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With the increasing focus on energy efficiency for the green transition, there is a desire to identify and reduce all losses possible. For rotary machines losses occur as friction at the bearings, and therefore frictionless magnetic bearings can be an important upgrade to high-speed machinery. However, as proven by Samuel Earnshaw in 1842 magnetic bearings are inherently unstable. Typically, magnetic bearings are stable in the lift (often vertical) direction. This results in instability in the radial direction and therefore some kind of stabilizing force is required in this direction [1].&lt;br/&gt;For a magnetic bearing, the radial stiffness determines the active force needed to keep the bearing stable. Thus there is a compromise in designing a magnetic bearing which can both lift as high a load as possible and at the same time have a low and constant radial instability [2-5].&lt;br/&gt;Here we describe a new magnetic bearing aiming to fulfill these requirements [6]. The bearing consists of two nested conical axi-symmetric permanent magnet rings with quadratic cross-sections, as shown in Fig. 1. This configuration is a superposition between a radially magnetized bearing and an axially magnetized bearing. The rotor and stator are centered on the same axis, and when load is applied on the rotor bearing this will move along the z-axis. The stator has a magnetization direction that is along the inwards normal to its surface, while the rotor has a magnetization direction outwards normal to its surface. Thus, there is repulsion between the rings. The idea behind this design is that the magnetic forces will both have a component providing a lifting force along the z-axis, but also a component along the radial direction, which might allow the stiffness along this direction to be controlled. Such a design of tilted conical rings of permanents magnets can easily be made of cubes of permanent magnets placed in a ring.&lt;br/&gt;We consider a design where there is a minimum distance between rotor and stator of 0.2 mm and where the conical rings that make up both the rotor and the stator have quadratic cross-section side lengths of 3 mm. The radius of the center of the tilted ring in the stator is fixed at 24 mm, while it is varied for the rotor. We consider different tilt angles of the stator/rotor rings in the bearing and we also consider ring geometries with twice the cross-section and different magnetization directions. The “single” bearing consists of two tilted concentric rings with opposite magnetization as described above, the “double” bearing has two magnets in each tilted ring, and finally the “alternating” bearing has the magnetization of one part of the ring reversed in both the stator and rotor. All configurations are illustrated in Fig. 1. We assume that the permanent magnets in the bearing have a remanence of 1.4 T and a relative permeability of 1.&lt;br/&gt;We determine the force and stiffness of the bearing using a magnetostatic finite element (FEM) axi-symmetric model in Comsol Multiphysics. As the simulations are 2D axi-symmetric a very high mesh resolution can be used, eliminating the known convergence issues with bearing models in 3D [7]. We have conducted a variation study to determine the optimal geometry of the bearing as function of the tilt angle and the bearing type. Tilt angles from 0 to 90 degrees in steps of 15 degrees were modelled. Combined with the three bearing types studied, this gives a total of 21 bearings modelled. For each of these we use an adaptive routine to determine the optimal rotor radius to a precision of 0.1 mm. We show that the optimal value of the rotor radius is almost identical for the different bearing geometries considered, with decreasing rotor radius for increasing tilt angle.&lt;br/&gt;Following this we determine which bearing would be easiest to control, i.e. has a constant stiffness value, but which also provides a high lifting force. We investigate this by calculating the relative standard deviation for both the axial force and stiffness. This is shown in Fig. 2 for all studied bearings which are stable in the axial direction.  Choosing the best bearing depends on the desired characteristics of the bearing. If one desires a bearing that is easy to control, i.e. has a mostly constant stiffness, for a 3 mm axial displacement the double geometry with a tilt angle between 60-75 degrees is clearly the best.  When a smaller axial displacement of 1 mm is considered, all bearings with a tilt angle of 60 degrees display almost constant stiffness. This is a remarkable property, as such a bearing is very easy to control with an active bearing. In this bearing the exact load on the rotor is not of importance, as long as it is between the minimum and maximum lift, and at the same time the stiffness is constant regardless of the load. We thus show that by varying the tilt angle and the bearing geometry it is possible to tailor a bearing to the desired characteristics and achieve an easily controllable bearing with a high lifting force.

