without the new repulsive force component . Furthermore, the magnitude of the new total potential force is bigger than its old value, hence there is a more possibility to push the robot away from the obstacle and make it easier to maintain a big clearance between the robot and the obstacles than before. IV. SIMULATION STUDY AND THE RESULTS OF PROPOSED The proposed modified APF based algorithm is mathematically modeled in MATLAB\u2122 simulation environment" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0001192_icemi46757.2019.9101435-Figure2-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0001192_icemi46757.2019.9101435-Figure2-1.png", + "caption": "Fig. 2 Structure of rolling bearing", + "texts": [ + " The principal component analysis technique with the ability about dimension reduction of high-dimensional vectors is used to solve the above problems in this paper, BP neural network with good classification performance is adopted to identify the faults of bearings, the existing diagnostic technique based on 1125 978-1-7281-0510-9/19/$31.00 \u00a92019 IEEE Authorized licensed use limited to: University of Exeter. Downloaded on June 22,2020 at 18:05:21 UTC from IEEE Xplore. Restrictions apply. wavelet packet energy features is used to compare with it. The diagnosis accuracy is improved by using the fault diagnosis technique based on principal component analysis from the experimental results, which can remove the redundant features. The research content of this paper is shown in Fig.1. Fig. 1 Research contents In Fig.2, the rolling bearing consists of inner race, outer race, ball and retainer [9].The outer race, which is usually fasten on the bearing base, does not turn round with the shaft. The inner race, which is connected with the drive shaft, rotates along with the rotation shaft. In addition to the relative sliding between internal and external grooves, there is also the rotation of the ball, which is an important part to convert sliding friction into rolling friction. The retainer is used to avoid friction or collision between adjacent rolling balls [10]" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0000356_iecon.2019.8927121-Figure4-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0000356_iecon.2019.8927121-Figure4-1.png", + "caption": "Fig. 4. Double winding induction RLM [12]", + "texts": [ + " If this is fulfilled, the rotary motion will not affect the generated magnetic field. But the linear speed of the machine will have an influence on the performance of the rotary movement, which must be taken into consideration [6], [8]. Such simplistic motor combination approach was proposed also by P. de Wit et al. Their motor had a very simple solid aluminum cylindrical mover [10]. Upon another approach the induction RLM can be built up also with a single stator stack having two sets of winding [11]. This constructional solution (given in Fig. 4) is more compact than that presented previously, but it is also more complicated. As it can be seen, it has an axial winding which generates a rotating circular magnetic field (1) and a ring-type winding which produces a traveling magnetic field (2), both inside a common stator iron core (5) and the housing (3). As in the case of almost all RLMs, the long cylindrical mover (4) can rotate, displace linearly and perform a compound helical movement. The rotor of this induction RLM, which has to interact with both of the magnetic fields, can be a solid, a cylinder with a twin iron and copper layer [11], or with crossed bars (as that shown in Fig. 4) [12]. K. Kluszczynski et al. presents in [12] the way as both above detailed induction RLM variants can be easily built up mainly by using components of on-shelf three-phase squirrel cage induction machines. In [13] T. Onuki et al. propose a multi stator induction RLM, composed of four commercial induction machine stators with a common mover. The control of the helical two DoF motion is assured by means of controlling the phase angle of the four voltages feeding the four sets of stator windings" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0000356_iecon.2019.8927121-Figure5-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0000356_iecon.2019.8927121-Figure5-1.png", + "caption": "Fig. 5. Series coupled PM RLM [14]", + "texts": [ + " The reason is that the PMs assure high efficiency and power density, and in a consequence compactness to the RLMs. This feature is very important since the main reason of applying such machines is the space reduction. Supplementary, PM RLMs usually have good servo characteristics, which makes them ideal for advanced industrial applications. One of the simplest series coupled PM RLM was proposed in [14]. Both parts of the cylindrical machine, those performing the linear and the rotational movements, are slotless PM actuators with concentrated windings having a common mover (see Fig. 5). Two optical encoders are integrated into the machine to measure the mover position on the two axes. This PM RLM has some advantages, as the lack of cogging due to the slotless structure. The fully axial-symmetric structure of the linear machine part enables an independency of the rotary motion. A basic drawback is its excessive length. A more compact PM RLM, with the two concentrically stacked actuators, was proposed by Jang et al. [15]. The mover comprises of two cylindrical Halbach arrays [16] of different patterns, separated by a cylindrical iron core piece (see Fig" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0002244_tnsre.2020.3042221-Figure3-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0002244_tnsre.2020.3042221-Figure3-1.png", + "caption": "Fig. 3. Experimental apparatus and setup. (a) Photo of the electrical stimulation containing surface electrodes; (b) The CAD graph of the experimental apparatus; (c) Schematic diagram of experimental posture. The ankle positions during experiments were shown in Fig. 3(c) with coordinate axis.", + "texts": [ + " (16) When the electrical stimulation current was delivered to the GAS and TA simultaneously, the total active stiffness Kest of ankle joint can be estimated as the sum of the active stiffness [30] induced by TA and GAS shown as Kest = Kest_T A + Kest_G AS (17) where Kest_T A and Kest_G AS are the active stiffness supplied by TA and GAS, respectively. Authorized licensed use limited to: Rutgers University. Downloaded on May 19,2021 at 04:49:23 UTC from IEEE Xplore. Restrictions apply. As shown in Fig. 3, the experimental apparatus mainly consisted of a direct drive brushless AC servo motor (DM1B045G, Yokogawa, Japan) with a servo driver (UB1DG3, Yokogawa, Japan), a motion controller (GTS-400-PV(G)PCI, Googoltech, Hong Kong), a torque sensor (AKC-205, 701st Research Institute of China Aerospace, Science and Technology Corporation, China), a mechanical footplate and supporter, a multi-channel functional electrical stimulator (P29632, Faisco, China), surface electrical stimulation electrodes (M2223, 3M, USA), a personal computer (PC), and a data acquisition card (USB-6341 DAQ card, National Instruments, USA) [25], [27]" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0003544_cta.2127-Figure14-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0003544_cta.2127-Figure14-1.png", + "caption": "Figure 14. Rotating position \u03b8 definition.", + "texts": [ + " It should be noticed that the measured mutual inductances of M31 or M32 in the different coil positions are obviously different. Thus, a question may be raised in here \u2013 can these measured parameters\u2019 values be validated? The answer is \u2018more or less\u2019. The single coil, for example, between 3C and 4C, 4C and 5C, or 5C and 6C is magnetically coupled by a different flux linkage depending on the angular \u03b8. Copyright \u00a9 2015 John Wiley & Sons, Ltd. Int. J. Circ. Theor. Appl. 2016; 44:1094\u20131111 DOI: 10.1002/cta Therefore, the flux-linkage associated mutual inductance to each single coil is given to cos \u03b8, as presented in Figure 14. As previously stated, each segment was 22.5\u00b0 apart on the armature for this 2-pole/16-slot machine. Hence, the mutual inductance M31 defined between 9C and 6C (three coils) and 5C and 6C (one coil) can be written as M31MAX\u00d7 cos 22.5\u00b0 with the current direction from 9C to 6C. Also, M31 of 9C\u20136C with 4C\u20135C and 3C\u20134C can be denoted as M31MAX\u00d7 cos 45\u00b0 and M31MAX\u00d7 cos 67.5\u00b0, respectively. Theoretically, these mutual inductances should be expected including the term of cos \u03b8 as M31MAX cos22:5\u00b0 M31MAX cos45\u00b0 \u00bc 0:924 0:707 \u00bc 1:306; M31MAX cos45\u00b0 M31MAX cos67:5\u00b0 \u00bc 0:707 0:383 \u00bc 1:846: Our measured value ofM31 for 5C\u20136C was 7" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0002540_iccr51572.2020.9344388-Figure2-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0002540_iccr51572.2020.9344388-Figure2-1.png", + "caption": "Figure 2. The humanoid robot in the study: (a) The simulation model; (b) The construction dimension; (c) The joint configuration and frames.", + "texts": [ + " Assuming that the offset \u0305 and landing location \u0305 of each cycle are constant, the walking period can be obtained as follows: = log \u0305 + \u0305\u0305 (18) It is noted that if the walking cycle of the robot is forced to be set to a constant value , the recursive relation of can be obtained: = \u2212 (19) where, = / is the scaling factor. Since > 1 , the sequence of will be unstable. So, extra control laws are needed to keep the robot walking on a desired cycle of time. The joint configuration of the simulation model is showed in the Fig. 2. The floating frame is set in the middle of the two \u210e 1 joints and is attached to the torso. The position and posture of the whole robot are described based on the floating coordinate and joint coordinate. Due to the walking study is carried out in the sagittal plane, the floating coordinate and actuated joint coordinate are respectively established as =[ , , ] (representing the x-axis position, z-axis position and rotation of the floating base within the sagittal plane as shown in Fig. 1) and = , , , , , , , , , , , , while the remaining joints are considered to be fixed" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0001717_icmae50897.2020.9178892-Figure5-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0001717_icmae50897.2020.9178892-Figure5-1.png", + "caption": "Figure 5. The principle of FDM technique.", + "texts": [ + " FDM (Fused Deposition modelling) or FFF (Fused Filament Fabrication) 3D printers is primarily used for ABS material processing. Plastics come in the form of a long filament wound around a spool. Employing a print head, a molten layer of plastic is deposited on the print bed, which then adheres. Once the first layer has been drawn, the print bed drops and a new layer is built on the previous layer. This is repeated several times, ultimately resulting in a 3D printed model. The principle of FDM technique is shown in Fig. 5. [13] Five samples of both types (porous and shell) were produced from the material ABSplus-P430 Ivory plastics using FDM technique for the standard tensile tests. The 3D printer Prusa i3 Mk2 was selected for the samples production. The melting temperature of the filament is 300 \u00b0C, while the nozzle temperature was 255 \u00b0C. The temperature of the basement was 100 \u00b0C. The thickness of the deposited layer was 0.254 mm given by the producer of the 3D printer device and the diameter of ABS filament was \u03d5 1" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0002521_ei250167.2020.9346917-Figure3-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0002521_ei250167.2020.9346917-Figure3-1.png", + "caption": "Fig. 3 Magnetic flux density distribution of the HS-PMSM", + "texts": [ + " Next, the machine\u2019s characteristics such as magnetic flux density distribution, magnetic flux line distribution, permanent magnet flux linkage, no load back electromotive force, electromagnetic torque, and loss are analyzed by finite element method and discussed. 2735 Authorized licensed use limited to: Miami University Libraries. Downloaded on June 15,2021 at 08:41:38 UTC from IEEE Xplore. Restrictions apply. The major parameters of HS-PMSM are listed in Table 1. While the distribution of magnetic flux density under load condition is shown in Figure 3, and the distribution of magnetic flux line under the same condition is shown in Figure 4. Figure 5 and 6 show the radial magnetic flux density in the machine gap. As shown in Figure 5, the red dot line and the yellow dash line represent for the flux density generated by the permanent magnets under no armature current condition and the flux density generated by the armature current under no permanent magnets exciting condition, respectively. While the blue solid line represent for the flux density generated by the armature current and permanent magnet together" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0000925_9781119477891.ch8-Figure8.16-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0000925_9781119477891.ch8-Figure8.16-1.png", + "caption": "Figure 8.16 Three\u2010layer blown film co\u2010extrusion die.", + "texts": [ + " The setup as it is shown here, is a mono\u2010extrusion instal lation. This means that we have one extruder that will feed one material or one blend of materials to the die, generating a mono\u2010layer film. Depending on the final film structure that we want, a co\u2010 extrusion installation can be used. In a co\u2010extrusion instal lation, a multi\u2010layer film can be produced by having multiple extruders feeding molten material to the die. In the die, these materials are put in layers on top of each other, before coming out of the die (see Figure\u00a08.16). The remaining of the installation is identical to mono\u2010extrusion. Today co\u2010extrusion blown film installations are avail able for multi\u2010layer films of 3, 5, 7, 9, or even more layers. We have to keep in mind that the different extruders have to be adapted to the type of material and the thick ness we want to have for each layer. We will need for instance a different size of extruder for a tie\u2010layer with a thickness of 5 \u03bcm versus a contact layer with a thickness of 100 \u03bcm. But also, different types of material require different types of screws" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0003921_coase.2014.6899477-Figure10-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0003921_coase.2014.6899477-Figure10-1.png", + "caption": "Figure 10. Velocity-based detection mechanism.", + "texts": [ + " (a)Robot (b)Velocity and Contact Force-based Safety Device (c) Details Contact Force Detection Plate Switch-off and Shaft-lock Mechanism Shaft A Velocity-based Detection Mechanism Linear Spring Y Contact Force-based Detection Mechanism Wire Rope Pulley Linear Spring Z Claw D Wheel Caster Claw D Wire Rope To Contact Force Detection Plate Shaft A Switch-off and Shaft-lock Mechanism Velocity-based Detection Mechanism Motor Wheel Linear Spring Z Velocity and Contact Force-based Safety Device Timing Belt (d)Three Claws B. shown in Fig.9. If the acceleration of Contact Force Detection Plate is so small that we can neglect the inertial force, we obtain the following motion equation: fF k y\u2206= , (1) where F is the detection contact force level, kf is the spring constant of Linear Springs Y, y\u2206 is the displacement from the natural length of Linear Springs Y. We can approximately set the detection contact force level on the basis of (1). 2) Velocity-based Detection Mechanism Fig. 10 shows the mechanism which mechanically detects the unexpected angular velocity of Shaft A. The damping torque by Rotary Damper and the spring torque by Linear Spring B act on Claw A, when Gear B is rotated by Gear A. As the velocity of Gear A (i.e. Shaft A) increases, the damping torque increases. Claw A rotates by the torque difference between the damping torque and the spring torque, and locks Plate A, if the velocity of Shaft A exceeds the detection velocity level. The detection velocity level is adjustable by changing the attachment position of Linear Spring B as shown in Fig. 10. If the mass and moment of inertia of Claw A are so small that we can neglect the inertial torque and the gravitational torque, we can obtain the following motion equation of Claw A: c s k r xv \u2206\u03c9 = , (2) where c is the damping coefficient of Rotary Damper, \u03c9 is the detection velocity level, s is the gear ratio of Gear A to Gear B, kv is the spring constant of Linear Spring B, r is the distance between the shaft axis of Rotary Damper and the attachment position of Linear Spring B, and x\u2206 is the displacement from the natural length of Linear Spring B" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0000355_ritapp.2019.8932800-Figure3-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0000355_ritapp.2019.8932800-Figure3-1.png", + "caption": "Figure 3. Quadrotors Used in the Study", + "texts": [ + " In this way, the pitch of the quadrotor is calculated at the desired target coordinate. \ud835\udc43\ud835\udc3c\ud835\udc47\ud835\udc36\ud835\udc3b = \ud835\udc43\ud835\udc43\ud835\udc43\ud835\udc52\ud835\udc5f\ud835\udc5f\ud835\udc5c\ud835\udc5f + \ud835\udc43\ud835\udc59\ud835\udc43\ud835\udc4e\ud835\udc50\ud835\udc50\ud835\udc62\ud835\udc5a\ud835\udc62\ud835\udc59\ud835\udc4e\ud835\udc61\ud835\udc52\ud835\udc51 + \ud835\udc43\ud835\udc37\ud835\udc43\ud835\udc60\ud835\udc5d\ud835\udc52\ud835\udc52\ud835\udc51 (8) Equation (9) shows how the roll of the quadrotor is computed. \ud835\udc4c\ud835\udc43, \ud835\udc4c\ud835\udc3c, \ud835\udc4c\ud835\udc37 are all PID constants that were retrieved in order to get a stable flight. \ud835\udc4c\ud835\udc52\ud835\udc5f\ud835\udc5f\ud835\udc5c\ud835\udc5f is the distance between the target and current coordinate while \ud835\udc4c\ud835\udc4e\ud835\udc50\ud835\udc50\ud835\udc62\ud835\udc5a\ud835\udc62\ud835\udc59\ud835\udc4e\ud835\udc61\ud835\udc52\ud835\udc51 is the sum of all \ud835\udc4c\ud835\udc52\ud835\udc5f\ud835\udc5f\ud835\udc5c\ud835\udc5f. \ud835\udc4c\ud835\udc60\ud835\udc5d\ud835\udc52\ud835\udc52\ud835\udc51 on the other hand is the speed of the quadrotor at the y-axis. In this way, the roll of the quadrotor is calculated at the desired target coordinate. Figure 3 shows the setup of the Crazyflie 2.0 drone used in this study. Reflective markers are attached to the drone for the motion capture system. The controllers can use a Bluetooth LE that allows mobile devices as a means for the main controller. Since the quadrotor is lightweight, its maximum payload is only 15 grams. This research is implemented in a game platform. This is also used as the experiment setup. The researchers developed the rules in the said games. a) Game of Drones is a non-contact type of game" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0000227_2050-7038.12261-Figure9-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0000227_2050-7038.12261-Figure9-1.png", + "caption": "FIGURE 9 A, Flux distribution with phase winding current. B, Flux distribution with the excitation winding current", + "texts": [ + " The excitation winding is placed at the tip of the stator poles of all the phases, and the sense winding is placed next to the excitation winding as shown in Figure 7A, and also the enlarged version of the SRM is shown in Figure 7B for quarter half of the stator periphery. The excitation winding is supplied with a sinusoidal voltage source of 5\u2010V peak and 30\u2010 kHz frequency. The current in the excitation winding is 12 mA as shown in Figure 8 and is equally distributed across all the stator poles of the machine. The total torque ripple caused by the excitation winding is 0.1% of the rated motor torque. The flux distribution in the machine due to the phase winding and the excitation winding is independently shown in Figure 9A,B, respectively. The sense voltage of phases A and C as well as the inductance derived from the sense voltages is shown in Figure 3. The actual inductance and the estimated inductance of phases B and D are shown in Figure 10. The effectiveness of inductance estimation from dynamic LUT at constant speed is demonstrated in Figure 10. The estimated inductance is very close to the actual inductance at constant speed. The inductance, the rate of change of inductance normalized with respect to speed, and the current reference of phase A are shown in Figure 11" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0003267_icra.2015.7139418-Figure1-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0003267_icra.2015.7139418-Figure1-1.png", + "caption": "Fig. 1. Quadrotor reference frames: rotor speeds provide moments on the vehicle dynamics.", + "texts": [ + " The state vector ~x \u2208 R6 is composed by positions and orientations variables of the vehicle, and their respective derivatives, in the Lie Group SE(3), such that: ~x = [ ~\u03c8 ~\u03c9 ] , where ~\u03c8 = [\u03c6, \u03b8, \u03c8] T , in rad (radians), is the Tait-Bryan orientation vector in SO(3), related to the world reference frame {W } \u2208 R3, and ~\u03c9 = [p, q, r] T , in rad/s, is the angular speed vector measured on the body reference frame of the vehicle {B}. The vehicle input vector is ~u = (\u21262)2 \u2212 (\u21264)2 (\u21263)2 \u2212 (\u21261)2 4\u2211 i=1 ( 1)i+1(\u2126i) 2 , where \u2126i represents the rotation speed (in rpm) applied to the i-th rotor. Rotor dynamics is neglected, since its acceleration is much larger than variations on the states of the vehicle. Figure 1 illustrates forces and moments acting on the vehicle. The time variation of the angular motion of the quadrotor can be described as: ~\u0307\u03c8 = H~\u03c9, (1) I ~\u0307\u03c9 + 1 2n\u03c1 ~\u03c9 T |~\u03c9|+ (~\u03c9\u00d7I~\u03c9) = l\u03b6~u, (2) where I stands for the inertial matrix of the flying robot, diagonal due to its symmetry; l is the distance between each rotor and the center of mass of the vehicle; \u03b6 is a positive scalar gain of the motor; n \u2248 3.3\u00d710 4 kg/m is the mean drag coefficient and \u03c1 \u2248 1.22 kg/m3 is the air density according to the International Standard Atmosphere (ISA)" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0001701_0954405420949757-Figure15-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0001701_0954405420949757-Figure15-1.png", + "caption": "Figure 15. Coordinate systems of pinion and four-axis CNC machine tools.", + "texts": [ + " All of u0k are linked to form a dash line in Figure 14. After All of u0k are obtained, unit vector of cutter axis is calculated as ( cos u0k, sin u0k, 0) in pinion coordinate system, which is denoted by (nx, ny, nz). Axial cutter position detecting Coordinate system of four-axis CNC machine tools is denoted by s= \u00bdO;X,Y, Z , pinion coordinate system is denoted by s1 = \u00bdO1;X1,Y1, Z1 . Origin O coincides with origin O1, Z axis and X1 axis have the same direction, X axis and Z1 axis have the opposite direction, as shown in Figure 15. Suppose that when the values of X axis and Z axis change from 0 to a positive value, X and Z axis sliding tables move along positive X and Z axis respectively. When the value of Y axis changes from 0 to a positive value, Y axis sliding table moves along negative Y axis. When axial cutter position detecting is performed, Y axis sliding table moves to make cutter axis and A axis intersect, then the value of Y axis is set to 0 in CNC system. X and Z axis sliding tables move to make edge detector and inner end of pinion contact, the distance between inner end and point O1 is denoted by d in pinion coordinate system shown in Figure 2, then the value of X axis is set to rc d" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0001877_icarm49381.2020.9195392-Figure8-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0001877_icarm49381.2020.9195392-Figure8-1.png", + "caption": "Fig. 8. Illustration of the maximum deformation rate measurement setup, VT-1 undergoing a large deformation (>97.6%) without fracture.", + "texts": [ + " Once again the prototyped wings are fixed to the edge of the micromanipulator clamp with the aerodynamic forces represented by a pin placed at 70% of the wingspan and chord for the estimation of maximum deformation rate spanwise and chordwise, respectively. The rate of maximum deformation differs from flexural stiffness in terms of wing displacement (\u0394) which is raised beyond 5% up until the applied force exceeds yield strength causing the test subject to undergo an evident deformation (bending). However, in contrast to the brittle carbon fibre wings introduced in the literature [37], the artificial PVC wings (VT-1,2) here offer a much larger plastic deformation range shown in Figure 8, i.e. the wings 978-1-7281-6479-3/20/$31.00 \u00a92020 IEEE 606 Authorized licensed use limited to: University of Brighton. Downloaded on October 22,2020 at 17:55:23 UTC from IEEE Xplore. Restrictions apply. do not break immediately after yielding. Therefore, VT-1 did not undergo a fracture even at maximum wing displacement (\u0394max=100mm) recording a deformation rate of >97.6% that is almost double the rate (57.9%) obtained on an artificial cicada wing [37]. This is particularly favourable in aerial applications due to the high demand for wing flexibility that could prevent irreversible fracture failures" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0002264_s12221-020-1017-z-Figure1-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0002264_s12221-020-1017-z-Figure1-1.png", + "caption": "Figure 1. (a) and (b) show a photograph and schematic diagram of a three channel CNC ring spinning machine, respectively; 1-3: slivers, 4- 6: back rollers, 7: condenser, 8: middle roller, 9: rubbing apron, 10: front roller. Vf and Vm denote the feeding speeds of the front and middle rollers, respectively, and Vb1, Vb2 and Vb3 denote the feeding speeds of the three back rollers. (c) illustrates the formation of a three component spun yarn using three channel CNC ring spinning.", + "texts": [ + " By tuning the feeding speeds from the back rollers, the compositions and structures in the axial direction of yarn can be modified. Furthermore, we conducted a quantitative analysis concerning the effects of the manufacturing parameters, such as the blending ratio, time period, spindle speed and twist coefficient, on the evenness and breaking strength of the yarn. At last, we used the orthogonal experimental design to optimize the manufacturing parameters. Our work paves the way for the development of manufacturing methods for the large-scale production of multifunctional textiles. Figure 1(a) shows a photograph of the developed CNC ring spinning frame. The CNC system consists of a touch screen and a programmable logic controller (PLC). The spinning system consists of multi-channel drafting, twisting and winding units, as demonstrated by Figure 1(b) and (c). The multi-channel drafting unit mainly consists of three coaxially coupled variable speed back rollers (denoted as #4\u2013#6 in Figure 1(b)). The three back rollers are mounted on three shafts that are driven by three independent servo motors (denoted as #1-#3 in Figure 2). The system response time is 0.001 s. Three fiber slivers with different colors, i.e., magenta, yellow and green are fed into the back rollers, respectively. The three slivers are subject to the same drafting ratio in the main draft zone, whereas they are subjected to different drafting ratios in the break draft zone. Since the speeds of middle and front rollers are fixed, the total draft ratio for each sliver is relevant only to the feeding speed of the corresponding back roller controlled via PLC" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0001895_1350650120964295-Figure4-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0001895_1350650120964295-Figure4-1.png", + "caption": "Figure 4. Two-dimensional sketch of a sphere-sphere contact.", + "texts": [ + " The problem of contact detection can be mathematically described as follows: Let B1 and B2 be the contacting geometrical bodies and S1 \u00bc @B1 and S2 \u00bc @B2 the respective surfaces. Si is defined by the implicit function Fi. The distance d 1Corporate Research, Robert Bosch GmbH Renningen, Renningen, Germany 2Department of Microsystems Engineering, University of Freiburg, Freiburg, Germany Corresponding author: Marius Wolf, Corporate Research, Robert Bosch GmbH Renningen, Robert-Bosch-Campus 1, 71272 Renningen, Germany. Email: marius.wolf2@de.bosch.com between the contact points P1 2 S1 and P2 2 S2 is a local extremum. If B1 \\ B2 6\u00bc 1, the bodies are in contact (e.g. in Figure 4) and the distance between P1 and S2 and between P2 and S1 must be a maximum for P1 2 B2 ^ P2 2 B1. If B1 \\ B2 \u00bc 1, the bodies are separated (e.g. in Figure 5) and P1 and P2 must be chosen such that d equals the global minimum. If d is defined positive for B1 \\ B2 6\u00bc 1 and negative for B1 \\ B2 \u00bc 1, it equals the theoretical maximal penetration of B1 and B2. For the force application point follows Pm \u00bc P1 \u00fe P2 2 (1) State of the art contact detection models An approach for universal contact detection commonly used in commercial MBS programs is based on discretisation of the contacting bodies\u2019 surface into a large amount of polygons and checking interference for each polygon", + " The notation rSP;w \u00bc xSP;wex;w \u00fe rSP;wer;w is used for the location of the sphere\u2019s centre with respect to Jw. er;w is set, so rSP;w 0. This approach is beneficial, because P1 and P2 always lay in the plane defined by rSP and the rotational axis of the second body. Therefore, the problem is reduced to a two-dimensional problem in the working coordinate system. Afterwards, the results must be transformed back into the global coordinate system Jg. In case of a sphere-sphere contact, the connecting vector between PSP1 and PSP2 is collinear with the connecting vector of the spheres\u2019 centre (see Figure 4). Therefore, the contact properties can be calculated via PSP1 \u00bc rSP1 \u00fe cSP SP rSP2 rSP1 krSP2 rSP1kRSP1 (2) PSP2 \u00bc rSP2 cSP SP rSP2 rSP1 krSP2 rSP1kRSP2 (3) d \u00bc RSP1 \u00fe RSP2 cSP SPkrSP2 rSP1k (4) with cSP SP \u00bc 1 if one sphere is concave \u00f0RSPi < 0\u00de 1 else (5) The outer normal n of the cone (see Figure 5) can be expresses as n \u00bc cSP COsinhCO ex;w \u00fe cSP COcoshCO er;w (6) with cSP CO \u00bc 1 if cone is concave 1 else (7) The contact points PCO and PSP are calculated by projecting the sphere\u2019s centre onto the cone along n (see Figure 5)" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0000556_robio49542.2019.8961839-Figure1-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0000556_robio49542.2019.8961839-Figure1-1.png", + "caption": "Fig. 1. Automated Roll-to-Roll fluidic self-assembly process. (A) Overview illustrating component dispensing, fluidic self-assembly based on surface tension, and recycling of excess components. (B) Principle of jet pump. (C) Illustration of surface-tension driven self-assembly using molten-solderbased-receptor.", + "texts": [ + " In addition, we use a novel lamination approach in the realization of flexible solid state lighting modules incorporating distributed inorganic light emitting diodes (LEDs) using the assembly system. Since this application requires not only self-assembly, but also the formation of multilayer interconnects, the lamination approach is crucial. 1057-7157 \u00a9 2015 IEEE. Translations and content mining are permitted for academic research only. Personal use is also permitted, but republication/redistribution requires IEEE permission. See http://www.ieee.org/publications_standards/publications/rights/index.html for more information. Fig. 1 illustrates the layout of automated roll-to-roll fluidic self-assembly system. The system contains two units: (i) roll-to-roll assembly unit (shaded in gray) involving motor, rollers, customized agitator, and polyimide web to regulate operation parameters such as web moving speed, web angle, agitation frequency and amplitude. (ii) component recycling and dispensing unit (shaded in blue) containing diaphragm pump, jet pump and dispensing nozzle to reintroduce and gently dispense unassembled components", + " The system operation includes four steps: (i) transporting the components to the assembly unit; a jet pump delivers the unassembled components upward into a narrow fluid channel (5 mm inner diameter tubing). Originally we installed a mechanical pump between the bottom of the chamber and the dispensing head to circulate the components, however, this lead to mechanical damage to the components and the pump. To prevent this damage, we use an indirect circulation approach coupling a mechanical pump (QV variable speed pump, Fluid metering, Inc., NY) with the customized jet pump as shown in Fig. 1(B). The jet pump requires a smaller diameter nozzle (1.48 mm2) to accelerate the carrier fluid into a desired direction. Typical velocity of carrier fluid in the upward fluid channel (20 mm2 tubing) is 25 cm/s to lift up the dense silicon components (density of 2.33 g/cm) to the dispensing head. (ii) dispensing the transported components on the substrate; the components are dispensed on top of the substrate by gravity. Gentle introduction of components to the receptor is important to reduce the effect of liquid flow which induces additional drag. The dispensing head is located 5 cm away from the polyimide web (50 \u03bcm thick, 5 cm wide) pointing not directly at the web. Since the components leave the narrow dispensing head and meet large volume of liquid, the velocity decreases and the components fall downward following a vertical path to the web by gravity. (iii) assembling the components on the substrate; the dispensed components are self-assembled on the advancing substrate. Fig. 1(C) shows the detailed attachment process. The substrates contain multiple pre-defined solder-coated receptors which are formed using a customized dip-coating of a low melting point solder on copper clad polyimide (50 \u03bcm thick, Pyralux LF series, DuPont, NC) films patterned using a microfabrication technique. Procedure to fabricate the substrate is described in the experimental section. The self-assembly process is accomplished in water at 80 \u00b0C where the solder (Indalloy #117, MP. 47 \u00b0C, Indium Corp" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0001564_lra.2020.3013894-Figure5-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0001564_lra.2020.3013894-Figure5-1.png", + "caption": "Fig. 5. Diagram of the DAM-based three-revolute joint manipulator.", + "texts": [ + ", theoretical bounds of the SGR), the DAM can cover the actuating area under the black dotted line, which can be expressed as follows: \u03c9 = \u03c90\u03c4s 4 1 \u03c4 , (6) where \u03c90 and \u03c4 s are the no load speed and the stall torque of the motor, respectively. In summary, through using the proposed variable gearing, the DAM can produce greater velocity and force than the JAM which is equipped with a power-equivalent motor. To show the effectiveness and potential of the proposed variable gearing of the DAM, a three-revolute joint planar manipulator was selected. Detailed model description and notation are shown in Fig. 5. Because the proposed variable gearing can increase either speed or force, it can be considered as a bi-objective optimization problem of maximizing velocity and force at the same time. The general bi-objective optimization formulation can be expressed Authorized licensed use limited to: Cornell University Library. Downloaded on August 16,2020 at 15:34:17 UTC from IEEE Xplore. Restrictions apply. as Find x1, x2, \u00b7 \u00b7 \u00b7 , xn Tominimize \u03b1f1(x) + (1\u2212 \u03b1)f2(x) subject to gj(x) \u2264 0 j = 1, 2, \u00b7 \u00b7 \u00b7 , J hk(x) = 0 k = 1, 2, \u00b7 \u00b7 \u00b7 ,K x (L) i \u2264 xi \u2264 x (L) i i = 1, 2, \u00b7 \u00b7 \u00b7 , n, (7) where \u03b1 is a weight coefficient (0 \u2264 \u03b1 \u2264 1); the superscripts (l) and (u) denote the lower and upper bounds of the design variables; f1 and f2 are the objective functions to be optimized simultaneously (force and speed, respectively, in this study); and gj and hk are the inequality and equality constraint functions, respectively" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0002861_978-3-030-55889-5-Figure2.24-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0002861_978-3-030-55889-5-Figure2.24-1.png", + "caption": "Fig. 2.24 Influence functions for a cantilever beam and an angled two-span beam", + "texts": [ + "5/ f , and all other points, which have the horizontal distance /2 from the main pole, swing by v = /2 \u00b7 tan \u03d5 upwards, and so H = P \u00b7 v = P \u00b7 l/(4 \u00b7 f ) = M/ f . In the case of a uniform load p this results in H = 2 \u00b7 p \u222b /2 0 x \u00b7 tan \u03d5 dx = p \u00b7 2 8 \u00b7 f = M f . (2.63) One concluding remark on cantilever beams or, more generally speaking, attached structural parts. Why is it that only the right side, the free end, but not the left part moves, when the spread \u0394w = 1 or tan \u03d5r = \u22121 is initiated, as in Fig. 2.24 a, b? To see also a displacement on the left would require the action of a force, actually two, \u00b1 f , at both sides of the joint, but then the right part would enter into a \u201cspin\u201d, because it is not supported. So, if one side is kinematic, the other side will not move. In the case of the angled two-span beam on \u201cball bearings\u201d in Fig. 2.24c, the bearings do not hinder axial rotations and so the right part starts to rotate freely, when the joint (Green\u2019s function for M on the left side of the support) is spread, Fig. 2.24 d, and the absence of forces means absence of displacements in the first span, the Green\u2019s function is zero. The same applies in reverse order to the moment on the right of the support. The two-span beam carries any load like a series of two single-span hinged beams. If the beams form an angle of 90\u25e6, it is obvious, but it also true for any other angle = 180\u25e6 in between. At 180\u25e6 the beam becomes unstable, since the beam will rotate about its longitudinal axis, and thus will shed off any load. If the structure is statically indeterminate, forces are required to spread the hinge, so system # 2 is not load free, but even in this situation the work W2,1 = 0 is zero, and Betti\u2019s theorem reduces as before to B (w1, w2) = W1,2 = 0 " + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0003974_wcica.2014.7053572-Figure1-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0003974_wcica.2014.7053572-Figure1-1.png", + "caption": "Figure 1. the model of bevel-up needle Then we can control the needle by change the peak. The coordinate transformation matr coordinate system of A and B can be describe", + "texts": [ + " The particle filter requires a lot of particles to approximate the probability density, which also led to the biggest problem of particle filter: a large amount of computation. In the second part of this paper, we describe a mathematical model of the flexible needle and the estimation problem. The third section describes the details and steps of the three filtering methods. The fourth part is simulation and comparison of the flexible needle estimate. This part establish system model of bevel-up flexible needle. Figure 1 shows the system model. The base of the probe is controlled by a stepper motor, which can rotate and 978-1-4799-5825-2/14/$31.00 \u00a92014 IEEE 5051 move forward. For the tip of the needle is bev a sideway extrusion on the needle, which mak accurate motion. 0 1 R p g \u23a1 \u23a4 = \u23a2 \u23a5 \u23a3 \u23a6 Wherein, p is the position vector of system B relative to A. Here 2 0l = , w coincides with the end of the needle. We use v and \u03c9 to describe the lin velocities of the end of the needle. v a three-dimensional vectors. Let TV v\u23a1= \u23a3 established the following relationship: ( )1 1 1 2 2V g g V u V u \u2228\u2212= = + [ ]1 0 0 1 0 0 TV \u03ba= [ ]2 0 0 0 0 0 1 TV = Where, 1u is the speed of the probe alon the angular velocity of rotation of the probe" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0002128_j.mechmachtheory.2020.104171-Figure5-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0002128_j.mechmachtheory.2020.104171-Figure5-1.png", + "caption": "Fig. 5. Schematic representation of surface roughness between shaft and RSSR using the Dupont et al. model. Adapted from Ref. [10] .", + "texts": [ + " [27] developed a model to describe the friction force according to this pre-sliding domain. They also prescribed a transition coefficient in order to relate the elastic relative motion ( z ) to total relative motion ( x ). Zimmermann et al. [10 , 35] have also reported valuable experimental tests that confirm the robustness of relations for the pre-sliding domain strongly. According to the Dupont et al. friction model, the contact roughness can be represented by springs and dampers, as shown in Fig. 5 . The Dupont et al. friction model can be expressed by the following equation, wherein ( . ) denotes time derivative. \u23a7 \u23aa \u23aa \u23a8 \u23aa \u23aa \u23a9 F f = \u03c30 z + \u03c31 \u0307 z + \u03c32 \u0307 x z ss ( \u0307 x) = F f \u03c30 = sig ( z ) \u03c30 ( F C + ( F S \u2212 F C ) e \u2212( \u02d9 x v s ) 2 ) \u02d9 z = \u02d9 x ( 1 \u2212 \u03b1( z, \u02d9 x) z z ss ( \u0307 x) ) (4) \u03b1( z, \u02d9 x) denotes the transition from purely elastic ( \u03b1 = 0 ) to purely plastic ( \u03b1 = 1 ) motion given by the following relation. \u03b1( z, \u02d9 x) = \u23a7 \u23aa \u23aa \u23aa \u23a8 \u23aa \u23aa \u23aa \u23a9 0 f or | z | \u2264 z ba and sgn ( \u0307 x) = sgn ( z ) \u03b1m ( " + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0000407_6.2020-1184-Figure6-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0000407_6.2020-1184-Figure6-1.png", + "caption": "Fig. 6 Coordinate system and curvature definition for inextensional deformation of a shell, after [21].", + "texts": [ + " (21) Finally, the propagation moment in opposite sense bending is given by: M\u2217+ = D11w f lat r\u2217x + 2 ( D12R r\u2217x + D22 ) (22) C. Bistability Bistability of various thin-walled tape geometries can be investigated using the analysis in [21, 22]. The present case of a partially flat tape spring is one side of an X boom described in [22]. Thus, the non-dimensionalised changes in curvature for the flat and arc regions are: \u2206\u03ba\u0302arc = C\u0302 2 1 \u2212 cos (2\u03b8) cos (2\u03b8) + 1 \u2212 2 C\u0302 2 sin (2\u03b8) \u2206\u03ba\u0302 f lat = C\u0302 2 1 \u2212 cos (2\u03b8) 0 2 sin (2\u03b8) (23) where C\u0302 is the non-dimensionalised imposed curvature and \u03b8 is defined as per Figure 6. Non-dimensionalised variables are denoted by hats. The non-dimensionalised strain energy is given by: U\u0302 = 1 2w ( (w \u2212 w f lat )\u2206\u03ba\u0302 T arcD\u0302\u2206\u03ba\u0302arc + w f lat (\u2206\u03ba\u0302 T flatD\u0302\u2206\u03ba\u0302flat) ) (24) where the non-dimensional terms are: U\u0302 = UR2 wD11 , D\u0302 = D D11 , C\u0302 = CR. Equation 24 produces an energy landscape in which stable geometries are found at local minima. The stable geometries can instead be found analytically at points where: \u2202U\u0302 \u2202\u03b8 = 0, \u2202C\u0302 \u2202\u03b8 = 0. (25) When there is no coupling between bending and twisting D16 = D26 = 0 Equation 25 can be solved to show a stable geometry exists at: \u03b8 = \u03c0/2, C\u0302 = D\u030212(w \u2212 w f lat ) wD\u030211 (26) The geometry described in Equation 26 is stable provided: \u2202U\u03022/\u2202\u03b82 > 0 (27) where \u2202U\u03022/\u2202\u03b82 is given in Appendix A" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0002816_thc-2010-0566-Figure2-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0002816_thc-2010-0566-Figure2-1.png", + "caption": "Fig. 2. Normal stresses.", + "texts": [ + " Unit of \u03c3 is N/m2, also called Pascal (Pa) The left cylinder is experiencing a stress of mg/\u03c0r2 whereas the right cylinder experiences half that much (mg/2\u03c0r2), therefore we can now say that the left cylinder will fail first. The situation we analyzed is called uniaxial state of stress because the cylinders are experiencing stress along one axis and the stress is called normal stress because the force is acting perpendicular to the 0928-7329/10/$27.50 2010 \u2013 IOS Press and the authors. All rights reserved cross-sectional area (Fig. 2A). Normal stress always tries to either pull, called tension, or push, called compression, a body (Fig. 2B). Material inside the cylinders as shown in Fig. 1 responds to this tensile stress and deforms by stretching itself resulting in increase in the length of the cylinder at the cost of a decrease in the diameter. This deformation is expressed in terms of normal strain, denoted by \u03b5 (epsilon). Strain is a concept where the deformation due to the applied stress is normalized by the initial dimensions of the body. \u03b5 = \u2206L L (2) where L is the original length and \u2206L is the increase in length (Fig. 3A)" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0000907_b978-0-12-813372-9.00006-3-Figure6.16-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0000907_b978-0-12-813372-9.00006-3-Figure6.16-1.png", + "caption": "FIGURE 6.16 A mechanical oscillator based on reflex reversal. Source: Recreated from McMahon, T. (1984). Muscles, reflexes, and locomotion (Princeton paperbacks). Princeton, NJ: Princeton University Press.", + "texts": [ + "15 The principle of reciprocal inhibition: (left) cocontraction and (right) inhibition when a force is applied to the wrist. Source: Recreated from McMahon, T. (1984). Muscles, reflexes, and locomotion (Princeton paperbacks). Princeton, NJ: Princeton University Press. McMahon (1984) modeled this reflex reversal mechanism as a damped pendulum that produces self-sustained oscillations. Without some injection of energy (e.g., a stimulus), the oscillations would slow and stop, but with stimulation, the oscillations are able to continue indefinitely. In the pendulum system depicted in Fig. 6.16, the \u201cstimulus\u201d takes the form of a gas jet released when the pendulum opens an electrical contact meant to simulate a \u201creceptor\u201d in the nervous system. The stimulus is only effective if the pendulum is already in motion because the stimulus is not strong enough to start the pendulum from rest. This pendulum mimics the reflex reversal mechanism seen in each limb, although it cannot account for any interlimb coordination (McMahon, 1984). Recently, mechanical oscillators have been used to improve stability in quasipassive robots", + " Mechanical oscillators can be used to model the complex dynamics of gait. Understanding interlimb coordination requires the concept of entrainment. Dutch astronomer Christian Huygens observed that when two pendulum clocks that have slightly different periods are placed on the same wall, they synchronize, assuming the wall is compliant enough to translate mechanical energy between the clocks (Willms, Kitanov, & Langford, 2017). McMahon (1984) illustrated the process of entrainment by analogy to a clock mechanism similar to that depicted in Fig. 6.16. Imagine that an external force shakes the entire enclosure left and right in a sinusoidal fashion. The external forcing causes the pendulum to swing with a period equivalent to the external force, even if the eigenfrequency of the pendulum is different from that of the external force. If the amplitude of the force is great enough, the pendulum will repeatedly open and close valves B and C. If valve A is also open\u2014 the conditions which would produce a self-sustained oscillation\u2014then, the amplitude increases but the frequency does not", + " Placing Huygen\u2019s clocks farther apart on the wall weakens the coupling between them, and, at some critical distance, the coupling becomes effectively zero and the clocks will swing at their own natural frequencies. Second, the coupled oscillators need to have relatively similar natural frequencies. If starting from a state of entrainment, and one of the pendula is shortened with respect to the other, then at some point the clocks will once again tick at their own natural frequencies. Lastly, entrainment depends on the strength of internal forces of the clock (e.g., gas pressure in Fig. 6.16) and the strength of the external forcing function (Willms et al., 2017). The range of relative frequencies the sinusoidal force may take, and still result in entrainment, is the zone of entrainment (Zalalutdinov et al., 2003). Continuing with the analogy of a sinusoidally forced clock, imagine the situation where the clock and external force have a constant frequency ratio (e.g., 1:1) such that the clock ticks once for each forcing cycle. If the shaking frequency is continuously lowered below the lower bound of the zone of entrainment, then the clock transitions back to its intrinsic frequency" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0001395_012029-Figure5-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0001395_012029-Figure5-1.png", + "caption": "Figure 5. Brushless HTS dynamo concept.", + "texts": [ + " In the past, brushless concepts based on magnetic iron core were used that employed current leads spanning room-temperature and cryogenic regions. These current leads created a large thermal load for the cryocooler. Moreover, such an exciter is too heavy for aerospace applications where high specific power is required. A possible option is to employ the HTS dynamo [9,11] developed by Victoria University of Wellington (VUW). Its weight is expected to be less than 10% of the weight of magnetic core exciter. Figure 5 illustrates the HTS dynamo concept that is capable of supplying more than a kA of current [12]. Figure 6 shows how such a dynamo may be integrated into the motor as a passive device for energising the field winding without current leads spanning crossing the cryostat wall. The permanent magnets (PM) for excitation are attached to the outer wall of the stationary bayonet within the air core of the motor. The superconducting elements in which DC voltage is induced is attached to the outer wall of the rotating bayonet" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0001971_icsgrc49013.2020.9232583-Figure4-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0001971_icsgrc49013.2020.9232583-Figure4-1.png", + "caption": "Fig. 4. Scanned Input File", + "texts": [ + " The printing speed, layer height and extrusion temperature with their chosen levels have been combined with each other in order to form the optimum parameter levels combination that helped the researchers ensure the accuracy of the set dimensions in the actual design. Table 2. Parameters and Their Range Per Trial Run TEMP Degree C HEIGHT mm SPEED mm/sec 1 210 0.3 50 2 210 0.1 50 3 230 0.3 50 4 230 0.1 70 5 230 0.3 70 6 230 0.1 50 7 210 0.3 70 8 210 0.1 70 187 Authorized licensed use limited to: Dalhousie University. Downloaded on November 01,2020 at 15:43:32 UTC from IEEE Xplore. Restrictions apply. After scanning the said file as shown on Figure 4, the 3D printed output files use a Netfabb Software to compile and to measure the scanned/captured images of the 3D object and a sample image of what Netfabb software looks like. The researchers used Minitab software to easily analyze the data and have a basis in rejecting the null hypothesis. Minitab software generated important data such as p-value, f-value, degrees of freedom, r-squared value and percent contribution. The data presented in tabular form with the factors and interactions as shown in Figure 5" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0000241_012039-Figure3-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0000241_012039-Figure3-1.png", + "caption": "Figure 3. The successive states of the hydraulic manipulator in the process of loading the assortment into the forest truck (time points t of the computer experiment are indicated): a \u2013 top view (projection on the XY plane); b \u2013 starboard view (projection on the YZ plane); c \u2013 rear view (projection on the XZ plane).", + "texts": [ + " This program allows, by changing the basic geometric, functional and inertial parameters of a hydraulic manipulator with a regenerative hydraulic drive, to imitate its operation in loading or unloading cycles, as well as to optimize the basic parameters of the regenerative hydraulic drive. In the simulation model, a typical process of loading assortments with a hydraulic manipulator into a timber truck is reproduced. The hydraulic manipulator grabs the assortment from the bundle of assortments, deviating 900 from the direction of the forest truck, and lowering the arrow to grab the assortment with a grab (figure 3). After that, the boom of the hydraulic manipulator turns upward and at the same time returns to the longitudinal axis of the logging truck. To raise the assortment through the side stops of the truck body, the boom of the hydraulic manipulator rotates with a margin up. To create such a trajectory in the simulation model, the time dependences of the corresponding flow rates of the working fluid in the hydraulic cylinders for lifting and turning the boom are calculated. A similar boom path was used when unloading a timber truck with a hydraulic manipulator", + " Thus, the recovered energy for the entire loading cycle is determined. During movement by the boom, the grade is rocked relative to the hinge of the gripper grip, which is reproduced with sufficient adequacy in the simulation model. In particular, when the boom stops above the Forestry 2019 IOP Conf. Series: Earth and Environmental Science 392 (2019) 012039 IOP Publishing doi:10.1088/1755-1315/392/1/012039 body of the forest vehicle at the time of 5.4 s, the grade continues to oscillate and in the model is released later, at the time of 6.0 s (figure 3). The effect of the swinging significant mass on the boom and hydraulic system is taken into account in the model, and this action affects the recovered energy. The main indicator characterizing the efficiency of the recovery system is ECP energy stored in a pneumohydraulic accumulator during one loading cycle (or ECP energy per unloading cycle): 0 , CPt CP PGA PGAE P t Q t dt (14) where tCP \u2013 duration of one loading cycle; PPGA \u2013 pressure in the pneumohydraulic accumulator; QPGA \u2013 the rate of flow of the working fluid in the pneumohydraulic battery" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0000435_022055-Figure2-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0000435_022055-Figure2-1.png", + "caption": "Figure 2. PRHM with jointed engagements [6]", + "texts": [ + " However, in this situation, side force, affecting the satellite wheel, increases. In the end, this measure will limit maximum fluid pressure. B) The increase of angle \u03bb (see figure 1) of holding of satellite wheel as much as it is possible. The restrictions are connected in some cases with interference (co-occurrence condition) of central wheels, in other cases with the risk of satellite wheel \u201cfalling out\u201d. C) The change of the form of the teeth \u2013 transformation from involute to jointed engagement [6] \u2013 see figure 2. ICMTMTE IOP Conf. Series: Materials Science and Engineering 709 (2020) 022055 IOP Publishing doi:10.1088/1757-899X/709/2/022055 For PRHM with such engagements the ratio of sliding speed Vs to the calculated circular speed V\u0440 in the engagement: 32 ZV V p s . (13) It means that only thanks to the decrease of sliding speed along teeth profiles losses will be by 5.125.1 4 k times smaller than in involute PRHM (12). Extra decrease of mechanical losses will be due to the decrease of friction coefficient f in the teeth contact close to surface" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0000623_012064-Figure4-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0000623_012064-Figure4-1.png", + "caption": "Figure 4. Geometric models of outer rings (a), roller bearings (b) and axle box housing (c).", + "texts": [], + "surrounding_texts": [ + "Currently, there are significant changes in the operating conditions of wheel sets of railway cars. The greatest contribution to these changes is made by an increase in the speeds of movement of the compositions and an increase in their weight [1\u20133]. At the same time, the decisive contribution to ensuring the reliability of wheelset operation is made by increasing the operational reliability of the wheels and elements of axle boxes \u2013 rollers and bearing rings [4\u20139]. One of the effective ways to improve them is to search for rational design options based on the use of mathematical models of their stress-strain state (SSS) and experimentally verified criteria for assessing their durability [10\u201312]. Under the conditions of cyclic operation of high-loaded bearing elements of structures, such a criterion, as a rule, is the intensity of stresses [13\u201322]." + ] + }, + { + "image_filename": "designv11_71_0002485_cac51589.2020.9327292-Figure4-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0002485_cac51589.2020.9327292-Figure4-1.png", + "caption": "Fig. 4: Torque balance under P2P4 fault (H < 0)", + "texts": [ + " Compared with the tiltable quad-rotor UAV established in [8], the influence of the position of opi along the axis zb in the body coordinate system is more considered. Therefore, (3) is revised as: ob pi = Rz((1 \u2212 i) \u03c0 2 ) [ L 0 H ]T , i = 1 . . . 4. (13) where H is the distance between the center of gravity ob to the axis of rotation opi in the zb direction, and H is positive while ob below opi (as Fig. 3 shown). Then, in the condition of P2 and P4 have failed, the roll channel of the aircraft needs to balance the moment generated by the component force of the propeller thrust Fpi in the tangential direction of the obopi axis along zb (as Fig. 4 shown). Similarly, when P1 and P3 have failed, the pitch channel needs the moment generated by the component force of the Fpi in the tangential direction of the obopi axis to balance. III. Controller design Based on the classical control theory, this paper linearizes the tilting quad-rotor model proposed in Section II. In the case of the fault of the opposing propeller, the transfer function is used to describe the linearized model of each attitude angle, and the controller is designed according to the characteristics of the transfer function" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0000105_012019-Figure3-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0000105_012019-Figure3-1.png", + "caption": "Figure 3. Planar mechanism with all members: a) 3D view of the model, b) Display of trajectory of characteristic mechanism points using the 'Trace Path' option.", + "texts": [ + " After establishing calculation of motion kinematic parameters of the mechanism, structural analysis can be performed due to the force induced by motion dynamics (acceleration and force in the bonds) or external acting forces on the members of the mechanism. After the calculation of the movement of the mechanism, it is possible to obtain time-based motion diagrams, trajectory of characteristic moving points of the members of the mechanism, speed vectors and accelerations for characteristic points, using the \u201eResults and Plots\u201c option. IRMES 2019 IOP Conf. Series: Materials Science and Engineering 659 (2019) 012019 IOP Publishing doi:10.1088/1757-899X/659/1/012019 Figure 3a) shows a 3D model of examples with all the geometry parameters of the mechanism members, and Figure 3b) shows an example of using the 'Trace Path' option to obtain the path of the required moving points of the mechanism, [5, 6, 7]. Figure 4 shows the speed change diagrams for the B and C points of mechanism in the time domain. Figure 5 shows the acceleration diagrams for the B and C points with the ability to display the normal and tangential components of acceleration of the moving points of the mechanism. IRMES 2019 IOP Conf. Series: Materials Science and Engineering 659 (2019) 012019 IOP Publishing doi:10" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0001145_s40430-020-02387-2-Figure3-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0001145_s40430-020-02387-2-Figure3-1.png", + "caption": "Fig. 3 Knee joint line diagram", + "texts": [ + " The parameters of the subjects\u2019 lower limb are given in Table\u00a01. A subject of mass 56.7\u00a0kg is considered for the study, and the inputs are scaled anthropometrically. The length of the damper keeps varying with the knee and so does the damping force, which in turn controls the torque at the knee joint. The upper part of damper is connected to the thigh by extending a small perpendicular pin of length s from the knee joint and the other end of damper is linked to the shank at a distance of b from the joint, K as shown in Fig.\u00a03. The torque at the knee joint is related to the damping force by Eq.\u00a0(4). The angular displacement is positive in the counterclockwise direction, and thus, the damping force always acts opposite to the motion which explains the negative sign in the torque equation. Journal of the Brazilian Society of Mechanical Sciences and Engineering (2020) 42:294 1 3 Page 5 of 14 294 Considering the triangle PQK in Fig.\u00a02, the angle can be calculated using the set of Eqs.\u00a0(5) and (6). A twin-rod MR damper operating in the shear mode is considered in this study", + " A total of 500 turns of copper coil of wire gauge 26 were wound on the web of the piston, and the sleeve was press-fit. The outer dimensions of the damper are selected based on the normal human knee size and shape constraints, and the dependent variables such as the number of turns of coil, wire gauge of coil and gap size are determined based on the outer geometry of damper and also using the previously published literature. The fabricated damper is shown in Fig.\u00a05. The fabricated damper in this study has a fully extended length of 0.24\u00a0m and a stroke of 50\u00a0mm. Thus, the value of b in Fig.\u00a03 is fixed to 0.22\u00a0m such that the damper can safely reach full knee extension and maximum knee flexion during walking. Similarly, the value of s is taken as 0.035\u00a0m. Further, initial computational studies conducted using just the knee joint damper showed that the shank does not reach full knee extension at this phase end. The use of a spring in compressed form can store some energy in the stance phase and assist the shank to full extension. (4) k = \u2212Fdb sin (5)ld = \u221a s2 + b2 \u2212 2bs sin k (6) = sin\u22121 ( s cos k ld ) The current prostheses such as stabilized knee also use springs for active knee extension [30]", + " Experimental testing of the MR damper was performed using the damper testing machine shown in Fig.\u00a06. The tests were conducted at varying amplitudes, frequencies and currents. Amplitudes of 5 to 10\u00a0mm in steps of 5\u00a0mm, frequencies of 0.5 to 2.5\u00a0Hz in steps of 0.5\u00a0Hz and currents of 0 to 1.2 A in steps of 0.4 A were considered for the experiments. The selected amplitude and frequency sets yield velocities ranging from 0.0157 to 0.235\u00a0m/s. In normal human walking, the knee joint reaches a maximum speed of 60\u00a0rpm [4]. For the considered knee damper geometry shown in Fig.\u00a03, this results in a maximum speed of 0.21\u00a0m/s. Therefore, the selected amplitudes and frequency data sets cover the entire range of normal human walking. For the given range of amplitudes, frequencies and current, a total of 60 experiments were conducted. Experimental results for a sample of data sets are shown in Fig.\u00a07. According to the Bouc\u2013Wen model, the damping force of the MR damper is given by Eq.\u00a0(8) [33]. Here, K is the stiffness and C is the damping coefficient, x is the piston displacement, x\u0307 is velocity, z is the hysteresis variable, is the scaling parameter and f0 is the force offset", + " This is later used in the differential equation of motion given by Eq.\u00a0(1) and the knee angle trajectories are evaluated. The PD plus CT controller minimizes the error difference between the desired and actual knee angle displacements. The structure of the model employed in Simulink is shown in Fig.\u00a09. The structure shown in Fig.\u00a09 is implemented in Simulink software. For the prosthetic limb to reach full extension, the spring stiffness, ks and initial length of the spring are set to 1300\u00a0N/m and 0.27\u00a0m, respectively. The total length PQ in Fig.\u00a03 in fully extended knee position is 0.2236\u00a0m, which is below the assumed initial length of the spring; thus, the spring remains in the compressed state till the end of the swing phase. The control parameters, Kp and Kd , are tuned to reduce the error difference between the controller-estimated torque and the desired knee torque. The controller-estimated torque and the desired torque are shown in Fig.\u00a010. The achievable range is defined as the range in which both the passivity and limitation constraints defined by Eqs" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0000829_ijthi.2020070107-Figure3-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0000829_ijthi.2020070107-Figure3-1.png", + "caption": "Figure 3. Structure characteristics and assembly model of NC turning centre", + "texts": [], + "surrounding_texts": [ + "Established\ufeffon\ufeffthe\ufeffbase\ufeffof\ufeffthe\ufeffgeometric\ufeffassembly\ufeffmodel,\ufeffNC\ufeffmachine\ufefftool\ufeffmotion\ufeffmodel\ufeffhas\ufeffbeen\ufeff used\ufeffto\ufeffdescribe\ufeffthe\ufeffmotion\ufeffof\ufeffthe\ufeffmachine\ufefftool.\ufeffStudents\ufeffare\ufeffinstructed,\ufefffirst\ufeffof\ufeffall,\ufeffto\ufeffimport\ufeffthe\ufeff assembly\ufeffmodel\ufeffof\ufeffNC\ufeffturning\ufeffcenter\ufeffto\ufeffprocessing\ufeffmodule\ufeffof\ufeffNX\ufeffsoftware,\ufeffand\ufeffto\ufeffallocate\ufeffin\ufeffthe\ufeff Machine\ufeffTool\ufeffBuilder\ufeff the\ufeffspecific\ufeffmoving\ufeffparts\ufeffof\ufeffNC\ufeffmachine\ufeff tool.\ufeffBesides,\ufeff it\ufeff is\ufeffnecessary\ufeff to\ufeff certify\ufeffthe\ufeffmutual\ufeffmovement\ufeffrelationship\ufeffbetween\ufeffthe\ufeffmoving\ufeffparts,\ufeffto\ufeffset\ufeffup\ufeffrange\ufeffand\ufeffdirection\ufeffof\ufeff each\ufeffmoving\ufeffcoordinate\ufeffaxis,\ufeffas\ufeffwell\ufeffas\ufeffthe\ufeffparameters\ufeffof\ufeffmachine\ufefftool\ufeffjoint\ufeffpoints.\ufeffFor\ufeffthis\ufeffpart,\ufeffit\ufeff is\ufeffonly\ufeffafter\ufeffthe\ufeffstudents\ufefffully\ufeffmaster\ufeffthe\ufeffkinematic\ufeffcontents\ufeffof\ufeffmotion\ufeffmodel\ufeffof\ufeffNC\ufeffmachine\ufefftool,\ufeff together\ufeffwith\ufeffits\ufeffspecific\ufeffoperation\ufeffparameters,\ufeffcan\ufeffthe\ufeffsimulation\ufeffmotion\ufeffmodes\ufeffof\ufeffNC\ufeffmachine\ufefftool\ufeff be\ufeffdetermined.\ufeffTake\ufeffthe\ufeffNC\ufeffturning\ufeffcenter\ufeffof\ufeffour\ufeffexperiment\ufeffcenter\ufefffor\ufeffinstance,\ufeffstudents\ufeffmust\ufeffread\ufeff firstly\ufeffthe\ufeffrelevant\ufeffinformation\ufeffand\ufeffoperation\ufeffmanual\ufeffof\ufeffthe\ufeffNC\ufeffmachine\ufefftool.\ufeffAnd,\ufeffthough\ufeffin-depth\ufeff understanding\ufeffof\ufeff the\ufeffgeneral\ufeffsituation,\ufeff the\ufeffperformance\ufeffparameters\ufeffand\ufeffoperation\ufeffprocess\ufeffof\ufeffour\ufeff machine\ufefftool,\ufeffstudents\ufeffcan\ufeffdesignate\ufeff2\ufeffkinematic\ufeffchains\ufeffin\ufeffour\ufeffNC\ufeffturning\ufeffcenter\ufeffand,\ufeffsubsequently,\ufeff define\ufeff the\ufeff sub-components\ufeffof\ufeff the\ufeffkinematics\ufeffmodel\ufeffof\ufeffmachine\ufeff tool.\ufeffWith\ufeff the\ufeff establishment\ufeffof\ufeff motion\ufeffrelationship,\ufeffthe\ufeffmotion\ufeffmodel\ufeffof\ufeffmachine\ufefftool\ufeffcould\ufeffbe\ufefffinally\ufefffinished\ufeffas\ufeffshown\ufeffin\ufeffFigure\ufeff4. After\ufeffthe\ufeffmachine\ufefftool\ufeffmotion\ufeffmodel\ufeffhas\ufeffbeen\ufeffestablished\ufeffwell,\ufeffstudents\ufeffwould\ufeffactively\ufeffengage\ufeff in\ufeffthe\ufeffstudy\ufeffof\ufeffNC\ufeffmachine\ufefftool\ufeffresearch\ufeffin\ufeffthe\ufeffreal\ufeffworld,\ufeffin\ufeffthe\ufeffreading\ufeffand\ufeffcontemplation.\ufeffIn\ufeff order\ufeff to\ufeffcomplete\ufeff the\ufeffvirtual\ufeffsimulation\ufeffsystem\ufeffof\ufeff the\ufeffNC\ufeffmachine\ufeff tool,\ufeff to\ufeffpursue\ufeffof\ufeffa\ufeffsense\ufeffof\ufeff accomplishment,\ufeffstudents\ufeffare\ufeffcultivated\ufeffto\ufeffstudy\ufeffsimultaneously\ufeffin\ufefftheory\ufeffand\ufeffpractice.\ufeffThe\ufeffprocess\ufeff of\ufeffestablishing\ufeffthe\ufeffentire\ufeffmachine\ufeffmovement\ufeffmodel\ufeffis\ufeffalso\ufeffthat\ufeffof\ufeffdeepening\ufeffthe\ufeffunderstanding\ufeffof\ufeff the\ufeffNC\ufeffmachine\ufefftools,\ufeffwhich\ufeffprepares\ufefffor\ufeffthe\ufeffsubsequent\ufeffpractical\ufeffoperation." + ] + }, + { + "image_filename": "designv11_71_0001411_ilt-03-2020-0083-Figure4-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0001411_ilt-03-2020-0083-Figure4-1.png", + "caption": "Figure 4 Cavitation photographs and cavitation boundary point definitions of the SG-LFS", + "texts": [ + " As the convergence is achieved, the distributions of f and F are obtained. Both the dimensionless pressure P and film pressure p can be obtained by equations (9) and (10), respectively: P \u00bc Ff (9) p \u00bc Ff po pc\u00f0 \u00de1 pc (10) The flow chart of calculation procedure is shown in Figure 3. 3.1 Cavitation and analysis definition A visual experiment of cavitation (Li et al., 2020) for the SGLFS under the operating conditions of pi = 0.6MPa, po = 0.1MPa, m = 0.032Pa\u00b7s@45\u00b0C and the rotating speed n = 2,500 r/min is shown in Figure 4. Effects of operating conditions Zhentao Li et al. Industrial Lubrication and Tribology According to Figure 4(a), the liquid film starts to rupture along the lower land-groove interface and reforms in the spiral groove. The reason is that the liquid film thickness is divergent at the lower land-groove boundary, but part of the liquid film cannot withstand the formation of negative pressure and begins to rupture. In the groove, the liquid film is squeezed and becomes convergent to the boundary of higher land-groove, the liquid film pressure increases and the ability to resist the formation of negative pressure is gradually enhanced, the liquid film reassembling and cavitation zone being generated. Effects of operating conditions Zhentao Li et al. Industrial Lubrication and Tribology During the analysis of cavitation induction, the initial rupture position in the spiral direction r1-r2 and the finial reformation position in the circumferential direction u 1-u 2 of liquid film are taken as the analysis positions. In Figure 4(c), boundary points O, V and U are defined as the inner starting end, outer starting end and finial reformation position of liquid film reformation boundary, respectively. 3.2 Effect of process coefficient c Figure 5 shows the changes of film pressure and relative density with the increase of process coefficient g from 0.5 104 to 4 104. Figure 5(a) shows that all the pressure peaks in the inner dam areas decrease to cavitation pressure and then enter the cavitation area at the inner dam-groove boundary with different pressure gradients" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0003971_robio.2014.7090583-Figure2-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0003971_robio.2014.7090583-Figure2-1.png", + "caption": "Fig. 2. The quadrotor testbed platform.", + "texts": [ + "10(\ud835\udc61) = ( \ud835\udc50\ud835\udc651(\ud835\udc61)\ud835\udc60\ud835\udc653(\ud835\udc61)\ud835\udc60\ud835\udc655(\ud835\udc61)\u2212\ud835\udc60\ud835\udc651(\ud835\udc61)\ud835\udc50\ud835\udc655(\ud835\udc61) ) \ud835\udc40 \ud835\udc621(\ud835\udc61) ?\u0307?11(\ud835\udc61) = ?\u0307?(\ud835\udc61), ?\u0307?12(\ud835\udc61) = \u2212\ud835\udc54 + \ud835\udc50\ud835\udc651(\ud835\udc61)\ud835\udc50\ud835\udc653(\ud835\udc61) 1 \ud835\udc40 \ud835\udc621(\ud835\udc61) (2) Note that the state variables are \ud835\udc657 = \ud835\udc65, \ud835\udc658 = ?\u0307?, \ud835\udc659 = \ud835\udc66, \ud835\udc6510 = ?\u0307?, \ud835\udc6511 = \ud835\udc67, \ud835\udc6512 = ?\u0307?. The notations \ud835\udc50 and \ud835\udc60 are the abbreviations for cosine and sine. In this section, the overall configuration of the quadrotor is described in detail. This consists of hardware structure, controller design, and sensor system. Our experimental quadrotor aircraft was built from locally available materials and low-cost electronic circuitry. Fig. 2 displays a prototype of quadrotor with a testbed platform. The platform restricts the aircraft to a fixed point in the treedimentional space. For roll and pitch control experiments, the test bed constructed from water pipes was used in order to let the quadrotor rotate in one direction at a time. A spindle motor in hard disk drive, shown in the inserted subfigure, was modified for yaw control movement. It allows the quadrotor spinning in a full circle when performing a test. Fig. 3 shows the overall hardware architecture of the quadrotor UAV" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0000923_ibcast47879.2020.9044556-Figure1-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0000923_ibcast47879.2020.9044556-Figure1-1.png", + "caption": "Fig. 1. Ball bearing geometry and the possible defects.", + "texts": [ + " Based on the fault classification in [17], bearing faults can be categorized into two types: 1) single-point faults, which are defined as a visible single fault; and 2) generalized roughness, which refers to a damaged bearing. A single-point fault produces a characteristic fault frequency that depends on the surface of the bearing that contains the fault. Because most rotating machines use rolling-element bearings that consist of an outer race and an inner race, single-point faults in a bearing considered in this study include the following: 1) outer-race fault (ORF); 2) inner-race fault (IRF); and 3) ball bearing fault (BBF), as shown in Fig. 1. The bearing fault-free case is denoted by BFF. A bearing fault introduces specific frequency components that depart from the normal distribution, which subsequently increases the kurtosis value. Fault-related torque oscillations at particular frequencies are often related to the shaft speed. The characteristic bearing fault frequency fc in different bearing faults is given by the following relationships [11]: fc = \u23a7\u23aa\u23aa\u23aa\u23aa\u23a8 \u23aa\u23aa\u23aa\u23aa\u23a9 fout = Nb 2 fr ( 1\u2212 db cos\u03b2 dp ) for ORF fin = Nb 2 fr ( 1 + db cos\u03b2 dp ) for IRF fball = dp db fr ( 1\u2212 ( db cos\u03b2 dp )2 ) for BBF (1) where fout is the ORF frequency, fin is the IRF frequency, fball is the ball fault frequency, db is the ball diameter, dP is the pitch ball diameter, Nb is the number of balls, \u03b2 is the ball contact angle (with the races), and fr is the mechanical rotor frequency" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0002386_012019-Figure7-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0002386_012019-Figure7-1.png", + "caption": "Figure 7 \u2013 Obstacles in the (x,y) and (r,z) planes", + "texts": [ + "1088/1757-899X/995/1/012019 This thesis ought to be victorious when the robot mobilizes towards a motionless or a target in motion and gradually attains it without experiencing any barriers in the path and traverse in a nonlinear path. Any obstacle in the path of the robot is treated as a sphere. As discussed earlier, the method of plane decomposition is used we use projections of the spherical obstacle in the 2 dimensional planes. The graphical representation of obstacles in both horizontal and vertical planes is as shown below: Figure 7 (a) and (b) depicts the obstacles in the (x,y) and (r,z) planes, respectively. In the (x,y) horizontal plane the velocity of the robot is defined by \u03c51= \u03c5r cos(\u03c6r). The angle formed by the robot with the (x,y) plane is as we know. The projection of the spherical obstacle on the horizontal plane is denoted by a circle named Obxyas shown in figure 8 (a). Now, the robot is moving towards the obstacle. At any given point in time (let us assume at time t1), the robot encounters an obstacle within its line of sight" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0001931_j.jmmm.2020.167474-Figure1-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0001931_j.jmmm.2020.167474-Figure1-1.png", + "caption": "Fig. 1. Hysteresis motor geometrical structure.", + "texts": [ + " Therefore, the angular movement of rotor\u2019s mesh is fulfilled easily, and then the discontinuity in the FEM equations of the rotor part that includes the J-A hysteresis model, is avoided and subsequently the eddy currents due to rotation are taken into account. Also, the J-A vector hysteresis model of the hard-magnetic material of the rotor is incorporated in the FEM. Then, the transient behavior of a typical hysteresis motor is studied during short period of over-excitation, using the proposed hybrid FEMBEM model and the results are compared with those of the frequencydomain FEM. Fig. 1 shows the geometrical structure of the studied motor. The stator includes a total of 36 slots and is made of nonoriented 3% Si steel sheets that the corresponding saturation curve can be found in literature [18]. The 3-phase stator winding has two poles and 236 turns in each phase coil and is fed by a 3-phase voltage source of 40 V with the frequency of 50 Hz. The rotor includes a solid magnetic ring of hard iron (Fe-Co 17%) with the electrical conductivity of 4.17 \u00d7 106 S/m [19]. The synchronous speed is 3000 rpm, and also the air\u2013gap is 0" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0001044_pesgre45664.2020.9070523-Figure7-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0001044_pesgre45664.2020.9070523-Figure7-1.png", + "caption": "Fig. 7 Flux Distribution due to (a) Fundamental current (b) Fundamental with 7th harmonics injection (c) Fundamental with 7th harmonics injection (Detailed).", + "texts": [ + " The flux distribution of BLDCM for fundamental current is shown in Fig. 6. Similarly, the eddy current loss including harmonics is = = n i iee fBKP 1 22 max (7) Where, Ke is the eddy current loss coefficient. Equations (6) & (7) indicate that hysteresis and eddy current loss are frequency dependent. Thus, current harmonics reduction ensures significant decrease of these losses. The proposed technique injects 7th harmonics in the system for performance improvement. The flux distribution without 7th harmonics injection is shown in Fig. 7 (a) where Figs. 7(b) and (c) shows the detailed flux distribution with 7th harmonics current 978-1-7281-4251-7/20/$31.00 \u00a92020 IEEE 3 Authorized licensed use limited to: University of Waterloo. Downloaded on June 01,2020 at 10:42:04 UTC from IEEE Xplore. Restrictions apply. excitation. The core loss is calculated using Ansoft-ANSYS software for both conventional and proposed control to prove the betterment in overall loss reduction. (c) Switching Loss: During inverter switching due to rise time, fall time and reverse recovery time of practical switches there are turn on and turn off loss in switches" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0000662_icus48101.2019.8995953-Figure4-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0000662_icus48101.2019.8995953-Figure4-1.png", + "caption": "FIGURE 4. Schematic drawing of the VSA system integrated with a reaction wheel.", + "texts": [ + " On the other hand, the benefits of the fast dynamics of the reaction wheel can be exploited with small sampling times. Therefore, we anticipate that the usage of the NMPC for VSA robots with reaction wheels will be limited to low-DOF systems for the foreseeable future. IV. CASE STUDY A. MODEL OF THE CASE STUDY SYSTEM In order to show the merits of the reaction wheel augmentation of VSA robots, we conducted a case study using a single-DOF VSA robot with a reaction wheel attached to it. The schematics of this system is shown in Fig. 4. The task is to throw a ball attached to the distal end of the robot to the farthest distance. The link is actuated by two servo-motors via two NEEs and the disk is actuated by a brushless DC motor. The dynamics of the systems can be obtained using the modeling formalism presented in subsection III-A. The inertia matrix M\u0303 for this system is M\u0303 = [ M11 Jw1 Jw1 Jw1 ] (15) withM11 = ml1(L l c,1) 2 +mw1 (L w c,1) 2 +mbL2c,b+ J l 1+ J w 1 + Jb, where ml1, J l 1,L l c,1 are the mass, the inertia about center of mass, and the distance from the pivot to center of mass of the link" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0000327_j.mechmachtheory.2019.103729-Figure1-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0000327_j.mechmachtheory.2019.103729-Figure1-1.png", + "caption": "Fig. 1. The structure of the classicl 3 \u00d7 3 \u00d7 3 order cube.", + "texts": [ + " Finally, the DOF of the Rubik\u2019s Cube mechanism based on the screw theory and the \u201cmodified K-G formula\u201d is solved. The Rubik\u2019s Cube has an artistic appearance, clever structure, and unique movement characteristics. Although it seems so simple and compact at first glance, it is actually ever-changing. It goes beyond the traditional mechanical concept and design method: (1) There are many components, kinematic pairs and loops in the Rubik\u2019s Cube mechanism. The structure of the classical 3 \u00d7 3 \u00d7 3 order cube is shown in Fig. 1 . The cube consists of about 30 loops, 54 kinematic pairs and 27 components, including 26 sub-pieces (6 center pieces, 8 corner pieces and 12 side pieces) and a 3-D cross. Each sub-piece includes a hidden inward extension that ensures the interlocking relationship between adjacent sub-pieces. Inward extensions on side and corner pieces are clips for mosaics. All sub-pieces are connected by the 3-D cross with three orthogonal axes. Specifically, six center pieces form rotational connections with six axes of the 3-D cross, and the remaining 20 sub-pieces are inlaid with clips to form a complete Rubik\u2019s Cube" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0003456_j.ifacol.2015.09.392-Figure2-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0003456_j.ifacol.2015.09.392-Figure2-1.png", + "caption": "Fig. 2. The simplified mechanical model of the skateboardskater system", + "texts": [ + " Namely, by means of this simple strategy, the skater could avoid the instability of the board. But this strategy can work if the direction of the rectilinear motion does not matter, namely, the skater does not have to avoid an object on the road and/or does not have to follow the desired path of the road. To investigate the case when the skater has to follow a predefined direction by its board, another control law is needed. In this study, we consider this latter case. 2. NON-HOLONOMIC MECHANICAL MODEL The mechanical model in question (see Figure 2) is based on Kremnev and Kuleshov (2008) and Varszegi et al. 1 See in Varszegi et al. (2014) Proceedings of the 12th IFAC Workshop on Time Delay Systems June 28-30, 2015. Ann Arbor, MI, USA Copy ight \u00a9 IFAC 2015 286 Position Control of Rolling Skateboard Balazs Varszegi \u2217 Denes Takacs \u2217\u2217 Gabor Stepan \u2217\u2217\u2217 \u2217 Department of Applied Mechanics, Budapest University of Technology and Economics, Budapest, Hungary (e-mail: varszegi@mm.bme.hu) \u2217\u2217 MTA-BME Research Group o D namics of Machines and V hicles, Budapest, Hungary (e-mail: takacs@mm", + " Namely, by means of this simpl strategy, the skater could void the instability of the board. But is strategy can work if the direction of the rectilinear motion does not matter, namely, the skater does not h ve to avoid an object on he road and/or does n t have to follow the desired path of the road. To investigate the case when the kater has to follow a predefin d direction by its board, another control law is neede . I this study, we consider this latter case. 2. NON-HOLONOMIC MECHANICAL MODEL The mechanical model in question (see Figure 2) is based on Kr mnev nd Kuleshov (2008) and Varszegi et al. 1 See in Varszegi et al. (2014) Proceedings of the 12th IFAC Workshop on Time Delay Systems June 28-30, 2015. Ann Arbor, MI, USA Copyright IFAC 2015 286 Position Control of Rolling Skateboard Balazs Varszegi \u2217 Denes Takacs \u2217\u2217 Gabor Stepan \u2217\u2217\u2217 \u2217 Department of Applied Mechanics, Budapest University of Technology and Economics, Budapest, Hungary (e-mail: varszegi@mm.bme.hu) \u2217\u2217 MTA-BME Research Group on Dynamics of Machines and Vehicles, Budapest, Hungary (e-mail: takacs@mm", + " Namely, by means of this simple strategy, the skater could avoid the instability of the board. But this strategy can work if the direction of the rectilinear motion does not matter, namely, the skater does not have to avoid an object on the road and/or does not have to follow the desired path of the road. To investigate the case when the skater has to follow a predefined direction by its board, another control law is needed. In this study, we consider this latter case. 2. NON-HOLONOMIC MECHANICAL MODEL The mechanical model in question (see Figure 2) is based on Kremnev and Kuleshov (2008) and Varszegi et al. 1 See in Varszegi et al. (2014) Proceedings of the 12th IFAC Workshop on Time Delay Systems June 28-30, 2015. Ann Arbor, MI, USA Copyright \u00a9 IFAC 2015 286 ositio o trol of olli g ate oar Balazs arszegi \u2217 enes Takacs \u2217\u2217 abor Stepan \u2217\u2217\u2217 \u2217 Department of Applied echanics, Budapest University of Technology and Economics, Budapest, Hu gary (e-mail: varszegi@mm.bme.hu) \u2217\u2217 TA-B E Research Group on Dynamics of achines and Vehicles, Budapest, Hungary (e-mail: takacs@mm", + " Namely, by means of this simple strategy, t e skater could avoid the instability of the board. But this strategy can work if the direction of the rectilinear motion does not matter, namely, the skater does not have to avoid an object on the road and/or does not have to follow the desired path of the road. To investigate the case when the skater has to follow a pre efi ed direction by its board, another control law is needed. In this study, we consider this latter case. 2. NON-HOLONO IC ECHANICAL ODEL The mechanical model in question (see Figure 2) is based on Kremnev and Kuleshov (2008) and Varszegi et al. 1 See in Varszegi et al. (2014) Proceedings of the 12th IFAC Workshop on Time Delay Systems June 28-30, 2015. Ann Arbor, MI, USA Copyright \u00a9 IFAC 2015 286 Balazs Varszegi et al. / IFAC-PapersOnLine 48-12 (2015) 286\u2013291 287 Position Control of Rolling Skateboard Balazs Varszegi \u2217 Denes Takacs \u2217\u2217 Gabor Stepan \u2217\u2217\u2217 \u2217 Department of Applied Mechanics, Budapest University of Technology and Economics, Budapest, Hungary (e-mail: varszegi@mm.bme", + " Namely, by means of this simple strategy, the skater could avoid the instability of the board. But this strategy can work if the direction of the rectilinear motion does not matter, namely, the skater does not have to avoid an object on the road and/or does not have to follow the desired path of the road. To investigate the case when the skater has to follow a predefined direction by its board, another control law is needed. In this study, we consider this latter case. 2. NON-HOLONOMIC MECHANICAL MODEL The mechanical model in question (see Figure 2) is based on Kremnev and Kuleshov (2008) and Varszegi et al. 1 See in Varszegi et al. (2014) Proceedings of the 12th IFAC Workshop on Time Delay Systems June 28-30, 2015. Ann Arbor, MI, USA Copyright \u00a9 IFAC 2015 286 Position Control of Rolling Skateboard Balazs Varszegi \u2217 Denes Takacs \u2217\u2217 Gabor Stepan \u2217\u2217\u2217 \u2217 Department of Applied Mechanics, Budapest University of Technology and Economics, Budapest, Hungary (e-mail: varszegi@mm.bme.hu) \u2217\u2217 MTA-BME Research Group o D namics of Machines and V hicles, Budapest, Hungary (e-mail: takacs@mm", + " Namely, by means of this simpl strategy, the skater could void the instability of the board. But is strategy can work if the direction of the rectilinear motion does not matter, namely, the skater does not h ve to avoid an object on he road and/or does n t have to follow the desired path of the road. To investigate the case when the kater has to follow a predefin d direction by its board, another control law is neede . I this study, we consider this latter case. 2. NON-HOLONOMIC MECHANICAL MODEL The mechanical model in question (see Figure 2) is based on Kr mnev nd Kuleshov (2008) and Varszegi et al. 1 See in Varszegi et al. (2014) Proceedings of the 12th IFAC Workshop on Time Delay Systems June 28-30, 2015. Ann Arbor, MI, USA Copyright \u00a9 IFAC 2015 286 Position Control of Rolling Skateboard Balazs Varszegi \u2217 Denes Takacs \u2217\u2217 Gabor Stepan \u2217\u2217\u2217 \u2217 Department of Applied Mechanics, Budapest University of Technology and Economics, Budapest, Hungary (e-mail: varszegi@mm.bme.hu) \u2217\u2217 MTA-BME Research Group on Dynamics of Machines and Vehicles, Budapest, Hungary (e-mail: takacs@mm", + " Namely, by means of this simple strategy, the skater could avoid the instability of the board. But this strategy can work if the direction of the rectilinear motion does not matter, namely, the skater does not have to avoid an object on the road and/or does not have to follow the desired path of the road. To investigate the case when the skater has to follow a predefined direction by its board, another control law is needed. In this study, we consider this latter case. 2. NON-HOLONOMIC MECHANICAL MODEL The mechanical model in question (see Figure 2) is based on Kremnev and Kuleshov (2008) and Varszegi et al. 1 See in Varszegi et al. (2014) Proceedings of the 12th IFAC Workshop on Time Delay Systems June 28-30, 2015. Ann Arbor, MI, USA Copyright \u00a9 IFAC 2015 286 ositio o trol of olli g ate oar Balazs arszegi \u2217 enes Takacs \u2217\u2217 abor Stepan \u2217\u2217\u2217 \u2217 Department of Applied echanics, Budapest University of Technology and Economics, Budapest, Hu gary (e-mail: varszegi@mm.bme.hu) \u2217\u2217 TA-B E Research Group on Dynamics of achines and Vehicles, Budapest, Hungary (e-mail: takacs@mm", + " Namely, by means of this simple strategy, t e skater could avoid the instability of the board. But this strategy can work if the direction of the rectilinear motion does not matter, namely, the skater does not have to avoid an object on the road and/or does not have to follow the desired path of the road. To investigate the case when the skater has to follow a pre efi ed direction by its board, another control law is needed. In this study, we consider this latter case. 2. NON-HOLONO IC ECHANICAL ODEL The mechanical model in question (see Figure 2) is based on Kremnev and Kuleshov (2008) and Varszegi et al. 1 See in Varszegi et al. (2014) Proceedings of the 12th IFAC Workshop on Time Delay Systems June 28-30, 2015. Ann Arbor, MI, USA Copyright \u00a9 IFAC 2015 286 (2014). The skateboard is modeled by a massless rod (between the front axle at F and the rear axle at R) while the skater is represented by a massless rod (between the points S and C) with a lumped mass at C. In this model, the connection between the skater and the board (at S) is assumed to be rigid", + " Due to the fact that the longitudinal axis of the skateboard is always parallel to the ground, one can choose four generalized coordinates to describe the motion: X and Y are the coordinates of the skateboard center point S in the plane of the ground; \u03c8 describes the direction of the longitudinal axis of the skateboard; and finally, \u03d5 is the inclination angle of the skater\u2019s body from the vertical direction. The geometrical parameters are the following. The height of the skater is denoted by 2h. The length of the board is 2l while m represents the mass of the skater. The parameter g stands for the gravitational acceleration. Here we model the skater\u2019s navigating effort as a linear PD controller, which applies a torque to the skateboard (see Figure 2). The rectilinear motion of the board is prescribed along the Y = 0 line, and the control torque is calculated via Mc (t) = \u2212PY (t\u2212 \u03c4)\u2212DY\u0307 (t\u2212 \u03c4) (1) where \u03c4 refers to the time delay, P and D represent the proportional and the differential control gains, respectively. The controller produces zero torque if the board follows the prescribed stationary path. Regarding to the rolling wheels of the skateboard, kinematic constraints can be formed. These constraining equations define the velocities vF and vR of the front point (F) and the rear one (R), respectively. The directions of these velocities depend on \u03d5 through \u03b4S, which is the so-called steering angle (see Figure 2). This angle can be expressed from the equation sin\u03b2 (t) tan\u03ba = tan \u03b4S (t) , (2) where \u03ba is the complementary angle of the so-called rake angle in the skateboard wheel suspension (for the derivation of this relation please see Kremnev and Kuleshov (2008) or Varszegi et al. (2014)). Here we also prescribe the longitudinal speed of the board, which is kept on the constant value V . The so-formed kinematic constraints can be written as A \u00b7 q\u0307 = A0 , (3) where qT = [ X Y \u03c8 \u03d5 ] , (4) A = [ sin\u03c8 \u2212 cos\u03c8 sin\u03d5 tan\u03ba \u2212 cos\u03c8 \u2212 sin\u03c8 sin\u03d5 tan\u03ba \u2212l 0 sin\u03c8 + cos\u03c8 sin\u03d5 tan\u03ba \u2212 cos\u03c8 + sin\u03c8 sin\u03d5 tan\u03ba l 0 cos\u03c8 sin\u03c8 0 0 ] , (5) A0 T = [ 0 0 V ] , (6) The equations of motion of non-holonomic systems can be determined by means of several methods" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0003912_amm.635-637.1355-Figure1-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0003912_amm.635-637.1355-Figure1-1.png", + "caption": "Fig. 1 Link coordinate of 6 DOF underwater robot", + "texts": [ + " An optimal solution are obtained from 8 sets of feasible solutions. The establishment of the robot kinematics equation In order to make the handling robot system has stronger function and complete more complicated tasks, the degrees of freedom are 6. The 6 joints are all rotational joints, respectively is the shoulder revolute joint, shoulder pitching joint, elbow pitching joint, arm rotating joint, wrist pitching movement joint, wrist rotary movement joint. A mechanical linkage coordinate system is build according to the D-H method as shown in Fig. 1. Each link parameters of the robot joint is 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 Ltd, www.scientific.net. (#525119300, Link\u00f6pings Universitetsbibliotek, Link\u00f6ping, Sweden-05/01/20,00:52:18) Table1 Link parameters of the robot Link number i 1ia \u2212 1i\u03b1 \u2212 id i\u03b8 Joint variable range Link parameters (mm) (\u00b0) (mm) (\u00b0) (\u00b0) (mm) 1 0 0 0 1\u03b8 -10~90 2 1a -90 0 2\u03b8 -90~30 1a =95 3 2a 0 0 3\u03b8 -100~30 2a =805 4 3a -90 4d 4\u03b8 -135~135 3a =182 5 0 90 0 5\u03b8 -60~60 4d =569 6 0 -90 6d 6\u03b8 -180~180 6d =450 According to the D-H parameters of the robot in Table 1, the forward kinematics equation of the handling robot with six degree of freedom can be written as follows: ( ) ( ) ( ) ( ) ( ) ( )0 0 1 2 3 4 5 6 1 1 2 2 3 3 4 4 5 5 6 6 0 0 0 1 x x x x y x x x z x x x n o a p n o a p T T T T T T T n o a p \u03b8 \u03b8 \u03b8 \u03b8 \u03b8 \u03b8 = = \uff081\uff09 The inverse solution of the kinematics of robot The end pose of the handling robot with 6 DOF has been given, scilicet , ,n o a and p is known, then, seeking joint variables 1 2 6, , ,\u03b8 \u03b8 \u03b8 is called inverse solution of the kinematics of the robot" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0000126_j.ifacol.2019.10.069-Figure11-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0000126_j.ifacol.2019.10.069-Figure11-1.png", + "caption": "Fig. 11. Collision at top layers when closing the dome built in the orientation shown in Fig. 10 (a).", + "texts": [ + " The three chosen slicing directions for slicing the dome are shown in Fig. 10. The green sections of the dome have an overhang angle less than 10\u00b0, and the sections with the red colour have overhang angles higher than 10\u00b0. The dome can be made on a 3 axis machine if the red surfaces have the support structures underneath them. However, the amount of material will be significant. Although a 5 axis machine provides the opportunity to address surface normal variations, closing the dome top for scenario in Fig. 10 (a) causes collisions at the top layers as shown in Fig. 11. For the build orientation in Fig. 10 (b), the lower layers have a high overhang angle (90\u00b0). In this case, if the nozzle maintains the tangency condition with respect to the surface, it will collide with the table when building the first layers, as shown in Fig. 12. Fig. 10 (b) orientation, and radial partitioning for the Fig. 10 (c) orientation. In order to avoid any collision at the first layers, a round bar can be used as the substrate and several layers be deposited at one end of the bar. The number and consequently the thickness of deposited layers needs to be more than the width of beads that forms the dome", + " Two solutions are available to eliminate this problem: (i) adding support structures for the problematic surfaces or (ii) partitioning the surface and introducing 2019 IFAC IMS August 12-14, 2019. Oshawa, Canada Hamed Kalami et al. / IFAC PapersOnLine 52-10 (2019) 230\u2013235 235 prevents the collision of the machine head to the table. After the dome is built, it just needs to be cut from the round bar. This is an interesting solution, but it may be difficult to expand upon. The geometry presented in Fig. 4 would be challenging. Fig. 11. Collision at top layers when closing the dome built in the orientation shown in Fig. 10 (a). Fig. 12. Collision of the nozzle to machine table when making dome in the orientation shown in Fig. 10 (b). Fig. 13. Building a dome from the end of a round bar by rotary production strategy. a) 2.5 axis layers at the end of the bar. b) Building the dome at the end of bar. c) The way this prevents collision on the machine. 3.2 Round partitioning of the thin wall dome Based on the orientation showed in Fig" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0000126_j.ifacol.2019.10.069-Figure12-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0000126_j.ifacol.2019.10.069-Figure12-1.png", + "caption": "Fig. 12. Collision of the nozzle to machine table when making dome in the orientation shown in Fig. 10 (b).", + "texts": [ + " However, the amount of material will be significant. Although a 5 axis machine provides the opportunity to address surface normal variations, closing the dome top for scenario in Fig. 10 (a) causes collisions at the top layers as shown in Fig. 11. For the build orientation in Fig. 10 (b), the lower layers have a high overhang angle (90\u00b0). In this case, if the nozzle maintains the tangency condition with respect to the surface, it will collide with the table when building the first layers, as shown in Fig. 12. Fig. 10 (b) orientation, and radial partitioning for the Fig. 10 (c) orientation. In order to avoid any collision at the first layers, a round bar can be used as the substrate and several layers be deposited at one end of the bar. The number and consequently the thickness of deposited layers needs to be more than the width of beads that forms the dome. This is shown in Fig. 13 (a) which 6 layers are deposited at the end of the bar by using just 2.5 axis movement. Then as Fig. 13 (b) shows, the dome starts being built from the side of the last bead on the bar", + " Two solutions are available to eliminate this problem: (i) adding support structures for the problematic surfaces or (ii) partitioning the surface and introducing 2019 IFAC IMS August 12-14, 2019. Oshawa, Canada Hamed Kalami et al. / IFAC PapersOnLine 52-10 (2019) 230\u2013235 235 prevents the collision of the machine head to the table. After the dome is built, it just needs to be cut from the round bar. This is an interesting solution, but it may be difficult to expand upon. The geometry presented in Fig. 4 would be challenging. Fig. 11. Collision at top layers when closing the dome built in the orientation shown in Fig. 10 (a). Fig. 12. Collision of the nozzle to machine table when making dome in the orientation shown in Fig. 10 (b). Fig. 13. Building a dome from the end of a round bar by rotary production strategy. a) 2.5 axis layers at the end of the bar. b) Building the dome at the end of bar. c) The way this prevents collision on the machine. 3.2 Round partitioning of the thin wall dome Based on the orientation showed in Fig. 10 (c), it is needed to be partitioned properly to be able to be built. Here, a 3 + 2 machine configuration is assumed" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0002923_9781119633365-Figure4.8-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0002923_9781119633365-Figure4.8-1.png", + "caption": "Figure 4.8 Two examples to illustrate the feedback configuration: (a) forward code D with a positive unity-gain feedback code and (b) forward code D/(1 \u2212\u00a0D) with a negative unity-gain feedback code.", + "texts": [ + " The second type of transfer code configuration is a feedback configuration, as shown in Figure\u00a04.7, where TCf is the forward transfer code and TCb is the feedback one. The input\u2010to\u2010output transfer code can be obtained as follows: V V TC TC TC o i f f b1 (4.6) The above equation is valid if and only if the forward power flows through TCf exclusively and the return power flows through TCb only. A practical example with the transfer codes of TCf\u00a0 =\u00a0 D and TCb\u00a0 = 1, and with a positive feedback (+) is shown in Figure\u00a04.8a, which yields Vo/Vi\u00a0=\u00a0D/(1 \u2212\u00a0D). Another example shown in Figure\u00a04.8b is with TCf\u00a0=\u00a0D/(1 \u2212\u00a0D), TCb\u00a0= 1, and a negative feedback (\u2212), which yields Vo/Vi\u00a0=\u00a0D. In the next chapter, we will use a buck converter with a positive unity feedback to derive a buck\u2013boost converter and then derive the buck converter from the buck\u2013boost converter with a negative unity feedback. These two examples can illustrate the feedback configuration shown in Figure\u00a04.7. The third type of transfer code configuration is a feedforward configuration, as shown in Figure\u00a04.9, in which the one in Figure\u00a04" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0000712_ssci44817.2019.9002649-Figure1-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0000712_ssci44817.2019.9002649-Figure1-1.png", + "caption": "FIGURE 1. A relay-aided massive MIMO cellular network.", + "texts": [ + " Results validate the effectiveness of the two methods. Notations: R represents real number,C represents complex number, vectors are denoted by italic lower case x, matrix is denoted by upper case bold A, \u2016.