**References:**

[1] J. P. Yonnet, IEEE Transactions on Magnetics, Vol. 17, p.1169-1173 (1981)&lt;br/&gt;[2] J. G. Detoni, Proceedings of the Institution of Mechanical Engineers, Part C: Journal of Mechanical Engineering Science, Vol. 228, p.1829-1844 (2014)&lt;br/&gt;[3] J. Sun, D. Chen and Y. Ren. , IEEE Transactions on Instrumentation and Measurement, Vol. 62, p.2960-2965 (2013)&lt;br/&gt;[4] L. Zhang, H. Wu, P. Li, Y. Hu and C. Song, Mathematical Problems in Engineering, Vol. 2019, p.4265698 (2019)&lt;br/&gt;[5] K. K. Nielsen, C. R. H. Bahl, N. A. Dagnaes, I. F. Santos and R. Bjørk, IEEE Transactions on Magnetics, 10.1109/TMAG.2020.3042957 (2020)&lt;br/&gt;[6] R. Bjørk and C. R. H. Bahl, “A conical passive magnetic bearing with constant stiffness”, submitted for publication (2021)&lt;br/&gt;[7] N. A. Dagnæs-Hansen and I. F. Santos., Proceedings of the Institution of Mechanical Engineers, Part C: Journal of Mechanical Engineering Science, Vol. 233, p.5280-5293 (2019)

**Images:**

![](https://assets.underline.io/uploads/markdown_image/1/image/da90a1119a35e4bf7055dcd0b30f5edb.jpg)
Fig.1: The bearing geometry considered. The stator, shown in green, and the rotor, shown in grey, each consists of two tilted rings of permanent magnets with a quadratic cross-section. Their direction of magnetization is also indicated. On the right is shown a cross-section of the bearing, with the three types of magnet configurations considered. The bearing illustrated here has a tilt angle of 30 degrees.

![](https://assets.underline.io/uploads/markdown_image/1/image/015e67cee587135f544e04ac077b1ccd.jpg)
Fig.2: The relative standard deviation of the stiffness in the z-direction in percent as function of the relative standard deviation of the force in the z-direction in percent.

INTERMAG 2021

 A constant stiffness magnetic bearing

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This Conference provides an outstanding opportunity for worldwide participants to meet their colleagues and collaborators and discuss developments in all areas of magnetism research.

With the increasing focus on energy efficiency for the green transition, there is a desire to identify and reduce all losses possible. For rotary machines losses occur as friction at the bearings, and therefore frictionless magnetic bearings can be an important upgrade to high-speed machinery. However, as proven by Samuel Earnshaw in 1842 magnetic bearings are inherently unstable. Typically, magnetic bearings are stable in the lift (often vertical) direction. This results in instability in the radial direction and therefore some kind of stabilizing force is required in this direction [1]. For a magnetic bearing, the radial stiffness determines the active force needed to keep the bearing stable. Thus there is a compromise in designing a magnetic bearing which can both lift as high a load as possible and at the same time have a low and constant radial instability [2-5]. Here we describe a new magnetic bearing aiming to fulfill these requirements [6]. The bearing consists of two nested conical axi-symmetric permanent magnet rings with quadratic cross-sections, as shown in Fig. 1. This configuration is a superposition between a radially magnetized bearing and an axially magnetized bearing. The rotor and stator are centered on the same axis, and when load is applied on the rotor bearing this will move along the z-axis. The stator has a magnetization direction that is along the inwards normal to its surface, while the rotor has a magnetization direction outwards normal to its surface. Thus, there is repulsion between the rings. The idea behind this design is that the magnetic forces will both have a component providing a lifting force along the z-axis, but also a component along the radial direction, which might allow the stiffness along this direction to be controlled. Such a design of tilted conical rings of permanents magnets can easily be made of cubes of permanent magnets placed in a ring. We consider a design where there is a minimum distance between rotor and stator of 0.2 mm and where the conical rings that make up both the rotor and the stator have quadratic cross-section side lengths of 3 mm. The radius of the center of the tilted ring in the stator is fixed at 24 mm, while it is varied for the rotor. We consider different tilt angles of the stator/rotor rings in the bearing and we also consider ring geometries with twice the cross-section and different magnetization directions. The “single” bearing consists of two tilted concentric rings with opposite magnetization as described above, the “double” bearing has two magnets in each tilted ring, and finally the “alternating” bearing has the magnetization of one part of the ring reversed in both the stator and rotor. All configurations are illustrated in Fig. 1. We assume that the permanent magnets in the bearing have a remanence of 1.4 T and a relative permeability of 1. We determine the force and stiffness of the bearing using a magnetostatic finite element (FEM) axi-symmetric model in Comsol Multiphysics. As the simulations are 2D axi-symmetric a very high mesh resolution can be used, eliminating the known convergence issues with bearing models in 3D [7]. We have conducted a variation study to determine the optimal geometry of the bearing as function of the tilt angle and the bearing type. Tilt angles from 0 to 90 degrees in steps of 15 degrees were modelled. Combined with the three bearing types studied, this gives a total of 21 bearings modelled. For each of these we use an adaptive routine to determine the optimal rotor radius to a precision of 0.1 mm. We show that the optimal value of the rotor radius is almost identical for the different bearing geometries considered, with decreasing rotor radius for increasing tilt angle. Following this we determine which bearing would be easiest to control, i.e. has a constant stiffness value, but which also provides a high lifting force. We investigate this by calculating the relative standard deviation for both the axial force and stiffness. This is shown in Fig. 2 for all studied bearings which are stable in the axial direction. Choosing the best bearing depends on the desired characteristics of the bearing. If one desires a bearing that is easy to control, i.e. has a mostly constant stiffness, for a 3 mm axial displacement the double geometry with a tilt angle between 60-75 degrees is clearly the best. When a smaller axial displacement of 1 mm is considered, all bearings with a tilt angle of 60 degrees display almost constant stiffness. This is a remarkable property, as such a bearing is very easy to control with an active bearing. In this bearing the exact load on the rotor is not of importance, as long as it is between the minimum and maximum lift, and at the same time the stiffness is constant regardless of the load. We thus show that by varying the tilt angle and the bearing geometry it is possible to tailor a bearing to the desired characteristics and achieve an easily controllable bearing with a high lifting force.

**References:**

[1] J. P. Yonnet, IEEE Transactions on Magnetics, Vol. 17, p.1169-1173 (1981) [2] J. G. Detoni, Proceedings of the Institution of Mechanical Engineers, Part C: Journal of Mechanical Engineering Science, Vol. 228, p.1829-1844 (2014) [3] J. Sun, D. Chen and Y. Ren. , IEEE Transactions on Instrumentation and Measurement, Vol. 62, p.2960-2965 (2013) [4] L. Zhang, H. Wu, P. Li, Y. Hu and C. Song, Mathematical Problems in Engineering, Vol. 2019, p.4265698 (2019) [5] K. K. Nielsen, C. R. H. Bahl, N. A. Dagnaes, I. F. Santos and R. Bjørk, IEEE Transactions on Magnetics, 10.1109/TMAG.2020.3042957 (2020) [6] R. Bjørk and C. R. H. Bahl, “A conical passive magnetic bearing with constant stiffness”, submitted for publication (2021) [7] N. A. Dagnæs-Hansen and I. F. Santos., Proceedings of the Institution of Mechanical Engineers, Part C: Journal of Mechanical Engineering Science, Vol. 233, p.5280-5293 (2019)

**Images:**

![](https://assets.underline.io/uploads/markdown_image/1/image/da90a1119a35e4bf7055dcd0b30f5edb.jpg)
Fig.1: The bearing geometry considered. The stator, shown in green, and the rotor, shown in grey, each consists of two tilted rings of permanent magnets with a quadratic cross-section. Their direction of magnetization is also indicated. On the right is shown a cross-section of the bearing, with the three types of magnet configurations considered. The bearing illustrated here has a tilt angle of 30 degrees.

![](https://assets.underline.io/uploads/markdown_image/1/image/015e67cee587135f544e04ac077b1ccd.jpg)
Fig.2: The relative standard deviation of the stiffness in the z-direction in percent as function of the relative standard deviation of the force in the z-direction in percent.