\u2016p represents p-norm (.)T represents transpose, (.)H represents Hermitian transpose, CN (\u00b7, \u00b7) represents complex normal distribution, E [\u00b7] represents expectation, diag(\u00b7 \u00b7 \u00b7) represents the diagonal matrix. II. SYSTEM MODEL Consider a single-cell multi-user multi-relay massive MIMO system, as shown in Fig. 1, where the BS is equipped with M antennas, together with N single-antenna relay stations, serving K single-antenna users. In this paper, two-hop relay is considered, and decode-and-forward (DF) and amplifyand forward (AF) protocols are employed in the uplink and downlink, respectively. Each RS can only communicate with one user, and the BS can simultaneously communicate with all users. Users connect to the BS or a RS using the nearby principle. With time division duplexing (TDD) operation, the BS obtains the downlink CSI through uplink training" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0002989_978-3-642-16214-5_9-Figure9.36-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0002989_978-3-642-16214-5_9-Figure9.36-1.png", + "caption": "Fig. 9.36. Damage to steel plates. a Extensive heat discolouration; b hot spots; c scoring, wear; d sinter transfer [9.30]", + "texts": [ + "35b): incorporation of oil-carbon residues as a result of high thermal load on the oil, \u2022 lining separation (Figure 9.35c): due to insufficient lining strength resulting from lack of resin (manufacturing error), \u2022 lining detachment: poor adhesion or chemical attack, i.e. due to rust under the lining, \u2022 lining wear: as a result of an unsuitable counter-running surface, abrasive foreign bodies, little operating clearance, \u2022 frictional martensite on sinter linings (Figure 9.35d): caused by excessive surface temperatures. The following damage types can be found in steel plates: \u2022 extensive heat discolouration (Figure 9.36a): high shifting work, long slipping times or not enough operating clearance lead to thermal overload due to excessive friction surface temperature, \u2022 hot spots (Figure 9.36b): high frictional power with short slipping times, frequently accompanied by unfavourable pressure conditions, leads to local thermal overloading. The hot spots are often distributed randomly. The cause can be inferred from the position of the hot spots. An even distribution indicates corrugated plates, uneven pistons or plate natural vibrations. If the hot spots are on the outer diameter, this suggests plate wobbling. \u2022 scoring, wear (Figure 9.36c): due to insufficient lubrication or abrasive foreign bodies, \u2022 matting: due to additive residues, \u2022 sinter transfer on sinter linings (Figure 9.36d): thermal overload due to lacking lubrication, insufficient operating clearance or poor design, \u2022 corrosion: resulting from water in the oil. The static clutch torque is supported/transmitted by the plate gearing in the inner and outer plate carriers. The pressure on the driving flanks (1) of the gearing has to be examined. (Note: the number in the parentheses, e.g. (1), refers to the algorithm shown in Figure 9.37.) The design of multi-plate clutches must guarantee that the torques are transmitted frictionally, i" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0001438_012062-Figure1-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0001438_012062-Figure1-1.png", + "caption": "Figure 1. 3D model and design diagram of the active mechanism.", + "texts": [ + " Exoskeletons that directly control movements in the joints of the limbs are more appropriate means of rehabilitation, in any case, with severe motor impairment. The advantage of robot therapy is a higher quality of training compared to classical rehabilitation due to their longer duration, the accuracy of repetitive cyclic movements, a constant training program, the availability of tools to evaluate the success of the classes with the possibility of demonstration to the patient [5]. The conceptual design of the system for the rehabilitation of the lower extremities is shown in Figure 1. The device consists of two mechanisms: a passive orthosis and an active parallel robot. The active 3- PRRR parallel robot proposed by Kong and Gosslen [6] and also known as Isoglyde has three degrees of freedom - translational movement along each axis. The mechanism consists of three kinematic chains \ud835\udc34\ud835\udc56\ud835\udc35\ud835\udc56 \ud835\udc36\ud835\udc56 \ud835\udc37\ud835\udc56. The position of the rehabilitation platform, which is an equilateral triangle \ud835\udc371\ud835\udc372 \ud835\udc373 centered at point P and the radius of the circumscribed circle R, is determined by linear displacements \ud835\udc92 = (\ud835\udc5e1\ud835\udc5e2 \ud835\udc5e3)" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0003554_gtindia2014-8203-Figure2-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0003554_gtindia2014-8203-Figure2-1.png", + "caption": "Figure 2. Casing of middle support.", + "texts": [ + " The blades of the fifth rotor wheel of five-stage intermediate pressure compressor of gas turbine engine was object of research in this paper (Figure 1). The casing of the engine middle support is located downstream the fifth compressor stage. High pressure compressor is located after the support casing. 1 Copyright \u00a9 2014 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 There are seven unevenly distributed racks of different cross-sections in the channel of support casing. (Figure 2). These racks are the cause of circumferential variation of the gas flow in gas-turbine engine passage, which leads to increased dynamic stresses in the fifth rotor wheel blades, as a consequence, to its breakage. The goal of this work was to reduce the dynamic stresses in the rotor blades of the intermediate pressure compressor fifth stage by means of the blades reprofiling and circumferential variation flow reduction. p static pressure, [Pa] p\u0305 relative static pressure \u03c6 circumferential pitch, [] GTE gas-turbine engine RW rotor wheel RWi number of rotor wheel IPC intermediate pressure compressor GV guide vane GVi number of guide vane IGV inlet guide vane HPC high pressure compressor CFD computational fluid dynamics FEM finite element method Method for dynamic stresses calculation in RW blades of the IPC fifth stage was created at the first stage [5]" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0000473_asjc.2283-Figure1-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0000473_asjc.2283-Figure1-1.png", + "caption": "FIGURE 1 Pioneer3-DX specifications", + "texts": [ + " N, x, j, and i denotes the kth input, the jth cluster center with the smallest distance in Equation. (2), the input data, and the ith dimension, respectively. After this step is completed, the algorithm again performs Step 2. This section describes the mobile robots and navigation control method. Navigation control is implemented by a behavior selector (BS). The proposed BS changes towards goal mode (TGM) or the wall-following mode (WFM) of the behavior of a robot, thus depending on the environmental conditions to achieve navigation control. Figure 1 shows the physical Pioneer3-DX, which is a compact differential-drive mobile robot. The mobile robot has numerous features, including reliability, high scalability, swift navigation, and programmability. In 22 a hybrid of the type-1 neural fuzzy network (T1NFN) and a functional link neural network called a functional neural fuzzy network (FNFN) is proposed for solving nonlinear control problems. A functional link neural network (FLNN) is adopted as the subsequent version of the T1NFN to improve accuracy, and the performance of a Takagi-Sugeno-Kang (TSK)-type NFN and the FNFN are compared" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0000663_appeec45492.2019.8994460-Figure4-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0000663_appeec45492.2019.8994460-Figure4-1.png", + "caption": "Fig. 4. Phasor diagram in the I/F acceleration (a) and torque characteristics used for the stability judgement (b).", + "texts": [ + " I/F acceleration The aforementioned procedure readily locates the rotor in the predetermined position, this step is intended to accelerate the motor from the static state to a minimum speed, at which SMO based on the back EMF could work accurately. In other words, during the acceleration process, the rotor position keeps unknown which means no speed feedback. The control strategy will be described in the synchronous rotating reference frame, just as the pre-location chapter, and the angle between the stator dq frame and the rotor drqr frame is defined as \u03b8, with the value positive when dq frame leads the drqr frame and vice versa, shown in Fig. 4(a). Apparently, the angle \u03b8 is 90\u00ba in the predetermined position of the prelocation step, which means that there is no torque applied on the motor. Combined with (6), the electromagnetic torque Tem can be expressed as: em fM q 3T = n\u03c8 i cos 2 (10) In the process of I/F acceleration, current and angle are given by direct command as shown in Fig. 3, which means that the current and speed can be appointed directly. Then if the q axis current iq increases from zero, in a similar way, the speed \u03c9 will increase from zero and correspondingly the electrical angle \u03b8 will also increase, resulting in the toque Tem increasing", + " It\u2019s noted that the final current command should be large enough to make \u03b8>0, which is helpful for the transition to speed feedback loop (step 3). In the I/F acceleration, without the position detection, it\u2019s vital to ensure the self-stabilization of the motor. Taking \u03b8 (0, ) as an example, if the q axis current keeps constant and the resistance torque TL increases, then \u03b8 increase and iqcos\u03b8 increase correspondingly, therefore the motor will generate bigger electromagnetic torque Tem than before and achieve motor balance again, referring to Fig. 4(b). However, when \u03b8 ( , ) increases caused by the increasing of TL, the Tem will decrease and thus become smaller and smaller than TL, leading to the out-of-step problem. A similar analysis can be conducted when \u03b8 (- , 0), in conclusion, the motor can maintain self-stabilization mechanism if \u03b8 ranges from to , but lose that if \u03b8 ranges from to or to . Therefore, in the practical application, the speed command should not increase too fast, otherwise the motor will enter the region where the motor couldn\u2019t maintain the self-stabilization mechanism" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0000050_expat.2019.8876524-Figure5-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0000050_expat.2019.8876524-Figure5-1.png", + "caption": "Fig. 5. a) Product insertion component; c) Packaging block component, b) and d) 3D models after printed and assembled.", + "texts": [ + " These supports were composed of six main components designed with the possibility of being easily disassembled, as can be verified in Fig. 4. The product placement component served as an auxiliary accessory for the pneumatic cylinder rod, which made the product placement inside the package, when it was needed in the required position, and add the closing system of the left side flap. The auxiliary lock packaging component served as an auxiliary accessory for the rod, of the pneumatic cylinder which made the block of the packaging, when it was in the required position. Fig. 5 shows the components developed with Solidworks\u00ae and the 3D model printed and assembled. For this project a mono-stable 5/2 solenoid valve was used, with a spring that holds it in its resting or closed position. The electric sensors allowed the position monitoring of the pneumatic cylinders in their several operating states. For this purpose an end-of-travel magnetic sensor was chosen, compatible with the pneumatic cylinder considered. C. 3D Printing Technology The 3D printing technology resides in the process of creating layer-to-layer objects using various materials until a three-dimensional shape is obtained [6]" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0002759_b978-0-12-804560-2.00011-0-Figure4.5-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0002759_b978-0-12-804560-2.00011-0-Figure4.5-1.png", + "caption": "FIGURE 4.5 Model of a planar reaction mass pendulum. The two mass points slide under the action of force fr , keeping thereby equal distance r from the center point R. Thus, the moment of inertia Mr2 is controllable via fr . Note that fr and mC are redundant control inputs with regard to the inertia moment mr2\u03c6\u0308 that ensures the inertial coupling with passive coordinate \u03b8 .", + "texts": [ + " The variable centroidal inertia model matches the variable, posture-dependent centroidal dynamics of a humanoid robot to a greater extent. This can be demonstrated with the so-called reaction mass pendulum (RMP) model [67,68,130]. The RMP is an underactuated system comprising a pendulum with a massless telescopic leg and a \u201creaction mass\u201d assembly at the tip of the pendulum. The mass assembly represents the total (constant) mass and the (varying) aggregate inertia of the humanoid in a concentrated form at the CoM. A planar version of the RMP model is shown in Fig. 4.5. The mass assembly consists of two constant mass points with variable relative position r , symmetric w.r.t. the rotation joint R at the tip of the leg (the \u201chip\u201d joint). Note that this joint is actuated via torque mC . The inertia moment around R is adjustable by varying the mass distance r . The general form of the equation of motion is the same as (4.19). We have[ Mp Mpa MT pa Ma ][ q\u0308p q\u0308a ] + [ cp ca ] + [ gp ga ] = [ 0 Fa ] , (4.26) where qp = \u03b8, qa = [ l r \u03c6 ]T denote the generalized coordinates, Fa = [ fs fr mC ]T stands for the generalized force, Mp = M(l2 + r2), Mpa = [ 0 0 Mr2 ] , Ma = diag [ M M Mr2 ] are the inertia matrix components, cp = 2M((ll\u0307 + rr\u0307)\u03b8\u0307 + rr\u0307\u03c6\u0307), ca =[ Ml\u03b8\u03072 Mr(\u03b8\u0307 + \u03c6\u0307)2 2Mrr\u0307(\u03b8\u0307 + \u03c6\u0307) ] are the nonlinear velocity-dependent forces, and gp = \u2212Mgl cos \u03b8 , ga = [ Mg sin \u03b8 0 0 ]T are the gravity-dependent forces" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0000419_s40313-019-00558-8-Figure1-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0000419_s40313-019-00558-8-Figure1-1.png", + "caption": "Fig. 1 Geometry of bubble between two elastic walls", + "texts": [ + " This property is employed in ultrasound radiography application ranges from stroke detection to blood volume and perfusion measurement (Meyer et\u00a0al. 2003). The rest of this paper is organized as follows. Section\u00a02 includes the model description. Section\u00a03 is devoted to controller design. The stability of the closed-loop system is guaranteed by the Lyapunov theory and also the effect of controller parameters on the performance of the closed-loop system is investigated. The simulation results including the case of parametric uncertainty are also presented. The geometry of the bubble between two elastic walls is shown in Fig.\u00a01. It is assumed that the bubble is placed between two elastic walls of finite thickness and also the bubble radius (R) is small compared to the distance of walls (d) . The space between the walls is filled with an incompressible liquid. The method of image sources and Lagrangian formalism are applied to obtain the dynamic of bubble\u2019s radial behavior between two elastic walls as follows (Doinikov et\u00a0al. 2012): where R,R0 and R\u0307 are the radius of the bubble, initial radius of bubble and bubble interface velocity, respectively" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0002541_iros45743.2020.9340742-Figure5-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0002541_iros45743.2020.9340742-Figure5-1.png", + "caption": "Fig. 5: Approach 2: Current pose of phone as the shared reference frame. The phone is mounted on the robot and its instantaneous coordinate frame links the virtual object location to its actual real-world coordinate on the map.", + "texts": [ + " Approach 2: Instantaneous pose of phone as the shared frame of reference In this approach, instantaneous pose of the phone is used to estimate the virtual object location on the map of the area (OGM). The phone is mounted on the robot with a fixed frame of reference in the robot\u2019s coordinate system in ROS (see Fig. 2b). The robot is localized on the map using adaptive Monte Carlo localization (AMCL), and the pose of the phone relative to the map is calculated based on the robot\u2019s pose. The transformation between map and phone frame is provided by the tf messages published by ROS. The phone frame, which is shared by both the robot and AR scene (see Fig. 5), is used to estimate the corresponding location of virtual objects in the OGM. The real-world location of the phone where the AR application is launched serves as the global coordinate frame of the phone in the AR scene. The AR application determines the pose of virtual objects based on this global frame of the phone. For computing the virtual object location relative to the robot\u2019s map, first the virtual object location in the global frame (globalPobject) needs to be transformed into the phone\u2019s current frame (pcPobject)", + " Finally, the absolute location of virtual object on the map (mapPobject) is calculated by transforming the virtual object location in the phone\u2019s coordinate frame in the robot system to the map\u2019s 4448 Authorized licensed use limited to: Central Michigan University. Downloaded on May 14,2021 at 18:09:26 UTC from IEEE Xplore. Restrictions apply. coordinate frame. The detailed steps are given below and the transformations in the AR system are shown in pink color and in robot system as dotted lines in Fig. 5. Step 1: Compute the virtual object location in the phone\u2019s current coordinate frame in the AR system (pcPobject)AR = ( globalTpc ) ( globalPobject ) . (1) Step 2: Transform the virtual object in the current left-handed coordinate system (LHS) of phone in the AR system to that of the right-handed coordinate system (RHS) of phone in the robot system. Q = (RX(90 \u25e6)) (pcPobject)AR , (2) (pcPobject)robot = EQ, (3) where RX denotes the rotation matrix corresponding to a rotation about the X-axis and E is the elementary row operator to swap the coordinate axis for LHS to RHS coordinate transformation given by E = 1 0 0 0 0 1 0 1 0 " + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0003943_icra.2014.6907115-Figure4-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0003943_icra.2014.6907115-Figure4-1.png", + "caption": "Fig. 4. Simulation results when the robot is pushed from behind on the center of the torso with a force of 150 N for 0.1 seconds. The dark blue circle indicates the location of the disturbance force, the orientation is perpendicular to the plane shown, out of the page.", + "texts": [ + " These simulations consist of three classes of disturbance: being pushed while standing still, dynamic walking, and being pushed while walking. In all of these simulations, the robot was initialized to begin the simulation as if it had just taken a step while walking (i.e. with one foot ahead and to the side of the other). The first simulation was used to verify that the controller was acting as expected in the presence of an external disturbance while the robot was stationary. A simulated force of 150 N was applied to the center of the torso in a direction out of the page for 1/10 second. The series of images in Figure 4 show the progression of the simulation at the transitions between the various states in the controller\u2019s state machine. The second simulation included a desired COM velocity, to dictate how fast and in what direction the robot should attempt to walk. In this simulation, the robot is statically stable until it reaches the Swing state, as the Lift state is augmented to move the swinging foot in the direction of desired motion before switching to tracking the FPE in the Swing state. A series of images, in Figure 6, shows the simulation at the transition points between states for one step of the walking cycle" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0002861_978-3-030-55889-5-Figure8.5-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0002861_978-3-030-55889-5-Figure8.5-1.png", + "caption": "Fig. 8.5 Here, the Lagrange multiplier is the Gateaux derivative j u of the non-linear functional J (u)", + "texts": [ + "105) and so the Lagrangemultiplier is identical with the vector g of the influence function. According to (8.105) the vector \u03bb is the gradient of the functional J (u) with respect to the right side f [5], \u2202 J \u2202 fi = \u03bbi , (8.106) and its plot serves as a sensitivity plot of the functional, see section 3.32. If the functional J (u) is nonlinear, the equations read K u = f K \u03bb = ju , (8.107) where jui := \u222b 1 0 J \u2032(u \u2212 s u;\u03d5i ) ds , (8.108) and J \u2032 is the Gateaux derivative of J (u) at the point u, see Fig. 8.5. So, the vector g = \u03bb of the functional depends on u, depends on where we are on the equilibrium path. This way \u201cinfluence functions\u201d can also be formulated for nonlinear problems, more precisely for the tangential stiffness matrix at a particular equilibrium point u, if one moves away from this point in \u201ctangential direction\u201d. This is also how goal oriented refinement methods have been applied successfully to nonlinear problems; for details, see [6, 7]. Birkh\u00e4user Basel Boston Berlin Chapter 9 Addenda Before we concentrate on details, let us briefly describe the approach to the equations of statics in this book" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0003903_aps.2014.6904982-Figure1-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0003903_aps.2014.6904982-Figure1-1.png", + "caption": "Fig. 1: Cross section of the bearing cartridge highlighting the modified cage and stationary interrogator.", + "texts": [ + " A frequency sweep centered around the resonant frequency of the cage circuit is fed to the receiver coil. The AC magnetic field produced by the current of the receiver coil is picked up by the cage circuit. At resonance, there will be maximum power transfer from the receiver coil to the cage circuit. This can be observed on a spectrum analyzer as a dip in S11 parameters the receiver coil. The cage of the angular contact ball bearing cartridge used in the turbocharger is made of polyetheretherketone (PEEK) which has exceptional material properties in terms of strength and thermal properties. Fig. 1 shows the construction of the sensor components and placement in the bearing cartridge. The turbocharger is powered by an industrial sized compressor and lubricated by an independent oil circulation system. Fig. 2 shows an overview of the test rig and the piping configuration. The temperature of the air that is being compressed by the turbocharger increases significantly creating a temperature gradient across the turbocharger from higher temperatures on the compressor side to lower temperatures on the turbine side" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0000713_j.matpr.2020.01.060-Figure1-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0000713_j.matpr.2020.01.060-Figure1-1.png", + "caption": "Fig. 1 Geometry of spherical shell element.", + "texts": [ + " [4] used first order shear deformation theory to carry free vibration analysis of functionally graded porous doubly curved shells. Shi et al. [5] have proposed a free vibration analysis of the functionally graded double curved shallow shell structures with general boundary conditions. 2. Mathematical Formulation 2.1 Spherical Shell Under Consideration A spherical shell having width a along x direction, breadth b in y direction and radii of curvature R1 and R2. The geometry of the spherical shell element is shown in Fig. 1. Based on certain kinematical assumptions, the following is the displacement field assumed for the present theory. ( ) ( ) ( ) ( ) ( ) ( ) ( ) ( ) ( ) ( ) ( ) ( ) 3 5 0 0 2 4 1 3 5 0 0 2 4 2 2 4 0 2 4 4 161 3 5 4 161 3 5 4 161 1 x x y y z z wz z zu x, y,z = u x, y - z + z x, y z x, y R x h h wz z zv x, y,z = v x, y - z + z x, y z x, y R y h h z zw x, y,z = w x, y + x, y x, y h h \u03c6 \u03c8 \u03c6 \u03c8 \u03c6 \u03c8 \u2202 + \u2212 + \u2212 \u2202 \u2202 + \u2212 + \u2212 \u2202 \u2212 + \u2212 (1) where, u, v, w are the displacements in x, y, z directions respectively" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0003562_icma.2014.6885964-Figure1-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0003562_icma.2014.6885964-Figure1-1.png", + "caption": "Fig. 1. The simplified structure of triple inverted pendulum", + "texts": [], + "surrounding_texts": [ + "A triple inverted pendulum, a highly sensitive system, is stabilized by using the VGLQR method, which demonstrates the potential of the control scheme. As shown in Figure 3, the triple inverted pendulum is consisted of a cart, rod1, rod2 and rod3, which are linked by hinges. The cart can move freely on the linear orbit, and the rods can rotate freely in plumb plane. Let the clockwise angle and moment in Figure 3 be in positive direction. And let x be the displacement of the cart, u the acceleration of the cart, \u03b8i the angle between rodi and vertical direction, Oi and Gi the linking point and centroid of rodi respectively, mi the mass of rodi, Ji the moment of inertia of rodi around Gi, li the distance from Oi to Gi, Li the length of rodi, ci the fricative coefficient of rodi around Oi, (i = 1, 2, 3). The equilibrium point of the triple inverted pendulum is that the rod1, rod2 and rod3 are all in upright position and the cart\u2019s position is designated as zero point. The mechanical parameters are described in Table 1. The model of the triple pendulum is represented as follow x\u0308 = u, (35) b1\u03b8\u03081 + a2L1 cos(\u03b82 \u2212 \u03b81)\u03b8\u03082 + a3L1 cos(\u03b83 \u2212 \u03b81)\u03b8\u03083 = (\u2212c1 \u2212 c2)\u03b8\u03071 + a2L1 sin(\u03b82 \u2212 \u03b81)\u03b8\u030722 + c2\u03b8\u03072 +a3L1 sin(\u03b83 \u2212 \u03b81)\u03b8\u030723 + a1g sin \u03b81 \u2212 a1 cos \u03b81x\u0308, (36) a3L1 cos(\u03b83\u2212\u03b81)\u03b8\u03081+a3L2 cos(\u03b83\u2212\u03b82)\u03b8\u03082+b3\u03b8\u03083= \u2212a3L1 sin(\u03b83\u2212\u03b81)\u03b8\u030721\u2212 a3L2 sin(\u03b83\u2212\u03b82)\u03b8\u030722+ c3\u03b8\u03072 \u2212c3\u03b8\u03073 + a3g sin \u03b83 \u2212 a3 cos \u03b83x\u0308, (37) a3L1 cos(\u03b83\u2212\u03b81)\u03b8\u03081+a3L2 cos(\u03b83\u2212\u03b82)\u03b8\u03082+b3\u03b8\u03083= \u2212a3L1 sin(\u03b83\u2212\u03b81)\u03b8\u030721\u2212 a3L2 sin(\u03b83\u2212\u03b82)\u03b8\u030722+ c3\u03b8\u03072 \u2212c3\u03b8\u03073 + a3g sin \u03b83 \u2212 a3 cos \u03b83x\u0308, (38) Taking x = [x, \u03b81, \u03b82, \u03b83, x\u0307, \u03b8\u03071, \u03b8\u03072, \u03b8\u03073], the (35)-(38) can be rewritten as the state-space form x\u0307 = f(x, u), y = x. (39) where f = [f1, f2, \u00b7 \u00b7 \u00b7 , f8]T . In the numerical experiment, we take Q = diag(10, 500, 500, 500, 0, 0, 0, 0), R = 1, and apply the VGLQR and LQR method to stabilize the triple inverted pendulum, and the initial state is taken as: x = 0, \u03b81 = \u03b83 = 0.0536(rad) and \u03b82 = \u22120.0536(rad), which is a difficult condition for control. The simulation results are shown in Figure 2(a)-(e). As can be seen from Figure 2(a)-(e), the VGLQR controller has better adaptiveness and robustness. In addition, The effectiveness of the VGLQR method has also been demonstrated by physical experiment. A triple inverted pendulum is successfully stabilized by using the present control scheme. The real-time responses of one physical experiment are shown in Figure 2(f)-(j). Remark Because the friction and other features of our nonlinear simulation model can approximate but are unable to duplicate exactly the physical system\u2019s behaviour, the real-time curves fluctuate in a small range and can not converge to zero(see Figure 2(f)-(j)). In practice, since the uncertain factors always exert on the system during the real-time control process, stabilizing the three pendulums around the unstable equilibrium state is a dynamic process, and the cart and rods fluctuate near the equilibrium point." + ] + }, + { + "image_filename": "designv11_71_0000632_s1064230719060029-Figure1-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0000632_s1064230719060029-Figure1-1.png", + "caption": "Fig. 1. A rotary platform with two degrees of freedom.", + "texts": [ + "OI: 10.1134/S1064230719060029 The problem considered here arose in connection with the development of a control system for a rotary platform that is installed on an orbiting spacecraft and serves to accurately orient the object fixed in it. The general view of the platform is shown in Fig. 1 [1]. The main component of the platform design is a twodegree suspension consisting of two rigid frames (external and internal) rotating around mutually perpendicular axes. The external frame rotates around an axis rigidly connected to the spacecraft, and the internal frame rotates around an axis rigidly connected to the external frame. The object is fixed in the internal frame and its orientation is controlled by a rotary platform. The frame\u2019s rotation is controlled by two independent electric drives with DC motors" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0003670_2014-01-1727-Figure5-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0003670_2014-01-1727-Figure5-1.png", + "caption": "Figure 5.", + "texts": [ + " flywheel), and an IVT KERS - Kinetic Energy Recovery and Storage - a regenerative braking system GWT - The vehicle weight plus the weight the vehicle is designed to carry APPENDIX The Mathcad analysis provides a set of parameters that accomplishes all of the stated objectives (note: all frictional losses are ignored): The GWT of the vehicle: The vehicle will be accelerated from zero to: The final kinetic energy of the vehicle is: (5) Tire OD: Calculate RPM of wheel at vf: (6) Time to accelerate from vo to vf: (7) Force required on the vehicle: (8) Torque on the tire: (9) Using torque and time we can calculate the power required: (10) The tire translates linear mass to angular inertia at the tire: (11) Angular velocity, momentum, torque, and inertia are delivered to the KERS system through the vehicle differential with a typical gear reduction: Gear reduction: From here, we will refer to the input of the KERS system as \u2018sub-c\u2019: The inertia is reflected as: (12) The final angular velocity is: (13) The torque delivered from the KERS system is: (14) Set the output velocity (\u03c9c) as a function of time: (15) The vehicle momentum reflected to the output of the system: (16) The general configuration of the Gramling/Martin [9] system is shown schematically in Figure 5. Variable-Inertia Flywheel Specification: Mass of flyweight(s): MSa := 30\u00b7kg; MSb := 30\u00b7kg Inertia of the flyweights about its center of mass (cm): Iacm := .0049\u00b7m2\u00b7kg; Ibcm := .0049\u00b7m2\u00b7kg Add body (fixed) inertia: Iaby := .764\u00b7m2\u00b7kg; Ibby := .764\u00b7m2\u00b7kg Set the initial moment of inertia for FWa and for FWb: Iao= 4\u00b7m2\u00b7kg; Ibo= 1\u00b7m2\u00b7kg Ratio FWa:diff - will always be 1:1 for this worksheet: na := 1 Ratio FWb:diff - initial velocity for FWb should be increased over the FWa velocity. An objective of the selection of certain parameters is to improve efficiency" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0000338_iecon.2019.8926794-Figure15-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0000338_iecon.2019.8926794-Figure15-1.png", + "caption": "Fig. 15. The experimental field Fig. 16. Initial position of Angle detection", + "texts": [ + " A library for using DMP is released on GitHub, and it is downloaded to this node and introduced to ROS. This node distributes the topic named \u201c/mpu6050 msg\u201d to \u201cservo node.\u201d \u2022 servo node (Self-made node) \u2013 \u201cServo node\u201d subscribes to the topic that is the output of an \u201cmpu 6050 node\u201d YAW axis, and the servo motor is controlled using it. In this experiment, a library group for using a MPU6050 DMP was made available to ROS. The purpose was to place the car on a rotating table and to measure angles using DMP. The experimental field used in this experiment is shown in Figure 15. In order to be able to measure angles using this field, a plate for positioning was attached at the top of the car. After that, a rotating table was aligned with this field and placed. At this time, as shown by the circle in Figure 16, the plate and this field were aligned vertically. This was the initial position of the car. After that, the car was measured by rotating 90\u00b0 to the left and right. These ROS nodes were made open, and the open source is managed by the public repository of GitHub. Anyone can be downloaded from the following URL" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0002816_thc-2010-0566-Figure16-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0002816_thc-2010-0566-Figure16-1.png", + "caption": "Fig. 16. Spherical storage tank.", + "texts": [ + " In simple external loading situation like tension, compression, torsion or bending we can calculate the maximum internal stresses using Eqs (1), (4), (6), (7), (9) and (10) and if these stresses are higher than the allowed stresses then we know that the structure will fail. In real situation two or more of these external loading takes place simultaneously, what happens then? One such simple example of combined loading is a spherical liquid storage tank, for example the heart. The tank is pressurized and if we look at a small square piece of material on the tank surface we can see that the square experiences tensile stresses on both sides as shown in Fig. 16. To handle these situations with combined loading we use Mohr\u2019s circle. Most of the materials have to be analysed 3-dimensionally but since they are geometric in nature, using the symmetry we can convert them to 2 dimensions like we did for the spherical storage tank (Fig. 16) Mohr\u2019s circle is a plot between the normal stress (\u03c3) on the horizontal or x-axis and the shear stress (\u03c4 ) on the vertical or y-axis (Fig. 17). Tensile normal stresses are positive and compressive are negative, for the shear stress we follow a sign convention shown in Fig. 18 where a very small volume element inside the bulk of the material is shown. Coming back to Fig. 17, the position of the center of the circle c is the average of the two normal stresses (\u03c3x and \u03c3y and reference point a can be plotted with \u03c4 xy on the y-axis and \u03c3x on the x-axis" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0000603_iccas47443.2019.8971558-Figure1-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0000603_iccas47443.2019.8971558-Figure1-1.png", + "caption": "Figure 1. The application of the TakoBot 2.", + "texts": [ + "00 \u24d2ICROS Authorized licensed use limited to: Auckland University of Technology. Downloaded on October 25,2020 at 08:37:21 UTC from IEEE Xplore. Restrictions apply. 219 results. The new prototype supposes to improve robot dexterity and reachability capabilities. In chapter II will explain the concept design, III chapter will explain kinematic kinetic formulation, IV chapter will present experiments and simulation results. II. CONCEPT DESIGN A. Application Analysis The intended application of the robot is shown in Fig.1. The manipulator is designed to be used in the agriculture field for harvesting, weeding and inspection operations if necessary. The intended working environment is highly constrained and by the time workspace will be different because of growing plants. To perform necessary tasks the robot should have the following functions: (1) Flexible dexterity: the robot should work in the confined workspace (2) Obstacle avoidance capability: the robot should avoid contact with solid surfaces and not collide. (3) Reachability: the robot should get the required position in spite of narrow space and obstacles. (4) Safety: the robot should be safe enough to avoid breaking any sticks of the plant. (5) Portability: the robot should be compact to be used as a tool for farming CNC platform (Fig.1). B. System Design Based on application analysis, the robot should have a slender structure with redundant DOFs. After consideration of previous TakoBot 1 prototype capabilities, we made changes in the design and actuation system as well. After numerous experiments with TakoBot 1, demonstrated some limitations related to the design. For example, a robot could not perform torsional motion during the work which led to the accumulation of strain energy inside of manipulator. Secondly, in case of bending it could not perform a pure bending shape" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0003878_s11434-014-0376-5-Figure6-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0003878_s11434-014-0376-5-Figure6-1.png", + "caption": "Fig. 6 (Color online) The skeletal structure of the forelimb and the plane five-link mechanism [14]", + "texts": [ + " All major bones of the leg are supposed to be in the longitudinal plane of the horse; 4. The mass of the leg, compared with the horse weight, is too light and ignored. According to the assumptions above, the skeleton-joint leg structure could be regarded as the link-revolute pair structure. Taking the skeletons as rigid bodies and joints as the revolute pairs with active drives, the forelimb and hind limb of the horse turn into a plane open-chain five-link mechanism. The corresponding relations are shown in Table 2 and Table 3. In Fig. 6 and Fig. 7, A is the forelimb foothold on the ground and B is the hind limb foothold. H represents the height of the shoulder joint in forelimb and the hip joint in hind limb. L refers to the horizontal distance from foothold A to the shoulder joint, or from foothold B to the hip joint. The forelimb and hind limb both can be equivalent to plane open-chain five-link mechanisms. Although the length of links and the postures of two mechanisms are different, the characteristics of two mechanisms are similar" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0001455_access.2020.3009089-Figure1-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0001455_access.2020.3009089-Figure1-1.png", + "caption": "FIGURE 1. Schematic diagram of the active balancing system.", + "texts": [ + " Then, based on the estimated signals, the barrel system was transformed into the pure integration chain form. Hence, the sliding mode control technology can be directly used. To attenuate the hydraulic uncertainties, a sliding mode controller was designed as the position tracking controller based on the new state variables. The closed loop system was proven to be exponentially stable in the sense of Lyapunov theory. II. DYNAMIC MODELS AND PROBLEM FORMULATION A. DECRIPSION OF THE ACTIVE BALANCING SYSTEM In this section, a simple but effective hydraulic active balancing system is designed. As shown in Figure 1, the active balancing system is mainly comprised of three accumulators and a two-stage hydraulic cylinder whose structure is the same as the actuator of the barrel servo system. VOLUME 8, 2020 131371 The three accumulators are placed in parallel, and the gas pressure of each accumulator is carefully designed such that each of them works mainly at specific barrel angle ranges. The gravitational torques within the angle 0-40 degrees, 40-60 degrees and 60-75 degrees are mainly balanced by one of the three accumulators" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0001752_022053-Figure2-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0001752_022053-Figure2-1.png", + "caption": "Figure 2. Healthy and Fault Gear Model.", + "texts": [ + " At last, the effects of dynamic load on healthy and pitted gear tooth were analyzed and the percentage of decrease in mesh stiffness was evaluated for the defected gear tooth pair. In the current investigation, we considered same physical parameters for both the driver and driven gear, which are listed in table 1. The gear quality has been taken as per DIN 3962-1 is DIN8 [12]. According to the Ref. [11], the gear tooth pits are designed in circular shape in order to enhance this work. The healthy and fault gear are modeled using Solidworks software as shown in figure.2. In faulty gear, all the pits are designed with 2 mm diameter and 0.8 mm depth. 3rd International Conference on Advances in Mechanical Engineering (ICAME 2020) IOP Conf. Series: Materials Science and Engineering 912 (2020) 022053 IOP Publishing doi:10.1088/1757-899X/912/2/022053 In [13], the gear tooth involute profile was generated from base circle and continued as a straight line up to the root circle. In addition to this the gear tooth fillet curve was ignored in order to replicate the single tooth as a cantilever beam as shown in figure 3" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0000967_0954406220916504-Figure9-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0000967_0954406220916504-Figure9-1.png", + "caption": "Figure 9. ESH chain of full vehicle model: (a) double-wishbone suspension mechanism and (b) multi-link suspension mechanism.", + "texts": [ + " The sum of three displacements in equation (16) associated with the roll motion generated by cornering is defined as D\u0302r such that D\u0302r \u00bc q3S\u03023 \u00fe q4S\u03024 \u00fe q5S\u03025. The infinitesimal displacement D\u0302r may be rewritten as D\u0302r \u00bc qrS\u0302r, where the unit line vector S\u0302r represents the instantaneous screw axis of the roll motion and qr is the magnitude of D\u0302r. Kinestatic analysis of full vehicle model When a car is steered round a corner, cornering forces caused by cornering act on the vehicle body to generate the strut spring forces. Referring to Figure 9, the spring force can be considered to produce the input force H through the virtual H-joint S\u0302H of the serial ESH chain. Consequently, the wrench w\u0302H is generated through the ESH chain and can be determined from equation (6) w\u0302H \u00bc H r\u0302 T HS\u0302H r\u0302H \u00f017\u00de where r\u0302H represents the column vector of the reciprocal Jacobian of the ESH chain corresponding to S\u0302H. By definition, r\u0302H is determined as a unit line vector which is perpendicular to the ground and passes through S-joint (see Figure 9). If the friction is negligible, the work done by the strut spring force is equal to the work produced by the virtual H-joint qH H \u00bc lsfs \u00f018\u00de From equations (15) and (18), the force H produced by the virtual H-joint is related to the strut spring force fs by H \u00bc fsr\u0302 T s S\u0302H \u00f019\u00de Substituting equation (19) into equation (17) yields w\u0302H \u00bc fs r\u0302 T s S\u0302H r\u0302 T HS\u0302H r\u0302H \u00f020\u00de This wrench w\u0302H can be considered to be acting on a virtual spring. Since r\u0302H is reciprocal to the unit line vectors S\u03021, S\u03022 and S\u03026 in equation (16), the small displacement qH of the virtual H-joint in ESH chain is obtained from equation (3) qH \u00bc 1 r\u0302 T HS\u0302H r\u0302 T H D\u0302r \u00f021\u00de The small change of the strut spring can be found by equating equation (21) with equation (15) ls \u00bc r\u0302 T s S\u0302H r\u0302 T HS\u0302H r\u0302 T H D\u0302r \u00f022\u00de For a strut spring with the stiffness ks, the relation fs \u00bc ks ls can be substituted into equation (20) to give w\u0302H \u00bc K lr\u0302H \u00f023\u00de where K \u00bc ks r\u0302 T s S\u0302H r\u0302 T HS\u0302H " + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0000045_ab5165-Figure3-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0000045_ab5165-Figure3-1.png", + "caption": "Figure 3. Illustrated fabrication process of the PDMS lens (a) fabricate a lens mold using concave mirror, (b) pour the PDMS to the mold, (c) geometric model of the PDMS lens, (d) pill off the PDMS lens from the mold, (e) bond the lens with the connecting brackets, (f) fabricated PDMS lens", + "texts": [ + " In the soft plano-convex lens shape, R is the radius of the curvature of the convex lens, a is the radius of the lens, n is the refractive index of the lens material, h is the height of the lens, and d is the thickness of the plane lens part, as noted in figure 2. Assuming R>>h in the proposed plano-convex lens, the focal length F of the lens is derived as (1) By assuming the constant volume and R>>h, the change in the focal length is then derived as (2) The plano-convex soft lens is made of PDMS silicone (SYLGARD 184 from Dow Corning). The silicone was prepared by mixing the base and the curing agent in 10:0.4 ratio. It is poured over lens mold. The mold for the PDMS lens part is shown in figure 3(a). It consists of a spherical concave mirror (radious of curvature : 35mm, diameter : 70mm) and a assembly part. The surface of the mold is coated with ER-200 Release Agent from SMOOTH-ON, in order to decrease adhesion between the PDMS lens and the mold. After, degassing in a vacuum chamber for 20 minutes at room temperature and thermal cross-linking at 333K for 8 hours, the lens structure is peeled off from the mold. Convex part of the fabricated lens has a diameter of 35mm and a height of 2.2mm, and plane part has a diameter of 45mm and a thickness of 1" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0002769_ijhvs.2019.102683-Figure3-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0002769_ijhvs.2019.102683-Figure3-1.png", + "caption": "Figure 3 Three-dimensional point contact tyre model", + "texts": [ + " The radial and torsional elasticity of each elastic torsion bar is modelled assuming linear stiffness and damping along the lateral (Ky, Cy) and vertical (Kz, Cz) axis and torsional stiffness and damping (Kt, Ct), as seen in Figure 2(c). The effective vertical, lateral and roll spring rates of the torsio-elastic suspension, however, are functions of the kinematics of the lateral suspension linkage. The tyre is modelled as a three-dimensional point contact model, as described in Crolla et al. (1990) and shown in Figure 3. This simple model is considered appropriate and computationally efficient for relative analyses of three suspension configurations. Non-linear contact, however, is considered to incorporate potential loss of tyre\u2013terrain contact. The equations of motion for the vehicle model can be expressed in the matrix form as: [ ]{ } [ ] 0,s d TM X F F F+ + + = (1) where ( ) ( ),[ ] diag , , , , ; , , , ; , , 1, , 4s s x y uj uj uj i i iM M m m I I m m I j f r m m I i = = = (2) is (23 \u00d7 23) diagonal matrix of model masses and mass moments of inertia, with M being the total vehicle mass, Ix and Iy are the roll and pitch mass moments of inertia of the sprung mass, and Iuj(j = f, r) is the roll mass moment of inertia of the front and rear unsprung masses" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0001734_s12206-020-0834-8-Figure6-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0001734_s12206-020-0834-8-Figure6-1.png", + "caption": "Fig. 6. Nao-H25: (a) physical prototype; (b) rob model.", + "texts": [ + " 3, a two-stage optimization method is established in this work, which will be evaluated in the following parts. In this section, to reveal the relationship between energetic cost and step parameter configuration, the energetic performance is studied by using the two-stage optimization. Besides, compared with other works, the energetic benefits of the proposed approach would be demonstrated. The basic parameters, as listed in Table 1, are set based on the physical characteristics of Nao-H25 robot, which contains 5 joints in each leg. The robot model is demonstrated in Fig. 6. The algorithm parameters and step parameter configurations are listed in Table 2. Particularly, since we do not discuss the energetic cost consumed by the start motion and stop motion, the energetic performance when walking from 50 mm to 550 mm (totally 500 mm) is analyzed here. Given the target distance (D), we first computed the energetic costs by using the PDC-BM. Using different step parameters listed in Table 2, the optimal velocity coefficients (\u03b2opt) and the corresponding minimal energetic costs (Eopt) are demonstrated in Figs" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0003275_aim.2015.7222530-Figure10-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0003275_aim.2015.7222530-Figure10-1.png", + "caption": "Figure 10. Configuration3: 7 segments placed in the region 0\u00b0 < y < 180\u00b0 in the best possible configuration.", + "texts": [], + "surrounding_texts": [ + "A 3-channel gauss meter (Lakeshore-Model 460) was used to measure the magnetic flux density. A torque gauge (HTG2-40 made by IMADA) with its respective torque sensor held the IPM at the centre of the system. The IPM was connected to the torque sensor via a green plastic connector that was prototyped using a 3D printer. The torque sensor and the probe tip of the gauss meter were mounted on plastic holders which were also fabricated with a 3D printer. Both the torque sensor and the probe tip of the gauss meter can be moved along the X and Z axes and the arrays of magnets can only be moved along the Y axis. These displacements are controlled by a micromanipulation system constructed of XYZ stages as shown in Fig. 13." + ] + }, + { + "image_filename": "designv11_71_0001225_s00170-020-05332-8-Figure12-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0001225_s00170-020-05332-8-Figure12-1.png", + "caption": "Fig. 12 (a) The Rubik robot designed in Webots simulator. (b) The simplified individual model. (c) All homogeneous robots are aggregated into a rectangular grid pattern", + "texts": [ + " As mentioned above, the proposed algorithm solves the problem that hollow shapes are difficult to form. Based on the selfassembly algorithm, a variety of simulated and hardwarebased experiments are demonstrated. We had some physical robots Rubik available to form small shapes. To prove the flexibility of the extended algorithm, we used a simulator to perform experiments that require more robots than we currently have. Driven by the needs of the self-assembly algorithm, an omnidirectional vehicle robot called Rubik is designed in the Webots simulator as shown in Fig. 12a. The robot integrates sensors and actuators such as infrared sensors, ultrasonic sensors, and servo motors. It is capable of the following functions: 1. The robot adopts a symmetrical cubic structure to ensure the tightness of the lattice-based system. 2. It can detect the presence of nearby robots through distance sensors. 3. It can send and receive signals through infrared sensors to interact with nearby robots. 4. It has basic information storage and computation capabilities. We have done the simulations by the Webots simulator. We simulated with a computer with processor Intel Core i76700K, Nvidia GeForce GTX 1060, and 6GB of memory. Considering the high dependence of software on the hardware environment, the simulation rate is negatively correlated with the number of individuals, so the simulation rate is significantly reduced under large-scale robots. Thus, when the number is large, we use a simplified model (see Fig. 12b) to reduce the resource consumption of the software. All homogeneous robots are aggregated into a rectangular grid pattern. The centerto-center distance of adjacent robot is 12 cm, and the moving speed of each robot is set to 1 cm/s as shown in Fig. 12c. All robots are given an identical program including a desired hollow shape and a self-assembly algorithm. Referring to the stratified mechanism, all individuals are initially aggregated and arranged in a regular rectangular grid pattern. For the two types of hollow shapes, different simulation experiments are carried out to verify the feasibility of the self-assembly algorithm. In the simulation of the hollow shape while its hole region located outside the initial shape, the state update method, follow-up motion method, and extended motion rules are used" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0000045_ab5165-Figure1-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0000045_ab5165-Figure1-1.png", + "caption": "Figure 1. Illustration of the proposed focal-tuning mechanism", + "texts": [ + " In order to provide six step piecewise control of the tunable lens, the locking mechanism is composed of an internal gear and an SMA driven latch. The design, fabrication process and experimental test for a prototype are described in the paper. For focal length tuning, the tunable lens module uses the proposed SMA driven iris diaphragm. In the focus tuning system, SMA spring is used due to its unique characteristics such as large deformation, high power density, lightweight and compact size [31]. When the SMA spring is actuated, by applying the electrical signal, the ring starts to rotate in the counter-clockwise direction, as shown in figure 1. Simultaneously, connecting brackets, which encircle the PDMS lens, slide along the slope and pull the lens outward in the radial direction which causes focal length change. Figure 2. Shape(Illustration) of the plano-convex lens To calculate the change in the focal length, the lensmaker's formula is used. In the soft plano-convex lens shape, R is the radius of the curvature of the convex lens, a is the radius of the lens, n is the refractive index of the lens material, h is the height of the lens, and d is the thickness of the plane lens part, as noted in figure 2" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0000967_0954406220916504-Figure1-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0000967_0954406220916504-Figure1-1.png", + "caption": "Figure 1. Spatial serial mechanism.", + "texts": [ + " Kinestatic relations of spatial mechanism A kinematic joint (or, pair) can be modelled by use of a unit twist S\u0302 which is expressed in the Plu\u0308cker\u2019s axis coordinates as S\u0302 \u00bc r s\u00fe hs s , where s is the unit direction vector of the axis of the joint, r is the position vector from the origin O to the joint axis and h is a pitch. A prismatic joint (P-joint) can be represented by a free vector as S\u0302i \u00bc si 0 2 R6 1, where si \u00bc X Y Z T is the unit direction vector of the P-joint. A revolute joint (R-joint) can be expressed by a line vector with zero pitch as S\u0302i \u00bc ri si si , where si is the unit direction vector of the joint axis. In Figure 1, the twist of the last link of an ndegrees-of-freedom (DOF) (14n46\u00de spatial serial mechanism10 can be expressed by T\u0302 \u00bc _q1S\u03021 \u00fe \u00fe _qnS\u0302n \u00bc Js _q \u00f01\u00de where _q \u00bc _q1 _qn T 2 Rn 1 is the joint velocity vector and Js S\u03021 S\u0302n h i is the screw-based Jacobian of the serial mechanism. The twist can be viewed as T\u0302 \u00bc lim t!0 D\u0302 t for an infinitesimal time interval t, where D\u0302 denotes an infinitesimal displacement of the last link of the serial mechanism. From equation (1), the infinitesimal displacement of the last link is given by D\u0302 \u00bc q1S\u03021 \u00fe \u00fe qnS\u0302n \u00bc Js q \u00f02\u00de where q \u00bc q1 qn T " + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0000881_s11665-020-04703-2-Figure4-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0000881_s11665-020-04703-2-Figure4-1.png", + "caption": "Fig. 4 Typical schematic diagram of a TA10A sliding against the fixed ball (a), 3D (b) and 2D (c) profiles of the wear scars", + "texts": [ + "22 g/cm3, respectively, in accordance with an ASTM Standard of B962-08 (Ref 24). Compared with those of SPS TA10A, Vickers hardness and mean density of SLM TA10A were observed to be significantly higher. Figure 3 depicts a typical cross-sectional morphology and XRD pattern of the fabricated TA10A samples. Microstructures of SLM TA10A can be observed in Fig. 3(a). As is apparent from Fig. 3(b), the SLM TA10A primarily consisted of TiAl alloy and silver, in concordance with the XRD diffraction peaks of the main phases. Figure 4(a) presents the schematic diagram of a TA10A sample sliding against a fixed Si3N4 ball of diameter 6 mm and hardness 15.2 GPa. As depicted in Fig. 4(a), the friction and wear properties of the fabricated material were examined at a wide range of temperatures using a ball-on-disk tribometer of model No. HT-1000. During the sliding process over 80 min, the friction coefficient of TA10A was measured using the controlled system of HT-1000. Values of the wear rateW can be calculated using the formula, W = A\u00c6P/(F\u00c6C) (Ref 25). Here, A, P, F, and C denoted the cross-sectional area, wear-scar perimeter, applied load, and sliding distance, respectively. Figure 4(b) and (c) illustrates the 3D and 2D profiles of the wear scars. A profiler stylus was used to slowly trace a wear scar along the straight line AA (see Fig. 4b), and the coordinate positions of the stylus were recorded to construct a 2D profile of the wear scar, as depicted in Fig. 4(c). The mean crosssectional area A of a wear scar was ascertained to calculate the wear rate W of TA10A. Figure 5 records the typical friction coefficients and wear rates of TA10A at a wide range of temperatures. Compared to the values for SPS TA10A, the friction coefficients and wear rates for SLM TA10A were significantly lowered. Low friction coefficient of approximately 0.28 and a wear rate of almost 1.51 9 10 4 mm3 N 1 m 1) were obtained for SLM TA10A at 450 C. In order to explain the friction and wear behavior of SLM TA10A, the primary wear mechanisms and the silver content of wear scars needed to be discussed further" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0002248_ichve49031.2020.9279899-Figure1-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0002248_ichve49031.2020.9279899-Figure1-1.png", + "caption": "Fig. 1. Geometric model of overhead conductors. (a) Aluminum Conductors Steel Reinforced. (b) Aluminum Conductors Aluminum-Clad Steel Reinforced. (c) Aluminum Conductors Carbon fiber composite core.", + "texts": [ + " And three kinds of overhead conductors are used for analysis. Compare with the cross section of overhead conductors, the length of it can be considered as infinite. The calculation problem can be regarded as a 2D problem. Meanwhile, natural convection and electromagnetic heat are considered in this model. The objects of this paper are Aluminum Conductors Steel Reinforce, Aluminum Conductors Aluminum-Clad Steel Reinforced and Aluminum Conductors Carbon fiber composite core. The structures of overhead conductors are shown in Fig. 1. The structural parameters of overhead conductors are shown in Table I TABLE I. THE STRUCTURAL PARAMETERS OF OVERHEAD CONDUCTORS Object Aluminu m Cross Section (mm2) Steel Cross Section (mm2) Outer Diameter of Reinforcing Core(mm) Outer Diameter of Conductors (mm) ACSR 400 51.90 9.20 27.60 ACSR/AS 400 51.90 9.20 27.60 ACCC/TW 402.40 50.26 8.0 24.80 Authorized licensed use limited to: University of Technology Sydney. Downloaded on May 20,2021 at 07:25:37 UTC from IEEE Xplore. Restrictions apply. The conductive material in the overhead conductors is aluminum, and the steel plays a supporting role" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0001931_j.jmmm.2020.167474-Figure2-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0001931_j.jmmm.2020.167474-Figure2-1.png", + "caption": "Fig. 2. Mesh generated within the stator and rotor parts.", + "texts": [ + " In the air\u2013gap region that connects the stator and rotor domains, the magnetic vector potential A satisfies the Laplace\u2019s equation. Then, by applying the Green\u2019s identity to the Laplace\u2019s equation the BEM equation is obtained in terms of A and its normal derivative \u2202A \u2202n [20,21]: ciAi + \u222b \u0393 A \u2202G \u2202n d\u0393 = \u222b \u0393 G \u2202A \u2202n d\u0393 (3) where \u0393 is the boundary of the air\u2013gap subdomain G is the Green function and c is the boundary factor dependent on the local curvature of \u0393 (ci = 0.5 on smooth boundaries). As shown in Fig. 2, the stator and rotor domains are meshed using three-node triangular finite elements, apart from the outer layers of the rotor domain where the special boundary layers are used to take into account more precisely the skin-effect. The whole mesh includes 9850 elements with the average quality of 0.762. The stator and rotor subdomains are treated by FEM and an approximate solution of equation (1) is derived using the Galerkin\u2019s method: \u222b \u03a9 (\u03c5\u2207A\u2207w \u2212 Jw)d\u03a9 \u2212 \u222b \u0393 w\u03c5G \u2202A \u2202n d\u0393 = 0 (4) where w is the linear shape function between each of the two adjacent boundary elements" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0001133_tasc.2020.2994518-Figure4-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0001133_tasc.2020.2994518-Figure4-1.png", + "caption": "Fig. 4. Geometry of the numerical model with planar (left) and bend stack (right). The modelled part is non-transparent. Copper-coloured cylinder is an electromagnet, silver element in the middle is the modelled stack. The environment is not shown. All dimensions are in mm.", + "texts": [ + " 3D numerical model is used to find the distribution of electric current in the stacks and the strength of trapped flux. The model is created to find discrepancies between an ideal stack with uniform material quality and the actual one, helping analyse the damage to the stack. Additionally it allows to quantify the performance by fitting material parameters. Magnetic Formulation package of Comsol Multiphysics is used. The package applies H-formulation [31]. The geometry of the models with planar and bend stack is shown in figure 4. Thanks to the symmetry of the system only its quarter has to be modelled. The stack is located in the centre of the electromagnet modelled as an externally imposed current density. Both the electromagnet and surrounding are treated as a material with magnetic permeability of 1. External boundaries have magnetic insulation condition imposed on them, meaning that the magnetic field is always parallel to them. Authorized licensed use limited to: UNIVERSITY OF BIRMINGHAM. Downloaded on June 14,2020 at 01:41:33 UTC from IEEE Xplore" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0000659_978-981-15-1616-0_21-Figure20-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0000659_978-981-15-1616-0_21-Figure20-1.png", + "caption": "Fig. 20 Velocity streamlines", + "texts": [], + "surrounding_texts": [ + "From Figs. 9, 15, and 21 of velocity contours, it can be observed that the velocity of exhaust gas is greater at the inlet and outlet pipe compared to the velocities in the subsequent chambers. Figures 10, 16, and 22 show the variation of pressure loss along the length of inlet tube, and Figs. 11, 17, and 23 show the pressure variation along length of outlet tube for corresponding velocities 39, 44, 49 m/s (Fig. 19). From the above data, in Table 4, it can be inferred thatwith increase in exhaust inlet velocity, the backpressure also increases linearly. One can also infer that pressure inside muffler decreases as it reaches exit. A CFD analysis result was verified with physical test results. There was little difference between CFD and experimental Design Analysis and Pressure Loss Optimization \u2026 105" + ] + }, + { + "image_filename": "designv11_71_0001895_1350650120964295-Figure6-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0001895_1350650120964295-Figure6-1.png", + "caption": "Figure 6. Two-dimensional sketch of a sphere-torus contact.", + "texts": [ + " For this special case, eq. (8-12) can be simplified to rPCO;w \u00bc rSP;w (14) xPCO;w \u00bc 0 (15) rPSP;w \u00bc rSP;w (16) xPSP;w \u00bc xSP;w \u00fe cSP CORSP (17) and d \u00bc RSP \u00fe cSP COxSP;w (18) While concave cones are used for modelling of outer ring flanges, contact pairings involving concave spheres and convex cones are not relevant for contact detection within REBs and are therefore not considered within this paper. In the working coordinate system Jw the sphere-torus contact can be reduced to a sphere-sphere contact (see Figure 6). The second sphere with Radius rTO is located at \u00f00;RTO; 0\u00de in terms of Jw and the contact is determined according to the procedure described in the previous section \u2018sphere-sphere\u2019. This procedure remains valid for either concave spheres or tori. Torus-cone problems can be reduced to finding the minimum distance between the torus\u2019s circle of revolution and the cone, because the contact points PCO; PTO and the centre of the torus\u2019s tube are in line. The circle of revolution is defined by the torus\u2019s major radius RTO" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0002314_cyber50695.2020.9279112-Figure1-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0002314_cyber50695.2020.9279112-Figure1-1.png", + "caption": "Figure 1. Schematic diagram of the 2-DOF Delta mechanism.", + "texts": [ + " If the pulley was changed to a fixed panel at the joint, the above shortcomings could be significantly improved. When two PAMs were assembled using a staggered structure, the joint had a large range of motion but a small joint output torque. This might reduce the payload of the system. Based on the above analyses on the advantages and disadvantages, a triangular actuation method was proposed in this paper which used only one PAM to drive the joint. A 2- DOF Delta mechanism was then developed and schematically illustrated in Fig. 1. Compared with the conventional 2-DOF Delta mechanism, two PAMs were utilized as actuators instead of motors. As shown in Fig. 1, the Delta mechanism included a base, a moving platform, and two parallel arms. Each arm consisted of a humerus, an elbow and a forearm. In the developed 2-DOF Delta mechanism, parallelograms were utilized to prevent the moving platform from rotating. A PAM was installed between the base and each humerus to provide driving torque. Detailed zoomed in illustration of the installation of the PAM was presented in Fig. 2. The PAM was installed on the diagonal of the parallelogram of the humerus. A rotary encoder was installed at point O to measure the rotation angle of the humerus \u03b8" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0000942_ilt-01-2020-0030-Figure1-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0000942_ilt-01-2020-0030-Figure1-1.png", + "caption": "Figure 1 Angular velocity of spinning and rolling", + "texts": [ + "htm Industrial Lubrication and Tribology \u00a9 Emerald Publishing Limited [ISSN 0036-8792] [DOI 10.1108/ILT-01-2020-0030] This research is supported by the National Natural Science Foundation of China (No. 51975543). Received 25 November 2019 Revised 20 February 2020 Accepted 23 February 2020 In addition to rotation and revolution, the ball slides and spins on the races under the action of gyroscopic moment. The spin to roll ratio is one of important parameters, which is a key factor to evaluate the wear performance of bearings. Figure 1 shows the relationship between sliding angular velocity v g and rolling angular velocity vR. x0y0z0 is a nutation coordinate system, x0 passes through the center of a ball and is parallel to x axis for xyz coordinate system. b 1 is the pitch angle of the ball, b 2 is the yaw angle and c j is the position angle of the ball (for No. j ball). Under the load, an angle called contact angle naturally forms between the ball and inner or outer raceways. At a low speed, the contact angle of inner race ai and that of outer race ae is approximately equal" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0002485_cac51589.2020.9327292-Figure2-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0002485_cac51589.2020.9327292-Figure2-1.png", + "caption": "Fig. 2: Definition of coordinate system for tiltable quad-rotor, include earth coordinate system, body coordinate system and propeller coordinate system", + "texts": [ + " Section II describes the kinetics model of the tilting quad-rotor under two conditions: normal and faulty. Then, Section III presents the controller design of the linearized model. Section IV present simulation results assessing the effectiveness of the control system we designed, and a prototype experiment is carried out in Section V. Finally, Section VI concludes this paper. II. System dynamics The system used in this paper can be viewed as the joint motion of five connected rigid bodies [8]. As shown in Fig. 2, the coordinate system is established: the earth fixed coordinate system S g(ogxgygzg), the body coordinate system S b(obxbybzb) and the four propeller coordinate systems S pi (opi xpi ypi zpi ), i = 1...4. Use Rg b \u2208 S O(3) to represent the transition matrix from the body coordinate system S b to the earth coordinate system S g: (in this paper, use s(\u00b7) to represent sin(\u00b7), c(\u00b7)to represent cos(\u00b7), and t(\u00b7) to represent tan(\u00b7)) Rg b = \u23a1\u23a2\u23a2\u23a2\u23a2\u23a2\u23a2\u23a2\u23a2\u23a3 c\u03b8c\u03c8 c\u03c8s\u03b8s\u03c6 \u2212 s\u03c8c\u03c6 c\u03c8s\u03b8c\u03c6 + s\u03c8s\u03c6 c\u03b8s\u03c8 s\u03c8s\u03b8s\u03c6 + c\u03c8c\u03c6 s\u03c8s\u03b8c\u03c6 \u2212 c\u03c8s\u03c6 \u2212s\u03b8 s\u03c6c\u03b8 c\u03c6c\u03b8 \u23a4\u23a5\u23a5\u23a5\u23a5\u23a5\u23a5\u23a5\u23a5\u23a6 " + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0000407_6.2020-1184-Figure3-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0000407_6.2020-1184-Figure3-1.png", + "caption": "Fig. 3 Geometry of a partially flat tape spring.", + "texts": [ + " However, tape springs in which there is a central flat region have been investigated to a lesser extend. The HIMast slit tube, [24, 27], is one example, and is designed to allow for embedded conductive elements and a greater root restraining area for increased stiffness. This paper proposes an energy minimisation method to determine the cross section shape, moment-rotation relationship and stable geometry for partially flat tape springs. In this paper a partially flat tape spring is represented by two arcs that tangentially join a central flat, shown in Figure 3. Only opposite sense bending is investigated in the present study because of the known torsional buckling in uniform tape springs. A Finite Element (FE) method is described to capture the moment-rotation behaviour and natural fold radius. Validation of the model is conducted with literature relating to uniform tape springs in the limit of the flat region tending to zero. Then, results from the energy model are compared with a combination of derived analytical expressions and the FE method. III. Energy model A", + " Moment-rotation behaviour The derived energy model has been show to capture the non-linear behaviour of tape springs in bending and is in good agreement with previous analytical models for tape springs of a uniform cross section radius. Figure 11 shows the moment-rotation behaviour of tape springs with increasing flat portions of the cross section. The arc length of all the tape springs remain at 26.88 mm. The cross section remains constant by reducing the included angle, \u03b1, of the tail sections, shown in Figure 3. The radii of the arcs is 14 mm and the tape springs have the mechanical properties of beryillium-copper. Figure 11 shows that the derived energy model does capture a reduction in the maximum snap-through moment with increasing initial flat cross section. Figure 12 shows the relationship between the maximum snap-through moment and the flat section width. The tapes with a larger flat section do not exhibit the snap-through behaviour and are highlighted in Figure 12. There is some discrepancy between the two methods but overall the reduction in snap-through moment is captured" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0000603_iccas47443.2019.8971558-Figure3-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0000603_iccas47443.2019.8971558-Figure3-1.png", + "caption": "Figure 3. Single-segment design comparison", + "texts": [], + "surrounding_texts": [ + "218\nI. INTRODUCTION\nDue to a great contribution to the investigation on continuum robots made them more applicable for industry and medicine. Today, bio-inspired robots from the elephant trunk, octopus arm and other animals with tentacles represent the great potential to be used almost in every aspect of our life. The slender design and hyper-redundant structure make continuum manipulators to adapt an unstructured environment and reachable in confined workspace easily. Traditional serial manipulators with a rigid structure and motorized joints cannot meet these requirements[1].\nAccording to the backbone structure, continuum manipulators can be divided into three groups: discrete, hard and soft continuum robots. Discrete continuum structure consists of a hard spacer disc and interconnected by universal joints or ball joints, and drives by wires. For instance, Tensor arm, Snake robot from OCRobotics, Elephant trunk and Emma type robots [2-5]. Hard continuum type of robots presents manipulators with an elastic support backbone bending feature, which made of rubber or other soft materials and separated by hard spacer discs. For example, Backbone, Catheter, OctArm and Air-OCTOR robots [6-9]. Such kind of robots actuates by cables, dielectric elastomer and by a pneumatic actuator. Among all actuation system, only cabledriven actuation generates high output force, which dramatically increases robot payload capacity.\nAnd the third type of continuum manipulator is soft continuum robots, which are made of hundred per cent soft\nmaterials such as rubber or silicone. These robots are safest and softest robots which can able to change the body size as well. There are many contributed investigations on these type of robots [10-13].\nDue to such great potential, wire-driven continuum manipulators attracted wide attention in a couple decades. There are many scholars and engineers brought great contribution to the development of continuum robotics field. For instance, I.D. Walker and Jones established a new kinematic model based on improved DH method [14]. Moreover, Han Yuan and Zheng Li also proposed a kinematic analysis of the continuum robot based on a static model [15]. Furthermore, X. Dong et.al. from Nottingham University proposed continuum robot design with twin pivot compliant joint and it moves by twin cable actuation which provides motion and high cable tension. Dong proposed kinematics based on cable length variation[16]. Recently, Tiangliang and et.all. developed a cable-driven redundant spatial manipulator with improved stiffness and load capacity by actuating fewer numbers of cables [17]. Furthermore, Julia Starke and et.al. developed a continuum robot arm with a helical tendon routing system, such kind actuation increased robot dexterity and obstacle avoidance capabilities. However, during the robot motion, torsional stress increased [18].\nThis paper presents a novel discrete wire-driven continuum manipulator with a sliding disc mechanism named TakoBot 2. This is a modified prototype of TakoBot 1 [1]. Based on gained experimental results from TakoBot 1 prototype, a modified prototype should have additional redundant degrees of freedom such as torsional motion and twist motion as well. Moreover, this paper explains forward and inverse kinematic formulation with validated simulation\n978-89-93215-18-2/19/$31.00 \u24d2ICROS\nAuthorized licensed use limited to: Auckland University of Technology. Downloaded on October 25,2020 at 08:37:21 UTC from IEEE Xplore. Restrictions apply.", + "219\nresults. The new prototype supposes to improve robot dexterity and reachability capabilities. In chapter II will explain the concept design, III chapter will explain kinematic kinetic formulation, IV chapter will present experiments and simulation results.\nII. CONCEPT DESIGN\nA. Application Analysis The intended application of the robot is shown in Fig.1. The manipulator is designed to be used in the agriculture field for harvesting, weeding and inspection operations if necessary. The intended working environment is highly constrained and by the time workspace will be different because of growing plants.\nTo perform necessary tasks the robot should have the following functions:\n(1) Flexible dexterity: the robot should work in the confined workspace\n(2) Obstacle avoidance capability: the robot should avoid contact with solid surfaces and not collide.\n(3) Reachability: the robot should get the required position in spite of narrow space and obstacles.\n(4) Safety: the robot should be safe enough to avoid breaking any sticks of the plant.\n(5) Portability: the robot should be compact to be used as a tool for farming CNC platform (Fig.1).\nB. System Design Based on application analysis, the robot should have a slender structure with redundant DOFs. After consideration of previous TakoBot 1 prototype capabilities, we made changes in the design and actuation system as well. After numerous experiments with TakoBot 1, demonstrated some limitations related to the design. For example, a robot could not perform torsional motion during the work which led to the accumulation of strain energy inside of manipulator. Secondly, in case of bending it could not perform a pure bending shape. In the new prototype, we added a passive sliding mechanism (Fig.4). Proposed sliding mechanism demonstrated the following benefits:\n1) Smart bending stress distribution: This ability helped to the manipulator to bend purely and it distributed spring potential energy to the whole segment of the manipulator. In comparison with previous, it required less torque and less load to the cables. 2) Torsional motion: passive sliding disc mechanism provided better dexterity ability in case bending helical motion. Helical motion increases torsional stress of the structure unless a disc does not generate yawing around the z-axis. In spite of the heavier weight of the TakoBot 2 was able to perform better and more accurate.\nPassive sliding disc mechanism works for all discs attached to the backbone except base and end-effector discs, which means total length always remains constant. Fig.4. demonstrates a sliding mechanism structure. According to the design, travelling distance of the disc is 10mm, while the disc diameter is 50mm, an average distance of the single section (between discs) about 35 mm. 3D printed discs connected by coil compression springs, such design provides stiffness to the manipulator. The spring constant could be variable depends on motor torque, in this prototype we used spring with constant 0.63 N/mm. Each section consists of four segments, end segments discs connected by four springs, while mid-section segments connected by 8 springs between discs.\nAuthorized licensed use limited to: Auckland University of Technology. Downloaded on October 25,2020 at 08:37:21 UTC from IEEE Xplore. Restrictions apply.", + "220\nC. Actuating mechanism Manipulator's tendons actuate by a linear lead shaft\nconnected to the stepper motor by a coupler. Tendons attached to both sides of specially designed inlets on screw housing and screw with housing travels along the shaft. 1mm steel cable utilized as a tendon.\nIII. KINEMATICS AND KINETIC FORMULATION\nA. Forward kinematic formulation Coordinate systems are set at every universal joint. The homogeneous coordinate transform matrices:\n\u03a30\u2192\u03a31 , H0,1= ( R u0,1 0 0 0 1 ), u0,1=( x0 y0 l0 )\n(1)\n\u03a3i-1\u2192\u03a3i, Hi-1, i= ( R ui-1,i 0 0 0 1 ) , ui-1,i=( 0 0 L ),\n(i=2, \u22ef, n)\n(2)\nR= Rz(\u03b8zi)Rx(\u03b8xi)Ry(\u03b8yi) (3)\nwhere, \ud835\udc65\ud835\udc650 and \ud835\udc66\ud835\udc660 are an initial position of the base. \ud835\udc45\ud835\udc45\ud835\udc65\ud835\udc65(\ud835\udf03\ud835\udf03\ud835\udc65\ud835\udc65\ud835\udc65\ud835\udc65) and \ud835\udc45\ud835\udc45\ud835\udc66\ud835\udc66(\ud835\udf03\ud835\udf03\ud835\udc66\ud835\udc66\ud835\udc65\ud835\udc65) are rotation matrices of ith universal joint that has two rotation angles \ud835\udf03\ud835\udf03\ud835\udc65\ud835\udc65\ud835\udc65\ud835\udc65 and \ud835\udf03\ud835\udf03\ud835\udc66\ud835\udc66\ud835\udc65\ud835\udc65, \ud835\udc45\ud835\udc45\ud835\udc67\ud835\udc67(\ud835\udf03\ud835\udf03\ud835\udc67\ud835\udc671) is a rotation matrix of the ith disk with a rotation angle \ud835\udf03\ud835\udf03\ud835\udc67\ud835\udc671 along the axial axis and L is a length between neighbouring universal joints. Three rotation matrices have:\nMultiplying the H-matrices successively, we get unit\nvectors and the position vector of the ith coordinate system; H0,i=H0,1H1,2\u22efHi-1, i= (\nii ji ki ui 0 0 0 1 ) (4)\nwhere, \ud835\udc62\ud835\udc62\ud835\udc65\ud835\udc65 is the position of the ith universal joint Ui (\ud835\udc56\ud835\udc56 =\n1,\u22ef , \ud835\udc5b\ud835\udc5b \u2212 1). The position vector \ud835\udc5d\ud835\udc5d\ud835\udc65\ud835\udc65 of the end-point P\ud835\udc5b\ud835\udc5b and position of sliding plates Pi (i=1, \u22ef, n-1) of the manipulator are obtained by, (pi\n1)=H0,i(0 0 li 1)T, (i=1, \u22ef, n) (5)\nwhere, \ud835\udc59\ud835\udc59\ud835\udc5b\ud835\udc5b is a fixed length between the nth universal joint and the most distal plate. Position vectors of 8 hole A0, B0, C0, D0 , A\u03020, B\u03020, C\u03020, D\u03020 at the base plate are determined as,\na0=( ax ay 0 ) , b0=( bx by 0 ) , c0=( cx cy 0 ) , d0=( dx dy 0 ) ,\na\u03020=( a\u0302x a\u0302y 0 ) , b\u03020=( b\u0302x b\u0302y 0 ) , c\u03020=( c\u0302x c\u0302y 0 ) , d\u03020=( d\u0302x d\u0302y 0 ) ,\n(6)\nPosition vectors of 4 hole A\ud835\udc65\ud835\udc65, A\u0302\ud835\udc65\ud835\udc65, C\ud835\udc65\ud835\udc65, C\u0302\ud835\udc65\ud835\udc65 at the ith plate\n(\ud835\udc56\ud835\udc56 = 1,\u22ef , \ud835\udc5b\ud835\udc5b) are obtained as\n(ai 1)=H0,i( ax ay li 1 ) , (a\u0302i 1)=H0,i( a\u0302x a\u0302y li 1 ) ,\n(ci 1)=H0,i( cx cy li 1 ) , (c\u0302i 1)=H0,i( c\u0302x c\u0302y li 1 ) , (i=1, \u22ef, n)\n(7)\nwhere, \ud835\udc59\ud835\udc59\ud835\udc65\ud835\udc65 is an axial length between the ith universal joint and the ith plate, which varies as the plate slides along rods, except \ud835\udc59\ud835\udc59\ud835\udc5b\ud835\udc5b.\nIn the same way, position vectors of 4 hole B, B\u0302\ud835\udc65\ud835\udc65, D\ud835\udc65\ud835\udc65, D\u0302\ud835\udc65\ud835\udc65 at the ith plate (\ud835\udc56\ud835\udc56 = 1,\u22ef ,\ud835\udc5a\ud835\udc5a) are obtained as,\n(bi 1)=H0,i( bx by li 1 ) , (b\u0302i 1 )=H0,i ( b\u0302x b\u0302y li 1) ,\n(di 1)=H0,i( dx dy li 1 ) , (d\u0302i 1 )=H0,i ( d\u0302x d\u0302y li 1) , (i=1, \u22ef, m)\n(8)\nB. Kinetic formulation Our continuum manipulator is divided by two segments.\nAuthorized licensed use limited to: Auckland University of Technology. Downloaded on October 25,2020 at 08:37:21 UTC from IEEE Xplore. Restrictions apply." + ] + }, + { + "image_filename": "designv11_71_0001603_s00202-020-01081-9-Figure8-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0001603_s00202-020-01081-9-Figure8-1.png", + "caption": "Fig. 8 Calculation of commutations number", + "texts": [ + " In order to quantify the benefits of the MP FC in limiting the commutation losses, intended as number of commutation, a methodology is proposed to count the commutations performed by the inverter switches per each control cycle and for the entire operating period as well. The counting procedure is based on the voltage vector index (i), each of them representing a specific switch state of the inverter by means of a triplet of binary digits. When a change of voltage index (i) occurs, denoting a change of inverter output voltage, the switch states are also changed. The consequent number of commutations occurred in the inverter corresponds to the number of digits changed in the two triplets that can be easily counted sampling time by sampling time. Figure\u00a08 illustrates the code steps which are applied to calculate the number of commutations, according to the voltage vector index monitoring as, just described. The implementing flowchart of the proposed MP FC is reported in Fig.\u00a09. To investigate the validity of the proposed MP FC scheme, its performance is compared with the MP DTC scheme which utilizes a cost function form expressed as in Eq. (53): Instead of using (16), it can be foreseen that the execution time taken by the MP DTC cost function is greater than the correspondent one of the proposed MP FC", + " For the MP FC with LMC control approach, the spectra of stator current components are shown in Fig.\u00a021, in which the \u03b1-component introduces a value of 7.386\u00a0A with THD of 2.22%, while the \u03b2-component shows a value of 7.264\u00a0A, and a THD of 2.53% of the fundamental. In the same manner, the THD values for the stator currents obtained during this test are given in Table\u00a03. The comparison between the three control procedures is also performed in terms of the number of commutations which are counted using the code illustrated in Fig.\u00a08. The comparison data are in Table\u00a04. The comparison confirms the suitability of MP FC, with and without LMC, for limiting the commutation losses in comparison with the MP DTC which inherently suffers by the high computational burden. The proposed MP FC control procedure has been implemented by a dSpace 1104 control board. The experimental tests have been carried out under the same working conditions adopted in the simulation tests. Experimental results confirm the validity and effectiveness of the proposed MP FC and its superiority over the MP DTC in limiting the ripples in torque, flux and current" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0000562_978-3-030-39216-1_2-Figure13-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0000562_978-3-030-39216-1_2-Figure13-1.png", + "caption": "Fig. 13. Generalized frequency responses for the norm (N)", + "texts": [ + " Figure 12 shows similar characteristics for different pathologies. Note that all the transitions of the graphs through the abscissa axis correspond to the transitions of the hodographs to the new quadrant. For healthy eyes (N), these are the two graphs with the largest amplitudes. A healthy eye has characteristics in the first, fourth, and third quadrants that are similar to the delay characteristics, what confirmed by Fig. 11. Characteristics of eyes with pathologies do not possess such properties and at the same time significantly differ from each other. Figure 13 shows six generalized frequency responses for norm (N). Note that, despite significant differences in amplitudes, these graphs remain similar to the graphs and hodographs for the delay in the same three quadrants, which were discussed above. The graphs of the generalized frequency characteristics of works for six eyes with glaucoma (GL), shown in Fig. 14, do not possess this property. These graphs show a significant departure from the delay in the first quadrant, which is also clearly seen in Fig", + " Scalability and Parallelization of Sequential Processing 291 The mappings between the information spaces, mentioned above, provide a wide range of options for transforming information during the processing. The most universal and effective is to transform the prior information and all the available raw data into the canonical form, combine information in this form and calculate the final result from the accumulated canonical information. Here the use of the canonical form as the main one allows transforming the processing scheme into a completely parallel one, Fig. 13. This scheme has all the advantages of the previous one. In addition, the transformation of various pieces of data and the prior information in the canonical form can be carried out independently and in parallel on different computers. When new data fragments appear, it will be necessary to convert them into the canonical form, add to the accumulated canonical information and recalculate the result using the updated canonical information. Fig. 12. General scheme of sequential updating in canonical form. Each data set is first transformed into canonical form. After that canonical partial information is added to the accumulated information. 292 P. Golubtsov The transformations presented in Fig. 13 are described by the following expressions: \u2022 Transformation of the initial prior information to the canonical form: x0;F0\u00f0 \u00de 7! ~u0; ~T0 \u00bc F 1 0 x0;F 1 0 ; \u2022 Transformation of individual piece of raw data into partial canonical information: yi;Ai; Si\u00f0 \u00de 7! ~ui; ~Ti \u00bc A i S 1 i yi;A i S 1 i Ai ; i \u00bc 1; . . .; k; \u2022 The composition of all the fragments of canonical information, which reflects parallel accumulation of information in canonical form: \u2022 The construction of the final result of estimation from the accumulated canonical information: uk; Tk\u00f0 \u00de 7" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0000356_iecon.2019.8927121-Figure14-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0000356_iecon.2019.8927121-Figure14-1.png", + "caption": "Fig. 14. Rotary-linear stepper motor [36]", + "texts": [ + " A similar machine, but with much longer stroke possibilities was developed by J.F. Pan et al. The main differences against the previously detailed rotary-linear SRM is that it has 2 two salient pole rotational stack, and 3 for the linear motion, as it can be seen in Fig. 13. The mover is a passive toothed cylinder. For independently controlling the two types of motion a simple but efficient net torque methodology based decoupling method was proposed [35]. Also stepper motors were suggested in rotary-linear variants [36], [37]. The RLM shown in Fig. 14 seems like the well-known teethed rotating stepper motor of variable reluctance type. Half of the stator poles are teethed axially, while the other ones on a perpendicular direction to this. The first mentioned set of poles are used to generate the rotational movement, and the other ones the linear one. As in the previously detailed case, all the poles have concentrated coils wound around them. The rotor comprises if two half teeth pitch shifted iron core stacks. They equally have teeth both on the circumference and axial direction to assure the dual motion for this RLM" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0000893_msf.982.75-Figure4-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0000893_msf.982.75-Figure4-1.png", + "caption": "Fig. 4. Comparison between the experimental and numerical deformed shapes when wrinkling occurs due to insufficient blank holder pressure - Stress distribution of different layers along the normal direction (S11)", + "texts": [ + " Thinning and tearing at Depth failure: The depth at which failure happened within the forming process was measured from the numerical values of the thinning rate curve of the laminate. Wrinkling: Find an appropriate method for the severity of wrinkling is difficult; a measure through observation was used to compare the severity of wrinkling of the formed cup. Delamination: In this investigation, observation measure was used to compare the severity of delamination of the formed cup. Effect of Blank-holder Force. Simulations and experimental results provided us a comprehensive overview of the effect of blank-holder force on the of the FML system formability. Fig. 4 shows cups formed at low blank-holder force of 6 kN, all specimen in simulation and experiment\u2019s shown the presence of wrinkling, demonstrating that a low blank-holder force wrinkling as the primary failure mode. Fig. 5 shows formed cups at the high blank-holder force of 6 kN; all formed cups showed very small or non-presence of wrinkling. Some Simple failed due to tearing or fracture, signifying that a high blank holder force increases the possibility of failure due to tearing and fracture and decreases the severity of wrinkling" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0003203_iecon.2014.7048554-Figure4-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0003203_iecon.2014.7048554-Figure4-1.png", + "caption": "Fig. 4. Illustration of the rotating hysteresis shape in the stator coordinate system; Illustration for \u03d1 = 40\u25e6", + "texts": [ + " Due to this fact and the requested use of a d-q-orientated hysteresis [6]\u2013 [11], a transformation must be implemented according to the following: Clarke Transformation ( I\u03b1 I\u03b2 ) = \u239b \u239c\u239d 2 3 \u2212 1 3 \u2212 1 3 0 1 \u221a 3 \u2212 1 \u221a 3 \u239e \u239f\u23a0 \u239b \u239dIa Ib Ic \u239e \u23a0 (1) Rotation Matrix( Id Iq ) = ( cos\u03d1 sin\u03d1 \u2212 sin\u03d1 cos\u03d1 )( I\u03b1 I\u03b2 ) (2) Inverse Rotation Matrix( I\u03b1 I\u03b2 ) = ( cos\u03d1 \u2212 sin\u03d1 sin\u03d1 cos\u03d1 )( Id Iq ) (3) Park Transformation [11], [12]. ( Id Iq ) = X \u239b \u239dIa Ib Ic \u239e \u23a0 (4) X = 2 3 \u239b \u239c\u239c\u239d cos\u03d1 cos ( \u03d1\u2212 2\u03c0 3 ) cos ( \u03d1+ 2\u03c0 3 ) \u2212 sin\u03d1 \u2212 sin ( \u03d1\u2212 2\u03c0 3 ) \u2212 sin ( \u03d1+ 2\u03c0 3 ) \u239e \u239f\u239f\u23a0 These transformations allow the analytical rotation of the hysteresis shape (see Fig. 4). 1) Choice of Bit Pattern: Due to the rotation of the hyssteresis shape the switching sectors rotate as well. Therefore these sectors must be steadily actualised while the hysteresis is rotating. Fixed angles to separate the sectors (according to the \u03b1\u03b2-coordinate system) are chosen, in which a specific space vector is switched. These angles are 30\u25e6, 90\u25e6, 150\u25e6, 210\u25e6, 270\u25e6, and 330\u25e6 (see Fig. 5). These angles remain constant also during the rotation of the hysteresis shape (see Fig. 6). 