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Conventional conveyance systems using wheels and belt conveyors occur friction and wear at the contact portions. In order to solve these problems, a noncontact conveyance system has been developed[1][2]. In this research, a permanent magnet repulsion levitation system and superconducting magnetic levitation system are adopted. The permanent magnet repulsion levitation system cause a large levitating force at the loading part. Superconducting magnetic levitation system cause a stable levitation at the carrier body. There is no influence of large load weight on the levitation of the carrier because the loading part and the carrier body are connected by the linear way. The carrier moves above the magnetic rail. The coils are installed on magnetic rail, and the magnetic field gradient is generated by the exciting coils. The carrier is propelled by the magnetic field gradient. By increasing the flux of the coil behind the HTS in the propulsion direction and reducing the flux of the HTS in the propulsion direction, magnetic field gradient is generated and the carrier propels. The coil pitch is optimized and propulsion force is examined by adding the exciting coils [3] [4] Fig.1 shows the experimental device[4]. The change of the magnetic field gradient is analyzed when changing the coil pitch by 5mm in the range of 25~50mm with the FEM analysis software JMAG-Designer[5]. The propulsion force is confirmed by the experimental results. The experimental condition is as follows; current of the coil I=2.0A and levitation gap g=23mm. The propulsion force is measured by the load cell installed in front of the carrier. From the analytical result, when the coil pitch is changed, the position of peak value of the flux density due to the magnetization and demagnetization on the magnetic rail changes, but the crest value of the magnetic field gradient is not changed. In the range of coil pitch 40~50mm, the excitation coils are separated. The flat section is formed between the positive peak value and the negative one in the magnetic gradient field. Thus, in the z-axis direction, the surface flux density of the HTS is the strongest at the center. Fig.2 (a) shows the average value and the maximum value of the propulsion force at each the coil pitch. At the coil pitch 50mm, both the average propulsion force and the maximum propulsion force are largest. As a result, the optimum value of the coil pitch under the HTS is 50mm. Center of the HTS keeps the largest flux from the permanent magnet. The coils are arranged on the magnetic rail at the coil pitch 50mm, as the diameter of the coil used in this research is 22mm, only the odd number coil in Fig.1 (b) will be installed. Thus, the excitation switching interval becomes large. Thus, as shown in Fig.1 (b), coils are installed at intervals of 25mm. At first, No.1 coil and No.3 coil in Fig.1 (b) are excited. Next, No.2 coil and No.4 coil are excited. In this way, the coils to be excited are switched, and coil pitch 50mm is realized. The case of exciting the coil between the two excited coils is considered.No.1 coil shown in Fig.1 (b) is magnetized, No.2 and No.3 coils are demagnetized. The flux density in the y-axis direction at that time is analyzed with the magnetic analysis software JMAG-Designer. In the range of x=-20~20, the flux density in the y-axis direction in three excitation coil case is smaller than that in two excitation coil case. As the magnetic rail is composed of Halbach array, the y-axis direction flux density is strong on the magnetic rail. Therefore, the change of the flux due to the excitation of the coil tends to occur in the y-axis direction. The propulsion force is compared when exciting two coils at intervals of 50mm with the propulsion force when three coils are excited at intervals of 25mm. No.1 coil in Fig.1 (b) is magnetized, No.2 and No.3 coils are demagnetized. The load weight up to 58.8N is placed on the carrier. Fig.2 (b) shows the maximum propulsion force for each load weight. Larger propulsion force is given when No.1, No.2 and No.3 coil in Fig.1 (b) are excited t at all load weight. This is because the range of the magnetic field gradient is increased by adding a coil to be excited. In the case of three coils excitation, as the change of flux density in the y-axis direction is large, it is necessary to consider the magnetic gradient field in the y-axis direction. In the case of two coils excitation, it is suitable to install the coils at intervals of 50mm so as to catch the center of the HTS. In addition, in the case of three coils excitation the propulsion force increases by generating a gradient in the flux density in the y-axis direction.

**References:**

[1]F. C. Moon, Superconducting levitation, John Wiley & Sons, New York, 1994. [2]J. Hull, M. Murakami, “Applications of bulk high-temperature Superconductors”, Proc. of IEEE, vol.92, 2004, pp1705-1718. [3]S. Ohashi, D. Dodo, “Influence of the Propulsion System on the Levitation Characteristics of the HTSC-Permanent Magnet Hybrid Magnetically Levitated System”, IEEE Transactions on Applied. Superconductivity. vol.17, no.2, 2007, pp2083-2086. [4]Y. Takaki, T. Sumida, S. Ohashi, “Improvement of Velocity Control in the Permanent Magnet-HTS Hybrid Magnetically Levitated Conveyance System”, IEEE Region 10 Conference (TENCON), 2016, pp3105-3108 [5] T. Sumida, Y. Kamitani, N. Yamada, S. Ohashi, “Propulsion Characteristics using Pinned flux of the HTS in the Permanent Magnet-HTS Hybrid Magnetically Levitated Conveyance System”, IEEE Transaction on Applied Superconductivity, 2016, Vol.26, No.4, 3600905.

**Images:**

![](https://assets.underline.io/uploads/markdown_image/1/image/b238f5b473b4485f39cd59d4d26df252.jpg)
Experimental device of the conveyance system

![](https://assets.underline.io/uploads/markdown_image/1/image/ca7ed8c3000c56c4b2c7fc0a1c756c36.jpg)
Propulsion force between HTS and the permanent magnet

 Optimization of the Excitation Method of the Propulsion Coils in the Permanent Magnet-HTS Hybrid Magnetically Levitated Conveyance System

1 Introduction Reciprocating engines generate torque only during the explosion process and this causes a large ripple in output torque. When the torque ripple is transmitted through the engine drive system, torsional vibration is induced in a drive shaft, resulting in uncomfortable vibration and noise. This rotational vibration is a problem not only in automobiles, but also in ships, cableways, and other transportation equipment. Various vibration suppression methods have been proposed to solve this problem. Matsuhisa et al. proposed a dynamic vibration absorber (DVA) with an additional mass that oscillates in the radial direction of the cableway to suppress rotational vibration caused by wind [1]. When the natural frequency of the added mass is set to about twice the natural frequency of the gondola, the Coriolis force generated by the radial vibration of the added mass has been shown to be effective in damping the rotational vibration. However, since the additional mass is passively vibrated by the centrifugal force of the gondola, itis not possible to control the magnitude of the Coriolis damping force. This paper proposes a new hybrid DVA to reduce rotational vibration of a rotor. The additional mass operates as the mover of the electromagnetic linear oscillatory actuator (LOA). The LOA actively oscillates the mover by current control and obtains the variable damping force caused by the Coriolis effect. LOAs have been applied to various industrial equipment due to their simple structure, direct drive, and high efficiency [2].They have also been applied to an active control device for automobiles [3]. In this paper, we describe the structure and mathematical model of the proposed hybrid DVA. Moreover, we evaluate numerically the vibration suppression effect under the active operation of the LOA. 2 Hybrid DVA Using Electromagnetic Oscillatory Actuator 2.1 Basic Structure and Operational Principle Fig. 1 shows the basic configuration and operational principle of the proposed hybrid DVA. 1-DOF LOA is embedded on the surface of a rotor and oscillates in the radial direction (r-direction). Excitation of the coil generates an electromagnetic force in the r-direction, which works to increase the mover stroke of the LOA. Consequently, the Coriolis force is produced in the circumferential direction (θ-direciton),which suppresses the rotational vibration in the θ-direction. Since the coil of the LOA rotates synchronously with the rotor, we need to discuss power supply. We assume a direct power supply from a battery mounted on the rotor surface. Therefore, we include the equivalent moments of inertia of the stator and the battery in the moment of inertia for the rotor. Our hybrid DVA can be also understood as a opposite phenomenon of a parametric excitation (parametric damper). Fig. 1(b) shows an example of the mover trajectories under parametric excitation and parametric damper. This figure clearly suggests that the vibration in the θ-direction increases after one oscillation period because of the positive energy provided by the parametric excitation. When the trajectory becomes reversed, the vibration decreases because of the negative energy provided by the parametric damper. The important point is that both cases require to satisfy the following relationship, ωr:ωθ = 2:1. (1) Here, ωr and ωθ are the natural frequency in the r- and θ-directions, respectively. 2.2 Equations of Motion for Hybrid DVA The Euler-Lagrange equation gives the following equations of motion in the r- and θ-directions, Jd2θ/dt2=m(l+r)d2θ/dt2+cθdθ/dt+2m(l+r)dθ/dt*dr/dt+kθθ=Tincosωt, (2) md2r/dt2+crdr/dt+krr=m(l+r)(dθ/dt)2+KtI+Kdr. (3) Here, the fourth term on the left-hand side of equation (2) is the Coriolis force, which works to decrease the vibration in the θ-direction. The first term on the right-hand side of equation (3) is the centrifugal force, which works to increase the vibration in the r-direction. The kinetic parameters of the hybrid DVA are determined as follows so that equation (1) holds, 2√(kθ/(J+ml2))=√((kr-Kd)/m). (4) 3 Reduction of Rotational Vibration under Active Control In this section, we solve the above equations of motion and verify the effectiveness of the proposed hybrid DVA through numerical simulation. Fig. 2 shows the rotational vibration of the rotor and the displacement of the mover with and without the LOA. In this simulation, The mover is oscillating passively until 4s. After that, Electromagnetic force generated by the LOA starts to oscillate the mover actively. Focusing on the envelopes of the waveform (blue solid line: without LOA, orange dashed line: with LOA), the active operation with LOA reduced the rotational vibration of the rotor by about 30%. In response to the decrease in the rotational vibration, the amplitude of the mover displacement is approximately doubled. We are able to conclude that the electromagnetic force generated by the LOA increases the displacement of the mover and the resulting damping effect by the Coriolis force also increases. In the final paper, we will describe the detailed design of the hybrid DVA and evaluate performances of the LOA through finite element analysis. Because this technique may also be applied to reduce torque ripple in AC motors, the rotational vibration by torque ripple will also be discussed.