2) Utilization of Zero Voltage Space Vectors: A reduction of the switching frequency can be achieved by using the zero voltage space vectors" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0002342_rcar49640.2020.9303289-Figure5-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0002342_rcar49640.2020.9303289-Figure5-1.png", + "caption": "Figure 5. Middle layer structure", + "texts": [ + " The integrated insole should be in direct contact with the bottom of the human foot and exert a force to assist the foot in lifting. In consideration of the comfort of wearing, this paper used a soft insole as the basis to design a three-layer integrated insole. As shown in figure 4, the three-layer structure from top to bottom is an upper flexible insole, an intermediate layer structural member and a lower flexible insole, which are connected by double-sided buckle rivets. The middle layer structure is the core of the three-layer integrated insole structure as shown in figure 5. It consists of a front-end structure and a back-end structure of the middle layer, and distributed sensors are also installed. The front-end of the middle layer is used to connect the driven-rope. The final 465 Authorized licensed use limited to: Univ of Calif Santa Barbara. Downloaded on May 19,2021 at 02:45:08 UTC from IEEE Xplore. Restrictions apply. action point of the driven-rope should be located at opposite ends of the insole. And the front-end of the middle layer is to determine the endpoint of the rope which is located at the outermost hole on either side of the structure" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0000045_ab5165-Figure4-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0000045_ab5165-Figure4-1.png", + "caption": "Figure 4. Illustrated fabrication process of tunable (a) 3d printing the rotational ring and attach the conductive pattern, (b) connect with the SMA springs and put into the case3, (c) connect SMA spring and passive springs to the latch, (d) fabricated locking system", + "texts": [ + " After, degassing in a vacuum chamber for 20 minutes at room temperature and thermal cross-linking at 333K for 8 hours, the lens structure is peeled off from the mold. Convex part of the fabricated lens has a diameter of 35mm and a height of 2.2mm, and plane part has a diameter of 45mm and a thickness of 1.2mm. Lastly, six connecting brackets are bonded to the plane lens part using DOWSILTM 7091 Adhesive Sealant from Dow Corning. (a) (b) (c) (d) (e) (f) 1 2 3 4 5 6 7 8 9 10 11 12 13 14 15 16 17 18 19 20 21 22 23 24 25 26 27 28 29 30 31 32 33 34 35 36 37 38 39 40 41 42 43 44 45 46 47 48 49 50 51 52 53 54 55 56 57 58 59 60 A cc pt ed M an sc rip t 3 Figure 4 shows the overall fabrication process of the tunable lens system. The rotational ring part is 3d printed with Polycarbonate that has a high glass transition temperature of 420K. The rotating part is composed of the rotational ring, the locking gear, and the conductive pattern made of aluminum tape. The conductive pattern is used for sensing the rotation angle of the rotational ring. Due to its inherent shape, when the system is rotated, the area of the conductive pattern, that faces the coil sensor, is varied", + " Fabricated tunable lens. (a) Whole structure, (b) Locking part, (c) Innter structure. To reduce the rotational friction, the thin ball bearing which has an inner diameter of 70mm and an outer diameter of 74mm is installed between the rotational ring and the case2. In the initial state, one SMA spring is pre-stretched from its initial length of 30mm to 55mm while the other spring remains in its initial length. The ends of two SMA springs are then fixed to the case3 and the rotational ring as shown in figure 4(b). The locking part is composed of the locking gear, the SMA spring, passive springs, and the latch. One side of the springs is parallelly connected to the latch, and the other side is connected to the case3. Parameters of the SMA springs which are used in the rotating part and the locking part are summarized in table 1, and in figure 5, the fabricated tunable lens is shown. Part Feature Specifications Rotating part Material Ni-Ti Wire diameter 500 \u03bcm Loop diameter 6 mm Initial length of the spring 30 mm Locking part Material Ni-Ti Wire diameter 500 \u03bcm Loop diameter 3 mm Initial length of the spring 4 mm 1 2 3 4 5 6 7 8 9 10 11 12 13 14 15 16 17 18 19 20 21 22 23 24 25 26 27 28 29 30 31 32 33 34 35 36 37 38 39 40 41 42 43 44 45 46 47 48 49 50 51 52 53 54 55 56 57 58 59 60 A cc ep te d nu sc Smart Mater" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0003741_itec-ap.2014.6941162-Figure2-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0003741_itec-ap.2014.6941162-Figure2-1.png", + "caption": "Fig. 2. The structure of the axial MFM-BDRM", + "texts": [ + " As the axial flux topology of the magnetic-field-modulated brushless doublerotor machine (MFM-BDRM) is superior in terms of torque density to the equivalent standard radial flux topology and the flat structure results in high utilization of axial space, it is particularly suitable for application in hybrid electric vehicles (HEVs). The axial MFM-BDRM system is shown in Fig. 1, which is composed of an ICE, a battery pack, two inverters, an axial MFM-BDRM, a conventional permanent magnet synchronous machine (PMSM) and a final gear. The axial MFM-BDRM realizes the function of speed decoupling while the conventional PMSM enables the torque decoupling from the ICE to the load. II. TOPOLOGY AND OPERATING PRINCIPLE Fig. 2 shows the configuration of the axial MFM-BDRM. The axial MFM-BDRM comprises three parts: a stator, a modulating ring rotor and a permanent magnet (PM) rotor. The modulating ring rotor, composed of evenly distributed ferromagnetic segments and nonmagnetic segments, is located between the stator and the PM rotor. The input shaft of the PM rotor is connected with the output shaft of the ICE while the output shaft of the modulating ring rotor, also as the revolving shaft of the PMSM, is linked to the load" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0002162_jsen.2020.3039865-Figure6-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0002162_jsen.2020.3039865-Figure6-1.png", + "caption": "Fig. 6. Analysis results of movement consistency, and the missing part in the figure represents that the subject\u2019s RoO in this group is 0.", + "texts": [ + " RoO was used to measure the movement consistency of the subjects, which refers to the range of position variation of a weight-bearing pattern between the three experiments. For example, between the three experiments, if a subject\u2019s normal walking pattern is located at 2, 3, and 2 from left to right, RoO is 1. If RoO is 0, movement consistency is considered excellent; If RoO is 1 or 2, movement consistency is considered to be good; If RoO is 3 and 4, movement consistency is poor. Analysis of the shank segments (Fig. 6) showed the movement consistency of Sub5 is poor, whereas Sub6 has a good movement consistency. We found that Sub6 has excellent movement consistency in group G3 and G5. At the same time, the analysis of the thigh segments showed that the movement consistency of the thighs amounts to 90%, which was 12.5% higher than 77.5% of the shanks. Statistic \u2212\u2192 \u03b8 0,1 (TABLE II) is the offset of the shanks during natural walking. We observed that 87.5% of the subjects had a left-leaning gait on the shanks. At the same time, the offset of the thighs during natural walking was calculated, and 75% of the subjects showed a right-leaning gait on the thighs" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0001846_eit48999.2020.9208283-Figure1-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0001846_eit48999.2020.9208283-Figure1-1.png", + "caption": "Fig. 1 Inventor model of BUV", + "texts": [ + " An electric vehicle topology was chosen as it meets most of the established criteria, in addition to being very energy-efficient and pollution-free. The BUV topology was modeled after a golf cart, which makes it very versatile (capable of several tasks, easy to drive even for novices, etc.). Several designs were laid out in Autodesk Inventor and the final design was chosen based on how well the various components could be accommodated, how easy it would be to fabricate and modify, and how well the BUV would perform. The design concept is shown in Fig. 1. III. MATERIALS AND FABRICATION After researching many vendors, decisions were made on which components to purchase and install and which components to fabricate using stock materials. In some cases, readymade components were modified. Guiding factors were availability, cost, and function. The frame was a simple ladder shape with multiple rungs and out-runners in order to be able to affix the axles, suspension, floor, etc. The footprint was to be 4\u2019 x 8\u2019, standard golf cart dimensions. Steel box tubing was used" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0000591_pierc19091602-Figure1-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0000591_pierc19091602-Figure1-1.png", + "caption": "Figure 1. BBLDCM with Outer Rotor.", + "texts": [ + " Taking the x-axis single-degree-of-freedom displacement of the rotor as an example, the radial suspension force Fx in the x-axis direction of the rotor is defined as control quantity u. Super twisting algorithm is adopted, and the second-order sliding mode is improved to realize the second-order sliding mode direct suspension force control. The control results show that the second-order sliding mode direct suspension force control is feasible and effective. BBLDCM embeds a set of suspension control windings in stator slots of the brushless DC motor, so that the suspension magnetic field and rotating magnetic field share a set of core magnetic circuit. Figure 1 is a schematic diagram of an external rotor BBLDCM. a1, a2, b1, b2, c1, and c2 are suspension windings of the motor. Each set of windings is composed of two windings in series. When a1 winding passes into the current in the direction shown in the figure, the magnetic force at air gap 1 decreases. Conversely, the magnetic density at air gap 2 increases, thereby breaking the balance of the air gap magnetic density on both sides of the rotor and generating the suspension force that displaces the rotor along axis +x" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0000431_dynamics47113.2019.8944673-Figure6-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0000431_dynamics47113.2019.8944673-Figure6-1.png", + "caption": "Fig. 6. Action of components of voltage space vector on amplitude and rotational speed of stator flux", + "texts": [ + " Integrating equation (24), the stator flux linkage of the AFPM motor in the fixed frame of reference is dtiRu )( . (27) If we take into account that during the switching period of inverter the voltage vector is constant, equation (27) can be written as 0 tidtRut . (28) Because the stator resistance is usually small, it can be neglected. Then, according to (28), the trajectory of the stator flux vector will move in the direction of the vector of the applied voltage (Fig. 5). At the instant of switching the stator flux linkage is 0t . Having studied the presented Fig. 6, we can conclude that to control the magnetic flux linkage vector of the AFPM motor stator, it is necessary to adjust the along voltage vector, and to control the rotation speed of the flux linkage vector, it is necessary to adjust the vertical component of the voltage vector. In the case of AFPM motor (as contrasted with induction motor), flux linkage will non-constant value when zero space vectors u0 and u7 are applied since the permanent magnets rotate with rotor. Therefore, the use of zero-voltage states to control the AFPM motor is impractical", + " Parameter Value Rated torque MN (N\u00b7m) 11 Power switches of voltage source inverter IGBT For the current model, it is assumed that the bandwidths of torque and magnet flux hysteresis controllers are 0.01\u00b7M* and 0.01 \u00b7\u03c8* respectively. When modeling, it is also assumed that when starting up, the rotor is in the starting position ( 0 ); therefore, the initial flow area is equal to 1 . The variation range of the reference torque is from 11 to -11 N\u00b7m at t = 0.175 s, and from -11 to 11 N\u00b7m at t = 0.175 s. After examining the Fig. 6 and (28), it is clear that the smaller the sample time, the smaller the flux and torque. For the model under study, the sample time TS = 10 \u03bcs was chosen. The torque characteristics is shown in Fig. 8. This figure shows that when a set torque changes, the reference torque very quickly follows the set torque. The reference torque is controlled within the bandwidth when the set torque is constant, i.e. when the system is in steady state. The magnetic flux of the stator, controlled within the bandwidth, is shown in the Fig" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0001140_1.i010790-Figure5-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0001140_1.i010790-Figure5-1.png", + "caption": "Fig. 5 Schematic of 5 UAVsA1;: : : ;A5 pursuing a target UAV swarm,", + "texts": [ + " If the Chebyshev center is not located at the centroid of the polygon formed by the pursuing UAVs, then unequal values of \u03bbi will be required. An algorithm to compute \u03bb1; : : : ; \u03bbn is as follows. For a given X, find \u03bb1; : : : ; \u03bbn: minimize Xn i 1 \u03bbi \u2212 1 n 2 subject to constraints \u03bbi > 0; Xn i 1 \u03bbi 1; Xn i 1 \u03bbiAi X (37) Wenow consider a scenariowhere the pursuing UAVs are carrying a rope line with which they wish to surround the target. The rope essentially forms an open chain as depicted in Fig. 5 (where the solid edges of the polygon represent the rope). In such a scenario, theUAVs mayneed to reorient the formation appropriately so that the formation approaches the target from the direction of the open end of the chain. More precisely, if h\u0302 represents the unit normal to the line joining the pair of UAVs that are carrying the open end of the chain, we would like the UAVs to apply the requisite accelerations that will enable them to move towards the target, while simultaneously rotating the formation so as to orient h\u0302 appropriately. In Fig. 5, the rope line is formed byUAVsA1A2A3A4A5, whileUAVsA1 andA5 carry the open ends of the rope. Therefore, the normal of interest in this figure is the normal to the line joining UAVs A1 and A5. For the purpose of this section, we assume that the sizes of the target circle and the polygon A1; : : : ; An formed by the pursuing UAVs are such that (i) the open edge of the polygon is of length greater than the diameter of the circle, and (ii) the perimeter of the polygon is greater than the circumference of the circle B", + " After substituting terms from (38),we see that in order for this to happen, we require that the accelerations of A1 and An satisfy the following equation: \u2212a1 sin \u03b41 \u2212 \u03b81n an sin \u03b4n \u2212 \u03b81n V\u03b8;1n 2Vr;1n\u2215r1n \u2212 K1 \u2212 r1nK2 \u03b81n \u2212 \u03b81n;d r1nK1 _\u03b81;nd (40) The above equation can then be written in terms of the control inputs u\u03b1;1, uV;1, u\u03b1;n, uV;n as follows: \u2212u\u03b1;1V1 cos \u03b11 \u2212 \u03b81n \u2212 uV;1 sin \u03b11 \u2212 \u03b81n u\u03b1;nVn cos \u03b1n \u2212 \u03b81n uV;n sin \u03b1n \u2212 \u03b81n V\u03b8;1n 2Vr;1n\u2215r1n \u2212 K1 \u2212 r1nK2 \u03b81n \u2212 \u03b81n;d r1nK1 _\u03b81;nd (41) If the control inputs are applied such that (36) and (41) are satisfied, then the pursuing UAVs will be able to surround the target with the rope at the appropriate orientation. The reference value \u03b81n;d would be equal (or close) to \u03b1X \u03c0\u22152 . Objective c: This requirement dictates that the distances between the UAVs need to remain constant throughout the engagement. We note that a n\u2013sided polygon can be triangulated into n \u2212 2 distinct triangles. As an example, the pentagon A1A2A3A4A5 in Fig. 5 can be triangulated into three trianglesA5A1A2,A5A2A3, andA5A3A4. It can be ensured that the pentagon will remain rigid if each of these three triangles remain rigid, or in otherwords, each of the seven sidesA1A2, A2A3, A3A4, A4A5, A5A1, A5A2, and A5A3 remain of constant length. A general n-sided polygon will remain rigid, if 2n \u2212 3 inter UAVdistances (between UAV pairs chosen based on a triangulation of the polygon) remain constant. LetE represent the set of these 2n \u2212 3 inter UAV-distances. The distance between a UAV pair Ai; Aj can be kept constant if the accelerations of UAVs Ai and Aj are such that _Vr;ij 0 (with the assumption that the initial conditions are such that Vr;ij 0)" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0000799_gncc42960.2018.9019083-Figure1-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0000799_gncc42960.2018.9019083-Figure1-1.png", + "caption": "Figure 1. XQ-6B flying-wing UAV", + "texts": [ + " The fast adaptation and robustness of the method are guaranteed. This paper is based on an in-depth study of the L1 control method. A lateral control law for flying-wing aircraft is established. The control reconfiguration effect of L1 adaptive controller is verified by regarding control surface failures as time-varying parameter and input disturbance. II. PROBLEM DESCRIPTION The research object of this paper is the XQ-6B flying-wing UAV designed and manufactured by the laboratory the author is working for. As shown in Fig. 1. Authorized licensed use limited to: Imperial College London. Downloaded on June 17,2020 at 17:42:11 UTC from IEEE Xplore. Restrictions apply. The control surfaces and their numbers are marked in Figure 1. The No.4 and No.5 control surfaces are arranged in an overlapping manner. When unilateral No.4 and No.5 control surfaces deflect a certain angle, the drag on both sides of the UAV will become asymmetrical. The UAV will be driven into a yaw motion. The two pieces of control surface are used as a rudder. According to the characteristics of the flying-wing UAV model and its faults, the model of the UAV can be represented as the following state-space equation: ( ) ( ) ( ) ( ) ( )T x t Ax t Bu t y t C x t \u03c3 = + + = (1) The uncertain parameters of the system spread across A, B and \u03c3 " + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0003627_2014-01-2146-Figure4-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0003627_2014-01-2146-Figure4-1.png", + "caption": "Fig. 4. Assumption diagram of airship balance.", + "texts": [ + " Operation of these diagrams result in control commands to inlet valves (out-fan) and outlet valves (out-valve). Diagram elements realize logical discrete control law according to alphabet of input/output values and logic of their dependency described in publication [4]. Ballonets control function described in previous section stay within air boost control task in airship body. Additional pitch control function introduced for them in statement requires development of mathematical model of its dependency on air masses stored in ballonets. Assumption diagram of airship balance is given in Fig. 4. Therewith airship basic moments are affected by masses: m0 - airship total mass without ballonets, m1 - stern ballonet mass, m2 - fore ballonet mass, determining corresponding gravity forces: G0, G1, G2. Besides, moments are affected by each force arm (l0, l1, l2) relative to center of buoyancy Ob. Airship pitch angle change leads to arm change (l0 *, l1 *, l2 *). Thus resulting in balance conditions change. According to diagram on Fig. 4, the ballonets masses and pitch angle mathematical dependency model, via total moment M\u03a3, is of the form: (3) Airship turn dynamics affected by moment (3) is determined by system of equations: (4) where \u03c9(t) - angular rate of turn. System of models (3) and (4) is realized in Simulinc environment as a diagram given on Fig. 5. Inputs are fed with airship masses. In0 - total mass, In1 - stern ballonet mass, In2 - fore ballonet mass. Modeled in the diagram output is airship pitch angle. Diagram on Fig" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0002161_icem49940.2020.9270955-Figure3-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0002161_icem49940.2020.9270955-Figure3-1.png", + "caption": "Fig. 3. FEA model of five-phase PMSM", + "texts": [ + "4 225sin n eI \u03b8 \u03c0 \u03b1 + (12) To sum up, two fault-tolerant currents can be obtained with single-phase open-circuit faults using the above method. In the same way, two fault-tolerant currents with two-phase and three-phase open-circuit faults can also be obtained. And the fault tolerance performances in the two cases are compared by simulation. In this section, simulation model is established to verify the above theory and experimental system is designed. To verify the performance of the fault-tolerant current, a five-phase surface-mounted PMSM model with 20 slots and 18 poles is built using the nite element analysis (FEA) as shown in Fig. 3. The winding of the motor adopts singlelayer distribution. The main parameters are listed in Table I. The simulation waveform with single-phase open-circuit fault is shown in Fig. 4. Fig. 4 (a) shows that when the angle is 0.937 , THD is the smallest, only 2.4%. Therefore, the optimal angle is THD=0.937 . By substituting the optimal angle into (8) and (12), the fault-tolerant currents in two cases of no round MMF limit and reconstruction of round Authorized licensed use limited to: Auckland University of Technology" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0002923_9781119633365-Figure5.7-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0002923_9781119633365-Figure5.7-1.png", + "caption": "Figure 5.7 (a) Buck and inverse buck and (b) boost and inverse boost.", + "texts": [ + " Similarly, based on observation and induction, the converters shown in Figures\u00a01.7 and 1.8 can be explained with two canonical switching cells, namely, Tee canonical cell and Pi canonical cell, as shown in Figure\u00a05.6. By exhaustively enumerating all (b)(a) 5 Synthesizing Process with\u00a0Graft Scheme98 (a) (b) of possible combinations of Zin, Zout, and Zx, and including LC network, source, and load, the converters shown in Figures\u00a01.7 and 1.8 can be derived. Additionally, the converters with extra LC filters and inverse versions of the converters can be derived. Figure\u00a05.7 shows the buck and boost converters and their inverse versions, in which the input\u2010to\u2010output transfer ratios shown in the bottom of the converters are corresponding to continuous conduction mode (CCM), and D is the duty ratio of the active switch. Moreover, the converters with the same higher step\u2010up voltage transfer ratio but with different circuit configurations, as shown in Figure\u00a05.8, can be developed. Although this approach can derive new PWM converters and is straightforward, it is still tedious and lacks of mechanism to explain the converters with identical transfer ratio but with different configurations" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0003499_muh-1404-6-Figure6-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0003499_muh-1404-6-Figure6-1.png", + "caption": "Figure 6. Bolted connections between column/bed and bed/base.", + "texts": [ + " Many research studies have been developed in the last decades to model the bolted connections in the machine tool; however, this paper is concerned with the method discussed in [10]. The contact between the bolts\u2019 cylindrical bodies and the holes in the corresponding substructures is defined as bonded, while that between the lower surface of the bolts\u2019 heads and the surface of the substructure is defined as a frictionless contact. Bolted connections between column/bed and bed/base are shown in Figure 6. Any machine tool center must contain moving parts to achieve the necessary motions for the cutting process, known as cutting speed, feed, and depth of cut. The number of moving parts and the type of motions corresponding to them differ according to the machine tool category. The change of the positions of the moving parts leads to a change in the static and dynamic performance of the machine tool, as well. There are two forms of position dependency. The first is that the position of some parts changes before the cutting process begins, but while the cutting is happening the position remains constant; for example, see the spindle head position in the case of open category milling machine tool shown in Figure 7" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0003488_msec2015-9243-Figure5-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0003488_msec2015-9243-Figure5-1.png", + "caption": "Figure 5: Layers of the deposit path and master geometry (top) and layers trimmed with a master surface", + "texts": [ + " During the calculation, the paths get trimmed with the previous defined boundary. In a dialog the user can influence the track offset and the tolerances used to calculate the equidistant tracks. The result of the whole procedure is visualized in Figure 2. To generate multiple layers, the initial layer is copied and moved in Z-direction. To avoid a waste of material (powder) and to reduce the effort in further post processing steps (e.g. milling), the resulting deposit path gets trimmed on a master geometry (see Figure 5). This master geometry is constructed using a CAD software on a basis of the CAD model of the non-worn part. As a last step, these trimmed tracks gets converted into a NC program using the post processor of LMDCAM2. Figure 6 shows a part being processed with a NC program generated by LMDCAM2. 3 Copyright \u00a9 2015 by ASME Downloaded From: http://proceedings.asmedigitalcollection.asme.org/ on 01/28/2016 Terms of Use: http://www.asme.org/about-asme/terms-of-use 5-AXIS PROCESSING The metallic bonding of the deposited material onto the surface of the substrate depends on an angle between the welding head and the surface itself" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0002586_edpc51184.2020.9388196-Figure13-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0002586_edpc51184.2020.9388196-Figure13-1.png", + "caption": "Fig. 13. Prototype parts, LBM-printed and finished test gear with coatings (left), FDM-printed gear body (right).", + "texts": [ + " This prevents damage to the coating in the tooth mesh due to excessive temperatures. It should also be noted that the resulting temperature distribution is similar to deep toothing with splash lubrication, which is typical for electric drives. This prevents an increased thermal load on the entire system. Based on the design and simulation of the new gear stage, prototype parts have already been produced using the manufacturing processes explained in section III. This enables proof of feasibility and quality of the components to be produced later. The prototypes are shown in Fig. 13. On the left side LBM-printed gears with the macro geometry of the pinion shaft are shown. After production, the gears are hardened and ground to achieve the required quality. The finishing of LMB gears does not show any noticeable differences to conventionally manufactured gears. Afterward, the gears are coated (see front gear Fig. 13 left), whereby a high surface quality could be proven with a measured coating thickness of 2.5 \u00b5m. Furthermore, the first test prints of the FDM wheel body have already been printed, which is shown in grey on the right in Fig. 13. This wheel body is designed in two parts. It will later be connected to the steel gear rim and the output shaft. Concluding the paper, extensive research work in various fields has been performed, as a basis for realizing nextgeneration powertrains: LBM manufacturing and processing of 20MnCr5 material, plastic gear bodies with specially designed infill gyroid structures, gear mesh excitation simulation with test bench validation [24], low-loss design of the gear mesh, optimized micro geometry [15], as well as coating of ground teeth of gear wheels12" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0003907_s0263574714000058-Figure2-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0003907_s0263574714000058-Figure2-1.png", + "caption": "Fig. 2. (Colour online) Illustration of a MANET topology of a swarm. Dashed lines represent the link quality between the pairs of robots, thinner arrows represent the force vectors regarding each chosen neighbor, and larger arrows represent the resulting force vectors that ensure MANET biconnectivity.", + "texts": [ + " In other words, the initial deployment was able http://journals.cambridge.org Downloaded: 01 Jul 2014 IP address: 155.97.178.73 to ensure that each exploring robot would be able to communicate with k neighbors from the same swarm, k \u03b5N, thus ensuring that the MANET is k-connected. After the initial deployment process is concluded, robots explore the environment while ensuring the same k-connectivity of the swarm by defining \u03c74 [t] as a set of attractive and repulsive forces.12 Let us consider the following illustrative example presented in Fig. 2 in which it is necessary to guarantee a biconnected network (k = 2). As it is possible to observe, robot 1 chooses robot 2 and 4 as its nearest neighbors since they are the nearest ones or the ones that present the higher signal quality. The link between robot 1 and 2 corresponds to the ideal situation, such that any attractive or repulsive force is necessary. However, robot 4 is too far away from robot 1, thus resulting in an attraction virtual force toward it. Robot 2 chooses robot 3 and 4 as its nearest neighbors since robot 1 has first chosen robot 2" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0001945_s00707-020-02847-9-Figure2-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0001945_s00707-020-02847-9-Figure2-1.png", + "caption": "Fig. 2 Square panel with clamped boundary condition", + "texts": [ + " This element consists of eight nodes with six degrees of freedom at each node. Its kinematics allow for the finite membrane strain, which is desirable for this proposed analysis. The analysis was performed in four steps. First, we carried out a structural static analysis to obtain the in-plane stress stiffness due to the prescribed in-plane shear deformation. The in-plane shear deformation was applied to one edge of the laminated composite panels horizontally, while the opposite edge was fully restrained with all degrees of freedom fixed, as depicted in Fig. 2. In the second, we made use of the static analysis performed in the first step to carry out eigen buckling analysis. The lowest eigenvalues yield the linear critical buckling load, such that: {Fcr} = \u03bb {F} (5) where {Fcr} is the critical linear buckling load, and \u03bb are the eigenvalues of Eq. (1), without taking into account the nonlinear stiffness matrix. Due to nonlinearity in the kinematics of the wrinkling, the eigenvalues do not necessarily predict the critical load (in-plane shear deformation) for the onset of wrinkling", + " [9] that for problems exhibiting stress concentration due to nonlinear geometrical effects mesh refinement can greatly influence our ability to obtain consistent and reliable wrinkling patterns. Accordingly, we explored the use of different mesh densities on the fidelity of the solution. The effect of mesh refinement on the wrinkling patterns at three different locations including near the corner of the panel, Fig. 4 Out-of-plane deformation along line AB Fig. 5 Out-of-plane deformation along line CD where maximum stress intensity occurs, was studied (see Fig. 2). A normalized element size ratio is defined by \u03b3 = en\u00d7n/e100\u00d7100, where en\u00d7n and e100\u00d7100 are the typical element sizes for mesh density of n, and 100 elements along each sides of the initially selected square panel. Six mesh densities corresponding to different ratios \u03b3 were generated. Table 1 summarizes the information concerning the different mesh refinements, where a larger value of \u03b3 indicates a finer mesh. In all cases, a horizontal 1 mm displacement is applied to investigate the wrinkling development due to in-plane shear deformation. It is assumed that the panel was made of five laminas with the orientations and material properties defined in Table 2. The out-of-plane deformations along three different directions, including the close vicinity of the corner region (see Fig. 2), are provided in Figs. 4, 5, and 6. For all cases considered, the wrinkle numbers and their half wavelength for different mesh refinement was the same. However, for the finer mesh (\u03b3 = 0.6), the wave crest and wave trough positions are switched, and upon the further mesh refinement their positions remained consistent. Therefore, we concluded that the selection of element size of \u03b3 = 0.6 would be sufficient for consistent and accurate development of wrinkles, and this mesh refinement was selected for the remainder of the work. 3.2 Effect of the ply angle One of the objectives of this work is to find the influence of the ply angles on wrinkle patterns. This will help us to avoid or suppress the wrinkles under the prescribed shear deformation. Accordingly, different orientations of lamina, as listed in Table 3, were examined. In all considered cases, it is assumed that the layers have the same thickness, material properties, and identical imperfections. The out-of-plane displacement along the diagonal line (EF in Fig. 2) of the panel for different layouts is provided in Figs. 7, 8, and 9. Figure 7 reveals that deploying the outer layers with 0\u25e6 orientation suppressed the wrinkles close to the vicinity of the panel corners, while at the middle section of the panel the maximum out-of-plane displacements were similar to those observed in layout-2. The number of wrinkles and their half-wave length were identical for these two considered layouts. Figure 8 shows that the wrinkling patterns of layout-3 and layout-4 were exact opposite of the wrinkling patterns of layout-2 and layout-1, respectively", + " 10a and d), while the maximum and minimum locations of the deformation had been switched as predicted in Figs. 7 and 8. Figures 10b and c illustrate similar out-of plane deformation behavior for layout-2 and layout-3. The wrinkling patterns of layout-5 and layout-6 are identical to layout-3 but with smaller amplitude, as depicted in Figs. 10e and f. This similarity can be explained with the aid of the major and minor principal stresses which are shown in Fig. 11 along the main wrinkle wave deformed almost at 45 \u25e6 (coinciding with line GH in Fig. 2) to X-axis. According to the tension field theory [1] in a two-dimensional stress field, wrinkles are formed along the major principal stress S1, and since the membrane cannot tolerate any compressive load they will wrinkle along the minor principal stress S2. Accordingly, Fig. 11 indicates that as we move from point G (Fig 2) toward point H, the major principal stress reaches a constant value with the growth in its amount at the corners. Concurrently, the minor principal stress reaches a negligible value along the wrinkle wave. It should be noted that for layout-1 and layout-2, the value of S2 around the top corner (point H) is not zero, while for the rest of the layouts this minor stress takes a small value around the lower corner. In other words, the wrinkles for layout-1 and layout-2 start to extend from the lower corner along line GH, where the minor principal stress is negligible, while for the other layouts, the wrinkle extension starts at a distance from the lower corner toward the top corner. In addition, it is observed that since layout-5 depicts higher resistance against wrinkling, the major principal stress at the corner of the panel was lower compared to the other layer configurations. In order to find out how the wrinkles develop and grow for layout-5, during a given load increment, the major principal stress contour and out-of plane displacement along line EF (see Fig. 2) are plotted in Figs. 12 and 13, respectively. It is observed that, increasing the shear displacement at the top edge of the panel, the major principal stress field starts to develop along the diagonal line of the panel and coincides with the wrinkle direction. Referring to Fig. 13, the amplitude and pattern of the wrinkles also varied in accordance with the major principal stress. It can be seen that for the first 10% of total applied in-plane shear displacement, the out-of plane deformation follows the seeded imperfection" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0003878_s11434-014-0376-5-Figure9-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0003878_s11434-014-0376-5-Figure9-1.png", + "caption": "Fig. 9 The force diagram of the i-th link", + "texts": [ + " Therefore, the forelimb model is chosen as the object of force analysis, which could be also applied to the hind limb. The horse body is keeping stationary when standing. Connecting with the horse body permanently, the scapula of the forelimb is also stationary, as showing in Fig. 8. The ground reaction force to the forelimb foot is F. To keep the leg equilibrium, the drive moments added to each joint are denoted by M1, M2, M3 and M4. Each link in the five-bar mechanism must be equilibrium by external forces and external moments. Figure 9 shows the force diagram of the i-th link. Mi-1,i represents the i-th link moment exerted by the (i-1)-th link, Mi?1,i is the i-th link moment exerted by the (i ? 1)-th link. Fi - 1,ix and Fi - 1,iy refer to the x component force and y component force of the Fi-1,i exerted by the (i-1)-th link on i-th link, Fi?1,ix and Fi?1,iy stand for the x component force and y component force of the Fi?1,i exerted by the (i ? 1)-th link on i-th link. Therefore, the balance constraint equations about forces and moments at i-th link are obtained, as the following shows: F * i\u00fe1;i \u00fe F * i 1;i \u00bc 0; \u00f06\u00de M * i 1;i \u00feM * i\u00fe1;i \u00fe F * i 1;i \u2018 " + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0000050_expat.2019.8876524-Figure3-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0000050_expat.2019.8876524-Figure3-1.png", + "caption": "Fig. 3. Laboratory prototype - Design in Solidworks \u00ae.", + "texts": [ + " For the complete execution of this process it was necessary to develop a laboratory base system that would consider all the necessary components, that would allowed the control of the entire process. Some of these were: 1) the auxiliary insert and lock packaging components; 2) Coils; 3) Connectors; 4) Pneumatic cylinders and base brackets; 5) Pneumatic cylinder locking components; 6) Electrovalve; Packaging; 7) Path guides; 8) Programmable Logic Controller; 9) Roller Base; 10) Sensors; 11) Laboratory base prototype; 12) System border. The numbering previously represented can be observed in the assembly Fig. 3 and Fig. 8. The cylinders base supports allowed the pneumatic cylinder to be kept in a resting position along the base of the system. Each pneumatic cylinder was fixed by two brackets and each were placed on the end of the smaller cylindrical outer surface. These supports were composed of six main components designed with the possibility of being easily disassembled, as can be verified in Fig. 4. The product placement component served as an auxiliary accessory for the pneumatic cylinder rod, which made the product placement inside the package, when it was needed in the required position, and add the closing system of the left side flap", + " The results allowed to conclude that the component would be capable to sustain the efforts predicted. Some of the results through this analysis and simulation, can be seen in Fig. 6 and Fig. 7. With the components and systems defined, the laboratory base prototype that would serve as the basis for controlling and supervising the system was assembled, shown in the Fig. 8. The installation of the electropneumatic process was guided by a model developed with the use of the 3D modelling program, Solidworks\u00ae (see Fig. 3), in order to try to ensure the minimum tolerances necessary for the correct operation of the process. To program the PLC SIMATIC S7-1516-3 PN/DP developed by Siemens, different programs standards can be used. The IEC 61131-3 includes several programming languages, namely the Structured Text (ST), Function Block Diagram (FBD), Ladder Diagram (LD), Instruction List (IL), Sequential Function Chart (SFC). The program TIA portal, developed by Siemens\u00ae, provides a set of tools and functions necessary for the implementation of automation tasks mounted on a single platform of cross-programs" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0000317_978-3-030-35699-6_38-Figure17-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0000317_978-3-030-35699-6_38-Figure17-1.png", + "caption": "Fig. 17. 3D diagram of a system with multiple coils [3]", + "texts": [ + " Moreover, in case of a non-deformable contact, the non-magnetic object would have to be very resistant to the very important instantaneous acceleration during the shock. In a further work, the structure of our coil gun will be optimized dividing the coil in several smaller coils activated sequentially. For example, scenarios using 2, 3 or 4 coils (each one having a half, a third or a quarter of the total turns of the initial coil), triggered sequentially by software or using a position sensor instead of a single coil as shown in Fig. 17 will be evaluated as in [2,7]. This will lead to have successively a maximum current on each coil when the rod is optimally placed in the coil, leading to a increased projectile speed. Eighteenth International Middle East Power Systems Conference (MEPCON), pp. 506\u2013511, December 2016. https://doi.org/10.1109/MEPCON.2016.7836938 2. Bencheikh, Y., Ouazir, Y., Ibtiouen, R.: Analysis of capacitively driven electromagnetic coil guns. In: The XIX International Conference on Electrical Machines - ICEM 2010, pp" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0000471_physreve.101.013002-Figure9-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0000471_physreve.101.013002-Figure9-1.png", + "caption": "FIG. 9. Schematic of a beam segment of length a0, width w, and thickness h undergoing bending deformation by an angle \u03b8 such that the section A deforms to A\u2032, the radius of curvature after deformation is R.", + "texts": [ + " The results presented in this work may provide some insights into the manufacturing process of paper and fabric sheets. ACKNOWLEDGMENTS This work was supported by Natural Sciences and Engineering Research Council (NSERC) of Canada through the Collaborative Research and Development grant (501467-16). S.P. would like to acknowledge funding through the Discovery Grants program from NSERC, Canada. We thank Shubham Agarwal for his inputs on the viscous bending model. APPENDIX A: DERIVATION OF VISCOUS BENDING FORCE Figure 9 shows a beam segment undergoing bending deformation by an angle of \u03b8 . During bending, viscous stresses are developed at the section in order to slow down the deformation process. For any point on the beam\u2019s section, viscous stress can be written in terms of dynamic viscosity (\u03bc) and strain rate (\u03b5\u0307) at that point of the section as \u03c3v = \u03bc\u03b5\u0307 (A1) 013002-7 For a point on the section at a distance y from the neutral axis (positive direction of y is shown in Fig. 9), the axial strain can be written as \u03b5 = y R . (A2) Hence the strain rate will be \u03b5\u0307 = \u2212yR\u0307 R2 . (A3) Using inextensibility of the neutral axis, it can be shown that \u03c3v = \u03bc y\u03b8\u0307 a0 , (A4) where \u03b8\u0307 is the rate of bending deformation of the segment. The viscous bending moment can be written as Mv = \u222b h/2 \u2212h/2 \u03c3vyw dy = \u03bc \u03b8\u0307w a0 h3 12 . (A5) Finally the viscous bending force can be calculated: fv = Mv a0 = ca\u03b8\u0307h2 12a0 , (A6) where ca = \u03bcwh/a0 is the phenomenological damping coefficient. APPENDIX B: ELASTIC MODEL VALIDATION Consider a thin film under uniaxial load without substrate" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0003499_muh-1404-6-Figure8-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0003499_muh-1404-6-Figure8-1.png", + "caption": "Figure 8. (a) Cutting mechanics, (b) axial and radial depth of cut.", + "texts": [ + " This integration leads to a realistic representation of the overall performance of the machine tool. Many works have been carried out on different aspects linked to cutting force prediction. Recently, researchers used the FEM to simulate the cutting process and hence predict the cutting forces and chip morphology. However, simulation of the cutting process using the FEM is not crucial to this paper. An analytical method has been used instead. A cutting force model was developed for conventional end-milling operations [14]. Figure 8 illustrates the cutting mechanics, and the axial and radial depth of cut. The expressions for the cutting force model were derived as follows. Fx=Fu[ (\u03b8e\u2212\u03b8s)\u2212Pf ( sin2\u03b8e\u2212sin2\u03b8s ) \u22120.5 (sin 2\u03b8e\u2212 sin 2\u03b8s) ] Fy= \u2212Fu[Pf (\u03b8e\u2212\u03b8s)\u2212 ( sin2\u03b8e\u2212sin2\u03b8s ) \u22120.5Pf (sin 2\u03b8e\u2212 sin 2\u03b8s) ] Meanwhile, Fu=Km.r.ft/tan\u03b2/2 ; ft= feedrate/(rpm \u2217 no.of tooth). Fx is the normal direction cutting force (N), Fy is the feed direction cutting force (N), r is the tool radius (mm), \u03b8s is the integrating start angle (rad), \u03b8e is the integrating end angle (rad), \u03b2 is the helix angle, ft is the feed per tooth, the proportional factor Pf was usually selected as 0" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0002989_978-3-642-16214-5_9-Figure9.4-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0002989_978-3-642-16214-5_9-Figure9.4-1.png", + "caption": "Fig. 9.4. Shifting mechanism of a 5-speed gearbox for front-transverse mounting (VW Golf III). 1 5th gear shift fork; 2 selector shaft detent; 3 5th gear lock; 4 connecting bar; 5 front selector rod; 6 rear selector rod; 7 selector lever; 8 relay lever; 9 selector bar bearing bush; 10 bearing plate; 11 shift lever bearing housing; 12 5th gear end stop; 13 1st/2nd gear end stop", + "texts": [ + " The variety of designs and combinations of internal and external shifting elements is virtually unlimited, so this section will restrict itself to describing the basic principles. Sections 12.1 and 12.2 explain some typical examples of existing designs. The morphological matrix gives an overview of shifting elements (Table 9.1). Table 9.1. Morphological matrix for shifting elements Parameters Configurations (shifting elements) External shifting system Linkage Multi-bar linkage Cable control Shift-by-wire \u2013\u2013\u2013\u2013\u2013\u2013\u2013\u2013\u2013 Example Single bar actuation [9.12] VW Golf III, Figure 9.4 Shift and select cable MB A-class [9.8] Automatic transmission ZF 6 HP 26, Figure 12.25 \u2013\u2013\u2013\u2013\u2013\u2013\u2013\u2013\u2013 Generate shifting force Mechanical, manual effort Electromechanical Electrohydraulic Electropneumatic Example Shift lever Transmission actuator, Figure 12.34 Electromagnetic clutch Gear selection and shifting force transformation Selector bars, levers Ball joint, four-bar linkage Selector shaft, turning shaft Gearshifting drum \u2013\u2013\u2013\u2013\u2013\u2013\u2013\u2013\u2013 Example 3-bar shift mechanism, Figure 9.3 VW Golf III, Figure 9.4 Figure 12.3 (ZF) Smart (Getrag), Figure 12.14 \u2013\u2013\u2013\u2013\u2013\u2013\u2013\u2013\u2013 Shifting Shift fork Swing fork Piston \u2013\u2013\u2013\u2013\u2013\u2013\u2013\u2013\u2013 \u2013\u2013\u2013\u2013\u2013\u2013\u2013\u2013\u2013 Example Figure 12.9 (MB) Figure 12.1 (VW), Figure 12.3 (ZF) Conventional automatic transmission, Figure 12.23 (MB) \u2013\u2013\u2013\u2013\u2013\u2013\u2013\u2013\u2013 \u2013\u2013\u2013\u2013\u2013\u2013\u2013\u2013\u2013 Frictional connection Single/Multicone Spreader ring Multi-plates Belt Sprag Example Cone synchronizer Porsche synchronizer Clutch, brake, Figure 6.30 Brake, Figure 6.35 Freewheel (AT), Figure 6.30 Positive engagement Dogs Pins Sliding gear Draw key \u2013\u2013\u2013\u2013\u2013\u2013\u2013\u2013\u2013 Example Figure 9", + " By selecting appropriate lever proportions, gearshift effort can be reduced at the expense of increasing the shift stroke (see also Figure 12.3). Kinematics is needed to describe the breakdown of gear changing into selecting and shifting movements. In order to avoid load changing reactions and vibrations in the gearshift lever and gearshift housing, the external gearshift system is suitably decoupled from the gearbox and the car body. The reverse gear is safeguarded against operating errors with a locking device, depending on the philosophy of the vehicle manufacturer. Figure 9.4 shows the external shifting elements, and some of the internal shifting elements of a passenger car with a transverse-mounted gearbox. The remote shifting is mechanical, using a four-bar linkage. The shifting arrangement shown was produced in series by VW in the Golf III until the mid-1990s. Instead of costly linkage kinematics, it is popular to use cable controls, especially for passenger cars with transverse-mounted gearboxes. Shifting and selecting forces are transmitted from the gearshift lever in the shift housing to the internal gearshift system [9" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0002728_978-3-642-03016-1_3-Figure3.16-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0002728_978-3-642-03016-1_3-Figure3.16-1.png", + "caption": "Fig. 3.16 1. Stable balancing with active control moment M in the middle joint (B) simulating the subtalar joint. 2. When the control moment is positioned higher (B1), the mechanical demands are increasing: M1 > M", + "texts": [ + " Those better conditions possibly were: restoration of a normal synergetic angulation in the ankle and knee, and less stump pain associated with such synergetic angulations. The basis for the hypothesis could be derived from the mathematical modeling and biomechanical analysis of the one-leg standing in the frontal plane (Pitkin 1970; Velikson et al. 1973). In the study, a subject performed the task of one-leg balance while standing on a skate. A corset prevented trunk movements, and arms were held tight to the torso. The dynamic of balancing in the frontal plane was modeled by two joined mathematical pendulums attached to the ground as shown in the picture (Fig. 3.16, 1) where free joint A represents the contact between the skate and ground; powered joint B stands for subtalar joint; C is the body COM. The dynamic of balancing was modeled by a system with two degrees of freedom: twolinked reverse pendulum whose segments had length l1and l2; masses m1\u0438 m2and whose balancing is provided by a control moment in joint B (Fig. 3.16, 1). A problem to be mathematically solved was whether it would be possible to stabilize the pendulum with a moment M(j, j, a, a)& & in its position of unstable equilibrium. The Lagrange equations for the model are: + m2 l1 l2 sin a (2ja \u2212 a2) \u2212 (m1 + m2)gl1 sin j + m2 gl2 sin(j\u2212a) = 0, j [\u2212 m2l2 + m2l1l2 cos a ] + a m2l1 \u2212 j m2 l1 l2 sin a \u2212 m2 gl2 sin(j\u2212a) = M. j [(m1 + m2) l1 + m2l1 \u2212 2m2 l1 l2 cos a] + a [\u2212 m2 l2 m2 l1 l2 cos a]2 2 2 22 2 We considered the task of stabilization by first approximation for the equivalent system where a = p + j + b (3", + " x3 x4 g g m l l+ ++ + \u00e9 \u00ea \u00ea \u00eb \u00e9 \u00ea \u00ea \u00ea \u00ea \u00ea \u00eb \u00e9 \u00ea \u00ea \u00ea \u00ea \u00ea \u00eb \u00f9 \u00fa \u00fa \u00fb \u00f9 \u00fa \u00fa \u00fa \u00fa \u00fa \u00fb \u00f9 \u00fa \u00fa \u00fa \u00fa \u00fa \u00fb \u00e6 \u00e7 \u00e8 \u00f6 \u00f7 \u00f8 \u00e6 \u00e7 \u00e8 \u00f6 \u00f7 \u00f8 \u00e6 \u00e7 \u00e8 \u00f6 \u00f7 \u00f8 \u00e6 \u00e7 \u00e8 \u00f6 \u00f7 \u00f8 \u00e6 \u00e7 \u00e8 \u00f6 \u00f7 \u00f8 To assure the solvability of the problem of stabilization, we proved the nonsingularity of the matrix W = {QPQ, K, PQn-1}, n = 4 W = V \u00b9 0. That condition is also sufficient for optimal stabilization (Boltyansky 1969) when the criterion got quality of the process is the functional I = aij xi xj + bij ui uj + dt ,\u00f2 \u00a5 t0 S n i,j=1 S n i,j=1 \u00e6 \u00e7 \u00e8 \u00e6 \u00e7 \u00e8 where \u03b1ij,\u03b2ij are coefficients. In particularly, the growth of I due to the growth of the ratio l2/ l1 tells us about the increase of work for control over stabilizing the equilibrium if a joint B with moment M shifts upward (M1) to position 2 (Fig. 3.16, 2). It was mathematically proven that asymptotic stability of the system can be provided if the power control function M is a linear combination of the both pendulums\u2019 angles and velocities, as is a case in the muscle control. The other result was that when the powered joint moves up (position B1), the system requires more mechanical work and power to be produced by the control function: M1 > M. It means that with the stiff prosthetic ankle and consequently decreased performance of the existing knee joint, the balancing is less efficient" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0000327_j.mechmachtheory.2019.103729-Figure16-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0000327_j.mechmachtheory.2019.103729-Figure16-1.png", + "caption": "Fig. 16. Intermediate loop mechanism attached to the front end loop.", + "texts": [ + " However, according to the analysis of the other two loops, the DOF of the components 2 \u2032 and 4 \u2032 involved in L(1) becomes 0, and two rotational DOFs are lost in the component 3 \u2032 , and only one movement freedom along the Z axis is left. So far, we have completed the over-constraint analysis of the three front-end loops and the constraint analysis between the three loops. Next, the constraint relationship of the intermediate loop mechanism has been analyzed. The intermediate loop is attached to the two front-end loops. This means that the mechanism shown in Fig. 16 will be analyzed, which is composed of the front-end loops L(1), L(2) and the intermediate loop L(8). Unlike the three-loop mechanism in the previous chapter, the third loop of the mechanism does not including higher rank component. That is, the highest rank component of the third loop (intermediate loop) is one in the front-end loop, which is used to add new constraints to the higher components in the front-end loop. For such a mechanism, first the intermediate loop L(8) is analyzed directly and the relative DOF between the components in the loop must be found" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0000327_j.mechmachtheory.2019.103729-Figure5-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0000327_j.mechmachtheory.2019.103729-Figure5-1.png", + "caption": "Fig. 5. Kinematic diagram of the Rubik\u2019s Cube unit.", + "texts": [ + " 1) Kinematic diagram of first configuration of Rubik\u2019s Cube unit mechanism Firstly, the sub-adjacency matrix of a motion unit mechanism is extracted from the adjacency matrix which is shown in formula (2) of the Rubik\u2019s Cube first configuration, and the sub-adjacency matrix is shown in formula (4) . C 1 = \u23a1 \u23a2 \u23a2 \u23a2 \u23a2 \u23a2 \u23a2 \u23a2 \u23a2 \u23a2 \u23a2 \u23a2 \u23a3 2 0 3 E 1 4 0 9 E 1 2 \u2032 E 1 3 \u2032 0 4 \u2032 0 9 \u2032 S 1 E 1 0 E 1 0 0 E 1 0 S 1 0 E 1 0 E 1 0 0 E 1 S 1 E 1 0 E 1 0 0 0 0 S 1 + R 1 E 1 0 0 0 0 E 1 0 S 1 + R 1 0 E 1 0 0 E 1 0 E 1 S 1 0 0 E 1 0 0 E 1 0 S 1 + R 1 S 1 S 1 S 1 S 1 + R 1 S 1 + R 1 S 1 S 1 + R 1 0 \u23a4 \u23a5 \u23a5 \u23a5 \u23a5 \u23a5 \u23a5 \u23a5 \u23a5 \u23a5 \u23a5 \u23a5 \u23a6 (4) Combining the positional relationship of the Rubik\u2019s Cube mechanism, a mechanism diagram of the unit mechanism in the aligned state is drawn as shown in Fig. 5 . It can be seen in the sub-adjacency matrix and the kinematic diagram of the Rubik\u2019s Cube unit that each movable component forms a spherical pair with the same 3-D cross component. So a common adjacency matrix C I 1 can be separated in the sub-adjacency matrix C 1 . C I 1 = \u23a1 \u23a2 \u23a2 \u23a2 \u23a2 \u23a2 \u23a2 \u23a2 \u23a2 \u23a2 \u23a2 \u23a2 \u23a3 2 0 3 0 4 0 9 0 2 \u2032 0 3 \u2032 0 4 \u2032 0 9 \u2032 S 1 0 0 0 0 0 0 0 S 1 0 0 0 0 0 0 0 S 1 0 0 0 0 0 0 0 S 1 0 0 0 0 0 0 0 S 1 0 0 0 0 0 0 0 S 1 0 0 0 0 0 0 0 S 1 S 1 S 1 S 1 S 1 S 1 S 1 S 1 0 \u23a4 \u23a5 \u23a5 \u23a5 \u23a5 \u23a5 \u23a5 \u23a5 \u23a5 \u23a5 \u23a5 \u23a5 \u23a6 (5) Thus the adjacency matrix C O 1 of the outer loop mechanism can be obtained" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0002759_b978-0-12-804560-2.00011-0-Figure4.2-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0002759_b978-0-12-804560-2.00011-0-Figure4.2-1.png", + "caption": "FIGURE 4.2 IP type models with BoS constraint lf = lh + lt (the foot length). (A) Nonlinear (constant-length) IP-on-foot model. The ankle torque (ma = 0) alters the GRM, thereby displacing the CoP and accelerating/decelerating the CoM. (B) Nonlinear (constant-length) IP-on-cart model. The CoM horizontal motion is determined by the displacement of the cart, associated with that of the CoP. (C) Variable-length, constant-height LIP-on-cart model. In all models, the GRF acts along the line connecting the CoP and the CoM.", + "texts": [ + " Also, the appearance of the gait is \u201cape-like,\u201d i.e. it is quite different from the erect human gait. The LIP model can be improved by adding a model of the foot. It is assumed that the foot is in multipoint/line contact with the supporting ground. The foot serves as an interface for the impressed/reaction forces that act via the contact points. Note that the resultant force is the GRF, f r . This force is applied at a specific point called Center of Pressure (CoP). This situation is depicted in Fig. 4.2A, where xp denotes the CoP. In the case of multiple links in contact with the environment, CoPs are defined for each link. These CoPs exist only within the Base-of-Support (BoS) boundaries of the links in contact. A number of studies in the field of biomechanics have clarified the important role of the CoP in human balance control [85,10,162]. The importance stems from the fact that the tangential components of the Ground Reaction Moment (GRM) are zero at this point. Thus, the location of the CoP within the BoS can be used as an indicator of contact stability", + " It is worth noting that in the early studies on humanoid robots, the special case of flat ground was exclusively considered and the term \u201czero-moment point\u201d (ZMP) was coined [158]. On flat ground, the ZMP is identical to the CoP/net CoP (in the cases of single/double stance). On nonflat ground, though, these points do not coincide [120]. In the following discussion, whenever ambiguity is avoidable, the terms CoP and ZMP will be used alternatively. Linearized IP-on-Foot Model A constant-length (nonlinear) IP model attached to a foot is shown in Fig. 4.2A. The GRF f r acts at the CoP, xp, along the line determined by the CoP and the CoM. The two components of this force result from gravity and the CoM acceleration. Note that the moment at the CoP is zero. Note also that the ankle joint is not anymore passive, as it was with the LIP model. The ankle torque ma = 0 alters both, the CoM acceleration and the CoP. The equation of motion of the simple IP is Ml 2\u03b8\u0308 = Mgl sin \u03b8 + ma. (4.7) The CoP is determined from the GRM balance equation as xp = \u2212 ma M(g + z\u0308g) = (4", + " In the special case of a fixed CoP, denoted as x\u0304p, the solution to the above equation can be obtained in closed-form. We have xg(t) = ( xg0 \u2212 x\u0304p ) cosh(\u03c9t) + vg0 \u03c9 sinh(\u03c9t) + x\u0304p (4.12) = 1 2 ( xg0 \u2212 x\u0304p + vg0 \u03c9 ) e\u03c9t + 1 2 ( xg0 \u2212 x\u0304p + vg0 \u03c9 ) e\u2212\u03c9t + x\u0304p. (4.13) This equation plays an important role in balance stability analysis (cf. Chapter 5). IP-on-Cart Model The ODE (4.11) clarifies that the CoM motion trajectory can be controlled via the forcing term, i.e. by appropriate position control of the CoP. This approach resembles the way of balancing a stick on the palm. The respective model is shown in Fig. 4.2B. This model is widely used in the field of control under the name \u201cIP-on-cart model.\u201d Note that the distance between the CoM and the cart (the CoP) is assumed constant. The equation of motion in the horizontal direction is of the same form as (4.11). Note, however, that the coefficient \u03c92 is not anymore constant; now we have \u03c9(t) = \u221a z\u0308g + g zg \u2212 zp . (4.14) The vertical CoP coordinate zp is usually assumed constant (e.g. zero on a level ground). Solving (4.11) for the CoP with the above expression for \u03c9, one obtains xp = xg \u2212 x\u0308g z\u0308g + g zg = xg \u2212 frx frz zg", + "15) Here frx and frz denote the two components of the GRF. This equation is used for balance control under the name \u201cZMP manipulation\u201d or \u201cindirect ZMP\u201d control [80,144], as explained in Section 5.4. LIP-on-Cart Model A variable \u03c9, as in (4.14), complicates the analysis since the solution cannot be obtained in such a simple form as that for the linearized IP model. To alleviate this problem, the taskbased (LIP) constraint zg = z\u0304g = const \u21d2 \u03c9 \u2261 \u03c9 = \u221a g/z\u0304g = const (4.16) is imposed. The resulting model is shown in Fig. 4.2C. Note that the form of the equation of motion, (4.11), will be preserved. Thus, any displacement of the cart (or equivalently, the CoP) will alter the horizontal CoM acceleration via the term \u03c92xp. This also implies respective changes in the direction of the GRF. We have frx Mg = xg \u2212 xp z\u0304g (4.17) and hence, of the GRM, i.e. (xg \u2212 xp)Mg = z\u0304gfrx . Note that with this model, the vertical component of the GRF equals the gravity force; frz = Mg. Furthermore, using (4.8) with z\u0308g = 0, the equation of motion can be represented as x\u0308g = \u03c92xg + ma Mz\u0304g ", + "32) where c \u2208 p is a set of constants, p denoting the number of polygon sides. The equation of the ith side4 is given by psirp = ci , i \u2208 {1,p}, psi denoting the ith row in Ps . The coupling problem pertinent to the equation of motion will be addressed as in Section 4.4.1, by restricting the CoM to move within a plane. In this case, the spherical IP model is referred to as the 3D LIP [54]. The equation of motion is then determined by two ODEs that are decoupled in x and y, as in (4.31). The 3D LIP model can thus be regarded as an extension of the LIP-on-cart model (cf. Fig. 4.2C) in 3D. Next, consider the extension of the planar IP-on-foot and the IP-on-cart models (shown in Fig. 4.2A and B, respectively) to 3D. The universal joint at the foot of the respective spherical IP (the ankle joint) is assumed actuated. The equation of motion of each of these models has the form of (4.10). Thus r\u0308g = \u03c92(rg \u2212 rp). (4.33) Here, \u03c9 = \u03c9IP = \u221a g/l for the constant-length (l = const) IP model, while \u03c9 = \u03c9 = \u221a g/z\u0304g for the constant-height (z\u0304g = const) IP model. In the case of a variable-stroke model without the CoM vertical-motion constraint, use \u03c9 = \u03c9(t) as in (4.14). The equation of motion, expressed in terms of the CoP coordinates, assumes then the form of (4.15), i.e. rp = rg \u2212 zg z\u0308g + g r\u0308g. (4.34) Furthermore, the LIP-on-cart model in Fig. 4.2C can also be extended to 3D by replacing the cart with a sphere rolling on a horizontal plane of constant height z\u0304g [23,22]. The tangential components of the GRM, mt = [ mx my ]T , are relevant. They can be obtained from the equation of motion as mt = MgS\u00d7 2 ( 1 \u03c92 r\u0308g \u2212 rg ) , (4.35) 3 A lower-case subscript is used to distinguish a 2D vector from a 3D one. 4 Recall that the curly inequality sign denotes componentwise operation. where S \u00d7 2 \u2261 [ 0 1 \u22121 0 ] . The subscript t signifies that the respective vector quantity is composed of tangential (i", + " \u2022 The representation with the mixed quasivelocity, VM , yields inertial decoupling between the CoM dynamics on one side and the angular momentum and joint-space partial dynamics on the other. The latter two components are inertially coupled; the coupling determines the coupling angular momentum conserving manifold in joint space. The tangent subspace to this manifold at q is called the angular momentum Reaction Null Space. \u2022 The angular component of the spatial momentum is conserved: \u2022 in the absence of external forces; \u2022 when the CRB centroid lies on the line of action of the external force (as with the IP models in Fig. 4.2). Using the appropriate representation of the angular momentum is important in balance control design, as will be shown in Chapter 5. Furthermore, note that with the exception of the flight phase during running and jumping, a humanoid robot is always in contact with the environment. The posture is then typically characterized as single/double stance, or multicontact. In the latter case, multiple interdependent closed loops are formed via the hand contacts, in addition to those at the feet. In this situation, the momentum control objectives have to be determined as constraint-consistent" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0002683_012034-Figure16-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0002683_012034-Figure16-1.png", + "caption": "Figure 16. Finite element model for modeling of the experiments.", + "texts": [ + " The shapes of the specimens and a basic cell are presented in Fig. 14. IOP Conf. Series: Materials Science and Engineering 986 (2020) 012034 IOP Publishing doi:10.1088/1757-899X/986/1/012034 Tests were carried out with Instron 8850 Axial-Torsion System. The photographs of the specimens during tests are shown in Fig. 15. FE modeling of uniaxial tension and compression was performed in SIMULIA Abaqus implicit solver with the same finite element models both for tension and compression. Constructed model is shown in Fig. 16. IOP Conf. Series: Materials Science and Engineering 986 (2020) 012034 IOP Publishing doi:10.1088/1757-899X/986/1/012034 The characteristics of the finite-element model are as follows: - element type \u2013 C3D4 (four-node linear tetrahedra); - number of elements \u2013 1 156 665; - number of nodes \u2013 273 733. During the simulation the bottom face of the specimen is fully fixed, the top face has a prescribed axial displacement. One run takes about five hours. On Fig. 17 \u2013 Fig. 18 there are curves obtained from the natural testing and FEA modelling" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0000710_masy.201900100-Figure7-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0000710_masy.201900100-Figure7-1.png", + "caption": "Figure 7. a) Constraints and b) loading conditions.", + "texts": [ + " Such an approach consists of two consecutive analyses: the former, performed by solving independently the thermal problem, allows achieving the temperature distribution by considering as starting condition an engine temperature of 180 \u00b0C; the latter considers the temperature distribution previously predicted as nodal thermal load. FE model consists of 194 104 nodes and 271 421 elements. With reference to Figure 1, brick elements, C3D4, C3D8, and C3D6 (from Abaqus elements library) have been used for the encapsulation component in red and for the rock wool layer in grey. Other parts have been modeled through shell elements, S4 and S3 (fromAbaqus elements library). Relatively to PA6GF30,material properties have been set according to the experimental tests previously described. For the car body (Figure 7), BH220 steel has been considered, while for the rock wool, a Young modulus of 20 MPa and a Poisson coefficient of 0.3 have been considered. Concerning the boundary conditions, body car has been constrained symmetrically has shown in Figure 7a. In addition, the engine cover has been investigated under an acceleration of 4 g applied along the z-axis, as shown in Figure 7b. Engine cover has been constrained in correspondence of A point, simulating the clip aimed to avoid the relative rotation between the fixed and movable counterparts. Other joints have been modeled according to the techniques presented in refs. [14,15]. 5. Results Figure 8a shows the temperature distribution over the engine encapsulation system. Temperatures are shown in \u00b0C. Figure 8b Macromol. Symp. 2020, 389, 1900100 \u00a9 2020 WILEY-VCH Verlag GmbH & Co. KGaA, Weinheim1900100 (2 of 4) Figure 5. Force versus displacement curve provided by the bending test. Figure 6. Strain versus time curve provided by creep test. shows von Mises stresses map affecting the encapsulation system, achieved by considering the temperature distribution as nodal thermal loads for the mechanical analysis. According to Figure 8b, the maximum stress has been predicted to be 61.73 MPa and concentrated nearby the modeled clip (A point of Figure 7b). Figure 9 shows the map of the nodal displacements along z-axis. The maximum displacement has been predicted to be 15 mm in the central part of the engine cover. Such a displacement may be too large leading it to touch the engine. Further investigation is needed. 6. Conclusions In this paper, an innovative engine encapsulation system has been designed, aimed to reduce the CO2 emissions. A thermostructural analysis, based on the FEmethod, has been performed in order to investigate on the structural response of the proposed engine encapsulation system under critical loading conditions" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0000499_icoecs46375.2019.8949996-Figure5-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0000499_icoecs46375.2019.8949996-Figure5-1.png", + "caption": "FIGURE 5. (a) Position trajectories of the multi-vehicle system; (b) Velocity trajectories of the multi-vehicle system.", + "texts": [ + "0367 0.3767 0.2374 0.7099 \u22120.2053 0.2833 0.7099 2.4161 . Solid blue lines in Figs. 5(a) and (b) respectively show the velocity and position trajectories of the multi-vehicle system from t = 0s to t = 40s. Using the star, asterisk, square and diamond to represent the states of each vehicle, the TVF of the multi-vehicle system under switching interaction topologies is illustrated at t = 10s using bold dash-dotted lines, at t = 25s using bold dashed lines and at t = 40s using bold dotted lines. From Fig. 5, we can observe that both the velocities and positions of the multi-vehicle system reach parallel rectangle formations and the parallel rectangles keep rotation. In Fig. 6, the switching signal, formation error, and coupling weights are displayed respectively. Fig. 6(a) shows that the interaction topology G\u03b4(t) of the multi-vehicle system switch every 1s among G1, G2, G3 and G4 randomly. From Fig. 6(b), the formation error of the multi-vehicle system converges to zero which means that the TVF is achieved" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0000050_expat.2019.8876524-Figure7-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0000050_expat.2019.8876524-Figure7-1.png", + "caption": "Fig. 7. Displacement graphical detail along the component.", + "texts": [ + " The development of the base brackets, by means of 3D printing technology, instead of a standard commercial connections was substantiated in the fact that it supports the teaching of this type of technology. The stresses considered were focused on the pressure exerted by the compressed air of the network in the brackets, 7 bar, and the simulation has been performed in a 3D modelling program: SolidWorks\u00ae and SolidWorks Simulation. The results allowed to conclude that the component would be capable to sustain the efforts predicted. Some of the results through this analysis and simulation, can be seen in Fig. 6 and Fig. 7. With the components and systems defined, the laboratory base prototype that would serve as the basis for controlling and supervising the system was assembled, shown in the Fig. 8. The installation of the electropneumatic process was guided by a model developed with the use of the 3D modelling program, Solidworks\u00ae (see Fig. 3), in order to try to ensure the minimum tolerances necessary for the correct operation of the process. To program the PLC SIMATIC S7-1516-3 PN/DP developed by Siemens, different programs standards can be used" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0001619_j.cja.2020.07.025-Figure1-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0001619_j.cja.2020.07.025-Figure1-1.png", + "caption": "Fig. 1 Helicopter slung load system configuration.", + "texts": [ + " Meanwhile, the effects of path constraints and the FCS on the optimization results are discussed at the end of this section. The high-order flight dynamic model is not suitable for this study because of the numerical computational cost of subsequent optimization procedures. Since a helicopter cannot fly at high speed with a slung load, the shock wave at the advancing blade and the dynamic stall at the retreating blade are neglected. The rotor force and moment calculation method for analytically integrating the blade force over the rotor disk are of enough precision in this study. As is shown in Fig. 1, the helicopter and the slung load are both 6-DOF rigid body models. The suspension system consists of a long pendant and a 4-cable sling set connected with each other through a swivel such that the yaw moment is not passed to the helicopter, which is described in detail by Gass- Please cite this article in press as: WANG L et al. Optimization of aerial slung load r 10.1016/j.cja.2020.07.025 away et al.11 A single sling in the suspension system is modeled as a stretch-only spring-damp system with 3-DOF, 1 stretching motion and 2 swing motions" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0003632_0954406214549786-Figure4-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0003632_0954406214549786-Figure4-1.png", + "caption": "Figure 4. Contact forces between worm and worm wheel (right-hand).", + "texts": [ + " Step 5: Perform a static finite element analysis to determine the rotation angle r of the master at UNIV OF CONNECTICUT on May 22, 2015pic.sagepub.comDownloaded from node around the x-axis when the torque T s\u00f0 \u00de is applied on MC. After that, the equivalent tangential meshing stiffness of worm and worm wheel can be calculated by equation (11) kc s\u00f0 \u00de \u00bc T s\u00f0 \u00de r2wg r \u00f011\u00de Step 6: Repeat the five steps above to get the meshing stiffness of the worm gear pair at different tilting angles. Axial stiffness of worm supporting bearings. Figure 4 depicts the contact forces between the worm and worm wheel.26 With the torque T s\u00f0 \u00de applied, the force components can be calculated by equations (12) to (14) Faw \u00bc Ftwg \u00bc T s\u00f0 \u00de=rwg \u00f012\u00de Frw \u00bc Frwg \u00bc Ftwg tan a \u00f013\u00de Ftw \u00bc Fawg \u00bc Faw tan \u00f014\u00de where Faw, Frw, and Ftware the axial, radial, and tangential contact forces applied on the worm, respectively. Fawg, Frwg, and Ftwg are the axial, radial, and tangential contact forces applied on the worm wheel, respectively. is the lead angle on the pitch cylinder of worm" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0000081_978-3-030-32710-1_10-Figure13-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0000081_978-3-030-32710-1_10-Figure13-1.png", + "caption": "Fig. 13 Rotation of a machining tool by means of OM (ABCDEF): a rotation coaxial with the module\u2019s symmetry axis; b eccentric rotation", + "texts": [ + " The novel concept of OD may be used also in many other fields. The examples of applications of OD and its new functional capabilities are established below, e.g. [10, \u2022 Example 1. Possibility of control over of physical-mechanical properties, geometrical shape of contact surface and displacement trajectory (Fig. 10). \u2022 Example 2. Travel of long objects in pipe (Fig. 11). \u2022 Example 3. Travel of object in pipe with possibility of vibroprotection and positioning (Fig. 12). \u2022 Example 4. Possibility of hole drilling in the end of pipe wall (Fig. 13). \u2022 Example 5. Possibility of battering in pipe of wall end (Fig. 14). \u2022 Example 6. OD can connect together, forming some novel mobile selfreconfigurable structures (swarm systems) for various applications (Fig. 15). Example 1 Example of control of geometrical shape of contact surface by the OD is given in Fig. 10. In this mode, in the motion of the OM 1 (ABCDEF) (Fig. 4) of the adaptivemobile parallel spatial robot (Fig. 10), each longitudinal displacement of the rear face ABC and frontal face DEF is preceded by alternating discrete rotations in both directions, with specified increment, relative to the direction of motion", + " 4) in the rods of the lateral faces of the clamping module. It involves coordinated change in the length of those rods at a command from control system 11 (Fig. 4), on the basis of the readings of sensors 10 (Fig. 4) mounted at points 7 (Fig. 4) of the frontal and rear faces of the clamping module and also sensors 5 (Fig. 4) and 6 (Fig. 4) at each of the rods of the lateral faces of the clamping module. Example 4 Rotation and supply of a machining tool (a drill or bit, for instance) by means of OM (ABCDEF) (Fig. 13). In this case, the tailpiece of the tool takes the form of a crankshaft with a rotating bush at the end that is incapable of axial motion. (The rotating bush is not shown in Fig. 13). The rotating bush is clamped by the radial limiters at points of the rods in the frontal face. Each limiter may be rigidly connected with one section of a three section bush intended for capture of the rotating tailpiece bush. The required force is determined on the basis of the sensors 4 in the frontal face. The radial limiters at the rear face are fixed at the internal contact surface, as in previous modes. Then, coordinated change in the length of the rods at the lateral faces brings the cutting section of the tool to the machining point, fixes the axis of tool rotation, and ensures the required cutting force, determined from the readings of sensors 3 at the rods of the lateral faces", + " In tool rotation, coordinated increase in length of the rods at the lateral faces ensures its longitudinal supply with specified force; the generation of impact and vibration effects is possible in combination with tool rotation. In that case, the spatial position, cutting force, and impact and vibration effects are monitored by means of sensors 9 and 10 (Fig. 4) at the radial limiters of the frontal face and sensors 3, 5, and 6 (Fig. 4) at the rods in the lateral faces. The axis of tool rotation may be coaxial with the symmetry axis of OM 1 (Fig. 13a) or eccentric (Fig. 13b). Example 5 Organization of impact and vibration effects by a slotting tool on the end surface (Fig. 14) of a tubular profile by means of OM 1 (Fig. 4). In thismode, the slotting tools are established at each point of the frontal face. (The clamping of the slotting tool\u2019s tailpiece is not shown in Fig. 14). The radial limiters of the rear face are fixed at the internal contact surface as in previous modes. Then, coordinated change in the length of the rods at the lateral faces brings the working sections of the slotting tools into contact with the end surface at the machining site, and machining begins at specified frequency, amplitude, and force" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0001653_er.5776-Figure1-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0001653_er.5776-Figure1-1.png", + "caption": "FIGURE 1 Schematic diagram of MFCs with major components shown. MFC, microbial fuel cells [Colour figure can be viewed at wileyonlinelibrary.com]", + "texts": [ + " Microorganisms in the anode chamber produce electrons (e\u2212) and protons (H+) by degrading organic matter, and biomass as the final product in the MFC.3 In recent years, research on alternative renewable energy sources has grown exponentially, such as solar, wind, tidal, geothermal, and Received: 3 April 2020 Revised: 26 June 2020 Accepted: 28 June 2020 DOI: 10.1002/er.5776 Int J Energy Res. 2020;1\u20139. wileyonlinelibrary.com/journal/er \u00a9 2020 John Wiley & Sons Ltd 1 biomass power generation, while MFCs (Figure 1) have the potential as alternative energy sources.4 MFCs have been drawn much attention as biosensors recently. MFC systems can also be utilized to treat wastewater5 and monitor the quality of wastewater with function of acting as an environmental sensor.6 Generally, a sensor is an electronic device used to detect certain properties in the surrounding. The measured quantities then are converted into signal voltage to a processor. Asghary et al7 used a dualchambered MFC as a power supply to construct an analytical device in monitoring the DNA immobilization and hybridization events", + " Therefore, the objective of the present study is to investigate the effect of electrode spacing on the performance of dual chamber MFCs with honeycomb flow straightener in the anode chamber. In this study, the main focus is to investigate the performance of MFC for possible use as a sensor in waste water treatment. The MFC of dual chamber with recirculation was fabricated and employed in the present study. The material of the system was made of acrylic. In the system, there are three parts: anode chamber, cathode chamber, and buffer chamber as shown in Figure 1. The dimension of the anode chamber is 7.0 cm \u00d7 5.0 cm \u00d7 8.0 cm, while that of the cathode chamber is 5.0 cm \u00d7 5.0 cm \u00d7 6.0 cm separated by a polymer exchange membrane (PEM; Nafion 117, Dupont, with surface area of 25.0 cm2, 5.0 cm \u00d7 5.0 cm). The dimension of the buffer chamber is 5.0 cm \u00d7 5.0 cm \u00d7 10.0 cm. The material for electrodes is graphite sheet with the dimension of 5.0 cm \u00d7 5.0 cm. In Figure 1, it is clear seen that a flow device such as honeycomb was installed in the anode chamber in order to produce a uniform flow in the anode chamber as the flow is recirculating in the circulating tube. The size of the honeycomb is 5.0 cm \u00d7 5.0 cm \u00d7 8.0 cm and assembled with plastic straws with inner diameter of 0.7 cm. Because the length of honeycomb has insignificant influence on the performance of MFCs, the dimension of honeycomb was chosen to be the same as the chamber. The anolyte was prepared with 50% waste water (from Luodon wastewater recycle center [24 450N, 121 470E], Ilan, Taiwan) and 50% sodium acetate medium content" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0003328_amm.761.329-Figure4-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0003328_amm.761.329-Figure4-1.png", + "caption": "Fig. 4 Equivalent stress of FDM nozzle using frequency 30 kHz to 40 kHz.", + "texts": [ + " The factor of safety (FoS) for this model was obtained by using this result. The calculations showed that the factor of safety of the model was 20.56, referring to the fact that the nozzle can withstand a frequency range of between 20 kHz to 30 kHz that will be transmitted from the ultrasound transducer. Even though the FoS is high, we observed that bending still took place on the nozzle due to the fact that the part has a thin thickness profile. Our observations also showed that no loose screws were found on the part after being subjected to vibrations. Fig. 4 shows the equivalent stress of FDM nozzle for the range of frequency between 30 kHz to 40 kHz. The highest value is 12.753 MPa and the lowest value is 6.464 MPa. The FoS for this model was found to be 18.8975, referring to the fact that the nozzle can withstand frequencies between 30 kHz to 40 kHz from the ultrasonic transducer. The FoS is higher than 20 kHz to 30 kHz because the ultimate tensile strength is also high. The ultimate tensile strength becomes higher because of the frequency applied is higher" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0001217_032005-Figure1-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0001217_032005-Figure1-1.png", + "caption": "Figure 1. Constructive (a) and design (b) schemes of the proposed connection.", + "texts": [ + " Therefore, the actual task is to study in detail the strength indicators of connections of operating parts and load-bearing structures of agricultural machines as the main criterion of durability [4, 5]. MIP: Engineering-2020 IOP Conf. Series: Materials Science and Engineering 862 (2020) 032005 IOP Publishing doi:10.1088/1757-899X/862/3/032005 Improving the strength and durability of bearing connections of operating parts of agricultural machines that are operated under complex loading conditions is possible by using the proposed connection design [6-8]. The design scheme of the proposed connection is shown in figure 1 (a) and consists of a bushing 4, parts 1 and 2, a bolt 3, and a nut 5. In parts 1 and 2, a cylindrical hole is provided for the height of the connected parts, and the bolt is installed with a gap on the side of the head. On the side of the nut, a bushing is installed to fit the bolt 3 without a gap. During the connection assembly the bolt 3 passes freely through the hole of the part 1, since its diameter exceeds the bolt's landing diameter. In the hole of the bushing 4, the specified bolt enters with some tension", + " As a result, there is an increase in the margin of wear-resistant strength and durability of the connection. In addition, in this case, connection boring of the connected parts is practically excluded, which, as known, is a rather labor-intensive technological operation. The mathematical model of the proposed connection can be represented as a differential equation of the elastic line of the rod [9-11]: . 2 2 2 2 yx q dz yd (z)EI dz d = (1) The calculated connection scheme implementing equation (1) is presented as shown in figure 1 (b). To establish the forces and moments shown in this figure, we solved the equation (1) using the method of initial parameters [7, 8]: ; ! z EI )(zR ! z EI )(zM ! z EI )(zR ! z EI )(zM ! z EI zR ! z EI zM )z( dz dy z EIy z EIy(z) b\u0441bcdd yaya 3232 32 0 3 1 2 1 32 32 00 \u2212 + \u2212 + \u2212 + \u2212 + ++++= (2) . !EI )(zR !EI )(zM !EI )(zR !EI )(zM !EI zR !EI zM )( dz dy EI(z) dz dy z bc z bc z d z d z a z a z 212 121 0 2 11 2 22 \u2212 + \u2212 + \u2212 + + \u2212 +++= (3) when =z , we obtain )( dRy +\u2212= . In this case according to (2) and (3) we obtain: ; 62 )( 32 aa d RM R z EI +=+\u2212 (4) " + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0003199_iicpe.2014.7115807-Figure5-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0003199_iicpe.2014.7115807-Figure5-1.png", + "caption": "Fig. 5. Torque change for input voltage change.", + "texts": [ + " From (13), it can be inferred that for a small change in the input volt\u2013time Vs \u0394t, the stator flux varies by a factor (say) \u0394\u03c8s. This change in the flux due to the input volt-time change is reflected as a change in the torque angle and the corresponding torque magnitude is shown in (17). \u0394Te = 3 2 * P 2 Lm LrLa \u03a8r \u03a8s+ \u0394\u03a8s sin\u0394\u03b4 (17) Equation 17 represents the change in torque as a function of changes in stator flux and torque angle. A power converter changes the stator flux by changing the input volt\u2013time, as shown in Fig. 5. The required output voltage vector is determined by the space vector position, the reference torque and the reference flux. The reference torque and flux values are estimated. With the input voltage, the line currents and the present switch position, the model estimates the actual flux with which the torque and speed of the induction motor are calculated. The error between the reference speed and the actual speed of the machine is given to a PI controller that generates the torque reference. The machine speed is also used to generate the reference flux, as shown in Fig" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0002119_icaccm50413.2020.9212874-Figure6-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0002119_icaccm50413.2020.9212874-Figure6-1.png", + "caption": "Fig. 6. Total deformation of CFRP propeller(case1).", + "texts": [ + " 1 2 3 4F mg F F F F (1) 2 i t iF K (2) Where m is the total mass (Quadcopter and payload) and g is acceleration due to gravity. Kt is the coefficients, depends upon propellers geometry and air density. If the value of m is 20 kg then the value of the total thrust force is calculated as follows. Total Thrust force (F)= mg= 20*9.81=196.2 N Thrust force required per propeller= F/4 =196.2/4 =49.05 N The calculated thrust force is approximately 50 N for a single propeller. Thus, the propeller is subjected to 50 N of thrust force to find out the deformation and stress due to the bending effect as shown in Figure 5. Figure 6, 7 and 8, 9 show the total deformation and stress for CFRP and GFRP materials, respectively. Table II presents the comparison of results for CFRP and GFRP. 60 Authorized licensed use limited to: Carleton University. Downloaded on May 30,2021 at 15:29:13 UTC from IEEE Xplore. Restrictions apply. A Quadcopter\u2019s propeller is subjected to horizontal bending due to collision of the propeller with obstacles or a wall. To calculate the bending deformation under the applied rotational moment about the rotation axis, the propeller has been fixed by both the tip as shown in Figure 10" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0000134_s12555-018-0710-9-Figure4-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0000134_s12555-018-0710-9-Figure4-1.png", + "caption": "Fig. 4. Conceptual QC structure.", + "texts": [ + " Therefore, technically it follows that R: (us1,us2,us3,us4) def = (LIVA \u25e6LI)(ur0,ur1,ur2,ur3) , (1) where ur0 def = R0(r0) and ur i def = Ri(ei), \u2200i \u2208 1,3, where ei def = ri \u2212 yi denotes the ith output feedback controller related control error. 3. QUADROCOPTER MODELS The conceptual internal structure of the QC has been illustrated in Fig. 3. The QC, in principal, is composed of the main body (\u2018QC body\u2019) and four motor\u2013blade pairs (\u2018S\u2019) connected to a power supply system (\u2018Z\u2019) acting altogether as the AC system. 3.1. Quadrocopter body model for simulation purposes Applying the Newton-Euler equations of motion to describe the QC mechanics in three dimensional space (Fig. 4) yields a simulation model Mp. The structure of the Mp is given by: \u03b1\u0308x = \u03b1\u0307y\u03b1\u0307z Ixy \u2212 Iz Ixy + l Ixy (\u2212Fc2 +Fc4) , (2a) \u03b1\u0308y = \u03b1\u0307x\u03b1\u0307z Iz \u2212 Ixy Ixy + l Ixy (Fc1 \u2212Fc3) , (2b) \u03b1\u0308z = l Iz 4 \u2211 i=1 (\u22121)iMrot i. (2c) where x,y,z denote the QC\u2019s coordinate frame in three dimensional space; \u03b1x,\u03b1y,\u03b1z \u2208 R are the Tait-Bryan (yaw, pitch and roll) angles; Iiidx ,\u2200iidx \u2208 {\u2018x\u2019, \u2018y\u2019, \u2018z\u2019}, denote the moment of inertia with respect to the basis of QC\u2018s coordinate frame. In order for the utility model to be welldefined it is crucial to identify the parameters of the derived structure, namely: pd def = [Ixy, Iz, l] T as well as the shape and parameter of the functions Fc i def = Fc i (ui,uz) and Mrot i def = Mrot i(ui,uz), \u2200i, e" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0003532_icl.2015.7318107-Figure1-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0003532_icl.2015.7318107-Figure1-1.png", + "caption": "Fig. 1. Class project model \u2013 gear-pump system; component and assembly modeling and assembly drawing", + "texts": [ + " While the students practice the CAD software to create virtual 3D models, they study about how to generate professional engineering drawing documents in which the detailed information about the model should be indicated. In this course, special class sessions are provided for the students to study and learn engineering graphics fundamentals such as orthographic projection views and dimensioning and tolerancing (GD&T). The class project model; gear pump system with seven independent components has been selected for the students to create all components/parts every time they practice CAD software and assemble them to create a complete assembly model and drawing as shown in Figure 1. 3) CAD team project \u2013 poster exhibition and competition: After practicing CAD software to create assigned virtual individual components/parts, assembly model and corresponding document drawings shown in the course textbook, the students are asked to apply their skills to create 3D solid CAD models of selected real objects such as a stapler, hand-held calculator, an electric hand dryer, among others. For this assigned work the students should measure the required dimensions of the assigned objects, create their own hand sketches and 3D solid models and compare the final model with the real object to evaluate their CAD models and proficiency" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0003051_icelmach.2018.8507049-Figure1-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0003051_icelmach.2018.8507049-Figure1-1.png", + "caption": "Fig. 1. Stator (a) and rotor (b) structures of the baseline machine.", + "texts": [ + " This duplex three-phase system features the possibility of operating the machine either in six-phase or in three-phase configuration. In fact, connecting the two winding sets in anti-series the winding structure does not change: the air gap Magneto-Motive Force (MMF) harmonic spectra produced in three-phase and in sixphase configurations are the same. The detailed design features and the experimental validation of the baseline machine can be found in [5], [18]. However, for the sake of completeness, Fig. 1 and Table 1 reports the stator and rotor structures and their main geometrical dimensions, respectively. Fig. 2 depicts the winding disposition in the stator slots for the baseline machine. The winding function analysis consists on the cumulative sum along the stator periphery of the turn numbers of a phase winding. For conventional distributed and tooth-wound winding layouts, the winding function has a zero average value as each coil presents a positive and a negative side, and a maximum amplitude corresponding to half conductors in series per phase [14]" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0003612_amm.664.158-Figure1-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0003612_amm.664.158-Figure1-1.png", + "caption": "Fig. 1 Typical representation of single-propellant system", + "texts": [ + " [6] obtained an optimal position for installing accumulator, but the analysis on parameters is not thorough and it neglects the influence of accumulator inertance. This paper starts with the establishment of dynamical model for propulsion system via the method of transfer matrix, and then a thorough analysis on the influence of parameters on natural frequency of propulsion system is carried out. At last, combining the analysis on the stability of the system, effective methods of suppressing the POGO are given. A typical single-propellant system is 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: 130.237.29.138, Kungliga Tekniska Hogskolan, Stockholm, Sweden-09/07/15,20:20:54) The models of two vital elements in propulsion system are expressed as follows: A. Suction line 1 2 2 1 sL cos(\u03b1l) - sin(\u03b1l) P P\u03b1l = Q \u03b1l Q sin(\u03b1l) cos(\u03b1l) sL (1) where P1 ,P2, Q1 and Q2 are pressure and volume flow at the top and end of the element, L is inertance of fluid in line, l is the length of the line, \u03b1 is wave number and s is Laplace variable" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0001985_intercon50315.2020.9220221-Figure1-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0001985_intercon50315.2020.9220221-Figure1-1.png", + "caption": "Fig. 1. Rotor free body diagram.", + "texts": [ + " Downloaded on November 02,2020 at 13:19:40 UTC from IEEE Xplore. Restrictions apply. of the rotor radial position from the characteristics of a real experimental prototype. This prototype works vertically and is supported on the lower end of the shaft by a pivoted bearing that prevents radial and axial movements, but allows angular movements. This type of system has two degrees of freedom for position control [12] and the rotor displacement occurs in the directions of x and y axes. Starting from the application of the second motion law to Figure 1, [11] developed a G(s) function that relates the control signal U(s) to the output position Y (s) of the rotor. The dynamics of the x-axis is equal to the y-axis and both are described by: G(s) = 3.68\u00d7 106 (s+ 91.51)(s\u2212 91.51) (1) Equation (1) show that the radial position model has a unstable pole, thus a controller is fundamental to the system operation. The first version of the ADRC controller was presented in chinese by its creator, Jingqing Han, in work entitled \u201cAuto disturbances rejection controller and its applications\u201d in 1998 [13]" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0003255_aim.2015.7222699-Figure1-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0003255_aim.2015.7222699-Figure1-1.png", + "caption": "Fig. 1. Underactuated biped robot model", + "texts": [ + " We use it as a virtual time parameter of the linearized system to ensure the control of the next landing position of the swing leg, and show that the generated trajectory becomes symmetric in the configuration space [8][9][10]. Through numerical analysis, we discuss the relationship between the gait symmetry and mechanical energy restoration. Furthermore, we numerically show that generation of a walking gait following stepping stones can be achieved on the assumption that the robot can visually perceive the road surface condition. Fig. 1 shows the model of an underactuated biped robot. This consists of two identical rigid leg frames and three point masses, and can exert the hip-joint torque, uH . The main difference from the commonly-used (point-footed) compasslike biped robot [11] is the CoM location of each leg frame. The CoM of the leg frames is positioned at a distance of w [m] from the central axis. In other words, the CoM is not on the hip-foot line. We set the anterior direction to positive. The inertia moments about the masses are ignored. Fig. 2 shows the equivalent models for Fig. 1. The robot consists of two identical leg frames that are not compliant or cannot bend during motion. Such bent leg frames with positive w have the effect to shift the robot CoM forward. Then the robot can easily overcome the energy (potential) barrier at mid-stance [12]. 978-1-4673-9107-8/15/$31.00 \u00a92015 IEEE 1184 Define \u03b81 and \u03b82 as the angular positions of the stance and swing legs with respect to vertical. Let \u03b8 = [ \u03b81 \u03b82 ]T be the generalized coordinate vector. The robot equation motion then becomes M(\u03b8)\u03b8\u0308 + h ( \u03b8, \u03b8\u0307 ) = SuH = [ 1 \u22121 ] uH " + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0001138_0954406220925836-Figure11-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0001138_0954406220925836-Figure11-1.png", + "caption": "Figure 11. 3T3R-AMP2AMP2AMP2 GPM.", + "texts": [ + " With no constraint exerting on the end-effector, only unconstrainted limbs are feasible to design mechanisms. When employing two R3RMRMR03R 0 3R D-H3limbs to connect the two sides of the end-effector, three groups of parallel revolute joints are included in the limb. To make sure the hybrid limbs can perform the necessary three-dimensional translation, the R3 revolute joins in the two limbs are not parallel. Then the 3T3R-AMP2AMP2AMP2 GPM with rotations about three directions is obtained, as shown in Figure 11. GPMs with AMP3 After arranging a single revolute joint in the middle of the moving platform, the AMP3 is constructed. The axis of the revolute joint is perpendicular to the platform, the end moving platform degenerates into the intermediate moving platform. The full-cycle rotation of the end-effector realized through the revolute joint is directly actuated by the middle limb, and the rotation is decoupled from other motions. By adding serial kinematic chains to the above-proposed mechanisms, the GPMs with high rotational performance by combining different articulated moving platforms can be obtained" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0000790_042080-Figure1-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0000790_042080-Figure1-1.png", + "caption": "Figure 1. Joint scheme with a movable longitudinal tie: 1-bumps; 2 - circular tape; 3- movable longitudinal binder.", + "texts": [ + " Soft hinges allow only bending of parts of the body and reduce the effort required for bending. They allow the movement of parts of the body in the required range of angles, and it is achieved by constructive features. The main idea is in the longitudinal forces transmission to the force elements located in the neutral surface or plane of the hinge. International science and technology conference \"FarEastCon-2019\" IOP Conf. Series: Materials Science and Engineering 753 (2020) 042080 IOP Publishing doi:10.1088/1757-899X/753/4/042080 Figure 1 shows a soft corrugated hinge, characterized in that the longitudinal tie of the hinge is made movable, and connected to transverse circular bands through elastic elements made, for example, of rubber. The currently used corrugated hinges in the glove do not have the proper flexibility. Ideally, the soft hinge of the spacesuit should not resist the movements of the astronaut's limbs. A soft corrugated hinge is proposed as a good practical solution. It is characterized by the longitudinal hinge of the hinge which is movable (Fig" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0002749_gt2009-59696-Figure3-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0002749_gt2009-59696-Figure3-1.png", + "caption": "Fig. 3 Exciter head shown attached to stator via stingers and load [10]", + "texts": [ + " They allow the bearing housing to move freely in the radial direction yet prevent pitch and yaw rotations and axial movement. Power is delivered to the rotor via a 65 KW (90HP) air turbine through a flexible coupling. Speed can be varied from 0 to 16k rpm. ISO VG 32 oil is supplied to the test bearing in varying quantities up to 90 liter/min. A range of static external loads are applied from a pneumatic loader to the test bearing for speeds ranging from 4k to 13k rpm. The static loader can apply a steady load up to 22 kN in the y direction. The hydraulic shaker connections shown in Fig. 3 provide dynamic loads to the bearing housing that are parallel and perpendicular to the static load. This, \u201cshake-the-bearing\u201d test-rig design was first proposed by Glienicke [14] in 1966. A pseudo-random waveform that includes all frequencies from 20-320 Hz in 20Hz intervals is the input signal to the hydraulic shakers. The amplitude and phase of the wave-form components are determined to minimize the peak force required from the shaker while providing adequate response amplitudes within the bearing. The shakers apply forces to the bearing housing via stingers. A stinger is rigidly attached to the bearing housing at one end and attached to the shaker via a load cell to measure the dynamic loads imposed by the shaker. As shown in Fig. 3, two piezoelectric accelerometers are attached to the bearing housing. The y direction identifies the applied static load direction, and eddy-current proximity probes are placed in the x and y axes of the bearing housing to measure rotor-bearing relative Childs GTP-09 3 nloaded From: http://proceedings.asmedigitalcollection.asme.org/ on 01/30/2016 Term displacement components. Two probes are placed in a plane at the drive end, and two are placed in a parallel plane at the non-drive end. Because measurements are taken in two parallel planes, both the pitch and yaw of the stator housing (relative to the rotor axis) can be measured and minimized prior to testing" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0000126_j.ifacol.2019.10.069-Figure5-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0000126_j.ifacol.2019.10.069-Figure5-1.png", + "caption": "Fig. 5. Collision of the nozzle to machine table. a) CAD model of the part and collision making layer. b) Collision on building process", + "texts": [ + " In order to make complex parts with a high overhang angle using DMD, the nozzle needs to be mounted onto a 5 axis system. As Fig. 3 shows, the nozzle can rotate to keep being tangent (or approximately tangent) to surface to build from previous layers. This eliminates the need of making support structure for walls having high overhang angle. However, 5 axis machines increase the complexity of the process and the possibility of a collision between the machine head, the substrate, or previously deposited layers. This is shown in Fig. 4 (collision due to previously deposited layers) and Fig. 5 (machine to deposition head collision). DMD nozzles are very expensive and very sensitive to any collision. Consequently, if any crash happens it would be very costly. The collision related issues cannot be considered only on a feature basis. As Fig. 4 shows if the part needs to be made in shown orientation, the nozzle collides to previously deposited layers when it starts to make the top inclined surface. In the 2.5 axis additive manufacturing machines, the nozzle just moves in the x-y plane. As all the previously deposited layers and the machine table are below the deposition head, the probability of crash is very low. On the other hand, in a 5 axis machine configuration, there are 2 rotary movements in addition to the 3 translational ones. So, the nozzle can tilt as much as the axes\u2019 limits allow, and it may hit the table. Fig. 5(a) shows a round part that at the shown layer, the tangent line is too horizontal so if the nozzle needs to be tangent to the surface, it crashes to machine table (Fig. 5(b)). Two solutions can be proposed to solve this problem, tilting the nozzle and sectioning of the surfaces. Addressing overhang angles is a critical matter for AM processes. This is defined as the angle between the slicing direction and the part surface (Ven et al. 2018). The slicing direction is the normal vector of the slicing planes (Eiliat and Urbanic, 2019). Fig. 6 illustrates the overhang angle for a thin wall. 2019 IFAC IMS August 12-14, 2019. Oshawa, Canada 231 232 Hamed Kalami et al. / IFAC PapersOnLine 52-10 (2019) 230\u2013235 For each machine parameter setting-material there is a maximum overhang angle", + " If the nozzle axis leans in N-F surface, the angle between N and the nozzle axis is called lead angle (Fig. 8 (a)). When nozzle turns in C-N surface, as Fig. 8 (b) shows it is called tilt angle. The partitioning algorithm is performed on a thin wall dome to determine feasible build solutions using by 5 axis DMD without introducing any collisions. The maximum overhang angle is assumed to be 10 degrees for this case study (Fig. 9). 2019 IFAC IMS August 12-14, 2019. Oshawa, Canada Hamed Kalami et al. / IFAC PapersOnLine 52-10 (2019) 230\u2013235 233 a) b) Fig. 5. Collision of the nozzle to machine table. a) CAD model of the part and collision making layer. b) Collision on building process For each machine parameter setting-material there is a maximum overhang angle. Exceeding this value causes the material collapse. In this situation either a support structure is needed, or the nozzle needs to rotate to be more tangential to surface to decrease the overhang angle. The former solution is the only solution for 2 \u00bd axis machines (i.e., all powder bed systems)" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0003208_9781118751992.ch4-Figure4.12-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0003208_9781118751992.ch4-Figure4.12-1.png", + "caption": "Figure 4.12 Illustration for Problem 4.3.", + "texts": [], + "surrounding_texts": [ + "1. P. G. de Gennes and J. Prost, The physics of liquid crystals (Oxford University Press, New York, 1993). 2. L. M. Blinov and V. G. Chigrinov, Electrooptical effects in liquid crystal materials (Springer-Verlag, New York, 1994). 3. A. J. Leadbetter, Structure classification of liquid crystals, in Thermotropic liquid crystals, ed. G. W. Gray (John Wiley & Son, Chichester, 1987). 4. R. M. Hornreich, Landau theory of the isotropic\u2013nematic critical point, Phys. Lett., 109A, 232 (1985). 5. I. Lelidis and G. Durand, Electric-field-induced isotropic\u2013nematic phase transition, Phys. Rev. E, 48, 3822 (1993). 6. R. B. Meyer, Phys. Rev. Lett., 22, 918 (1969). 7. A. I. Derzhanski, A. G. Petrov, and M. D. Mitov, J. Phys., (Paris), 39, 273 (1978). 8. D. Schmidt, M. Schadt and W. Z. Helfrich, Naturforsch, A27, 277 (1972). 9. E. L.Wood, G. P. Bryan-Brown, P. Brett, A. Graham, J. C. Jones and J. R. Hughes, Zenithal bistable device (ZBD) suitable for portable applications, SID Intl. Symp. Digest Tech. Papers, 31, 124 (2000). 10. J. S. Patel and R. Meyer, Flexoelectric electro-optics of a cholesteric liquid crystals, Phys. Rev. Lett., 58, 1538 (1987). 11. G. Chilaya, Cholesteric liquid crystals: optics, electro-optics, and photo-optics, Chirality in liquid crystals, ed. H.-S. Kitzerow and C. Bahar (Springer, New York, 2001). 12. L. Komitov, S. T. Lagerwall, B. Stenler, and A. Strigazzi, Sign reversal of the linear electro-optical effect in the chiral nematic phase, J. Appl. Phys., 76, 3762 (1994). 13. J. S. Patel and S.-D. Lee, Fast linear electro-optic effect based on cholesteric liquid crystals, J. Appl. Phys., 66, 1879 (1987). 14. R. B. Meyer, L. Liebert, L. Strezelecki, P. Keller, J. Phys. (Paris) Lett., 36, L69 (1975). 15. J. W. Goodby, R. Blinc, N. A. Clark, S. T. Lagerwall, M. A. Osipov, S. A. Pikin, T. Sakurai, K. Yoshino, B. \u017dek\u0161, Ferroelectric liquid crystals: Principle, properties and applications, Ferroelectricity and related phenomena, Vol. 7 (Gordon and Breach Publishers, Amsterdam, 1991) 16. N. A. Clark, S. T. Lagerwall, Submicrosecond bistable electro-optic switching in liquid crystals, Appl. Phys. Lett., 36, 899 (1980). 17. S. Garoff and R. B. Meyer, Electroclinic effect at the A-C phase change of a chiral smectic liquid crystal, Phys. Rev. Lett., 38, 848 (1977); Phys. Rev. A., 19, 388 (1979)." + ] + }, + { + "image_filename": "designv11_71_0000033_s00202-019-00846-1-Figure5-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0000033_s00202-019-00846-1-Figure5-1.png", + "caption": "Fig. 5 2D FEM model for electromagnetic performance assessments", + "texts": [ + " A new two-dimensional magnetostatic FEM is proposed to estimate the steady-state performance of the PMSG in [33]. A magnetostatic model of a PMSG is important because it can then be used as a starting point in a transient analysis or evaluation of a new generator design [33]. Fig. 3 3D model of PMSG Complete View StatorCoilN40SH-Magnets Rotor Fig. 4 Manufactured parts of PMSG 1 3 In the case that the number of poles of the machine is high (as in the application of which 2D electromagnetic details are given in Fig.\u00a05), 2D analytical solutions and FEM provide highly accurate solutions for the magnetic field distribution and machine performance. In this paper, 2D FEM transient solver was chosen to analyze the magnetic fields, energy, force/torque, power loss, core loss, speed, and flux of a model at various time steps. There are some limitations and assumptions for 2D FEM transient solver as follows [34]: \u2022 Magnetic field solution is proportional to changing air gap/pole arc ratio and radius of the selected cylindrical cutting plane; \u2022 Mutual and leakage fluxes in the end-winding cannot be taken into account; \u2022 It is not used for the analysis of the materials with different cross sections; \u2022 Rotational motion can be cylindrical or non-cylindrical" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0002861_978-3-030-55889-5-Figure1.19-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0002861_978-3-030-55889-5-Figure1.19-1.png", + "caption": "Fig. 1.19 Temperature effects", + "texts": [ + "131) We can further assume wT to be calculated on the statically determinate structure, wT (x) = \u03b1T \u0394T h x2 + a x + b , (a, b are constants) (1.132) since any necessary corrections can be treated separately. The principle of virtual forces for a deflection wT , we place a point load P\u2217 = 1 at the source point x , reads G (\u03b4w\u2217, wT ) = 1 \u00b7 wT (x) \u2212 \u222b l 0 E I \u03b4w\u2217\u2032\u2032 w\u2032\u2032 T dx = 0 , (1.133) or with w\u2032\u2032 T = \u03b1T\u0394T/h wT (x) = \u222b l 0 \u03b4M\u2217 \u03b1T \u0394T h dx . (1.134) where \u03b1T \u223c 10\u22125 (steel, concrete) is the temperature coefficient of the material, \u0394T is the temperature difference between the top and bottom fibers of the beam and h is the beam height, see Fig. 1.19. In the same way, the temperature displacement in longitudinal direction is uT (x) = \u222b l 0 \u03b4N \u2217 \u03b1T T dx , (1.135) where T is the increase in temperature. 