**References:**

[1] L.D. Viet, N.D. Anh, H. Matsuhisa, Journal of Sound and Vibration, Vol. 330, no. 9, pp. 1904–1916 (2011) [2] M. Kato, Y. Kono, K. Hirata, and T. Yoshimoto, IEEE Transactions on Magnetics, Vol. 50, no. 11, 8206404 (2014) [3] F. Kitayama, K. Hirata, N. Niguchi, and M. Kobayashi, Sensors, Vol. 16, No. 377 (2016)

**Images:**

![](https://assets.underline.io/uploads/markdown_image/1/image/560279a19b259c411514e6dd264efe2c.jpg)
(a) Basic structure of the proposed hybrid DVA. (b) The trajectories of the mover under parametric excitation and damper.

![](https://assets.underline.io/uploads/markdown_image/1/image/f9bd317da50aeaf568c06373bc992f63.jpg)
Rotational vibration of the rotor and the displacement of the mover with and without the LOA.

 Reduction of Rotational Vibration Using Coriolis Force Generated by Electromagnetic Oscillatory Actuator Moving in Radial Direction

I Introduction A soliton is a wave that maintains its shape while it propagates at a constant velocity. The soliton exists in various nonlinear systems such as fiber optics, biology, magnetics, and electric circuit. In particular, Hirota proposed an LC ladder circuit, in which inductors and capacitors are connected like a ladder [1], [2]. Hirota experimentally proved the magnetic soliton was excited on the LC ladder circuit by using the capacitor whose capacitance is an exponential function of voltage. The observed magnetic soliton propagates along the LC ladder circuit at a constant velocity while maintaining the shape [3]. In other words, the magnetic soliton corresponds to a kind of rotating magnetic field. The rotating magnetic field is produced by a three-phase alternating current (AC) and is essential for AC motors. The current technology requires to use Pulse Width Modulation (PWM) inverters for generating the rotating magnetic field. On the other hand, the magnetic soliton can be generated by applying a pulsed voltage only at one point on the LC ladder circuit, which is expected to simplify the PWM inverter and its control algorithm [4]. This paper proposes an LC ladder circuit consisting of linear capacitors and nonlinear inductors. we also show the magnetic soliton can be excited if the inductance of the inductor becomes an exponential function of the current. However, the coil inductance wound on soft magnetic materials becomes symmetric with respect to the positive and negative currents, not exponential. To solve this problem, we propose a permanent magnet flux biased inductor (PMFB inductor) in which a permanent magnet is mounted into the magnetic circuit and the operating point of the iron core is biased. Finite element analysis (FEA) reveals that the inductance varies exponentially with the current. Finally, we solve numerically the differential equation of the LC ladder circuit consisting of the PMFB inductors and capacitors, and show the magnetic soliton propagates at a constant speed in the LC ladder circuit. II Permanent Magnet Flux Biased Inductor Fig. 1(a) shows the basic structure of the PMFB inductor. It consists of a stator core, a mover core, a coil, and a permanent magnet (PM). Its magnetic circuit is axisymmetric. An axially magnetized permanent magnet is embedded in the stator core. Here, we explain the reason why the inductance of the PMFB inductor changes exponentially as follows. The flux by PM mainly flows through the stator core as shown by the solid blue line. This leads to biasing the B-H curve of the stator core. When a positive current is applied, the flux by current strengthens the flux by PM. This results in magnetic saturation of the stator core and the inductance decreases with the current. When a negative current is applied, the flux by current weakens the flux by PM. This results in eliminating the magnetic saturation of the stator core and the inductance increases with the current. Therefore, exponential inductance characteristics is obtained when we design the shape of the PM, stator and mover cores property. Fig. 1(b) shows the analyzed relation between the flux and current obtained through FEA. The slope of this curve corresponds to the inductance. The solid line shows the bias flux provides asymmetric characteristics for positive and negative current. The dashed line is an approximation using the exponential function shown below, I = a(1-e-b(Φ-ΦPM)) (1) Here, a and b are the parameters and ΦPM is the bias flux by PM. The exponential function is a good approximation of the I-Φ property. This result implies the magnetic soliton can be excited on the LC ladder circuit consisting of the PMFB inductors and linear capacitors. III Numerical Results and Discussion In this chapter, we numerically solve the differential equation of the nonlinear LC ladder circuit mentioned in the previous chapter, and presents the successful excitation and propagation of the magnetic soliton. Kirchhoff’s voltage law gives the following differential equation, d2Φn / dt2=a/C(2e(-bΦn)-e(-bΦn-1)-e(-bΦn+1)) (2) Here, Φn is the flux through the nth inductor and C is the capacitance of the capacitor. The internal resistance of the PMFB inductor and capacitor is neglected. Note that this differential equation has a similar form to the one presented in [1] and [2]. Fig. 2 shows the numerical result of the differential equation in Equation (2). The number of LC units is 26 and the capacitance is 100µF. The start and end of the LC ladder circuit are connected and periodic. One magnetic soliton excited at 0s propagates at a constant speed in the positive and negative directions while maintaining its shape. Since the LC ladder circuit is periodic, the two magnetic solitons collide at a certain time. The waveform before and after the collision clearly shows the solitons are not affected by the collision. This is one of the most important properties of solitons. These results concludes magnetic solitons has a potential to be applied as a kind of the rotating magnetic field. In the final paper, we will describe the detailed design procedure for the PMFB inductor because this digest paper only presents the straight type PMFB inductor, not an annular one. Moreover, we will visualize and evaluate the rotating magnetic field produced by the magnetic soliton through FEA.