1.12 The Complete Equation 31 We can now assemble all these single parts in one formula, Mohr\u2019s equation, 1 \u00b7 \u03b4 = \u222b M\u2217 M E I dx + \u222b N \u2217 N E A dx + \u222b M\u2217 \u03b1T \u0394T h dx + \u222b N \u2217 \u03b1T T dx + \u2211 i F\u2217 i Fi ki\ufe38 \ufe37\ufe37 \ufe38 force springs + \u2211 j M\u2217 j M j k\u03d5 j\ufe38 \ufe37\ufe37 \ufe38 moment springs \u2212 \u2211 k F\u2217 k w\u0394 k \ufe38 \ufe37\ufe37 \ufe38 sup. settlements \u2212 \u2211 l M\u2217 l tan \u03d5\u0394 l \ufe38 \ufe37\ufe37 \ufe38 sup. rotations . (1.136) The \u03b4 on the left side stands for a displacement or a rotation" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0003270_icpe.2015.7167862-Figure3-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0003270_icpe.2015.7167862-Figure3-1.png", + "caption": "Fig. 3. Concept diagram of inequality demagnetization", + "texts": [ + " To be more specific, the portion of demagnetization is determined according to direction of rotate. This is because, when a motor is operating, the magneto-motive force from the supplied power affects the q-axis, not the d-axis [6]. In this study, we performed equivalent modeling in the light of demagnetization conditions. Fig. 2 shows magnetization shape diagram according to demagnetization ratio of equality demagnetization pattern. 2015 KIPE B. Inequality Demagnetization Pattern of PM Inequality demagnetization pattern is equally demagnetized at all magnets, as shown in Fig. 3. However, it is different demagnetization ratio between Npole and S-pole. When demagnetization in permanent magnet of motors is occurred, each permanent magnet has a little difference on demagnetization ratio. Therefore, we determined irreversible demagnetization pattern having different demagnetization ratio between N-pole and S-pole. Fig. 3 shows a concept diagram of inequality demagnetization. In addition, Fig. 4 shows magnetization shape diagram according to demagnetization ratio. Ratio of irreversible demagnetization patterns is N-pole 50%, S-pole 60% and N-pole 50%, S-pole 70%. Fig. 5 shows a weighted demagnetization pattern. One of N-pole and S-pole is demagnetized. Demagnetization pattern of this model cannot be occurred in real conditions. However, it has a special meaning for BEMF harmonic characteristics analysis. Because the BEMF waveform is directly influenced by flux linkage variance of each phase, we selected weighted demagnetization model" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0001211_012001-Figure4-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0001211_012001-Figure4-1.png", + "caption": "Figure 4. Road train trajectories for two cases of power distribution between the units: 1 \u2013 100/0; 2 \u2013 25/75.", + "texts": [ + " The friction model described in [16] is used to calculate this resistance moment. The mathematical model of the active road train with EMT is implemented in Matlab/Simulink dynamic mathematical simulation environment. The effect of trailer wheel activation on road train dynamics was assessed using the results of simulation of going into corner and subsequent turning with a fixed position IASF-2019 IOP Conf. Series: Materials Science and Engineering 819 (2020) 012001 IOP Publishing doi:10.1088/1757-899X/819/1/012001 of the steering wheel and constant speed. Figure 4 shows simulation results for two power distribution cases (100/0 and 25/75). As can be seen from Figure 4, road train trajectories differ significantly. The turning radius of the tractor with passive semitrailer is 19.18 m, but it is 17.05 m with active semitrailer. The main reason for differences is the variation of wheel slip angles when torque is applied to semitrailer wheels. Figure 5 shows the comparison of wheel slip angles of active and passive semitrailers (for wheels 1, 3 and 6). Figure 6 shows the variation of tractor trajectory curvature for four cases of power distribution between the tractor and semitrailer" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0001031_icimia48430.2020.9074843-Figure11-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0001031_icimia48430.2020.9074843-Figure11-1.png", + "caption": "Fig. 11 Rotary Linear Low Complexity Motor [14 -15]", + "texts": [ + " Future research is required in this motor for making the size of the motor further smaller. The geometry of these machines is a modular structure having an arbitrary number of modules with poly-phase winding. The stator is a grossly hollow cylindrical shape and composed of two similar modules. Each module has one ferromagnetic core and a winding is shown in Fig. 10. The isotropic structure of the ferromagnetic core is placed as facing the air gap [14 -15]. The rotor has a cylindrical hollow shaped and is placed inside the stator as shown in Fig. 11. It consists of a ferromagnetic core, and an even number of magnets are arranged in alternate sequence. These magnets have characteristics of radial magnetization, hence half of the magnets have outward magnetization direction and inward magnetization direction for others. When the stator module \u2018A\u2019 is energized, the magnetic field is generated. The generated magnetic field induces the eddy currents in the magnet. The eddy currents, in turn, causes the eddy flux which is outwards the main flux. The outward direction of the eddy flux combines with the outward magnetization direction produces the rotational torque" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0002216_icem49940.2020.9270977-Figure6-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0002216_icem49940.2020.9270977-Figure6-1.png", + "caption": "Fig. 6: Example of an on-load field solution obtained using RFOA. The rotor current are computed outside the FEA and imposed along the q-axis of the adopted reference frame. Notice that the rotor flux linkage space vector lies along the d\u03bb-axis of the oriented reference frame.", + "texts": [ + " A satisfying description of this formulation is contained in [5]. However, this approach becomes more complicated when the machine is skewed, where it is necessary to consider several 2D slices and add further constraints to the FE problem equations [6]. The analysis technique adopted in this work is well described in [3], [7]. The machine is simulated using the magneto-static formulation of the FEA problem, where both the stator and rotor currents are field sources. An example of an on-load field solution obtained using the RFOA technique is reported in Figure 6. Basically, in a certain time instant, the stator current is imposed in the chosen reference frame, named d\u03bbq\u03bb in Figure 6, with the components isd and isq. The rotor current is computed using the procedures described in [3], [7] and then imposed along the q-axis of the reference frame. The procedure computes the rotor current so that the q-axis rotor flux linkage results equal to zero; in such a way, the adopted reference frame for simulations is exactly the d\u03bbq\u03bb, oriented with the rotor flux space vector, and all the vector quantities are referred to this reference frame. From the field solution in Figure 6, it is possible to derive the stator and rotor flux linkage space vector in d\u03bbq\u03bb, \u03bb\u03bbs and \u03bb\u03bbr ; the torque and the rotor angular frequency are derived using the model equation [8]: Tdq = 3 2 p (\u03bbsdisq \u2212 \u03bbsqisd) ; \u03c9sl = p PJr Tdq (14) This analysis technique allows to easily consider the presence of closed rotor slot and the rotor skewing. For this purpose, the procedure outlined in [9] has been followed. Different 2D FE slices are considered and each problem is solved using the static formulation, saving a long computation time" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0002324_iceca49313.2020.9297379-Figure5-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0002324_iceca49313.2020.9297379-Figure5-1.png", + "caption": "Figure 5: Configuration of motors, their direction of rotation and orientations of axis for hovering mode.", + "texts": [ + " Thus the dynamic model of the aircraft is divided into two modes: Hovering mode and Gliding mode. The drone uses four motors for propulsion and control. It can be maneuvered by changing the speed of these motors. Brushless DC motors are used in this drone whose angular velocities can be varied by varying frequency of the current supplied with the help of ESC (Electronic Speed Controller). Figure 4 shows the direction of rotation and the configuration of four motors when v iewed from the top. The following figure 5 shows the configuration of motors, their direction of rotation and the orientations of axis for hovering mode. Let, x, y and z be three perpendicular axis attached to the fixed frame as shown in figure 5. Roll axis, Pitch axis and Yaw axis be three perpendicular axis attached to the frame of the drone as shown in figure 5. T1, T2, T3 and T4 be the thrusts developed by the motors 1, 2, 3 and 4 respectively. The length between motors 1 and 2 (and so between motors 3 and 4) be 2*L1 and the length between motors 1 and 3 (and so between motors 2 and 4) be 2*L2. ax, ay, and az be accelerat ions of the drone on x, y and z axis respectively. \u03d5, \u03b8 and \u03c8 be Pitch, Roll and Yaw angles respectively. \u03c91, \u03c92, \u03c93 and \u03c94 be angular velocit ies of motors 1, 2, 3 and 4 respectively. M1, M2, M3 and M4 be the moments generated by motors 1, 2, 3 and 4 respectively" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0000126_j.ifacol.2019.10.069-Figure13-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0000126_j.ifacol.2019.10.069-Figure13-1.png", + "caption": "Fig. 13. Building a dome from the end of a round bar by rotary production strategy. a) 2.5 axis layers at the end of the bar. b) Building the dome at the end of bar. c) The way this prevents collision on the machine.", + "texts": [ + " In this case, if the nozzle maintains the tangency condition with respect to the surface, it will collide with the table when building the first layers, as shown in Fig. 12. Fig. 10 (b) orientation, and radial partitioning for the Fig. 10 (c) orientation. In order to avoid any collision at the first layers, a round bar can be used as the substrate and several layers be deposited at one end of the bar. The number and consequently the thickness of deposited layers needs to be more than the width of beads that forms the dome. This is shown in Fig. 13 (a) which 6 layers are deposited at the end of the bar by using just 2.5 axis movement. Then as Fig. 13 (b) shows, the dome starts being built from the side of the last bead on the bar. Fig. 13 (c) shows how this solution No Start Insert geometry Create surface model 3 slicing directions Partition the part surface Max overhang angle Building orientation for each section Merge toolpaths Make toolpaths for each section Is there any collision? Choose least time & support material for n orientations End Is there any collision? Yes No Yes Make toolpath 2019 IFAC IMS August 12-14, 2019. Oshawa, Canada 233 234 Hamed Kalami et al. / IFAC PapersOnLine 52-10 (2019) 230\u2013235 prevents the collision of the machine head to the table. After the dome is built, it just needs to be cut from the round bar. This is an interesting solution, but it may be difficult to expand upon. The geometry presented in Fig. 4 would be challenging. Based on the orientation showed in Fig. 10 (c), it is needed to be partitioned properly to be able to be built. Here, a 3 + 2 machine configuration is assumed. In order to prevent collisions, the part is split radially, as shown in Fig. 13. The building sequence is shown by numbers in the illustration. The nozzle should be tangent to the surface during production. Top edge of each sub-section is the substrate for starting the next section. The solution to prevent collisions is to build the part at the end of a plate. The plate is located vertically in the machine table. Therefore, when making the first layer of the first section, the trunnion table rotates 90\u00b0 to put the plate in a horizontal orientation. As the deposition of the layers progress, the plate turns vertically, and the dome gradually forms (Fig", + " / IFAC PapersOnLine 52-10 (2019) 230\u2013235 235 prevents the collision of the machine head to the table. After the dome is built, it just needs to be cut from the round bar. This is an interesting solution, but it may be difficult to expand upon. The geometry presented in Fig. 4 would be challenging. Fig. 11. Collision at top layers when closing the dome built in the orientation shown in Fig. 10 (a). Fig. 12. Collision of the nozzle to machine table when making dome in the orientation shown in Fig. 10 (b). Fig. 13. Building a dome from the end of a round bar by rotary production strategy. a) 2.5 axis layers at the end of the bar. b) Building the dome at the end of bar. c) The way this prevents collision on the machine. 3.2 Round partitioning of the thin wall dome Based on the orientation showed in Fig. 10 (c), it is needed to be partitioned properly to be able to be built. Here, a 3 + 2 machine configuration is assumed. In order to prevent collisions, the part is split radially, as shown in Fig. 13. The building sequence is shown by numbers in the illustration. The nozzle should be tangent to the surface during production. Top edge of each sub-section is the substrate for starting the next section. Fig. 14. Splitting the surface dome into subsections. Fig. 15. Collision of the nozzle to the table when making section 2. The solution to prevent collisions is to build the part at the end of a plate. The plate is located vertically in the machine table. Therefore, when making the first layer of the first section, the trunnion table rotates 90\u00b0 to put the plate in a horizontal orientation" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0000105_012019-Figure1-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0000105_012019-Figure1-1.png", + "caption": "Figure 1. View of the path of selected points of the Whitworth mechanism.", + "texts": [ + " This program cannot perform detailed motion analyzes, such as determining velocity and acceleration of individual members of the mechanism, but is extremely useful in cases where it is necessary to see how each member of the mechanism is connected and what their mutual mobility is. By entering different input parameters and relationships between members, it is possible to determine the function of movement and make appropriate changes to improve the functioning of the mechanism itself. Also, in this way, it is easy to check whether the dimensions and path of the mechanism are obtained by analytical or graphical method properly defined. Figure 1 shows Whitworth's mechanism with drawn paths of selected moving points. If there is a fault in the design of the mechanism or dimensions, e.g. if the length of the sliding path is not long enough and the slider will slip out of it, the program will report an error, stop with the simulation of the motion, and display the location on the mechanism where irregularities occur. Once the members of the mechanism have been formed, it is necessary to define the driving members, as well as the input parameters such as the speed of the drive member, and the dimension of individual members" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0001693_052026-Figure1-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0001693_052026-Figure1-1.png", + "caption": "Figure 1. Potato tuber chopper: 1 - motor shaft; 2 - a vertical knife; 3 - horizontal knife; 4-cylindrical housing, which at the same time is a receiving hopper; 5 - contradictions; 6 - disks with knives; 7 - unloading hopper;8 - throws.", + "texts": [ + " To study the process of potato tubers, an experimental shredder was used [1], [7]. In the installation, the vertical knives are mounted rigidly at an angle of 90\u00b0 to the axis of the horizontal plane of the horizontal AGRITECH-III-2020 IOP Conf. Series: Earth and Environmental Science 548 (2020) 052026 IOP Publishing doi:10.1088/1755-1315/548/5/052026 knives, with the formation of vertical windows formed from above by the plane of the horizontal knife, below by the plane of the disk, and left and right by the planes of vertical knives, figure 1. The countercutter has the ability to interact with horizontal and vertical knives. The tuber, falling from the loading hopper onto the disk and, when interacting with the contradiction, is evenly distributed on its end surface and is processed. Horizontal and vertical knives exert a force on the product, resulting in its separation into pieces in the shape of a parallelepiped, one of the faces of which corresponds to the profile of the windows formed by the surfaces of the parts of the cutting apparatus" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0000253_2019045-Figure2-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0000253_2019045-Figure2-1.png", + "caption": "Fig. 2. Schematic diagram of determining tooth surface temperature distribution.", + "texts": [ + "When the temperature of gear is changed, the theoretical tooth profile and the practical tooth profile are not superposition. On this basis,Wang [11] proposed a method for calculating the thermal deformation of helical gear. Present experimental measurement technology can get the meshing point temperature value of gear. Reference [1] gives a simplified formula for temperature distribution along gear radial direction, which can be applied in engineering. Based on this, the specific process of the paper is shown in Figure 1. 2 Determination of tooth surface temperature field Figure 2 showed the schematic diagram of determining tooth surface temperature distribution. The temperature values of instantaneous meshing point 1, 2, \u2026, n are firstly measured and fitted. Together with the given temperature formula along gear radial direction, temperature field is determined. 2.1 Measurement and fitting of temperature of instantaneous meshing points on the tooth surface Using the infrared technology, the miniature thermocouple automatic recorder and the simulated heat source infrared technology to measure the temperature of instantaneous meshing points on the tooth surface" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0003618_ijvsmt.2015.067521-Figure1-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0003618_ijvsmt.2015.067521-Figure1-1.png", + "caption": "Figure 1 (a) Mechanism in the vehicle (b) Mechanism CAD model (see online version for colours)", + "texts": [ + " Moreover, the approach followed here considers not only the performance analysis but also the feasibility analysis, which usually is not presented in the works about mechanism suspension. The mechanism topology description and the mobility analysis are presented in Section 2. A kinematic and a kinetostatic mechanism model is developed and described in Section 3. In order to evaluate the mechanism feasibility and performance, a co-simulation was implemented in MATLAB/Simulink and CarSim softwares, and the results and analysis are explained in Section 4. Finally, conclusions are listed in Section 5. The parallel mechanism, shown in Figure 1, is generated by the topological synthesis method (Hess-Coelho, 2007; Malvezzi and Coelho, 2014) and performs three independent motions. The first active motion sets the camber angle, which is achieved by the action simultaneously of the actuators 2 and 3. The second active motion sets the toe angle, obtained only by the movement of the actuator 3. The third motion can be either passive or active. In the passive motion, it is assumed that the car does not perform evasive manoeuvres or cornering" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0003721_1.4879277-Figure8-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0003721_1.4879277-Figure8-1.png", + "caption": "Fig. 8. Two experiments are shown where the concept of negative mass is useful: (a) a helium balloon reaching its (upward) terminal speed, and (b) centrifugal buoyancy forces in a rotating vessel filled with fluid.", + "texts": [ + " For example, a helium balloon that is released will rise upwards and reach a terminal speed that can be calculated by using the well-known result for terminal speed,7 vT \u00bc ffiffiffiffiffiffiffiffiffiffiffi mg=b p and using the modulus of the effective (negative) mass of the 999 Am. J. Phys., Vol. 82, No. 10, October 2014 Notes and Discussions 999 This article is copyrighted as indicated in the article. Reuse of AAPT content is subject to the terms at: http://scitation.aip.org/termsconditions. Downloaded to IP: 129.89.24.43 On: Sun, 05 Oct 2014 09:30:52 balloon [(see Fig. 8(a)]. Another application, similar to what we have discussed in this paper, deals with buoyant centrifugal forces. A mass m inside a rotating centrifuge filled with a fluid will experience an outward centrifugal force given by8 Fc \u00bc m mf x2R, where x is the angular velocity and R is the distance from the axis of rotation. When the density of the particle is lower than that of the fluid, the particle will behave as having a negative effective mass and the centrifugal force will be Fc \u00bc meffx 2R; (6) with meff given by Eq" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0000910_rcar47638.2019.9043930-Figure1-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0000910_rcar47638.2019.9043930-Figure1-1.png", + "caption": "Figure 1. 3D model of quadcopter tilting rotor", + "texts": [ + " Subsequently, the controller\u2019s parameters tuning process using the particle swarm algorithm is designed. Finally, some simulations are executed to prove the effectiveness and the robustness of the proposed control law and conclusions are made. Define the North-East-Down(NED) inertial frame as { } E E EX Y Z , the Forward-Right-Down(FRD) body frame as { } B B BX Y Z . Every rotor can rotate alone its arm axis, { } Ri Ri RiX Y Z is the rotor frame attach to thi rotor group, where RiX points outward alone the arm direction, RiZ points downward along the spinning axis, as shown in Figure 1. Define 3BA SO R as the direction cosine matrix of rotation from frame A to frame B. Therefore, the rotation matrix that represents the orientation of the inertial frame w.r.t body frame is defined as follows, EB c c s s c c s c s c s s s c s s s c c c s s s c s s c c c R where , and corresponds to roll, pitch and yaw angle respectively. s , c represents the trigonometric functions sin( ) and cos( ) . The rotation matrix between thi rotor frame and body frame can be represented as, = (1 ) ( ) 2iB Z XR i R R R where 0 ( ) 0 0 0 1 Z c s s c R 1 0 0 ( ) 0 0 X c s s c R i is the tilt angle of the rotor" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0001206_1350650120929500-Figure3-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0001206_1350650120929500-Figure3-1.png", + "caption": "Figure 3. Schematic diagram of the spindle vibration test scheme.", + "texts": [ + " All data are measured and sent to the central control room of the test station. In the control room, all kinds of operations on site are controlled by microcomputer, and test data are obtained. The test objects are 1# and 2# bearings, which are four-pad tilting pad bearings. The structural schematic diagram is shown in Figure 2. The structural parameters and operating conditions of tilting pad bearing are shown in Table 1. During the test process, vibration data of rotor in two directions are tested. The schematic diagram of vibration test scheme is shown in Figure 3, where the coordinate origin O is the center of bearing block and the coordinate origin Oj is the center of journal. The vibration of the rotor is the vibration of the rotor relative to the bearing block, which is measured by the eddy current sensor fixed on the bearing block. The fixed positions of the eddy current sensor are shown as points 1 and 2 in Figure 3. The reference frame is the bearing block. Absolute coordinate system x0oy0 is the Cartesian coordinates system. In additional, the rotor speed should also be measured by eddy current sensor. An axial groove is opened on the journal, with a width of 6\u201310mm, Driver motor Gas turbine gniraeb#2gniraeb#13# bearing 4# bearing Exciter end Turbine end Coupling Figure 1. Schematic diagram of the test bed structure. a length of 20\u201340mm, and a depth of 0.5\u20131mm. Install eddy current sensor according to the test method of rotor vibration, and adjust the gap voltage to about 8\u201312V" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0001156_cis-ram47153.2019.9095819-Figure1-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0001156_cis-ram47153.2019.9095819-Figure1-1.png", + "caption": "Fig. 1. Overview of a powered knee prosthesis with series-elastic actuator", + "texts": [ + " Secondly, torque controller is developed based on the sensor and, finally, one above-knee amputee walks on the PKP-SEA and conventional prosthesis at the walking speed of 0.8m/s. This preliminary study shows that the series elastic actuator with light-weight and small-size torque sensing elasticity and controller generates walking locomotion, though its behavior looks difference from the one with the conventional prosthesis due to lack of training. 321978-1-7281-3458-1/19/$31.00 c\u00a92019 IEEE Authorized licensed use limited to: UNIVERSITY OF BIRMINGHAM. Downloaded on June 15,2020 at 12:43:24 UTC from IEEE Xplore. Restrictions apply. II. HAREDWARE Figure 1 shows a powered knee prosthesis with a serieselastic actuator (PKP-SEA). A PKP-SEA is composed of DC blushless motor (70W), harmonic gear (1:100 gear ratio), and an elastic spring in series. The detail of the SEA is described in the next section. Comparison of weight, height and range of motion with other representative commercial products are shown in Table I. Ottobock 3R80 is a passive joint with a damper and spring. \u00d6ssur Rheo knee3 and Ottobock Genium are quasi-passive prostheses composed of computercontrolled dampers providing more stable and efficient walking gait" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0002190_00207721.2020.1853272-Figure4-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0002190_00207721.2020.1853272-Figure4-1.png", + "caption": "Figure 4. ThedigraphG2 includingone leader and four followers.", + "texts": [ + " It is verified that all the conditions of Theorem 3.3 are satisfied. According to the definition of , we can calculate = diag(1, 2,\u22124, 1). In view of Theorem 3.3, system (3) can reach weighted synchronisation. In the meantime, there are \u2212x3(\u221e) = 2x2(\u221e) = 4x1(\u221e) = 4x4(\u221e) and \u2212v3(\u221e) = 2v2(\u221e) = 4v1(\u221e) = 4v4(\u221e). Figures 2 and 3 show the evolutions of parameters \u03b8i, positions xi, and velocities vi for i = 1, 2, 3, 4, which is in accordance with Theorem 3.3. Example 4.2: Assume the interaction topology is selected as the digraphG2 in Figure 4, where the cooperative\u2013competitive interactionweights\u03c9ij aremarked besides the edge (j, i). Choose \u03b10 = 1 4 , b1 = b2 = 1 and b3 = b4 = 0, it is easy to see that the digraph G2 Figure 2. Evolutions of parameters \u03b8i for i = 1, 2, 3, 4 over the interactively balanced digraph G1. Figure 3. Evolutions of positions xi and velocities vi for i = 1, 2, 3, 4 over the interactively balanced digraph G1. satisfies Theorem 3.7. By calculation, we can obtain = diag(1, 2,\u22121, 2). According to Theorem 3.7, there are \u03b81 \u2192 \u2212\u03b83 \u2192 \u03b80 = 1 4 and \u03b82 \u2192 \u03b84 \u2192 2\u03b80 = 1 2 " + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0000959_0954406220917995-Figure4-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0000959_0954406220917995-Figure4-1.png", + "caption": "Figure 4. Different wing models used in present study.", + "texts": [ + " In this study, it can be seen that discretization time-step size Dt\u00bc 0.001 s the most suitable. Comparisons of different corrugation patterns and inclination angles during gliding flight This part mainly researches the contributions of different corrugation patterns to aerodynamic forces. Different corrugated wing models with different corrugation patterns/inclination angles and a flat-plate wing model are used to identify the effect of corrugated wing models. We constructed five corrugated wings based on former research31,32 shown in Figure 4. Corrug-1 has two valleys near the leading edge while the left region is flat. Corrug-2 has the same valleys with corrug-1 but has a cambered part. Corrug-3, 4, and 5 are covered with valleys with different inclination angles. These six wing models have the same chord length and spanwise length undergoing the same flapping kinematics. Table 1. The computed cases descriptions in the present study. Case Cell numbers Minimum grid size Time step-Dt (s) D CL/ D CD Grid-1 1.80 106 0.67%c 1.00 10 3 11", + " The CD of plate is always the minimum in all cases when AOA is less than 25 and CD of corrug-2 has the lowest value when AOA is higher than 35 . Corrug-3 rises rapidly than other three cases when AOA exceed 35 . Plate shows good aerodynamic performance when AOA is no more than 25 , which means corrugated wings have no advantage than plate at small AOA under present simulation conditions. To further research the effect of corrugation configurations on aerodynamic performances during steady gliding, we performed simulations with different valley inclinations (shown in Figure 4). Figure 5(c) gives the CL of plate and three corrugated wings with different inclinations. Overall, the CL of plate tends to be higher than corrugated wings, and the CL of Corrug-3, 4, and 5 have the similar shapes, especially a sudden fall at AOA\u00bc 35 . As the results indicate, the inclinations have little impact on the CL when AOA is smaller than 15, but diminishes CL with inclination increasing for AOA above 15 . It can be concluded that the inclination cannot improve the aerodynamic performances of corrugated wing model", + " However, the lift force coefficients do not possess the same values for the corrugated wings. The reason might be that the asymmetric of the wing model around the pitching axis, i.e. two sides of the corrugated wings are not the same. Effects of inclination angles on hovering performances The previous section discussed the effects of various corrugation patterns. This section continues to research the effect of corrugation on aerodynamic forces in hovering flight in the term of inclination angles. The corrugated wings used in this part are described in Figure 4. Figure 9 shows the time history of the aerodynamic force coefficients. Overall, the lift and drag coefficients curves have similar shapes as those in the previous section (shown in Figure 7). Lift force curves have one positive peak (shown in Figure 9(a)), while drag curves meet two positive peaks in a downstroke or upstroke. The first CL peak of corrug-5 is bigger than the second one, which is different from corrug3 and 4. The reason leads to this discrepancy is still unclear. CL of corrugated wings is slightly lower than that of flat-plate, about 5%. For the drag coefficients, one noticeable feature can be seen, where they change little with the increasing of inclination angles, as shown in Figure 9(b). The CD of corrug-5 is bigger than that of corrug-3 and 4 around t\u00bc 4.25T and 4.6T. One possible reason contributes to the difference at t\u00bc 4.6T is bigger smoothing part before valleys as shown in Figure 4. The results presented in Figure 9 show that the inclination angle has finite effect on the aerodynamic. Figure 10 shows the flow field around the flapping wing in terms of the spanwise vorticity. At t\u00bc 4.25T, which reaches to the minimum AOA in one cycle, the wing rotates clockwise around the pitching axis and continues moving downward at the same time. At this moment, there are vortices attached to the upper surface and shed from the trailing edge into the wake. The strength of the vortex decreases from left to right at t\u00bc 4" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0000435_022055-Figure3-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0000435_022055-Figure3-1.png", + "caption": "Figure 3. Joining-up in series of two sections of PRHM [7]", + "texts": [ + "1 4 k times smaller than in involute PRHM (12). Extra decrease of mechanical losses will be due to the decrease of friction coefficient f in the teeth contact close to surface. The restriction of the use of jointed engagement is their disconjugacy hence vibro-activity. The dynamics of such hydraulic machine needs researching. Let us note that the use of such engagements is possible only in schemes M < N: i.\u0435. : 2\u00d74, 4\u00d76, 6\u00d78. D) One more way of increasing efficiency is joining two sections of PRHM in series [7] see figure 3. In this case enlarged passageway channels unload satellite wheels, locking work chambers at the moment of minimum intensity of environment displacement. Dead centres are present only in PRHM of group 1\u00d71\u2013P. Coping with these centres require significant complication of the mechanism of PRHM. ICMTMTE IOP Conf. Series: Materials Science and Engineering 709 (2020) 022055 IOP Publishing doi:10.1088/1757-899X/709/2/022055 This method gives the increase of theoretical efficiency by 1.3 \u00f7 1.4 times without the increase of mechanical loss" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0001895_1350650120964295-Figure5-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0001895_1350650120964295-Figure5-1.png", + "caption": "Figure 5. Two-dimensional sketch of a sphere-cone contact.", + "texts": [ + " The distance d 1Corporate Research, Robert Bosch GmbH Renningen, Renningen, Germany 2Department of Microsystems Engineering, University of Freiburg, Freiburg, Germany Corresponding author: Marius Wolf, Corporate Research, Robert Bosch GmbH Renningen, Robert-Bosch-Campus 1, 71272 Renningen, Germany. Email: marius.wolf2@de.bosch.com between the contact points P1 2 S1 and P2 2 S2 is a local extremum. If B1 \\ B2 6\u00bc 1, the bodies are in contact (e.g. in Figure 4) and the distance between P1 and S2 and between P2 and S1 must be a maximum for P1 2 B2 ^ P2 2 B1. If B1 \\ B2 \u00bc 1, the bodies are separated (e.g. in Figure 5) and P1 and P2 must be chosen such that d equals the global minimum. If d is defined positive for B1 \\ B2 6\u00bc 1 and negative for B1 \\ B2 \u00bc 1, it equals the theoretical maximal penetration of B1 and B2. For the force application point follows Pm \u00bc P1 \u00fe P2 2 (1) State of the art contact detection models An approach for universal contact detection commonly used in commercial MBS programs is based on discretisation of the contacting bodies\u2019 surface into a large amount of polygons and checking interference for each polygon", + " Afterwards, the results must be transformed back into the global coordinate system Jg. In case of a sphere-sphere contact, the connecting vector between PSP1 and PSP2 is collinear with the connecting vector of the spheres\u2019 centre (see Figure 4). Therefore, the contact properties can be calculated via PSP1 \u00bc rSP1 \u00fe cSP SP rSP2 rSP1 krSP2 rSP1kRSP1 (2) PSP2 \u00bc rSP2 cSP SP rSP2 rSP1 krSP2 rSP1kRSP2 (3) d \u00bc RSP1 \u00fe RSP2 cSP SPkrSP2 rSP1k (4) with cSP SP \u00bc 1 if one sphere is concave \u00f0RSPi < 0\u00de 1 else (5) The outer normal n of the cone (see Figure 5) can be expresses as n \u00bc cSP COsinhCO ex;w \u00fe cSP COcoshCO er;w (6) with cSP CO \u00bc 1 if cone is concave 1 else (7) The contact points PCO and PSP are calculated by projecting the sphere\u2019s centre onto the cone along n (see Figure 5). It follows PCO \u00bc xPCO;w ex;w \u00fe rPCO;w er;w and PSP \u00bc xPSP;w ex;w \u00fe rPSP;w er;w with rPCO;w \u00bc rSP;wtanhCO \u00fe xSP;w tanhCO \u00fe \u00f0tanhCO\u00de 1 (8) xPCO;w \u00bc rPCO;w tanhCO (9) rPSP;w \u00bc rSP;w cSP CORSPcoshCO (10) xPSP;w \u00bc xSP;w \u00fe cSP CORSPsinhCO (11) and d \u00bc RSP cSP CO xPCO;w xSP;w sinhCO (12) Equations (8-12) are invalid, if the projection of rSP misses the cone (e.g. is too far left in Figure 5). This occurs for hCO < p=2 ^ xSP;w rSP;w < tanhCOj j (13) which should not be the case for roller-flange contacts. If the flange does not have an opening angle (hCO \u00bc p=2), the cone becomes a disc and numerical issues (tanhCO \u00bc 1) occur for eq. (8) and eq. (9). For this special case, eq. (8-12) can be simplified to rPCO;w \u00bc rSP;w (14) xPCO;w \u00bc 0 (15) rPSP;w \u00bc rSP;w (16) xPSP;w \u00bc xSP;w \u00fe cSP CORSP (17) and d \u00bc RSP \u00fe cSP COxSP;w (18) While concave cones are used for modelling of outer ring flanges, contact pairings involving concave spheres and convex cones are not relevant for contact detection within REBs and are therefore not considered within this paper" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0003438_icinfa.2015.7279530-Figure9-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0003438_icinfa.2015.7279530-Figure9-1.png", + "caption": "Fig. 9, Test rig of QPZZ-II.", + "texts": [ + " The figure of IMF2 also presents a same character, but it isn\u2019t obvious. Other figures could hardly find any useful information. The result shows that two methods are in agreement with each other. So we draw a conclusion that the new method is effective for the extraction of the interest IMF. A test rig named QPZZ-II is used in this test. The faulty bearing is a roller bearing with having outer race faulty. Three acceleration sensor is installed in three directions\uff0cwhich their location be expressed as the figure 9. In test, a small load is used which can present a clear shock signal. And shaft speed is 20Hz. We draw its default signal as figure 10. Next, we apply the new method to extract the interest IMF of the vibration signal. There are nine IMFs after the EMD, and the first third IMFs are shown in figure 11-13. Then we compute their kurtosis values. And the result from the first four IMFs are shown in the table II. IV. CONCLUSION The paper has been proposed a new method about selecting the interest IMF" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0003275_aim.2015.7222530-Figure21-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0003275_aim.2015.7222530-Figure21-1.png", + "caption": "Figure 21. Components of the slider-crank mechanism. A. Platform, B. Connecting rod, C. Piston, D. Laser retlective surface, E. Spring holder, F. TPM case, G. Platform supporter. IJ is the angle form by the external magnetic system and the x axis.", + "texts": [ + " The comparison between the analytical model and experimental results for 'tz exerted on the [PM by the external magnetic system produced by configurations 2 and 3 is presented in Figs. 19-20. The peak torque of 10 mNm on the [PM was possible when the external magnetic field was created by the array of magnets set in Configuration2 (see Fig. 19). For this reason, we used the Configuration2 to actuate the drug release mechanism. Its details are presented in the following subsection. We fabricated the slider-crank mechanism in a plastic material (ABS) with a 3D printer. All its components are depicted in Fig. 21. Since the IPM is connected to the crank of the slider-crank mechanism, the piston will release the drug from the reservoir when the IPM is rotated around the z axis by the external magnetic field (Fig. 22). The crankshaft torque T to balance the discharge force F can be expressed as follows [20]: T = FRsina(l +\ufffdcosa) (2) L By using the law of cosines, we can also express the crank angle, a, as a function of x (i.e., the position of point B) as _ (R Z +X Z _L Z ) a = cos 1 (3) 2Rx Due to practical reasons, we placed a helical spring to measure the piston force F by using the Hooke's law: F = Kilx (4) ilx represents the displacement of the spring and K is the stiffness of the spring. K was measured as l.59 [N/mm]. In our experiments, we manually rotated the arc-shaped magnets and the cubic IPM rotated at the same time, compressing the spring when the crank angle changed from 180\u00b0 to 0\u00b0 and extending the spring when the crank angle changed from 0\u00b0 to -180\u00b0. A laser (opto N COT 1700 by Micro-Epsilon) was used to measure the stroke x as shown in Fig. 23. The beam of the laser reflects on the reflective surface that is connected to the piston as shown in Fig. 21. The laser reading was used to estimate F and the crank angle a given by (3) and (4), respectively. Once F and a are estimated, we use (2) to estimate the torque delivered to the crankshaft by the force F. The slider-crank mechanism was fabricated with R=3 mm and L=9 mm. Fig. 24 shows the piston force generated as the external magnetic system rotates one full cycle, compressing the helical spring in the left hand side of the curve and extending it on the other half of the curve. The experimental peak force of 3" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0001588_aim43001.2020.9158991-Figure8-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0001588_aim43001.2020.9158991-Figure8-1.png", + "caption": "Figure 8 Wire arrangements", + "texts": [ + " One stepper motor (Trinamic Motion GmbH) rotates a low frictional lead shaft (6 mm lead) to slide the lead screw, by which two wires (1 mm steel cable) are oppositely pulled or eased. This mechanical design allows holding tension even the motor turns off. Here we consider a way to control the end-point position of the manipulator. It is assumed that information about wire tensions is only available, which is obtained by the proposed PtM. The eight wires that go out of the pre-tension mechanism are allocated as shown in Fig.8; the wires \ud835\udc4e and ?\u0302? controlled by the motor-a are allocated along a vertical X-axis, the wires c and ?\u0302? controlled by the motor-c are allocated along a horizontal Y-axis, which induces that the motor-a takes a principal role to move the end-point along the vertical axis, and the motor-c takes the one for the horizontal Y-axis. The wires \ud835\udc4e , ?\u0302? and c, ?\u0302? are connected to the distal end of the first section (see Fig.1). While the wires b and b\u0302 driven by the motor-b and the wires d and d\u0302 driven by the motor-d, which are connected to the distal end of the second section (see Fig", + " For example, if one wishes to move the end-point along Y-axis, the motor-a is a principal motor to be driven. (2) Compensatory driving (C-driving): The above primitive driving will induce a deviation along the vertical axis, which should be compensated in actual works of the manipulator. For example, the manipulator will take a load at the end-point, which will move the end-point downward that should be compensated. In the following section, we theoretically explain the compensatory driving denoted above. The pre-tension spring receives 2f (see Fig.8), therefore: 2\ud835\udc53\ud835\udf0e = \ud835\udc58\ud835\udc5d\ud835\udc62\ud835\udc5d\ud835\udf0e , (\ud835\udf0e = \ud835\udc4e, \ud835\udc4f, \ud835\udc50, \ud835\udc51) 2\ud835\udc53\ud835\udf0e\u0302 = \ud835\udc58\ud835\udc5d\ud835\udc62\ud835\udc5d\ud835\udf0e\u0302, (?\u0302? = ?\u0302?, \ud835\udc4f\u0302, \ud835\udc50,\u0302 \ud835\udc51\u0302) (1) where \ud835\udc62\ud835\udc5d\ud835\udf0e , \ud835\udc62\ud835\udc5d\ud835\udf0e\u0302 are compression length of the pretension spring of which spring constant is \ud835\udc58\ud835\udc5d (Fig.10). Let \ud835\udc91\ud835\udc5b = (\ud835\udc5d\ud835\udc5b\ud835\udc65 \ud835\udc5d\ud835\udc5b\ud835\udc66 \ud835\udc5d\ud835\udc5b\ud835\udc65)\ud835\udc47 be the position vector of the end-point and \ud835\udf53 = (\ud835\udf19\ud835\udc4e \ud835\udf19\ud835\udc4f \ud835\udf19\ud835\udc50 \ud835\udf19\ud835\udc51)\ud835\udc47 be rotation angle vector of the 4 motors. Let \ud835\udc7b\ud835\udc5a = (\ud835\udc47\ud835\udc4e \ud835\udc47\ud835\udc4f \ud835\udc47\ud835\udc50 \ud835\udc47\ud835\udc51)\ud835\udc47 be a torque vector loaded on the motor axes. The virtual work principle gives an identity; \u0394\ud835\udc40\ud835\udc54\ud835\udc47\u2206\ud835\udc5d\ud835\udc5b = \ud835\udc47\ud835\udc5a \ud835\udc47 \u2206\ud835\udf19 (2) where \u0394M is a perturbed mass loaded on the end-point and \ud835\udc54 = (0 0 \ud835\udc54)\ud835\udc47 is the gravitation acceleration vector" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0001133_tasc.2020.2994518-Figure2-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0001133_tasc.2020.2994518-Figure2-1.png", + "caption": "Fig. 2. Bending setup.", + "texts": [ + " The plates are covered with Kapton to ease the release of the stack. The assembly is then placed in a vacuum chamber together with a vessel containing previously deareated Stycast. After that the vacuum chamber is evacuated and Stycast is poured on the stack. Then the air is let in again to the chamber and Stycast fills the space between the tapes. Then that the stack is left for curing and released after that. After the first run of measurements the sample is bend to obtain C-shaped stack. The bending setup used is shown in figure 2. The stack is placed on a bending pipe. The pipe lies on three pads driven by screws threaded into U-shaped frame. During the bending the stack is warmed up to 100oC using hot air. At this temperature Stycast becomes elastic. The pipe with the stack is then pushed by screws towards a bending rod attached to the frame. When the desired shape of the stack is obtained the setup is allowed to cool down. The stack becomes rigid again and the process is finished. After bending the stack is magnetised and the trapped magnetic flux is measured again" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0000623_012064-Figure5-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0000623_012064-Figure5-1.png", + "caption": "Figure 5. The location of the rollers in the front bearing box.", + "texts": [], + "surrounding_texts": [ + "Currently, there are significant changes in the operating conditions of wheel sets of railway cars. The greatest contribution to these changes is made by an increase in the speeds of movement of the compositions and an increase in their weight [1\u20133]. At the same time, the decisive contribution to ensuring the reliability of wheelset operation is made by increasing the operational reliability of the wheels and elements of axle boxes \u2013 rollers and bearing rings [4\u20139]. One of the effective ways to improve them is to search for rational design options based on the use of mathematical models of their stress-strain state (SSS) and experimentally verified criteria for assessing their durability [10\u201312]. Under the conditions of cyclic operation of high-loaded bearing elements of structures, such a criterion, as a rule, is the intensity of stresses [13\u201322]." + ] + }, + { + "image_filename": "designv11_71_0000039_j.ymssp.2019.106423-Figure13-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0000039_j.ymssp.2019.106423-Figure13-1.png", + "caption": "Fig. 13. Local zoom at clearance location.", + "texts": [ + " In order to validate the proposed simultaneous identification algorithm, a clearance test-bed which is the cantilever with adjustable clearance value and effective stiffness is applied to conduct the experimental testing, and the complete experimental testing system is shown in Fig. 12. The experimental testing system mainly includes clearance test-bed, excitation module (signal generator, power amplifier and vibration exciter), sensor module (impedance head and eddy current displacement sensor), data acquisition module (LMS data acquisition system) and data analysis module (computer, LMS Test-Lab data acquisition and analysis software). For the clearance test-bed, the local zoom at clearance location is shown in Fig. 13. It shows that clearance studied in paper means the gap existing between the contact head of small cantilever and the cantilever, and the clearance value can be adjusted by adding the slice with different thickness. Additionally, in this clearance test-bed, there are two kinds of stiffness at the clearance location, namely the bending stiffness of the small cantilever kb (in the frequency range 0\u2013 300 Hz, the dynamic stiffness of the small cantilever remains essentially constant and close to its bending stiffness, it is verified by the Ansys analysis and experiment test [1,36]) and contact stiffness, and effective stiffness is essentially the series equivalent stiffness of contact stiffness and bending stiffness" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0001775_0278364920955242-Figure6-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0001775_0278364920955242-Figure6-1.png", + "caption": "Fig. 6. (Left) Schematic of an articulated body in contact with a fixed surface. All Cartesian measurements are made with respect to an inertial reference frame fOg. Body may be in contact at more than one surface. We assume contact to occur at each surface at a single point. Only the ith contact point is shown here. (Right) Let ni denote the contact normal at the point of contact, tX , i denotes one of the two orthogonal tangent directions, and FC, i is the contact force to be determined.", + "texts": [ + " The belted ellipsoids shown in Figure 5(b) represent the contour traced out by the effective mass value as u\u0302 is varied while keeping both the chosen point on the end-effector, and the configuration of the manipulator constant. In the same lines as (4), an effective inertia in u\u0302 direction can be defined with respect to an orientating task at the endeffector. In the following subsection, we use the previous results to derive the governing equations for a system of articulated bodies in a multi-point contact scenario. Consider an articulated body B in contact with a fixed surface at time t as shown in Figure 6. Suppose B is in contact with the environment at mC locations. We assume that the contact patch at each touching surface is modeled as a point contact, although this is not a restrictive assumption as discussed later. Thus, the full contact configuration between the body and the fixed environment is described by mC point contacts. Let xi be the position of contact point ci located on B measured with respect to an inertial frame fOg. Body B is defined to be in contact at point ci if it is collocated with a point on the fixed surface at time t" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0001057_0021998320920920-Figure8-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0001057_0021998320920920-Figure8-1.png", + "caption": "Figure 8. 0xz in the binders with warps and wefts shown for context.", + "texts": [ + " This can be illustrated by contours of the stress in the direction of the applied load, xx, which is in the global coordinate system, as shown in Figure 7. Label A in the figure demonstrates where load from the binder is transferred to the nearby wefts when the binder transitions between travelling along the x-axis and through the thickness. When the tow path of a binder transitions from being aligned with the load to not being aligned with the load, much of the load is transferred to other tows by shear stresses. Figure 8 shows the 0xz concentrations that result in the binders in these regions, and the location the 0xz concentrations in the binders match the locations where xx in the binder dropped significantly. It should also be noted that the most severe 0xz occurs within the volume of the binder, not on the surface, as seen in the cutaway of the binder in Figure 8. The high HTT in these regions are dominated by the large magnitude of 0xz. The regions with the second most severe stresses in the binders occur where z-aligned binders come close to wefts, such as indicated by label B in Figure 5. For all of these regions, the failure index corresponding to transverse tension, HTT, is highest, as shown in Figure 6. The stress component that contributes the most to HTT where the binders traverse the textile thickness is 0zz. Figure 9 shows 0zz in the binder, and the locations of 0zz concentrations can be seen to match the locations of local maxima of Hmax, shown in Figure 5" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0001523_iwcmc48107.2020.9148486-Figure1-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0001523_iwcmc48107.2020.9148486-Figure1-1.png", + "caption": "Fig. 1. Network architecture of the tethered UAV airborne system", + "texts": [ + " The lightweight NSA BBU supports both the connection to the public NSA core network through satellite links and the connection to the local sinking emergency NSA core network, so 5G NSA terminals or 4G terminals not only support public network communication using public network SIM cards, but also support private network connection for emergency communication using private network SIM cards. The airborne RRU is mounted on the UAV platform to realize the urgent deployment of emergency communications. The network architecture of the tethered UAV airborne system is shown in Figure 1. The scheme uses a tethered UAV as a flight platform to provide signal coverage for emergency communication service. The flying height of the UAV can be adjusted between 0 and 200 meters according to the demand, and continuous power supply is provided through a photoelectric hybrid cable connected to the ground retractable box. \u2162. UAV PLATFORM This article selects a long-endurance tethered UAV system, as shown in Figure 2, which can mount on a variety of equipment, widely used in emergency communication, fire rescue, environmental protection and military scientific research and other fields" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0001225_s00170-020-05332-8-Figure16-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0001225_s00170-020-05332-8-Figure16-1.png", + "caption": "Fig. 16 The hardware of the robot Rubik. (a) Robot Rubik. (b) Microcontroller. (c) Buck module. (d) Distance module. (e) Communication module. (f) Motion controller. (g) Deceleration motor", + "texts": [ + " Although the number of robots is small, we believe that the results of these experiments verify the feasibility of the self-assembly algorithm. Based on the simulation robot Rubik, we constructed 12 physical modular robots Rubik. Each of them is capable of communication, computation, and moving. We built the bottom motion platform with three omnidirectional wheels placed at an angle of 120\u00b0. It can achieve the motion in any direction in the plane, but we only need to use four of the specific directions. As shown in Fig. 16, modular robot hardware mainly includes 6 components, and each module is listed in Table 1. All robots are aggregated and arranged in a regular grid pattern initially. After pretreatment, each robot gets its location differentiates into different states. Referring to the selfassembly stratified mechanism, a motion-chain is generated at the edge of the current configuration. Then, individuals within the motion-chain start to move into the desired shape orderly. As shown in Fig. 17, we realized a simple \u201chat\u201d shape formation" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0003275_aim.2015.7222530-Figure13-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0003275_aim.2015.7222530-Figure13-1.png", + "caption": "Figure 13. Experimental set up consisting of measurement instruments and the array of arc-shaped magnets.", + "texts": [ + " The IPM was connected to the torque sensor via a green plastic connector that was prototyped using a 3D printer. The torque sensor and the probe tip of the gauss meter were mounted on plastic holders which were also fabricated with a 3D printer. Both the torque sensor and the probe tip of the gauss meter can be moved along the X and Z axes and the arrays of magnets can only be moved along the Y axis. These displacements are controlled by a micromanipulation system constructed of XYZ stages as shown in Fig. 13. The comparison between analytical model and experimental results for the Bx produced by radially and tangentially magnetized arc-shaped magnets is presented in Figs. 14-15. The comparison between the analytical model and experimental results for the total Bx generated by Configurations 2, 3 and 4 is presented in Figs. 16-18. These experimental results for Bx generated by Configurations 2, 3 and 4 validate the analytical models that we use for the optimization of the magnetic system. However, we found a small difference in Fig" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0001961_j.matpr.2020.09.086-Figure3-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0001961_j.matpr.2020.09.086-Figure3-1.png", + "caption": "Fig. 3. The fragment of the turbine impeller.", + "texts": [ + " For example, dimensional electrochemical treatment used to remove burrs in hard-to-reach places, to form artificial roughness and differential elimination of the thickness difference of the material of the cooled shells (Fig. 1), can be applied (after appropriate adaptation) and for aligning the microgeometry of a narrow interscapular canal of variable section of rotor parts and assembly units of a turbo-pump assembly (Fig. 2). At the same time, it becomes possible to combine refinement with predetermined quality indicators of second-order surfaces of narrow (less than 5 mm) turbine interscapular channels (Fig. 3), which are currently processed only at the input and output edges, without developing technologically difficult to access internal zones surfaces of the blades [1]. In Fig. 1, caverns for turbulence of the cooler flow are visible on the bottom of the fins, which are performed by the laborious uipment plants, electroerosion method (it takes several hours to complete one part, depending on the dimensions). This operation is preceded by laborious deburring along the edges of the ribs, which is weakly mechanized (again, within a few hours or two or three shifts on one shell, depending on the size)" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0001654_tmag.2020.3019082-Figure1-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0001654_tmag.2020.3019082-Figure1-1.png", + "caption": "Fig. 1. 4-phase SRM; (a) conventional coil arrangement, (b) proposed coil arrangement, and (c) inverter.", + "texts": [ + " Some coil arrangements have been investigated for 3- phase SRMs, and those for 4-phase SRMs cannot be seen [5]- [8]. In this paper, a new coil arrangement for a 4-phase SRM with 12- and its multiple slot stators is proposed. In addition, the performance of the proposed coil arrangement is compared with that of the conventional coil arrangement using a 24-slot stator. The torque ripple and radial force are compared, and the cause of their difference are discussed. Finally, the motor efficiency is compared. A. Conventional and new coil arrangements Fig. 1(a) shows the conventional 4-phase SRM with conventional coil arrangements. The number of poles and slots is 6 and 8, respectively, and 2 patterns of coil arrangement are available. However, these coil arrangements cannot be expanded to a 12-slot stator due to its magnetic symmetry. On the other hand, Fig. 1(b) shows the 4-phase SRM with a new coil arrangement. The number of poles and slots is 9 and 12, respectively. These SRMs are driven using an asymmetric inverter shown in Fig. 1(c). These coil arrangements are compared in terms of the torque, radial force performance, and motor efficiency using a 24-slot stator. B. Comparison Model Fig. 2 and Table I show the comparison model of a 24/18 4- phase SRM with conventional or new coil arrangements shown in Table II, where these are designed for a traction motor of a heavy construction equipment. The whole size of these SRMs is decided by the application, and each dimension is designed so that the torque at 1500 rpm can be maximum", + " Torque Comparison In order to evaluate the torque characteristics of 3 models, 2-D finite element analysis is conducted, where a commercial software (JMAG-Designer (JSOL Corporation)) is used. A half model shown in Fig. 2 is used, where the mesh shape is a triangle (Fig. 3), and slide meshes are applied to the air gap. The rotor position is applied so that the rotation speed will be at 1500 rpm, and pulse voltages are applied to each phase, where the DC-link voltage is 550 V. In order to apply pulse voltages on the coils, the electrical circuit shown in Fig. 1(c) is coupled with the electromagnetic analysis. The leading angle and application width of the applied voltage are adjusted to increase the torque. The maximum phase current is 270 A, and a conventional hysteresis control is applied to 3 models so that the phase current can be lower than 270 A. The leakage flux of the coil end is ignored, and the torque is calculated using a nodal force method. Figs. 5(a), (b), and (c) show the torque waveforms of Models A, B and C, respectively, where ta, tb, tc, and td are torque due to the A-, B-, C-, and D-phase coils, respectively, and T is the total torque" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0002059_ecce44975.2020.9236256-Figure16-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0002059_ecce44975.2020.9236256-Figure16-1.png", + "caption": "Fig. 16. Photos of the prototype HE3 machine. (a) Stator and winding. (b) Salient rotor.", + "texts": [ + " It shows that the on-load EIV and VPR vary with the change of \u03b8a, as the permeability of the stator and rotor laminations varies when \u03b8a is changed. Therefore, both the no-load induced and the AC current induced voltage pulsations are related to \u03b8a and thus the onload induced voltage pulsation. It can be found that the HE4 machine with E-core stator has the highest on-load EIV and VPR among those HESFMs over the whole range of \u03b8a due to abundant harmonics and higher amplitude of low order harmonics. IV. EXPERIMENTAL VALIDATION As illustrated in Fig. 16, a prototype HE3 machine is designed and fabricated for further verification of the previous FE analyzes. The prototype machine consists of a stator with 12-slots and a salient rotor with 10-poles. The DC field winding of the prototype machine has twelve DC field coils, in which they are connected in series. In addition, each field coil contains 10 turns, which is lower than that in the previous FE calculations to facilitate the winding process of the prototype machine. 4686 Authorized licensed use limited to: Carleton University" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0001455_access.2020.3009089-Figure3-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0001455_access.2020.3009089-Figure3-1.png", + "caption": "FIGURE 3. Diagram of the barrel servo system.", + "texts": [ + " The curves of the gravitational torque and the balance torque provided by the active balancing system are depicted in Figure 2. As can be seen, although not totally compensated, the unbalanced torques are kept at small arrange throughout the whole working area. It is also should be noted that the active balancing system is totally isolated from the barrel servo system, which make the dynamics of the barrel servo system significantly simplified. B. DYNAMICS OF THE BARREL SERVO SYSTEM The diagram of the barrel servo system is given in Figure 3. As can be seen, the structure and the placement of the actuator are the same as that of the active balancing system, thus the synchronous of the two actuators are guaranteed physically. By using the Newton\u2019s Second Law, the mechanical dynamics of the barrel system can be given as J \u03b8\u0308 + B\u03b8\u0307 + Td = Fl sin\u03b2 (1) where J is the equivalent moment of inertia; B is the equivalent viscous damping coefficient; Td is the equivalent load torque including the unbalanced gravitational torques, external disturbances, nonlinear frictions and other unmolded dynamics" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0003714_esda2014-20232-Figure7-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0003714_esda2014-20232-Figure7-1.png", + "caption": "Figure 7: Boundary conditions and geometry: a) Rotor, b) Housing, c)Pedstall, d) Stand, and e) Load cell arm", + "texts": [ + " However, the analysis has been carried out on some important parts according to the need and importance which is more time and cost effective. The effect of the other parts are only considered in boundary conditions of the important components during the analysis. In some cases, transient analysis is needed to assess critical situations. The static load analysis in is done in 1 second period of time which is equivalent to a time step. Mode shape analysis is also performed for the structural components for predicting critical damage locations. The model of the rotor is shown in figure 7(a). The purpose of this analysis is to provide an estimation of the shaft behavior in different operating conditions in order to confirm its mechanical performance. In this regard, the natural frequencies of the rotor are obtained. As well, the modal analysis and comparison of the results with operating frequency range is performed. Static analysis is needed to check the rotor displacement caused by the weight. Also, transient analysis of the rotor is crucial to check its dynamic behavior to confirm the structure mechanical strength. As shown in figure 7(c), the pedestals are designed to support the water brake by two cylindrical bearings. For these two parts, modal analysis and static analysis are performed to ensure their proper mechanical performance. It should be noted that due to the nonrotating nature of the pedestal, the dynamic analysis is not required. As illustrated in figure 7(d), stand is a component that connect the dynamometer to the foundation. This assembly needs to be analyzed in terms of the weld resistance and tensileshear performance as well as vibration. Here, modal and static structural analysis are presented. The load cell arm, which is illustrated in figure 7(e), transmits the load from the dynamometer housing to a load cell. A transient analysis is carried out considering the instantaneous loading that the housing faces during the test. Some results are shown in figures 8-12 in order to verify that the system is well designed and to ensure that the design goals are met in a repetitive cycle. As illustrated in the figures, results show that the final design is well accomplished and no defects are found in the water brake components. Downloaded From: http://proceedings" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0001064_iccais46528.2019.9074699-Figure2-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0001064_iccais46528.2019.9074699-Figure2-1.png", + "caption": "Fig. 2: Single serial branch.", + "texts": [ + " There are one wrench and one link in each serial chain and they are connected by sphere joint. The revolute joints are actuated by the motors and the sphere joints are passive. The body frame \u03a3O and \u03a3E locate at the geometrical centers of the base platform and end-effector respectively. In this paper, the coordinate vectors are defined w.r.t. the base frame by default. The coordinate position vectors of the revolute and sphere joint centers are denoted as Bi, Ti, and Ai(i = 1, 2...6) respectively. In Fig. 2, \u03a3Wi and \u03a3Li are the wrench and link frames. wi, li and ai represent the vectors of\u2212\u2212\u2192 BiTi, \u2212\u2212\u2192 TiAi and \u2212\u2212\u2192 EAi respectively. cwi and cli are the centers of mass (C.O.M.) of the wrenches and links respectively. The joint space vector \u03b8 = [\u03b81, \u03b82, ..., \u03b86]T denotes the angle positions of the active revolute joints. The workspace pose of the end-effector is represented by \u03c7E = [ hT ,\u03c6T ]T , where h = [x, y, z] T and \u03c6 = [\u03b1, \u03b2, \u03b3] T . The velocity of the end- effector is expressed as vE = [ h\u0307T ,\u03c9T ]T \u2208 R6\u00d71" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0001117_kem.841.327-Figure1-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0001117_kem.841.327-Figure1-1.png", + "caption": "Fig. 1 Meshed model of the specimen to be simulated.", + "texts": [ + " For every 7 grams of Epoxy resin added, 1 gram of hardener was added. Finite Element Modelling. ANSYS ACP is a package intended for composite modelling. It consists of a Pre-processing tool where various properties regarding the material, fibres are compiled. The composite Specimen is modelled according to the ASTM D638 and ASTM D790 specimen size in a CAD modelling software and then imported into the ANSYS PrePost in the WORKBENCH. The finite element model for tensile test created in the PrePost is shown in Fig. 1. A mesh element size was 2 mm was chosen for the layers. Face mesh was applied in order to produce less warping and better aspect ratio of the elements. The FEM Model was solved for Equivalent Stresses, Equivalent Elastic strain in the specimen and equivalent stress in the GFRP and sugarcane fibres. Non-linear analysis was conducted in 1000 steps of computation. Also, the model was solved considering large deflection in the specimen. Loading Conditions. A CAD model of metal clamps was imported into the simulation as shown in Fig" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0002324_iceca49313.2020.9297379-Figure4-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0002324_iceca49313.2020.9297379-Figure4-1.png", + "caption": "Figure 4: Top View", + "texts": [ + " III. DYNAMIC MODELLING As to all VTOL aircraft, this design has two modes of operation: Hovering and Gliding. Thus the dynamic model of the aircraft is divided into two modes: Hovering mode and Gliding mode. The drone uses four motors for propulsion and control. It can be maneuvered by changing the speed of these motors. Brushless DC motors are used in this drone whose angular velocities can be varied by varying frequency of the current supplied with the help of ESC (Electronic Speed Controller). Figure 4 shows the direction of rotation and the configuration of four motors when v iewed from the top. The following figure 5 shows the configuration of motors, their direction of rotation and the orientations of axis for hovering mode. Let, x, y and z be three perpendicular axis attached to the fixed frame as shown in figure 5. Roll axis, Pitch axis and Yaw axis be three perpendicular axis attached to the frame of the drone as shown in figure 5. T1, T2, T3 and T4 be the thrusts developed by the motors 1, 2, 3 and 4 respectively" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0001905_case48305.2020.9216801-Figure2-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0001905_case48305.2020.9216801-Figure2-1.png", + "caption": "Fig. 2. FOV diagram in the x-z plane", + "texts": [ + " Its various components are described below g1 bounds = vt \u2264 vmax, \u2212vt \u2264 vmin (11) g2 bounds = \u03c6t \u2264 \u03c6max, \u2212\u03c6t \u2264 \u03c6min (12) g3 bounds = \u03b3t \u2264 \u03b3max, \u2212\u03b3t \u2264 \u03b3min (13) g4 bounds = v\u0307t \u2264 v\u0307max, \u2212v\u0307t \u2264 v\u0307min (14) g5 bounds = \u03c6\u0307t \u2264 \u03c6\u0307max, \u2212\u03c6\u0307t \u2264 \u03c6\u0307min (15) g6 bounds = \u03b3\u0307t \u2264 \u03b3\u0307max, \u2212\u03b3\u0307t \u2264 \u03b3\u0307min (16) g7 bounds = \u03c8\u0307t \u2264 \u03c8\u0307max, \u2212\u03c8\u0307t \u2264 \u03c8\u0307min (17) The penalties formed with gibounds are given below \u03bb1 = max(0, vmin \u2212 vt) +max(0, vt \u2212 vmax) (18) \u03bb2 = max(0, \u03c6min \u2212 \u03c6t) +max(0, \u03c6t \u2212 \u03c6max) (19) \u03bb3 = max(0, \u03b3min \u2212 \u03b3t) +max(0, \u03b3t \u2212 \u03b3max) (20) \u03bb4 = max(0, v\u0307min \u2212 v\u0307t) +max(0, v\u0307t \u2212 v\u0307max) (21) \u03bb5 = max(0, \u03c6\u0307min \u2212 \u03c6\u0307t) +max(0, \u03c6\u0307t \u2212 \u03c6\u0307max) (22) \u03bb6 = max(0, \u03b3\u0307min \u2212 \u03b3\u0307t) +max(0, \u03b3\u0307t \u2212 \u03b3\u0307max) (23) \u03bb7 = max(0, \u03c8\u0307min \u2212 \u03c8\u0307t) +max(0, \u03c8\u0307t \u2212 \u03c8\u0307max) (24) We use finite difference ut+1\u2212ut \u2206t to model the derivatives on the control variables. Fig. 1 shows the FOV of a downward-facing camera rigidly attached to the body of FWV. The projection of FOV on the x\u2212 y plane can be modeled as the quadrilateral PQRS. We derive the FOV constraints by projecting the FOV prism to x\u2212 z and y\u2212 z plane. The projection on the x\u2212 z plane is shown in Fig.2 Let (xtargett , ytargett ) be the position of the target vehicle on the x\u2212 y plane and the difference between the x-coordinate of target vehicle and FWA ~dx = xt \u2212 xtargett . For the target be in FOV, we must have \u2212\u2212\u2192 AD < ~dx < \u2212\u2212\u2192 CD (25) From Fig.2, we have \u2212\u2212\u2192 CD = zt tan(\u03b8x + \u03b3t) (26) \u2212\u2212\u2192 AD = zt tan(\u2212(\u03b8x \u2212 \u03b3t)) (27) Therefore, the first half of the FOV constraint can be written as g1 fov : zt tan(\u2212(\u03b8x \u2212 \u03b3t)) < ~dx < zt tan(\u03b8x + \u03b3t) (28) Similarly, by projecting the FOV prism along the y \u2212 z plane, we can derive the following the other half of the FOV constraint. g2 fov : zt tan(\u2212(\u03b8y \u2212 \u03c6t)) < ~dy < zt tan(\u03b8y + \u03c6t) (29) 2\u03b8x and 2\u03b8y are the camera\u2019s angle of view along x and y axis respectively. The difference between the y-coordinate of target and FWA be ~dy ,i" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0001775_0278364920955242-Figure11-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0001775_0278364920955242-Figure11-1.png", + "caption": "Fig. 11. Schematics of the three experimental setups designed for studying (a) single-point planar (TP1), (b) two-point planar (TP2), and (c) single-point spatial collisions (TS1). Each setup consists of passive multi-joint pendulums hanging from a fixed structure. The pendulums are allowed to collide by dropping them from different initial configurations. Joint positions, velocities, and accelerations are measured at a rate of 4 kHz using high-precision optical joint encoders.", + "texts": [ + " Instead, if one only considered the effective mass of the rigid body RF (Figure 10(c)), the inertia of the colliding body would be underestimated. Lastly, we note that although the previous inferences were drawn for the multi-point collision scenario, identical inferences can be drawn for the multi-point steady contact scenario regarding the generalized accelerations and change of generalized momenta. To test the accuracy of the CSR model, we devised several setups consisting of pendulums with serial kinematic chain structures as shown schematically in Figure 11. The fabricated setups are shown in Figure 12. In each experiment, a pair of pendulums hangs from a rigid and grounded frame, close enough such that they collide often when swinging. Figure 13 shows a few frames from a video of the pendulums colliding on one of the devised setups. The left pendulum is dropped in frame 1, and collision occurs in frame 3. To ensure that the pendulums are rigid and joint friction is negligible, the links are constructed out of aluminum, and steel precision ball bearings are used at each joint" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0001380_s1068798x20060076-Figure2-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0001380_s1068798x20060076-Figure2-1.png", + "caption": "Fig. 2. Saw unit with circular blade motion: (1) saw blades; (2) upper hinge with corrective mass; (3) elastic elements; (4) lower hinge with corrective mass; (5) lateral support; (6) pulley; (7) upper shaft; (8) lower shaft; (9) shaft bearings; 10) cam; (11) lift mechanism for upper shaft.", + "texts": [ + "eywords: saw module, saw assembly, dynamic stability, corrective masses DOI: 10.3103/S1068798X20060076 A fundamentally new saw system is undergoing development and plant testing at present; versions for cutting logs and double-edge beams are under design (Fig. 1). The machine consists of two main mechanisms: one for introducing the workpiece in the cutting zone; and the other for removing the finished product from the cutting zone (Fig. 2). The system includes six identical saw modules (Fig. 3). Sharpening the tool in the saw module decreases its mass. In addition, in order to decrease costs, long blades are replaced by short blades. In other words, the number of teeth is decreased, since coating with hard alloy of stellite type to increase tooth hardness considerably increases the cost of the blade. 45 With variation in mass of the saw modules, their dynamic characteristics are impaired, and the operational stability of the blades is decreased" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0001850_icra40945.2020.9196951-Figure1-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0001850_icra40945.2020.9196951-Figure1-1.png", + "caption": "Fig. 1. The pseudo-rigid-body model of the robotic catheter with two sets of tri-axial actuation coils subject to the surface contact constraint. The contact frame C is chosen such that its origin is located at the contact point of the catheter and the z-axis is the outward normal of the tissue surface. x0 denotes the contact point position in surface frame, which remains static during heart surface motion. The spatial frame S is given as shown.", + "texts": [ + " In this paper, we present the analysis of the contact stability of the MRI-actuated robotic catheter under cardiac tissue motions. Unlike the above studies, the contact force control schemes proposed in this paper are formulated without using additional force sensors, and the contact force control algorithms are proposed based on the calculation of the contact force actuation Jacobian. In the Pseudo-Rigid-Body (PRB) model, the catheter is modeled as a series of pseudo-rigid links connected by elastic joints (Fig. 1) [23]\u2013[26]. Each joint is modeled as a spherical joint which is parametrized by rotation angles in a set of axis angle representations \u03b8i = [\u03b8i1,\u03b8i2,\u03b8i3] T [27], which implies non-ordered rotations. The twist of the i-th joint is given as \u03bei(\u03b8i) = [ \u2212\u03b8i\u00d7qi \u03b8i ] , (1) where qi is the initial position of the i-th joint in spatial frame. The shape of the catheter under the n+ 1-link PRB model can then be described by the joint angle vector \u03b8 = [\u03b8 T 1 ,\u03b8 T 2 , . . . ,\u03b8 T n ] T \u2208 C \u2282 R3n, where C denotes the configuration space of the catheter robot [27]", + " (12) At a given joint configuration \u03b80, A(\u03b8) and N(\u03b8) can then be approximated at \u03b80 using Taylor\u2019s theorem as A(\u03b8) \u2248 A(\u03b80) + A\u2032(\u03b80)\u2206\u03b8 and N(\u03b8) \u2248 N(\u03b80) + N\u2032(\u03b80)\u2206\u03b8 , where A\u2032(\u03b80)= \u2202A/\u2202\u03b8 |\u03b8=\u03b80 2 and N\u2032(\u03b80)= \u2202N/\u2202\u03b8 |\u03b8=\u03b80 . Then (12) can be linearized as JT C fc \u2248(A(\u03b80)+A\u2032(\u03b80)\u2206\u03b8)u\u2212 (N(\u03b80)+N\u2032(\u03b80)\u2206\u03b8)\u2212K\u03b80. (13) Since the joint displacement \u2206\u03b8 is small based on the quasistatic assumption, (13) can be approximated as JT C fc \u2248 A(\u03b80)u\u2212N(\u03b80)\u2212K\u03b80. (14) The contact force-actuation Jacobian Jcu at joint configuration \u03b80 is then calculated as Jcu = d fc/du\u2248 JT c \u2020A(\u03b80). Consider a given target catheter tip position on the tissue surface. This target position is assumed to be static in the surface frame, as shown in Fig. 1. The goal of the contact force control is to improve the stability of the contact between the catheter and the tissue surface under heart surface motion, namely, maintaining static catheter tip positioning at the desired location on the tissue surface with adequate normal contact force, despite cardiac motion. In this section, first a quasi-static contact force control algorithm which computes a set of actuation currents for a desired normal contact force and target tip position for a given instantaneous surface configuration is introduced (Algorithm 1)" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0003897_1.4856595-Figure1-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0003897_1.4856595-Figure1-1.png", + "caption": "Fig. 1. A rigid double pendulum consisting of two point masses attached to a massless rigid rod.", + "texts": [ + "200 On: Sun, 05 Oct 2014 17:27:25 particle systems are subject to constraints, the standard model must be supplemented in some manner in order to deal with the unknown constraint forces.3 In a paper published in this journal in 1982, Stadler4 drew attention to the fact that the principle of angular momentum is routinely applied to systems that do not consist only of point masses, but contain constraints as well. As an illustrative example he considered planar motion of a rigid double pendulum, consisting of two point masses P1 and P2 fastened to a massless rigid rod and pivoting about a fixed point O in an inertial frame of reference (see Fig. 1). The correct solution to this problem can be obtained in several different ways and in particular by a direct application of the principle of angular momentum. However, there is a subtle logical point that is usually overlooked: to apply the principle of angular momentum as a theorem, one would need to show that the rigid rod can be replaced by a set of internal forces that satisfy the collinearity condition. Thus, Stadler4 supposed that the constraints were enforced by internal forces between the two particles and a third particle P0 fixed at the pivot" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0000829_ijthi.2020070107-Figure5-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0000829_ijthi.2020070107-Figure5-1.png", + "caption": "Figure 5. Tool path simulation of the typical part in NC turning centre", + "texts": [], + "surrounding_texts": [ + "During\ufeffthe\ufeffwhole\ufeffprocess,\ufeffthe\ufeffstudents\ufeffinvolve\ufeffpersonally\ufeffin\ufeffthe\ufeffrunning\ufeffof\ufeffthe\ufeffentire\ufefftask\ufeffof\ufeffNC\ufeff machining\ufeffsimulation.\ufeffOn\ufeffthe\ufeffone\ufeffhand,\ufeffthey\ufeffare\ufeffinstructed\ufeffby\ufeffus\ufeffteachers\ufeffto\ufeffaccomplish\ufeffsmoothly\ufeff the\ufefftask;\ufeffon\ufeffthe\ufeffother\ufeffhand,\ufeffthey\ufeffare\ufeffassisted\ufeffwhen\ufeffencountering\ufeffdifficulties\ufeffin\ufeffhandling\ufeffthe\ufeffprocess\ufeff of\ufeffNC\ufeffmachining\ufeffsimulation\ufeffsystem.\ufeffAnd\ufeffit\ufeffwould\ufeffbe\ufeffexemplified\ufeffas\ufefffollows.\ufeffWhile\ufeffthe\ufeffstudents\ufeff involve\ufeffpersonally\ufefffrom\ufeffthe\ufeffvery\ufeffbeginning,\ufeffthey\ufeffcould\ufeffgradually\ufeffunderstand\ufeffand\ufeffmaster\ufeffthe\ufeffbasic\ufeff requirements\ufeffof\ufeffeach\ufefflink\ufeffof\ufeffthe\ufeffproject,\ufeffand\ufeffthe\ufeffcrucial\ufeffpoints\ufeffof\ufeffthe\ufeffwhole\ufeffcognitive\ufeffprocess." + ] + }, + { + "image_filename": "designv11_71_0003255_aim.2015.7222699-Figure2-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0003255_aim.2015.7222699-Figure2-1.png", + "caption": "Fig. 2. Equivalent biped models with bent- and arc-leg frames", + "texts": [ + " 1 shows the model of an underactuated biped robot. This consists of two identical rigid leg frames and three point masses, and can exert the hip-joint torque, uH . The main difference from the commonly-used (point-footed) compasslike biped robot [11] is the CoM location of each leg frame. The CoM of the leg frames is positioned at a distance of w [m] from the central axis. In other words, the CoM is not on the hip-foot line. We set the anterior direction to positive. The inertia moments about the masses are ignored. Fig. 2 shows the equivalent models for Fig. 1. The robot consists of two identical leg frames that are not compliant or cannot bend during motion. Such bent leg frames with positive w have the effect to shift the robot CoM forward. Then the robot can easily overcome the energy (potential) barrier at mid-stance [12]. 978-1-4673-9107-8/15/$31.00 \u00a92015 IEEE 1184 Define \u03b81 and \u03b82 as the angular positions of the stance and swing legs with respect to vertical. Let \u03b8 = [ \u03b81 \u03b82 ]T be the generalized coordinate vector" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0001793_022006-Figure2-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0001793_022006-Figure2-1.png", + "caption": "Figure 2: Coordinates of all structural loads measurements.", + "texts": [ + " Project data recorded from four reference commercial turbines was used to validate the load-prediction algorithm. 1. Prototype model of a 6-MW turbine erected in North Germany, 2. Prototype model of 2- MW turbine located in China, 3. One offshore 5-MW Senvion turbine located at Nord Sea, which are part of the Research at Alpha Ventus (Rave) offshore research project are used. All the structural loads in all reference turbines are measured in accordance with IEC TS 61400-13 and the coordinate system used for the measurement of loads is shown in Figure 2. More technical details from the reference turbines are not described in this paper due to the clause of confidentiality. Table 1 and Table 2 describes the list of available SCADA signals and structural data in the reference turbines. Table 3 describes the number of available concurrent SCADA and structural data available from each reference turbine for validation. The red stars in the table indicates that a particular signal is not measured in the respective turbine. The Science of Making Torque from Wind (TORQUE 2020) Journal of Physics: Conference Series 1618 (2020) 022006 IOP Publishing doi:10" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0001099_j.promfg.2020.02.102-Figure1-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0001099_j.promfg.2020.02.102-Figure1-1.png", + "caption": "Figure 1. Elements of sustainable production exemplified by additive manufacturing and investment casting for three applications.", + "texts": [ + " The localized production minimizes transport and hence emissions are reduced. Drawbacks of the technology are the low production rates, and especially in the case of powder-bed based methods limited part dimensions and increased safety requirements. For conventional technologies such as investment casting the same principles may be applied \u2013 sustainability is improved by using recycled material as well as optimized and energy efficient heat treatment processes to manufacture multifunctional lightweight products. Figure 1. Elements of sustainable production exemplified by additive manufacturing and investment casting for three applications. This paper assesses the sustainability of two manufacturing technologies, laser beam melting and investment casting (Fig. 1). Examples of application include topology-optimized grippers and stochastic as well as defined cellular metal structures. Materials investigated are stainless steel and alloys of aluminum and titanium. Laser beam melting is a very versatile additive manufacturing technology and is used as base process in this paper. The starting point is a 3D CAD model, which is virtually sliced into thin layers with a layer thickness of approx. 45 \u00b5m. Based on this data the physical part is built. A thin layer of metal powder, stored in a dose chamber, is uniformly distributed across the working area by a coater blade (Fig", + " Drawbacks of the technology are the low production rates, and especially in the case of powder-bed based methods limited part dimensions and increased safety requirements. For conventional technologies such as investment casting the same principles may be applied \u2013 sustainability is improved by using recycled material as well as optimized and energy efficient heat treatment processes to manufacture multifunctional lightweight products. This paper assesses the sustainability of two manufacturing technologies, laser beam melting and investment casting (Fig. 1). Examples of application include topology-optimized grippers and stochastic as well as defined cellular metal structures. Materials investigated are stainless steel and alloys of aluminum and titanium. Laser beam melting is a very versatile additive manufacturing technology and is used as base process in this paper. The starting point is a 3D CAD model, which is virtually sliced into thin layers with a layer thickness of approx. 45 \u00b5m. Based on this data the physical part is built. A thin layer of metal powder, stored in a dose chamber, is uniformly distributed across the working area by a coater blade (Fig" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0000830_0954410020911832-Figure1-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0000830_0954410020911832-Figure1-1.png", + "caption": "Figure 1. Coordinates system diagram.", + "texts": [ + ",16 was developed based on the following assumptions: 1. The principle axis of inertia of the combination is parallel to those of the forward and aft parts. 2. Only the roll constraint moment is considered between the forward and aft body. 3. The forces and moments, except those in the roll channel, are considered to apply to the combination. Projectile 7DOF dynamic model Firstly, the inertial reference frame (IRF) OXYZ, the velocity reference frame (VRF) OX2Y2Z2, and the no-roll reference frame (NRRF) O are established, as shown in Figure 1. The direction of OX corresponds to the launch direction, OY vertically up, and OZ to the right. The direction of OX2 axis corresponds to that of the projectile\u2019s velocity vector ~v, and the direction of O axis is the same as the axis-direction of projectile ~x. Both of the VRF and NRRF can be obtained by rotating and panning IRF. To obtain NRRF, the IRF rotates along OZ axis, and then along O . Similarly, when , and O are substituted with v, v, and OY2, the VRF can be obtained. v, v are the pitch and yaw angles of the velocity vector, respectively", + " Clearly, the dynamic model of a dual-spin projectile considering wind is almost the same as the model derived when ignoring wind in this format. The main difference caused by atmospheric wind in the dynamic model is its effect on the aerodynamic forces and the moments of the projectile and canards. Uncontrollable forces and moments with atmospheric wind The angles of attack and sideslip , are the angles between the vector of the axis direction of the projectile and that of the velocity of the projectile in pitch and yaw, respectively. As is shown in Figure 1, the VRF can be obtained by rotating the NRRF angle of attack (AOA) a along O axis, and then rotating angle of sideslip along OY2 axis. According to the definition of , and the conversion relation of VRF and NRRF, , can be computed by equation (4), and the transfer matrix between the VRF and NRRF can be written as shown in equation (5). sin \u00bc cos v sin sin v cos v cos\u00f0 v\u00de sin \u00bc cos sin\u00f0 v\u00de= cos \u00f04\u00de ANV \u00bc cos cos cos sin sin sin cos 0 sin cos sin sin cos 2 64 3 75 \u00f05\u00de ~vr \u00bc vrx2 vry2 vrz2 \u00bc v wx2 wy2 wz2 \u00f06\u00de Taking wx2,wy2,wz2 as the components of the wind in the VRF, the relative velocity vector to the wind in VRF can be calculated by equation (6)", + " CD drag force coefficient for the whole projectile CL lift force derivative coefficient for the whole projectile ClpA spin damping moment coefficient for the aft body CMp Magnus moment coefficient for the whole projectile CMq damping moment coefficient for the whole projectile CM static moment derivative coefficient for the whole projectile CM _ pitch damping moment coefficient due to the change rate of the angle of attack C cN normal control force derivative coefficient for steering canards CNp Magnus force coefficient for the whole projectile d reference length e natural base Fc normal control force due to the steering canards g gravity acceleration i imaginary unit Ix axial moment of inertia Iy transverse moment of inertia of the projectile lCG distance between the point of mass and the acting point of the control force due to canards m,mA,mF mass of the total projectile, aft, and forward body Mc control moment due to steering canards p rolling speed of the aft and forward body q, r pitch and yaw rates expressed in the no-roll reference frame s trajectory arc length S reference area v velocity magnitude of projectile composite mass center wx,wz range wind and crosswind wx2,wy2,wz2 wind speed expressed in the velocity reference frame w wind vertical to velocity in the com- plex form defined as wy2 \u00fe iwz2 , angle of attack and sideslip r, r relative angle of attack and sideslip to the wind CE magnitude of the total equivalent deflection angle roll angle E direction angle of the equivalent canard model air density L direction of fire , Euler pitch and yaw angles v, v pitch and yaw angle of the velocity vector complex angle of attack defined as \u00fe i A aft body f fast arm of epicyclical motion F forward body s slow arm of epicyclical motion Appendix 1 According to the definition and coordinate transforming relations shown in Figure 1, it can be approximately expressed that \u00bc \u00fe , where the complex axis-direction vector is recorded as \u00bc \u00fe i ; the complex AOA is recorded as \u00bc \u00fe i ; the complex velocity vector is recorded as \u00bc v \u00fe i v. The derivative of equation (16) with respect to time can be expressed as \u20ac \u00bc \u20ac v \u00fe i \u20ac v i wv 1bz\u00f0 _pA pA _vv 1\u00de \u00fe \u00bd\u00f0by \u00fe bC\u00de _v\u00fe ibz _pA \u20ac i \u00fe _ \u00bd\u00f0by \u00fe bC\u00de\u00f0v wx2\u00de \u00fe ibzpA \u00fe bC CEvre i E\u00f02 _vr \u00fe vri _ E\u00de \u00f037\u00de As the parameters corresponding to aerodynamic terms (such as bx, by, bz, kzz) are small quantities (usually less than 0" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0000316_icems.2019.8921951-Figure10-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0000316_icems.2019.8921951-Figure10-1.png", + "caption": "Fig. 10 Third-order natural frequencies of rotor system", + "texts": [], + "surrounding_texts": [ + "In the design of high-speed motors, the rotor's operating speed needs to avoid the first-order critical speed. When the working speed is close to the first-order critical speed, in order to prevent the resonance phenomenon, it can be considered to change the critical speed of the motor rotor by changing the rotor structure. Therefore, it is necessary to analyze the influence of the change of the rotor system structure on the critical speed. The rotor system structure is divided into several parts as shown in Fig. 12, where La is the bearing mating length of the motor. When the motor bearing is selected, La is basically determined; Lb is the non-effective length inside the motor, and the end of the winding The length is related to the design of the ventilation structure; Lc is the electromagnetic effective length of the motor. After determining the electromagnetic scheme of the motor, Lc is basically determined; Ld is the rotor shaft elongation. Therefore, this section mainly studies the influence of the axial elongation Ld of the rotor system on the critical speed. Figure 13 shows the effect of shaft elongation on the critical speed of the rotor system. It can be seen from the figure that when the shaft elongation Ld is less than 120mm, the first-order critical speed of the rotor system is less affected. When the shaft elongation is greater than 120mm, the first-order critical speed of the rotor decreases more obviously, when the rotor shaft extends. The lower the first- order critical speed, the lower the rotor stiffness, and the rotor of the motor may develop from a rigid rotor to a flexible rotor. Therefore, the shaft should be prevented from extending too long to prevent resonance of the rotor system. IV. CONCLUSION In this paper, a 120kW, 24000r/min high-speed permanent magnet synchronous motor is taken as the research object. The rotor strength analysis and critical speed of high-speed permanent magnet motor are calculated and analyzed. The following conclusions are drawn: the key factors affecting rotor vibration mainly include rotor stiffness and rotor. In the two aspects of critical speed, the maximum deformation of the rotor is calculated based on the analytical method and the finite element method. The bending stiffness of the rotor is less than 10% of the air gap value of the motor, which can meet the rigidity requirement. The first-order critical speed of the rotor system is much larger than that of the rotor. The working speed will not cause resonance, which can ensure the safe and stable operation of the rotor of the motor; the first-order critical speed of the rotor system decreases with the increase of the rotor shaft extension, and the excessive extension of the shaft should be avoided to make the rigid rotor develop into a flexible rotor. V. ACKNOWLEDGMENTS This work was supported in part by the National Natural Science Foundation of China (nos. 51777048 and 51407050) and the University Nursing Program for Young Scholars with Creative Talents in Heilongjiang Province (UNPYSCT2016163). The authors would like to thank the anonymous reviewers for their valuable comments and suggestions that strengthened this paper. VI. REFERENCES [1] Tian Dong. Analysis of dynamic characteristics of high speed permanent magnet motor rotor [D]. Shenyang University of Technology, 2016. [2] Chen Yan. Vibration Processing and Cause Analysis of High Speed Motor[J]. Motor Technology, 2010, (06): 37-40. [3] S. Jumayev, M. Merdzan, KO Boynov, et al. The Effect of PWM on Rotor Eddy-Current Losses in High-Speed Permanent Magnet Machines [J]. IEEE Transactions on Magnetics, 2015, 51(11): 1-4. [4] Kim S, Kim Y, Lee G, et al. A Novel Rotor Configuration and Experimental Verification of Interior PM Synchronous Motor for High-Speed Applications [J]. IEEE Transactions on Magnetics, 2012, 48(2): 843- 846. [5] HNPhyu, NLHAung and J.Quan. Investigation of the Effect of Winding Structure and MMF Harmonics on the Rotor Eddy Current Loss of High Speed Permanent Magnet Motor[C]. 2016 IEEE Region 10 Conference (TENCON), Singapore, 2016: 204-209. [6] [S. Neethu, S. Pal, AK Wankhede, et al. High-Performance Axial Flux Permanent Magnet Synchronous Motor for High Speed Applications [C]. IECON 2017 43rd Annual Conference of the IEEE Industrial Electronics Society, Beijing, China, 2017: 5093-5098. [7] X. Song, J. Fang, B. Han. High-Precision Rotor Position Detection for High-Speed Surface PMSM Drive Based on Linear Hall-Effect Sensors [J]. IEEE Transactions on Power Electronics, 2016, 31(7 ): 4720-4731." + ] + }, + { + "image_filename": "designv11_71_0002161_icem49940.2020.9270955-Figure10-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0002161_icem49940.2020.9270955-Figure10-1.png", + "caption": "Fig. 10. Prototype and control board. (a) Prototype. (b) Control board.", + "texts": [ + " Especially under non-adjacent opencircuit conditions, the advantage of elliptical MMF is more prominent. Therefore, under the open-circuit conditions of five-phase PMSM, the torque performance of elliptical MMF is better than that of the traditional reconstructed round MMF. Authorized licensed use limited to: Auckland University of Technology. Downloaded on December 21,2020 at 15:26:05 UTC from IEEE Xplore. Restrictions apply. To verify the proposed approach, a five-phase PMSM and its control board was designed and manufactured. The prototype and control board are shown in Fig. 10. The prototype is a 20-slot/18-pole machine with open-end windings and surface-mounted PMs in Fig. 10(a), and the parameters are listed in Table II. The control board is shown in Fig. 10(b), which is composed of power supply unit, speed and torque signal processing unit, voltage and current sampling unit, main control unit and driving unit. The main control unit adopts TMS320F28335 minimum system, which outputs PWM control signals. The current sampling unit uses an integrated current Hall sensor like ACS724 of Allegro to convert the current signal into linear voltage signal. The speed and torque signal processing unit adopts the high-speed optocoupler 6N136 made by Toshiba Corporation of Japan to convert 5V to 3" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0003895_icems.2014.7013984-Figure1-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0003895_icems.2014.7013984-Figure1-1.png", + "caption": "Fig. 1. The overall structure of the PSRM", + "texts": [ + " In Section III, three methods of thrust force characteristics measurement for the PSRM are discussed. Experiments are implemented, and results are analyzed in Section IV. Conclusion remarks are presented in Section V. As one kind of the direct drive planar motors, PSRMs are regarded as the combination of two linear switched reluctance motors along the X direction and Y direction, and each linear switched reluctance motor is based on the \u2018straightened-out\u2019 version of a 6/4-pole SRMs [1]. The overall structure and the specifications of the PSRM are shown in Fig. 1 and Table I, respectively. The PSRM is composed of X moving platform, Y moving platform, stator base, stator sets, linear guides, linear encoders, and so on. Each moving platform consists of two sets of three movers, and each set of three movers is responsible for each direction. One mover is composed of a set of laminated silicon steel blocks and a winding. Structure of the X moving platform is presented in Fig. 2. The stator sets are a mesh structure which is constructed from small pieces of laminated steel blocks, as shown in Fig" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0002338_1369433220971728-Figure1-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0002338_1369433220971728-Figure1-1.png", + "caption": "Figure 1. Deployable structure consists of two SLEs with clearance in joint A.", + "texts": [ + " In short, the dynamic behavior of deployable structures with clearance is very sensitive to the joint conditions, even slight change to the clearance size or friction coefficient will lead to a transition of the mechanism\u2019s response from regular to irregular or vice versa. The necessity to control the dynamic response of a mechanism with clearance depends on the specific design and requirements. Modeling of joint clearance and contact force in planar deployable structures Kinematic description of scissor deployable structures with joint clearance The linear array deployable structure consists of many substructures with the same geometrical properties, and each sub structure consists of a pair of bars which are connected by a pivot at the midpoint. Figure 1 shows the deployable structure composed of two scissor units, where joint A is considered as the imperfect joint and Mxy is the global coordinate system. Bar 1 and bar 2 can rotate around the pin B, and similarly, bars 3 and 4 rotate with each other at the pin E. Bar 1 is hinged at the fixed point M and ui(i = 1, 2, 3, 4) represents the angle between bar i and axis x in the global coordinate system. Hinges o1 and o2 (similarly, o3 and o4) are the midpoint of bars 1 and 2, respectively. RB and RJ are used to indicate the radius of the bearing and journal, respectively, and c=RB RJ is the clearance value, as shown in the upper left of Figure 1. The journal and bearing can move relative to each other unconstrained due to the revolute clearance, which leads to randomness of the system. According to the size of the clearance circle, the motion of the shaft in the bearing can be divided into three different modes: the free flight, the impact, and the contact mode, as shown in the lower left of Figure 1. The clearance exists in the real joint to ensure the relative movement between the connecting components, which is also the source of the impact forces resulting the friction and wear, uncertainty in motion position, and degradation of the system performance. A fundamental requirement of dynamic analysis of a deployable structure with clearance is to accurately determine whether the bearing boundaries and journal are in contact with each other. In Figure 2, points Pb and Pj represent the centers of bearing and journal, respectively, and ebj denotes the eccentricity vector connecting the centers of bearing and journal", + " In the case that bar 1 in the deployable structure rotates with constant angular velocity, equations (28) and (29) are used to control the system for continuous contact, and equation (27) is used to obtain the required input torque for maintaining the desired motion. In this section, the deployable structure consisting of two SLEs is used as an illustrative example to demonstrate how a revolute clearance joint affect the dynamics of a mechanism and validate the control procedure presented in the preceding sections, where joint A is a real joint, as depicted in Figure 1. In Figure 1, each bar has the same length and mass, 2 m and 5.68 kg, respectively, and the mass of the slide is 5 kg. The initial deployment angle u1 is set as 30 , and the bar 1 rotates at a constant angular speed of 150 rpm. The journal and bearing centers concentric at the beginning of the simulation, and the initial conditions of the simulation can be obtained from the movement of the ideal mechanism. Also, the numerical parameters used in dynamic solution are listed in Table 2. The steps of the computational methodology for dynamic analysis in planar deployable structure with clearance and friction are shown as follows: (1) Define the initial configuration of the system and the initial conditions for the simulation" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0002338_1369433220971728-Figure4-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0002338_1369433220971728-Figure4-1.png", + "caption": "Figure 4. Local deformation of deployable structure during impact.", + "texts": [ + " For the general clearance joint, the positional relationship between potential contact points should meet the following geometric constraint conditions: d= r Q b r Q j , \u00f05\u00de nj 3 nb = 0 d 3 nb = 0 , \u00f06\u00de where r Q b and r Q j are the position vectors of potential contact points in the global coordinate systems, nj and nb are the normal vectors of the possible contact points (Li et al., 2018). When the potential contact points are determined, then the contact penetration and relative contact velocity can be obtained. Figure 4 shows the position relationship between the bearing and the shaft during impact. For convenience of display, the second unit is not drawn, where the contact region is indicated in blue. The impact points are denoted as Qb and Qj, their position vectors in the global coordinates are represented as rQb and r Q j . _rQb and _rQj are the velocity of the contact points, which can be obtained by differentiating the position vectors with respect to time: r Q b = rb +RBn r Q j = rj +Ajs P j +RJn ( \u00f07\u00de _rQb = _rb +RB _n _rQj = _rj + _Ajs P j +Rj _n _n= _at 8< : \u00f08\u00de where rb and rj are the geometric centers of joint A and bar 2, sP j is the position vector of the center of the journal in the body-fixed coordinate o1x1y1" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0002989_978-3-642-16214-5_9-Figure9.47-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0002989_978-3-642-16214-5_9-Figure9.47-1.png", + "caption": "Fig. 9.47. Parking lock with radially acting locking pawl (see also Figure 12.22). 1 Gearshift lever on the transmission; 2 axis selector shaft; 3 detent plate (P park position, R reverse gear position, N neutral position, D drive position); 4 detent spring; 5 compression spring; 6 connecting rod; 7 output shaft; 8 parking lock wheel; 9 return spring; 10 pawl; 11 locking roller; 12 guide; 13 roller bearing", + "texts": [ + " A parking lock mechanically locks the transmission output shaft against the transmission housing. In the case of commercial vehicles and buses with spring actuator parking brakes, the parking lock is not utilised as an additional element to keep the vehicle stationary. In this case, the transmission selector lever lacks a P position. The basic principle of parking locks is the radially acting locking pawl. The details of its actual design vary. With respect to the locking elements, we differentiate between: \u2022 systems with a locking roller (Figure 9.47) and \u2022 systems with a locking cone (Figure 9.49). Concerning activation we distinguish between: \u2022 mechanical designs (Figure 9.47) and \u2022 electric designs, usually in connection with e-shifting (Figure 9.49). In mechanical designs, the parking lock is activated by means of a Bowden cable connecting the selector lever unit in the vehicle to the transmission, while in electric designs this is done, for example, with button activation. In the case of mechanical activation, the locking action is initiated by the driver moving the selector lever to engage the park position. The transmission shown in Figure 12.22 will be used as an example to illustrate the design and function of a mechanically activated parking lock of a conventional automatic transmission. The parking lock shown in Figure 9.47 has a radial locking pawl. Moving the selector lever 1 to the park position P has the following effects: 1/ The detent plate 3 rotates about the axis 2 of the selector shaft in the same direction as the gearshift lever, until the roller on the detent spring 4 engages in the park position P. 2/ The connecting rod 6 linked to the detent plate, moves the roller 11 (which runs on it) in the guide 12 parallel to the output shaft 7. 3/ At the end of the guide, the roller runs over a roller bearing 13 fixed in the housing, pressing upwards against the sloping back of the pawl 10", + " With a trailer, the requirements must be fulfilled for a 12% slope. Fulfilment of the disengagement requirement/actuating force: for heavy vehicles and corresponding ratio conditions between the tyres and the parking lock wheel, pressure between the locking roller (or locking cone) and the pawl can be very high on a 30% slope. The geometry and choice of material or coating must guarantee sustainable disengagement forces even under these conditions. Related to the axis of the selector shaft (2 in Figure 9.47), typical values lie between 10 and 20 Nm. The formation of frictional martensite due to excessive pressure between the parking roller (cone) and the pawl should be avoided. Fulfilment of operational safety: the parking lock should be considered from the standpoint of comfort, but even more so from that of safety. The interaction of the geometry of the cams on the detent plate (Roostercomb) and the hardness of the detent spring contributes heavily to shifting feel. The friction in the entire system \u2013 from the selector lever unit in the vehicle, through the Bowden cables, to the detent plate and the active roller of the detent spring \u2013 must be selected such that only defined conditions are possible" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0000757_mepcon47431.2019.9007981-Figure2-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0000757_mepcon47431.2019.9007981-Figure2-1.png", + "caption": "FIGURE 2. System architecture of the proposed PS-UA scheme in PS LTE system.", + "texts": [ + " That is, increasing bias value leads to the CRE for the corresponding BS j, and therefore, the offloading of more users to the corresponding tier. In this paper, we also discuss which bias factor will be suitable in the sense of mPC load balancing and PS prioritybased user association. 9778 VOLUME 4, 2016 B. eICIC ABS RATIO MAXIMIZATION The mPC is moving to provide a high data rate to the users connected to it. In order to offload more users to an mPC, bias is introduced. Although, by increasing bias value more users are associated with the mPC as shown in Fig. 2 by grey circle for which the radius of the grey circle increases as the bias value increases. This will also result in more interference with the offloaded users located in the CRE because of getting high power from the neighbor eNB. Thus, in order to reduce this interference, eICIC was introduced by third-generation partnership project (3GPP) in Rel. 10, by sending an ABS from eNB j during the highly interfering RBs. In this paper, our objective is to find the optimum value \u03c5\u2217j of ABS ratio \u03c5 = n/TABS that can associate a higher number of PS users with the mPC based on their connection priority for load balancing" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0001691_j.oceaneng.2020.107700-Figure1-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0001691_j.oceaneng.2020.107700-Figure1-1.png", + "caption": "Fig. 1. The schematic diagram of a modularized floating structure.", + "texts": [ + " It greatly simplifies the arrangement of thrusters and the control scheme, so that one can realize the design concept of using as fewer thrusters as possible and control as more motions as possible. Of course, the idea also can save a great deal of investment in engineering applications because of the simple layout. This paper is organized as follow. In the next section, the mathematical model of the multi-modular floating structure with vertical propellers is derived. Then the control process is derived to get the thrusts of the propellers. Afterwards, numerical simulations are given to verify the effectiveness of the control method. At last, a conclusion is drawn. As shown in Fig. 1, the modularized floating structure consists of N semi-submersible modules which are connected by hinged connectors. Between two adjacent modules, two hinged connectors are mounted. An anchor chain is installed on each module to prevent the whole floating structure from drifting. Global coordinate system (X,Y, Z) is set of which X \u2212 Y plane is on the still water surface and Z axis is perpendicularly upwards. A local coordinate system (xi, yi, zi,\u03b1i, \u03b2i, \u03b3i) is set at the mass center of the ith module", + " Symbol \u0394\u03c9 is a frequency step. In the right side of Eq. (1), vector fh is the hinged connector force vector that acts on the mass centers of modules of the floating structure. To get the hinged connector force fh, we need firstly to derived the hinged connector load fh that is the internal forces of hinge connector. The hinged connector is regarded as a rigid constraint of the two adjacent modules in some degrees of freedom. In this paper, there are two hinged connectors deployed between two modules as shown in Fig. 1, which limits the related surge, sway, heave, roll and yaw motions, and releases pitch motion. However, the arrangement of the hinged connectors makes the floating structure existing over constrained in x, y, z, \u03b1 and \u03b3 directions. For simplify, we suppose that connector 1 shown in Fig. 1 constrains the related movements in x, y, z directions and connector 2 only constrains the related movements in x, z directions, making the two adjacent modules only exist related pitch motion. According to the virtual work principle, the constraint equation of the ith and (i+1)th module is Dk [ xT i , x T i+1 ]T = 0 (5) where the displacement constraint matrix Dk can be written as Dk = [ Dk,A,Dk,B ] (6) where Matrix Dk is the constraint relationship between the kth and (k+1)th module, where vectors ck,c0 = [c1 k,c0, c2 k,c0, c3 k,c0] T and ck,j,h = [cx k,j,h, c y k,j,h, cz k,j,h] T are respectively the locations of the mass center of the kth module and the center of the jth joint connector (two joint connectors between two adjacent modules) in the global frame under static state" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0001292_6.2020-3042-Figure3-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0001292_6.2020-3042-Figure3-1.png", + "caption": "Fig. 3 Transition from typical to inverted von K\u00e1rm\u00e1n vortex street.", + "texts": [ + " 2 Inverted von K\u00e1rm\u00e1n vortex street (velocity profile indicates a momentum surplus). Another regime studied by Lai and Platzer [14], Lewin and Haj-Hariri [15], and Young [16] was the vortex-pair shedding that represents the transition from the drag producing wake to the thrust producing inverted von K\u00e1rm\u00e1n 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 Ju ne 2 8, 2 02 0 | h ttp :// ar c. ai aa .o rg | D O I: 1 0. 25 14 /6 .2 02 0- 30 42 vortex street. Young and Lai [17] concluded that this type of wake structure, represented in Figure 3, was caused by the interaction between bluff-body type natural shedding from the trailing edge and the motion of the airfoil. An additional type of vortex shedding characterized by an asymmetrical vortex shedding appears when the airfoil is subjected to substantial reduced frequencies and plunge amplitudes, in other words, relatively high Strouhal numbers. Bratt [18] identified this type of vortex shedding. However it was only Jones et al. [19] who meticulously studied this phenomenon experimentally and numerically, concluding that its genesis was due to the closeness of the shed vortices and that it was mostly an inviscid occurrence" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0002137_b978-0-12-821841-9.00009-8-Figure5.5-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0002137_b978-0-12-821841-9.00009-8-Figure5.5-1.png", + "caption": "FIGURE 5.5", + "texts": [ + " After the waste is blended, it will be tested for physical, chemical, and biological characteristics. Most of the testing is carried out by using Spectrophotometer DR/4000 that is based on SMEWW, which is shown in Table 5.6. There are eight parameters tested in this study. Those parameters are pH, BOD, COD, oil and grease, temperature, nitrogen, and phosphorus. The electrodes that have been used in this MFC are made from graphite. The graphite is cut into two sizes, 603 60 and 1103 50, which are shown in Fig. 5.5. Then, two holes are drilled on the graphite plate. This hole is used to insert two Flowchart of experimental design. Heads, skins, tails, and intestinal organs of fish. The fish waste was blended and weighed according to the composition. stainless steel screw holders so that the electrode is in a hanging position. Then, the electrode is tied to the container lid using a bolt and nut. The results gathered from the experiments are analyzed by using Statistical Package for the Social Science (SPSS) Statistic 17" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0003929_scis-isis.2014.7044649-Figure6-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0003929_scis-isis.2014.7044649-Figure6-1.png", + "caption": "Figure 6. Equipment for varying angles", + "texts": [ + " The riding robot makes the forwarding velocity ms increase or decrease incrementally in eight steps to protect the safety of the rider on starting and stopping. IV. EXPERIMENT Angles The control of the riding robot using equipment for varying angles allows the riding robot to migrate by changing the direction of the smartphone within 30\u00b0 from the left and right with respect to the heading angle of the riding robot. In the experiment, the equipment for varying angles consists of a servo motor RX-10 (ROBOTIS) and a controller board CM700 (ROBOTIS), which are mounted on the riding robot and a smartphone holder as shown in Fig. 6. The smartphone communicates with the riding robot through Bluetooth wireless communication. The riding robot migrates based on the heading angles of the smartphone transmitted from the riding robot. The heading angle of the smartphone, obtained using equipment for varying angles, rotates from the heading angle of the riding robot ( to ) in 30 at an angular velocity of 26.6 /s. The heading angle of the smartphone keeps the heading angle at for 2 s, rotates it from to in 60 at an angular velocity of 26" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0002171_icem49940.2020.9270890-Figure1-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0002171_icem49940.2020.9270890-Figure1-1.png", + "caption": "Fig. 1: 2D Structure of PM Synchronous Motor.", + "texts": [ + " Finally, the paper is concluded in Section IV. An 8-pole surface-mounted PMSM with concentrated winding is chosen as the case study for implementation of simultaneous ITSC faults in different phase windings. The reason that a concentrated winding structure is considered here, is simply because it is easier for this modeling to be implemented on. The same procedure can be applied on a motor with distributed winding with a few modifications in the flux equations. A 2D view of this PMSM is depicted in Fig. 1 while motor parameters are listed in Table I. Some of these parameters are obtained from the manufacturer datasheet and the rest from a few FEM Simulations which are explained in [15]. Each phase winding consists of 4 (number of pole-pairs) coils denoted by a1,2,3,4, b1,2,3,4, and c1,2,3,4, which are connected in series. It is assumed that an ITSC fault is present in the first coils of each of the three-phase windings, splitting the coil into one faulty part and one healthy part. The fault severity \u00b5a,b,c is defined as the ratio of number of the shorted turns to the total number of turns per coil" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0003479_mipro.2015.7160432-Figure2-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0003479_mipro.2015.7160432-Figure2-1.png", + "caption": "Fig. 2. Body fixed frame with aerodynamic angles, pitch rates and moments", + "texts": [ + " 1). Secondary control surfaces are intended to relieve excessive control load and include trim and spring tabs. Auxiliary control surfaces improve the aircraft performance characteristics and include flaps, spoilers, speed brakes and slats. These control surfaces can be controlled by a pilot or by an autopilot as given in Fig. 1. The aim of all mentioned control surfaces is to control the airplane height, speed (V), roll (p), pitch (q) and yaw (r) rates, and moments (L, M and N) as presented in Fig. 2. On a modern high speed aircraft, the pilot\u2019s control joystick and pedals are not directly linked with the control surfaces. Appropriate systems are in between and augment the pilot control commands applied on the joystick and pedals. Large aircrafts have systems that convert pilot control commands into electrical signals. They are sent via electrical or optical cables to various actuators which can be electromechanical, pneumatical, electrical or hydraulic devices. Actuators then move the appropriate control surfaces to the desired direction" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0001057_0021998320920920-Figure6-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0001057_0021998320920920-Figure6-1.png", + "caption": "Figure 6. Contours of the mode corresponding to Hmax for binder in unit-cell 2 with warps and wefts shown for context.", + "texts": [ + " The maximum failure index provides a metric of severity of the stress state. The regions of severe stresses tended to form near two types of features of the tow architecture. First, the most severe stresses in the binder developed where the binder transitions between travelling along the x-axis and travelling through the thickness, such as label A in Figure 5. In these regions where the tow path of the binder transitions and the most severe stresses occur, transverse tension is the mode corresponding to the maximum failure index, as shown in Figure 6. The high failure index for transverse tension is due to a severe 0xz that occurs. Where the binders travel along the x-axis, the binders carry a significant amount of load, since the fibers align with the applied load. This can be illustrated by contours of the stress in the direction of the applied load, xx, which is in the global coordinate system, as shown in Figure 7. Label A in the figure demonstrates where load from the binder is transferred to the nearby wefts when the binder transitions between travelling along the x-axis and through the thickness", + " It should also be noted that the most severe 0xz occurs within the volume of the binder, not on the surface, as seen in the cutaway of the binder in Figure 8. The high HTT in these regions are dominated by the large magnitude of 0xz. The regions with the second most severe stresses in the binders occur where z-aligned binders come close to wefts, such as indicated by label B in Figure 5. For all of these regions, the failure index corresponding to transverse tension, HTT, is highest, as shown in Figure 6. The stress component that contributes the most to HTT where the binders traverse the textile thickness is 0zz. Figure 9 shows 0zz in the binder, and the locations of 0zz concentrations can be seen to match the locations of local maxima of Hmax, shown in Figure 5. When the binders are most closely aligned with the z-axis, the local z-axis of the binders most closely aligns with the global x-axis, which is the direction of load. Consequently, the 0zz concentrations form where load is transferred from the binders to nearby wefts" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0002861_978-3-030-55889-5-Figure2.14-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0002861_978-3-030-55889-5-Figure2.14-1.png", + "caption": "Fig. 2.14 Rules for pole-plans", + "texts": [ + " An \u201cacademic\u201d proof of this principle is the following: \u2022 If E I or E A change in a member, the associated forces f + are equilibrium forces, see Chap. 5. \u2022 The forces f + are orthogonal to all rigid-body motions \u2022 The influence functions for N , M, V are rigid-body motions The consequence is: The internal forces 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. The plan of these poles we call pole-plan. The following rules apply to the construction of pole-plans, see Fig. 2.14: line: (i) \u2212 (i, j) \u2212 ( j), e.g.: (1) \u2212 (1, 2) \u2212 (2). 6. The secondary poles (i, j), ( j, k), (i, k) of three segments I, J, K lie on one line: (i, j) \u2212 ( j, k) \u2212 (i, k), e.g.: (1, 3) \u2212 (1, 4) \u2212 (3, 4). All the other poles can be found with the help of so-called trace lines. A trace line or locus is the straight line on which the pole must lie according to rules #3 to #6. \u2022 The intersection of two trace lines of the same pole marks the location of the pole. \u2022 If two trace lines run parallel and aim at the same pole, they intersect at infinity" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0001064_iccais46528.2019.9074699-Figure3-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0001064_iccais46528.2019.9074699-Figure3-1.png", + "caption": "Fig. 3: Measuring system for the PKM", + "texts": [ + " By taking the objective function as the square of the output error between Authorized licensed use limited to: UNIVERSITY OF BIRMINGHAM. Downloaded on May 10,2020 at 05:32:41 UTC from IEEE Xplore. Restrictions apply. the real plant and the mathematical model subject to the same desired path, the nonlinear optimization technique can be used to derive the dynamic parameters of the PKM in the CLIM [12]. To avoid solving forward kinematics numerically, The pose of \u03a3E can be measured by the photogrammetry sensor which is the output of the PKM. Accordingly, the forward kinematics of the PKM can be eliminated. As shown in Fig. 3, the photogrammetry sensor is a dual camera-based visual sensor. The position coordinates of magnet reflectors in the field of view of photogrammetry sensor can be measured. By attaching the reflectors on the end-effector and the base platform respectively, the pose of \u03a3E w.r.t. \u03a3O could be measured. The transformation matrix from the \u03a3E to \u03a3O, OTE , can be derived from Eq. 5. OTE = CT\u22121 O CTE , (5) where CTO and CTE are the homogeneous transformation matrices from the sensor frame to the base frame and from the sensor frame to the end-effector frame respectively" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0000315_icems.2019.8921590-Figure10-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0000315_icems.2019.8921590-Figure10-1.png", + "caption": "FIGURE 10. (a) Photograph of the assembled OMT with polarizations. (b) Photograph of the OMT before and after the assembly.", + "texts": [ + " In order to keep a finite thickness during the manufacturing process, a cusp at the intersection (Fig. 9) of the two-miter bends is truncated at a width of 0.1 mm. In order to be able to machine the 4-section transformer and the 3-miter bends on a cut length of 6.33 mm (represents the full width of the band 1 waveguide), a standard 2 mm-diameter end-mill was used. III. EXPERIMENTAL VERIFICATION OF THE OMT The structure of the OMT presented in Fig. 4 is realized by superimposing three aluminum blocks made of Aluminum 6061 with a T6 temper (see Fig. 10(a) and 10(b)). The three parts are machined using a CNC milling machine and two dowel pins are used for alignment. The machining required for the manufacturing of this OMT is standard for a CNC machine with tolerance within reach of a production-type workshop. The dowel pins and the features of the turnstile and bends have +/\u22120.01 mm tolerances which is the tightest tolerance in the OMT. The rest of the features have +/\u22120.02 mm and +/\u22120.1 mm tolerances. The RF design is optimized so that a high level of tolerance is not required in the manufacturing" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0001137_012001-Figure1-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0001137_012001-Figure1-1.png", + "caption": "Figure 1. Experiment setup", + "texts": [ + " Kulkarni and Wadkar[7] investigated the distributed defects of the ball bearing on the vibration response. The level of vibration varies with the increase of surface deformation. When the speed increase, the vibration level reduced. As the load increases, the surface deformation and the vibration level increase. Machado et al.[8] investigated the wear of journal bearing on rotating system in time domain. In this study, a test rig for measuring rolling contact surface deformation is used as shown in Figure 1. The rig is consisting of roller specimens, stepper motor, linear guide, and four ball bearings. The test rig is put directly under Alicona Infinite Focus Microscope (IFM) therefore the surface scanning can be carried out directly. Load is applied to top roller. During surface characterization, the load is removed and there is an access hole to allow the microscope to scan the surface of the specimen. Both roller specimen is supported by two ball-bearings, but the top roller is free to move in the vertical direction, and the movement in the horizontal direction is limited by linear guides and the holders" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0003921_coase.2014.6899477-Figure13-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0003921_coase.2014.6899477-Figure13-1.png", + "caption": "Figure 13. Experimental setup.", + "texts": [ + " We developed a cylinder-shaped robot to verify the effectiveness of the new safety device. Fig.12 shows the developed robot. The diameter is 0.6[m] and the height is 0.7[m]. The mass is approximately 25[kg]. The robot has two wheels, two motors, one safety device, and two casters. Each wheel is controlled by each motor. The safety device is connected to Shaft A of the right wheel via a timing belt. We conducted the following experiments by using the developed robot. A. Contact Force-based Detection Mechanism Fig. 13 (a) shows the experimental setup. We measured the contact force between the robot and a force sensor (Mini45 SI-580-20, ATI Industrial Automation), while the robot moved and collided with the force sensor at a low speed (0.05[m/s] (= 0.5[rad/s] at Shaft A)). The sampling frequency was 100[Hz]. We experimented by using the detection velocity level of 3.84 [rad/s] and the detection contact force level of 22.2 [N] as an example to check whether the safety device achieves the function. Five trials were conducted", + " The differences between the detection contact force level and the experimental values are attributed to the output force errors of Linear Springs Y, friction in the safety device, among others. B. Velocity-based Detection Mechanism Next, for safety reasons, we mounted the robot on a spacer and attached some markers on Gear A (i.e. Shaft A), Claw A, and Plate A of the safety device. Then, we measured the velocity of Shaft A, the motion of Claw A, and the motion of Plate A by using a motion capture system (HAS-500, DITECT Corporation) while increasing the velocity of Shaft A by the motor (Motor 1) (see Fig. 13(b)). The sampling frequency of the motion capture system was 100 [Hz]. The detection velocity level was 3.84 [rad/s]. Five trials were conducted. Fig. 16 shows the typical example of the experimental results. In Fig. 16, the time when Claw A locked Plate A is indicated by an arrow. Fig. 16 indicates that the velocity of Shaft A was approximately the detection velocity level at the time when Claw A locked Plate A. Furthermore, Fig. 16 indicates that the velocity of Shaft A ultimately became zero" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0003558_ijahuc.2014.064425-Figure5-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0003558_ijahuc.2014.064425-Figure5-1.png", + "caption": "Figure 5 The overlapping of two readers: (a) deployment of two readers and (b) grid graph for (a) (see online version for colours)", + "texts": [ + " The value of each element in the product matrix MD \u22c5 MC could be either 0 or 1. If the value is 0, the corresponding grid is not covered by the reader. If the value is 1, the corresponding grid is within the coverage area, and is covered by the reader. Suppose that the reader matrices for readers 1 and 2 are 1 2 and ,r rM M respectively. When two readers 1 and 2 are deployed, the deployment matrix MD is the summation of matrices 1 2 and .r rM M Then, we can find the effective coverage area by multiplying matrices MC and MD. From Figure 5(b), we can see the effective coverage area is the double-coloured area. From Figure 6, we can see that 28 elements have a value of 1. This means that 28 grids are covered by one RFID reader (either reader 1 or 2). We can also see that 12 elements have a value of 2. That indicates that 12 grids are covered by two RFID readers (both readers 1 and 2). There are two important factor, coverage and overlapping rates, to evaluate the effectiveness of the RFID reader deployment. The coverage rate denotes the rate of total covered area to the whole deployment area" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0000873_compumag45669.2019.9032824-Figure1-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0000873_compumag45669.2019.9032824-Figure1-1.png", + "caption": "Fig. 1. Concept of homogenization and implementation of voltage constraint", + "texts": [ + " Surprisingly, none of the cited works have ever utilized the special spiral-like rolled geometry of coils. That is, homogenization has always been made as if the turns were separate (closed) current carrying loops (this is called the nested-loop approximation [10]). In the H-formulation, for instance, the conductor current is weakly prescribed as an integral constraint, and the coil voltage is evaluated afterwards via an averaging in the coil\u2019s cross-section [2], [4]. The aim of this work is to introduce a new homogenization technique for rolled conductors (Fig. 1). It is developed with HTS tape coils in mind, but foil transformer and battery designers may benefit from it, too. Beyond the speed-up of simulations, the proposed method makes coil voltage available in the A-V formulation in a simple manner. Moreover, this homogenization technique may be extended to the thermal analysis as well. This paper presents the theory, a feasibility study, and the first results of a test problem concerning the dc characterization of an HTS coil. We focus on the electromagnetic modeling of an HTS tape coil in the dc or \u201cslow transient\u201d regime", + " However, when the tape-level eddy currents have to be reconstructed too, e.g. for determining the ac losses, the methods presented in [14] or [15] may be subsequently used. Further, if some local phenomena (e.g. quench) has to be modeled with enhanced 2 accuracy, the hybrid type discretization of [8] can be applied. Here we confine ourselves to pancake coils, but the application of the method on other 3-D geometries is straightforward. A. Homogenized Coil Model The densely wound tape coil is substituted by a bulk medium (see Fig. 1) via the constitutive law, J = \u03c3\u0304E, (1) in which J denotes the current density, E the electric field intensity, and \u03c3\u0304 the homogenized conductivity tensor. Obviously, \u03c3\u0304 must be diagonal in the local frame (\u03be\u03b7z) that fits the spiral path of the tape, moreover it vanishes in the direction perpendicular to the tape, thus \u03c3\u03be\u03be = 0 (see Fig. 2). After transforming \u03c3\u0304 into the\u2014more familiar\u2014cylindrical frame (r\u03d5z), radial and off-diagonal elements appear, too. In the expressions of \u03c3\u0304 (Fig. 2), p denotes the pitch of the spiral, and \u03c3eq = f\u03c3s, where f is the filling factor and \u03c3s(E,B, T ) the electric conductivity of the HTS film", + " Faraday\u2019s law can be written for the latter as\u222e C\u222aC0 E \u00b7 dl = \u2212 d\u03a6 dt , (14) where \u03a6 is the total magnetic flux of S \u222a S0. Assuming that the flux of S0 is negligible to that of S, we can write (14) approximately as\u222b C [\u2212\u2202tA\u2212\u2207V ] \u00b7 dl\u2212 u \u2248 \u2212 d dt \u222b C A \u00b7 dl. (15) Canceling and applying the gradient theorem we arrive at U = V2 \u2212 V1. (16) Just like the tape, its terminals are not manifested in the homogenized model either. Therefore it seems natural to extend the constraint (16) to the whole inner and outer cylinder walls (cf. Fig. 1) as the boundary conditions (8)-(9) for V . Although voltage excitation is the simplest to realize, current I may also be prescribed via the (homogeneous) normal current density on \u03931 and \u03932, similar to (7). In this case one can easily obtain the resulting coil voltage as (16) afterwards. We study the HTS coil described in [2] so that we can compare the results. The inner/outer radii and height of the coil are r1 = 29.5 mm, r2 = 39.5 mm, h = 4.35 mm (this is equal to the tape width). The number of turns is N = 40, hence the pitch, p = 250 \u00b5m, and min\u2126c (p/r) \u2248 6" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0000642_s12206-020-0134-3-Figure12-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0000642_s12206-020-0134-3-Figure12-1.png", + "caption": "Fig. 12. Designed RRPR linkage for the first example.", + "texts": [ + " The first example involves a set composed of the first four Table 2. Given poses for the first example. Pose iX iY iq 0 0.0 0.0 0.0\u00ba 1 1.5 0.8 10.0\u00ba 2 1.6 1.5 20.0\u00ba 3 2.0 3.0 60.0\u00ba Fig. 8. Circle-point curve for the RR dyad for the first example. poses proposed by Chang et al. [9]. The parameters associated with each pose are shown in Table 2. According to the procedure shown in Table 1, there were obtained the graphical results shown in Figs. 8-11. The numerical results produced by the approach are shown in Table 3. Finally, Fig. 12 shows the resulting linkage assembled in pose 0. Evaluation of the performance A careful analysis of the obtained results allows to realize that: (1) All the locations reached by the coupler point remain on a single and continuous trajectory, see Fig. 13. (2) The coupler point remains on one single branch. (3) The sequence of occurrence of the locations reached by the coupler point is in the desired order. From the foregoing observations, one may conclude that example 1 is a successful design. This second example involves a set composed of the five poses proposed by Chen et al" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0003851_e2014-02130-2-Figure2-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0003851_e2014-02130-2-Figure2-1.png", + "caption": "Fig. 2. Pendulum described by: (a) spherical coordinates, (b) fixed spherical coordinates.", + "texts": [ + " These points are also equilibrium points. During motion of the pendulum, while reaching the equilibrium i.e. at the point \u03b8 = 0 and its neighbourhood the mass moment of inertia is equal to zero, hence numerical integration of this system is impossible. 2.1.2 Fixed spherical coordinates To solve the problem of a singular point in the equilibrum point, we change coordinate system to fixed spherical coordinates [22]. Angle \u03b2 describes rotation about axis x and angle \u03b1 is between pendulum and Y Z plane, as shown in Fig. 2. The cartesian coordinates are described as follow: xA = h sin\u03b1, yA = h cos\u03b1 sin\u03b2, zA = \u2212h cos\u03b1 cos\u03b2 and the kinetic energy is given in the following form: Ek = 1 2mh 2(cos2 \u03b1\u03b2\u03072 + \u03b1\u03072), Ep = mgh(1\u2212 cos\u03b1 cos\u03b2). The singular point is for \u03b1 = \u2212\u03c02 or \u03b1 = \u03c0 2 . As long as \u03b1 < \u03c0 2 numerical integration is possible. 2.2 Model of the beam Solid beam can be considered as a system of 3 points with mass distributed in the following manner: 16 at the ends and 2 3 in the middle [23]. Let us consider velocity of each point in the following form: V 2A = x\u0307 2 A + y\u0307 2 A + z\u0307 2 A, V 2 B = x\u0307 2 B + y\u0307 2 B + z\u0307 2 B, V 2 C = x\u03072C + y\u0307 2 C + z\u0307 2 C " + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0003544_cta.2127-Figure2-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0003544_cta.2127-Figure2-1.png", + "caption": "Figure 2. Wheatstone bridge transient state.", + "texts": [ + " The Wheatstone bridge reaches its steady state after the switch has been closed, and then R2, R3, and R4 are adjusted until V=0. Thus, a steady state can be obtained when the bridge is balanced by R1 R2 \u00bc R3 R4 . Copyright \u00a9 2015 John Wiley & Sons, Ltd. Int. J. Circ. Theor. Appl. 2016; 44:1094\u20131111 DOI: 10.1002/cta The transient state of the Wheatstone bridge occurs after opening the switch at t=0, and then a current denoted as i1 flows through the inductor L1. This transient current starts from its initially steady value I1 and then decays to I1 = 0. Figure 2 explains for this transient state. Hence, the inductor L1 yields a voltage source L1 di1 dt during its transient period, and the transient i1 keeps in the same direction as in the steady-state condition. The bridge voltage is therefore defined as V \u00bc L1 di1 dt R1 \u00fe R3\u00f0 \u00dei1: (2) Because V= (R2 +R4)i1, it is suitable to substitute for i1 as V \u00bc L1 di1 dt R1 \u00fe R3 R2 \u00fe R4 V \u00bc L1 di1 dt R1 R3 \u00fe 1 R2 R4 \u00fe 1 V : With R1 R3 \u00bc R2 R4 , the equation simplifies to V \u00bc 1 1\u00fe R1 R2 L1 di1 dt : (3) Copyright \u00a9 2015 John Wiley & Sons, Ltd" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0003442_1708-5284.12.2.189-Figure8-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0003442_1708-5284.12.2.189-Figure8-1.png", + "caption": "Fig. 8. Structure pulley.", + "texts": [], + "surrounding_texts": [ + "We notice that this same example or desired path (spiral) is studied by Vafaei et al. (2011).\nWith the same representation, in the example (2), we present the spiral trajectory for kinematic model\nfigure 6-a and spiral trajectory represented in work space figure 6-b, the cables lengths figure7-a and the cables tensions figure7-b values of eight cable-based robots necessary to follow spiral desired trajectory.", + "6. The dynamic response of our system with a PID controller In this section, we begin by presenting the dynamic equation of the robot with five and eight cables and its state-space representation. Then, the response will be simulated in closed loop with PID controller (Vadia et al., 2003).\n6.1. Dynamic Model of the End Effector The dynamic model of the actuator is expressed by the following relationship:\n(12a)\nWhere: m: is the mass matrix. X\u0308: is the acceleration vector of the end-effector. FR = (FRx FRy FRz)T: is the resultant force of the all\ntensions applied to the cables. Where:\n(12b)\n6.2. The dynamic comportment of the motors The dynamic comportment of the motor which represented on figure (8) is expressed by the following equation:\nm\nm\nm\nx\ny\nz\nFR0 0\n0 0\n0 0\n\n\n \n\n\n \n\n\n \n\n\n = && && &&\nx\nRy\nRz\nF\nF\n\n\n \n\n\n \nm X FR\n\u2022\u2022 =\n(13)\nWith:\n(14)\nJmat\nJ\nJ\nJ\nJ\nJ\n=\n\n\n \n1\n2\n3\n4\n5\n0 0 0 0\n0 0 0 0\n0 0 0 0\n0 0 0 0\n0 0 0 0\n\n\n \n=Cmat\nC\nC\nC\nC\nC\n1\n2\n3\n4\n0 0 0 0\n0 0 0 0\n0 0 0 0\n0 0 0 0\n0 0 0 0 5\n\n\n \n\n\n \nJ C rT\u03b2 \u03b2 \u03c4 \u2022\u2022 \u2022 + = \u2212 .", + "We consider that all the rays of the pulley are the same:\nri = r (i = 1.2...5 or 8), \u03c4 (\u03c41 \u03c42, ...\u03c4i)\u03a4 is the vector of the torques applied by the motors. t(t1,t2,\u2026ti)T is the vector of tension cables. \u03b2 is the angle of rotation of the pulley. So:\n(15)\nWhere Li0 are the initial lengths of the cables:\nSo\ni = 1,...,5 or 8 (16)\nby subtracting successively (13) with respect to time, we get:\n(17)\nSubstituting (12) we obtain:\n\u03b2 \u03b2 \u03b2\u2022\u2022 = \u2202 \u2202 + \u2202 \u2202 d dt x x x x& &&\n& &\u03b2 \u03b2\n\u03b1 \u03b1\n= \u2202\n\u2202 = \u2212 x x r\n1\n1 1 1 1 cos( ) cos( ) cos( ) sin( ) si\u0398 \u0398 n( )\ncos( ) cos( ) cos( ) sin( ) sin( )\nco\n\u03b1\n\u03b1 \u03b1 \u03b1\n1\n2 2 2 2 2 \u0398 \u0398\ns( ) cos( ) cos( ) sin( ) sin( )\ncos( ) co\n\u03b1 \u03b1 \u03b1\n\u03b1\n3 3 3 3 3\n4\n\u0398 \u0398\ns( ) cos( ) sin( ) sin( )\nsin( ) cos( ) si\n\u0398 \u0398\n\u0398\n4 4 4 4\n5 5\n\u03b1 \u03b1\n\u03b1 n( ) sin( ) cos( )\u03b1 \u03b1 5 5 5 \u0398\n\n\n \n\n\n & & & x y z \n\n\n \n\n\n \n\u03b2\n\u03b2 \u03b2\n\u03b2\n=\n\n\n \n\n\n =\n\u2212 \u2212 1 2 10 1 201 ( ) ( )\n( )\nX\nX\ni X\nr\nL L\nL\nM\nL\nL Li i\n2\n0\nM\n\u2212\n\n\n \n\n\n .\nL Aix Aiy Aizi0 2 2 2= + +( ) ( ) ( )\nt r\nJ C= \u2212 \u2212 \u2022\u2022 \u20221 ( ).\u03c4 \u03b2 \u03b2\n(18)\nFinally, the set of equations of the dynamic model can be expressed in a standard form for robotic systems (19):\n(19)\nWhere:\n(20a)\nAnd\n(20b)\ni=1,\u20265 or 8 (20c)\n7. Control law and architecture This section presents our control architecture and the low level control has been ensured by means of the implementation.\nOf a Proportional- Integrated and Derivative (PID) controller based on the overall system Cartesian dynamics equations of motion (17). (Vadia et al., 2003) in this paper the PID controller gains are determined the error using a Matlab programmed simulation to achieve reasonable performance for the trajectories.\nThe establishment of the control law along x, y and along z is:\n\u03c4\n\u03c4 \u03c4\n\u03c4\n= \u22c5 \u22c5\n\n\n \n\n\n \n1\n2\ni\nN X X S X J d\ndt X C X X( , ) ( )( )\n\u2022 \u2022 = + \u03b4\u03b2 \u03b4 \u03b4\u03b2 \u03b4\nM r m S X J X = +* ( ) \u03b4\u03b2 \u03b4\nX t M X N X X M X S X \u2022\u2022 \u2212 \u2022 \u2212= \u2217 + \u2217 \u2217( ) ( ) ( , ) ( ) ( )1 1 \u03c4\nt r\nJ d\ndt X X\nX X C X X= \u2212 + \u2212 \u2022 \u2022\u20221 \u03c4 \u03b4\u03b2 \u03b4 \u03b4\u03b2 \u03b4 \u03b4\u03b2 \u03b4 \u2022 " + ] + }, + { + "image_filename": "designv11_71_0002190_00207721.2020.1853272-Figure7-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0002190_00207721.2020.1853272-Figure7-1.png", + "caption": "Figure 7. The cooperative\u2013competitive weighted digraph G3.", + "texts": [ + " Evolutions of parameters \u03b8i for i = 1, 2, 3, 4 over the interactively balanced digraph G1. Figure 3. Evolutions of positions xi and velocities vi for i = 1, 2, 3, 4 over the interactively balanced digraph G1. satisfies Theorem 3.7. By calculation, we can obtain = diag(1, 2,\u22121, 2). According to Theorem 3.7, there are \u03b81 \u2192 \u2212\u03b83 \u2192 \u03b80 = 1 4 and \u03b82 \u2192 \u03b84 \u2192 2\u03b80 = 1 2 . In addition, x1 \u2192 \u2212x3 \u2192 x0, x2 \u2192 x4 \u2192 2x0, v1 \u2192 \u2212v3 \u2192 v0, v2 \u2192 v4 \u2192 2v0, which is aligned with Figures 5 and 6. Example 4.3: Consider system (3) over the cooperative\u2013competitive weighted digraph G3 of Figure 7, where the oscillators 1, 2, 3 form the 1st leader group, the oscillators 4, 5 constitute the 2nd leader group, and the remaining agents are the followers. In view of the cooperative\u2013competitive interaction weights \u03c9ij marked besides the edges (j, i), we can obtain Gm1 , Gm2 , and GF are interactively balanced. For the follower 6, there are two paths starting from the agents with the smallest label numbers of Gm1 and Gm2 , namely, the first path (1, 3), (3, 6), and the second path (4, 6). We can compute that the interaction weightproduct associated with the first path is \u03c931\u03c963 = (\u22122) \u00d7 4 = \u22128, and the interaction weight-product corresponding to the second path is \u03c963 = \u22128", + " By the same calculation, we can obtain the followers 6, 7 are also weight-product balanced. Moreover, it is not hard to calculate \u0306 = diag(1,\u22122,\u22122, 1, 13 ,\u22128, 2, 8). According to Theorem 3.14, we have x1(\u221e) = \u22121 2x2(\u221e) = \u22121 2x3(\u221e), x4(\u221e) = 3x5(\u221e), and\u22121 8x6(\u221e), 12x7(\u221e), 1 8x8(\u221e) \u2208 co{x1(\u221e), x4(\u221e)}. For the velocities of oscillators, we have the same conclusion. Figure 8 illustrates the state evolution of eight oscillators. It is seen that Figure 8 is in accordance with the above results. In order to verify Theorem 3.15, we only reset \u03c912 = 1 2 in Figure 7 and the rest remains unchanged for simplicity. As a result, the digraph Gm1 formed by the 1st leader group is not interactively balanced but all Ci = \u00b11 in G\u0302m1 , G\u0302m2 , and Ci = 1 in G\u0302F , which means that the agents of the system can achieve the weighted containment tracking motion. Here, the evolutions of xi and vi(i = 1, 2, . . . , 8) of the agents withGm1 not being interactively balance are presented in Figure 9. Obviously, the oscillators 1,2,3 in the 1st leader group achieve the state stability and the followers 6,7,8 reach weighted containment tracking motion" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0001073_s12083-019-00858-5-Figure1-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0001073_s12083-019-00858-5-Figure1-1.png", + "caption": "Fig. 1 Processes of falling down without protection", + "texts": [ + " For \u2200(\u03b4pB, \u03c9B) \u2208 S1(\u03b4pB, \u03c9B), if [ (f G B )T(\u03c4G B )T ] [ \u03b4pB \u03c9B ] \u2264 0 (4) the external contact state of the robot is invariant, where S1(\u03b4pB, \u03c9B) is a micro offset set for linear displacement and angular displacement of robot under no environmental interference. Equation 4 describes the necessary and sufficient condition for maintaining stable contact of a humanoid bipedal walking on the ground with large friction coefficient, as well as the necessary condition for maintaining a stable state on the ground with a small friction coefficient or non flat ground. Figure 1 shows the falling process of a humanoid robot. Since the distance between the robot\u2019s Center of Gravity (COG) and the landing point is short, the robot is simplified as a particle and described by a first-order 2-D inverted pendulum, and do the falling locomotion under gravity. After falling to the ground, the robot keeps in a stable state. Its ZMP is a certain point in the support polygon, which is determined by contact points of robot\u2019s limbs. The resultant force of the reaction passes the ZMP, and the stable state satisfies (3) and (4)" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0001057_0021998320920920-Figure13-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0001057_0021998320920920-Figure13-1.png", + "caption": "Figure 13. Contours of maximum Hashin failure index, Hmax, for warps in unit-cell 2 with binders and wefts shown for context.", + "texts": [ + " The difference is likely due to the greater distortion of the wefts\u2019 tow path and warps\u2019 cross-sectional shape due to the presence of the much larger binders in the plain orthogonally woven textile model. This suggests that the locations of important stress concentrations can have a strong dependence on parameters of the tow architecture and insights attained for a twill orthogonally woven textile may not be valid for other types of 3D textiles, even other types of orthogonally woven textiles such as plain orthogonally woven.25,26 Figure 13 shows contours of the maximum index of Hashin\u2019s criteria, Hmax, for the warps in unit-cell 2. Throughout the warps, the failure indices for Hashin\u2019s criteria remained significantly lower than in the binders and wefts, and highly localized concentrations of Hmax only developed in locations where a binder came close to a warp just after the path of the binder transitioned to be approximately aligned with the z-axis, such as the locations in the callouts in Figure 13. Throughout all of the warps, 0xx remained the largest stress component, and 0xx was the largest contributor to failure index corresponding to LT failure, HLT, which was the maximum index for all locations except the region surrounding and including the concentrations shown in Figure 13. Where the concentrations of Hmax (refer to equation (5)) occurred, the largest index corresponded to transverse tension, HTT, which was dominated by shear stresses. The largest contributor of the shear stresses was 0yz. Figure 14 shows contours for 0yz within the warps, and the locations of 0yz concentrations match the locations of the Hmax concentrations that were shown in Figure 13. As mentioned before, a previous study considered a plain orthogonally woven textile composite with larger binders and a smaller textile thickness.25,26 However, the warps in the twill orthogonal weave studied in this paper experienced very different stress states compared to the plain orthogonal weave. In the plain orthogonal weave,25,26 the warps were shown to experience severe longitudinal shear stresses, but the twill weave considered herein did not exhibit nearly as severe shear stresses. The difference is again probably due to the larger binders in the plain orthogonal weave causing much more distortion of the tow paths" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0000126_j.ifacol.2019.10.069-Figure16-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0000126_j.ifacol.2019.10.069-Figure16-1.png", + "caption": "Fig. 16. (a) Building the sectioned dome on the machine. (b) Built part on substrate plate.", + "texts": [ + " The nozzle should be tangent to the surface during production. Top edge of each sub-section is the substrate for starting the next section. The solution to prevent collisions is to build the part at the end of a plate. The plate is located vertically in the machine table. Therefore, when making the first layer of the first section, the trunnion table rotates 90\u00b0 to put the plate in a horizontal orientation. As the deposition of the layers progress, the plate turns vertically, and the dome gradually forms (Fig. 16). The benefit of the radial partitioning over the rotary tool path is that by increasing the number of partitions in Fig. 14, the part can be produced on a 3+2 axis machine which makes the complexity of the process diminish significantly. For a 3 + 2 solution, 10 sub-sections are required. However, if lead or tilt angles are required to address detailed part quality issues, the tool paths can be modified for each construction plane. If there are issues with the surface finish between sections, stock can be added, and the rough surface machined to a plane", + " Top edge of each sub-section is the substrate for starting the next section. Fig. 14. Splitting the surface dome into subsections. Fig. 15. Collision of the nozzle to the table when making section 2. The solution to prevent collisions is to build the part at the end of a plate. The plate is located vertically in the machine table. Therefore, when making the first layer of the first section, the trunnion table rotates 90\u00b0 to put the plate in a horizontal orientation. As the deposition of the layers progress, the plate turns vertically, and the dome gradually forms (Fig. 16). The benefit of the radial partitioning over the rotary tool path is that by increasing the number of partitions in Fig. 14, the part can be produced on a 3+2 axis machine which makes the complexity of the process diminish significantly. For a 3 + 2 solution, 10 sub-sections are required. However, if lead or tilt angles are required to address detailed part quality issues, the tool paths can be modified for each construction plane. If there are issues with the surface finish between sections, stock can be added, and the rough surface machined to a plane" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0001588_aim43001.2020.9158991-Figure7-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0001588_aim43001.2020.9158991-Figure7-1.png", + "caption": "Figure 7 Slider parts and assembled view", + "texts": [ + " The compressional springs are equipped along the slider pass. The PtM yields wire-tension and provides tension information that is obtained by measuring the compression length of springs by the linear potentiometers. (Fig.6). Figure 6. Fabricated PtM device All of the components are made of ABS. The guide shaft and slide potentiometer length are 60mm. We used 50mm helical compression springs with constant 1.8 N/m and with a 35% deflection rate. This means the traveling distance of the PtM slider is about 15-20 mm (Fig 7). Spring linear potentiometer PtM slider 461 Authorized licensed use limited to: University of Wollongong. Downloaded on August 11,2020 at 10:23:43 UTC from IEEE Xplore. Restrictions apply. Fig.9 shows an actuating unit; 4 units are equipped in the actuating part. One stepper motor (Trinamic Motion GmbH) rotates a low frictional lead shaft (6 mm lead) to slide the lead screw, by which two wires (1 mm steel cable) are oppositely pulled or eased. This mechanical design allows holding tension even the motor turns off" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0000136_012076-Figure1-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0000136_012076-Figure1-1.png", + "caption": "Figure 1. System of forces implemented in grinding area between rollers", + "texts": [ + " Due to the rolls design it is possible to achieve various methods of material destruction: compression, stretching and abrasion [4-6]. Using rolls in the form of Reuleaux Triangle profile allows achieving a combination of different grinding mechanisms [7]. This article describes the analysis of the material destruction in built-up construction of roll grinder. The paper is given a research of plates stress-strain state of combined rolls under load. \u0421ompletely reversing load is realized in the grinding area between the rollers in the form of RT-profile. Figure 1 shows vectors and directions of forces. The grinding area is shifted about spin axis of rolls; when RT-profile rotates. Owing to this effect, crushed material is mixed, and forces act on this material from different points and directions and with different intensity. In this way various methods of material destruction work together: compression, stretching, abrasion, impingement attack, constantly changing the forces directions operating on the material, change of numerical values and force vectors [8, 9]" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0002361_ffe.13405-Figure1-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0002361_ffe.13405-Figure1-1.png", + "caption": "FIGURE 1 Shape and dimensions of bearing test specimen, in mm: (A) test bearing NU206, (B) Type A defect, (C) Type B defect", + "texts": [ + " In each test condition, stress intensity was analysed via the finite element method (FEM) to evaluate the phenomenon, considering both rolling-contact and fitting conditions. Moreover, crack-growth resistance of bearing steel was measured under different stress ratios, with investigation of crack-opening/closing behaviour. The results offer novel criteria for creating a fracture map for rolling bearings. A cylindrical, JIS-NU206, roller bearing was used for fatigue-testing. The steel grade of the specimen bearing was JIS-SUJ2, the main components of which were 1% carbon and 1.5% chromium. Figure 1A displays the shape and dimensions of the test bearing. Two kinds of artificial defects were introduced onto the outer-ring raceway. Type A defects were semicircular slits, simulating early stage flaking, with depths ranging from 0.05\u20130.15 mm (cf. Figure 1B). On the other hand, Type B defects were through-thickness incisions, replicating deeper cracks, with depths ranging from 1.0\u20131.8 mm (cf. Figure 1B). Both defect types were generated by electron discharge machining (EDM). Crack-growth testing was also conducted on JISSUJ2, the chemical composition of which was 1.02C0.27Si-0.39Mn-0.022P-0.008S-1.39Cr (mass %), balanced with Fe. Rectangular plate-specimens (cf. Figure 2A) were fabricated from a round bar of 75 mm in diameter, with care taken to avoid the central region (cf. Figure 2B). The plates were heat-treated at 1113 K for 1 h, then quenched in an oil bath and finally tempered at 473 K for 1" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0000311_ecce.2019.8912581-Figure1-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0000311_ecce.2019.8912581-Figure1-1.png", + "caption": "Fig. 1. Structure and phase current diagram of the 12/10 six-phase DCbiased HEVRM. (a) Machine structure. (b) Stator currents.", + "texts": [ + " In part II, the structure of DC-biased HEVRM prototype and the basic drive system introduced, and the mathematical equations in stationary reference frame and rotating reference frame are given. In part III, to simplify the calculation, a MTPA current control strategy based on polynomial fitting method is shown to improve the average torque. Furthermore, a harmonic current suppression strategy is introduced to enhance the effect of the MTPA control strategy. Experimental results are carried out in part IV. II. BASIC DRIVE SYSTEM AND MATHEMATICAL EQUATIONS The profile of the 12/10 six-phase DC-biased HEVRM prototype is shown in Fig. 1 (a). The PMs are inserted in the center of stator tooth, and the integrated field and armature windings are wounded on the stator tooth. Hence, the permanent magnetic circuit and the field winding excitation magnetic circuit are in series. The six-phase stator current waveforms are shown in Fig. 1 (b). The phase currents can be expressed as follows: 1 1 1 cos( ) cos( ) cos( ) A ac e dc B ac e dc C ac e dc i I I i I I i I I \u03b8 \u03b8 \u03b3 \u03b8 \u03b3 = + = \u2212 + = + + (1a) 2 2 2 cos( ) cos( ) cos( ) A ac e dc B ac e dc C ac e dc i I I i I I i I I \u03b8 \u03b8 \u03b3 \u03b8 \u03b3 = \u2212 = \u2212 \u2212 = + \u2212 (1b) where Iac is the magnitude of the AC current component, Idc is the DC-biased current component, \u03b8e is the electrical angle, \u03b3 is the phase difference between adjacent phases. The stator windings are divided into two winding groups, in phase A1, B1 and C1, the DC-biased current component is positive, in phase A2, B2 and C2, the DC-biased current component is negative" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0003634_icems.2014.7013745-Figure5-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0003634_icems.2014.7013745-Figure5-1.png", + "caption": "Fig. 5. Magnetic field distribution when inner motor is energized alone", + "texts": [ + " In the PMRM, because of the doubly salient structure of outer motor, magnetic circuit of outer motor will be different when outer rotor rotates to a different position. Magnetic field distribution of PMRM when permanent magnets act alone is shown in Fig. 3 and Fig. 4. As Fig. 3 shows, when outer rotor rotates to this position, magnetic flux will go through two aligned teethes of stator and outer rotor. When there are no aligned teethes, magnetic flux will go through adjacent teethes as Fig. 4 shows. Magnetic field distributions of PMRM when inner rotor or stator armature windings energized alone are shown in Fig. 5 and Fig. 6. motor will increase or decrease when +D axis or \u2013D axis currents of inner motor act alone. By comparing Fig. 5(c) with Fig. 5(d), it can be seen that when +Q axis currents act alone, magnetic fluxes of outer motor little change. From Fig. 5, it can be obtained that D axis currents of inner motor has much stronger effect on magnetic field distribution of outer motor than Q axis currents of inner motor. As Fig. 6(a), Fig. 6(b) and Fig. 6(c) show, stator teeth of phase C and outer rotor teeth align completely. When positive currents of phase C act alone, magnetic fluxes of inner motor decrease obviously. Meanwhile, when negative currents of phase C act alone, magnetic fluxes of inner motor are little changed. As Fig. 6(d), Fig. 6(e) and Fig", + " 6(f) show, stator teeth of phase C and outer rotor teeth align incompletely. When positive currents or negative currents of phase C act alone, magnetic fluxes of inner motor are little changed. In the actual operation process, when single phase of outer motor is energized, stator teeth of this phase and outer rotor teeth align incompletely. Based on the situation above, it can be obtained that the influence that outer motor acts on inner motor is much weaker than the influence that inner motor acts on outer motor from Fig. 5 and Fig. 6. Therefore, the rests of the paper focus on the analysis of influence that inner motor acts on outer motor. The numerical results of electromagnetic coupling effect on magnetic flux that inner motor acts on outer motor are shown in Fig. 7, Fig. 8, Fig. 9 and Fig. 10. Current shown in Fig. 7 and Fig. 8 is Q axis currents of inner motor, magnetic flux indicates that created by inner motor, and angle indicates mechanical angle that outer rotor rotates counterclockwise. Fully unaligned teeth position indicates 0 degree and fully aligned teeth position indicates 14 degree" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0002204_icem49940.2020.9270855-Figure4-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0002204_icem49940.2020.9270855-Figure4-1.png", + "caption": "Fig. 4. Flux lines of FSCW-DPM machine with various PM excitations: (a) \u03b8=90 deg, rotor PM only. (b) \u03b8=90 deg, stator PM only. (c) \u03b8=0 deg, dual PM. (d) \u03b8=90 deg, dual PM. (e) \u03b8=180 deg, dual PM. (f) \u03b8=270 deg, dual PM.", + "texts": [ + " Downloaded on May 15,2021 at 05:01:08 UTC from IEEE Xplore. Restrictions apply. Fig. 3. Simulated air-gap flux densities by PMs of the FSCW-DPM machine. (a) Waveforms. (b) Harmonics. It is well known that harmonics in armature winding MMF and PM exciting fields having the same spatial order, same rotating speed and direction can interact to generate non-zero torque. For the analyzed 12-slot, 7-pole pair, single-layer FSCW machine, by comparison in Fig. 2 and Table II, the working harmonics that can produce torque can be derived in Table III. Fig. 4 shows the no-load flux line distributions with different PM excitation conditions. As shown in Fig. 4 (a), when the rotor PM excited, the flux lines pass through the air gap, and through the stator teeth as much as possible, then go across the air gap again, and return to the rotor PM pass through the rotor salient poles. When only stator Halbach PM excited (Fig. 4 b), one part of the flux lines emitted from the sides of the PM passes through the stator yoke, links with coils and then returns to the rotor through the air gap, and the other part returns directly to the rotor through the air gap. Figs. 4 (c)-(f) show the flux line distributions at four rotor positions when dual PMs excited. The flux lines basically form a loop through the adjacent stator teeth, and the feature of a very short magnetic circuit is kept. (a) (b) Fig. 5 shows the no-load phase A flux-linkage waveforms with different PM excitations, the six marked points correspond to the flux line distributions in Fig. 4. The phases of the flux-linkage induced by stator/rotor PM are the same, and perfect sinusoidal waveforms are obtained. In addition, the introduction of stator PM makes the amplitude increase from 0.075wb to 0.09wb, by 20%. Fig. 6 reveals the no-load phase A back-EMF waveforms and harmonics with different PM excitations. The phases of the back-EMF waveforms are the same. Besides, the comparison of the waveforms shows that the introduction of stator PM makes the back-EMF amplitude increase from 175 V to 210 V, 20% higher back-EMF value being obtained" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0003702_cca.2014.6981367-Figure1-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0003702_cca.2014.6981367-Figure1-1.png", + "caption": "Fig. 1. Single-track vehicle model.", + "texts": [ + " In this paper, for front steering vehicles, we propose a combined robust model predictive controller comprised of MPC with side-slip angle restriction [14] and SMC [11]. MPC generates an optimal reference control input satisfying the restriction of the side-slip angle, then SMC effectively suppress the external disturbance to achieve the reference. The advantage of our controller is verified through comparative simulation with previous methods including nonlinear MPC using nonlinear tire model. II. VEHICLE DYNAMICS Figure 1 depicts the model of a front steering vehicle described as a single-track model. V is the velocity, (x, y) is the gravity center, \u03b8 is the attitude angle, \u03b2 is the side slip angle, \u03b3 is the yaw rate, \u03b4 is the steering angle and Lf , 978-1-4799-7409-2/14/$31.00 \u00a92014 IEEE 328 Lr are front / rear wheel base, respectively. Fy,f and Fy,r represent front / rear side force, respectively. yd(x) represents the desired path. We use the subscript f for front wheel and r for rear wheel (i = f, r). Assuming the vehicle velocity is constant and the load shift of vehicle can be ignored, the equation of vehicle dynamics are represented by dx dt =V cos(\u03b8 + \u03b2) =: fx, (1) dy dt =V sin(\u03b8 + \u03b2), (2) d\u03b8 dt =\u03b3, (3) mV d(\u03b8 + \u03b2) dt =Fy,f cos(\u03b2 \u2212 \u03b4) + Fy,r cos\u03b2, (4) Iz d\u03b3 dt =LfFy,f cos \u03b4 \u2212 Fy,rLr =: f\u03b3 , (5) where m and Iz denote the vehicle mass and the inertia moment" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0002222_icstcc50638.2020.9259658-Figure2-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0002222_icstcc50638.2020.9259658-Figure2-1.png", + "caption": "Fig. 2. General N-Trailer-System with car-like tractor.", + "texts": [ + " Two vehicle models are described: the fixed-frame model, which describes the vehicle movement on a plane and is employed in the simulation, and the offset model, used in the control algorithm. The derivation of kinematic vehicle models is vastly described in literature, e.g. [12]. In this contribution, a compact vector notation [2] is used for the models. The vehicle model 978-1-7281-9809-5/20/$31.00 \u00a92020 IEEE 489 Authorized licensed use limited to: Univ of Calif Santa Barbara. Downloaded on May 21,2021 at 03:33:25 UTC from IEEE Xplore. Restrictions apply. of a General N-Trailer-System with a car-like tractor in a fixed Cartesian reference system is shown in Fig. 2. Since the number of trailers is N , the total number of bodies is m = N + 1. The front axle of the car-like tractor is steered and indexed by 0. The position of the axle centers on the plane are p0 . . .pm, with the components (pi,1, pi,2), and the orientation of the axles is described by \u03b80 . . . \u03b8m. L1 is the wheelbase, L2 . . . Lm is the distance from the hitch point to the wheel axle of the trailers. The parameter D was previously defined and has to be specified for D1 . . . Dm\u22121. By introducing the tangential vector \u03c4 i = cos \u03b8ie1+sin \u03b8ie2 and the normal vector \u03bdi = \u2212 sin \u03b8ie1 + cos \u03b8ie2, and after short calculations, the kinematic model equations follow as p\u03071,1 = v1 cos \u03b81 (1) p\u03071,2 = v1 sin \u03b81 (2) \u03c91 = (v1/L1) tan \u03b40 (3) \u03c9i = (1/L1)[v1\u3008\u03c4 1,\u03bdi\u3009 \u2212 i\u22121\u2211 j=2 (Lj +Dj)\u03c9j\u3008\u03bdj ,\u03bdi\u3009 \u2212D1\u03c91\u3008\u03bd1,\u03bdi\u3009], i = 2, " + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0003626_978-90-481-9707-1_27-Figure84.6-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0003626_978-90-481-9707-1_27-Figure84.6-1.png", + "caption": "Fig. 84.6 For each execution of one motion primitive, there are two segments of the path corresponding to two circular arcs of angle ?, so that the distance between the UAV and the UGVs is less or equal to r when the UAV is on these segments. In this figure, the dashed curve is the path of the UAV, and the solid curves are the segments of the path in which convoy protection is provided. The UAV and the convoys are drawn at the times when the UAV enters and exits these segments", + "texts": [ + " Using a similar approach as in the static convoys, case, it is possible to determine the minimum number of UAVs required. In this case, on the path of each execution of one motion primitive, there are two segments of the path when the distance between the UAV and the UGVs is less than or equal to r . Positions of UAV and UGVs and hence their distance can be uniquely characterized by the arc angle as d. /. Convoy protection is provided by one UAV for two circular arcs of angle ? for each execution of one motion primitive, where d. ?/ D r . Refer to Fig. 84.6 for an example. Similar to the multi-UAV coordination approach in the previous section, one can use a timing strategy to schedule the UAVs such that, at any time, one of the UAVs is inside the convoy circle. First, note that the minimum number of UAVs required to provide continuous convoy protection can be obtained by the following corollary, which directly follows from the fact that, for each motion primitive, the length of the path in which one UAV stays inside the convoy circle is 2R ?, while the length of the entire path for the motion primitive is 2R\u02c7" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0001442_j.matpr.2020.06.022-Figure1-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0001442_j.matpr.2020.06.022-Figure1-1.png", + "caption": "Fig. 1. The frame element.", + "texts": [ + " (2), we get M\u00bd x2 Xf gsinxt \u00fe \u00bdK Xf gsinxt \u00bc 0 \u00f06\u00de Since sin x t\u2013 0, the above equation can be reduced to x2\u00bdM fXg \u00fe \u00bdK fXg \u00bc 0\u00f07\u00de Substituting k =x2 the above equation may be rewritten as \u00bdK k\u00bdM \u00bd Xf g \u00bc 0 \u00f08\u00de Equation (8), the eigenvalue problem, is solved by the inverse iteration scheme to find out the natural frequencies and the mode shapes of the given structure. Eigenvalue gives the value of natural frequencies, and the eigenvectors give the mode shape for a particular eigenvalue. 2.2. The FEM methodology The methodology used for the vibration analysis of the machine tool structures has been described below:- 1. The machine tool structure is assumed to be discretised into \u2018n\u2019 number of \u20182-noded frame elements\u2019. The value of \u2018n\u2019 depends on the shape and complexity of the structure. 2. As shown in Fig. 1, the 2-noded frame element has been obtained by combining a rod element and a beam element [1]. Each of its two nodes is associated with two linear displace- Please cite this article as: J. Garg and S. Bala Garg, A simplified methodology for tures, Materials Today: Proceedings, https://doi.org/10.1016/j.matpr.2020.06.0 ments and one rotational deformation. Consequently, every frame element has six degrees of freedom. The shape functions for the frame element were obtained by combining the shape functions of the beam and the rod elements, and are given below:- N1 = (L \u2013 x) /L N2 = 1 \u2013 (3x2 / L 2) + (2x3 / L 3) N3 = x \u2013 (2x2 / L) + (x3 / L2) N4 = x / L N5 = (3x2/ L 2) \u2013 2 (x3/ L 3) N6 = \u2013 (x2/ L) + (x3/ L 2) In the matrix from [N] = {N1 N2 N3 N4 N5 N6} Where [N] is the matrix of shape functions for a frame element" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0002268_j.apples.2020.100032-Figure3-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0002268_j.apples.2020.100032-Figure3-1.png", + "caption": "Fig. 3. Initial situation and numerical reference results (a) Initial model in exploded view (b) components to be optimized (c) Numerical reference results of the existing wheel component (deformations).", + "texts": [ + " Due to the structure esulting from the five load cells and the rotating load (bending force F B ) f the measuring wheel, a cyclic symmetry for the optimized structure hould be aimed at. Accordingly, only a fifth of the structural compoent has to be considered for the further optimization process, since the urther geometry is obtained by mirroring. In order to map the circulatng load, the measuring wheel component is analyzed in 4\u00b0 rotating load teps. In this way a sufficient accuracy is achieved. The connections to he other add-on components are represented by bolts idealized as beam lements and are subjected to a defined preload force. Fig. 3 shows the initial situation before the optimization. Fig. 3 a llustrates the design of the WFT in exploded view. The components o be considered for the structural optimization are colored in red. In first step, the integration of functions can already eliminate various onnecting elements, allowing an initial weight reduction and reduced ssembly effort. Fig. 3 b shows the components to be optimized in an nlarged form. Reference simulations were carried out to obtain values for the perissible stresses and deformations. By changing the material from aluinum to Ti6\u20134 (see Leuders et al. (2013) ), a stiffness advantage due o the higher Young\u2019s modulus is already achieved independently of he subsequent structural optimization. The safety factor to be observed as calculated from the stresses. The permissible deformation and thus he required stiffness of the component can be taken directly from the umerical simulations of the reference component. The results are dislayed in a standardized form. A value of 1 corresponds to the max- i o t u s t w I c v r a f a T t 3 t a n t l s t r r v o l b t e c s t v i t m d A t t o i a m F f mum permissible stress / displacement in the reference part, a value f 0 corresponds to no occurring deformation. The reference results for he occurring deformations are shown in Fig. 3 c. For the numerical sim- lations with the software Abaqus, the components were meshed with ufficiently small square tetrahedral (C3D10). Due to the small deformaions a geometrical linear calculation was carried out. The bolt preloads ere determined based on the real bolt preloads given by the Kistler nstrumente Gmbh . In order to reduce the simulation time, all screw onnections were simplified by truss elements with bolt loads. In a preious simulation it could be shown that no significant change of the esults occurs by this assumption. The load values and the lever arm are ssumed in relation to the test specifications for measuring wheels. The resulting stresses are low, with the exception of the bolt interace and minor stress increases due to notch effects. The deformation nalysis ( Fig. 3 c) defines a first restriction for the optimization process. he optimized wheel component may have at most the deformations of he existing wheel component. . Optimization process and result With knowledge of the relevant boundary conditions and constraints, he basis for the optimization process is established. In order to systemtically optimize the structure, the non-design space resulting from conections to surrounding components is first defined. In this way, the iniial model for the optimization is determined" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0002342_rcar49640.2020.9303289-Figure3-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0002342_rcar49640.2020.9303289-Figure3-1.png", + "caption": "Figure 3. Solidworks model of rope drive mechanical module", + "texts": [ + " The calf wearing module mainly binds the rope drive module and the spring module to the calf. The four modules of the flexible ankle-assist robot prototype are related to each other and also play their own roles. The rope drive mechanical module is equipped with a drive motor and transmission. It is the power source for all modules and plays an important role. According to the compact structure requirements of the rope drive mechanical module, the characteristics of the integrated rope and the steering gear structure, we design the mechanical structure which is shown in figure 3. The structure consists of a connecting plate, a reel, and a pulley. The connecting plate is used to connect the steering gear and the reel. At the same time, a slot is designed at the rear for inserting the Velcro tape and then connecting with the calf wearing module. In addition, the connections and tightness of the rope are also a consideration when designing. The connection between the two ends of the rope, one end is connected with the reel, which is not easy to remove frequently. The other end is connected to the foot wearing module" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0000253_2019045-Figure8-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0000253_2019045-Figure8-1.png", + "caption": "Fig. 8. Helix diagram of left tooth surface after deformation.", + "texts": [], + "surrounding_texts": [ + "Taking the measured temperature gear as an example [1], the material of gear is 45 steel and the coefficient of linear expansion l=11.6e 6c 1, room temperature t0= 25 \u00b0C. Table 1 shows other parameters of the gear. The tooth surface temperature distribution measured by Zhengzhou Institute of machinery [1] is shown in Figure 9. Because the installation error andmachining error of double helical gear are not considered, the gear tooth temperature distribution of two tooth surfaces is the same. According to Figure 9, the temperature equation along reference circle axial with a linear velocity of 130m/s is deduced. 0 b 8 t \u00bc 68:5 8 < b 33 t \u00bc 0:16b\u00fe 67:22 33 < b 53 t \u00bc 0:2b\u00fe 65:9 53 < b 73 t \u00bc 0:1b\u00fe 71:2 73 < b 93 t \u00bc 0:25b\u00fe 60:25 93 < b 108 t \u00bc 0:24b\u00fe 61:18 108 < b 123 t \u00bc 0:1b\u00fe 76:3 123 < b t \u00bc 88:6 8>>>>>>< >>>>>: \u00f09\u00de According to formulas (4), (5), (7), (8). the tooth surface before and after thermal deformation are shown in Figures 10\u201313 (where, the x axis describes longitudinal direction, the y axis describes tooth thickness direction and the z axis describes radial direction). Because the installation error and machining error of double helical gear are not considered, the thermal deformation of Figures 12 and 13 is the same. Take left tooth surface as an example, the difference of the tooth surface coordinates before and after thermal deformation is shown in Figure 14, the mean value of the difference of the tooth surface coordinates before and after thermal deformation is shown in Table 2. According to reference [1], Table 3 showed the modification of thermal deformation, which corresponds to the deformation of tooth thickness. In this paper, the diameter of reference circle is 199.36mm, which is close to 200mm in Table 4. Therefore, the modification can be taken as 0.009mm. In Table 3, the mean value of the difference of the tooth surface coordinates along tooth thickness direction is 0.009627mm. The error is one order of magnitude smaller than them. When the installation/machining error of gear exists, the meshing path of two tooth surfaces change. Reference [12] gave the meshing path of left and right tooth surfaces under installation/machining error of gear (Fig. 15). The temperature of instantaneous contact points on the left/ right tooth surface is assumed that still satisfies formula (9). The thermal deformation of two tooth surfaces is recalculated, the mean value of the difference of tooth thickness direction without and with considering installation error and machining error of gear is shown in Table 4. Linear velocity/(m/s) Modification of thermal deformation Gear diameter/mm 100 150 200 250 300" + ] + }, + { + "image_filename": "designv11_71_0001329_robosoft48309.2020.9116019-Figure10-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0001329_robosoft48309.2020.9116019-Figure10-1.png", + "caption": "Fig. 10 Target object (silicone tube) and task plate", + "texts": [ + " 8 was prepared for the manipulation, sensible CAVS was attached to each finger of the gripper, and the length in the longitudinal direction of the CAVS was set to 25 mm. As shown in Fig. 4, the longitudinal direction exhibited a larger ratio of friction change. Therefore, the finger was constructed so that the longitudinal direction of the finger could correspond to the longitudinal direction of CAVS\u2014as the longitudinal direction of the finger corresponded to the gravitational direction for most of the operation. The experimental setup is shown in Fig. 9. We used a robotic arm (UR5) equipped with the gripper, with sensible CAVS at the tip. As shown in Fig. 10, we prepared the task plate containing a cylinder (diameter: 70 mm, height: 30 mm), and two pins (diameter: 4 mm, height: 17 mm) for hooking. The target object was a tube with hooks attached to either end (Fig. 10). Its diameter was 8 mm and its length was 320 mm. Initially, one hook was attached to the pin on one side. The task was for the robot system to pick up the tube, train it around the cylinder, and attach the hook on the other end of the tube to the pin on the other side of the cylinder. Fig.12 Overview of the manipulation The main procedure for the test was as follows: Step 1. Grasp the tube Step 2. Slide the fingertips along the tube to the position required for the subsequent operation Step 3" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0000204_ecce.2019.8912828-Figure2-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0000204_ecce.2019.8912828-Figure2-1.png", + "caption": "Fig. 2. Conductors considering skin depth. (a) rectangular wire (b) round wire", + "texts": [ + "00 \u00a92019 IEEE 3137 ( ) ( ) ( ) ( ) 2 1 3 sinh 2 sin 2 sinh sin , 2 cosh 2 cos 2 cosh cos t DC zR R \u03d5 \u03be \u03c8 \u03be \u03be \u03be \u03be \u03be\u03d5 \u03be \u03be \u03c8 \u03be \u03be \u03be \u03be \u03be \u03be \u2212 = + + \u2212= = \u2212 + (2) where, f is the frequency, \u03bc0 is the permeability of conductor, \u03c3 is the conductivity of conductor, hc0 and bc0 are the height and width of a turn respectively, zt and za are the number of layer in row and column respectively, b is the width of the slot, and \u03be is the reduced conductor height. To explain the AC resistance of round wire instead of rectangular wire like Fig.1 (b), the equivalent conductivity of conductor was derived from (3) and (4) based on the resistances considering skin depth as shown in Fig 2. ( )( ) ( )( ) ( ) 22 0 0 22 0 0 0 0 0 2 2 4 rectangular c c round c c c c c l R h h l R D D h b D \u03c3 \u03b4 \u03c0 \u03c3 \u03b4 = \u2032 \u2212 \u2212 = \u2212 \u2212 = = (3) where, Rrectangular and Rround are the resistance considering skin depth for rectangular wire and round wire respectively, \u03c3' is the equivalent conductivity of conductor, \u03b4 is the skin depth, and Dc0 is the diameter of round wire. Assuming the Rrectangular and Rround are the same, the equivalent conductivity of conductor can be defined as (4). ( )tan4 rec gular roundR R \u03c0\u03c3 \u03c3\u2032 = = (4) By replacing the conductivity in (1) into equivalent conductivity in (4), the analytic solution for AC resistance of series connected round wire can be derived by modifying (1) into (5)" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0003565_ipec.2014.6869805-FigureI-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0003565_ipec.2014.6869805-FigureI-1.png", + "caption": "Fig. I. Image of power factor in mechanical system.", + "texts": [ + " 1 represents apparent power and active power in mechanical system. Robot should have had apparent power to do the work, while only some amount of apparent power is used in a motion as a active power. The part of power, which is not used for the motion, is called reactive power. Relationship between active power, apparent power and reactive power is shown on Fig. 2 as analogic representation with electrical circuit. Force provides effort to perform motion, velocity is flow in the system, and energy is transferred to the load as shown in figure. In contact motion case, load is present as impedance Z. A. Active power As in electrical system, active power in mechanical system should explain the amount of power used in the work. From the analogy, active power in mechanical system can be defined as p Pact !X, 1 {T T Jo pdt. (1) (2) In real system, active mechanical power Pact should be avail able from the integral of instantaneous mechanical power p. f and x are instant force and instant velocity respectively. In contrast to active power Pact, a system should have some amount of power they apply for work" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0000033_s00202-019-00846-1-Figure6-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0000033_s00202-019-00846-1-Figure6-1.png", + "caption": "Fig. 6 Field distribution on the entire surface of the PMSG. a Magnetic flux density and b active flux paths", + "texts": [ + " Also, the 2D XY problem can be reduced to scalar form, and the magnetic vector potential A\u20d7 is reduced to the z-axis scalar AZ [14, 22, 36]: In relation to that, in the case that the current flowing in the stator winding is modeled as stranded coils; the current densities for stranded coils are given by [34]: (10) 1 \ud835\udf070 \u2207 \u00d7 (\u2207 \u00d7 A\u20d7) \u2212 \u2207 \u00d7 M\u20d7 = \u2212\ud835\udf0e\u2207\ud835\udef7 \u2212 \ud835\udf0e dA\u20d7 dt + \ud835\udf0ev \u00d7 (\u2207 \u00d7 A\u20d7) (11)\ud835\udf0e dA\u20d7Z dt \u2212 1 \ud835\udf070 \u22072AZ = \u2212\ud835\udf0e\u2207\ud835\udef7 + \u2207 \u00d7 M\u20d7 1 3 where nc is the number of turns, Sc is the cross-sectional area of the coil, i(t) is the current per turn. In the study carried out, ANSYS Maxwell 2D package is used in which finite element analysis is employed to prove the correctness of the values calculated numerically. Magnetic analysis instead of quantitative techniques is preferred to determine the structure of designed prototype generator. Magnetic flux density distribution in the air gap and the stator yoke is presented in Fig.\u00a06 as a result of derived simulations. In order to see the flux densities on the machine, the flux density distribution in the whole section of the machine (Fig.\u00a06a) has been obtained with a 10\u00b0 rotor rotation. Here, the rotor yoke density is around 0.9\u00a0T in the center of the pole spring and 2.06\u00a0T for the opposite sides of the rotor. Also, the flux density of the magnet is around 0.83\u00a0T. Magnet flux density is an important parameter for demagnetization analysis. The average flux density value on the entire surface is 1.3\u00a0T, while the air-gap flux density is around 0.8\u20131\u00a0T. In addition, stator teeth flux density is 1.5\u00a0T. As the saturation point of silicon steel sheet used was 1.7\u00a0T, the sheet did not face saturation. Thus, the hysteresis losses of the machine will be very low. Figure\u00a06b shows the magnetic flux and flux path formed in magnets. The leakage of (12)Js = \u2212 1 0 \u22072AZ = nc Sc the magnet and the air gap through the slots has not been observed. Also, the amount of leakage flux between adjacent magnets in the opposite direction is not large. In today\u2019s research and development processes, the simulation of dynamic systems allows predictions and concept decisions related to the final product to be made at an early stage. This not only involves the modeling, simulation and testing of individual structural components or modules, but also requires the interplay of a large number of functions (with simulation models and hardware components from various domains) to build up the full system" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0002047_ecce44975.2020.9235956-Figure6-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0002047_ecce44975.2020.9235956-Figure6-1.png", + "caption": "Fig. 6. Illustrations of the central line offsets.", + "texts": [ + " From the figures, it can be shown that after applying phase shift elimination, the phase shift between PM and DC produced flux-linkages can be significantly reduced. Fig. 5(d) shows the variations of phase shift angles with and without elimination against different PM ratios. It can be seen that the phase shift elimination illustrated in (11) is still valid for different PM ratios. Although the phase shift angle cannot be reduced to zero, it is much smaller compared with the original . In fact, there will be offsets between the PM (DC) produced flux and the geometric central line of PM (iron) pole. If the offsets are considered (shown in Fig. 6(a)), (10) will be changed to (16), where is the PM field angle offset and is the DC field angle offset. + + ( \u2212 ) = 180\u00b0 (16) The amplitudes of and are quite small, which can also be obtained in Fig. 5(d). Therefore, the method of phase shift elimination in (10) is applicable in Type I AS-CP-HEFRMs. For Type II, the initial rotor position is shown in Fig. 1(b). It should be noticed that the number of rotor pole is even and the offsets of central lines of magnetic field and poles are neglected. In only PM excitation operating mode, if only the DC bias and fundamental harmonic are considered, will 17 Authorized licensed use limited to: UNIVERSITY OF CONNECTICUT", + " When the rotor pole number is odd in Type II, flux density across Coil 4 in PM only operating mode turns to be: = \u2212 cos( \u2212 ) (22) Thus, the flux-linkage of Coil 4 can be expressed as: = ( \u2212 cos( \u2212 )) (23) Therefore, the phase flux-linkage of PM excitation will be: = 2 \u2212 2 sin sin (24) Correspondingly, the phase flux-linkage of DC excitation will be: = \u22122 \u2212 2 sin sin (25) In hybrid operating mode (flux-enhancing), the flux-linkage of phase A for odd-number-rotor machines in Type II can be written as: = 2 ( \u2212 ) \u22122 ( sin + sin ) sin (26) From (21) and (26), it can be seen that Type II machines with even-pole-number rotors exhibit a bipolar phase fluxlinkage, while odd-rotor-pole-number machines exhibit a bipolar phase flux-linkage with DC bias, as shown in Fig. 7(b). If central line angle offsets are taken into consideration (shown in Fig. 6(b)), the expressions of the phase fluxlinkages of hybrid excitations (flux-enhancing) will be: For even rotor pole number in Type II: = \u22122 ( sin( + ) + sin( + )) sin (27) For odd rotor pole number in Type II: = 2 ( \u2212 ) \u22122 ( sin( + ) + sin( + )) sin (28) It can be seen from (27) and (28) that the central line offsets only have influence on the amplitude of the phase fluxlinkage but not on phase angles of PM and DC excitations. Consequently, Type II can be defined as an \u2018asymmetric topology and symmetrically excited\u2019 AS-CP-HEFRM" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0000253_2019045-Figure7-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0000253_2019045-Figure7-1.png", + "caption": "Fig. 7. Helix diagram of left tooth surface after deformation.", + "texts": [], + "surrounding_texts": [ + "Taking the measured temperature gear as an example [1], the material of gear is 45 steel and the coefficient of linear expansion l=11.6e 6c 1, room temperature t0= 25 \u00b0C. Table 1 shows other parameters of the gear. The tooth surface temperature distribution measured by Zhengzhou Institute of machinery [1] is shown in Figure 9. Because the installation error andmachining error of double helical gear are not considered, the gear tooth temperature distribution of two tooth surfaces is the same. According to Figure 9, the temperature equation along reference circle axial with a linear velocity of 130m/s is deduced. 0 b 8 t \u00bc 68:5 8 < b 33 t \u00bc 0:16b\u00fe 67:22 33 < b 53 t \u00bc 0:2b\u00fe 65:9 53 < b 73 t \u00bc 0:1b\u00fe 71:2 73 < b 93 t \u00bc 0:25b\u00fe 60:25 93 < b 108 t \u00bc 0:24b\u00fe 61:18 108 < b 123 t \u00bc 0:1b\u00fe 76:3 123 < b t \u00bc 88:6 8>>>>>>< >>>>>: \u00f09\u00de According to formulas (4), (5), (7), (8). the tooth surface before and after thermal deformation are shown in Figures 10\u201313 (where, the x axis describes longitudinal direction, the y axis describes tooth thickness direction and the z axis describes radial direction). Because the installation error and machining error of double helical gear are not considered, the thermal deformation of Figures 12 and 13 is the same. Take left tooth surface as an example, the difference of the tooth surface coordinates before and after thermal deformation is shown in Figure 14, the mean value of the difference of the tooth surface coordinates before and after thermal deformation is shown in Table 2. According to reference [1], Table 3 showed the modification of thermal deformation, which corresponds to the deformation of tooth thickness. In this paper, the diameter of reference circle is 199.36mm, which is close to 200mm in Table 4. Therefore, the modification can be taken as 0.009mm. In Table 3, the mean value of the difference of the tooth surface coordinates along tooth thickness direction is 0.009627mm. The error is one order of magnitude smaller than them. When the installation/machining error of gear exists, the meshing path of two tooth surfaces change. Reference [12] gave the meshing path of left and right tooth surfaces under installation/machining error of gear (Fig. 15). The temperature of instantaneous contact points on the left/ right tooth surface is assumed that still satisfies formula (9). The thermal deformation of two tooth surfaces is recalculated, the mean value of the difference of tooth thickness direction without and with considering installation error and machining error of gear is shown in Table 4. Linear velocity/(m/s) Modification of thermal deformation Gear diameter/mm 100 150 200 250 300" + ] + }, + { + "image_filename": "designv11_71_0002989_978-3-642-16214-5_9-Figure9.45-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0002989_978-3-642-16214-5_9-Figure9.45-1.png", + "caption": "Fig. 9.45. Single-sided plates (SSP). 1 Inner plate, one-sided lining; 2 outer plate, one-sided lining", + "texts": [ + " Design measures to compensate the rotational pressure include: \u2022 stronger return spring: inexpensive, but reduces the effective shifting pressure, \u2022 ball valve for clutch discharge: low effort, but unsuitable for controlled clutches and overlapping shiftings, since the function is influenced by speed and oil temperature, \u2022 spring-controlled clutch discharge valve: relatively high effort, unsuitable for controlled clutches and overlapping shiftings, since the function is influenced by speed and oil temperature, \u2022 shift pressure-controlled discharge valve: functionally reliable, but very high effort, also unsuitable for controlled clutches and overlapping gearshifting, \u2022 pressure compensation chamber (Figure 9.41, rotational pressure compensation B 2 and E 11): requires a lot of installation space, yet always functions and can also be used for controlled clutches and overlapping shiftings. As opposed to conventional lined plates with double-sided friction lining (DSP = Double-Sided Plates, Figure 9.44), Single-Sided Plates (SSP) have only one friction lining alternating between the inner and outer plates (Figure 9.45). By increasing the available volume of steel for thermal storage, the load capacity of the clutch can be increased at equal installation space. However, SSPs are more expensive than DSPs, they increase the danger of assembly faults and tend to wobble at high rotational speeds due to their low weight. In the following, a few questions of detail concerning multi-plate clutches will be addressed. Drag torque is defined as the loss torque at open shifting plates through-flowed or dipped in oil. By means of the relative motion between the inner and outer plates, the oil found between them is sheared" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0003883_pomr-2014-0016-Figure3-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0003883_pomr-2014-0016-Figure3-1.png", + "caption": "Fig. 3. Normal force characteristic", + "texts": [ + " For a single panel the hydrodynamic force is calculated using the formula: (4) where: RHD \u2013 hydrodynamic resistance, CF, CP \u2013 friction and pressure force coefficients, VT, Vn, vT, vn \u2013 tangent and normal velocity vectors and absolute values, respectively. Definitions of VT and Vn, are given as: Vn = (V \u00b7 n)n; VT = V \u2013 Vn (5) On the hull of the lifesaving module, a number of bearing nodes were selected which were the objects of action of the ramp pressure forces. The value of the force is defined by the linear function of the crossing distance of the basic (undeformed) ramp surface by the bearing node, see Fig. 3 and Fig. 4. This linear function has been limited by the maximal pressure reaction, introduced to limit modelling of the bearing pressure forces to the range observed in real conditions (for instance due to the loss of stability of the structure, or exceeding the yield point). The pressure force acting on the ramp: (6) where: Espr \u2013 modulus of elasticity [N/m], \u0394s \u2013 distance between the bearing node and the base surface, see Fig. 2, n \u2013 unit normal vector Unauthenticated Download Date | 3/30/18 11:31 AM 37POLISH MARITIME RESEARCH, No 2/2014 The rolling friction force is a linear function of the pressure force" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0002154_j.matpr.2020.10.555-Figure1-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0002154_j.matpr.2020.10.555-Figure1-1.png", + "caption": "Fig. 1. CAD model of Y-s", + "texts": [ + " The details of DED parameters used for the deposition of the structure are given in Table 3. The TS for boundary wall was more than the TS for area filling because higher TS will increase the production rate and also improve the surface finish, but only up to a certain limit. The bead height and bead width, obtained from the DoE, for parameter A was 3.1 mm and 7.16 mm respectively and the same for parameter B was 3.8 mm and 9.6 mm respectively. The CAD model of hybrid Y-shape hybrid frame was developed using SolidWorks 2017 and different views of the model is shown in Fig. 1. The CAD model was sliced by an indigenously designed algorithm. Total number of layer created after slicing was 126 for hape hybrid frame. the maximum height of 250 mm in Z-direction as shown in Fig. 2. The steps involved in slicing of the CAD model is shown in Fig. 3 for better understanding of slicing operation. The CAD model was sliced in multiple layers with an targeted layer height of 2 mm. The targeted height of the layer deposition was obtained through desirability optimization technique using response surface methodology" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0002055_ecce44975.2020.9236026-Figure3-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0002055_ecce44975.2020.9236026-Figure3-1.png", + "caption": "Fig. 3. 3D-airgap machine: (a) concept drawing, (b) exploded view, and (c) 3D-FEA model.", + "texts": [ + " The electromagnetic torque produced in an axial flux machine can be expressed as, Taxial = 4Nphkw(A 2 ro \u2212A2 ri)BgaAa (3) where A2 ro, A 2 ri, Bga, Aa are rotor outer diameter, rotor inner diameter, airgap flux density, electrical loading, respectively. The output torque increases quadratically with respect to the outer diameter. In a 3D-airgap machine, the axial flux component is basically recycling the structural components along with the end turn from the radial flux section. Hence, a volumetric integration can significantly boost the electromagnetic torque. A conceptual drawing for the proposed 3D-airgap machine is shown in Fig. 3(a). The proposed machine is an electromagnetic and structural integration of both radial flux and axial flux machine with a new airgap structure which is continuous starting from radial direction and continues in the axial direction. In a 3D-airgap machine, airgaps for flux flow exist in both radial and axial directions. In this concept, both the radial flux and axial flux machines are developed within the same frame, as shown in Fig. 3. Magnets are placed in the end-plate, acting as the rotor for the axial 36 Authorized licensed use limited to: Carleton University. Downloaded on May 28,2021 at 09:22:22 UTC from IEEE Xplore. Restrictions apply. flux component. The magnetization can be oriented axially or in Halbach configuration. Axially laminated yoke with nonmagnetic teeth is placed beneath the stator core of the radial flux machine. The axial flux path is in the orthogonal direction to the radial flux path as shown in Fig. 3. An exploded view of the 3D-airgap machine is presented in Fig. 3(b). The axial flux machine\u2019s outer diameter has to be adjusted such that the flux lines from the two machines do not interfere or overlap with each other, as presented in Fig. 3(c). Otherwise, it may introduce unexpected harmonics in the torque. Moreover, the axial flux machine\u2019s inner diameter also needs optimization to circumvent the generation of eddy current in the shaft and resulting bearing current. Given that the flux loops from the radial and axial components do not interfere with each other, the electromagnetic torque from the 3D-airgap machine can be expressed as T3D = Trad + Taxial = 4Nphkw[(ro \u2212 ri)LBgA+ (A2 ro \u2212A2 ri)BgaAa] (4) The torque of the 3D-airgap machine is the summation of that of the radial and axial machine" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0003270_icpe.2015.7167862-Figure19-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0003270_icpe.2015.7167862-Figure19-1.png", + "caption": "Fig. 19. 10-pole 12-slot IPM-type BLDCM", + "texts": [ + " In case of demagnetization of non-demagnetization and equality demagnetization, 2nd and 4th harmonics are not appeared. However, in the inequality demagnetization, the more rate of demagnetization of N-S pole is a large difference, 2nd harmonic is occurred, and 4th harmonic is not occurred. In case of N-pole 50% demagnetization result of weighted demagnetization pattern, 2nd harmonic is relatively more appeared. However, 4th harmonic is not still appeared. In addition, in case of 70% demagnetization of N-pole, harmonics of 2nd and 4th are remarkably appeared. C. 10pole-12slot Fig. 19 shows the 10-pole 12-slot IPM-type BLDCM. 