**References:**

[1] R. Hirota and K. Suzuki, Journal of the Physical Society of Japan., Vol. 28, pp. 1366–1367 (1970) [2] K. Muroya and S. Watanabe, Journal of the Physical Society of Japan., Vol. 50, pp. 2762–2769 (1981) [3] D.L. Sekulic, M.V. Sataric, M.B. Zivanov, and J.S. Bajic, Electronics and Electrical Engineering., Vol. 121, pp. 53–58 (2012) [4] M. Kato, S. Lee, and K. Hirata, IEEE Transactions on Magnetics., to be published, DOI:10.1109/TMAG.2020.3018118 (2020)

**Images:**

![](https://assets.underline.io/uploads/markdown_image/1/image/a078e0e6a099c620cb22dd40ed96d8d2.jpg)
(a) Basic structure of PMFB inductor. (b) Analyzed relation between flux linkage and current.

![](https://assets.underline.io/uploads/markdown_image/1/image/1980322477f343c6154555d623e5e3d0.jpg)
Numerical result of magnetic soliton excited on the nonlinear LC ladder circuit.

 Numerical Analysis of Magnetic Soliton Excited on Nonlinear LC Ladder Circuit Array Using Permanent Magnet Flux Biased Inductor

I. INTRODUCTION Environment recognition is a crucial task for autonomous mobile systems such as robots and drones, and vison information obtained by a camera is important. These systems require a high spatial awareness while undergoing various disturbances in an unknown environment. However, there are some problems: degradation of recognition accuracy caused by vibration and increase in size of the structure using a multi-degree-of-freedom (DOF) driving mechanism. Therefore, we focused on a multi-DOF actuator. Various types of multi-DOF actuators have been proposed, which use principles based on synchronous motors [1], stepping motors [2]. Nevertheless, the conventional actuators have a large structure and control device, for high torque density and continuous rotation. For applying to a driving source of a camera attitude control system, an actuator that can be configured with a few components and driven by a simple control device is required. Therefore, a multi-DOF voice coil actuator (DOF-VCA) is discussed in this paper, which has high responsiveness because it has only a few components. Kim et al. proposed a two-DOF-VCA with four-coils and permanent magnets [3,4]. This actuator can generate two-DOF motion driven by two H-bridge circuits. Moreover, our group proposed a three-DOF-VCA with eight-coils driven by three H-bridge circuits [5]. However, further simplification of a structure and control device is required for robotic eyes in humanoid robots. In this paper, we propose novel multiple-DOF-VCAs: two-DOF actuator driven by a three-phase half-bridge circuit and three-DOF actuator driven by a four-phase half-bridge circuit. The proposed actuators can be driven by a smaller number of switching elements and coils compared to the conventional actuators. The proposed magnetic structure and driving principle are described. The torque characteristics are clarified by a magnetic field analysis using a three-dimensional finite element method (3D-FEM), and it is shown that the proposed actuator can be driven by the proposed principle. II. PROPOSAL OF NOVEL STRUCTURE AND PRINCIPLE Figs. 1(a) and (b) show the proposed two-DOF-VCA driven by three-phase and three-DOF-VCA driven by four-phase, respectively. The mover consists of three-coils arranged at intervals of 120 degrees. The stator consists of an outer stator yoke and an inner stator with a yoke and permanent magnets. This actuator can be driven two-DOF by combined torques generated in each coil. The coils are connected at a neutral point, and the coils are excited in a balanced state. Thus, the actuator can be driven by only a three-phase half-bridge circuit. The proposed three-DOF-VCA has a structure that coils and permanent magnets are arranged at intervals of 90 degrees, and the mover consists of four-coils. This actuator can be driven by combined torques in each coil connected at a neutral point and can be operated by a four-phase half-bridge circuit. III. TORQUE CHARACTERISTICS The evaluated models are designed for verifying the proposed principle. The static torques are calculated by the 3-D FEM when the mover is forcibly rotated. The torque characteristics of the proposed VCAs are shown in Fig. 2. Fig. 2(a) shows the calculated torque of the proposed two-DOF-VCA when the current is applied to generate torque in the x-axis direction. The torque in the y- and z-axis directions are not generated, and the torque in the x-axis can be only obtained. The torque is maximum at the origin position and decreases by increasing the rotational angle. This is because the centers of poles in permanent magnets and coils are close together, and the magnetically attracted torque decreases. Fig. 2(b) shows the torque characteristics of the proposed three-DOF-VCA. From the analyzed results, it is found that the torque can be generated by the proposed principle. The decrease of the torque by rotation is also occurred due to the same reason in that of the two-DOF-VCA. The maximum torque in the z-axis is times compared with that of the x- and y-axis. This is because the permanent magnets are arranged with tilts of 45 degrees. From this result, it is suggested that the maximum torque in each axis direction can be adjusted by changing the tilt angle of the permanent magnets. IV. CONCLUSION The novel multi-DOF-VCAs were proposed in this paper, which can be driven by a small number of switching elements and components. The structure and principle were shown, and it was clarified that the actuators can generate torque, through 3-D FEM analysis.