10-pole 12-slot IPM-type BLDCM has characteristic from fractional-slot unlike 6-pole 9-slot and 8-pole 12- slot. Namely, this model is different pole-slot combination ratio. In addition, reason of this model selection is to confirm equal BEMF harmonic characteristics from another model of different pole-slot combination ratio. Table IV shows the specification of 10-pole 12-slot IPM-type BLDCM. In addition, Fig. 20 shows the concept figure of 10-pole 12-slot model applying to equality, inequality, and weighted demagnetization patterns" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0002527_iros45743.2020.9341259-Figure4-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0002527_iros45743.2020.9341259-Figure4-1.png", + "caption": "Fig. 4: Collision Cone drawn with respect to the UAV velocity vu.", + "texts": [ + " Therefore, during the intermediate stage the UAV travels in a time optimal collision free manner till the line of sight 5686 Authorized licensed use limited to: Rutgers University. Downloaded on May 16,2021 at 23:16:53 UTC from IEEE Xplore. Restrictions apply. between the UAV and the goal becomes clear. At the post hazard stage the UAV has already avoided the collision and has to find the time optimal path to the goal. As described in the Introduction a collision cone [10] based approach with modifications to find the optimum heading, will be used for the collision avoidance. Fig. 4 shows a typical collision situation represented in the form of a collision cone. The collision cone is drawn with respect to the UAV. At the hazard stage the distance between the UAV and the obstacle is given as d. As in Fig. 4 the two tangents from the obstacle\u2019s position to the UAV\u2019s posterior safety boundary could be taken as P1 and P2. Let \u03b3 be the angle \u2220UOP2 and \u03c8 be the angle between OU and the relative vector ~vou. Definition 1: We can predict a future collision situation when the condition \u03c8 < \u03b3 is satisfied. In order to avoid the collision a heading change which satisfies \u03c8 > \u03b3 should be made by the UAV. There will be many ~vou vectors which would satisfy the condition \u03c8 > \u03b3, however, the most optimum vector according to Theorem 1 would be the vector pointing at point P2", + " A potential collision has been detected with two obstacles O1 and O2. Let Ru be the radius of the UAV, r1 be the radius of O1 and r2 be the radius of O2. If r1 > r2, the UAV\u2019s radius will be taken as Ru + r1 when drawing the collision cone. In a multiple collision scenario collision cones should be drawn for all the obstacles. The difference between the multiple obstacle scenario and the single obstacle scenario is the hazard stage. The other two stages remain similar. It is important to note that the vou vector in Fig. 4 could be directed at OP1 direction or OP2 direction. However, in a single obstacle scenario the relative velocity vector could only be directed at one direction unless the vector is right on OU . The main reason behind this selection is the cost function (5). The angle between vou and OP2 is much smaller than that of vou and OP1. The smaller angle will result a low cost value according to (7). However this selection criteria will not be valid to a multiple collision situation. Therefore, heading change of the relative velocity vector in both clockwise and anticlockwise directions should be considered in a multiple collision scenario" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0002495_ias44978.2020.9334770-Figure1-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0002495_ias44978.2020.9334770-Figure1-1.png", + "caption": "Fig. 1. The structure of measurement system", + "texts": [ + " In order to improve computation accuracy, a hybrid algorithm which combines SA method and NMS technique together is proposed to optimize parameters. Comparison between simulated results and measured ones verifies the validity of the proposed model. II. THE VECTOR HYSTERESIS MEASUREMENT OF SMC MATERIALS The 3-D magnetic property test system consists of a main excitation magnetic circuit, winding units, a power circuit unit, control arithmetic units, and magnetic measurement sensing units. And the overall structure of the system is shown in Fig. 1. Fig. 2 depicts the measurement process of 3-D magnetic characteristics. The NI-PXIe embedded controller is used as the computing core, supplying analog control signal to three high power amplifiers. The excitation voltage signal is selected by the impedance matching loop and the resonant circuit to obtain the large excitation current which is inputted into the main excitation magnetic circuit, the core of the measurement system. Any spatial trajectory of the 3- D magnetic field of the specimen can be generated through the core structure of the yoke" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0003680_s0894-9166(14)60051-3-Figure4-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0003680_s0894-9166(14)60051-3-Figure4-1.png", + "caption": "Fig. 4. A (4,2) chiral single-layered ZnONS subjected to axial load.", + "texts": [ + " Since three bond lengths (ri j 1, ri j 2, ri j 3) and three bond angles (\u03b8i j 1, \u03b8i j 2, \u03b8i j 3) are associated with each atom (See Fig.3), Eq.(2) can be rewritten as Et = \u2211 ij 1 2 K\u03c1 \u2211 k (drijk)2 + \u2211 ij 1 2 C\u03b8 \u2211 k (d\u03b8ijk)2+ \u2211 ij 1 2 C\u03c9 [ 1 3 \u2211 k (d\u03b2ijk) ]2 (4) The constant 1/2 in the first term of Eq.(4) is to ensure that the bond stretching energy is considered only once. By considering the force F as two perpendicular components, fa along the bond r3 and fp perpendicular to it, one can utilize the following geometrical relations to compute them (See Fig.4): fp = F cos (\u03c0 6 \u2212 \u0398 ) (5) fa = F sin (\u03c0 6 \u2212 \u0398 ) (6) in which the chiral angle, \u0398, is computed as \u0398 = arccos ( 2n + m 2 \u221a n2 + nm + m2 ) (7) The following relations are the force and moment equilibrium equations for the system represented in Fig.5, fp sin ( \u03b83 2 ) \u2212 fa cos ( \u03b83 2 ) = K\u03c1dr1 (8) fp cos (\u03b1 2 )(r1 2 ) = C\u03b8d\u03b83 + C\u03b8d\u03b82 cos (\u03a8) (9) where C\u03b8d\u03b83 and C\u03b8d\u03b82 cos (\u03a8) are the moments caused by bond angle variation in plane r1-r2, and by d\u03b83 in plane r1-r3, respectively. \u03a8, the torsion angle between planes r1-r2 and r1-r3, can be obtained as cos (\u03a8) = tan (\u03b83/2) tan (\u03b82) (10) \u03b82 and \u03b83 are geometrically related as cos (\u03b82) = \u2212 cos ( \u03c0 n + m ) cos ( \u03b83 2 ) (11) in which \u03c0/(n + m) is the angle between the bond r3 and the plane r1-r2" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0000350_iecon.2019.8926936-Figure1-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0000350_iecon.2019.8926936-Figure1-1.png", + "caption": "Fig. 1. Geometry of IPM motor.", + "texts": [], + "surrounding_texts": [ + "Keywords\u2014Internal Permanent-Magnet Synchronous Motor, Flux-Weakening.\nI. INTRODUCTION\nInternal Permanent magnet machine is one of the most common typologies employed in the design of power train for electric transportation. It shows: high torque density, high efficiency (due to the absence the magnetizing current and rotor circuit) and high flux-weakening capability [1].\nIn power train application flux-weakening control strategy is a critical point because electrical machines operate in a wide speed range above the constant torque region. The fluxweakening definition is due to the fact that the stator magnetic field is used to weaken the permanent-magnet flux linked with the stator coil. This aim is achieved by increasing the negative d-axis current component till the characteristic current of the machine for an ideal infinite speed range when the characteristic current match the outer limit of the current circle.\nMany papers deal with the optimal magnetic flux quantity to obtain an optimal machine behavior, that can be\nsubstantially synthetized in a machine design that force the characteristic current on the edge of the current limit circle in the id-iq plane, thus avoiding MTPV region to obtain the maximum exploitation of the power train.\nFor electric vehicle application this simple statement must fulfill different duty types with corresponding different current limits, therefore a trade-off has to be found.\nIn this paper the characteristic current is designed to be in the middle of two current limits corresponding to the continuous duty type (S1) and the short-time duty (S2).\nAccording to this design approach a large volume of rare earth permanent-magnets must be employed achieving high machine performance and efficiency, but also high permanent magnet linkage flux.\nWith this design approach, during flux-weakening operation at high speed it is necessary to counteract the magnet Back-EMF with an appropriate d-axis magnetomotive force, in particular a high negative d-axis current has to be fed even if the torque set point is quite low or at least equal to zero.\nDifferent flux-weakening control strategies have been developed in the last years as shown in [2], [3], nevertheless even if this control strategies are strongly proven in motor operation, few studies and examples of what happen during fast transitions of the torque set point and the external load at high speed have been deeply investigated.\nThe control strategy for an electrical vehicle is usually based only on a torque control loop and the driver adjusts the vehicle speed acting on the throttle (torque control). During brake or other fast torque transitions, the torque set point could be moved from its maximum value to zero and, if the motor is in the flux-weakening region, the motor should move its working point either in generating region or in a zero-torque condition.\nDuring the transition of torque set point, the current angle error due to iron losses changes its sign [4] making difficult to manage the energy-flow transition from the motor to the battery to avoid battery unsafe operation or overcharge. Moreover, the vehicle behavior could be heavily affected by the braking torque generated by the motor, especially for motorcycles also a short time braking torque could cause drive injures.\n978-1-7281-4878-6/19/$31.00 \u00a92019 IEEE 2682", + "The author of [5] proposes a flux-weakening control algorithm based on the increase of the id current error, once the voltage limit is reached. This id error is used to decrease the vq axis voltage set point. The control scheme is substantially a modification of the classical feed-forward vector control method. This solution allows a smooth transition from MTPA to flux-weakening control strategies and a fast response. In addition, this method does not strongly rely on the motor parameter and LUT of the magnetic model.\nOther control approaches [6] implement the fluxweakening strategy employing the outer voltage loop to generate an appropriate id current to keep the voltage vector within the range limits imposed by the Voltage DC BUS. The control scheme is aimed to the voltage maximum exploitation, also in this case no specific strategy is deeply investigated to manage the fast torque transition.\nIn paper [7], an enhanced control strategy for high speed permanent-magnet synchronous machines is proposed for electric vehicle application. This control strategy operates from low up to high speeds region employing the same approach of the outer voltage loop adding some additional block with also the MTPV equations. Even in this case fast torque transition issue is not pointed out and no tests on a real vehicle are presented. Other similar control strategies are presented in [8],[9],[10].\nThis paper shows a modified flux-weakening control strategy to appropriately manage fast torque transitions and/or braking torque issue. Two control schemes are presented: the first based on the outer voltage loop and the second on [5] solution. Experimental results of the two control strategies are presented in this paper.\nThe paper is organized as follows: Section II recalls theoretical permanent-magnet synchronous motor control equations and control strategies; Section III describes the main critical issues of common flux-weakening strategies which has been investigated also via experimental tests; Section IV shows the experimental results of the improved control methods and Section V describes experimental test bench and compares machine operating region profiles of the improved methods.\nII. PM VECTOR CONTROL EQUATIONS\nAccording to d-q rotor reference frame, where the d-axis corresponds to the North permanent-magnet (PM) magnetic axis, the steady-state voltage equations of permanent-magnet synchronous motor PMSM are expressed as follows:\nvd= Rsid- eLqiq\nvq= Rsiq+ e PM+ eLdid\nwhere id, iq, vd, vq are d- and q-axis components of stator currents and motor terminal voltages, PM is the PM flux, Rs is the stator resistance, Ld and Lq are d- and q-axis selfinductances. If the motor has no saliency the selfinductances are equals.\nThe terms - eLqiq and + eLdid are the motional terms which couples d-q axis components. These two contributions could be compensated in the control strategy in order to have two independent axes.\nThe torque equation of PM motor is expressed in (2):\nT= 3\n2 p[ PMiq+ Lq-Ld idiq]\nwhere p is the pole pair number. The Fig. 2 shows the machine operating regions: the region I is the constant torque and the region II is the constant power one. Inside the region I the operating limits are imposed by current circle limit, see Eq. (3), while the limit of region II is given by both voltage and current limit circle. The voltage limit circle in the id-iq plane becomes an ellipse, see Eq. (4) and Eq. (5) where m is rotor mechanical speed.\nid 2 + iq 2 \u2264 iLim 2 (3)\nvd 2 + vq 2 \u2264 vLim 2 (4)\nLd 2 iLim+\nPM\nLd\n2\n+ Lq 2iq\n2 \u2264 vLim\np m\n2\n(5)\nCurrent and voltage limits depend on the inverter kVA rating and the battery voltage state. For a given machine, the voltage" + ] + }, + { + "image_filename": "designv11_71_0003975_icinfa.2014.6932788-Figure2-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0003975_icinfa.2014.6932788-Figure2-1.png", + "caption": "Fig. 2. The logarithmic quantizer.", + "texts": [ + " , n) is the output measured by each fault estimator and A, B, D, E, Ci ,Di (i = 1, 2, \u00b7 \u00b7 \u00b7 , n) are all known real matrices with appropriate dimensions. In this paper, we focus on the distributed fault estimator design for system (1). In the distributed scheme, each estimator provides a fault estimation by using its local measurement and the information of neighbors. In the networked structure, the signal must be quantized before transmission due to the limited bandwidth resource. Therefore, in this paper, the logarithmic quantizer in [15] is used to quantize the measurement signal. Its operating principle is depicted in Figure 2 and the set of quantisation levels is defined by: Uj = {\u00b1u (j) i : u (j) i+1 = \u03c1ju (j) i , i = 1, 2, 3 \u00b7 \u00b7 \u00b7} \u222a {u(j) 0 } \u222a {0} 0 < \u03c1j < 1, u (j) 0 > 0 (2) As advocated in [16], the corresponding quantizer Qj can be depicted by the following piecewise function: Qj(\u00b5) = \u03c1iju (j) 0 , 0, \u2212Qj(\u2212\u00b5), if u (j) i 1+\u03b4j < \u00b5 \u2264 u (j) i 1\u2212\u03b4j , \u00b5 > 0 if \u00b5 = 0 if \u00b5 < 0 (3) where scalar \u00b5 is the input signal and Qj(\u00b5) is the corresponding quantized output; \u03c1j is called the quantization density and \u03b4j = (1\u2212\u03c1j)/(1+\u03c1j) is derived as the maximum error coefficient of quantizer Qj , then the quantization error can be defined as: eQj = Qj(\u00b5)\u2212 \u00b5 = \u2206j\u00b5,\u2206j \u2208 [\u2212\u03b4j , \u03b4j ]" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0001057_0021998320920920-Figure9-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0001057_0021998320920920-Figure9-1.png", + "caption": "Figure 9. 0zz in the binders with warps and wefts shown for context.", + "texts": [ + " It should also be noted that the most severe 0xz occurs within the volume of the binder, not on the surface, as seen in the cutaway of the binder in Figure 8. The high HTT in these regions are dominated by the large magnitude of 0xz. The regions with the second most severe stresses in the binders occur where z-aligned binders come close to wefts, such as indicated by label B in Figure 5. For all of these regions, the failure index corresponding to transverse tension, HTT, is highest, as shown in Figure 6. The stress component that contributes the most to HTT where the binders traverse the textile thickness is 0zz. Figure 9 shows 0zz in the binder, and the locations of 0zz concentrations can be seen to match the locations of local maxima of Hmax, shown in Figure 5. When the binders are most closely aligned with the z-axis, the local z-axis of the binders most closely aligns with the global x-axis, which is the direction of load. Consequently, the 0zz concentrations form where load is transferred from the binders to nearby wefts. The transfer of load is clearly shown by contours of xx in Figure 7, such as at label B in the figure" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0001775_0278364920955242-Figure7-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0001775_0278364920955242-Figure7-1.png", + "caption": "Fig. 7. Consider two articulated bodies, B1 and B2 in contact at one or more points. At the ith contact location, the bodies have points 1ci and 2ci, respectively, that are collocated (shown as yellow stars). The contact force applied by B1 on B2 is equal and opposite to the force applied by B2 on B1. Hence, we choose the relative displacement between the points 1ci and 2ci as the contact-space coordinates.", + "texts": [ + " Then, the overall dynamics of the system can be projected into the subspace complementary to the contact space by multiplying with NT C , which yields NT C A\u20acq + b + g\u00f0 \u00de= NT CG \u00f014\u00de because NT C JT C FC = 0. In other words, the contact forces do not affect the null-space motion instantaneously. While we have developed the CSR model so far assuming contact between articulated bodies and the fixed environment, it can be easily extended to the case of two or more articulated bodies in contact. Consider the case shown in Figure 7 where two bodies B1 and B2 are in contact at one or more points. Then, at the ith contact, points 1ci on body B1 and 2ci on body B2 are collocated. In this case, the kinematic conditions for collision, steady contact, and separation depend on the relative motion of the points 1ci and 2ci. Therefore, the contact-space coordinates are also most intuitively defined in terms of the relative motion as xC, i = 2xi 1xi \u00f015\u00de By convention, we choose the normal vector ni to be in the direction from body B1 to B2" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0000126_j.ifacol.2019.10.069-Figure6-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0000126_j.ifacol.2019.10.069-Figure6-1.png", + "caption": "Fig. 6. Overhang angle", + "texts": [ + " 5(a) shows a round part that at the shown layer, the tangent line is too horizontal so if the nozzle needs to be tangent to the surface, it crashes to machine table (Fig. 5(b)). Two solutions can be proposed to solve this problem, tilting the nozzle and sectioning of the surfaces. Addressing overhang angles is a critical matter for AM processes. This is defined as the angle between the slicing direction and the part surface (Ven et al. 2018). The slicing direction is the normal vector of the slicing planes (Eiliat and Urbanic, 2019). Fig. 6 illustrates the overhang angle for a thin wall. 2019 IFAC IMS August 12-14, 2019. Oshawa, Canada 231 232 Hamed Kalami et al. / IFAC PapersOnLine 52-10 (2019) 230\u2013235 For each machine parameter setting-material there is a maximum overhang angle. Exceeding this value causes the material collapse. In this situation either a support structure is needed, or the nozzle needs to rotate to be more tangential to surface to decrease the overhang angle. The former solution is the only solution for 2 \u00bd axis machines (i", + " In this situation either a support structure is needed, or the nozzle needs to rotate to be more tangential to surface to decrease the overhang angle. The former solution is the only solution for 2 \u00bd axis machines (i.e., all powder bed systems). Algorithmically, the tool path generation is simple; however, more production time is required to build the component, more material is consumed in order to build the support structures, and post processing (mostly machining) is required to remove the support structures (Jiang et al. 2018). Fig. 6. Overhang angle The focus of this paper is to address process planning issues for complex thin wall parts having high overhang angles. This is a major issue for bead deposition based AM technologies in industry nowadays. However, large, complex thin walled structures lend themselves to the DMD process. As a case study, two process planning solutions for a thinwalled dome by bead deposition based AM processes is explored. The proposed solution targets determining appropriate partitions for the dome surface and obtaining suitable build directions for each section" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0002993_msec2017-2796-Figure5-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0002993_msec2017-2796-Figure5-1.png", + "caption": "FIGURE 5: CASE STUDY OF A TI-6AL-4V PART", + "texts": [ + " TABLE 2: DATA USED FOR BUILD COST AND TIME MODEL Variable Value Source Layer thickness 0.03 mm J.Mezzetta et al [4] (2016) Scan speed 1250 mm/s Laser spot size 20 \u03bcm Hatch spacing 0.07 mm Density 4.43g/mm3 Porosity 99.5% Recoating time 20 s Machine datasheet Material cost 300 $/kg Supplier\u2019s website Energy cost 0.18 $/kWh Hydro Quebec[13] Indirect cost 53.35 $/h Ruffo (2006) [5] Energy consumption 162.13 kWh/kg Baumers (2012) [6] 5. CASE STUDY: The following section presents a study case of a Ti-6Al-4V part (Figure 5). Mechanical properties, surface roughness, support structure, time and build cost evaluation of three different build orientations are presented. A mechanical properties evaluation module was developed under MATLAB, which can accurately estimate the part performance for different build orientations as shown in Figure 6 using Equation 3 considering the applied load direction. Table 3 presents the results in terms of (%) relative to wrought reference (WR + x %). A surface roughness evaluation module was developed allowing determining the surface finish of each surface and evaluating the average roughness of the part for different build orientation using Equation 4 and 5" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0001581_s00202-020-01082-8-Figure1-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0001581_s00202-020-01082-8-Figure1-1.png", + "caption": "Fig. 1 Schematic view of the experimental setup", + "texts": [], + "surrounding_texts": [ + "A four-pole synchronous motor has been coupled to the studied induction machine to rotate at synchronous speed, and a variable three-phase power supply is connected to its stator terminal as seen in Figs.\u00a01 and 2. Since the induction machine rotates at a synchronous speed, the corresponding slip is zero and the equivalent circuit of the rotor opens as shown in Fig.\u00a03. Therefore, the rotor losses do not affect the calculations. On the other hand, since the machine is rotated by another motor, the windage and the friction losses are supplied by the driver (that here is the synchronous motor) and thus the losses effects are minimized on the test results. From the equivalent circuit shown in Fig.\u00a03, it is apparent that by measuring the stator current Is and the input voltage Vvar , the sum of the magnetizing reactance XM and the stator leakage reactance Xls is obtained as follows The stator winding resistance Rs is obtained from the usual DC test, and its value is 8.8 \u03a9. Meanwhile, the locked-rotor experiment is done, by which the sum of the stator and rotor leakage inductances is derived. In the following sections, the above system is simulated using Multiphysics FE package COMSOL Software, and the magnetizing and leakage inductances are calculated, separately." + ] + }, + { + "image_filename": "designv11_71_0001861_ijat.2020.109871-Figure7-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0001861_ijat.2020.109871-Figure7-1.png", + "caption": "Figure 7 Action diagram of abrasives and spherical iron powder (see online version for colours)", + "texts": [ + " Therefore, an abrasive grain size of 6 \u00b5m has a significant effect of removing cutter marks in cutting, whereas 1 \u00b5m has a slightly reduced polishing efficiency, but yields a satisfactory polished surface. Figure 6 shows the surface roughness Rz for spherical iron powder particle sizes of 2 and 75 \u00b5m, with abrasives of size 1 and 6 \u00b5m. When the particle size of the spherical iron powder was 75 \u00b5m, the surface roughness improved with an abrasive particle size of 1 \u00b5m. Schematic diagrams of the action of the abrasive and spherical iron powder shown are shown in Figure 7. When the particle size of the spherical iron powder is small (2 \u00b5m), the abrasive grains and the spherical iron powder mix randomly and move as the tool rotates. exercise. During that movement, the abrasive grains interfere with the workpiece and remove material. On the contrary, when the particle size of the spherical iron powder is large (75 \u00b5m), many abrasives adhere and roll around the iron powder, such that in addition to the grinding action due to rotation, the spherical iron powder has a rolling grinding action. This is considered to improve the surface roughness. However, the effect of improving the surface roughness is small when the abrasives size is 6 \u00b5m. The SEM images in Figure 7 show the interior of polishing pastes with abrasives of various sizes and constant abrasive mass; with 6 \u00b5m abrasives, the number of abrasives particles smaller by a factor of 36 compared to that with 1 \u00b5m abrasives. This is thought to be because the number of abrasives that adhere to the spherical iron powder and have a polishing action (the number of effective abrasives) has decreased significantly. Therefore, it is considered that the polishing efficiency is low with 6 \u00b5m abrasives. Figure 8 shows the change in surface roughness with respect to the number of revolutions when using abrasives of size 1 \u03bcm and spherical iron powder particles of size 75 \u03bcm, which are the conditions with the highest polishing rate" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0002923_9781119633365-Figure9.11-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0002923_9781119633365-Figure9.11-1.png", + "caption": "Figure 9.11 Zero-voltage turn-on lossless snubbers with energy recovery.", + "texts": [ + " Each of them employs an auxiliary active switch with passive components to form an active lossless snubber, which is used only to create condi tion for achieving ZVS at the switch turn\u2010on transition. Thus, these ZVS convert ers still possess the same steady\u2010state characteristics as those of their PWM converter counterparts. However, the soft\u2010switching PWM converters require aux iliary power switch, increase control complexity in a certain level, and need to establish a soft\u2010switching feature for wide line and load ranges. Figure\u00a09.11 shows an equivalent PWM converter with a ZVS soft\u2010switching cell. This topology differs from a conventional PWM converter by an additional auxil iary switch S1, a resonant inductor L1, a resonant capacitor C1, which includes the output parasitic capacitance of the power switch SM, and a freewheeling diode D1. This resonant branch L1\u2013S1\u2013C1 is active only during a short switching transition time to create a ZVS condition for main switch SM. A ZVS feature can be always performed for one switching direction by utilizing the current source to charge and discharge the resonant capacitor placed across the switch", + "10 Passive soft-switching boost converter with NZCS and NZVS. achieve ZVS. When switching transition is over, the circuit simply acts as a con ventional PWM converter. As a consequence, the converter can achieve soft switching while preserving the advantages of the PWM converters. It should be mentioned that ZVS can be implemented in many ways. From the circuit topological point of view, each active soft\u2010switching converter with ZVS can be viewed as a variation of the equivalent circuit shown in Figure\u00a09.11. By incorporating this type of resonant network, it creates a resonance to achieve ZVS. The ZVS concept can be extended to generate different types of soft\u2010switching converters. Figures\u00a0 9.12\u20139.14 show typical ZVS\u2010PWM topologies. These active ZVS\u2010PWM converters can achieve ZVS on the main switch without significantly increasing voltage and current stresses of the switch and can achieve soft switch ing on the auxiliary switch. Figure\u00a09.12a shows a circuit diagram of the ZVS\u2010PWM soft\u2010switching boost con verter" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0000058_icmae.2019.8881011-Figure8-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0000058_icmae.2019.8881011-Figure8-1.png", + "caption": "Figure 8. Position No 2", + "texts": [ + " ANALYSIS OF RESULTS AND DISCUSSION For the verification of the results, an algorithm with the calculations described above was performed and validated using MatLab\u00ae SimMechanics tool, considering the robot dimensions described in Table III, the simulation in figure 5 and the controller constant K of 10. (6) (7) (8) (9) The initial position of the robot is specified as illustrated in figure 6, with angle at the joints equal to zero. Different trajectories are plotted in order to verify the calculations raised in the inverse kinematics algorithm, starting with the values shown in Table IV, and the results obtained in figure 7. For the verification of the direct kinematics algorithm, the values from Table V and the results obtained in figure 8 are stipulated. Finally, a helical trajectory is traced to verify the behavior of the platform in the face of variations in time. From what is obtained the behavior of figure 9 in angular positions, figure 10 for positions in X, Y and Z, and figure 11 for positions in. . Figure 11 shows a 5-degree oscillation due to the kinematics propagation error present in the platform, therefore a more robust control is necessary, which implies a higher computational cost. IV. CONCLUSIONS The kinematics of the parallel robots are represented using the Jacobian, which is obtained in a homologous way, but inverse of the serial robot, therefore, the notation changes between a serial configuration to a parallel" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0003276_icphm.2015.7245045-Figure2-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0003276_icphm.2015.7245045-Figure2-1.png", + "caption": "Fig. 2. The bearing test stand. (a) Picture of the bearing test stand. (b) Schematic diagram of the test stand.", + "texts": [ + " Finally, we make use of the trained classifier and the diagnostic information of the test set to test the discriminability of the extracted sparse features. The flowchart of fault diagnosis is shown in Fig. 1. To verify the effectiveness and efficiency of the proposed method, we adopt a group of rolling element bearing data, our data is provided by Bearing Data Center of Case Western Reserve University [23] in our experiments. This data set has been widely used in rolling bearing community and considered as a benchmark. The test stand of rolling element bearings is shown in Fig. 2. The test stand of rolling element bearings has three parts. The main component is a three-phase induction motor (Reliance Electric 2HP IQPreAlert motor). A dynamometer is connected to the motor through a torque sensor by self-aligning coupling. An accelerator is mounted on the motor housing at the drive end for sensing vibration. Vibration signals are collected using a 16-channel DAT (Digital Audio Tape) recorder at a sampling frequency of 12 KHz. Single point faults are introduced to the bearings using electro-discharge machining with fault diameters of 0" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0001055_012011-Figure2-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0001055_012011-Figure2-1.png", + "caption": "Figure 2. Model of PMSG studied a fractional slot of 8 slot /18 PMSG", + "texts": [ + " In this research, we focus only to find a feasible stator slot opening width and height of shoe in a fractional (8 pole/18 slot) of PMSG to reduce cogging torque and to minimize of unbalance magnetic pull (UMP) effect. The cogging torque of this type of PMSG fractional is compared and presented. IOP Conf. Series: Materials Science and Engineering 807 (2020) 012011 IOP Publishing doi:10.1088/1757-899X/807/1/012011 In this paper, we consider with the inset magnet rotor type. The initial design is 8 pole / 18 slot of PMSG. The geometry model of shoe height and slot opening width of the PMSG is presented in Figure 1. The result of simulation of magnetix flux distribution illustrated in Figure 2. From the Figure.1, the different slot opening and stator shoe height values are applied to a fractional slot of 8 poles/18 slots of PMSG. In this paper, three steps of slot opening width of 1.60 mm (the largest), 1.55 mm and 1.50 mm (the smallest)). For stator shoe, authors investigated six steps of shoe height 2.3 IOP Conf. Series: Materials Science and Engineering 807 (2020) 012011 IOP Publishing doi:10.1088/1757-899X/807/1/012011 mm (highest), 2.2 mm, 2.1 mm, 2.0 mm 1.9 mm, 1.80 (shortest). The values of slot opening width and shoe height combination are presented in Table 1" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0003656_amr.1095.859-Figure2-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0003656_amr.1095.859-Figure2-1.png", + "caption": "Figure 2 The longitudinal tensile specimen size for T-joints", + "texts": [ + " The shielding gas argon was supplied from each side of the stringer in back of molten pool. The metallographic specimens were prepared after welding, and the microstructure of T-joints was observed by optical microscope. The micro-hardness distribution of T-joints was tested by Wilson Hardness TukonTM2500-6. The circumferential and longitudinal tensile properties for T-joints were tested by Z100 electronic universal testing machine, and the circumferential and longitudinal tensile specimen size is respectively shown as figure 2 and figure 3. The tensile fracture surface was observed aimed to analyze the fracture morphology by SEM. The thermal conditions and composition distribution of molten pool for laser welding is different in different parts of the weld, so different crystallizing morphologies appear in different parts of the weld. The heat affected zone of T-joints of 2060-T8/2099-T83 Aluminum Lithium alloy by laser welding is very narrow. There is fine grain zone between the heat affected zone and the weld, and there is very fine columnar grain comprising the fine grain zone, as shown in figure 4-a and figure 4-b" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0002314_cyber50695.2020.9279112-Figure2-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0002314_cyber50695.2020.9279112-Figure2-1.png", + "caption": "Figure 2. Zoomed in details of the installation of the PAM in the humerus.", + "texts": [ + " Compared with the conventional 2-DOF Delta mechanism, two PAMs were utilized as actuators instead of motors. As shown in Fig. 1, the Delta mechanism included a base, a moving platform, and two parallel arms. Each arm consisted of a humerus, an elbow and a forearm. In the developed 2-DOF Delta mechanism, parallelograms were utilized to prevent the moving platform from rotating. A PAM was installed between the base and each humerus to provide driving torque. Detailed zoomed in illustration of the installation of the PAM was presented in Fig. 2. The PAM was installed on the diagonal of the parallelogram of the humerus. A rotary encoder was installed at point O to measure the rotation angle of the humerus \u03b8. Based on the cosine theorem, \u03b8 could be calculated as: 22 2 0 ( ) arccos . 2 a b L H P ab (1) where a and b were the side lengths of the parallelogram, L0 was the initial length of the PAM, P was the driving pressure of the PAM, and H(P) denoted the nonlinearity between the pressure of the PAM and its resultant contraction. The current length of the PAM could then be calculated as L=L0-H(P)" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0002347_j.compstruct.2020.113512-Figure5-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0002347_j.compstruct.2020.113512-Figure5-1.png", + "caption": "Fig. 5. Boundary conditions and snap-through driving load on a driving patch.", + "texts": [ + " To maintain the connection between the adjacent patches, the panel model is assembled by applying multiple point constraints to the nodes along each linking edge of all patches. The maximum temperature for curing the prepreg carbon fibres is 135 \u00b0C. By cooling down the panel from the elevated temperature to a lower temperature 32 \u00b0C and restricting the central node of the panel, the initially flat grid panel is first deformed into the primary stable state. This stable config- uration is then actuated to another stable state by applying a small point load on the central node of a driving patch depending on concave/convex patch curvature, as illustrated in Fig. 5. During the actuation, the central node of the driving patch should be restricted in x and y direction, and the middle nodes of the straight sides of the driving patch should be fixed in z direction for a while to help the patch to buckle. By applying a small point force at the central node of the driving patch, the driving patch is subjected to enough out\u2010 of\u2010plane deformation so that its strain energy reaches the peak point. After removing the load and constraints while keeping the fixed constraint at the central node of the panel, the panel will remain in the new state if the driving patch does not return to the original state" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0003783_0954406215583522-Figure1-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0003783_0954406215583522-Figure1-1.png", + "caption": "Figure 1. Geometry of POS214 series by manufacturer IRAUNDI.2", + "texts": [ + " They are used for orientation purposes in wind turbine generators, solar trackers or cranes, to mention a few examples.1 There are several types of slewing bearings; the selection of the correct type of bearing is made on the basis of its particular characteristics such as accuracy, operational speed, and rotation resistance, amongst others. In manufacturer catalogues, the slewing bearings are classified in different series for different bearing types, ring geometries, ball diameters, gears, and so on. As an illustrative example, Figure 1 shows POS214 series by manufacturer IRAUNDI S.A. (Bergara, SPAIN), which corresponds to four contact point slewing bearings with balls of \u00d820mm and external gear.2 The rings of the slewing bearing are bolted to the supporting structures; for example, in the case of a wind turbine generator blade bearing, one ring is bolted to the hub and the other one to the blade.3 As the rings and the supporting structures must fit together, generally the first selection of the slewing bearing is carried out based on its mean diameter", + " Having thus preselected the bearing based on its size, it must be verified that it can withstand the operational loads: the axial load Fa, the radial load Fr and the overturning moment M (see Figure 3). In the case of slewing bearings, as it can be deduced from its aforementioned applications, the rotation range and speed is small so the acting loads Fa\u2013Fr\u2013M can be considered to be static. Therefore, the validity of the preselected bearing is verified based on its static loadcarrying capacity. For such purpose, manufacturer catalogues provide selection curves similar to the ones shown in Figure 4, which correspond to POS214 series by IRAUNDI (see Figure 1);2 there are eight curves, one for each of the bearings of the series (see Figure 2). The input data for the curves in Figure 4 are the acting loads M and Fa,eq, where M is the overturning moment and Fa,eq is an equivalent axial load, calculated with a mathematical expression that combines the acting axial and radial loads Fa and Fr. These loads M and Fa,eq are generally multiplied by a magnification factor, whose value depends on the requirements (type and severity of operation, rigidity, Department of Mechanical Engineering, University of the Basque Country (UPV/EHU), Bilbao, Spain Corresponding author: Mikel Abasolo, Department of Mechanical Engineering, University of the Basque Country (UPV/EHU), Alameda Urquijo s/n, 48013 Bilbao, Spain" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0002275_978-981-15-9180-8_15-Figure10-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0002275_978-981-15-9180-8_15-Figure10-1.png", + "caption": "Fig. 10 Comparison of the antimicrobial activity of the G/PV Abiocide antimicrobial nanocomposite films against E. coli and S. aureus", + "texts": [ + " These films had improved antibacterial actions against gram positive and gramnegative bacteria with polymeric biocide loading of 1, 5, and 10wt%. Antimicrobial activity for the loading level of 1, 5, and 10 wt% against gram negative bacteria E. coli was 92.0%, 95.8%, and 97.1%, respectively and 92.3, 99.6, and 99.7% against the gram positive S. aureus was observed. These results revealed that graphene/PVA thin films can be used in the hygienic, drinking and food package applications [43]. A comparative study of antimicrobial activity of these nanocomposites is shown in Fig. 10. Shang, explained fabrication of PVA hybrid films via solution intercalation method with varying filler (functionalized graphene sheets (HDA-GSs)) contents ranging 0\u201310 wt%. Owing to the two main factors, analysis showed the gas barrier properties: thermal stabilities of these composites and polymer chain-segment immobility and detour ratio. Improved barrier properties, reduced solvent uptake, chemical resistance, and flame insulator of polymer hybrids get eased from hindered ways through these nanocomposites [44] Chao et al" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0003915_amm.607.405-Figure2-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0003915_amm.607.405-Figure2-1.png", + "caption": "Fig. 2 The inherent frequency's first order", + "texts": [ + " 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: 128.210.126.199, Purdue University Libraries, West Lafayette, USA-08/07/15,09:33:40) From the modal analysis results, we can optimize the gearbox's structure and avoid its inherent frequency. In this section, we use ANSYS to get the inherent frequency and natural vibration modal of the housing and the gear system. Fig 2 shows the inherent frequency's first order, table2 and 3 respectively shows the 1-10 order inherent frequency and modal. Modal testing instruments include force hammer, YFF-6-025 force sensor, B&K 4382 acceleration sensor, DLF-8 four in one functional amplifier, INV306U-5260 data acquisition instrument, etc. The test frame is shown in Fig. 3. While testing the modal of the marine gearbox, we use the method of single point excitation and multiple point response, and use the complex modal GLOBAL method for modal fitting" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0001266_med48518.2020.9183117-Figure5-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0001266_med48518.2020.9183117-Figure5-1.png", + "caption": "Fig. 5. Dynamic of the drone in simulation", + "texts": [ + " They can be computed as follows T\u03b1 = G\u2020u, where G\u2020 is the pseudo-inverse of matrix G. Then, the actuators inputs are computed as Ti = \u221a (Ti sin(\u03b1i))2 + (Ti cos(\u03b1i))2 (20) \u03b1i = tan\u22121 ( Ti sin(\u03b1i) Ti cos(\u03b1i) ) (21) The nonlinear model and the controller were validated numerically and experimentally for the Quad \u00d7 configuration. The control scheme used in simulations is illustrated in Figure 4. In our numerical validation, the mechanical model is generated using the Simscape toolbox from MatlabTM. This allows us to consider the whole dynamics of the system. Figure 5 illustrates the performance of the aerial vehicle to reach the desired values. In this simulation the quadcopter starts in a classical form (horizontal attitude) and finishes vertically when tilting theirs arms, i.e., the angles for the arms are starting at 0\u25e6 and finishing at +90\u25e6 for the first two arms and -90\u25e6 for the 2 others. The experimental results have been obtained using an innovative platform developed in our laboratory. It is composed by a ROS system with a RaspberryPi and a Navio2 Shield" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0003943_icra.2014.6907115-Figure2-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0003943_icra.2014.6907115-Figure2-1.png", + "caption": "Fig. 2. Image showing the planes defined in [14] and [13]. The robot\u2019s right foot (on the left side of the image) has just lifted during walking. The green plane is the 3D FPE plane, while the red plane is the GFPE plane.", + "texts": [ + " This means that the GFPE plane defined in [14] can be considered a special case of the 3D FPE plane, where none of the links are rotating or moving relative to one another. Since this special case is only valid when the pre-disturbance robot is at rest or translating purely linearly (difficult to achieve in the real world), the plane from the 3D FPE work is used in the remainder of this work. This allows the robot to respond to disturbances not only in the restricted cases of the GFPE plane, but also when it is already moving. The difference between the two planes is shown in Figure 2, just after the robot has lifted its right foot (on the left, in the figure) while walking: the 3D FPE plane has already responded to this change by rotating towards the lifted foot, but the GFPE plane remains aligned with the current direction of linear translation. Once the robot starts to fall towards the lifted foot, the GFPE plane approaches the 3D FPE plane, as the COM begins to move in that direction. Once the horizontal direction of the plane has been calculated, a number of other values must be calculated and then projected onto the plane, to determine the position of the FPE within the plane" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0000765_012008-Figure1-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0000765_012008-Figure1-1.png", + "caption": "Figure 1. Finite Element Mesh of the FDM Model.", + "texts": [ + " SOLID90 element type with 20 nodes was used to generate the mesh of 0.4mm element size. The size of the model was selected to study the results using a single line generation of the part without using infill. The filament part was printed on a bed of size 15 (length) X 10 (width) X 4 (thickness) mm. The bed size was chosen very small to understand the heat transfer behavior between the deposited filament and the bed. The model has 25 layers of material deposition that equals to 10 mm of the part height. Figure 1 shows the meshed solid model of the fabricated part and the heated bed. ICAMPC IOP Conf. Series: Materials Science and Engineering 764 (2020) 012008 IOP Publishing doi:10.1088/1757-899X/764/1/012008 The element birth and death concept in ANSYS R19.2 Workbench was used to activate and deactivate a particular element at a given time and load condition. The birth and death concept has been successfully implemented in simulating temperature gradients in welding and melting of material [20\u201322]. The element birth and death method is applied to the FDM process to better replicate the material deposition process" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0001588_aim43001.2020.9158991-Figure9-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0001588_aim43001.2020.9158991-Figure9-1.png", + "caption": "Figure 9. Actuating unit", + "texts": [ + " (Fig.6). Figure 6. Fabricated PtM device All of the components are made of ABS. The guide shaft and slide potentiometer length are 60mm. We used 50mm helical compression springs with constant 1.8 N/m and with a 35% deflection rate. This means the traveling distance of the PtM slider is about 15-20 mm (Fig 7). Spring linear potentiometer PtM slider 461 Authorized licensed use limited to: University of Wollongong. Downloaded on August 11,2020 at 10:23:43 UTC from IEEE Xplore. Restrictions apply. Fig.9 shows an actuating unit; 4 units are equipped in the actuating part. One stepper motor (Trinamic Motion GmbH) rotates a low frictional lead shaft (6 mm lead) to slide the lead screw, by which two wires (1 mm steel cable) are oppositely pulled or eased. This mechanical design allows holding tension even the motor turns off. Here we consider a way to control the end-point position of the manipulator. It is assumed that information about wire tensions is only available, which is obtained by the proposed PtM" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0001998_j.mechrescom.2020.103624-Figure1-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0001998_j.mechrescom.2020.103624-Figure1-1.png", + "caption": "Fig. 1. (a) Studied cases of the nonuniform beams with the linear width, sinusoidal width, and radical width. (b) Illustration of the assembly of nonuniform beams under different postbuckling modes. (c) Side view of the postbuckling deformation configurations of the nonuniform beams subjected to bilateral constraints. (d) Symbols defined in the theoretical modelling of the bi-walled beams.", + "texts": [ + " For xample, it is desirable to maximize the harvested electrical power or energy harvesting purposes, while the objective in sensing ap- lications is to match the response amplitude to classified damage vents. Given the difficulties, reported in previous studies, of preisely controlling the postbuckling mode transitions, we introduce he nonuniform beam assembly approach to demonstrate a prorammable postbuckling system response. Three beam shapes are investigated, linear, sinusoidal and radial, as shown in Fig. 1 (a). Each individual beam \u2018k\u2019 is fully defined y its length L k , elastic modulus E k , net gap between the bilateral onstraints h k ( x ) thickness t k ( x ), width b k ( x ), nonuniform cross secion area A k ( x ) and moment of inertia I k ( x ). Fig. 1 (b) illustrates the onuniform beam assembly system that will exhibit different bucking mode transitions for the different beams under an external ap- lied load on the system. Fig. 1 (c) provides the side view of the ostbuckling shape configurations for the nonuniform beams sub- ected to bilateral constraints. The change in the buckling modes ith increasing displacement can be demonstrated in four stages: 1) original shape, (2) point contact, (3) line contact, and (4) snaphrough. The beams switch from (1) to (2) when the buckling limit s exceeded under the axial displacement. Increasing the displace- ent, the deformation is more severe, and the point contact is enarged to a line contact. Further, the deformed beam snaps to a igher buckling mode. .1. Geometry variation \u2013 single beam analysis The postbuckling response of a bi-walled single elastic beam ith nonuniform cross section under axial force control was studed and investigated [14] . For a single beam, the system consists of lamped-clamped nonuniform beam positioned between two biateral fixed frictionless walls as presented in Fig. 1 (d) The beam ill be subjected to axial displacement , and the beam has a ength L , elastic modulus E , net gap between the bilateral contraints h ( x ) thickness t ( x ), width b ( x ), nonuniform cross section rea A ( x ) = b ( x ) t ( x ) and area moment of inertia I(x ) = b(x ) t 3 (x ) 12 that ary continuously along the axial direction. Small deformation asumptions are adopted in Euler\u2013Bernoulli beam theory to model he beam under consideration. According to the conducted studies [ 21 , 22 ], the simple local alance of moments equation that governs the buckling behavior w f o a W n s t s w s g a n q r E i c 2 w s fl e b 0 0 0 w M o m 2 w 2 w c s here C 1 and C 2 are the unknown integral constants, and 1 ( X ) and 2 ( X ) represent the linearly independent special solutions for diferent cross section area configurations" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0003343_pamm.201410025-Figure1-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0003343_pamm.201410025-Figure1-1.png", + "caption": "Fig. 1: 3DOF Robot", + "texts": [ + " KGaA, Weinheim 1 Spatial Path Planning and Dynamic Modeling replacemen A geometric path as well as the robots mathematical model are the foundations for the optimization. There are many different ways to define geometric paths, like line - curve combinations, polynomials, or splines. Latter ones are used in the present paper to define a path zE(\u03c3) = \u2211 j djR d j (\u03c3) in world coordinates as a function of a scalar path parameter \u03c3. Rdj are the Bernstein polynomials of degree d representing the basis functions. The control points dj determine the path\u2019s shape, shown in Fig. 1. The inverse kinematics provides the joint angles q(\u03c3) = f(zE(\u03c3)) as a function of the endeffector coordinates. A dynamic model of the robot is necessary on the one hand to be able to simulate the robot system and on the other hand to compute the feed-forward torques used for optimization and control. Its calculation is done with the help of the Projection Equation (see [4]) , resulting in the equations of motion M(q)q\u0308 + g(q, q\u0307) = Q. This section shows an approach to calculate an optimal movement on a geometrical defined path", + " A gradient z\u2032o fulfilling all restrictions is stored for this point (i, j) and the cost function is adapted to Wi,j = Wo + \u2206\u03c3/ \u221a zi,j . After completing this procedure for every point in the discrete [\u03c3, z] plane, the optimal velocity trend can be calculated iteratively zi+1,opt = zi,opt + z\u2032i,opt\u2206\u03c3. With zopt on hand the time vector is obtained by ti = \u2211i k=1 \u2206tk, with \u2206tk = \u2206\u03c3/zk,opt. This leads to a path parameter trend over time \u03c3(t) representing the optimization result. For validation purposes the algorithms are implemented on a Catalyst CRS robot, shown in Fig. 1. The optimization results for a straight line in space (l \u2248 0.74m) are provided by Fig. 5 in terms of an optimal velocity profile zopt and \u03c3(t). Measurements of the motion quantities are shown in Fig. 6, and of the appropriate torque and torque derivatives in Fig. 7. One can see that most of the time the velocity restrictions are active. In the acceleration and deceleration phases jerk and acceleration constraints are active. The cycle time for this experiment results to tE \u2248 1.1s while a cycle time of tE \u2248 0" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0003883_pomr-2014-0016-Figure1-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0003883_pomr-2014-0016-Figure1-1.png", + "caption": "Fig. 1. Simplified picture of the system for evaluating people from the ship: a) side view of the ship, b) stern view of the ship with driving system components, c) shaft cross section with chain hoists and boats, and the visible ramp/door opening mechanism, d) top view of the boat placed on the hoist in the shaft with marked nodes of its seating on the hoist [3]", + "texts": [ + " \u201cSafe Abandoning of Ships - improvement of current Life Saving Appliances Systems\u201d, which was carried out in years 2004-2009 and the participant of which was the Faculty of Ocean Engineering and Ship Technology, Gdansk University of Technology. Within the framework of this project a number of concepts of novel LSA systems were developed, as described in [1, 2, 3] and [4]. In two of those systems which look most interesting an open stern ramp/door was applied, over which the launched lifesaving boats or other watercraft units with people inside are to slide into water in the direction opposite to the motion of the endangered ship. One of these systems, shown in Fig. 1 and described in detail in [3], provides opportunities for fast and safe evacuation of people using only gravitational forces. The rescued people get simultaneously into all lifeboats situated on different decks, after which a system is started to lower the boats at a controlled speed. When the boats reach the level of the ship\u2019s slip, they are hooked in, one by one, and freely slide down the ramp to the water. In the second system, developed by Fassmer and schematically shown in Fig. 2, the watercraft units are situated in rows on the lower deck" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0002176_icsmd50554.2020.9261740-Figure3-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0002176_icsmd50554.2020.9261740-Figure3-1.png", + "caption": "Fig. 3 Schematic of the contact relationship between the roller and the middle\\late localized defect.", + "texts": [ + " Here, the inner ring is fixed on the wheelset and the outer ring rigidly connects with the axle box. The time-varying displacement excitation and the time-varying stiffness excitation of the failure bearing can be described by an improved analytical model proposed by Liu et al. in Ref. [1]. When a localized defect of the bearing is in the middle or late stage of plastic deformations at the propagated defect edges, the schematic of the contact relationship between the roller and the defect region are shown in Fig.3. It can be seen that there are some visible differences of the relative radial displacement and the contact stiffness between the roller and the outer race. Therefore, the impact interaction between the roller and the defect region can be described as the comprehensive effect of the time-varying displacement excitation and the time-varying contact stiffness excitation induced by the localized defect on the outer race. For a localized defect with the cylindrical surface edges, the time-varying displacement excitation of the localized defect can be represented as a piecewise function [13]: ( )( )( ) ( ) ( )( )( ) ( ) ( ) 1 o 0 o 1 1 o 2 2 o 2 o 3 sin 0" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0001647_j.compstruct.2020.112870-Figure4-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0001647_j.compstruct.2020.112870-Figure4-1.png", + "caption": "Fig. 4. (a) Global coordinate system of the bio-ITCLC sandwich panel; (b) FE-predicted model.", + "texts": [ + " The other one is the supporting cores\u2019 zero or negative Poisson\u2019s ratio mostly in one direction, which is possibly used as references for a flexible skin with single morphing. However, most of the reported support cores may be hard to be employed in applications such as a flexible skin where multi\u2010directional and combined morphing functions are required. Thus, designing a new type Nomenclature bio\u2010ITCLC core biomimetic Isosceles Trapezoid Corrugated Lattice Cellular core part1, part2 and part3 components of one unit of the bio\u2010ITCLC core in the Fig. 3 x, y and z\u2010axis global coordinate system in Fig. 4a kft and kfn tangential stiffness and normal stiffness of the skin Fcr buckling load of the skin H and S Hoffman failure coefficient and strength of the skin kfz s and \u0394kf \u03c3 z\u2010axis structure stiffness and initial buckling stress stiffness of the skin \u03bb ratio kfz s/\u0394kf \u03c3 tf thickness of the skin tp2 and hp2 thickness and height of the part2 hp1=3 and ap1=3 height and upper\u2010side length of the part1 and part3 kx#p1=3/k y# p1=3/k z# p1=3 x/y/z\u2010axis stiffness of the part1 and part3 kx#p2 /k y# p2 /k z# p2 x/y/z\u2010axis stiffness of the part2 kx#f /ky#f /kz#f x/y/z\u2010axis stiffness of the skin Ry z p1=3(R y x p1=3) When the coefficient jRy z p1=3j (jRy x p1=3j) increases, the stiff- ness ky#p1=3 decreases and simultaneously the stiffness kz#p1=3 (kx#p1=3) increases", + " The ratio between the total load ffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi Fx#2 f \u00fe Fy#2 f \u00fe Fz#2 f q corresponding to the skin\u2019s fracture position and the stress gauge\u2019s projected area in the direction of the total loadffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi Fx#2 f \u00fe Fy#2 f \u00fe Fz#2 f q is the strength S. The bio\u2010ITCLC sandwich panel is made up of the part1, part2, part3 and top/bottom skins. In the Cartesian coordinate system as shown in Fig. 4a, the orientation of fibers in the part1 and part3 is along the y\u2010 axis direction, in the part2 is along the z\u2010axis direction, and in the top/ bottom skins is along the x\u2010axis direction. Thus, the constitutive material parameters of four components of the bio\u2010ITCLC sandwich panel are listed in Table 2. The normal pressure load F is prescribed for the upper surface of the top skin, whereas the lower surface of the bottom skin is fixed support. Also, Frictional algorithm, which gives the critical sliding and separating forces, is used as contact constrains between top/bottom skins and the bio\u2010ITCLC core. The FE\u2010predicted model can be seen in Fig. 4b. In the commercial software ANSYS, the mechanical enterprise is used to calculate the buckling load Fcr , x/y/z\u2010axis stiffness kx#f /ky#f /kz#f and strength S. For the buckling load Fcr ; the iteration computation is ended when the z\u2010axis deformation of the top skin appears an obvious inflection point. The total load of the top skin, corresponding to the inflection point time, is the buckling load Fcr . For the x\u2010axis stiffness kx#f (y\u2010axis stiffness ky#f ) (z\u2010axis stiffness kz#f ) (strength S), the iteration computation is ended, when the x\u2010axis deformation (y\u2010 axis deformation) (z\u2010axis deformation) (Hoffman failure coefficient) of the top skin, represented by dx#f (dy#f ) (dz#f ) (H as shown in the formula (3)) reaches a reference value" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0002923_9781119633365-Figure3.29-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0002923_9781119633365-Figure3.29-1.png", + "caption": "Figure 3.29 (a) Two diodes with common node \u00a0, (b) P-type grafted diode (PGD), and degenerated PGD under (c) VX\u00a0> VY, (d) VX\u00a0< VY, and (e) VX\u00a0= VY.", + "texts": [ + " However, it might require two blocking diodes, DB1 and DB2, to block the voltage difference between Vx and Vy, as shown in Figure\u00a03.28b, namely, N\u2010type grafted diode (NGD). If voltages Vx\u00a0> Vy, diode DB1 is always in forward bias, and it can be shorted, as shown in Figure\u00a03.28c. On the contrary, if voltages Vx\u00a0< Vy, diode DB2 can be shorted, as shown in Figure\u00a03.28d. If voltages Vx\u00a0= Vy, both DB1 and DB2 can be shorted, and there is only D12 left, as shown in Figure\u00a0 3.28e. Similarly, two diodes share a common node P, as shown in Figure\u00a03.29a, which can be grafted and degenerated into the configurations shown in Figure\u00a03.29b\u2013e. The one shown in Figure\u00a03.29b is named as P\u2010type grafted diode (PGD). Figure\u00a03.25 illustrates the derivation of boost\u2010buck (\u0106uk) converter with TGS. It can be also derived with a grafted diode, as shown in Figure\u00a03.30, in which the boost and buck converters connected in cascade and with switch S2 in the return 3 Fundamentals70 (a) (b) (c) (d) (e) path are shown in Figure\u00a03.30a. It can be seen that diodes D1 and D2 are with a common node N, and they can be replaced with an NGD, as shown in Figure\u00a03.30b. Since the reverse bias voltages VX and VY of the original diodes D1 and D2, respectively, are all equal to the voltage across capacitor C2, blocking diodes DB1 and DB2 are no longer needed, and the circuit can be simplified to the one shown in Figure\u00a03" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0001303_s00202-020-01035-1-Figure3-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0001303_s00202-020-01035-1-Figure3-1.png", + "caption": "Fig. 3 Magnetic flux distribution in the investigated VPM machine", + "texts": [ + " In this section, the presented analytical model is used to study the no-load/on-load magnetic flux density, cogging torque, electromagnetic torque, back-electromotive force, self-inductance and mutual inductance of a prototype VPM machine. The results of the analytical method are then verified by the results of the finite element method and experimental tests. (51)L = a IA (52)M = N AB IB The VPM machine parameters are given in Table\u00a01. 2-D finite element method is applied to calculate the investigated machine performance. Magnetic field distribution in the VPM machine is represented in Fig.\u00a03. No-load comparison of analytical and numerical results of radial/tangential flux density and cogging torque of the investigated machine is shown in Figs.\u00a04 and 5, respectively. On-load analytical and numerical comparison of radial flux density and self-/mutual inductance in the studied machine is shown in Figs.\u00a06 and 7, respectively. This section first presents the manufacturing stages of the various parts of the proposed VPM machine and then describes the experimental tests. The materials and dimensions of the studied machine parts are listed in Table\u00a02" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0002220_icee50131.2020.9261061-Figure2-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0002220_icee50131.2020.9261061-Figure2-1.png", + "caption": "Fig. 2. Machinery fault simulator components and 14 sampling modes", + "texts": [ + " A collection of six data sets related to the unbalanced fault on six discs, a dataset related to the combination of two unbalanced faults on the discs 1 and 4, three dataset related to the defect of the ball bearings, three datasets related to the fault of outer ring of the bearing and finally a healthy dataset (a total of 14 modes) were provided and used in the present work. Ten trials were conducted from each of 14 modes, each in five seconds of duration per test. Different components of the machinery fault simulator and the 14 modes of sampling process are shown in Fig. 2. In this paper, different methods were used in feature extraction, reduction of data dimensions and classification phases. In order to extract the effective features, the EMD decomposition method was first used to obtain the features for the IMFs of the signal. In the feature extraction section, a number of features that had been proven in previous studies were used. PCA method was adopted for data dimension reduction in consequence of the high dimensions of the EMD decomposition and a large number of features, obtained due to the number of sensors" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0002923_9781119633365-Figure9.2-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0002923_9781119633365-Figure9.2-1.png", + "caption": "Figure 9.2 Proper locations of snubber inductor L1 for PWM converters to achieve NZCS: (a) connected to a freewheeling diode and (b) connected to an active switch.", + "texts": [ + "1 shows the waveforms of switch gate signal VG, switch current IDS, and switch voltage VDS, illustrating a zero\u2010current switching mechanism. However, according to Figure\u00a09.1, the main requirement of NZCS turn\u2010on or turn\u2010off is to slow down the current rising time or falling time, respectively, during switching transition. This NZCS feature can be achieved by placing a snubber inductor L1 in series with the loop containing the active switch or freewheeling diode so that the switch current is limited to rise or fall slowly at switching transition. Figure\u00a09.2 shows an equivalent circuit of a PWM converter with possible locations of the snubber inductor L1. Thus, the rate of di/dt of the switch current is restricted, and the switching transition is close to zero\u2010current transition, namely, NZCS or ZCT. Snubber inductor L1 will store energy that needs to be recovered ultimately and delivered to the output load or the input voltage source to maintain lossless opera tion. In this case, the snubber cell must provide a conduction path when a free wheeling diode or a switch turns off" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0001140_1.i010790-Figure11-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0001140_1.i010790-Figure11-1.png", + "caption": "Fig. 11 Engagement geometry between a point and an arbitrarily shaped object.", + "texts": [ + " The guidance laws developed in the preceding sections assumed that the shape of the intruder swarm is bounded within a circle. In cases where the intruder swarm has a somewhat elongated shape, then the circular approximation can represent an over-approximation to the shape of the swarm, and this can be therefore inadequate. In this section, we show how the guidance laws can be modified when the shape of the intruder swarm is arbitrary. We provide a brief background result on the collision cone between a point and an arbitrarily shaped object [44]. Refer Fig. 11. As shown in [44], a point object A is on a collision course with an arbitrarily shaped objectB if the condition y < 0; Vr;b < 0 is satisfied, where y and Vr;b are defined as follows: y V\u03b8 cos \u03b8b \u2212 \u03b8 \u2212 Vr sin \u03b8b \u2212 \u03b8 2 V2 \u03b8 V2 r \u2212 sin2 \u03c8 2 (53) Vr;b Vr cos \u03b8b \u2212 \u03b8 V\u03b8 sin \u03b8b \u2212 \u03b8 (54) In the above equations, \u03c8 represents the angle subtended at A by the tangent linesAT1 andAT2 toB.T1AT2 represents the conical hull ofB, relative toA, and \u03c8 thus represents the angle of this conical hull. When there exist more than two tangent lines toB, thenAT1 andAT2 represent the pair of tangents that are such that B is completely contained within the sector T1AT2" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0003773_j.jweia.2015.01.001-Figure5-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0003773_j.jweia.2015.01.001-Figure5-1.png", + "caption": "Fig. 5. (a) Schematic trajectory for a particle with \u03a9{1 showing the initial and final horizontal and vertical velocities, trajectory (solid line), and projected trajectory (dashed line) originating at the virtual release height \u03b6v above the actual release height. (b) Schematic trajectory for a particle with \u03a941 showing same details but with the virtual release height below the actual release height.", + "texts": [ + " Even the largest boundary layer wind tunnels are not this tall and, as such, are unable to achieve a steady state trajectory for even a small grain of sand. As shown above, following an initial adjustment the debris particle will travel in a straight line of slope S\u00bc ffiffiffiffiffi \u03a9 p . Therefore, in the far field, the flight trajectory can be described by a straight line that appears to come from a height given by the release height plus a virtual release height offset \u03b6v where if \u03b6v is positive then the virtual release height is above the actual release height and is below the actual release height if \u03b6vo0. This is shown schematically in Fig. 5. The horizontal flight distance for large \u03b6 is, therefore, given by \u03c7 \u00bc \u03b6 \u03b6vffiffiffiffiffi \u03a9 p \u00f021\u00de where \u03b6 is measured from the release height and is positive upwards. For\u03a9o1 the initial slope is less than the final slope (\u03a9o ffiffiffiffiffi \u03a9 p ), whereas for \u03a941 the initial slope is greater than the final slope. Intuitively, one might therefore expect that, for \u03a9o1, the virtual release height would be above the actual release height (\u03b6v40), as shown in Fig. 5(a). Conversely, for\u03a941 the initial trajectory slope is greater than the final slope and one would expect the virtual release height to be below the actual release height (\u03b6vo0) as shown in Fig. 5(b). To test this hypothesis a series of 600 numerical flight simulations over 6 orders of magnitude in \u03a9 were conducted and the virtual release height offset was calculated by fitting a line through the flight trajectory late in the flight. A plot of the calculated virtual release height offset as a function of\u03a9 is shown in Fig. 6(a). For \u03a941 the virtual release height is negative, that is, the trajectory projects back to below the actual release height as shown in Fig. 5(b) and as hypothesized in the previous paragraph. Conversely, for \u03a9{1 the virtual release height is above the actual release height (see Fig. 5a). However, for 0:3o\u03a9o1, the virtual release height is below the actual release height, that is, \u03b6vo0. Therefore, for this range of \u03a9, the trajectory slope must initially steepen beyond the steady state slope and then approach the steady-state trajectory from a steeper rather than shallower slope. This implies that the trajectory has an inflection. This is shown schematically in Fig. 6(b). Also, Fig. 6(a) shows that for large \u03a9, \u03b6v approaches a constant whereas, for small \u03a9, \u03b6v remains as a function of \u03a9" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0003337_9781118886397.ch1-Figure1.9-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0003337_9781118886397.ch1-Figure1.9-1.png", + "caption": "Figure 1.9 Three-dimensional vectors in rectangular form.", + "texts": [ + " Cartesian coordinates in three dimensions are members of the set \u211d3 and are similar to those in two dimensions, with the addition of the vertical z-axis and the unit vector k\u0302 in the positive z-direction. We have already entered into the third dimension with the differential surface vectors. The unit vectors can be written as i\u0302 = (1, 0, 0) (1.35) j\u0302 = (0, 1, 0) (1.36) k\u0302 = (0, 1, 0) (1.37) The Cartesian coordinates of a three-dimensional vector (a, b, c) are defined as B = (a, b, c) where a and b are the x and y dimensions of the projection of the vector B on the x\u2013y plane and c is the height of the end of the vector B above the x\u2013y plane. The vector B = (4, 5, 6) is shown in Figure 1.9, the vector A = (4, 5) is the projection of the vector B on the x\u2013y plane. The magnitude of a vector B = (a, b, c) in three-dimensional Cartesian coordinates is B = \u2016B\u2016 =\u221aa2 + b2 + c2 (1.38) The magnitude of the vector B = (4, 5, 6) is B = \u221a 42 + 52 + 62 = 8.775 In the three-dimensional Cartesian coordinate system, a differential length vector dl may be defined from the differential lengths dx, dy, and dz as dl = dx\u0302i + dy\u0302j + dzk\u0302 (1.39) In the three-dimensional Cartesian coordinate system, three differential area vectors dSx, dSy, and dSz, may be defined from the differential lengths dx, dy, and dz as dSx = dy dz dSy = dx dz dSz = dx dy \u23ab\u23aa\u23ac\u23aa\u23ad (1" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0001842_mic50194.2020.9209618-Figure1-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0001842_mic50194.2020.9209618-Figure1-1.png", + "caption": "Fig. 1. Mobile robot and its dynamic target tracking problem setup.", + "texts": [ + " The model free actor-critic reinforcement learning algorithm and its associated mathematical derivations are detailed in section III. Section IV illustrates computer experiments for different scenarios that reflect the effectiveness of the proposed method followed by conclusion and future work presented in section V.978-1-7281-8386-2/20/$31.00 c\u00a92020 IEEE Authorized licensed use limited to: Middlesex University. Downloaded on November 06,2020 at 06:01:39 UTC from IEEE Xplore. Restrictions apply. Fig. 1 shows the setup of the problem that we address in this manuscript. The robot is characterized by its 2D position (xk, yk) and orientation \u03b8k \u2208 [\u2212\u03c0, \u03c0) with respect to the global coordinate frame X-Y at discrete time index k = 0, 1, . . . , where time t = kT with T > 0 being the sampling time. The time-varying position of the target is denoted by the position vector p [d] k = [x [d] k , y [d] k ]T \u2208 R2. Let the position and orientation (pose) of the robot is denoted by the vector qTk \u2261 [xk, yk, \u03b8k] \u2208 R2 \u00d7 S1", + " Without loss of generality, assume that the robot tracks the target trajectory while maintaining line-of-sight distance (also known as the safe distance between the robot and the target) d > 0. We emphasize that the mathematical models of both robot and the target are not required in the proposed control technology. For illustration, let us assume that the robot (rear-wheel drive) follows the discrete-time kinematic model described by xk+1 = xk + T\u03bdk cos \u03b8k, (1a) yk+1 = yk + T\u03bdk sin \u03b8k, (1b) \u03b8k+1 = \u03b8k + T \u03bdk L tan \u03b3k, (1c) where \u03b3k \u2208 (\u2212\u03c02 , \u03c0 2 ) is the steering angle of the font wheel/castor of the robot (see Fig. 1), and \u03bdk \u2208 R is the linear speed. Without loss of generality, this work considers a target modeled by an integrator. The discrete-time kinematic model of the target is given as p [d] k+1 = p [d] k + Ts u [d] k , (2) with u [d] k \u2208 R2 being the control input vector (velocity) that defines the trajectory (considered random in this work) of the target. In case, the robot modeled by (1) is described by a differential drive mobile robot, then its left-wheel velocity \u03bdL,k and right-wheel velocity \u03bdR,k are related by \u03bdk = 1 2 (\u03bdR,k + \u03bdL,k), (3a) \u03c9k = 1 ` (\u03bdR,k \u2212 \u03bdL,k) = \u03bdk L tan(\u03b3k) (3b) where \u03c9k = \u03bdk L tan \u03b3k is the angular velocity of the robot\u2019s center (xk, yk). Let us define the error vector as ek = [ \u03c1\u0303k, \u03b8\u0303k ]T = [\u221a x\u03032 k + y\u03032 k, \u03b8\u0303k ]T (4) where \u03b8 \u2032 k = atan2 ( y [d] k \u2212 yk, x [d] k \u2212 xk ) and \u03c1\u0303k = \u221a x\u03032 k + y\u03032 k is the Euclidean distance (position error) at time instant k as shown in Fig. 1 with x\u0303k = x [d] k \u2212 xk \u2212 d cos \u03b8 \u2032 k, y\u0303k = y [d] k \u2212 yk \u2212 d sin \u03b8 \u2032 k, and \u03b8\u0303k = \u03b8 \u2032 k \u2212 \u03b8k is the orientation error. The control problem can then be formally stated as follows: Find \u03bdk and \u03b3k such that ek \u2192 0 as k \u2192\u221e subject to (1) and (2). Reinforcement learning is a promising machine learning technique that is widely used for solving problems of many cyber-physical systems that pertain to goal-directed learning from interaction [18]. In this paper, tracking a mobile target using a wheeled mobile robot is formulated as a reinforcement learning problem, where the goal of the robot is to track the trajectory of a target while maintaining the safe-distance d > 0" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0000825_s41315-019-00114-2-Figure9-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0000825_s41315-019-00114-2-Figure9-1.png", + "caption": "Fig. 9 Control workflow of the Arduino robot", + "texts": [ + " We need to stick our motor shield above the Arduino Uno board (Fig.\u00a08). The L293D motor shield is able to drive servo motors, stepper motors or DC motors. In the case of Arduino robot, the L293D motor drive shield is driving 2 DC motors installed on 2WD Smart Robot Car Chassis Kit. The DC motor has no polarity. There is no positive and no negative. If we reverse our motor connections, the motor\u2019s direction of movement will change. The overall workflow of Arduino robot platform is illustrated with Fig.\u00a09. The Arduino Uno reads in the distance signals of HC-SR04 sensor and the (optional) commands 1 3 from BT06 Bluetooth module. It controls the DC motor by sending the PWM (pulse width modulation) signals to the L293D motor drive shield board. The debug information is sent to host computer through Arduino Uno board\u2019s #0 (RX) and #1 (TX) serial port. Image capturing and video streaming is supported by the attached mobile phone. The images and video signals are captured by the android phone and sent to the host computer under WIFI connection" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0001138_0954406220925836-Figure1-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0001138_0954406220925836-Figure1-1.png", + "caption": "Figure 1. Three types of articulated moving platforms: (a) AMP1; (b) AMP2; (c) AMP3.", + "texts": [ + " Articulated moving platform about one rotation Generally, if the end-effector of a mechanism has one rotational DOF about a defined direction, there should exist at least two parallel (or collinear) rotational axes or one internal rotational axis within the connected link. According to different characteristics of performing rotations, the articulated moving platforms can be classified into three types, i.e. AMP1, AMP2, and AMP3. AMP1: the rotation is achieved by two coaxial revolute joints, and the two revolute joints can only be arranged at the two ends of the end-effector, as shown in Figure 1(a). If the coaxial revolute joints are assembled on the moving platform of a mechanism, the end-effector will possess a local DOF. As a result, the coaxial revolute joints can only be installed on the fixed base and actuated by electric rotating machinery. AMP2: the rotation is realized through two parallel revolute joints and one planar translation, which is perpendicular to the axis of the revolute joint, as drawn in Figure 1(b). Taking the planar four-bar linkage as an example, the output motion of the end-effector is a rotation about a fixed axis. One of the two revolute joints and the two-dimension translation can be constructed by RRR, PRR, RPR, or PPR structures, where R and P signify the revolute joint and the prismatic joint, respectively. For a 3-DOF six-bar planar linkage, the rotational axis is changeable within the workspace resulting from two planar translational motions. In summary, to guarantee the mechanism possess an expected rotational performance, at least one of the two connected limbs can perform a planar translation perpendicular to the axis of the rotation. AMP3: when arranging a single revolute joint at the intermediate moving platform of the mechanism, the high rotation capacity in one direction is obtained, as depicted in Figure 1(c). The motion is actuated by the limb, which is directly connecting the revolute joint, and the connecting link is named as the intermediated moving platform. The distinguishing feature of this kind of articulated moving platform is that the rotation is decoupled from other motions. Articulated moving platform about two or three rotations After combining the above designed articulated moving platforms, the rotations in two or three directions can be derived. When two relatively independent rotations are required, the feasible platforms are composed of AMP1[AMP2, AMP2[AMP2, AMP1[AMP3, and AMP2[AMP3, where the symbol \u2018\u2018[ \u2019\u2019 indicates the combination of platforms" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0000642_s12206-020-0134-3-Figure3-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0000642_s12206-020-0134-3-Figure3-1.png", + "caption": "Fig. 3. Geometry of the RR dyad.", + "texts": [ + " N T j j j j j\u03b8 X Y j N\u00ba =r r L (1) However, in order to reduce the complexity of the synthesis approach, it is more useful to define the set of prescribed poses with respect to the first pose, that is: { } ( ) ( )0 0 01 0 , , , , , , 1,2, , . N T T i i i i i i i i i i x y X X Y Y for i N f f q q \u00ba - = = - - \u00ba - = p p r r L (2) Thus, for a set of four prescribed poses 3,N = whereas 4N = if we have five poses to be visited. We will begin the synthesis process with the derivation of the design equations associated with the RR dyad. To this end, Fig. 3 shows the linkage in two finitely separated poses, namely, pose 0, and arbitrary pose i. Basically, point B changes from 0B to ,iB point P changes from 0P to ,iP and point A remains fixed. The design equation for the RR dyad is based on the fact that distance between points A and iB must remain constant. This can be mathematically formulated as: ( ) ( ) ( ) ( )0 0i i- \u00d7 - = - \u00d7 -b a b a b a b a (3) where from the geometry of Fig. 3, the position of point B at the i-th pose can be computed as follows: 0/ 0 0ii B P i i i i\u00ae\u00ba = + = +b r p p R b\u03c1 (4) being 0 i\u00aeR a matrix that describes the rotation of coupler link 2, from pose 0 to pose ,i which is given by: 0 cos sin sin cos i i i i i \u00ae -\u00e9 \u00f9 = \u00ea \u00fa \u00eb \u00fb R f f f f (5) It should be noted that position vector 0b moves with body 2. Hence, vector 0b becomes 0 0i i\u00ae= R b\u03c1 at i-th pose. Substituting Eq. (4) into Eq. (3) and, after some algebra, it is obtained the following design equation for the RR dyad: ( ) ( )0 0 0 02 2 0T T i i i i i\u00ae \u00ae+ + \u00e9 - - \u00f9 =\u00eb \u00fbp p R b a I R b p (6) Eq" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0002485_cac51589.2020.9327292-Figure6-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0002485_cac51589.2020.9327292-Figure6-1.png", + "caption": "Fig. 6: Physical image of aircraft body: 1. Battery 2. Motor 3. Steering engine 4. Steering engine drive circuit 5. Electronic speed control", + "texts": [ + " From the relationship between (21)-(23) transfer function input and control system input (\u03c9i, \u03b1i), combined with the assumption of small disturbance linearization, we can get:\u23a1\u23a2\u23a2\u23a2\u23a2\u23a2\u23a2\u23a2\u23a2\u23a2\u23a2\u23a2\u23a2\u23a3 u\u03c6 u\u03b8 u\u03c8 F \u23a4\u23a5\u23a5\u23a5\u23a5\u23a5\u23a5\u23a5\u23a5\u23a5\u23a5\u23a5\u23a5\u23a6 = A \u23a1\u23a2\u23a2\u23a2\u23a2\u23a2\u23a2\u23a2\u23a2\u23a2\u23a2\u23a2\u23a2\u23a3 c\u03b11 CT\u03c9 2 1 c\u03b13 CT\u03c9 2 3 s\u03b11 CT\u03c9 2 1 s\u03b13 CT\u03c9 2 3 \u23a4\u23a5\u23a5\u23a5\u23a5\u23a5\u23a5\u23a5\u23a5\u23a5\u23a5\u23a5\u23a5\u23a6 (25) where F is the total pulling force produced by the propeller, and A satisfies: A = \u23a1\u23a2\u23a2\u23a2\u23a2\u23a2\u23a2\u23a2\u23a2\u23a2\u23a2\u23a2\u23a2\u23a3 0 0 \u22121 1 1 \u22121 0 0 0 0 1 1 1 1 0 0 \u23a4\u23a5\u23a5\u23a5\u23a5\u23a5\u23a5\u23a5\u23a5\u23a5\u23a5\u23a5\u23a5\u23a6 Then A\u22121 can meet the conversion from the output of the controller to the input of the controlled object. IV. Control system simulation Generally, the maximum angular velocity of the quad-rotor is about 3rad/s, and the natural frequency \u03c9n is smaller than the maximum response velocity of the airframe. When P2 and P4 have failed, the natural frequency \u03c9n,\u03c6 and \u03c9n,\u03c8 of the designed roll angle and yaw angle channel are 2.5rad/s, the natural frequency \u03c9n,\u03b8 of the pitch angle channel is 2, and the damping ratio \u03c2 is 0.5. According to the real object (Fig. 6), the parameters of the aircraft are measured as shown in Tab. I. Authorized licensed use limited to: Miami University Libraries. Downloaded on June 16,2021 at 05:42:48 UTC from IEEE Xplore. Restrictions apply. 7477 The aircraft can change the position of the center of gravity by increasing the bottom counterweight, and H = 0.05m during the simulation. The calculated controller coefficients of each channel are shown in Tab. II. A step input (0.5rad) is given to each channel at the zero time, the response curve of the system (Fig" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0002144_icem49940.2020.9270848-Figure4-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0002144_icem49940.2020.9270848-Figure4-1.png", + "caption": "Fig. 4. Geometry and 3D flux density plot of the double sided axial flux machine from the three dimensional FEA-simulation.", + "texts": [ + " In this section the simulation setup is introduced. The analytical model (ACM) is validated by a 3D finite element analysis (FEA) simulation. The considered EM is a doublesided axial flux machine with concentrated windings. The winding configuration is set to five slots and four poles. This leads to a higher winding factor of the fundamental component [19]. For a preliminary verification of the analytical model, the machine design is quite simple and no optimization of the machine geometry is conducted. A 3D non-linear FEA simulation (see Fig. 4) is solved with a stationary PARDISO solver. A parameter sweep over one mechanical period of the stator is calculated. The material properties of the rotor disk are determined by the material of machining steel 9S20 (see [25]) and the magnets are simulated as NdFeB - magnets (BMN-35SH). The remaining components are set as air. All material properties are listed in Tab. III. In Fig. 5 the airgap flux density and its frequency spectrum of the FEA simulation and the analytical model are compared. The signals are displayed normalized" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0002119_icaccm50413.2020.9212874-Figure1-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0002119_icaccm50413.2020.9212874-Figure1-1.png", + "caption": "Fig. 1. Quadcopter schematic diagram.", + "texts": [ + " The propeller model has been tested under the static loading and vibrations. The static loading has been applied considering bending in two different directions and the modal analysis has been performed for the vibrational frequencies. The obtained results have been analyzed and compared for two different materials namely, carbon fiber reinforced polymer (CFRP) and glass fiber reinforced polymer (GFRP). I. INTRODUCTION A Quadcopter is a type of unmanned aerial vehicle (UAV). It consists of four propellers at the end of the arms of the Quadcopter. Figure 1 shows the schematic representation of a Quadcopter. The battery provides the required amount of power to the motors, which rotate the propellers [1, 2,]. The rotation of these propellers generates the thrust forceF1, F2, F3 and F4.These propellers push the air in downward direction and the Quadcopter frame in upward direction. If the payload increases the thrust force generated by propellers try to bend the propeller in the vertical upward direction. Most of the propellers are failed due to the excess loading or by the collision (accident)" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0001334_s40435-020-00661-8-Figure1-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0001334_s40435-020-00661-8-Figure1-1.png", + "caption": "Fig. 1 Rotor forces and moments of a typical quadrotors", + "texts": [ + " However, in variable-pitch quadrotors the rotor forces and moments are created by changings the collective pitch through the pitch link at the rotor hub. Therefore, the constant speed for all rotors expected during any operations, and hence, the controls that generate a single force or moment is described as collective pitch, differential pitch, differential roll, and differential yaw provided by the blade collective pitch angle. To model the variable-pitch rotor, the forces generated by each rotor are calculated by blade element theory (BET). The forces and moments of each rotor are illustrated in Fig.\u00a01. This theory determine the forces of the blade element at ( rb, b ) and then integrates over the blade radius from root cutout to the blade effective radius, BR , where B is blade tip loss factor. Unlike hover, in forward flight, the rotor loads have to be averaged over a rotor revolution to obtain the steady state forces and moments. The aerodynamic load on the blade element provides a shear load and a moment at the blade root. In a conventional rotor, the flapping hinge at the blade root avoids much of this load from being transferred to the fuselage", + " The equations of motion of the fuselage are those of a rigid body, formulated (15) CMy = B2 a 2 ( 1 4 0 y \u2212 1 4 y z \u2212 1 8 B2 1c + 1 3 B y 0 + 1 4 B2 y 1 ) (16) CMz = B2 a 2 ( B2 Cd 0 4a \u2212 1 2 2 0 \u2212 1 2 2 z + Cd 2 2a 2 0 + Cd 0 4a 2 x + Cd 0 4a 2 y + Cd 2 2a 2 z + 1 4 Cd 1 a 0 ( 2 y + 2 x ) \u2212 1 8 B2( 2 1c + 2 1s ) + B2 Cd 2 8a ( 2 1c + 2 1s ) + Cd 2 4a B2 2 0 + Cd 2 6a B4 2 1 + Cd 2 4a 2 0 ( 2 x + 2 y ) + Cd 1 3a B ( 0 \u2212 z ) + 0 z a ( 1 \u2212 Cd 2 ) + Cd 1 4a B2 0 + Cd 1 5a B3 1 + 1 8a B2Cd 2 ( 2 x + 2 y ) 2 1 \u2212 1 4 B2 1 ( 0 \u2212 z) + Cd 2 2a B2 0 1 + 1 6a BCd 1 1 ( 2 x + 2 y ) \u2212 Cd 2 2a B2 z 1 + 2Cd 2 5a B3 0 1 + Cd 1 6a B x ( 1s \u2212 y ) \u2212 1 3 B 0 ( 0 \u2212 z ) + 2Cd 2 3a B 0 ( 0 \u2212 z) \u2212 1 6 B 0 ( 1s x + 1c y) + Cd 2 3a B 0 ( 1s x 0 \u2212 1c y ) \u2212 1 8 B2 1 ( 1s x + 1s x + 1c y + 1c y ) + Cd 2 4a B2 1 ( 1s x \u2212 1c y ) + Cd 2 3a B 0 1 ( 2 x + 2 y )) (17) ( rb, ) = 0 + rb R ( 1c cos b + 1s sin b ) (18) 0 = x tan TPP + CT 2 \u221a 2 x + 2 0 in the body-fixed coordinate system are given by Eq.\u00a0(19). Based on Fig.\u00a01, the longitudinal equations of motion of fuselage are obtained as: and the lateral-directional equation of motions are: where the fuselage forces and moments are given by (Fig.\u00a03): (19) XF + 4\u2211 i=1 XMi = m(u\u0307 + qw \u2212 vr + gsin\ud835\udf03) ZF + 4\u2211 i=1 ZMi = m(w\u0307 + pv \u2212 qu \u2212 g cos\ud835\udf19 cos \ud835\udf03) MyF + 4\u2211 i=1 Myi \u2212 ZM1zM1 + ZM2lM2 + ZM3lM3 + ZM4lM4 \u2212 4\u2211 i=1 XMi hM \u2212 XFhF + ZFlF = Iyyq\u0307 + ( Ixx \u2212 Izz ) pr (20) MxF + 4\u2211 i=1 Mxi + 4\u2211 i=1 YMihM \u2212 ZM4yM4 + ( ZM1 + ZM3 ) yM1 + ZM2yM2 + YFhF + ZFyF = Ixxp\u0307 + ( Izz \u2212 Iyy ) qr YF + 4\u2211 i=1 YMi = m(v\u0307 + ru \u2212 wp \u2212 g sin\ud835\udf19 cos \ud835\udf03) MzF + 4\u2211 i=1 Mzi + XM1yM1 \u2212 XM2yM2 + XM3yM3 + XM4yM4 + YM1lM1 \u2212 YM2lM2 \u2212 YM3lM3 \u2212 YM4lM4 \u2212 XFyF \u2212 YFlF = Izzr\u0307 + ( Iyy \u2212 Ixx ) pq and the fuselage moments are approximated as MxF = MyF = MzF = 0 " + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0003451_s00707-015-1403-6-Figure10-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0003451_s00707-015-1403-6-Figure10-1.png", + "caption": "Fig. 10 Motion of the NNC model with velocities of equal intensities in a moving system x\u0304 O\u0304 y\u0304", + "texts": [ + "; when x\u0307 = 0(y\u0307 = 0) and \u03d5\u0307 = 0 it will be x\u03070 sin \u03d50 + y\u03070 cos\u03d50 = 0 which means that the system does not rotate around the axis Mz and the particle M is moving along the straight line\u2014Fig. 9b: x sin \u03d50 + y cos\u03d50 = 0 where E = T = 1 2 (2m) ( x\u03072 + y\u03072 ) = const. Finally, if x\u0307 = 0 (y\u0307 = 0) and \u03d5\u0307 = 0, the system is at rest, so it will be E = T = 0 \u2261 0 = const. Furthermore, let us observe the motion of the considered model with respect to the moving coordinate system. Let the described model move in the system O\u0304 x\u0304 y\u0304 that is moving relative to the stationary system Oxy according to the well-known law x0 = B(t)\u2014Fig. 10. The equation of nonholonomic constraint relative to the system Oxy reads F = ( x\u0307 \u2212 B\u0307 ) sin \u03d5 + y\u0307 cos\u03d5 = 0, (5.44) and the differential equations of motion are 2mx\u0308 = \u03bb sin \u03d5, 2my\u0308 = \u03bb cos\u03d5, Jz \u03d5\u0308 = 0. (5.45) Differentiating the equations of constraints (5.44) with respect to time, it is obtained dF dt = ( x\u0308 \u2212 B\u0308 ) sin \u03d5 + ( x\u0307 \u2212 B\u0307 ) \u03d5\u0307 cos\u03d5 + y\u0308 cos\u03d5 \u2212 y\u0307\u03d5\u0307 sin \u03d5 = 0, and substituting x\u0308 and y\u0308 from the differential equations of motion (5.45), it follows (for m = 1 2 ) that dF dt = ( \u03bb sin \u03d5 \u2212 B\u0308 ) sin \u03d5 + ( x\u0307 \u2212 B\u0307 ) \u03d5\u0307 cos\u03d5 + (\u03bb cos\u03d5) cos\u03d5 \u2212 y\u0307\u03d5\u0307 sin \u03d5 = 0 wherefrom it is obtained \u03bb = B\u0308 sin \u03d5 \u2212 ( x\u0307 \u2212 B\u0307 ) \u03d5\u0307 cos\u03d5 + y\u0307\u03d5\u0307 sin \u03d5+ = B\u0308 sin \u03d5 \u2212 [( x\u0307 \u2212 B\u0307 ) cos\u03d5 + y\u0307 sin \u03d5 ] \u03d5\u0307 = B\u0308 sin \u03d5 \u2212 ( \u02d9\u0304x cos\u03d5 \u2212 \u02d9\u0304y sin \u03d5 ) \u03d5\u0307" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0000566_ab694f-Figure1-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0000566_ab694f-Figure1-1.png", + "caption": "Figure 1. The experimental setup.", + "texts": [ + " In previous works, this task has been performed by analysing the frequency of the sound emitted by the spinner [4], using stroboscopic lights, video analysis or commercial photogates [5]. In the present work, we use an alternative method to perform this task using a reed switch connected to an Arduino board. The main advantages of solutions based on the Arduino board are the low-cost of the components, the simplicity of operation, and the precision obtained [6\u20138]. We measure the angular speed of the spinner as a function of time after giving it an initial speed, and compare the experimental results with a theoretical model. Figure 1 shows the experimental setup used for the acquisition of data. The central bearing of the spinner is fixed to a vertical axle and a small magnet is placed in one of the spinner blades. A reed switch (Soway RS-02) is placed just below the spinner in order to be operated by the magnetic field of the magnet. The reed switch is connected the GND pin of the Arduino board (ground connection) and to the digital pin 12. When the magnet passes near the switch, the switch is closed and the digital pin from the Arduino board go to LOW" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0001072_sceecs48394.2020.171-Figure5-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0001072_sceecs48394.2020.171-Figure5-1.png", + "caption": "Figure 5: Front view Design of the model", + "texts": [], + "surrounding_texts": [ + "Articulated and SCARA robots are well recognized in the field of industrial robotics because of combined linear and rotary motion resulting in ease of maneuverability for complex tasks. But in industries where high work load is handled and requires high stability, Gantry robot is preferred. This may be understood by the following table [1]. In industries, the already existing gantry robots have either their base axis (X) fixed or they are moving on the overhead railings which limits the workspace of the robot as the location to pick and place an object is fixed. This project aims to make a movable gantry robot which can move on the floor and can access the part available at the floor space. By this, it can pick the object from any location and can place it on any other location, thereby having a greater workspace. Moreover, any sudden changes in picking and placing position can be encountered easily. Motor driven wheels are mounted on the base of the robot for its locomotion. II. CONCEPTUAL DESIGN As we discuss, the Gantry robot consist of minimum three motions. Each is in linear motion. This linear motion are perpendicular to each other which moves in X, Y and Zdirection. This X & Y-direction are on horizontal plane whereas the Z-direction is on vertical plane. The volume created by the X, Y & Z-axis in which the end effector of robot will move is known as the working envelope. 20 20 IE EE In te rn at io na l S tu de nt s' C on fe re nc e on E le ct ric al ,E le ct ro ni cs a nd C om pu te r S ci en ce (S C EE C S) 9 78 -1 -7 28 1- 48 62 -5 /2 0/ $3 Authorized licensed use limited to: Auckland University of Technology. Downloaded on June 02,2020 at 14:03:08 UTC from IEEE Xplore. Restrictions apply. SCEECS 2020 In this project, the motion of X, Y & Z-direction motion is achieve by the lead screw (Threaded rod) and coupler (nut) arrangement. The one end of the lead screw will be attached to the motor. So that the rotation motion will take place in the threaded rod and we kept the nut fixed. As the rod rotate the nut will move forward or backward according to the direction of rotation of the rod (i.e. clockwise or anticlockwise). Same technique is used for z-direction but in vertical plane. The motion of the nut will be upward and downward according to the rotation of rod. Figure 1: Design of Gantry Robot from top view This is the design of the structure from top side where all the motion of 3-axis are being shown. A coupling is used to connect motor shaft of 4mm to the 8mm of the lead screw. There is a fourth degree of freedom which rotates in the X-Y plane. To make the end effector more flexible according to the application. In this project we are attaching a fixed arm at the end-effector of Gantry robot. Authorized licensed use limited to: Auckland University of Technology. Downloaded on June 02,2020 at 14:03:08 UTC from IEEE Xplore. Restrictions apply. SCEECS 2020 III. CONTROLLING OF GANTRY ROBOT We are using the NodeMCU as the micro-controller. The NodeMCU has the functionality in which he can act as a server. NodeMCU has the station(STA)mode in which it can connect with the existing wifi-network and act as a HTTP server with the IP address assign to that network.[5] It act as a HTTP server only for the device which are connect to the same wifi network A. Interface of client with micro-controller. We are using the NodeMCU as the micro-controller. The NodeMCU has the functionality in which he can act as a server. NodeMCU has the station(STA)mode in which it can connect with the existing WiFi-network and act as a HTTP server with the IP address assign to that network.[5] It act as a HTTP server only for the device which are connect to the same WiFi network. HTTP stands for the Hypertext Transfer Protocol. It is standard application protocol which functions as requestresponse protocol between client and server. HTTP client help to send and receive the HTTP request and HTTP response respectively from the HTTP Server. Web-server of the NodeMCU will open when we put the IP address of the network in the address bar. Then the HTTP server page is open where we can send and receive the request of the client. The user is going to control the motion of the stepper, DC and servo motor using this web-page which is created on the HTTP server. BY clicking on the buttons which is created using the HTML. B. Interface between micro-controller and the actuators Now after receiving the data from the client. Microcontroller will instruct accordingly to the actuators. 1) Interfacing of Stepper motor with Node-MCU The amount of power required by the Stepper motor to work is 12V-1.5ampere which is provided by external power source. We also need a circuit which can control the rotation of motor perfectly and easily. Hence, we need an Hbridge circuit to control the stepper motor with Node-MCU i.e. micro-controller, H-bridge circuit is the combination of transistor or mosfets. So, here we are using the A4988 circuit. Using this IC we can control the motor by just 2-pins. One for the direction of the motor and other pin for the steps. This driver provide the full-step, haft-step, quarter-step, eight-step and sixteenth-step resolutions. This stepper motor is connect with a lead screw to move the complete articulated arm in X, Y and Z-coordinates. 2) DC Motor The DC motor is used to move the complete Gantry Robot from one location to another. So, to control the dc motor through NodeMcu we are using the driver IC- L298N. Authorized licensed use limited to: Auckland University of Technology. Downloaded on June 02,2020 at 14:03:08 UTC from IEEE Xplore. Restrictions apply. SCEECS 2020 As similar to stepper motor the DC-motor require the amount of power is 12V,1 or 2 ampere which is provided by external power source. Here, also we need H-bridge circuit to control and L298N is a dual channel H-bridge motor driver IC which can able to drive a pair of DC motor. In L298N driver IC there is a voltage drop of nearly 2V due to the switching transistor in H-bridge circuit. The servo motor is used to move the articulated arm which is attached to the end-effector of a Gantry Robot. The servo motor is controlled by the node-MCU pin. The motor is also connect to the 2 Jaw-gripper which is the end-effector of the articulated arm. The 2-jaw gripper is used to pick up the object and placed it at the required position. The gripper can be classified as coarse grippers or precise grippers. The coarse gripper is found where the components can be pick or placed with relatively little accuracy, whereas, precise gripper are used where the accuracy is essential or should have accurate holding position and orientation. 4) Gripper : Gripper is mounted as the end effector of the robot to hold and release the object. There are various factors while selecting a gripper which are as follows \u2022 Object to be handled \u2022 Sequence of operation \u2022 Mechanism of robot \u2022 Environment of workspace The first and foremost consideration while selecting a gripper is the object which is to be handled. There are following parameters related to object which influence the selection of gripper \u2022 Size of object \u2022 Number of potential gripping surface \u2022 Geometry of gripping surface \u2022 Distance of gripping surface Considering the above points, a two jaw V-Shaped gripper is chosen for the application. The gripper is actuated by the servomotor and a pairs of pinion. When the servomotor is actuated it drives the pinion in the same direction of rotation, whereas the gear in mesh with pinion is driven in opposite direction making the jaw of the gripper to come close and hold the object. To release the grip, the motor is rotated in opposite direction. IV. RESULT AND DISCUSSION A mobile gantry robot for pick and place application will be much more suitable for the industries having application for storing and dispatching of the goods in warehouses as well as in industries having assembly lines operation where an object has to picked up from a particular assembly line and is need to be placed at some other location. By using the controller, one can easily control the movement of the robot and will be much easier and flexible than any other robot which could be used for the same application. This robot will be more stable while picking and placing the objects because of its sturdy and symmetrical structure, thus enabling it to lift heavy objects as well. V. CONCLUSION \u201cMobile Gantry Robot for Pick and Place Application\u201d is an industry oriented project aiming specifically on pick and place application in the industries which came out as a solution looking at the problem in other pick and place robots as well as conventional robots. This project will ease out the various operation of the industries as far as the pick and place application is concerned. Authorized licensed use limited to: Auckland University of Technology. Downloaded on June 02,2020 at 14:03:08 UTC from IEEE Xplore. Restrictions apply. SCEECS 2020 VI. REFERENCES [1] https://www.linearmotiontips.com/when-do-you-need-a-gantry- robot/#ampshare=https://www.linearmotiontips.com/when-do-you-needa-gantry-robot/ [2] http://www.sagerobot.com/gantry-robots/ [3] Surinder Pal,\u201dDesign and remote control of a Gantry Mechanism\u201d, Office of graduate studies Texas A & M University, August 2017. [4] Mesfin B., Muluken T., Samson D., Abayneh T.,\u201d Design of Gantry Robot\u201d, Bahidar University of Techonolgy,2014. [5] https://www.electronicwings.com/nodemcu/http-server-on-nodemcuwith-arduino-ide [6] https://www.engineersgarage.com/esp8266/nodemcu-esp8266-steppermotor-interfacing/ [7] http://www.ece.montana.edu/seniordesign/archive/FL13/GantryRobot/in dex.html [8] https://www.macrondynamics.com/job-stories/gantry-systems-overview [9] https://www.machinedesign.com/motion-control/five-heavy-dutygantry-alternatives [10] https://howtomechatronics.com/tutorials/arduino/how-to-controlstepper-motor-with-a4988-driver-and-arduino/ Authorized licensed use limited to: Auckland University of Technology. Downloaded on June 02,2020 at 14:03:08 UTC from IEEE Xplore. Restrictions apply." + ] + }, + { + "image_filename": "designv11_71_0002099_s12206-020-1029-z-Figure3-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0002099_s12206-020-1029-z-Figure3-1.png", + "caption": "Fig. 3. ME conditions: (a) one-dimensional; (b) two-dimensional.", + "texts": [ + " 2(c) shows the three-dimensional FE condition, where the ends of all four cable force vectors could form a tetrahedron when the other ends of all four vectors are translated the common point. The one- and two-dimensional ME conditions are illustrated in Figs. 3(a) and (b). One-dimensional ME condition can be expressed as 2 1 ( ) 0,i i z i r f = \u00d7 =\u2211 2 1 ( ) 0,i i x i r f = \u00d7 =\u2211 2 1 ( ) 0 .i i y i r f = \u00d7 =\u2211 (7) This condition implies that all cable forces should lie on the -x y plane as shown in Fig. 3(a) (or each of all cable forces intersects both the x axis and the y axis) while the moment about the z axis generated by all cable forces is in equilibrium. Note that all those cable forces satisfying one-dimensional ME condition should also be balanced. Fig. 3(a) shows the case that all those two-dimensional cable forces are balanced by two-dimensional force constraint. Two-dimensional ME condition can be expressed as 3 1 ( ) 0, ti i x i r f = \u00d7 =\u2211 3 1 ( ) 0, ti i y i r f = \u00d7 =\u2211 3 1 ( ) 0 . ti i z i r f = \u00d7 =\u2211 (8) This condition implies that all cable forces should intersect the common axis (the \u02c6tz axis) as shown in Fig. 3(b) while the moment about both the \u02c6tx axis and the \u02c6ty axis, which are generated by all cable forces, should be in equilibrium. Note again that all the cable forces are also balanced by threedimensional force constraint as shown in Fig. 3(b). For the three-dimensional ME condition, all moment vectors ( ,i i im r f= \u00d7 for i = 1, 2, 3, 4) generated by the cable forces should form a tetrahedron, similarly to the three-dimensional FE condition. Note again that all those cable forces satisfying the three-dimensional ME condition should also satisfy the three-dimensional FE condition, with or without force constraint wrenches imposed on the CDPM. 2.2 Background on synthesis of pTqR type CDPMs This study focuses on the synthesis of the pTqR type n-DOF CDPM structures which are driven by n +1 cables, where pT and qR denote the p-DOF translational motion and the q-DOF rotational motion, respectively" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0002355_j.egyr.2020.11.179-Figure2-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0002355_j.egyr.2020.11.179-Figure2-1.png", + "caption": "Fig. 2. Magnetizing fluxes of BEESMs. (a) BEESM1. (b) BEESM2.", + "texts": [ + " Two kinds of indings are placed on the stator, one is dc field winding, and the other is ac armature winding. The field winding s concentrated and the armature winding is distributed. One field winding is corresponding one salient pole and two alient poles for the BEESM1 and BEESM2, respectively. The field windings of both machines can be connected n parallel or series. The current directions of the ends of adjacent field windings are opposite. 2.2. Operation principle When a dc current is injected into the field winding, the magnetizing flux of BEESM1 and BEESM2 would be enerated, as shown in Fig. 2. The polarity of air-gap flux density under neighboring field windings are magnetized s opposite polarity \u2018N\u2019 and \u2018S\u2019. Ignoring the saturation of iron core, the amplitudes of magnetomotive force (MMF) generated by the field winding an be calculated as F f = N f i f /2 (1) here N f is the number of turns of one coil for the field winding and I f is the field current in the coil. The MMF distributions of both machines are given in Fig. 3(a). The polarity of rotor salient poles depends on the directions of the magnetizing flux generated by the excitation winding" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0003878_s11434-014-0376-5-Figure16-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0003878_s11434-014-0376-5-Figure16-1.png", + "caption": "Fig. 16 (Color online) Adjusting angles of leg joints to achieve the best supporting posture", + "texts": [ + " Therefore, the overall evaluation M is still used to evaluate the leg posture. In Fig. 15, the variation of the overall evaluation M is given and the minimum value of the overall evaluation M could be found, while the value of h1 is 70.7 and h2 is 71.8 . At the same time, the leg posture should satisfy the physical constraints, the values of h3 and h4 could be calculated by the constraint equations (Eq. (9), Eq. (10)) according to h1 and h2. The value of h3 is 103.8 while the value of h4 is 115.6 (Fig. 16). The posture making the overall evaluation M minimum is regarded as the best supporting posture (Fig. 16). Under the actual constraints, the leg could not be at the dead-point state. However, the best supporting posture will be considered as the nearest posture of the dead-point state. In this study, the horse forelimb modeled as a five-link mechanism was analyzed. The inverse dynamic method was used to calculate limb joint forces and moments. Changing the ground reaction force and limb postures, the variations of limb joint forces and moments were calculated and the dead-point supporting effect of the leg structure help improve high loading capacity of horse standing" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0002531_epe50722.2020.9305650-Figure1-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0002531_epe50722.2020.9305650-Figure1-1.png", + "caption": "Fig. 1. FSCW-PMSM cross sections.", + "texts": [ + " The interest was to find the best candidate for a wind energy conversion system that can easily be adapted to various working parameters such as rated power and wind speed. II. GENERATOR DESIGN Rated working parameters for the studied electric machine are presented in Table I. According to these specifications, a fractional slot concentrated winding synchronous generator with 18 stator teeth/coils and different number of rotor pole pairs (from 6 to 12) has been designed with the main dimensions presented in Table II and cross sections in Fig.1. The rotor speed has been adopted different for each generator structure, according to the pole pairs number, so that the voltage frequency to be the same (f = 20 Hz). The total permanent magnet volume used for each structure has been kept constant as well, preserving the same polar coverage factor. The stator geometry is basically the same for all cases, except the connections between the 18 coils that were changed accordingly to the phasor diagrams presented in Fig. 2. TABLE I. GENERATOR RATED WORKING PARAMETERS Name of the parameter Value Rated power 3" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0003880_scis-isis.2014.7044812-Figure8-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0003880_scis-isis.2014.7044812-Figure8-1.png", + "caption": "Fig. 8. Dead-lock with a moving obstacle: (a) Simulation snapshot, (b) Total potential force variation, (c) Enlarged total potential force with oscillation", + "texts": [ + " Then, no velocity command is generated to move the robot forward and dead-lock is taken place. Fig. 7(b) explains the total potential force variation with the travelling distance of the robot in the Y-direction. It shows that the total potential force doesn\u2019t vary at the beginning and it starts to change when the robot detects the obstacle. When robot reaches the obstacle total potential force is decreasing and it becomes zero around 7m distance in Y-direction. Simulation study done for a dynamic obstacle moving towards the robot is shown in Fig.8. Initially, robot starts moving towards the goal and as the obstacle reaches the robot, it starts to move back away from the goal with oscillations. In this unexpected motion, robot has first come to a local minimum position and since the obstacle is moving towards the robot, the total potential force, forces the robot to move backward. As shown in Fig.8 (b), total potential force becomes zero for a moment and as the obstacle comes closer to the robot total potential force shows an oscillation. As a result, the robot 978-1-4799-5955-6/14/$31.00 \u00a92014 IEEE 262 starts to move back-ward with back and forth oscillation. Zoom-in graph in Fig.8(c) explains the variation of the total repulsive force for the backward motion of the robot in Ydirection. Though total potential force shows a positive value opposite to the goal direction, the robot moves backward. Therefore the robot never reaches its goal position until obstacle goes away from its path. In order to avoid this kind of motion in dynamic environments and dead-lock problem in static environments, motion planning algorithm is modified. To have a proper motion in such environments the proposed modified APF has shown the performance well" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0003642_cdc.2014.7039598-Figure2-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0003642_cdc.2014.7039598-Figure2-1.png", + "caption": "Fig. 2. Drawing of the gripper of the PERA", + "texts": [ + " We finally motivate the present PH approach for impedance grasping control with an example of an end-effector system. A. End-effector system To gain more insight into the role of the PH system (7), and the impedance grasping control strategy (19), we consider a class of standard mechanical systems with a constant mass-inertia matrix. The system is given by the gripper (end-effector) of the Philips Experimental Robot Arm (PERA), [14]. A picture of the PERA is shown in Figure 1 while a drawing of its gripper is shown in Figure 2. The gripper consists of a shaft actuated by the motor of the gripper which is attached to the fingers via cables. When the shaft moves left the gripper closes, and when it moves right the gripper opens. The gripper is controlled via scripts developed in Matlab R\u00a9 with a sampling time of 10ms. The model of the gripper in the PH framework consists of a mass mg, interconnected by a nonlinear spring, and a linear damping dg > 0. The states of the system are x = (q, p)>, where q is the displacement between the two tips of the gripper, and p is the generalized momenta of the system" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0000356_iecon.2019.8927121-Figure11-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0000356_iecon.2019.8927121-Figure11-1.png", + "caption": "Fig. 11. The modular iron core of the rotary-linear SRM [33]", + "texts": [ + " The RLM based on its most typical variant, the switched reluctance motor (SRM), is also a potential candidate to provide two DoF movement. This is due to its numerous appealing features, as easy-to-manufacture rugged and robust construction, low maintenance needs, simple power converter, wide speed range, etc. [32]. The rotary-linear SRM proposed in [33] is a simple but efficient combination of the rotational and linear SRMs. It has all the pluses of the SRMs: simple, but robust construction, low manufacturing and maintenance costs, high reliability and fault tolerance, and easy control. In Fig. 11 the iron core of this machine can be seen. The stationary part is comprising of 3 equally shifted usual 8 poles SRM stators. Each of the stator poles have a concentrated coil wound around. The mover is also constructed in a modular way. It has usual 6 poles SRM rotor stacks placed on the shaft. Their number depends only on the length of the intended axial displacement. The rotation or linear movement of this RLM can be achieved by the adequate supplying sequence of the stator coils. Upon another approach of the rotary-linear SRM, mainly similar iron core pieces taken from the already known rotary or S S N N Ferromagnetic poles Permanent magnets linear SRMs are used" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0003879_ijvd.2014.065716-Figure1-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0003879_ijvd.2014.065716-Figure1-1.png", + "caption": "Figure 1 Single-end full-scale inertial dynamometer", + "texts": [ + " The main step is to develop the dynamic model of braking torque change against the influence of speed, brake actuation pressure and brake interface temperature, for different brake pedal travels. It is important to find the functional relationship between pressure, speed and temperature in order to create possibilities for adjusting the brake actuation pressure upto the level that provides stable, but at the same time, maximum brake performance. In order to find and establish that correlation, the disc brake has been previously tested using single-end full-scale inertial dynamometer (see Figure 1). The tested disc brake belongs to the front axle of passenger car (Florida 1.4) with static load of 730 kg. The disc brake consists of the solid disc with effective radius of 101 mm, floating calliper with piston diameter of 48 mm and disc pads friction surface area of 32.4 cm2. The single-end full-scale inertial dynamometer has been used in order to provide data in controlled testing conditions in accordance with the selected testing methodology. It is very important because friction material characteristics critically affect overall brake performance and its change in a braking cycle" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0002989_978-3-642-16214-5_9-Figure9.1-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0002989_978-3-642-16214-5_9-Figure9.1-1.png", + "caption": "Figure 9.1 shows internal shifting elements for engaging gearwheels into the power flow. A distinction is made between positive locking clutches (e.g. dog clutch) and friction clutches (e.g. multi-plate clutch).", + "texts": [ + "3 (ZF) Smart (Getrag), Figure 12.14 \u2013\u2013\u2013\u2013\u2013\u2013\u2013\u2013\u2013 Shifting Shift fork Swing fork Piston \u2013\u2013\u2013\u2013\u2013\u2013\u2013\u2013\u2013 \u2013\u2013\u2013\u2013\u2013\u2013\u2013\u2013\u2013 Example Figure 12.9 (MB) Figure 12.1 (VW), Figure 12.3 (ZF) Conventional automatic transmission, Figure 12.23 (MB) \u2013\u2013\u2013\u2013\u2013\u2013\u2013\u2013\u2013 \u2013\u2013\u2013\u2013\u2013\u2013\u2013\u2013\u2013 Frictional connection Single/Multicone Spreader ring Multi-plates Belt Sprag Example Cone synchronizer Porsche synchronizer Clutch, brake, Figure 6.30 Brake, Figure 6.35 Freewheel (AT), Figure 6.30 Positive engagement Dogs Pins Sliding gear Draw key \u2013\u2013\u2013\u2013\u2013\u2013\u2013\u2013\u2013 Example Figure 9.1b, d, e, f Figure 9.1c Figure 9.1a Motorcycle gearbox \u2013\u2013\u2013\u2013\u2013\u2013\u2013\u2013\u2013 Shifting by hand entails more than bringing gearwheels into the power flow. An exact and smooth-running operation of the gearshift lever is needed. This involves the interaction of external shifting with internal shifting elements such as detent devices, guides and synchronizers (Section 9.2). Passenger cars must have attributes such as short shift strokes, a fluid shifting process, a crisp, sporty shifting feel and low shifting forces. The simplest type of shift system is sliding gears (Figure 9.1a). The gearwheels are not constantly meshed but are shifted into the power flow as needed. Sliding gears are used for reverse gear in both passenger car and commercial vehicle transmissions. Unsynchronized constant-mesh transmissions are often found in commercial vehicle transmissions. The constant-mesh gear pairs run on rolling bearings and have a positive locking connection to the transmission shaft via a sliding dog sleeve (gearshift sleeve) (Figure 9.1b). The gears are prevented from disengaging (gear dropout) by undercut dogs (Figure 9.2). Gearshifting is always made up of a selecting movement and a shifting movement. The selecting movement selects the gearshift sleeve to be shifted, and the shifting movement moves the gearwheel into the power flow. Figure 9.3 shows an example of this for a synchromesh gearbox with direct shifting (gearshift lever on the gearbox housing, e.g. in commercial vehicles) by three selector bars. The gearshift lever 1 and the ball joint 2 serve to select the gear and transmit the manual effort" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0003270_icpe.2015.7167862-Figure13-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0003270_icpe.2015.7167862-Figure13-1.png", + "caption": "Fig. 13. 8-pole 12-slot IPM-type BLDCM", + "texts": [ + " In equality demagnetization case, only size of fundamental harmonic is changed according to demagnetization ratio. In addition, 2nd and 4th harmonics responsible for waveform asymmetric are almost zero. In inequality and weighted demagnetization case, fundamental harmonic is decreased as being proportion to demagnetization ratio. However, 2nd and 4th harmonics have comparatively large value unlike equality demagnetization. When motor is working, because, flux linkage variance of N-pole and S-pole is different, 2nd and 4th harmonics of BEMF are occurred. B. 8-pole 12-slot Fig. 13 shows the 8-pole 12-slot IPM-type BLDCM. The reason why 8-pole 12-slot model is selected, although this model have the same pole-slot combination ratio such as 6-pole 9-slot it is selected for confirmation in the same BEMF harmonic characteristics. Table III shows the specification of 8-pole 12-slot. In addition, Fig. 14 shows the concept figure of 8-pole 12-slot model applying to equality, inequality, and weighted demagnetization patterns. Fig. 14 (a) shows the equality demagnetization concept of 8-pole 12-slot model" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0001516_acc45564.2020.9147724-Figure1-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0001516_acc45564.2020.9147724-Figure1-1.png", + "caption": "Fig. 1. Each satellite is equipped with an electromagnetic actuation system consisting of three orthogonal coils. The relative positions of the satellites are controlled by the interaction of equal-and-opposite electromagnetic forces produced by the actuation systems.", + "texts": [ + " Next, let Fij be a reference frame whose orientation depends on the direction of rij . Specifically, let iij , jij , and kij be the orthogonal unit vectors of Fij , and Fij is aligned such that rij \u00b7 iij = ||rij ||. Thus, resolving rij in Fij yields [||rij || 0 0]T. Let Rij : [0,\u221e) \u2192 SO(3) be the rotation matrix from Fij to FI. It follows that rij = Rij [||rij || 0 0]T. Since rij = \u2212rji, it follows that Fji is aligned such that iji = \u2212iij . We assume that jji = \u2212jij and kji = kij , which implies that Rji = Rij\u03a6, where \u03a6 , diag(\u22121,\u22121, 1). See Fig. 1 for a depiction of vectors and reference frames. Each satellite is equipped with an electromagnetic actuation system, which is used to generate intersatellite forces. Let the vector Fij denote the normalized electromagnetic force per unit mass applied to satellite i by satellite j, and let Fij \u2208 R3 denote Fij resolved in Fij . The intersatellite forces are modeled by Fij = 3\u00b50 4\u03c0||rij ||4 Uij , (1) where \u00b50 is the permeability constant and Uij = 2AiAj \u2212BiBj + CiCj \u2212AiBj \u2212AjBi AiCj \u2212AjCi (2) is the intersatellite control of satellite i between satellite i and j, and Ai, Bi, Ci \u2208 R are the actuator controls" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0002221_iccad49821.2020.9260553-Figure3-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0002221_iccad49821.2020.9260553-Figure3-1.png", + "caption": "Fig. 3. Kinematic diagram of the mobile robot", + "texts": [], + "surrounding_texts": [ + "Keywords: Fuzzy logic; the human-machine interface; mobile robot; ultrasonic sensors; and infrared.\nI. INTRODUCTION\nRobotics is a set of disciplines (mechanical, electronic, automatic, Computer). It is subdivided into two types: industrial robots and mobile robots. Industrial robots are generally fixed, they are used in numerous industrial applications: mechanical assembly, welding, painting ... Mobile robots are not fixed; they are classified according to locomotion in walking robots, with wheels, Caterpillars ... etc. Moreover, they can be classified according to the field of application in the military, laboratory, industrial, and service robots [1]. Mobile robots are a special case in robotics. Their interest lies in their mobility \", intended to fulfill difficult tasks (example: transport of heavy loads) and they work even in hostile environments (nuclear, marine, space, firefighting, surveillance ...).\nOn the other hand, the particular aspect of mobility imposes a technological complexity (sensors, actuators, energy) and methodological such as the processing of information using artificial intelligence techniques or particular processors (vectorial, cellular). The autonomy of the mobile robot is a faculty that allows it to adapt or make a decision to achieve a task even in an environment little known or unknown.\nThe A.G.V (Automated Guided Vehicle) is a special case of a mobile robot dedicated to purely industrial applications. It is sometimes called automatic trolley and it is equipped with automatic guiding equipment that is inductive, optical,\nelectromagnetic, or other. This type of vehicle is able to follow predefined and programmable paths or to plan its trajectories according to the type of guidance and navigation used. The A.G.V is presented in quite varied aspects, whether by shape, size, or weight.\nThe aim is the mobile robot develops control laws and makes decisions based on knowledge of its environment. The data that comes from different sensors embedded in a mobile robot are sometimes inaccurate, unreliable, and sometimes missing which affects the objective it has to achieve [2]. Mobile robot control is classified in the category of problems that are too complex. These systems do not have an accurate detection capability. Artificial intelligence techniques based on fuzzy logic are considered a very interesting solution for non-linear systems where it is difficult to create a mathematical model.\nEarly work has grown very rapidly due to the development of artificial intelligence techniques to ensure autonomous navigation of a mobile robot, such as fuzzy logic [3], and the artificial neural network (ANN) which is very popular among heuristic approach methods because of its parallel learning and processing capability [4] 5]. In addition to these two approaches, various searches have used algorithms inspired by nature such as Colony Optimization Ant (CAO) [6], [7], Genetic Algorithm [8], Particle Swarm Optimization [9, 10], Cuckoo Search [11], Grey Wolf Optimizer [12\u201314], and Dragonfly Algorithm [15]. As well as various combinations of these methods, including GA-Fuzzy Logic [16], GAArtificial Bee Colony [17], and the Fuzzy-Wind Driven Optimization algorithm [18], have proved to be interesting techniques to solve different aspects related to this issue. In particular, fuzzy logic control has proved to be efficient for the navigation of mobile robots, and the research proposed in [19] deals with the autonomous navigation of a mobile robot with obstacle avoidance. It remains to be verified that the proposed method is well optimized concerning computation time. Since the problem of fuzzy rule extraction is solved using a genetic algorithm. In [20], the authors propose a behavior-based fuzzy controller method in which the developed fuzzy controller can activate multiple behaviors at the same time. The controller follows a switching strategy that allows the robot to activate the correct behavior according to its situation in the environment.\n978-1-7281-6999-6/20/$31.00 \u00a92020 IEEE\nAuthorized licensed use limited to: Univ of Calif Santa Barbara. Downloaded on May 19,2021 at 19:47:56 UTC from IEEE Xplore. Restrictions apply.", + "Fuzzy logic provides a better way to automate human expertise, thus saving time and memory space, which gives a considerable speed to its inference engines by contributing to conventional methods. A fuzzy controller gives superior results than conventional controllers, and sometimes even better results than the human operator. Fuzzy logic has shown its effectiveness in managing uncertainty and incompleteness of data, making it a robust, simple, and adequate tool for dealing with these problems [21].\nThe rest of this paper is organized as follows. The next section describes the platform used in the implementation. The third section describes the kinematic model of the robot. Section IV presents the development of a fuzzy controller for avoiding obstacles. Simulation results that prove the performance of the proposed approach are given in Section V. Section VI presents the experimental test performed. The remarks and perspectives are presented through the conclusion in Section VII.\nII. MOBILE ROBOT: SPUTNIK\nSputnik is the next generation of Dr. Robot\u2019s line of research and development robots. Responding to requests for a robot that had the speed and payload of X80 with the development versatility of the DRK8080, a robot with the complexity expected of a research tool. In addition to all the features offered in the model X 80 robot, there are software modules, included with SPK-1, enabling remote access and control of the robot via Internet protocol.\nAs presented in Fig.1 Sputnik is a balance of speed and strength with portability and precision. The wheel-based platform\u2019s two 12V DC motors each supply 300 oz.-inches (22kg.cm) of torque to the robot\u2019s 18 cm (7 in.) wheels, yielding a top speed over 1 m/s (3.3 ft/s). Two high-resolution (1200 count per wheel cycle) quadrature encoders mounted on each wheel provide high-precision measurement and control of wheel movement. Weighing only 6.1 kg (13.5 lb.), the system is light, but it can carry an additional payload of 10 kg (22 lb).\nThe diagram presented in Fig.2 shows the topical scenario. The Sputnik is a wireless networked robot. It connects to the wireless AP or router via IEEE 802.11b/g network. The host PC (or called server PC) running the Sputnik control program could connect to this network:\nNetwork cable: connect the host PC to one of the LAN ports on the back of the router\nWireless: to connect the host PC to the wireless router, and configure the host PC\u2019s wireless settings using the default wireless configuration setting.\nIII. KINEMATIC MODEL OF MOBILE ROBOT\nThe wheeled mobile robot considered as a planer rigid body that rides on two wheels whose movement is obtained by acting on the speed of each wheel [23]. It can move towards any point in the plan by using the rotational speeds of two wheels. A set of equations shows that it is a system of multivariate equations. The pair (v,\u03c9) is the input control encompassing the linear and angular velocities. The parameters x, y, and \u03b8 are the coordinates and the orientation of the mobile robot respectively.\nThe kinematic model of a unicycle type robot is usually described by a simple non-linear model as shown in (1). xy\u03b8 cos \u03b8 0sin \u03b8 00 1 . v\u03c9 (1)\nIV. FUZZY CONTROL FOR MOBILE ROBOT.\nThe problem is to avoid the collision of the robot with the obstacles present in the environment. Does not care about the direction the robot takes to avoid the obstacle nor the memorization of the obstacles encountered.\nThe definition of a problem in fuzzy logic is composed of three main parts:\nAuthorized licensed use limited to: Univ of Calif Santa Barbara. Downloaded on May 19,2021 at 19:47:56 UTC from IEEE Xplore. Restrictions apply.", + "\u2022 Definition of variables: what inputs, outputs are available and relevant.\n\u2022 Partitioning the definition domains of these variables, we then create fuzzy subsets (near, very close, or far).\n\u2022 Definition of the rules which will give behavior to follow for each of the Situations.\nObstacles avoidance must necessarily be based on sensors giving information on the environment. We will use ultrasonic sensors and infrared range sensors here. A sensor allows us to record the distances of obstacles in the right, left, and front zones (see Fig.4).\nThe rules will, therefore, have three input variables (Front Distance, Right Distance, and Left Distance). To move the robot, the robot is supplied with an angle value. In this case, we choose first to turn then to advance with a constant speed. The angle variable is given in degrees, in negative value to turn left and positive to turn right (see Fig.5).\nThe partitioning variables domains used are shown in Fig.6. The front, left, and right variables will be identical (the robot is not right-handed or left-handed and thus possesses an exact symmetry).\nThe Membership function of the output variable angle is shown in Fig.7. The angle variable will allow giving the robot the future direction of navigation. In order not to give too many variations in the movement of the robot (go and return on the same axis) the range of definition of the angle variable will be: -50\u00b0 to 50\u00b0.\nThe fuzzy mechanism is done by three steps, fuzzification, inference, and defuzzification, as seen in Fig. 8. The fuzzy controller is considered as control device composed of a set of fuzzy rules to achieve the goals. In the fuzzification block, we define for example a fuzzy set A in a universe of discourse C defined by its membership function \u03bcA(c). For each \u03bc(c) representing the degree of membership of c in C, membership functions assigned with linguistic variables are used to fuzzify physical quantities. Next, in the inference block, fuzzified inputs are inferred to a fuzzy rules base. Rules base that maps input variables to output variables are represented in (2). IF c is A AND\u2026AND c is A THEN b is B (2)\nWhere i=1\u2026N. N is the number of rules in a given fuzzy rule base, c . . . c are the input variables which are the sensor data of the mobile robot, A \u2026 . A are the input fuzzy sets, B is the output fuzzy set and b is the output variable. The response of each fuzzy rule is weighted according to the degree of membership of its input conditions. The inference engine provides a set of control actions according to fuzzified inputs. Since the control actions are in a fuzzy sense. Hence, a defuzzification method is required to transform fuzzy control actions into a crisp value of the fuzzy logic controller. Mamdani fuzzy inference is the most commonly seen inference method.\nThe fuzzy knowledge base is the set of fuzzy rules.\nOur system has eight rules:\n\u2022 Going straight if the front is free.\n\u2022 Going straight if can go anywhere.\n\u2022 Turn a little left if the obstacle near on the right.\n\u2022 Turn a little right if the obstacle near on the left.\n\u2022 Turn right if the front is not free and the right is free.\n\u2022 Turn left if the front is not free and the left is free.\nAuthorized licensed use limited to: Univ of Calif Santa Barbara. Downloaded on May 19,2021 at 19:47:56 UTC from IEEE Xplore. Restrictions apply." + ] + }, + { + "image_filename": "designv11_71_0003802_icolim.2014.6934343-Figure2-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0003802_icolim.2014.6934343-Figure2-1.png", + "caption": "Fig 2. Geometry of the IA V and of the bundle (left). The computational domain (right).", + "texts": [ + " TEST SETUP We have considered and built a representative geometry of a basket of an IA V: this is made of a box that is I meter by 2 meters with a thickness of 10 cm. The IA V has a banister with a height of 1 meter and has a square section of 4 cm. A sketch of the basket is depicted in Fig 1. L. Barbieri et al. \u2022 Insulated aerial vehicles and high voltage live working in Italy: tests and studies to assess the safety aspects We have considered both a single bundle and a two bundle configuration. The first one is depicted in Fig 2. The bundle is made of three cilindric conductors each one placed on a vertex of an equilateral triangle with edges of 40 cm. Each conductor has a diameter of 4 cm. The bundle is 12 meter long. Each end is connected to a large sphere with a diameter of 2 meters to avoid any boundary effect. The conductors are bended by their own weight: the lowest point is I meter lower than the two end points. We have considered various reciprocal positions between the IA V and the conductors, see Fig 3. We have considered three sets of configurations: 1", + " Also with other cases we get small variations with a maximum of 12%. In all cases, Usa is lower without the IAV. This is due to the fact that the most important parameter for the discharge inception and propagation is the electric field on the power line. In fact the inception of all discharges takes place on the conductors. The maximum electric field on the bundle is more influenced by the ground than by the IAV. To support this, we have run some simulations using a parallel finite element code [5]. A representation of the mesh we have used is depicted in Fig 2: the number of elements varies with the mutual positions of the IA V and the power line, however it is approximately 800000 elements. have computed the ratio E / U50 that outlines the electric field produced by one kV of difference of potential. This parameter is a function of the geometry only. As we can see, the configurations with the IA V generate a weaker electric field, as a consequence, the inception voltages are higher. If we fix h, the position of the IA V has a small influence on U50 " + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0001019_j.precisioneng.2020.04.003-Figure7-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0001019_j.precisioneng.2020.04.003-Figure7-1.png", + "caption": "Fig. 7. Top and front views of the 3-V coupling showing only the threaded and truncated cylinders in a displaced state with constraint 2 disengaged.", + "texts": [ + " More importantly, these bearings do not have friction from seals. Combining two antifriction vees with a friction vee is useful in this context to simplify the calculation of the limiting CoF to \u03bc \u00bc tan \u03b1 of the friction vee. See the Appendix for the full derivation. L.C. Hale Precision Engineering 64 (2020) 200\u2013209 The unit experiment, which was repeated over multiple configurations of the coupling geometry and loading for a given condition of the contacting surfaces, involves gently resting the coupling on five of six constraints, as Fig. 7 demonstrates, and observing whether friction is low enough for it to slide to center as a one-degree-of-freedom mechanism or not. To be consistent with earlier work, the sliding direction is identified by the disengaged constraint, number 2 in the example of Fig. 7. Mainly as a matter of convenience, all experiments either had constraints 1\u20132 set to one inclination angle \u03b11-2 and the remaining constraints 3\u20136 set to a second angle \u03b13-6, or had all constraints set to the same angle, i.e., geometrically symmetric. This maintains a left-right symmetry for all tests and gives two observations for a single configuration of the coupling with constraint 1 or 2 being initially disengaged. The observed outcome: S for sliding, N for non-sliding or V for the verge of sliding when it is ambiguous, was recorded in a data table under the heading C1 or C2, indicating the initially disengaged constraint", + " Appendix The equilibrium equation (1) describes a 3-V kinematic coupling with five of six constraints engaged under the following restrictions and assumptions: 1) Three constraint planes, each containing a pair of constraint lines, are tangent to and equally spaced on a base cylinder of radius R, which for this experiment had a vertical axis with respect to gravity, see Fig. 6. 2) The base plane defined by the instantaneous centers for three vee constraints is perpendicular to the base cylinder. 3) Either constraint 1 or 2 as identified in Fig. 7 is engaged and has a CoF \u03bc1-2 and inclination angle \u03b11-2. 4) The remaining four constraints 3\u20136 are engaged and have the same CoF \u03bc3-6 and inclination angle \u03b13-6. 5) The external load including the nesting force is parallel to the base cylinder and distributed such that constraint 1 or 2 carries proportion \u03c9 and constraints 3\u20136 carry 1 \u2013 \u03c9. Clearly, \u03c9 is a dimensionless variables. 6) For each vee, the distance between the instantaneous center and the points of contact is r or in dimensionless form \u03c1 \u00bc r/R", + " 7) All the geometry defined above applies for the coupling virtually close to full engagement so that small angle approximations may be used and the shapes of contact surfaces do not make a substantial difference. The derivation of (1) begins by identifying the one-degree-of-freedom motion of the coupling and the sliding directions at the five engaged constraints. Two axes of rotation admissible for constraints 3\u20136 are apparent in Fig. 13: axis A is out of the page and identical to the instantaneous center demonstrated in Fig. 7; and axis B passes through the instantaneous centers for constraints 3\u20134 and 5\u20136. A fifth constraint at 1 or 2 ties rotation about axes A and B to a fixed ratio. Fig. 14-a shows the fifth constraint undergoing a unit displacement along its sliding direction and its components about axes A and B. Combining this with lever arms from Fig. 13, Equations (2) and (3) give corresponding rotations about axes A and B. \u03b8A\u00bc cos \u03b11 2 3 R (2) \u03b8B\u00bc 2 sin \u03b11 2 3 R (3) L.C. Hale Precision Engineering 64 (2020) 200\u2013209 The sliding directions at constraints 3\u20136 are less apparent, involving components in the plane of the vee and out" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0001579_0954407020945823-Figure1-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0001579_0954407020945823-Figure1-1.png", + "caption": "Figure 1. Structure of the multi-stage gear-type AT. AT: automatic transmission.", + "texts": [ + " Novel multi-stage gearbox working principle The hydraulically controlled multi-stage gearbox automatic gear shifting system designed in this study is mainly composed of an automatic gear shifting actuator and an automatic gear shifting hydraulic system. The novel multi-stage gearbox is mainly composed of transmission shafts and gears. The structure of the proposed hydraulically controlled multi-stage gear-type AT for vehicles is illustrated. The main functional components of the automatic gear shifting device and automatic gear selection device are designed in detail with an analysis of the novel structure of AMT shifting principle and transmission mechanism. The structure of the multi-stage gearbox is shown in Figure 1. The driving shaft is connected to the engine through coupling and rotates together. Meanwhile, the gear on the driving shaft is connected to the driving shaft through join keys. All the gears on the driving shaft simultaneously rotate with the driving shaft. The gears on the driven shaft are connected to the driven shaft through join keys. The number of gears on the driven shaft corresponds to the gears on the driving shaft, and all the gears on the driven and driving shafts mesh correspondingly" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0002917_b136378_7-Figure7.14-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0002917_b136378_7-Figure7.14-1.png", + "caption": "Fig. 7.14 Two-dimensional, cylindrical glucose sensor. Glucose diffuses to the sensor only in direction z whereas oxygen can reach the electrode by diffusion through the walls of the cylinder of length L", + "texts": [ + " The second problem is encountered when the electrode is used in vivo. The tissue or blood concentration of oxygen, which is the cosubstrate, is low and at high glucose concentrations the current becomes limited by the availability of oxygen. An elegant solution to this problem has been proposed (Gough et al., 1985). By design, the mass transport of oxygen has been increased, relative to that of glucose, by cylindrical diffusion into the enzyme layer and the transport of glucose restricted to the linear diffusion through the distal end of the sensor (Fig. 7.14). In other words, the area through which oxygen reaches the electrode has been made larger relative to the area for the access of glucose. The most important outcome of this design has been the elimination of the dependency on oxygen. The sensor responds only to glucose even in the case of low concentration of oxygen in the bulk. For this type of sensor, a catalase-free glucose oxidase must be used. In such a case, the hydrogen peroxide produced by the reaction with oxygen remains in the selective layer and can be detected by oxidation according to the reaction H2O2 = O2 + 2e\u2212+ 2H+ This sensor uses cylindrical microelectrode geometry (Fig. 7.14) for which the diffusion\u2013reaction reaction is written in spherical coordinates, similar to (2.24). \u2202CS \u2202 t = DS \u2202 2CS \u2202 r2 \u00b1 \u211cpHvmaxCS (CS + Km) (7.27) The sign in front of the reaction term is positive only for hydrogen peroxide. Also, the function \u211cpH can be made constant by operating the sensor in a medium of high buffer capacity. This is clearly a distinct advantage compared to the potentiometric sensors in which the buffer capacity represented a major interference. The initial and boundary conditions depend again on the model and on the operating conditions" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0003756_amm.509.86-Figure1-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0003756_amm.509.86-Figure1-1.png", + "caption": "Fig. 1. Clutch cover assembly and disc assembly diagram", + "texts": [ + " Reliability design is defined a project or product ability to fulfill the required function within the prescribed period of time and conditions and different from conventional design. The design 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: 136.186.1.81, Swinburne University, Hawthorn, Australia-09/07/15,08:47:49) reliability is one of the design goals and design variables are considered as random variables. This study researched on passenger vehicle clutch (shown in fig. 1) and proposed a set of system engineering analysis method (shown in fig. 2) using the modern reliability design method. The specific steps are as follows: 1) Collect the failure data and classify such as slipping, chatter, no release, damage, high pedal effort, etc. 2) Analyze the causes of clutch failures. The causes analysis of failures was carried out using FTA and FMEA combined with the mechanical design principles and function requirements. 3) Analyze causes of key components & parts. The key factors influencing the failure and components & parts performance were studied to find the corresponding relation of failure, system and parts design" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0003102_978-1-4419-1126-1_16-Figure16.3-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0003102_978-1-4419-1126-1_16-Figure16.3-1.png", + "caption": "Fig. 16.3 Illustration of the swimming micro robot\u2019s tail divided into three sections (Copyright 2007 IEEE.)", + "texts": [ + " In order to approximate the traveling wave one has to form the desired time functions g \u00f0d\u00de k \u00bc Gk sin\u00f0Ot Fk\u00de, by setting the phases and amplitudes of the input signal of the actuators. The phases Fi for all i \u00bc 1; n and amplitudes Gi for all i \u00bc 1; n of the actuator\u2019s input signal are derived from the solution of the beam model. When the motion and the geometry of the beam is small enough (for example a PZT bimorph with the size of 10 1 0.1 mm3) one can solve the beam\u2019s equations analytically. Otherwise a numerical model of FEA coupled with CFD or an equivalent mass spring model [82] should be used (Fig. 16.3). The theoretical model of a beam divided into n sections. The field equation of each section is: m1 @2w1\u00f0x; t\u00de @t2 \u00fe q1\u00f0x; t\u00de \u00fe K\u03021 @4w1\u00f0x; t\u00de @x4 \u00bc 08x \u00bc \u00bd0; a1L .. . 8i \u00bc 1; 2; ::: n mn @2w2\u00f0x; t\u00de @t2 \u00fe qn\u00f0x; t\u00de \u00fe K\u0302n @4wn\u00f0x; t\u00de @x4 \u00bc 08 x \u00bc \u00bdanL; L (16.2) m1 \u00bc Pn j\u00bc1 rijAij is the distributed mass of each elastic domain. In each domain, designated by index i, there are n layers. Each layer, designated by index j, has different cross sectional area, Aij (in the case of a rectangular cross section Aij\u00bc bijtij), and density, rij" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0003227_chicc.2015.7260767-Figure2-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0003227_chicc.2015.7260767-Figure2-1.png", + "caption": "Fig. 2: Multi-followers cross location", + "texts": [ + " Based on the bearing angle measurement, the position of the leader is determined by the multi-followers cross location and a consensus-based distributed extended information filter is developed for optimizing the estimating result in noise environment. Finally, sufficient conditions to achieve formation tracking with bearing-only measurement are presented. Let M1 represents the leader and Mi (i = 2, ..., N) represent followers. The planar coordinates of the leader and followers are denoted by p1 = [p1x(k), p1y (k)] and pi = [pix(k), piy (k)] (i = 2, ..., N) respectively, and the observed leader\u2019s bearing angle of Mi are \u03b2i(k)(i = 2, ..., N). Fig. 2 shows that the algebraic relation between the leader and followers and algebraic equations are given by tan\u03b2i(k) = p1y (k)\u2212 piy (k) p1x(k)\u2212 pix(k) , i = 2, ..., N (8) Next, from (8), we can get pi1x(k) = piy (k)\u2212pix(k)tan\u03b2i(k)\u2212pjy (k)+pjx(k)tan\u03b2j(k) tan\u03b2j(k)\u2212tan\u03b2i(j) , (9) pi1y (k) = (pjx(k)\u2212 pix(k)) tan\u03b2i(k) tan\u03b2j(k) tan\u03b2j(k)\u2212 tan\u03b2i(k) \u2212 pjy (k) tan\u03b2i(k) + pjy (k) tan\u03b2j(k) tan\u03b2j(k)\u2212 tan\u03b2i(k) , (10) i = j; i = 2, \u00b7 \u00b7 \u00b7 , N ; j = 2, \u00b7 \u00b7 \u00b7 , N. However, in practical environment, due to the existence of noise, direct mathematical calculation has large error" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0001931_j.jmmm.2020.167474-Figure13-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0001931_j.jmmm.2020.167474-Figure13-1.png", + "caption": "Fig. 13. The distribution of magnetic flux density obtained by time-domain FEM-BEM.", + "texts": [], + "surrounding_texts": [ + "In addition to the hybrid FEM-BEM in time-domain, the frequencydomain FEM is also employed for modeling of the hysteresis motor at steady-state. In frequency domain, the complex effective reluctivity is used for elliptical approximation of the hysteresis loop regarding the rotor hard material [9,23]. The complex effective reluctivity is specified as follows: \u03c5eff(|B| ) = \u03c5c(|B| )ej\u03b1h(|B|) (10) where \u03c5c is the absolute value and \u03b1h is the phase difference between magnetic field intensity and magnetic flux density. Both \u03c5c and \u03b1h depend on the magnetic flux density and are estimated from the B-H hysteresis loops of Fe-Co 17% (shown in Fig. 3) as: \u03c5c(|B| ) = Hrms/Brms (11) \u03b1h(|B| ) = arcsin( \u222e HdB/(2\u03c0HrmsBrms)) (12) where Brms and Hrms are rms values of B and H waveforms, respectively. Fig. 4 shows the variations of the absolute value \u03c5c and hysteresis angle \u03b1h, as a function of magnetic flux density. In fact, in the presence of eddy currents and frequency dependency of the hysteresis loop have been neglected in the frequency-domain FEM. Under these conditions, the hysteresis loop shape depends only on the magnitude of magnetic field at steady-state operation." + ] + }, + { + "image_filename": "designv11_71_0003544_cta.2127-Figure6-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0003544_cta.2127-Figure6-1.png", + "caption": "Figure 6. Connection of measuring self-inductance for the 7-coil equivalent circuit of the DC machine.", + "texts": [ + "5 hp, 1440 rpm, 16 slots, and 48 segments with lap winding, was especially for this research work. The so-called 7-coil or 5-coil equivalent circuits on the Copyright \u00a9 2015 John Wiley & Sons, Ltd. Int. J. Circ. Theor. Appl. 2016; 44:1094\u20131111 DOI: 10.1002/cta DC machine were established in [1]; the DCIB circuit is needed to be modified for each measurement. To measure the self-inductance of the six sub-coils in series in the machine for the 7-coil equivalent circuit, the DCIB is connected as illustrated in Figure 6. Figure 6 shows that the 16 armature sub-coils are connected with each other to form a closed loop. In order to measure the self-inductance of the six sub-coils in series for this 7-coil circuit, this connection to the DCIB was made for this purpose. It can be seen that the segments from 1C to 3C and from 9C to 11C are shorted out to carry zero current in this circuit. The series-connected coils 9C\u20133C and 11C\u20131C have 1A current flowing through them from the supply V1. Thus, the DCIB measures the self-inductance of two sets of the six sub-coils in parallel. In this diagram, the field circuit is not included, but it carries a 0.2A DC field current during the measurement. In order to establish the paralleled inductances being measured in Figure 6, it is necessary to first look at the direction of the flux produced by a current flowing over the sub-coils, referring to Figure 7 Using a similar concept to the convention in a generalized machine, the m.m.f. of a coil is along its axis. Therefor, the currents in the sub-coils 9C\u20133C and 11C\u20131C both produce m.m.f. and flux vertically, and they are of course mutually coupled. The physical meaning of the 7-coil and the 5-coil equivalent circuits [1, 2] is best visualized with the help of Figures 8 and 9, which is based on the commutating process on the DC machine over one revolution" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0000019_chicc.2019.8865756-Figure2-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0000019_chicc.2019.8865756-Figure2-1.png", + "caption": "Fig. 2: Local NED and body coordinate systems", + "texts": [ + " Therefore, from (36) one knows that V\u0307 (t) \u2264 0 if \u03beT (t)\u03a6i\u03be(t) + \u03b32\u03b2i(\u03c3,\u0393) \u2264 0, which implies that \u03bei(t) will converge into the set S1 = {\u03bei : |\u03beTi \u03a6i\u03bei| \u2264 \u03b32\u03b2i(\u03c3,\u0393)}, and this yields that the error {e\u0304i(t) : i = 1, 2, \u00b7 \u00b7 \u00b7 , N} will converge to a small ball around the origin. This completes the proof. To illustrate the effectiveness of the proposed FTC ap- proach, a simulation study on a practical UAVs formation is provided in this section. A swarm consisting of four Kyosho 260 unmanned helicopters (see Fig. 1 and Fig. 2) have been used for simulation testing and analysis. The communica- tion topology graph and desired formation configuration of the four UAVs are given in Fig. 3 and Fig. 4, respectively. trol problem has been investigated for a class of leader- follower MASs subject to actuator faults. A decentralized state observer and fault estimator based formation control protocol is proposed, together with the sufficient conditions under which the desired formation can be reached. It is proved that formation errors of all the following agents can converge to a small set around the origin" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0002119_icaccm50413.2020.9212874-Figure4-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0002119_icaccm50413.2020.9212874-Figure4-1.png", + "caption": "Fig. 4. Fixed support (case1).", + "texts": [ + " Ansys is based on the theory of FEA, which provides the user friendly interface to check the product performance under various boundary conditions. Following are the main objectives of the present study. A. Propeller bending due to thrust force (case1) B. Propeller bending due to rotational velocity (case2) C. Vibration analysis (case 3) The propellers are subject to bending in the vertical direction due to applied thrust force F. To analyze the propeller under thrust the propeller is fixed at the point where it connected to the motor shaft as shows in Figure 4. 1 2 3 4F mg F F F F (1) 2 i t iF K (2) Where m is the total mass (Quadcopter and payload) and g is acceleration due to gravity. Kt is the coefficients, depends upon propellers geometry and air density. If the value of m is 20 kg then the value of the total thrust force is calculated as follows. Total Thrust force (F)= mg= 20*9.81=196.2 N Thrust force required per propeller= F/4 =196.2/4 =49.05 N The calculated thrust force is approximately 50 N for a single propeller. Thus, the propeller is subjected to 50 N of thrust force to find out the deformation and stress due to the bending effect as shown in Figure 5" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0000137_sled.2019.8896287-Figure5-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0000137_sled.2019.8896287-Figure5-1.png", + "caption": "Fig. 5. Stator and rotor design.", + "texts": [ + " Subsequently, after some time steps, the control angle is set according to the MTPA requirement. V. SIMULATION RESULTS In order to verify the proposed approach a real-time simulation environment based on an ezDSPTMS320F2812 has been used. The control algorithm replicates exactly that of an actual drive, except that the motor is simulated by its differential model. The model of the motor uses a FEM mapping to simulate correctly the real behavior of the machine, the saturation and the cross-coupling effects. The simulated machine is shown in Fig. 5, while its rated values are in Table I. The load is simulated by the and parameters of the experimental set-up presented in the next section ( =0.3 and =0.2). Fig. 6 compares the behaviors of the actual , currents, computed by the motor model with , , , , the control feedbacks used during the open-loop start-up i.e. calculated using the reference angle. Initially, after the alignment, the actual and the feedbacks currents are equal. Later, during the acceleration, both , and , maintains the level imposed by the high band-width control, while the actual currents vary with the same oscillatory behavior (see Fig" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0000830_0954410020911832-Figure3-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0000830_0954410020911832-Figure3-1.png", + "caption": "Figure 3. Actual and equivalent canard models in the tail view.", + "texts": [ + " The air-relative velocity respect to the NRRF can be expressed as vr vr vr 2 64 3 75 \u00bc ANV vrx2 vry2 vrz2 2 64 3 75 \u00bc ANV v wx2 wy2 wz2 2 64 3 75 v wx2 wy2 wz2 \u00f0v wx2\u00de wy2 \u00f0v wx2\u00de wz2 2 64 3 75 \u00f08\u00de Substituting vr , vr , vr into the vector statements of moments illustrated in McCoy,24 the uncontrollable moments can be expressed as: Overturning moment: MON \u00bc 0:5 vrSdCM 0 vr vr T; Spin damping moment of the aft body: MSDN \u00bc 0:5 vrSd 2pAClpA 1 0 0 T ; Magnus moment: MMN \u00bc 0:5 Sd2\u00f0pAC A mp \u00fepFC F mp \u00de 0 vr vr T ; Pitch damping moment: MPDN \u00bc 0:5 Sd2\u00f0Cmq \u00fe CM _ \u00devr r tan q r T . To sum up, the total uncontrollable moment can be expressed as For simplicity and universal description of different actuators with canards, the actual canard model and the equivalent canard model established with respect to the NRRF are shown in Figure 3, in which F denotes the roll angle of the canards. Canards 1 and 3 in the actual canard model shown in Figure 3 as the real pitch canards are defined, and canards 2 and 4 are the real yaw canards. Similarly, canards A and B in the equivalent canard model illustrated on the righthand side of Figure 3 are defined as the equivalent pitch canards, and canards C and D are the equivalent yaw canards. The deflection angle of the yaw and pitch canards of the actual canard model are recorded as AY, AP; those of the equivalent canard model are recorded as EY, EP respectively. Considering that the canard force vectors of the actual and equivalent model should be identical, the relation between the deflection angles of these two models can be written as EP EY \u00bc cos F sin F sin F cos F AP AY \u00f010\u00de The magnitude of the total equivalent deflection angle CE and the direction angle E are defined as follows CE \u00bc ffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi 2EP \u00fe 2 EY q E \u00bc arc tan 2\u00f0 EY, EP\u00de \u00f011\u00de The full range arc-tangent function of arc tan 2\u00f0a, b\u00de is defined as arc tan 2\u00f0a, b\u00de \u00bc tan 1\u00f0a=b\u00de, if b4 0 tan 1\u00f0a=b\u00de \u00fe , if b5 0 =2, if a4 0 and b \u00bc 0 3 =2, if a5 0 and b \u00bc 0 8>>>< >>>: The canard force and moment with atmospheric wind can be created by equivalent canards expressed in the NRRF without wind as follows Fc Fc Fc 2 64 3 75\u00bc Sv2rC cN 2 C cA \u00f01\u00feK 2CE\u00de=C c N CE cos E CE sin E 2 64 3 75\u00fe 0 r r 2 64 3 75 8>< >: 9>= >;; Mc Mc Mc 2 64 3 75\u00bc Sv2rC cNlCG 2 0 CE sin E CE cos E 2 64 3 75\u00fe 0 r r 2 64 3 75 8>< >: 9>= >; \u00f012\u00de where c represents the relative angle of attack and sideslip, respectively" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0002805_012042-Figure1-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0002805_012042-Figure1-1.png", + "caption": "Figure 1.Four-bar linkage in the global coordinate system [1]", + "texts": [ + " The rest of this paper starts from Section 2 that details the position analysis of a four-bar linkage in brief, objective function, and the constraint handling technique. The design problems and optimiser are presented in Section 3. The design results are detailed in Section 4. The conclusions and discussion of the study are summarized in Section 5. A model of a four-bar linkage in this study is composed of four simple links connected with 4 simple joints. A variety of linkage types occur when assigning anyone link to be a frame. This linkage has one degree of freedom, which can operate with one actuator. The kinematic diagram of this linkage is shown in Fig. 1. The position analysis of the four-bar linkage uses a trigonometric formula for the relation of linkage lengths r1, r2, r3, andr4 and its other parameters, which is found in standard mechanics of machinery textbooks as mentioned in [1, 8]. The coupler point (P) in the global coordinate can be expressed as xP = xO2 + r2cos(2 + 1) + L1cos(0 + 3+ 1) (1) yP = yO2 + r2sin(2 + 1) + L1sin(0 + 3 + 1) wherexO2and yO2 are the coordinates position of the O2 in the global coordinates [8].The angles0,3, 4, and inside the kinematic diagram relate the known link lengths r1, r2, r3, and r4 at any crank angle (2) by law of cosine" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0002279_asemd49065.2020.9276061-Figure1-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0002279_asemd49065.2020.9276061-Figure1-1.png", + "caption": "Figure 1. Magnetic flux path. (a) Typical U-shaped segmental stator SRM flux path (b) inner stator flux path (c) outer stator flux path", + "texts": [ + " In order to solve the problem of large torque ripple, it is necessary to decouple the magnetic field generated by inner and outer stator windings, so that the inner and outer stator windings can be separately controlled. II. STRUCTURE OF THE PROPOSED DSSRM In order to realize the decoupled the magnetic fields between the inner and outer stator, a parallel magnetic circuit is needed to make the magnetic field produced by the inner stator repel the magnetic field produced by the outer stator. The flux path of the U-shaped stator segmental SRM is shown in Fig.1 (a), it can be seen that the magnetic field generated by the stator is closed through the stator teeth, two adjacent rotor teeth and the stator yoke, which is characterized by good magnetic circuit independence. According to the characteristics of the magnetic circuit in parallel, and the characteristics of the Ushaped magnetic circuit of inside and outside of the stator flux path respectively as shown in Fig.1 (b) and (c), it can be seen that the magnetic field generated by the inner stator passes through the closing of the inner teeth, rotor yoke and inner stator teeth of the two adjacent rotors, while the magnetic field generated by the outer stator passes through the closing of the outer teeth, rotor yoke and outer stator teeth of the two adjacent This work was supported by National Natural Science Foundation of China-NSFC-ASRT (Chinese-Egyptian) Cooperative Research project under Grant 5191101616, Key R&D Program of Jiangsu Province (Industry Foresight and Common Key Technologies)-Priority projects-subject under Grant BE2018001-3, the Project funded by China Postdoctoral Science Foundation under 2020M671642, and the Xuzhou Social Development-Basic Application Research Project under Grant KC18064" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0001626_ccdc49329.2020.9164274-Figure2-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0001626_ccdc49329.2020.9164274-Figure2-1.png", + "caption": "Fig 2. Aerodynamic layout and rudder surface distribution Analyze the aerodynamic characteristics of canard, flap and elevator.Figure 3 shows the pitching moment coefficient and lift drag ratio of three rudder surfaces at Mach 0.2 and 300m. By comparison, under the same rudder deflection, the pitching moment of the flaps is larger, while that of the canard is smaller.", + "texts": [ + " In order to control the plane deviation on the scheduled flight trajectory, aircraft can match the canard rudder, elevator and rudder flap. By adopting the lift mode that keeps the Angle of attack unchanged, the lift force required to make the track change is completely provided by the direct lift force generated by the control surface, so as to ensure that the carrier-borne aircraft can quickly complete the trajectory control under high control accuracy. 2.2 Dynamics Analysis In this paper, the research object is the Admire aircraft, whose aerodynamic layout and rudder surface distribution are shown in Figure 2.The two outer rudders behind the wings are elevators and the two inner rudders are flap rudders. The influence of canard rudder on lift can be neglected because the lift force produced by canard rudder is small and the rudder deflection has little influence on lift force. However, the lift force produced by the flaps and ailerons is larger, and the deflection of the rudders have a greater impact on the lift force of the body. When the lift force of the aircraft changes, the pitching moment changes with it The canard rudders with less influence on lift are selected as the balance rudders surface to offset the pitching moment and keep the attitude unchanged" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0000605_978-81-322-3965-9_9-Figure3-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0000605_978-81-322-3965-9_9-Figure3-1.png", + "caption": "Fig. 3 Electron transfer mechanism from the microbes to the anode", + "texts": [ + " The voltage of a MFC is thus the difference between the potential of the cathode and that of the anode and can reach themaximum theoretical limit of 1.2 V [0.8 V \u2013 (\u22120.4 V)]. Typical redox gradients with respect to SHE can be found in Brock Biology of Microorganisms 13th edition (Madigan et al. 2010). At the anode when microbes form a colony they start to metabolize sugar and other nutrients, which act as fuel to the microorganisms. The process of metabolism generates highly reduced biomolecules with extra electrons attached to them. These extra electrons are then transferred to the anode generally in one of the following three ways (Fig. 3): (1) Electrons can be transferred directly from the microbe\u2019s wall to the anode, (2) Mediator assisted electron shuttling from the biomolecule to the anode and (3) Electron transfer throughmicrobial nanowires or conductive appendages grown by the microbes (Gorby et al. 2006). A schematic of networks of nanowires in microbes can be seen in Fig. 4. Sometimes electrons are carried directly from the respiratory enzyme to the electrode when the microbe cells expose the redox active proteins on the surface of the electrode" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0002324_iceca49313.2020.9297379-Figure1-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0002324_iceca49313.2020.9297379-Figure1-1.png", + "caption": "Figure 1: Isometric View", + "texts": [ + " Solid wings (2R and 2L) are mounted such that they remain perpendicular to the frame (1). During take-off and hover, the aircraft will be aligned such that frame (1) remains parallel and the wings (2R and 2L) remain perpendicular to the ground. During glide, the wings (2R and 2L) will remain parallel to the ground. The same 4 BLDC motors (M1, M2, M3 and M4) will provide thrust for the aircraft to glide and will be used to maneuver the aircraft. The wings contain the rectangular shape and the airfoil consists of Clark Y shown in Fig.1. The dimensions include the total length of 1000m, width of 200 mm, and max height of 20 mm. For aircraft to make a naturally stable flight, the CG (Center of Gravity) is kept ahead of the CP (Center of Pressure of the 978-1-7281-6387-1/20/$31.00 \u00a92020 IEEE 105 Authorized licensed use limited to: East Carolina University. Downloaded on June 15,2021 at 10:48:38 UTC from IEEE Xplore. Restrictions apply. main wing). Keeping CG nearer to CP increases the risk of flipping the aircraft when there is a sudden gust of wind" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0001701_0954405420949757-Figure2-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0001701_0954405420949757-Figure2-1.png", + "caption": "Figure 2. Discrete points of face line on axial section.", + "texts": [ + " Homogeneous transformation matrix from cutter coordinate system to pinion coordinate system is expressed as M10 = (M (1) m1M (2) m1) 1Mm0 \u00f08\u00de Relative velocity of the locus r in pinion coordinate system is expressed as v= \u2202M10 \u2202t r \u00f09\u00de Suppose that r1 is the locus of pinion conjugated tooth surface, according to gear mesh theory, r1 can be determined by following equations r1 =M10r v (M10n)=0 \u00f010\u00de When the variable s1 is eliminated by equations above, r1 is determined by t and u1, which can be written as r1(t, u1). After r1 is obtained, unit normal vector of pinion conjugated tooth surface is expressed as n1 =M10n \u00f011\u00de Coordinates calculation of discrete points on tooth crest line Discrete points planning on axial section An irregular quadrilateral ABCD is obtained when tooth surface is projected to axial section, tooth crest line corresponds to face line, as shown in Figure 2. Two endpoints of face line is denoted by A and B, whose coordinates are written as (xA, 0, zA) and (xB, 0, zB). Discrete points are obtained by evenly dividing face line, whose number is denoted by m. The i-th point is denoted by Pi, whose coordinates are written as (xi, 0, zi). xi and zi are calculated as xi = xA + (xB xA)(i 1)=(m 1) zi = zA + (zB zA)(i 1)=(m 1) \u00f012\u00de Discrete points on tooth crest line A circle is formed when point Pi rotates around Z1 axis, the circle intersects tooth surface of convex side and concave side at P0i and P00i in Figure 3, whose coordinates are denoted by (x0i, y 0 i, z 0 i) and (x00i , y 00 i , z 00 i )", + " 0 \u00f014\u00de This problem is solved by particle swarm algorithm, number of particles is set to 50, number of iterations is set to 10,000, solving convergence condition is f\\ 10 4mm. When subscript value i varies from 1 to m, discrete points on tooth crest line of convex side are calculated in turn, then discrete points on tooth crest line of concave side are calculated by the same way. Tangent vectors of discrete points Point Of is the intersection of Z1 axis and extension line of segment BA, which is shown in Figure 2. Let r0i and r00i be the vectors defined by O1P 0 i ! and O1P 00 i ! , f0i and f00i be the vectors defined by OfP 0 i ! and OfP 00 i ! , n0i and n00i be unit normal vectors of convex side and concave side at P0i and P00i . Unit normal vectors of face cone at P0i and P00i are denoted by c0i and c00i , which are calculated as c0i = f0i 3 (r03 f0i)= f 0 i 3 (r03 f0i)j j c00= f00i 3 (r03 f00i)= f 00 i 3 (r03 f00i)j j \u00f015\u00de Let t0i and t00i be unit tangent vectors of tooth crest line at P0i and P00i , which are calculated as follow t0i = c0i 3 n0i= c 0 i 3 n0ij j t00i = c00i 3 n00i= c 00 i 3 n00ij j \u00f016\u00de c0i, c 00 i , n 0 i, n 00 i , t 0 i and t00i are shown in Figure 4", + " Suppose that when the values of X axis and Z axis change from 0 to a positive value, X and Z axis sliding tables move along positive X and Z axis respectively. When the value of Y axis changes from 0 to a positive value, Y axis sliding table moves along negative Y axis. When axial cutter position detecting is performed, Y axis sliding table moves to make cutter axis and A axis intersect, then the value of Y axis is set to 0 in CNC system. X and Z axis sliding tables move to make edge detector and inner end of pinion contact, the distance between inner end and point O1 is denoted by d in pinion coordinate system shown in Figure 2, then the value of X axis is set to rc d. Circumferential cutter position detecting When circumferential cutter position detecting is performed, a convex side tooth crest edge is selected as detecting datum. Suppose that pinion is installed on chunk as the position relationship shown in Figure 16, when rotation angle of A axis is kept unchanged, X, Y and Z axis sliding tables move to make edge detector and the convex side tooth crest line contact, the cutter position determined by X, Y and Z axis sliding tables is called circumferential cutter position" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0002048_ecce44975.2020.9235328-Figure1-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0002048_ecce44975.2020.9235328-Figure1-1.png", + "caption": "Fig. 1: FEA Model (a) 2-D, (b) 3-D; Designed prototype (c) stator, and (d) rotor.", + "texts": [ + " Finally, the experimental validation has been done by using an innovative testing setup in the rotor of a 5-phase permanent magnet assisted synchronous reluctance motor (PMaSynRM). II. PM AXIAL TEMPERATURE ESTIMATION BY A-LPTM Lumped parameter-based thermal analysis has been widely used to estimate the electrical machine's real-time temperature due to its less implementation complexity, less cost, and less computation time. Also, it is possible to use during the loading condition, and it can be modeled for almost all types of machines. Fig. 1 shows the FEA models and designed prototype (stator and rotor) of a 5-phase PMaSynRM [23], which is used in this study to develop the A-LPTM model. The machine specifications are given in the Table I. TABLE I: PMaSynRM SPECIFICATIONS Parameter Value Design Parameter Value Rate speed 1800 rpm Stack length (mm) 65 mm Rated power 3 kW Slot/ Pole 25/4 Rated current 15.17 A Rotor outer radius 95 mm Rated Voltage 67 V Air gap length 0.7 mm Rated Torque 15.45 Magnet Material NdFeB Phase Number 5 Core Material 50PN470 The core lamination structure prevents the heat flow along the axial direction because it has a very low thermal conductivity of air (\u2248 0", + " 3, shows an overview of the novel A-LPTM, which is used in this paper. The model is a quasi 3D model, and the nodes are considered functions of the position along the axial length (x) along which they span, as shown in the figure. In this model, current sources represent the heat sources generated by the corresponding loss, and these are the input of the thermal model. Copper loss, core loss, and magnet loss have mainly been considered in this formulation. FEA software is used to calculate these losses. Fig. 1(a) and (b) present the 2-D and 3-D model of the machine used in the FEA software. Copper loss is calculated based on the input RMS current and phase resistance. Core loss and magnet eddy current loss are directly calculated from FEA. A specific loading condition (RMS phase current and rotational speed) and losses corresponding to that loading condition are applied as input to the LPTM, which is developed in a MATLAB script. Fig. 3(a) and (b) show the estimated temperature by R-LPTM and A-LPTM at 5 A 1125 RPM and 15A 1800 RPM, respectively", + " ANALYSIS OF MAGNET AXIAL TEMPERATURE VARIATION EFFECT To investigate the effect of the magnet axial temperature variation on the machine's electromagnetic performance, four different loading conditions have been considered in table II. Fig.4 indicates the minimum, maximum, and average (radial temperature) temperature point of a magnet axial temperature distribution curve. The process algorithm to analyze the effect of axial temperature variation on a machine performance is shown in Fig. 7. At first, both the 2-D and 3-D FEA model of the 5- phase PMaSynRM have been designed. A five-phase winding configuration and machine design parameters are shown in Fig. 1 and Table 1, respectively. Then the electrical losses are calculated by electromagnetic analysis for the specified loading condition shown in Table II. Per-phase dc resistance is needed in FEA to calculate the copper loss, which is measured from the machine prototype. Also, for core loss calculation in FEA, loss coefficient (eddy, hysteresis, and excess) are obtained by providing different frequency vs loss curve of the core material (S_50PN470). In the next step, these loss data are used as input to the FEA thermal model and obtain the magnet axial temperature distribution" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0003617_ipec.2014.6870118-Figure4-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0003617_ipec.2014.6870118-Figure4-1.png", + "caption": "Fig. 4. Two-phase 1M model.", + "texts": [ + " This is because the ends of the mover core exist in case of the LIM. Fig. 3 shows operational impedance loci obtained by the DC decay testing method for a-n, b-n, c-n, a-b, b-c nd c-a winding terminals. The details of the DC decay testing method are shown in [3] . From the figure, it is con -firmed that Xaijs) = Xhn(js) = Xen(js) and Xhe(js) < Xah(js) = J('a(js) . This result agrees with the magnitude relation of mutual reactances guessed from the structure of Fig. 2. III. Simple Two-Phase Model Fig. 4 shows a circuit model of a two phase rotational type induction motor. Here, rl and rz are primary and secondary resistances. II and Iz are primary and secondary leakage inductances. Llmd and LZmd are primary and secondary d-axis inductances based on the main magnetic flux. Llinq and LZmq are primary and secondary q-axis inductances based on the main magnetic flux. \ufffd'e is the rotor electrical angle, and OJ,.e is the angular velocity. In this paper, LIM is regarded as a two phase induction motor with the following conditions", + " 3) In af3 stational reference frame shown in FigA, the permeance of the magnetic path of a-axis main magnetic flux differs from that of f3-axis main magnetic flux. As a result, a-axis mutual inductance md is not equal to fJ-axis one mq- The reasons of \\) and 2) can be explained from the symmetric nature of the operational impedances of a phase, b-phase and c-phase (XanCis) '\" XbnCis) '\" XenCis\u00bb. The reason of 3) can be done from the asymmetric nature of the operational impedances between a-b, b-c and c-a terminals (XhcCi s) < XahCi s) '\" XcaCi s\u00bb. The voltage equation of the circuit model of Fig. 4 is as follows [5]: k/ ]= [z'k/ ] = [R'k/ ]+ p([L'k/ ])+ p([M'k/ ]) where p is a differential operator (p = dldt), k/ ]= [vla vl,Li V2a , v2/V [i,,/ ]= [ila il,Li i2a , . 'V 12,Li (\\) (2) (3) * mel cos ere mq sin ere 2 e . 2 e md cos re + mq sm re -md sin ere mq cos ere -\ufffd (md -mq )sin 2ere \u2022 2 e 2 e md sin re + mq cos re If the following coordinate transformation, 1 0 [c]= 0 1 0 0 0 0 0 0 cos ere sin ere 0 0 - sin ere cos ere is applied to (\\), (\\) can be changed as follows: [c ][v\"p' ] = [C ][Z'][i\"p' ] [c]k/ ] = [c][z'Icr " + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0002759_b978-0-12-804560-2.00011-0-Figure4.7-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0002759_b978-0-12-804560-2.00011-0-Figure4.7-1.png", + "caption": "FIGURE 4.7 Model of a serial-link manipulator on a free-floating base. (A) Left: main coordinate frames and system parameters. (B) Right: representation as a composite rigid body (CRB) (all joints locked).", + "texts": [ + " Also, the so-called \u201cmanipulator inversion task\u201d was introduced wherein the end link remains fixed in inertial space while the orientation of the floating base varies in a desirable way. This behavior is achieved via a specific null space: that of the coupling inertia matrix. Later it was shown that the same null space can also play an important role for balance control of a humanoid robot subjected to an external disturbance [102,101,168]. Consider a free-floating serial-link chain in zero gravity comprising an n-joint manipulator arm mounted on a rigid-body satellite, as shown in Fig. 4.7A. The satellite represents the floating base of the system. It is assumed that the base is not actuated (there are no thrusters or RWs) while the manipulator joints are. Thus, the system is underactuated. The linear momentum of the system is expressed uniquely as p = n\u2211 i=0 Mi r\u0307 i = M r\u0307C, (4.70) Mi and r i denoting the mass and CoM position of Link i; M stands for the total mass and rC is the position of the system CoM. The expression for the angular momentum depends on the point of reference. In the field of space robotics, expressions w", + " The CRB is characterized also by an inertia tensor defined as IC(q) \u2261 n\u2211 i=0 ( I i + Mi[r\u00d7\u2190\u2212 iC ][r\u00d7\u2190\u2212 Ci ] ) \u2208 3\u00d73. (4.73) Hereby, r\u2190\u2212 iC = r i \u2212rC and q = (XB, \u03b8) denotes the generalized coordinates of the floating-base system, XB standing for the 6D position of the floating base. At this point it is instructive to note that the states of the manipulator joints and the floating base are derived from the joint angle encoder and the inertial measurement sensor (IMU) readings, respectively. Sensor fusion techniques may increase the accuracy of the latter [126]. Furthermore, as shown in Fig. 4.7B, the CRB coordinate frame {C} is conveniently attached to the system CoM. Since the CRB is not a real body, the coordinate axes are chosen to be parallel to those of the base-link frame {B}. With this choice, the spatial velocity of the CRB can be expressed by VM (defined in Section 2.11.4), with components CoM velocity vC = r\u0307C and the angular velocity of the base link \u03c9B . As an alternative expression, the so-called system spatial velocity VC = [ vT C \u03c9T C ]T will be introduced here. Angular velocity \u03c9C is referred to as the system angular velocity5; its meaning will be clarified in Section 4" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0003488_msec2015-9243-Figure8-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0003488_msec2015-9243-Figure8-1.png", + "caption": "Figure 8: Simulation interface", + "texts": [ + " LMDCAM2 therefore provides a function to smooth the normal vectors in respect to the geometry of the surface (see Figure 7, bottom). The user is able to define a radius in which the smoothing will be applied. Afterwards, LMDCAM2 captures the geometry of the part lying in the circle with the given radius around the point. This geometry is afterwards fitted with a plane via least square fitting whose normal vector is used as the smoothed normal vector for the regarding point. For predicting the machine movement and checking for collisions for a given tool path, LMDCAM2 features a simulation interface (see Figure 8). A model of the machine can be loaded into LMDCAM2 and is mapped into the software. This model has to include all kinematic data (like pivot points for the different axes). The axes of the machine can now be addressed by LMDCAM2 and the model will get rendered depending on the state of these axes. For each point of the trajectory, an inverse kinematic transformation is applied to calculate the state of the axes from the Cartesian coordinates and the normal vector for each point. Certainly, the simulation interface is still under development" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0001511_jsyst.2020.3006990-Figure4-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0001511_jsyst.2020.3006990-Figure4-1.png", + "caption": "Fig. 4. Examples of the paths obtained by the 3PA with the CSCS configuration.", + "texts": [ + " , (n+ 1)/2; 2) the solution of the three-point Dubins path from (a2k\u22121, \u03b82k\u22121) to a2k+1 through a2 k provides the orientations \u03b82 k and \u03b82k+1 as well; 3) another three-point Dubins path from (a2k+1, \u03b82k+1) to the next two viewpoints is obtained and the process is repeated to cover all the viewpoints; 4) when the number of viewpoints is even, the last segment will only be a two-point path; 5) the paths have one of the following four structures: CSCS,CSCC,CCCS, and CCCC but when the distance between every pair of viewpoints on the path is greater than 2rmin, the optimal path will have a CSCS form; Authorized licensed use limited to: Carleton University. Downloaded on August 04,2020 at 01:46:27 UTC from IEEE Xplore. Restrictions apply. 6) the midpoint bisects the turning arc. The illustrative examples in Fig. 4 represent two possible forms of the CSCS configuration. The LAA was introduced in [22], as receding-horizon-based control scheme. The idea behind this algorithm is to use the solution of the 3PA to determine only the path and orientation up to the middle viewpoint. Consequently, the solution for (a1, \u03b81), a2, and a3 is used to determine \u03b82. Note that the choice of \u03b82 will be heavily influenced by the location of a3 in the 3PA Dubins path solution. Once \u03b82 is known, the tour can be extended by solving another 3PA problem starting from (a2, \u03b82)" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0000885_s42947-020-0303-x-Figure2-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0000885_s42947-020-0303-x-Figure2-1.png", + "caption": "Fig. 2. Hamburg Wheel Tracking Test (a) HWTD device and (b) HWTD test results.", + "texts": [ + " The Hamburg wheel tracking test was conducted in accordance with the AASHTO T 324-11 [24] test standard. In the HWTD test, a steel wheel (0.71 kN) with 200 mm diameter and 47 mm width moves across (52\u00b12 passes per minute) a pair of asphalt mixture specimen submerged in water at approximately 50\u00b0C [25-27]. Typically, 20,000 number cycles of wheel passing is allowed over the specimens, and rut depth is recorded through a linear variable displacement transducer (LVDT). The HWTD test setup is shown in Fig. 2(a). A typical rut depth vs. number wheel passes (Fig. 2(b)) contains a post compaction consolidation point, creep slope, and stripping slope, which are defined as the inverse of the rate of deformation in their corresponding region. Post-compaction consolidation is defined as the rut depth at 1,000 load cycles. Creep slope is the inverse of the rutting slope after post-compaction consolidation but before the stripping inflection point. Creep slope can be used to evaluate the rutting potential of HMA mixes. Stripping inflection point is the point at which the creep slope and stripping slope-intercept" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0003208_9781118751992.ch4-Figure4.11-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0003208_9781118751992.ch4-Figure4.11-1.png", + "caption": "Figure 4.11 Schematic diagram showing the electroclinic effect in the smectic-A*.", + "texts": [ + " Regarding the bistability, on the one hand, it is good because it enables multiplexed displays of the ferroelectric liquid crystal on passive matrices; on the other hand, the bistability is a problem because it makes it difficult to produce gray scales. Another issue with SSFLC is that it is more challenging to achieve uniform orientation in SSFLC than in nematic liquid crystals. As discussed in Section 3.1, there is no ferroelectricity in chiral smectic-A crystals (smectic-A consisting of chiral molecules, denoted as smectic-A*). In a cell geometry of smectic-A* liquid crystal, as shown in Figure 4.11(b), at zero applied field, the liquid crystal director is perpendicular to the smectic layers. The transverse dipole moment has equal probability of pointing any direction in the smectic layer plane because of the unbiased rotation of the molecule along its long molecular axis. When temperature is lowered toward the smecticA*\u2013smectic-C* transition, short length-scale and time-scale domains with smectic-C* order form, because the tilt of the director away from the layer normal direction does not cost much energy", + " This effect of field-induced tilt of liquid crystal molecules in smectic-A* is known as the electroclinic effect. The same as in smectic-C*, the spontaneous polarization is perpendicular to the plane formed by the liquid crystal director n! and the smectic layer normal a!. Therefore, in the induced smectic-C* structure, the n ! a ! plane is perpendicular to the applied field. When a DC electric field pointing down is applied, smectic-C* structure, say, with positive tilt angle is induced as shown in Figure 4.11(a). When a DC electric field pointing up is applied, smectic-C structure with negative tilt angle is induced as shown in Figure 4.11(c). As shown in Section 4.3.2, when the tilt angle is \u03b8, the spontaneous polarization P ! s is given by Equation (4.45), and the electric energy density is \u2212P ! s E ! = \u2212cE\u03b8. The free energy density is [15] f = fo + 1 2 a T \u2212Tc\u00f0 \u00de\u03b82 + 1 4 b\u03b84\u2212cE\u03b8: \u00f04:52\u00de The tilt angle as a function of the applied electric field E can be found by minimizing the free energy: \u2202f \u2202\u03b8 = a T \u2212Tc\u00f0 \u00de\u03b8 + b\u03b83\u2212cE\u22610 \u00f04:53\u00de When the applied field is low, the tilt angle is small, and the cubic term in Equation (4.53) can be neglected" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0002759_b978-0-12-804560-2.00011-0-Figure4.8-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0002759_b978-0-12-804560-2.00011-0-Figure4.8-1.png", + "caption": "FIGURE 4.8 External wrench f E = 0, mE = 0 acting on the CRB. Left (A): The angular momentum is conserved when the line of action of f E passes through the CRB centroid. Right (B): When the line of action of the external force does not pass through the CoM, the force induces an angular momentum with rate of change in proportion to the distance between C and E.", + "texts": [ + " The CRB motion induced by the impressed external wrench depends on the line of action of the force component. A special case deserves to be mentioned. Let the external wrench, acting at point E, represent a pure force, i.e. FE = [ f T E 0T ]T . When the line of action of this force goes through the CRB centroid, the angular momentum will be conserved since the moment component of FC = T T\u2190\u2212 CE FE is identically zero; mC = 0. On the other hand, when the line of action is apart from the centroid, a change in the angular momentum will be induced (cf. Fig. 4.8), so d dt lC \u2261 mC = \u2212[r\u00d7\u2190\u2212 CE ]f E = 0. Assuming an external force of a unit magnitude, the magnitude of the rate of change of the angular momentum will be in proportion to the distance \u2016r\u2190\u2212 CE \u2016. This simple relation plays an important role in balance stability of humanoid robots (cf. Chapter 5). Finally, to obtain the system dynamics expressed in terms of the centroidal quasivelocity, adjoin the decoupled CRB dynamics (4.149) to the reduced-form dynamics (4.133). We have[ MC 0 0 M\u0302\u03b8B ][ V\u0307C \u03b8\u0308 ] + [ CC c\u0302\u03b8B ] = [ 0 \u03c4 ] + \u23a1\u23a3 T T\u2190\u2212 CE J\u0302 T E \u23a4\u23a6FE, (4" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0001346_tee.23184-Figure2-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0001346_tee.23184-Figure2-1.png", + "caption": "Fig. 2. Representation of 6/4 PM-SRM with the AHB converter", + "texts": [ + " Phase torque developed in the motor can be found by partial derivation of coenergy with respect to the rotor position: Tk = \u2202W \u2032 k (ik , \u03b8) \u2202\u03b8 ik =const (2) where W \u2019k is the coenergy of phase k th and can be obtained from the partial integration of flux linkage with respect to the phase current: W \u2032 k = \u222b \u03bbk (ik , \u03b8)ik \u03b8=const (3) Assuming magnetic linearity, phase torque can be simplified as follows: Tk = 1 2 i 2 k dLk d\u03b8 \u2212 1 2 \u03d52 m dRgk d\u03b8 + N ik d\u03d5mk d\u03b8 (4) In this equation, \u03c6m is the flux going out from the magnet, \u03c6mk represents the magnet flux coupled with the k th winding; Lk and Rk are inductance of the phase winding and the reluctance seen from the kth winding, respectively; N is the turn number of a phase winding. The total torque developed in the machine is the sum of phase torques: Tm = \u2211 k=A,B ,C ... Tk (5) Like in SRMs, the PM-SRMs also require power electronic circuit for continuous rotation. Figure 2 displays the three-phase PM-SRM with an Asymmetric Half Bridge (AHB) Converter, which is often accepted to be the classical topology for SRMs. 2.2. The DITC Figure 3 illustrates the simple block diagram of the DITC system. The DITC is constructed on generating gate signals from a hysteresis controller whose input is defined by the error T between the reference torque T ref and the motor torque T m . The performance of the DITC algorithm is strongly depending on accuracy of the estimated parameters such as phase currents and rotor position", + " FEM tool only calculates nonlinear total torque developed in the motor and phase currents to the Simulink while receiving phase voltages and rotor position to Simulink. An s-function block is employed for this purpose. Once phase currents and required gate signals for speed and torque control are known, corresponding phase voltages are obtained from the operation of AHB converter to establish a loop between FEM tool and Matlab/Simulink. All analyses for SRM and PM-SRM have been carried out under the following same conditions: The supply voltage has been fixed to 300 Vdc. By referring Fig. 2, unaligned and aligned positions of rotor according to the phase A is 0\u25e6 and 45\u25e6, respectively. The phase torque is positive between 0\u201345\u25e6 whereas it is negative between 45\u201390\u25e6, when the motor rotates in clockwise direction. The turn-on and turn-off angles of the phase A transistors have been respectively set to 5\u25e6 and 41\u25e6 to produce positive torque. The required turn-on and turn-off angles for other phase switches have been obtained by considering Phase B has \u221230\u25e6 and phase C has +30\u25e6 phase shift with phase A" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0001072_sceecs48394.2020.171-Figure2-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0001072_sceecs48394.2020.171-Figure2-1.png", + "caption": "Figure 2: Structure of Gantry robot (without gripper)", + "texts": [], + "surrounding_texts": [ + "Articulated and SCARA robots are well recognized in the field of industrial robotics because of combined linear and rotary motion resulting in ease of maneuverability for complex tasks. But in industries where high work load is handled and requires high stability, Gantry robot is preferred. This may be understood by the following table [1]. In industries, the already existing gantry robots have either their base axis (X) fixed or they are moving on the overhead railings which limits the workspace of the robot as the location to pick and place an object is fixed. This project aims to make a movable gantry robot which can move on the floor and can access the part available at the floor space. By this, it can pick the object from any location and can place it on any other location, thereby having a greater workspace. Moreover, any sudden changes in picking and placing position can be encountered easily. Motor driven wheels are mounted on the base of the robot for its locomotion. II. CONCEPTUAL DESIGN As we discuss, the Gantry robot consist of minimum three motions. Each is in linear motion. This linear motion are perpendicular to each other which moves in X, Y and Zdirection. This X & Y-direction are on horizontal plane whereas the Z-direction is on vertical plane. The volume created by the X, Y & Z-axis in which the end effector of robot will move is known as the working envelope. 20 20 IE EE In te rn at io na l S tu de nt s' C on fe re nc e on E le ct ric al ,E le ct ro ni cs a nd C om pu te r S ci en ce (S C EE C S) 9 78 -1 -7 28 1- 48 62 -5 /2 0/ $3 Authorized licensed use limited to: Auckland University of Technology. Downloaded on June 02,2020 at 14:03:08 UTC from IEEE Xplore. Restrictions apply. SCEECS 2020 In this project, the motion of X, Y & Z-direction motion is achieve by the lead screw (Threaded rod) and coupler (nut) arrangement. The one end of the lead screw will be attached to the motor. So that the rotation motion will take place in the threaded rod and we kept the nut fixed. As the rod rotate the nut will move forward or backward according to the direction of rotation of the rod (i.e. clockwise or anticlockwise). Same technique is used for z-direction but in vertical plane. The motion of the nut will be upward and downward according to the rotation of rod. Figure 1: Design of Gantry Robot from top view This is the design of the structure from top side where all the motion of 3-axis are being shown. A coupling is used to connect motor shaft of 4mm to the 8mm of the lead screw. There is a fourth degree of freedom which rotates in the X-Y plane. To make the end effector more flexible according to the application. In this project we are attaching a fixed arm at the end-effector of Gantry robot. Authorized licensed use limited to: Auckland University of Technology. Downloaded on June 02,2020 at 14:03:08 UTC from IEEE Xplore. Restrictions apply. SCEECS 2020 III. CONTROLLING OF GANTRY ROBOT We are using the NodeMCU as the micro-controller. The NodeMCU has the functionality in which he can act as a server. NodeMCU has the station(STA)mode in which it can connect with the existing wifi-network and act as a HTTP server with the IP address assign to that network.[5] It act as a HTTP server only for the device which are connect to the same wifi network A. Interface of client with micro-controller. We are using the NodeMCU as the micro-controller. The NodeMCU has the functionality in which he can act as a server. NodeMCU has the station(STA)mode in which it can connect with the existing WiFi-network and act as a HTTP server with the IP address assign to that network.[5] It act as a HTTP server only for the device which are connect to the same WiFi network. HTTP stands for the Hypertext Transfer Protocol. It is standard application protocol which functions as requestresponse protocol between client and server. HTTP client help to send and receive the HTTP request and HTTP response respectively from the HTTP Server. Web-server of the NodeMCU will open when we put the IP address of the network in the address bar. Then the HTTP server page is open where we can send and receive the request of the client. The user is going to control the motion of the stepper, DC and servo motor using this web-page which is created on the HTTP server. BY clicking on the buttons which is created using the HTML. B. Interface between micro-controller and the actuators Now after receiving the data from the client. Microcontroller will instruct accordingly to the actuators. 1) Interfacing of Stepper motor with Node-MCU The amount of power required by the Stepper motor to work is 12V-1.5ampere which is provided by external power source. We also need a circuit which can control the rotation of motor perfectly and easily. Hence, we need an Hbridge circuit to control the stepper motor with Node-MCU i.e. micro-controller, H-bridge circuit is the combination of transistor or mosfets. So, here we are using the A4988 circuit. Using this IC we can control the motor by just 2-pins. One for the direction of the motor and other pin for the steps. This driver provide the full-step, haft-step, quarter-step, eight-step and sixteenth-step resolutions. This stepper motor is connect with a lead screw to move the complete articulated arm in X, Y and Z-coordinates. 2) DC Motor The DC motor is used to move the complete Gantry Robot from one location to another. So, to control the dc motor through NodeMcu we are using the driver IC- L298N. Authorized licensed use limited to: Auckland University of Technology. Downloaded on June 02,2020 at 14:03:08 UTC from IEEE Xplore. Restrictions apply. SCEECS 2020 As similar to stepper motor the DC-motor require the amount of power is 12V,1 or 2 ampere which is provided by external power source. Here, also we need H-bridge circuit to control and L298N is a dual channel H-bridge motor driver IC which can able to drive a pair of DC motor. In L298N driver IC there is a voltage drop of nearly 2V due to the switching transistor in H-bridge circuit. The servo motor is used to move the articulated arm which is attached to the end-effector of a Gantry Robot. The servo motor is controlled by the node-MCU pin. The motor is also connect to the 2 Jaw-gripper which is the end-effector of the articulated arm. The 2-jaw gripper is used to pick up the object and placed it at the required position. The gripper can be classified as coarse grippers or precise grippers. The coarse gripper is found where the components can be pick or placed with relatively little accuracy, whereas, precise gripper are used where the accuracy is essential or should have accurate holding position and orientation. 4) Gripper : Gripper is mounted as the end effector of the robot to hold and release the object. There are various factors while selecting a gripper which are as follows \u2022 Object to be handled \u2022 Sequence of operation \u2022 Mechanism of robot \u2022 Environment of workspace The first and foremost consideration while selecting a gripper is the object which is to be handled. There are following parameters related to object which influence the selection of gripper \u2022 Size of object \u2022 Number of potential gripping surface \u2022 Geometry of gripping surface \u2022 Distance of gripping surface Considering the above points, a two jaw V-Shaped gripper is chosen for the application. The gripper is actuated by the servomotor and a pairs of pinion. When the servomotor is actuated it drives the pinion in the same direction of rotation, whereas the gear in mesh with pinion is driven in opposite direction making the jaw of the gripper to come close and hold the object. To release the grip, the motor is rotated in opposite direction. IV. RESULT AND DISCUSSION A mobile gantry robot for pick and place application will be much more suitable for the industries having application for storing and dispatching of the goods in warehouses as well as in industries having assembly lines operation where an object has to picked up from a particular assembly line and is need to be placed at some other location. By using the controller, one can easily control the movement of the robot and will be much easier and flexible than any other robot which could be used for the same application. This robot will be more stable while picking and placing the objects because of its sturdy and symmetrical structure, thus enabling it to lift heavy objects as well. V. CONCLUSION \u201cMobile Gantry Robot for Pick and Place Application\u201d is an industry oriented project aiming specifically on pick and place application in the industries which came out as a solution looking at the problem in other pick and place robots as well as conventional robots. This project will ease out the various operation of the industries as far as the pick and place application is concerned. Authorized licensed use limited to: Auckland University of Technology. Downloaded on June 02,2020 at 14:03:08 UTC from IEEE Xplore. Restrictions apply. SCEECS 2020 VI. REFERENCES [1] https://www.linearmotiontips.com/when-do-you-need-a-gantry- robot/#ampshare=https://www.linearmotiontips.com/when-do-you-needa-gantry-robot/ [2] http://www.sagerobot.com/gantry-robots/ [3] Surinder Pal,\u201dDesign and remote control of a Gantry Mechanism\u201d, Office of graduate studies Texas A & M University, August 2017. [4] Mesfin B., Muluken T., Samson D., Abayneh T.,\u201d Design of Gantry Robot\u201d, Bahidar University of Techonolgy,2014. [5] https://www.electronicwings.com/nodemcu/http-server-on-nodemcuwith-arduino-ide [6] https://www.engineersgarage.com/esp8266/nodemcu-esp8266-steppermotor-interfacing/ [7] http://www.ece.montana.edu/seniordesign/archive/FL13/GantryRobot/in dex.html [8] https://www.macrondynamics.com/job-stories/gantry-systems-overview [9] https://www.machinedesign.com/motion-control/five-heavy-dutygantry-alternatives [10] https://howtomechatronics.com/tutorials/arduino/how-to-controlstepper-motor-with-a4988-driver-and-arduino/ Authorized licensed use limited to: Auckland University of Technology. Downloaded on June 02,2020 at 14:03:08 UTC from IEEE Xplore. Restrictions apply." + ] + }, + { + "image_filename": "designv11_71_0002989_978-3-642-16214-5_9-Figure9.32-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0002989_978-3-642-16214-5_9-Figure9.32-1.png", + "caption": "Fig. 9.32. Tribological system of a multi-plate clutch: lined plate (shown with internal gearing), oil film and counter-running surface (steel plate, shown with external gearing)", + "texts": [ + " Multi-plate clutches will be understood in this section as clutches active in the shifting process, flown through with oil and activated by pressure oil. The multiplate brake is a special form of the multi-plate clutch. The following pages will first formulate the requirements of multi-plate clutches and then describe the basics of the shifting process, these considerations providing the basis for presenting the main features of design. A further section is concerned with the tribological system composed of the lined plate, oil film and steel plates as counter-running surfaces (Figure 9.32). Since the torque is transmitted through friction, the friction coefficient \u03bc between the friction surfaces has a large influence on the system\u2019s behaviour. Special oils, ATF oils (Automatic Transmission Fluid), were developed for use in automatic transmission with various gear ratios. The behaviour of friction and thus the transmission of torque can be influenced decisively with the type of friction lining and oil used. The section \u201cLayout and Design of Multi-Plate Clutches\u201d is concluded with design recommendations and information on detailed questions" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0001031_icimia48430.2020.9074843-Figure8-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0001031_icimia48430.2020.9074843-Figure8-1.png", + "caption": "Fig 8. Rotary Linear Induction Motor: 1-horizontal conductors, 2- vertical conductors, 3- housing, 4-rotor, 5- magnetic circuit [12]", + "texts": [ + " Three-phase RLIM requires two different types of windings, one is distributed three-phase winding and the second is a three-phase ring winding. Three phases distributed winding produces a rotating magnetic field, responsible for rotary motion. Three-phase ring winding generates traveling magnetic field, causes for linear motion. There are two possible designs for RLIM based on stator [12] is discussed as follows: Design 1: Design 1 has two types of windings which are placed on the stator. Horizontal conductors of a distributed winding and vertical conductors of a ring winding as shown in Fig.8. Both ring and distributed windings are placed on the common stator which has the same magnetic circuit. Both ring and distributed windings are perpendicular to each other. When the stator is energized, these two windings generate both rotating and traveling fields along the axes. A resultant two-dimensional magnetic field is generated across the air gap between the stator and rotor. The tubular rotor is designed for interacting with the resultant magnetic field. The interaction of two fields produces both rotational torque and linear force simultaneously" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0003921_coase.2014.6899477-Figure9-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0003921_coase.2014.6899477-Figure9-1.png", + "caption": "Figure 9. Setting of the detection contact force level. (a) Large (b)Small", + "texts": [ + " (a)Robot (b)Velocity and Contact Force-based Safety Device (c) Details Contact Force Detection Plate Switch-off and Shaft-lock Mechanism Shaft A Velocity-based Detection Mechanism Linear Spring Y Contact Force-based Detection Mechanism Wire Rope Pulley Linear Spring Z Claw D Wheel Caster Claw D Wire Rope To Contact Force Detection Plate Shaft A Switch-off and Shaft-lock Mechanism Velocity-based Detection Mechanism Motor Wheel Linear Spring Z Velocity and Contact Force-based Safety Device Timing Belt (d)Three Claws B. shown in Fig.9. If the acceleration of Contact Force Detection Plate is so small that we can neglect the inertial force, we obtain the following motion equation: fF k y\u2206= , (1) where F is the detection contact force level, kf is the spring constant of Linear Springs Y, y\u2206 is the displacement from the natural length of Linear Springs Y. We can approximately set the detection contact force level on the basis of (1). 2) Velocity-based Detection Mechanism Fig. 10 shows the mechanism which mechanically detects the unexpected angular velocity of Shaft A" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0002189_icem49940.2020.9271053-Figure10-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0002189_icem49940.2020.9271053-Figure10-1.png", + "caption": "Fig. 10. Efficiency map (a) proposed method (b) transient analysis (c) difference between proposed method and transient analysis.", + "texts": [ + " The AC copper loss according to torque and speed are shown in Figs. 9. Both of copper loss by proposed method and transient analysis are similar for all the range of loads. Moreover, the efficiency map obtained using the proposed method is compared that obtained using the transient analysis. Figs. 10 (a) and (b) show the efficiency map of the proposed method and the transient analysis, respectively. The two efficiency maps look almost the same to each other. To confirm this, the difference between the two efficiency maps is shown in Fig. 10 (c). The maximum efficiency difference between the two methods is 0.55%p, which is not a large difference. Also, an average of absolute values of the efficiency difference is 0.08%p, which means that efficiency obtained using the two methods is not significantly different in most regions. As shown in Figs. 10 (a), (b) and (c), it can be concluded that the performance of the motor considering AC resistance using the proposed method can be enough well predicted compared that using the transient analysis", + " From aforementioned advantage, the proposed method can be applied when designing not only the traction motor for EV but also the synchronous motor which have high AC copper loss. To verify that the proposed method is cost-efficient, the computation time of the proposed method using magnetostatic analysis will be compared with that of the transient analysis. When comparing each other, the same solver will be used for both of analysis. Moreover, the efficiency results by the load test of the prototype will be contained to find out the results of Fig. 10 are valid. [1] M. Lim, S. Chai, J. Yang and J. Hong, \"Design and Verification of 150-krpm PMSM Based on Experiment Results of Prototype,\" in IEEE Transactions on Industrial Electronics, vol. 62, no. 12, pp. 7827-7836, Dec. 2015. [2] G. Pellegrino, A. Vagati, B. Boazzo and P. Guglielmi, \"Comparison of Induction and PM Synchronous Motor Drives for EV Application Including Design Examples,\" in IEEE Transactions on Industry Applications, vol. 48, no. 6, pp. 2322-2332, Nov.-Dec. 2012. [3] J. de Santiago et al" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0001334_s40435-020-00661-8-Figure2-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0001334_s40435-020-00661-8-Figure2-1.png", + "caption": "Fig. 2 Rotor forces and moments in non-rotating hub-fixed coordinate system", + "texts": [ + " However, in stiff rotors, such as those on a typical quadrotor, however, these moments are transferred into the fuselage, and thus, they have to be calculated in this work. Each rotor has a rotational speed of where the direction of rotors are either clockwise or counterclockwise. The azimuthal angle of a rotor blade ( b ) is set to zero at the aft of the rotor disc, and increases in the direction of rotation. The four rotors have the same manner in their pitch and blade planform, except that two rotors are designed to rotate clockwise and the other two are counter-clockwise. Based on Fig.\u00a02, using r = rcg + ri , the hub velocity in the body-fixed coordinate system (\u2026)B is obtained as: (1) ( Pr I )B = ( P rcg I )B + ( P ri B )B + B IB \u00d7 rB i 1 3 where Pr I denotes the time derivative of the position vector r when the observer is located in inertial coordinate system. Also, IB present the angular velocity of body-fixed frame with respect to inertial coordinate system. Using ( P rcg I )B = (u, v,w) , ( P ri B )B = 0 , B IB = (p, q, r) , a n d rB i = ( xi, yi, zi ) , the hub speed components in body-fixed frame are corresponded to: (2) \ufffd Pr I \ufffdB = vB hub = \u23a1\u23a2\u23a2\u23a3 u \u2212 ryi + qzi v + rxi \u2212 pzi w \u2212 qxi + pyi \u23a4\u23a5\u23a5\u23a6 where ri is the coordinate of the ith rotor in body-fixed coordinate system. In this paper, the direction of body-fixed coordinate system is the same as non-rotating hub-fixed coordinate system, and thus, vB hub = vH hub : For the blade element position of r\ufffd = r + rb , the velocity components of the blade element in the direction shown in Fig.\u00a02 is obtained as: (3)vH hub = \u23a1\u23a2\u23a2\u23a3 vxH vyH vzH \u23a4\u23a5\u23a5\u23a6 = \u23a1\u23a2\u23a2\u23a3 u \u2212 ryi + qzi v + rxi \u2212 pzi w \u2212 qxi + pyi \u23a4\u23a5\u23a5\u23a6 1 3 (4) uT = rb + vxH sin b + (\u22121)mvyH cos b uP = R + vzH + (\u22121)mrbp sin b \u2212 rbq cos b where rotor inflow ratio is a function of blade azimuth ( b ) and radial distance ( rb ) . Also, m = 0 denotes counterclockwise rotor (rotor 1 and 3) and m = 1 is used for clockwise rotors (rotor 2 and 4). Based on Fig.\u00a02, the blade angle of attack ( b ) , the inflow angle ( ) , the elemental lift and drag is therefore corresponded to: In this work, the airfoil drag coefficient is Cd = Cd0 + Cd1 b + Cd2 2 b to ensure that the realistic flight condition for quadrotors, a is the blade lift curve slope, and c is blade chord. According to Fig.\u00a02, the forces normal and tangential to the rotor disc are written as: Using Eq.\u00a0(8) and Fig.\u00a02, the elemental forces and moments in the non-rotating hub plane (\u2026)H are obtained as: and Overall, the average rotor forces and moments are obtained by b 2 2 \u222b 0 BR \u222b 0 (\u2026)drbd b and therefore: (5) b = \u2212 (6) = tan\u22121 uP\u2215uT (7) dL = 1 2 ac ( u2 T + u2 P ) drb dD = 1 2 cCd ( u2 T + u2 P ) drb (8) dFz = dL cos + dD sin dFx = dL sin \u2212 dD cos (9) \u23a1\u23a2\u23a2\u23a3 dXM dYM dZM \u23a4\u23a5\u23a5\u23a6 = \u23a1\u23a2\u23a2\u23a3 dFx sin b (\u22121)mdFx cos b \u2212dFz \u23a4\u23a5\u23a5\u23a6 = \u23a1\u23a2\u23a2\u23a3 \u2212dH dY \u2212dT \u23a4\u23a5\u23a5\u23a6 (10) \u23a1\u23a2\u23a2\u23a3 dMx dMy dMz \u23a4\u23a5\u23a5\u23a6 = \u23a1\u23a2\u23a2\u23a3 \u2212(\u22121)mdFzrb sin b dFzrb cos b (\u22121)mdFxrb \u23a4\u23a5\u23a5\u23a6 1 3 The rotor solidity is = bc\u2215 R where b is the number of blade" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0002811_978-94-007-6046-2_71-Figure3-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0002811_978-94-007-6046-2_71-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 R 0", + "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 R b attached to this body, with respect to a given reference inertial frame R 0, 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 R b with respect to R 0, 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_71_0002561_picc51425.2020.9362467-Figure1-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0002561_picc51425.2020.9362467-Figure1-1.png", + "caption": "Fig. 1. Schematic Diagram of FNC System", + "texts": [ + " This paper is organized as follows. In section II, the System Configuration is presented. The analysis of the linear model of the system is done in section III and the compensation scheme is detailed in section IV. Section V deals with the frictional modeling of the system. The estimator design is given in Section VI and the friction controller design is given in Section VII. The simulation results are presented in Section VIII and concluding remarks are given in Section IX. A typical FNC system, shown in Fig. 1 makes use of a flexible joint configuration with a flex seal and a pair of orthogonally placed actuators which control the deflection of the nozzle in the pitch and yaw directions. The actuator comprises a BLDC motor and a ball screw mechanism. The system is usually connected to the launch vehicle and mounted on the nozzle [1]. During the flight, the autopilot commands the flight control system and actuators to elongate or retract in order to achieve the required nozzle deflection and by extension, the desired engine rotation [2]" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0000896_0954406220912005-Figure5-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0000896_0954406220912005-Figure5-1.png", + "caption": "Figure 5. Kinematic schemes of the ith leg of the: (a) RSRP; (b) RSPR; (c) RSPP; (d) RSG; (e) RS[P]R; and (f) RS[P]P wheels.", + "texts": [ + " On the other hand, by taking the derivative of (7) the following expression can be obtained _ti \u00bcMm\u00feN _q \u00f014\u00de where M \u00bc lTs\u00f0 2 \u00fe #\u00de L2s 2 L3s 3 lTc\u00f0 2i \u00fe #\u00de \u00fe L2c 2 L3c 3 \u00f015\u00de and N \u00bc Rc1s i 0 Rc1c i 0 \u00f016\u00de Finally, the kinematic model of the wheel that relates the velocities of the actuated joints and the operational velocities is as follows _ti \u00bc J _q \u00f017\u00de where the Jacobian matrix of the wheel J can be computed in the following way J \u00bcMK 1L\u00feN \u00f018\u00de Jacobian matrices can be used to evaluate configurations by means of kinematic performance indices and to determine the singular configurations of the mechanism.25 In the rest of the mechanism it is possible to obtain the equations of velocity (11), (14), and (17), differing only in the elements of matrices K, L, M, and N. Figure 5(a) shows the kinematic scheme of the RSRP wheel. In this case, the proximal link and the rim are connected by the prismatic joint P. To make the traction link move in a plane orthogonal to the axis of the wheel it is necessary the axis of motion of joint P be perpendicular to that of Rr, and at the same time the axis of R2 and Rr be parallel between them. For this mechanism, the unknown variables of equation (4) are the extension of the prismatic joint L3 and the angle of the traction link 2. The solution of this equation is given by the following expressions L3 \u00bc P ffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi P2 \u00fe L2 2 Q2 1 Q2 2 q \u00f09\u00de where P \u00bc Q1 cos\u00f0 i \u00fe \u00de \u00feQ2 sin\u00f0 i \u00fe \u00de. The angle of the axis of the prismatic joint of the proximal link is given as shown in Figure 5(a). The orientation of the traction link is given by the following expression 2 \u00bc atan2\u00bdQ2 \u00fe L3 sin\u00f0 i \u00fe \u00de,Q1 \u00fe L3 cos\u00f0 i \u00fe \u00de \u00f020\u00de where atan2 is the four-quadrant inverse tangent. For the velocity analysis, the vector of passive joint velocities is defined by m \u00bc !2 _L3 T . Jacobian matrices K and L are given by K \u00bc L2s 2 c\u00f0 i \u00fe \u00de L2c 2 s\u00f0 i \u00fe \u00de \u00f021\u00de and L \u00bc Q2 L3s\u00f0 i \u00fe \u00de c i=t 1 Q1 \u00fe L3c\u00f0 i \u00fe \u00de s i=t 1 \u00f022\u00de Jacobian matrices that relate the joint velocities and the velocity of the end of the traction link are M \u00bc lTs\u00f0 2 \u00fe #\u00de L2s 2 c\u00f0 i \u00fe \u00de lTc\u00f0 2i \u00fe #\u00de \u00fe L2c 2 s\u00f0 i \u00fe \u00de \u00f023\u00de and N \u00bc Rc1s i \u00fe L3s\u00f0 i \u00fe \u00de 0 Rc1c i L3c\u00f0 i \u00fe \u00de 0 \u00f024\u00de The kinematic scheme of the RSPR wheel is presented in Figure 5(b). In this case, the traction link and the proximal link are connected by a prismatic joint. The conditions to make the traction link move in a plane orthogonal to rotation axis of the wheel are the same as those of the RSRP wheel. For this kinematic chain the unknowns of the vector loop equation (4) are the extension of the prismatic joint L2 and the angle of the proximal link 3. The solution of this equation is given by the following expressions 3 \u00bc 2 arctan Q1 ffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi Q2 1 \u00feQ2 2 L2 3sin 2 q L3 sin Q2 0@ 1A \u00f025\u00de and L2 \u00bc L3 cos 3 Q1 cos\u00f0 3 \u00fe \u00de \u00f026\u00de In this case, the vector of passive joint velocities is defined as m \u00bc _L2 ", + " Jacobian matrices K and L are given by K \u00bc c\u00f0 i \u00fe \u00fe \u00de c\u00f0 i \u00fe \u00de s\u00f0 i \u00fe \u00fe \u00de s\u00f0 i \u00fe \u00de \u00f033\u00de and L \u00bc Q2 \u00fe L2s\u00f0 i \u00fe \u00fe \u00de L3s\u00f0 i \u00fe \u00de c i=t 1 Q1 L2c\u00f0 i \u00fe \u00fe \u00de \u00fe L3c\u00f0 i \u00fe \u00de s i=t 1 \u00f034\u00de Jacobian matrices that relate joint velocities and the velocity vector of the traction link are M \u00bc K \u00f035\u00de and N \u00bc N1 0 N2 0 \u00f036\u00de where N1 \u00bc lTs\u00f0 i \u00fe \u00fe \u00fe #\u00de L2s\u00f0 i \u00fe \u00fe \u00de\u00fe L3s\u00f0 i \u00fe \u00de Rc1s i and N2 \u00bc lTc\u00f0 i \u00fe \u00fe \u00fe #\u00de\u00fe L2c\u00f0 i \u00fe \u00fe \u00de L3c\u00f0 i \u00fe \u00de \u00fe Rc1c i. The rest of the mechanisms have similar solutions to those of the previously presented mechanism. In the case of the RSG wheel (see Figure 5(d)), the solution of the passive joint variables is given by 3 \u00bc 2 arctan Q1 ffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi Q2 1 \u00feQ2 2 L2 2sin 2 q L2 sin Q2 0@ 1A \u00f037\u00de and L3 \u00bc Q1 \u00fe L2 cos\u00f0 3 \u00fe \u00de cos 3 \u00f038\u00de It can be observed that this pair of equations are the same that (25) and (26) of the mechanism RPPR when \u00bc 0. On the other hand, the position of the end of the traction link is given by ti \u00bc \u00f0lR=L3\u00deR \u00fe I2 1 ci\u00fe 1r1i \u00f039\u00de where lR is the length of the line from the intersection of bi and ci to the end of the traction link, and g is the angle between such line and vector ci as can be seen in Figure 5(d). Jacobian matrices L and N are the same as those of mechanism RSPR, and matrices K and M are given by K \u00bc L2s\u00f0 3 \u00fe \u00de L3s 3 c 3 L2c\u00f0 3 \u00fe \u00de \u00fe L3c 3 s 3 \u00f040\u00de and M \u00bc lRs\u00f0 3 \u00fe \u00de L3s 3 c 3 lRc\u00f0 3 \u00fe \u00de \u00fe L3c 3 s 3 \u00f041\u00de The kinematic scheme of mechanism RS[P]R (RUGR) is presented in Figure 5(e). In this mechanism, the relationship between joint variables and matrices K, L, and N is identical to the case of the RSPR wheel. The position of the end of the traction link is given by equation (39), and matrix M is given by the following expression M \u00bc 0 lRs\u00f0 3 \u00fe \u00de L3s 3 0 lRc\u00f0 3 \u00fe \u00de \u00fe L3c 3 \u00f042\u00de In the case of the RS[P]R (RUGP) wheel, the expressions that relates the passive and actuated joints are given by equations (31) and (32) of the RSPP wheel. Jacobian matrices K and L are given by equations (33) and (34), respectively" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0002226_icem49940.2020.9270688-Figure1-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0002226_icem49940.2020.9270688-Figure1-1.png", + "caption": "Fig. 1: Four quadrant feeding in redundancy operation. One (remaining) inverter feeds two opposing quadrants of the stator winding. The other two quadrants are idling since the faulty inverter is disconnected.", + "texts": [ + " above, we get a fault tolerance against a single inverter fault, which is a relatively frequent failure [12]. This fault tolerance is achieved by disconnecting the faulty inverter from the machine and continuing operation only with the remaining inverter. Here, two 3-phase inverters feeding two parallel branches of the stator winding are assumed. In case of a single inverter fault, redundancy operation with feeding of a single parallel branch results in a sector-wise feeding of the stator winding as shown in Fig. 1. In this redundancy operation additional subharmonics in the magnetic air gap field of the stator winding will occur [13] and cause significant eddy currents in the solid rotor yoke [14]. In this paper, the design of a wind generator with tooth coil winding is treated with regard to a redundant stator feeding by two voltage link inverters and assuming sinusoidal stator currents. First of all, suitable tooth coil windings are identified concerning the rotor eddy current losses in the solid rotor yoke", + " Three concentrated windings with m = 3, which provide completely fed fundamental winding schemes in redundancy operation, and one concentrated winding with m = 6, which will not maintain the fundamental winding scheme in redundancy operation. A single layer, integer slot distributed winding with m = 3, q = 1 is added as a benchmark. The pole counts of the candidates are chosen in such a way that comparable slot pitches are reached, i.e. candidates with concentrated windings have a higher pole count than the candidate with the distributed winding. Moreover, the chosen pole counts ensure that the redundancy mode consists of two fed quadrants opposite to each other for the 3-phase variants (Fig. 1). This pattern is preferred, because radial and tangential stresses cancel out, resulting in zero integral translational force. Force compensation may be reached by any point symmetric pattern, but this pattern is also loss-optimal concerning the additional rotor yoke eddy current losses arising from sectorial feeding as shown in [14]. The candidates are assessed regarding the analytically calculated eddy current losses in the rotor yoke in normal and redundancy operation. Eddy current losses in the permanent magnets are relatively small due to segmentation (120 segments per pole) and therefore neglected in the analytical calculation", + " Therefore, higher order space harmonics are negligible for the total rotor yoke eddy current losses and not visible in the accumulated loss graphs in Fig. 3 - 7. The spectral distribution of rotor yoke eddy current losses under circumferentially intermittent feeding of the stator winding in redundancy operation for C2 is shown in Fig. 4. Due to the sectorial feeding of the stator winding many additional harmonics occur, especially subharmonics with long wavelengths. The largest wavelength is half of the machine circumference and covers one fed and one idling quadrant (Fig. 1). The calculation of the winding factors is treated in [13] and based on [23]. The distribution of winding factors is shown in Fig. 4 and differs from the completely fed machine (Fig. 3) in the occurrence of side bands around the usual harmonics. This effect is explained in [13] as the result of a convolution of the original winding spectrum and the spectrum of a masking function. Since only two of four quadrants are fed, the machine delivers only half of the nominal torque and the amplitude of the stator field is halved regarding the operating wave as well as all harmonics" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0003880_scis-isis.2014.7044812-Figure1-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0003880_scis-isis.2014.7044812-Figure1-1.png", + "caption": "Fig. 1. Overview representation of potential forces", + "texts": [ + " The potential field can be treated as a landscape with several mountains generated by the obstacles and valleys where the lowest valley point represents the goal point. The APF in an environment is composed of two fields, attractive field generated by the goal and repulsive field by the each obstacle. When the robot immersed in the potential field, attractive force and repulsive force acts on the robot, guides the robot towards the goal point. This combination of two forces is dedicated to guide the robot in a safe path while keeping it away from the obstacles. Fig. 1 represents the attractive potential and repulsive potential force distribution on the robot for a single obstacle existing in the environment. Attractive force that is produced by the goal can be expressed using the Gaussian function for APF as in (1). (1) Where, is the maximum value of the attractive force at any instance and is a constant that defines the width of the distribution. The parameter is defined as Euclidean distance between the robot and the goal. The repulsive potential force can be described as in (2)" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0000294_icems.2019.8921576-Figure4-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0000294_icems.2019.8921576-Figure4-1.png", + "caption": "Fig. 4. Temperature distribution of the motor with rated operation: (a) 3D temperature contour, (b) temperature distribution on the central section.", + "texts": [ + " Considering the convective heat dissipation at the end windings, a three-dimensional full motor temperature field model including the end fluid domain is established. The Authorized licensed use limited to: Middlesex University. Downloaded on July 18,2020 at 14:52:08 UTC from IEEE Xplore. Restrictions apply. temperature of the inlet water is 15 \u2103 and the flow rate is 20 L/min. The gaps between the joint surfaces of the stator yoke to the housing is 0.03mm and that of the slot insulation to the stator core is 0.1mm respectively [14-15]. The temperature distribution of the motor under the normal operation is shown in Fig. 4. The maximum temperature appears in the windings, and about 118 \u2103. The insulation grade of the windings is H whose limited operating temperature is 180 \u2103. Considering the margin, the 170 \u2103 is set to the limited temperature of the motor in the following analysis. When the MMF compensation is only adopted in M1, the maximum temperature appears in the windings of M1 and the temperature increases rapidly. The winding temperature reaches the rated temperature in about 6 minutes and gets the limited temperature in about 11 minutes, as shown in Fig" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0000603_iccas47443.2019.8971558-Figure8-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0000603_iccas47443.2019.8971558-Figure8-1.png", + "caption": "Figure 8. End-section kinetic structure", + "texts": [], + "surrounding_texts": [ + "221\nThe first segment is located by the distal part and the second segment is located by the proximal part. The first segment is operated by 2 pairs of 2 wires; total 4 wires. One pair of 2 wires is controlled by one motor that pulls one wire and push the other wire in the same length by using the pulley. While, the second segment is operated by 4 pair of 4 wires, total 8 wires. Therefore the second segment is controlled by 4 motors The second segment has m units and the first segment has n-m units. 4 pairs of wires are labeled by \ud835\udc4e\ud835\udc4e and ?\u0302?\ud835\udc4e, \ud835\udc4f\ud835\udc4f and ?\u0302?\ud835\udc4f, \ud835\udc50\ud835\udc50 and ?\u0302?\ud835\udc50, d and ?\u0302?\ud835\udc51,\nEquilibrium in moments at Un belonging to the first segment is (Sa,n-fa)(an-an-1)\u00d7(an-un)+(Sa\u0302,n-fa\u0302)(a\u0302n-a\u0302n-1)\u00d7(a\u0302n-un)+\n(Sc,n-fc)(cn-cn-1)\u00d7(cn-un )+(Sc\u0302,n-fc\u0302)(c\u0302n-c\u0302n-1)\u00d7(c\u0302n-un)++\nmw(pn-un)\u00d7g= ( 0 0 0 )\n( (9)\nwhere, an-an-1= an-an-1\n|an-an-1| , etc. m\ud835\udc64\ud835\udc64 is a payload applying at the\nend-point and g is the gravity acceleration vector. Equilibrium in moments at Ui , (i=m+1, \u22ef, n-1), belonging to the first segment is (Sa,n-fa)(an-an-1)\u00d7(an-un)+\n(Sa\u0302,n-fa\u0302)(a\u0302n-a\u0302n-1)\u00d7(a\u0302n-un)+ (Sc,n-fc)(cn-cn-1)\u00d7(cn-un )+ (Sc\u0302,n-fc\u0302)(c\u0302n-c\u0302n-1)\u00d7(c\u0302n-un)+\nmp \u2211 (pk-ui) n\nk=i+1\n\u00d7g= ( 0 0 0 )\n(10)\n(10)\nwhere. fa, fa\u0302, fc, fc\u0302 are wire tensions, Sa,i, Sa\u0302,i, Sc,i, Sc\u0302,i , (i=m+1, \u22ef, n) are spring tensions of the ith unit. \u201c\u00d7\u201d means a cross product and \u201c|*|\u201d, means the modulus of a vector \u2217. \ud835\udc5a\ud835\udc5a\ud835\udc5d\ud835\udc5d is the mass of one unit including the plate, the rod and the universal joint.\nThe spring tensions are obtained as, Sa,i\uff1dk(L-|ai-ai-1|), Sa\u0302,i\uff1dk(L-|a\u0302i-a\u0302i-1|),\nSc,i=k(L-|ci-ci-1|), Sc\u0302,i=k(L-|c\u0302i-c\u0302i-1|),)\n(11)\nwith spring coefficient k. Equations (9) and (10) contain 3(nm) equations including 4(n-m) -1variables of the n-m universal joints angles \u03b8xi, \u03b8yi, \u03b8zi, (i=m+1, \u22ef, n) and slide length of plates \ud835\udc59\ud835\udc59\ud835\udc56\ud835\udc56 (i=m+1, \u22ef, n-1). Equilibrium in force at ith plate (i=m+1, \u22ef, n-1) is,\n[-Sa,i+1(ai+1-ai) +Sa,i(ai-ai-1)+-Sa\u0302,i+1(a\u0302i+1-a\u0302i)+ Sa\u0302,i(a\u0302i-a\u0302i-1)-Sc,i+1(ci+1-ci)+Sc,i(ci-ci-1)-Sc\u0302,i+1(c\u0302i+1-c\u0302i))\nSc\u0302,i(c\u0302i-c\u0302i-1)+(n-i)mpg ]\u2219 (pi-ui)=0\n(12)\n(12) provide n-m-1equations. Combined it with (9) and (10), we obtain 4(n-m)-1 equations, which suffices in number to solve for 4(n-m)-1 variables; \u03b8x,i, \u03b8y,i , \u03b8z,i (i=m+1,\u22ef,n) and li (i=m+1, \u22ef, n-1) for a given set of wire tensions fa, fa\u0302, fc, fc\u0302.\nEquilibrium in moments at Um , the universal joint located at the most distal position belonging to the second segment is\n-Sa, m+1(am+1-am)\u00d7(am-um)+(Sb,m-fb)(bm-bm-1)\u00d7(bm-um)\n-Sa\u0302,m+1(a\u0302m+1-a\u0302m)\u00d7(a\u0302m-um)+(Sb\u0302,m-fb\u0302) (b\u0302m-b\u0302m-1) \u00d7(b\u0302m-um) -Sc, m+1(cm+1-cm)\u00d7(cm-um)+(Sd,m-fd)(dm-dm-1)\u00d7(dm-um)\n-Sc\u0302,m+1(c\u0302m+1-c\u0302m)\u00d7(c\u0302m-um)+(Sd\u0302,m-fd\u0302) (d\u0302m-d\u0302m-1) \u00d7(d\u0302m-um)\n+ (mw(pn-um)+mp \u2211 (pk-um) n-1\nk=m+1\n) \u00d7g= ( 0 0 0 )\n(13)\nFor the second segment, we can derive similar equations as (10), (11) and (12) by replacing {ai, a\u0302i, ci, c\u0302i} with {bi, b\u0302i, di, d\u0302i}, {Sa,i, Sa\u0302,i, Sc,i, Sc\u0302,i} with {Sb,i, Sb\u0302,i, Sd,i, Sd\u0302,i} for i=1, \u22ef, m-1 in (10) and for i=1, \u22ef, m in (11) and (12).\nAs a result, we obtain 4m equations included by (13), which suffices in number to solve for 4m variables; \u03b8x,i, \u03b8y,i, \u03b8z,i and li (i=1, \u22ef,m) for a given set of wire tensions \ud835\udc53\ud835\udc53\ud835\udc4f\ud835\udc4f, \ud835\udc53\ud835\udc53?\u0302?\ud835\udc4f, \ud835\udc53\ud835\udc53\ud835\udc51\ud835\udc51, \ud835\udc53\ud835\udc53?\u0302?\ud835\udc51.\nWire tensions fa, fa\u0302, fc, fc\u0302, fb, fb\u0302, fd, fd\u0302. are determined according to 4 motors\u2019 angles \u03d5a, \u03d5b,\u03d5c,\u03d5d As\nfa=kp ( \u03bb (\u03d5p+\u03d5a) 2\u03c0 -nL+ \u2211|ai-ai-1| n\ni=1\n) ,\nfa\u0302=kp ( \u03bb (\u03d5p-\u03d5a) 2\u03c0 -nL+ \u2211|a\u0302i-a\u0302i-1| n\ni=1\n) ,\nfc=kp ( \u03bb (\u03d5p+\u03d5c) 2\u03c0 -nL+ \u2211|ci-ci-1| n\ni=1\n) ,\nfc\u0302=kp ( \u03bb (\u03d5p-\u03d5c) 2\u03c0 -nL+ \u2211|c\u0302i-c\u0302i-1| n\ni=1\n) ,\nfb=kp ( \u03bb (\u03d5p+\u03d5b) 2\u03c0 -nL+ \u2211|bi-bi-1| m\ni=1\n) ,\nfb\u0302=kp ( \u03bb (\u03d5p-\u03d5b) 2\u03c0 -nL+ \u2211|b\u0302i-b\u0302i-1| m\ni=1\n) ,\nfd=kp ( \u03bb (\u03d5p+\u03d5d) 2\u03c0 -nL+ \u2211|di-di-1| m\ni=1\n) ,\nfd\u0302=kp ( \u03bb (\u03d5p-\u03d5d) 2\u03c0 -nL+ \u2211|d\u0302i-d\u0302i-1| n\ni=1\n) ,\n(14)\nwhere, \ud835\udf19\ud835\udf19\ud835\udc5d\ud835\udc5d is a motor rotation angle to generate a pretension, \ud835\udf06\ud835\udf06 is a lead of the screw rod and \ud835\udc58\ud835\udc58\ud835\udc5d\ud835\udc5d is the spring constant of the pretension spring.\nAuthorized licensed use limited to: Auckland University of Technology. Downloaded on October 25,2020 at 08:37:21 UTC from IEEE Xplore. Restrictions apply.", + "222\nC. Inverse kinematic solution\nAccording to given set of variables \ud835\udf03\ud835\udf03\ud835\udc65\ud835\udc65,\ud835\udc56\ud835\udc56, \ud835\udf03\ud835\udf03\ud835\udc66\ud835\udc66,\ud835\udc56\ud835\udc56 , \ud835\udf03\ud835\udf03\ud835\udc67\ud835\udc67,\ud835\udc56\ud835\udc56 (\ud835\udc56\ud835\udc56 = 1, \u22ef , \ud835\udc5b\ud835\udc5b) and \ud835\udc59\ud835\udc59\ud835\udc56\ud835\udc56 (\ud835\udc56\ud835\udc56 = 1, \u22ef , \ud835\udc5b\ud835\udc5b \u2212 1), we calculate the end-point position by Eq. (6),\n(pn 1 ) =H0,n ( 0 0 ln 1 ) (in jn kn rn 0 0 0 1 ) ( 0 0 ln 1 ) = (knln+rn 1 )\n(15)\nTaking a total differentiation of pn=knln+rn with respect to \u03b8x,i, \u03b8y,i , \u03b8z,i (i= 1, \u22ef, n) and li (i=1, \u22ef, n-1) and also motor angles \u03d5a, \u03d5b, \u03d5c, \u03d5d,\n\u2206pn= \u2202pn \u2202v \u2206v+ \u2202pn \u2202\u03d5 \u2206\u03d5 (16)\nWhere, v=(\u03b8x1,\u03b8x2, \u22ef,\u03b8xn, \u03b8y1,\u03b8y2, \u22ef,\u03b8yn ,\u03b8z1,\u03b8z2, \u22ef,\u03b8zn, l1 , l2 , \u22ef,ln-1 )\n\u2208R4n-1 and \u03d5=(\u03d5a, \u03d5b, \u03d5c, \u03d5d ). \u2202pn \u2202v \u2208R3\u00d74n-1 and \u2202pn \u2202\u03d5 \u2208R3\u00d74. Whereas, let w=(w1, w2, \u22efw4n-1)T=04n-1 represents the 4n-1 equations provided by Eqs. (9)(10)(12)(13), which also includes \u03b8x,i, \u03b8y,i , \u03b8z,i (i=1, \u22ef, n) , li (i=1, \u22ef, n-1) and also motor angles \u03d5a, \u03d5b, \u03d5c, \u03d5d. Taking a total differentiation for w=04n-1 as well, we have,\n\u2206w= \u2202w \u2202v \u2206v+ \u2202w \u2202\u03d5 \u2206\u03d5=04n-1 (17)\nwhere, \u2202w \u2202v \u2208R(4n-1)\u00d7(4n-1) and \u2202w \u2202\u03d5 \u2208R(4n-1)\u00d74 . Since \u2202w \u2202v is a square matrix, we can solve (19) with respect to the vector \u2206\ud835\udc63\ud835\udc63 as,\n\u2206v=- (\u2202w\n\u2202v ) -1 \u2202w \u2202\u03d5 \u2206\u03d5 (18)\nSubstituting (18) into (16), we have\n\u2206pn=\u2202pn \u2202v (\u2202w \u2202v ) -1 \u2202w \u2202\u03d5 \u2206\u03d5+ \u2202pn \u2202\u03d5 \u2206\u03d5=\n( \u2202pn \u2202\u03d5 - \u2202pn \u2202v (\u2202w \u2202v ) -1 \u2202w \u2202\u03d5 )\u2206\u03d5=J\u2206\u03d5\n, which can be solved for \u2206\ud835\udf19\ud835\udf19, by using a generalized inverse of the Jacobian J\u2208R3\u00d74\n\u2206\u03d5=J\u2020\u2206pn+P\u22a5(J)\u03a8 (19)\nwhere, J\u2020\u2208R4\u00d73 is a generalized inverse of \ud835\udc71\ud835\udc71 and P\u22a5(J)\u2208R4\u00d74 is a null projection operator of J, and \u2206\u03d5N\u2208R4 is a correction of \ud835\udf53\ud835\udf53 so as to minimize a positive scalar potential \ud835\udf11\ud835\udf11 by making use of a redundant actuation.\nWe use J\u2020=JT(J JT)-1 and P\u22a5(J)=I-J\u2020J. Eq.(19) provides a variation of motor angles \u2206\ud835\udf53\ud835\udf53 for a given position and direction variation \u2206pn . Applying the Euler method, we have the following variational equation, \u03c6+(\u2202\u03c6/\u2202\u03d5)\u2206\u03d5N=0\n,which is solved by\n\u2206\u03d5N=\u03c6 (\u2202\u03c6/\u2202\u03d5)(\u2202\u03c6/\u2202\u03d5)T (\u2202\u03c6 \u2202\u03d5)\nT\n( 20) As a candidate of \u03c6 , we take \u03c6=knz\n2 , where, k\ud835\udc5b\ud835\udc5b\ud835\udc67\ud835\udc67 is the z component of kn: the unit vector of the end-point orienting an axial direction. It means that the axial direction the endpoint takes on a horizontal plain as far as possible while keeping a designated position.\nAuthorized licensed use limited to: Auckland University of Technology. Downloaded on October 25,2020 at 08:37:21 UTC from IEEE Xplore. Restrictions apply.", + "223\nIV. EXPERIMENTS AND SIMULATION TakoBot 2 kinematics computed by Wolfram Mathematica software. As a robot controller, Arduino UNO was utilized, and for stepping motors, we used SilentStick TMC2208 motor drivers to provide sufficient power to the motors. During the experiments, manipulator demonstrated basic motions and helical motion in both directions. According to the conducted experiments, passive sliding mechanism increased robot bending stress tolerance and acted smoother than the previous prototype. Likewise, the mechanism discovered new horizons for continuum robot applications and imagination about continuum manipulators.\nV. REFERENCES\n[1] A.Yeshmukhametov, K. Koganezawa, Y. Yamamoto, \u201cDesign and kinematics of cable driven continuum robot arm with universal joint backbone\u201d,IEEE ROBIO 2018 conference proceeding, Kuala \u2013Lumpur, Malaysia. 2018. [2] V.C Anderson, R.C. Horn: Tensor Arm Manipulator Design, Mech. Eng 8998), 54-65. 1967 [3] R. Buckingham, A. Graham, \u201cNuclear snake-arm robots\u201d, Industrial Robot: An International Journal, Vol.39 Issue: 1,pp 6-11, https://doi,org/10.1108/01439911211192448, 2012. [4] M.W. Hannan and I.D.Walker, \u201cKinematics and the\nImplementation of an Elephant\u2019s Trunk Manipulator and other Continuum Style Robots\u201d, Journal of Field Robotics, pp 45-63,February 2003. https://doi.org/10.1002/rob.10070\n[5] Graham, R., Bostelman, R., \u201cDevelopment of EMMA Prototype suspended from the 6 m RoboCrane Prototype.\u201d Proc. ANS Seventh Topical Meeting on Robotics and Remote Systems,\u201d Augusta, GA, April 27-May 1, 1997. [6] I.A. Gravagne ; C.D. Rahn ; I.D. Walker, \u201cLarge deflection dynamics and control for planar continuum robots\u201d, IEEE/ASME Transactions on Mechatronics, Volume: 8 , Issue: 2 , June 2003. [7] Camarillo, D., Milne, C., Carlson, C., Zinn, M., Salisbury, J.: Mechanics modeling of tendon driven continuum manipulators. IEEE Transactions on Robotics (accepted for publication) (2008) [8] S. Neppalli, B. Jones, W. McMahan, V. Chitrakaran, I. Walker M. Pritts, M. Csencsits, C. Rahn, M. Grissom, \u201cOctArm - A Soft Robotic Manipulator\u201d, Proceedings of the 2007 IEEE/RSJ International Conference on Intelligent Robots and Systems San Diego, CA, USA, Oct 29 - Nov 2, 2007 [9] Jones, B., Walker, I.: Kinematics for multisection continuum robots. IEEE Transactions on Robotics 22(1), 43\u201355 (2006) [10] Haitham E., Usman M., Zain M.,Sang-Goo J., Muhammad U., Elliot W.H., Alisson M.O.,Jee-Hwan R., \u201cDevelopment and Evaluation of an Intuitive Flexible Interface for Teleoperating Soft Growing robots\u201d, 2018 IEEE/RSJ IROS, Madrid,Spain, October 1-5,2018. [11] A. Benouhiba, K. Rabenorosoa, P. Rougeot, M. Ouisse, N. Andreff, \u201c A multisection Electro-active Polymer based milli-Continuum Soft Robots. 2018 IEEE/RSJ IROS, Madrid,Spain, October 1-5,2018. [12] B. Ouyang, H. Mo, H. Chen, Y. Liu, D. Sun, \u201cRobust Model-Predictive Deformation Control of a Soft Object by Using a Flexible Continuum Robot, 2018 IEEE/RSJ IROS, Madrid,Spain, October 1-5,2018. [13] R. Kang, D.T. Branson, T. Zheng, E. Guglielmino, D.G. Caldwell, \u201cDesign, modelling and control of a pneumatically actuated manipulator inspired by biological continuum structures\u201d, Bioinspiration&Biomimetics 8 (2013) 14pp. [14] Bryan A. Jones, Ian D. Walker, \u201cA New Approach to Jacobian Formulation for a Class of Multi-Section Continuum Robots\u201d, Proceedings of the 2005 IEEE, ICRA, Barcelona, Spain, April 2005. [15] Han Yuan and Zheng Li, \u201cWorkspace analysis of cabledriven continuum manipulators based on static model\u201d, Robotics and Computer \u2013 Integrated Manufacturing. 49 (2018) 240-252. [16] X. Dong, M. Raffles, S. Gobos-Gusman, D. Axiente, J. Kell, \u201c A Novel Continuum Robot Using Twin-Pivot Compliant Joints: Design, Modelling and Validation, Journal of mechanism and robotics, February 16, 2015.\n[17] T. Liu, Z. Mu, H. Wang, W. Xu, Y. Li, \u201c A Cable-Driven Redundant Spatial Manipulator with Improved Stiffness and Load Capacity\u201d, 2018 IEEE/RSJ IROS, Madrid, Spain, October 1-5,2018\n[18] J. Starke, E. Amanov, M.Taha Chihaoui, J. Burgner-Kahrs, \u201cOn the Merits of Helical Tendon Routing in Continuum Robots\u201d, 2018 IEEE/RSJ IROS, Vancouver, BC, Canada, September 24-28, 2017.\nAuthorized licensed use limited to: Auckland University of Technology. Downloaded on October 25,2020 at 08:37:21 UTC from IEEE Xplore. Restrictions apply." + ] + }, + { + "image_filename": "designv11_71_0001949_j.mechmachtheory.2020.104139-Figure5-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0001949_j.mechmachtheory.2020.104139-Figure5-1.png", + "caption": "Fig. 5. Forces and moments acting on ball: (a) Conformal ball CVT, and (b) Non-conformal ball CVT.", + "texts": [ + " For analysis, contact between ball and disks in two ball CVTs is considered in the reference frames shown in Fig. 4 . It is set so that the force and moment direction that the ball exerts on the disk match the reference frames. The direction in which the ball inhibits the motion of the input disk and induces motion of the output disk is the direction of the force reference frames, and the direction of the moment is set based on \u03c9 ib and \u03c9 ob , indicated in red in Fig. 2 ; this is the direction of the moment that the ball exerts on the disk. Fig. 5 is a free body diagram for the input disk, output disk and one ball for each ball CVT. On the ball, the force equilibrium equation and the torque equilibrium equation are: F T c = 2 F N sin (27) F T i r ib \u2212 F T o r bo \u2212 F T c r c + M s i ,C cos ( \u03b1 + \u03b3 ) \u2212 M s o ,C cos ( \u03b1 \u2212 \u03b3 ) + M s c ,C cos ( \u03c0 2 \u2212 \u03b3 ) \u2212 T BL = 0 (28) F T i r ib \u2212 F T o r bo \u2212 F T c r c \u2212 M s i ,NC cos ( \u03b1 + \u03b3 ) + M s o ,NC cos ( \u03b1 \u2212 \u03b3 ) \u2212 M s c ,NC cos ( \u03c0 2 \u2212 \u03b3 ) \u2212 T BL = 0 (29) where T BL is torque loss generated by the bearing; a different equation can be used depending on the bearing type", + " 81 [ ( 2 B \u2212 1 ) ( 1 \u2212 0 . 5 C r o ) \u0303 \u03c9 o ] 0 . 67 G 0 . 53 0 . 201 \u00d7 ( \u02dc \u03c1e q x o )0 . 134 ( 1 \u2212 0 . 61 e \u22120 . 73 \u03b5 o ) (75) The contact model lets us obtain traction coefficients and spin momentum coefficients by integrating the shear stresses over the contact area. \u03bci = \u02dc ax i \u0303 ay i 1 \u222b 0 d R 2 \u03c0\u222b 0 \u02dc \u03c4i b x Rd \u03c8 (76) \u03bco = \u02dc ax o \u0303 ay o 1 \u222b 0 d R 2 \u03c0\u222b 0 \u02dc \u03c4o b x Rd \u03c8 (77) Because the z direction of the reference frame in Fig. 4 and the direction in which the spin momentum is defined in the free body diagram in Fig. 5 are different according to each CVT and contact point, the spin momentums are defined as follows. \u03c7i,NC = \u2212\u03c7i,C = \u02dc ax i \u0303 ay i \u0303 rCV T 1 \u222b 0 d R 2 \u03c0\u222b 0 ( \u02dc ax i \u0303 \u03c4i b y cos\u03c8 \u2212 \u02dc ay i \u0303 \u03c4i b x sin\u03c8 ) R 2 d \u03c8 (78) \u03c7o,C = \u2212\u03c7o,NC = \u02dc ax o \u0303 ay o \u0303 rCV T 1 \u222b 0 d R 2 \u03c0\u222b 0 ( \u02dc ax o \u0303 \u03c4o b y cos\u03c8 \u2212 \u02dc ay o \u0303 \u03c4o b x sin\u03c8 ) R 2 d \u03c8 (79) where the following coordinate transformation rule is used { X = Rcos\u03c8 Y = Rsin\u03c8 , 0 \u2264 R \u2264 1 , 0 \u2264 \u03c8 \u2264 2 \u03c0 (80) The traction coefficients and spin momentum coefficients between ball and cavity can be obtained by a similar process" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0001929_int.22304-Figure5-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0001929_int.22304-Figure5-1.png", + "caption": "FIGURE 5 Geometrical modeling of the mobile platform", + "texts": [ + " BS\u2010ZNN, bipolar\u2010sigmoid ZNN; FTNN, finite\u2010time neural network; L\u2010ZNN, linear activation ZNN; P\u2010ZNN, power activation ZNN; TCNN, time controlling neural network; ZNN, Zhang neural network [Color figure can be viewed at wileyonlinelibrary.com] In this section, we apply TCNN dynamics (9) as a solution law to realize the repeatable motion planning of mobile manipulator (three wheels platform). The structure of the proposed mobile manipulator is detailed in literature.31 The geometry of the mobile platform is visualized in Figure 5. (1) C: the center pointer of the mobile flatform; Its coordinate in world coordinate axis is x y z, ,c c c; (2) L: distance between Swedish wheels and C; (3) r : radius of the wheels; (4) \u03b80: orientation of the mobile platform; \u03b8\u03070 is the velocity of \u03b80; (5) \u03c6 \u03c6 \u03c6\u02d9 , \u02d9 , \u02d91 2 3: the velocity of each rotation wheel. For simplicity of expression, some matrixes are proposed to express part parament of the mobile manipulator P x y= [ ; ]c c c , A r L r L r L= [ (3 ), (3 ), (3 )]\u2215 \u2215 \u2215 , B r \u03b8 \u03c0 \u03b8 \u03c0 \u03b8 \u03b8 \u03c0 \u03b8 \u03c0 \u03b8 = 2 3 \u2212sin( ) \u2212sin 3 \u2212 sin 3 + cos( ) \u2212cos 3 \u2212 \u2212cos 3 + " + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0002108_s40799-020-00405-5-Figure2-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0002108_s40799-020-00405-5-Figure2-1.png", + "caption": "Fig. 2 Six-axis load head: a SOLIDWORKS [6] model; b built device", + "texts": [ + " The device must be very stiff so as to not distort the interconnected system dynamics. Both the high stiffness and sensitivity are obtained by the use of highsensitivity uniaxial load cells as constituents. The device must also have uniform sensitivity in all axes. The proposed device meets these requirements. As mentioned in the previous section, a Stewart platform consists of a base plate, and a top platform, connected by six struts. A SOLIDWORKS [6] model, and a picture, of the actual devices discussed here are shown in Fig. 2. The load heads consist of two hexagonal plates, connected by six commercially available 1000 lb uniaxial load cells (LC1011K [26]); other types of individual load cells may be used depending on the requirements of an application. In the interest of making the load heads as light as possible, the plates are made out of aluminum and, as shown in Fig. 3, circular holes are cut at their center. The hole in the bottom plate is smaller merely to fit a bolt to connect it to an electrical bushing (see Fig" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0000763_s00202-020-00967-y-Figure2-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0000763_s00202-020-00967-y-Figure2-1.png", + "caption": "Fig. 2 Exploded view of single-phase stack: a TFM; b HFM", + "texts": [ + " The proposed HFM topology is a structurally modified version of TFM topology. So, the performance of HFM is compared with that of TFM. In the literature, two topologies of TFM, without pole shaping [7,9,18] and with pole shaping [16,19], are reported. However, the comparison study between these two topologies is not available. Hence, both TFM topologies (without and with pole shaping) are considered to compare with HFM. The single-phase stack of TFM and the proposed HFM is shown in Fig. 1, and its exploded view is given in Fig. 2. Three-phase topology with outer rotor configuration is considered in this paper. The complete machine can be formed by combining three single-phase stacks with 120\u25e6 electrical angle between them. Each phase stack consists of p number of pole pairs. A 6-pole pair design is considered in this section to explain the structure of machines. Soft magnetic composite (SMC) is assumed as the core material to avoid the lamination difficulties in the manufacturing of TFM and HFM. In a single-phase stack of TFM, the stator has 12 poles, and the rotor has 24 poles. The stator is a hollow-shaped cylinder with 6 poles projected outwards at each side of the cylinder as shown in Fig. 2a. A ring-shaped coil is placed between the projected poles. The rotor has two layers of surface-mounted magnets with 12 magnets in each layer. The span between the stator poles is twice of the span between the rotor poles as shown in Fig. 1a. Due to this, only half the number of magnet\u2019s flux can link with the stator coil at maximum. In HFM, the number of poles in stator and rotor is 12. The stator is made of a hollow cylinder core with 6 outwards projected poles (I-core) at themiddle of the cylinder as shown in Fig. 2b. At both sides of the projected poles, a ring coil is inserted, and in between two I-cores, a C-core is placed. Also, both the coils are enclosed by C-cores. This C-core can be supported at both ends by non-magnetic frames. The rotor consists of 12 surface-mounted magnets in a single layer. In this topology, the stack length of stator is higher than that of rotor. For a p-pole pair design, the overall difference between the structures of TFM and HFM is given in Table 1. The magnetic flux path of one pole pair of TFM and HFM is shown in Fig" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0003449_20140313-3-in-3024.00239-Figure1-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0003449_20140313-3-in-3024.00239-Figure1-1.png", + "caption": "Fig. 1. A manipulator with 2-DOF", + "texts": [ + " Thus the incremental cost can be given by the following \u2206V (x(t0)) = \u222b t0+T t0 \u03d5\u2217(x) dt (21) where, \u03d5\u2217(x) is related to u\u2217. In theory, the difference of cost estimated by critic network at tth0 and (t0 + T )th instant should be same as \u2206V (x(t0) in (21) when the network is trained properly. Hence, we learn the weight in such a way, so that the difference between \u2206V (x(k)) and (V (P(k) \u2212 V (P(k + T )) is minimized. This must be noted that the constraints should be maintained while minimizing the error norm. Let us consider a manipulator of two DOF given in Fig. 1 which is taken as an example from Craig (2009). M(q)q\u0308+C(q, q\u0307) = \u03c4 \u2212 g(q) (22) Where, M(q) = ( m1l 2 C1 + Izz1 +m2l 2 1 +m2l 2 C2 +Izz2 + 2m2l1lC2 cos(q2) ) ( m2l 2 C2 + Izz2 +m2l1lC2 cos(q2) ) ( m2l 2 C2 + Izz2 +m2l1lC2 cos(q2) ) ( m2l 2 C2 + Izz2 ) and C(q, q\u0307) = ( \u22122m2l1lC2 q\u03072sin(q2)q\u03071 \u2212m2l1lC2 q\u030722sin(q2) ) (m2l1lC2 q\u03071sin(q2)) And the mass of link 1 & 2 are given by m1 and m2 respectively. For the given manipulator configuration equation (22) can be represented in state space form [ m11 m12 m21 m22 ] [ q\u03081 q\u03082 ] + [ c11 c12 c21 c22 ] [ q\u03071 q\u03072 ] = [ \u03c41 \u03c42 ] \u2212 g(q), (23) where mij and cij , i = 1, 2; j = 1, 2 are the elements of M and C respectively" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0000347_iecon.2019.8927574-Figure2-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0000347_iecon.2019.8927574-Figure2-1.png", + "caption": "Fig. 2. Schematic of the stator structure of flux reversal PM machine.", + "texts": [ + " Although some of the presented structures are really potential for the aforementioned applications, their complex structures cause them not to be widely produced; some examples of these structure are stator-PM machines, AxialField PM (AFPM) machines, Transverse Flux PM Machine (TFPM), Claw Pole machines, Spoke-Type PM machines. On the other hand, these structures can have higher efficiency, as well as compact sizes in comparison with the traditional PM machines. Stator-PM machines are an attractive structure due to placing both magnets and coils in the stator structure only. Flux Switching PM (FSPM) machines (Fig. 1) and Flux Reversal PM (FRPM) machines (Fig. 2) are two common stator-PM structures [2]-[4]. Placing the PMs in the stator structure makes these structures an appropriate candidate for transportation applications because they have salient and robust rotor without PMs or windings; good thermal dissipation as the magnets located on stator; and the armature field has little influence on the PM due to the PM field and the armature field are in parallel [5]. For instance, linear FSPM machines could be an interesting solution for the railway application because both PMs and windings are positioned on stator and translator does have a very simple and cheaper structures in compared to the other typical types [5], [6]" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0002861_978-3-030-55889-5-Figure3.49-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0002861_978-3-030-55889-5-Figure3.49-1.png", + "caption": "Fig. 3.49 Principal stresses in a cellular beam with fixed vertical edges", + "texts": [ + " The larger value uy means a larger part of the load will flow through the center element. Vice versa, if the material is less solid, the neighboring zones will hinder the propagation of the dislocation and so only a small part of the load will flow through the soft element. How stiff zones in a plate attract loads can be seen in Fig. 3.48, where the uniform load is mostly carried by the more solid zones of the plate, h = 40 cm, versus h = 20 cm for the inclusions. In the limit, if the inclusions have zero stiffness, E = 0, we see results as in Fig. 3.49, where all stresses are forced to flow around the holes. 222 3 Finite Elements Drop panels produce similar effects, since they cause the bending moments to migrate to the column capitals. Note, the jump in the plate thickness makes that the bending moments myy in a horizontal cross section\u2014perpendicular to the discontinuity\u2014jump while the moments mxx are not affected, see Fig. 3.50. In a vertical cross section through the center it would just be the opposite. In FE-analysis any linear functional J (uh) = gT f is the scalar product of the nodal values g of the influence function and the vector f of the equivalent nodal forces" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0003634_icems.2014.7013745-Figure4-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0003634_icems.2014.7013745-Figure4-1.png", + "caption": "Fig. 4. No-load magnetic field distribution (another position)", + "texts": [ + " Therefore, it is necessary to analyze by FEM in order to figure out detailed electromagnetic coupling of PMRM. III. ANALYSIS BY FINITE ELEMENT METHOD In order to analyze electromagnetic coupling of PMRM quantitatively, finite element method (FEM) is used. In the PMRM, because of the doubly salient structure of outer motor, magnetic circuit of outer motor will be different when outer rotor rotates to a different position. Magnetic field distribution of PMRM when permanent magnets act alone is shown in Fig. 3 and Fig. 4. As Fig. 3 shows, when outer rotor rotates to this position, magnetic flux will go through two aligned teethes of stator and outer rotor. When there are no aligned teethes, magnetic flux will go through adjacent teethes as Fig. 4 shows. Magnetic field distributions of PMRM when inner rotor or stator armature windings energized alone are shown in Fig. 5 and Fig. 6. motor will increase or decrease when +D axis or \u2013D axis currents of inner motor act alone. By comparing Fig. 5(c) with Fig. 5(d), it can be seen that when +Q axis currents act alone, magnetic fluxes of outer motor little change. From Fig. 5, it can be obtained that D axis currents of inner motor has much stronger effect on magnetic field distribution of outer motor than Q axis currents of inner motor" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0003479_mipro.2015.7160432-Figure4-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0003479_mipro.2015.7160432-Figure4-1.png", + "caption": "Fig. 4. Fully distributed FCS architecture [5]", + "texts": [], + "surrounding_texts": [ + "Continuously increasing requirements for aircraft and air transport safety along with operational demands for reliability, performance, efficiency and costs, are shifting the focus of recent development to distributed systems. The massive voting architecture proposed by Airbus [4] suggests to allocate the task of control laws and logic between flight control computers and control surface actuator nodes as shown in Fig. 5. Flight control computers and actuator nodes are connected via an advanced data communication network developed by Airbus. Flight control computers calculate the control laws and proprietary commands for control surface actuator nodes, which are then broadcast as messages over the communication bus. Actuator nodes are equipped with flight control remote modules, and perform massive voting upon receiving the messages from many flight control computers. The massive voting architecture resides upon digital communication technologies. New smart actuator technologies are explored for particular system application. Fault handling in the system proposed from Airbus is resolved within the actuator nodes. A high degree of fault detection as well as a fault location is demonstrated, both due to the large number of nodes [5]. A distributed FCS architecture is presented also for accessing fault handling and redundancy managing on the JAS39 Gripen [6]. The proposed system included 16 nodes. Various simulations showed that distributed sensor nodes meet fault detection coverage of 99% for both transient and permanent faults. The proposed system used a triggered multi master broadcast bus with Time division multiple access communication. As a result, the failure on any node cannot jeopardize communication by sending data outside the dedicated time slot, resulting in a fail silent system. Power line communications (PLC) have been proposed for distributed aircraft control systems in [7]. The PLC communications approach eliminates the need for digital data bus wiring by modulating the data on power cables that are installed between the flight control computer and control surface actuators. Although technology is promising and widely used in other applications, vehicular control systems are not usually installed with PLC systems. For aircraft\u2019s FCS, there are many requirements that make implementation of PLC difficult, such as using negative return wires on the power bus instead of chassis return as usual. The problem arises from selective frequency fading or multipath fading. Furthermore, as a general system design safety rule requires, primary and secondary flight surfaces must remain independent. More than one network must be used to reduce wiring, and also for tail surface reliability. Communication speed requirements for various standards must be met, and also to ensure reliability with a given number of remote units. From many other aspects, PLC has to be further developed for aircraft use and its usability is yet to be explored. Decentralisation is entering other aircraft subsystems, with the development of larger and more complex aircraft. Smart components are proposed for a decentralised fuel management system [8] and microcontrollers are embedded in the pumps, valves and sensors (Fig. 5). Proposed system components make its own decisions during various fuel operations, depending on the performed action. They share a time-triggered bus for communication. When a smart component reaches a decision, it transmits it over the bus. For safety, all system components retain a copy of the state vector that describes the system state. Laboratory and real scale testing have been performed proving such a distribution is possible and that the new system can be adaptable to faults." + ] + }, + { + "image_filename": "designv11_71_0001566_012084-Figure5-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0001566_012084-Figure5-1.png", + "caption": "Figure 5. Dynamic finite element model.", + "texts": [ + " The average contact surface pressure of the joint surface can be written as F P S (6) In the formula, P is the joint surface pressure in MPa; S is the joint area in mm2. The normal dynamic stiffness and tangential dynamic stiffness of the single bolt joint obtained from (3), (4), (5), and (6) are shown in Table 2. For finite element modeling, the 20-noded element 186 is selected as the element type and the element length is set as 0.5. Since the spring element is to be established at the corresponding node, it is necessary to define a hard point before free meshing. The dynamic finite element model is established as shown in Figure 5. IWMSME 2020 IOP Conf. Series: Materials Science and Engineering 892 (2020) 012084 IOP Publishing doi:10.1088/1757-899X/892/1/012084 In the virtual material layer method, the asperities between the two surfaces are assumed equivalent to a layer of virtual material with a certain thickness. The virtual material layer and the component surface were coupled by a fixed connection [17]. The equivalent schematic diagram is shown in Figure 6. The property parameters of the virtual material layer include three material parameters and one structural parameter, namely the elastic modulus, Poisson's ratio, density, and thickness, and four additional parameters are calculated as follows" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0001156_cis-ram47153.2019.9095819-Figure3-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0001156_cis-ram47153.2019.9095819-Figure3-1.png", + "caption": "Fig. 3. frequent domain of PKP-SEA", + "texts": [ + " Then the angle of the edge of a leaf spring is calculated as follows: (2) We tested this spring mechanism with a PKP-SEA. The experimental result shows that the sensor succeeded to measure the torque, and the resolution of the sensor is adjusted passively in relation to the applied force (Fig. 8). When its displacement of spring is small ( ), the torque changes almost linearly, while it dramatically increases or decreases when its displacement becomes large ( ). is estimated based on these data with 8th order polynomial function described as a black solid line in Fig.8. From the diagram shown in Fig. 3, we can obtain the following model in the time domain: (3) With the electrical model, the equations governing the electrical power consumption can be obtained from (3) and (4): (5) (6) where im and vm are the current and voltage of the geared motor, Rm is the winding resistant of geared motor, and kt is the torque constant of the geared motor. With these equations, we can control the torque of the knee joint by applying the voltage based on the required reference torque. III. METHODOLOGY A. Subject and Experimental Setup The preliminary experiment is conducted with one aboveknee amputee (male, 45 yrs, 172cm, 62kg, left above knee amputation)" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0000356_iecon.2019.8927121-Figure2-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0000356_iecon.2019.8927121-Figure2-1.png", + "caption": "Fig. 2. Two DoF actuator topologies [6]", + "texts": [ + " Due to their particular movement, the RLMs require also specific linear-rotary bearings. These must avoid the shaft to rotate smoothly and with low friction simultaneously with the straight-line movement of the machine. There are three basic types of electrical machines with two DoF: the surface (planar) motors with a flat structure [3], the spherical motors [4] and the rotary-linear ones, having a cylindrical construction [5]. Generally, combined rotary-linear movements can be achieved by means of three actuator topologies, as shown in Fig. 2 [6]. To assure the complex movement, the most facile solution is to use two parallel coupled conventional rotary and linear actuators. By means of this, the two motions are perfectly independent. The main drawback of this approach is the high space requirement, because the linear motor must move also the entire rotary motor. The series coupled structure has two separate stators with a common shaft, and therefore it is simpler. The distance between the two stators must be at minimum the linear stroke of the RLM" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0001895_1350650120964295-Figure8-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0001895_1350650120964295-Figure8-1.png", + "caption": "Figure 8. Occurrence of multiple contact points for torustorus contact. For better visualisation, the force application points Pm are plotted.", + "texts": [ + " To which points the algorithm converges highly depends on the given start value. Similar problems can be encountered for cone-torus contact pairings in the semi-analytical approach and by Gupta.7 In this case, however, only a one-dimensional function is minimized, which makes handling of false solutions due to convergence issues less complex. Thirdly, the contact between two tori may occur simultaneously at two points, if their rotational axes are coplanar and the tori overlap each other considerably (e.g. in Figure 8). This generates two problems: First, the gradient method10 can only detect one pair of contact points, whereas for the semianalytical approach penetration can be calculated for each extrema and therefore all contact points are detected. Second, the torus of the roller and the tori of the flange are often parallel at the initial time step of MBS. This may lead, depending on the overlap, to multiple contact points. When the roller skews in the following time steps, the contact now only occurs in the proximity of one of the previous contact points" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0001411_ilt-03-2020-0083-Figure1-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0001411_ilt-03-2020-0083-Figure1-1.png", + "caption": "Figure 1 Structure diagram of the friction pair of the SG-LFS", + "texts": [ + " So far, most research studies on cavitation focuses on the cavitation boundary, effects of cavitation on the sealing performances and its evolutionary characteristics. However, few studies analyze the effect mechanism of various factors on cavitation induction. In this study, a mathematical model based on the JFO boundary is established and a relative density characterizing the cavitation degree is defined to investigate the effects of operating conditions including the process coefficient, lubricant viscosity and cavitation pressure on the cavitation induction. The friction pair of downstream pumping SG-LFS is shown in Figure 1, the surface of the stationary ring is groove-free and that of the rotating ring is fixed with evenly distributed logarithmic spiral grooves. With a relative rotation, the hydrodynamic effect is produced in the both groove areas and land areas, and the hydrostatic effect is generated in both the inner dam area and outer dam area (Basu, 1992). Due to their combined actions, the friction pair is in a state of fluid lubrication and the SG-LFS realizes a non-contacting operation. In Figure 1, the land area is located between the adjacent grooves, the inner dam area between both the inner radii of the sealing face and the spiral groove and the outer dam area between both the outer radii of them. In Figure 1(a), h0 and hg represent the thickness of groovefree area and the groove depth, respectively. Because of the groove, the thicknesses of groove-free area and groove area are different, the value of the former is h0 and that of the latter is h01hg. In Figure 1(b), ri and ro are the inner and outer radii of the sealing face. rg1 and rg2 are the inner and outer radii of the spiral groove. a represents the spiral angle. uG and u L represent the circumferential angles of the groove and land. pi and po are the inner and outer pressures, respectively. For the SG-LFS, the geometrical parameters are shown in Table 1. Based on the following assumptions: the lubricant between the friction pair is Newtonian and its flow is laminar; both the thermal wedge and thermal distortion are neglected; and the angular misalignment and fluid inertial are neglected, the mass-conservation average flow Reynolds equation in polar coordinate is expressed as follows (Pinkus and Lund, 1981): 1 r @ @r r rh3 m @p @r ", + " To indicate the cavitation degree of each point, a relative density \u00ab r is introduced and it is defined as the ratio of the equivalent density r e to the density of pure liquid phase r liq. To realize the JFO boundary, an F-f cavitation algorithm (Payvay et al., 1999) is adopted . The cavitation index F is defined as 0 and 1 in the cavitation and liquid-film zones, respectively. Then the variable f at the liquid film zone is greater than zero and less than zero at the cavitation zone, its expression is defined as follows: f \u00bc p pc\u00f0 \u00de= po pc\u00f0 \u00de f \u00bc \u00ab r 1 ( (2) In Figure 1(b), the spiral grooves are uniformly distributed and the one-cycle computational domain is bounded by two periodic boundaries \u2039 and \u203a. Because of the irregularity of spiral curves, a coordinate transformation method (James and Potter, 1967) is used to transform the irregular region to a fanshaped regular one. The transformation function is expressed as follows: j\u00bc r h \u00bc u 1=tana\u00f0 \u00deln r=ri\u00f0 \u00de ( (3) Substituting equations (2) and (3) into equation (1) and performing a dimensionless treat, the universal governing equation in the new coordinate system yielded is written as follows: @ @j j @ Ff\u00f0 \u00de @j 1 a j @ Ff\u00f0 \u00de h 3 @h a j @ @j j @ Ff\u00f0 \u00de @h a j @ @h j @ Ff\u00f0 \u00de @j 1 a2 11\u00f0 \u00de @ @h @ Ff\u00f0 \u00de j @h " + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0001564_lra.2020.3013894-Figure1-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0001564_lra.2020.3013894-Figure1-1.png", + "caption": "Fig. 1. Bioinspired variable gearing systems: (a) Pennate muscle [11], (b) Pneumatic actuator array [12], (c) Spatially distributed actuation of a finger [21], (d) Distributed actuation mechanism [13].", + "texts": [ + " In Section II, the concept of bioinspired variable gearing in the DAM is explained. Then, the SGR is defined and mathematically expressed, based on the leadscrew actuation. Section III describes the simulation-based optimization that can determine the maximum performance of the DAM-3R manipulator. Then, the effects of the SGR are analyzed with a comparison of an equivalent JAM-3R manipulator. The conclusion follows in Section IV. Pennate muscles move with varying gear ratios by changing the pennation angle (\u03b8 in Fig. 1(a)), depending on the external load [11]. Based on this observation, a biomimetic pneumatic actuator array was proposed in [12]. Considering the above action of pennate muscles, the pneumatic actuator array can change the pennation angle (\u03b8m in Fig. 1(b)), which leads to a change in the gear ratio (Fig. 1(b)). Similarly, as shown in Fig. 1(c), the fundamental movement of human fingers, such as stretching or expanding, can be accomplished through contraction and relaxation of the flexor digitorum profundus, flexor digitorum superficialis, and extender digitorum [21]. Concurrently, sophisticated and dexterous movements can be achieved by the spatially dispersed actuation of opponens pollicis over the finger rather than by the lumped actuation. In [20], the spatially distributed actuation of opponens pollicis was implemented in the DAM by controlling a slider that moves along a link (Fig. 1 (d)). Ultimately, both systems control the force vector for variable gearing. However, it should be noted that the pneumatic actuator array change the direction of force to adjust the architectural gear ratio [12], whereas the DAM changes the point of application of the force [20]. This study will quantitatively investigate the DAM from the viewpoint of continuously variable gearing. In this study, a leadscrew was used to implement the linear movement of the slider along the link. Considering the offset between a hinge joint and slider (h in Fig" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0002204_icem49940.2020.9270855-Figure12-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0002204_icem49940.2020.9270855-Figure12-1.png", + "caption": "Fig. 12. Stator and rotor PM eddy current distribution with rated current. (a) Rotor PM part. (b) Stator PM part.", + "texts": [ + " 0 60 120 180 240 300 360 -240 -180 -120 -60 0 60 120 180 240 V o lt ag e ( V ) Position (Electrical deg) Rotor PM Stator PM Dual -16 -12 -8 -4 0 4 8 12 16 Current C u rr en t (A ) Voltage Fig. 11. Comparison of voltage and current waveforms at rated current. For the proposed machine, in addition to the mechanical loss, the main losses contain copper loss, iron loss, and PM eddy current loss in both stator and rotor parts. The stator/rotor PM eddy current distributions, and iron loss, PM loss waveforms at rated current are shown in Fig. 12 and Fig. 13. The rotor PM loss is high, which may cause the over high temperature rise of the rotor. However, it can be effectively suppressed by the PM segmentation which is mature. What is more, double-layer instead of single-layer winding which can greatly suppress the subharmonics of the armature MMF will also reduce the PM loss obviously. The loss and efficiency during rated operation are summarized in Table V. an efficiency of 93.9% is obtained. The introduction of PM in stator slot increases torque capability, power factor and efficiency" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0001035_s002565441907001x-Figure2-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0001035_s002565441907001x-Figure2-1.png", + "caption": "Fig. 2. Fig. 3.", + "texts": [ + " Therefore, the optimal control is with two switching points, including two sections on which |M| = const. A control is implemented with three sections of the change in the moment M: intensive acceleration (when m(t) = const = m0), a section with a parabolic change in the modulus of the kinetic moment (m(t) is a linear function of time, decreasing from m(t) = m0 to m(t) =\u2212m0) and intensive deceleration (when m(t)= const =\u2212m0). The switching points are t1 =55 s and t2 = 145 s. The results of mathematical modeling of the dynamics of the optimal turn are presented in Figs 2\u20135. Fig. 2 shows graphs of the changes in angular velocities in the coordinate system \u03c91(t), \u03c92(t), \u03c93(t) in time related to the spacecraft (the variables \u03c9i are given in degrees / s). Fig. 3 reflects the nature of the change in the components of the quaternion \u039b(t), which determines the current orientation of the spacecraft during the ongoing rotational maneuver: \u03bb0(t), \u03bb1(t), \u03bb2(t), \u03bb3(t). Fig. 4 shows the dynamics of changes in the components p1(t), p2(t), p3(t) of the vector p in time (the values of pi, like \u03bbj , are dimensionless quantities)" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0002281_smc42975.2020.9283493-Figure3-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0002281_smc42975.2020.9283493-Figure3-1.png", + "caption": "Fig. 3. The obstacle (left) and the sphere-swept volume (right).", + "texts": [], + "surrounding_texts": [ + "manufacturing efficiency, we are working on an FDM (Fused Deposition Modeling) printer called DEXTER which has dual Selective Compliance Articulated Robot Arms (SCARA). We develop a path planner to produce time-optimal motion paths for each of DEXTER\u2019s arms while guaranteeing that the arms do not collide with one another. We present a collision-free path planner for DEXTER\u2019s arms using an improved Sampling-Based Model Predictive Optimization (SBMPO) based on A* type optimization by adding efficient collision determination and a new type of cost function. The simulation results show that the improved SBMPO can be used to efficiently generate smooth collision-free paths and trajectories with bounded velocity and acceleration.\nKeywords\u2014 Trajectory planning, Dual-arm 3D printer,\nCollision detection, SBMPO\nI. INTRODUCTION\nThree-dimensional (3D) printing is an additive manufacturing (AM) technique in which a component is built by adding many thin layers of material one on top of the other. It is a revolutionary technology that enables the manufacture of customized parts, prototypes and intricate geometry objects more efficiently than traditional manufacturing techniques[1]. Characteristics of industrial robotics make robot-based 3D printers attractive alternatives to traditional gantry-style 3D printing [2]. Three advances afforded by robot-based printers are:\n1. Multi-direction printing. In contrast to a conventional printer which can only print layers perpendicular to the build direction, articulated robot arms can orient the print nozzle in multiple directions thereby enabling the printing of supportfree layers. [3][4].\n2. Secondary Operations. Besides printing, a robotic arm printer with different end-effector tools can also assemble, drill and weld on the printed part [5].\n3. Limitless Range. A robotic printer with a mobile base can potentially print parts of unlimited size [6].\nThe never-ending pursuit of increased throughput led the manufacturing community to consider multi-robot systems. Similarly, a multi-robot AM system can expand the advances noted above. Perhaps most importantly, a multi-robot printer can greatly improve the efficiency of all types of AM tasks. However, printing a part using multiple robots requires the production of collision-free constrained trajectories for each of the robots [7]. Single manipulator motion planning with obstacle avoidance has been researched widely in the past decades with researchers developing algorithms to address problems such as avoiding static or moving obstacles, path planning and time optimal trajectory planning [8-14]. Here we should mention the difference between these three concepts. Path planning is a purely geometric concept which produces a series of intermediate via-points along the end-effector path moving from point A to point B without a specified time law in task space or configuration space. Trajectory planning focuses on finding a time law to fulfill the motion along the geometric path. Finally, motion planning includes both path planning and trajectory planning.\nThe most common method for motion planning is the decoupled approach [8] which can be separated into a highlevel and lower-level planner stages. The high-level planner focuses on computing a collision-free path in task space or configuration space but ignoring the dynamics of the robot. The most popular methods are artificial potential fields (APFs) [9], RRT [10] and PRM [11]. After finding the collision free path, the planner must design a trajectory along the path \u2013 path tracking stage. A time optimal trajectory along the geometric path is determined whereby the manipulator dynamics and actuator constraints are considered [8]. The methods include Genetic Algorithm and Harmony search algorithm.\nAll of these methods are applicable to multi-manipulator systems, but the solutions are more complex due to the higher number degrees-of-freedom of the combined robots and the fact that each arm is an obstacle to be avoided by the other. These additional complications while difficult define our interest in creating collision free motion plans for multimanipulator AM systems. [15-17]\nShow in Fig. 1, DEXTER [18] is a robot-based printer with dual Selective Compliance Articulated Robot Arms (SCARA).\n978-1-7281-8526-2/20/$31.00 \u00a92020 IEEE 2389\nAuthorized licensed use limited to: Carleton University. Downloaded on June 17,2021 at 05:12:46 UTC from IEEE Xplore. Restrictions apply.", + "This platform is the example system for developing and testing our planner. In this system, each arm has two links and three stepper motors to position the nozzle at the end of each arm in 3-degree of freedom (DOF). Each of Dexter\u2019s arms must print in collaboration with the other, or one arm can print while the other\u2019s nozzle is replaced with alternative tooling to perform secondary operations. We must use a path planner to produce motion paths for each of DEXTER\u2019s arms while guaranteeing that the arms do not collide with one another.\nOur approach simplifies the original SMBPO motion planer by employing a purely A* algorithm while retaining SBMPO\u2019s sampling approach and algorithmic structure. Due to SBMPO\u2019s open structure, we change the cost function to improve the speed of sampling and performing collision checking. In addition, this new cost function has the ability to achieve time-optimal path planning. Consequently, the dualmanipulator system can achieve collision-free and time optimal motion planning with bounded joint velocity and acceleration.\nThis paper is organized as follows: Section II presents a collision-free motion planning problem formulation for a single 2-DOF manipulator and gives a brief discussion of the SBMPO strategy to solving the single manipulator problem. Section III provides the implementation of the improved SBMPO algorithm in both single and dual manipulator system simulations. Finally, the conclusions are discussed in section IV.\nSampling-Based Model Predictive Optimization (SBMPO) is a motion planning algorithm that utilizes kinematic and dynamic models to directly generate trajectories using a heuristic cost function. Only the initial and final position and velocities of the trajectory are specified. Based on the A* algorithm, the core process of SBMPO is to predict the next system state from the current system state and sampled control inputs [20]. SBMPO can produce motion trajectories directly and generate a smooth path in task space. In addition, using the concept of an implicit grid data structure, the planner is able to escape from dead-lock positions. Fig. 2 shows the block diagram of the SBMPO trajectory planning strategy.\nThe following are the main steps of SBMPO [21]:\n1. Select the vertex with highest priority in the queue\n2. Sample the input space\n3. Add a new vertex to the graph\n4. Evaluate the new vertex cost\n5. Repeat 2-4 for B number of successors\nPrevious papers [21][26] have verified the completeness in manipulator motion planning with SBMPO and its advantages in solving local minimal problem. Like most planners, SBMPO requires modification before it can be used in actual field environments. We added the following features to the original SBMPO:\n1. Effective detection of positions that cause collision of the manipulator with obstacles in the workspace.\n2. Decreased computation time especially in complex working spaces, such as those that produce dead lock positions.\n3. Motion planning in an environment containing moving obstacles.\nIn short, we upgrade the SBMPO algorithm with an obstacle-avoidance function and a new cost function for timeefficient path planning in the face of obstacles.\nIn a typical robot motion scenario, obstacles can be represented by polygons including convex polygons and concave polygons. To ensure a safe distance between the robot and the obstacles, we employ a sphere of radius r to sweep the inner primitive forming the Minkowski sum [23] of the sphere and the primitive. The radius r is determined by the safety distance (SD) that depends on customers\u2019 request.\nWith regard to obstacle avoidance, in each single sampling process the shortest distance between the manipulator (not only the end-effector but also the points along the manipulator\u2019s links) and the obstacle has to be compared with the safety distance. If the distance is equal or larger than the safety distance, the vertex is valid. But, if the distance is shorter than\n2390\nAuthorized licensed use limited to: Carleton University. Downloaded on June 17,2021 at 05:12:46 UTC from IEEE Xplore. Restrictions apply.", + "the safety distance, the vertex should be removed. The process is shown in the Fig.4.\nInstead of selecting the vertex with the highest priority in the queue as in a standard A* cost function we provide a new cost function (Eq. 1) which combines the distance and the velocity of the end-effector. In path planning, a path (shown in Fig. 5 connects the start vertex \ud835\udc630 to the desired destination \ud835\udc63\ud835\udc54 can be defined by a set of important via-points \ud835\udc631, \ud835\udc632, \u2026 , \ud835\udc63\ud835\udc56 , \u2026 , \ud835\udc63\ud835\udc54 along the end-effector path. Each point with the information of Cartesian coordinates ( \ud835\udc65\ud835\udc56 , \ud835\udc66\ud835\udc56 ) and velocity (vi).\nThe cost function of uniformed Cartesian coordinate and\nvelocity can be defined as follows,\n\ud835\udc39(\ud835\udc63\ud835\udc56) = \ud835\udefc\ud835\udc3b(\ud835\udc63\ud835\udc56) + \ud835\udefd\ud835\udc3a(\ud835\udc63\ud835\udc56) + \ud835\udefe\ud835\udc49(\ud835\udc63\ud835\udc56). (1)\nIn above equation, \ud835\udc39(\ud835\udc63\ud835\udc56) defines the cost of each vertex (\ud835\udc63\ud835\udc56) generated by the sampling process and \ud835\udc3b(\ud835\udc63\ud835\udc56) is the heuristic value which stands for the estimated distance between\nthe current vertex to the target destination in the task space. The\nestimated distance can be calculated as,\n\ud835\udc3b(\ud835\udc63\ud835\udc56) = \u221a(\ud835\udc65\ud835\udc54 \u2212 \ud835\udc65\ud835\udc56) 2 + (\ud835\udc66\ud835\udc54 \u2212 \ud835\udc66\ud835\udc56) 2 . (2)\n\ud835\udc3a(\ud835\udc63\ud835\udc56) is the traveled distance from the start position to the current position or,\n\ud835\udc3a(\ud835\udc63\ud835\udc56) = \u2211 \ud835\udc3a\ud835\udc56 \ud835\udc56 1 , (3)\nWhere \ud835\udc3a\ud835\udc56is the actual distance between vertex \ud835\udc63\ud835\udc56\u22121 and \ud835\udc63\ud835\udc56. \ud835\udc49(\ud835\udc63\ud835\udc56) is the velocity element, which is defined as,\n\ud835\udc49(\ud835\udc63\ud835\udc56) = \ud835\udc51\ud835\udc61 \u2217v(\ud835\udc63\ud835\udc56), (4)\nWhere v(\ud835\udc63\ud835\udc56 ) is end-effector velocity of the vertex \ud835\udc63\ud835\udc56 . The weighting factors \ud835\udefc, \ud835\udefd and \ud835\udefe must meet the following\nconditions,\n{ \ud835\udefc, \ud835\udefd \u2265 0 \ud835\udefe \u2264 0\n\ud835\udefc + \ud835\udefd + \ud835\udefe = 1 . (5)\nThe path generated by task space (workspace) sampling focuses strictly on the end-effector position and results is the shortest path to the goal, but problems occur when there are more than one possible joint configuration for a given task position. In our 2-DOF system, each task space position has 2 possible joint configurations: denoted elbow up and elbow down. Moreover, after the path is found and the joint trajectory is calculated using inverse kinematics, the resulting motion is generally not smooth and may not even be possible given manipulator joint limits and limits imposed by system dynamics. Therefore, sampling in task space appears more suitable for mobile robots or free workspace environments. To avoid these issues, sampling in configuration space (joint space) is a better choice. Each sample in joint space represents one configuration in task space. However, finding a path to the goal through joint space sampling requires more time than sampling in task space due to the nonlinear transformation of obstacles from task space to configuration space. Finally, neither of the sampling methods produce zero speed at the start and goal positions. To generate smoother end-effector paths with acceptable acceleration we sample in the second derivative of the configuration space (i.e.: joint acceleration space) and integrate to update the joint velocity and position.\nThe sampling process is modified as follows. First, the inputs are defined as 2-tuples of small changes in bonded angle\naccelerations (\ud835\udf031\u0308, \ud835\udf032\u0308 ). Then, generating joint velocities and joint angle by integration of the form:\n(\ud835\udf03?\u0307?)\ud835\udc5b\ud835\udc52\ud835\udc64 = (?\u0307?\ud835\udc56)\ud835\udc50\ud835\udc62\ud835\udc5f\ud835\udc5f\ud835\udc52\ud835\udc5b\ud835\udc61 + \u222b \ud835\udf03?\u0308? \ud835\udc51\ud835\udc61 (\ud835\udc56 = 1,2) (6) (\ud835\udf03\ud835\udc56)\ud835\udc5b\ud835\udc52\ud835\udc64 = (\ud835\udf03\ud835\udc56)\ud835\udc50\ud835\udc62\ud835\udc5f\ud835\udc5f\ud835\udc52\ud835\udc5b\ud835\udc61 + \u222b \ud835\udf03?\u0307? \ud835\udc51\ud835\udc61 (\ud835\udc56 = 1,2) (7)\nAfter the new joint velocities are computed, new joint angles can be similarly found. Finally, the end-effector velocity and coordinates in task space are computed using forward kinematics for 2-DOF manipulator,\n(\ud835\udc65 \ud835\udc66 ) = ( cos (\ud835\udf031) cos (\ud835\udf031 + \ud835\udf032) sin (\ud835\udf031) sin (\ud835\udf031 + \ud835\udf032) ) (\ud835\udc3f1 \ud835\udc3f2 ) (8)\nThe Jacobian of the 2DOF manipulator is,\n2391\nAuthorized licensed use limited to: Carleton University. Downloaded on June 17,2021 at 05:12:46 UTC from IEEE Xplore. Restrictions apply." + ] + }, + { + "image_filename": "designv11_71_0000852_012079-Figure1-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0000852_012079-Figure1-1.png", + "caption": "Figure 1. Mechanism 3D model (a) and its prototype (b).", + "texts": [ + " The purpose of this paper is to show how this method can be applied for a new mechanism, which structure can be defined as a parallel-serial one. The paper is organized as follows. First, the description of the mechanism is presented and the workspace analysis problem. Next, the different types of workspaces are calculated for the mechanism. Finally, the discussion section summarizes the results and gives several recommendations. The 3D model of the considered mechanism and its working prototype are presented in figure 1. This mechanism has a parallel-serial structure. The planar parallel mechanism with three degrees of freedom is placed on a plate 1 (figure 1, a) which can move along a coordinate axis Z, perpendicular to the plane of this planar mechanism. Though the parallel mechanism has three degrees of freedom, it has four equal kinematic chains, each with rotational joints 2 on both ends and a linear drive 3 between them. A redundant kinematic chain is added to improve mechanism rigidity and decrease the number of its singular configurations. One more movable link 5 is placed on a top platform 4 of the planar parallel mechanism. This link comprises an instrument and has a rotational degree of freedom which axis 6 lies in the plane of a parallel mechanism. Thus, the mechanism has five degrees of freedom in total. This mechanism is designed to perform operations of selective laser sintering, and its output link (instrument) is a laser beam unit. The prototype (figure 1, b) has an output link model represented by a steel shaft with the same weight and dimensions as a real instrument. ToPME-2019 IOP Conf. Series: Materials Science and Engineering 747 (2020) 012079 IOP Publishing doi:10.1088/1757-899X/747/1/012079 As the introduction said, a workspace is one of the most important characteristics of the mechanism. For the given mechanism, it is easy to find the workspace dimension along the Z direction, because it is fully defined by the limits of the corresponding linear guides" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0001208_s1560354720030065-Figure1-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0001208_s1560354720030065-Figure1-1.png", + "caption": "Fig. 1. Physical implementation of the Suslov problem.", + "texts": [ + " We think it would be interesting to consider a more general case of a reduced system and construct an extended variety of attractors. First, it will more clearly show the relationship between the Bilimovich and Suslov problems, and second, it will be possible to see the diverse structure of existing attractors. Next, we want to offer the reader a logical continuation of the research conducted in [15], and show the influence of viscous friction on the behavior of the system. 1. EQUATIONS OF MOTION 1.1. Equations of Motion in the General Case Consider the motion of a system consisting of several bodies (Fig. 1): \u2013 a rigid body with a fixed point that cannot rotate in the body-fixed direction e: (\u03c9,e) = 0, (1.1) where \u03c9 is the angular velocity; \u2013 n material points, each of which has a mass mi and moves inside the rigid body according to a given law ri(t). A detailed derivation of the equations of motion is presented in [15]. We introduce two coordinate systems to describe the motion of this rigid body: \u2013 a fixed coordinate system Oxyz centered at a fixed point O; \u2013 a moving coordinate system Ox1x2x3, chosen so that the vector e has coordinates e = (0, 0, 1) in the system Ox1x2x3" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0000987_0954406220915499-Figure2-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0000987_0954406220915499-Figure2-1.png", + "caption": "Figure 2. The schematic of intermediate steering transmission mechanism (r1 is the effective length of the first rocker; r2 is the effective length of intermediate rocker; r3 is the effective length of the second rocker; g1, g2, and g3 are the distances between the pivot of the rocker and the transition lever and the fixed end of the rocker arm; and ai (i\u00bc 1,2) is the distance between the hinge point of the ith axle knuckle arm and vertical rod and kingpin).", + "texts": [ + " The left trapezoid arm causes the right wheel to rotate around the kingpin through the tie rod and the right trapezoid arm. Hence, the steering of the right wheel is in coordination with that of the left wheel. At the same time, the steering knuckle arms of the second axle move through the first transition rod, intermediate rocker, second transition rod, second rocker, and posterior longitudinal rod. Hence, the steering of the wheels of the second axle is in coordination with that of the first axle. A schematic of the intermediate steering transmission mechanism is shown in Figure 2. The wheel alignment parameters considered in this paper are the caster angle a, the steering axis inclination angle , the camber angle c, and the toe-in angle . The schematic of the wheel alignment parameters is shown in Figure 3. The steering axle not only needs to ensure the vehicle has a stable steering function, but also needs to make the steering vehicle wheel have a self-righting effect to ensure the vehicle can travel stably in a straight line. The caster angle can ensure a stable self-righting performance under high-speed driving conditions; The steering axis inclination angle can ensure a stable self-righting performance under low speed and heavy load conditions" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0001349_j.promfg.2020.05.094-Figure2-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0001349_j.promfg.2020.05.094-Figure2-1.png", + "caption": "Fig. 2. Energy partitioning in DED", + "texts": [ + " When a high-density powder stream is irradiated by a laser beam, part of the energy of the laser beam is absorbed and reflected by the powders [28]. The leftover energy pass through the powder stream and hits the substrate. Meanwhile, the powders, carried with irradiated energy, are delivered into the molten pool [30]. Then, the substrate\u2019s temperature raises due to heat conduction, and at the same time, part of the energy is lost due to heat sinking effect, such as radiation and convection [31]. The energy distribution during DED is show in Figure 2. Lumped capacity modeling is a simple but practical method for solving heat transfer problems. In this model, the temperature of the solid body is considered a function of the time only, assuming the temperature is spatially independent and uniformly distributed. This model ignores the spatial variables, such as temperature gradients within the solid, and time is the only independent variable. The transient temperature response of the solid can be determined according to the energy balance law. Based on this, the energy absorbed by the substrate (Qs, J) can be calculated with Equation (1): s l rs f a rp radiation convectionQ Q Q Q Q Q Q Q\u03b7= \u2212 \u2212 + \u2212 \u2212 \u2212 (1) Where, Ql is the laser energy (J), Qrs is the energy reflected by the substrate (J), Qf is the latent heat energy (J), Qa is the energy absorbed by the powders (J), \u03b7 is the powder catchment efficiency, expressing the ratio of the part weight to the total mass of powder fed into the chamber, Qrs = (1- w\u03b2 )\u00d7 (Ql - Qa), w\u03b2 is the absorptivity of workpiece" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0000292_pee.2019.8923437-Figure2-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0000292_pee.2019.8923437-Figure2-1.png", + "caption": "Fig. 2. Steady state equivalent circuit of the BDFRM", + "texts": [ + " The synchronous speed of the BDFRM is \u03c9sync = \u03c9p/pr for \u03c9s = 0 by analogy to a classical 2pr pole synchronous turbo-machine with a DC supplied field winding. If \u03c9s>0, then the BDFRM operates in super-synchronous speed mode, and in sub-synchronous regime for \u03c9s<0 (i.e. with an opposite phase sequence of the secondary to the primary winding). The development of the BDFRM detailed space vector model can be found in [11]. The corresponding phasor diagram and the characteristic reference frames are depicted in Fig. 2. The respective voltage equations in an arbitrary frame rotating at \u03c9 have been derived in [2] and are reproduced here for convenience of analysis: r rr r p p pp p d v R i j dt \u03c9= + + \u03bb \u03bb , (5) ( )r rr r s s r ss s d v R i j dt \u03c9 \u03c9= + + \u2212 \u03bb \u03bb , (6) where the primary and secondary stator winding flux linkages can be represented in the following forms: r r rp psp p sL i L i \u2217= +\u03bb , (7) r r rs pss s pL i L i \u2217= +\u03bb . (8) In these equations, Lp is the 3-phase primary inductance, Ls is the secondary winding counterpart, Lps is the primary to secondary mutual inductance, Rp and Rs are the windings resistances, whereas '*' denotes the complex conjugates of the primary and secondary current vectors ipr and isr in the respective reference frames as illustrated in Fig. 2. Note that (5)-(8) are written in two different reference frames rotating at \u03c9 and \u03c9r-\u03c9, respectively. This is one of the most distinguishing features of the space vector theory developed for the BDFRM [2]. In keeping with this, note that the current isr * = isp is a referred secondary current to the primary reference frame (but modulated in a frequency, and not common turns ratio sense). Similarly, ipr * = ips is a magnetically coupled primary current rotating at \u03c9-\u03c9p in the secondary reference frame [2], [11]" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0003773_j.jweia.2015.01.001-Figure1-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0003773_j.jweia.2015.01.001-Figure1-1.png", + "caption": "Fig. 1. (a) Velocity diagram for a particle showing the wind speed U, particle velocity (u;w) and velocity of the air relative to the particle V. (b) Force diagram showing the drag force acting in the direction of the relative velocity and the weight force acting down.", + "texts": [ + " Compact debris is any object for which all length to width ratios are approximately one (as opposed to rod like debris which has one long and two short dimensions and plate like debris which has two long dimensions and one short dimension). Additionally, compact debris has a negligible lift coefficient and negligible rotational inertia. As such, the compact debris flight equations can be developed from the aerodynamic drag equation and the two dimensional equations of motion for a particle in a gravitational field. Consider a particle moving horizontally with velocity u and vertically with velocity wo0 (taking up as positive) in a steady uniform wind field of horizontal velocity U and zero vertical velocity (see Fig. 1a). The resulting drag force acts in the direction of the relative velocity while the weight force acts vertically downward (see Fig. 1b). Ignoring the buoyancy force acting on the particle then the resulting equations for the time variation of vertical and horizontal particle velocity are given by d2x dt2 \u00bc du dt \u00bc \u03c1CDA 2m \u00f0U u\u00de ffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi \u00f0U u\u00de2\u00few2 q \u00f02\u00de and d2z dt2 \u00bc dw dt \u00bc \u03c1CDA 2m w\u00f0 \u00de ffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi \u00f0U u\u00de2\u00few2 q g \u00f03\u00de where x and z are the horizontal and vertical coordinates, \u03c1 is the density of air, m is the mass of the particle, A is the cross sectional area of the particle (assumed constant), CD is a drag coefficient (assumed constant), and g is the gravitational acceleration constant" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0000990_cyber46603.2019.9066711-Figure3-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0000990_cyber46603.2019.9066711-Figure3-1.png", + "caption": "Fig. 3. Design of the bionic robotic fish", + "texts": [ + " When the excitation stimulus is less than the minimum excitation lowd , the oscillator stops oscillating. With the gradual enhancement of the excitation signal, the oscillation frequency increases and the amplitude increases gradually. The oscillator oscillates gradually. When the excitation signal exceeds the maximum value, the oscillator will not stop, and will continue to drive the bionic robot fish with the highest frequency and the highest amplitude. We designed a bionic robotic fish which was made in imitation of pike. It is shown in Fig. 3. The robotic fish has three joints in its posterior part of the body, two pectoral fins, and one caudal fin. There are total six servo motors to drive them. To control the robotic fish, the CPG model in (4) has been applied to regulate these six servo motors. To control the robotic fish shown in Fig. 3, we adopted the CPG model proposed in (4). A driving signal d was set for driving both sides of the robotic fish. It increased from 0 to 5 continuously. The output signals are shown in Fig. 4 where 1 2 0.3b b= = , the other parameters of the CPG model are shown in Table 1 [16, 17]. The neurons of the robotic fish joints and caudal fin oscillate sequentially. The frequency and amplitude of the neurons increase with the driving signal d . From joint 1 to joint 4, control signals for each joint servo motor are shown in Fig" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_71_0003381_icdcs.2015.46-Figure3-1.png", + "original_path": "designv11-71/openalex_figure/designv11_71_0003381_icdcs.2015.46-Figure3-1.png", + "caption": "Fig. 3. Sensing model of a camera sensor[26]", + "texts": [ + " In this section, we will show the problem studied in this paper at the end of this section. In this paper, a 3-D space is considered. Camera sensors are deployed to monitor a bounded region R, which is the target field (i.e., stadium). Camera sensors are fixed above region R and they can rotate, i.e., they can pan and tilt, as Fig.1 shows. Each sensor is quantified by its field of view (FoV) and typically, an FoV is represented by its angle of view \u03b8 and depth of view H . The depth of view presents us the area of the visual scene that is acceptably sharp [26]. As Fig. 3 shows, each sensor has a sensing range from Hmin to Hmax, corresponding to front-face and back-face of FoV respectively. And it has a field of view (FoV) of angle \u03b8, \u03b8h horizontally and \u03b8v vertically. For PTZ cameras, panning and tilting allow cameras to rotate around their axis to adjust FoV horizontally (pan) and vertically (tilt). PTZ cameras in this paper can pan (horizontal) freely, i.e., with an angle of 2\u03c0, and can tilt (vertical) with an angle of \u03c0. Thus, each camera can have a large FoV" + ], + "surrounding_texts": [] + } +] \ No newline at end of file