**References:**

[1] K. Kahlen, I. Voss, C. Priebe, and R. De Doncker, "Torque Control of a Spherical Machine With Variable Pole Pitch", IEEE Transactions on Power Electronics, Vol. 19, No. 6, pp. 1628-1634, 2004. [2] G. S. Chirikjian and D. Stein, "Kinematic Design and Commutation of a Spherical Stepper Motor", IEEE/ASME Transactions on Mechatronics, Vol. 4, No. 4, pp. 342-353, 1999. [3] H. Kim, H. Kim, D.Ahn, and D. Gweon, "Design of a new type of spherical voice coil actuator", Sensors and Actuators A: Physical, Vol. 203, pp. 181-188, 2013. [4] H. Y. Kim, H, Kim, D.-G. Gweon, and J. Jeong, "Development of a Novel Spherical Actuator With Two Degrees of Freedom", IEEE Transactions on Mechatronics, Vol. 20, No. 2, pp.532-540, 2015. [5] A. Heya, K. Hirata, and N. Niguchi, "Dynamic Modeling and Control of Three-Degree-of-Freedom Electromagnetic Actuator for Image Stabilization", IEEE Transactions on Magnetics, Vol. 54, No. 11, 8207905, 2018.

**Images:**

![](https://assets.underline.io/uploads/markdown_image/1/image/a311d6e64e9682e948dfa950b8094323.jpg)
Fig. 1. Proposed multi-DOF-VCAs

![](https://assets.underline.io/uploads/markdown_image/1/image/c9809c9091f2c198de95fa3470c5986a.jpg)
Fig. 2. Analysis results

 Proposal of Novel Multiple-Degree-of-Freedom Voice Coil Actuator

I.INTRODUCTION Tubular flux-switching permanent magnet machine (TFSPMM) inherits the advantages of tubular machine and flux-switching machine with simple structure, strong robustness and no radial force[1]. However, it is not practical to make a TFSPMM entirely from axially stacked silicon steel sheet in machine manufacturing, because the flux will pass vertically through the silicon steel at the yoke of stator and mover and cause eddy current loss. Soft Magnetic Composite (SMC) material can solve the problem of three-dimensional permeability due to its magnetic isotropy[2]. However, TFSPMM with a magnetic core made only by SMC have low thrust due to low permeability and saturation magnetic density of the material. Moreover, core loss of the SMC is very high at low frequency. In order to change the above situation, hybrid magnetic core combined silicon steel sheet and SMC is proposed in this paper, which combines the high permeability of silicon steel and the magnetic isotropy of SMC. In hybrid magnetic core, the silicon steel sheet and SMC have a parallel and series relationship in the magnetic field, so the distribution of the two materials will affect the performance of the machine, including the core loss. Therefore, core loss reduction is important in hybrid magnetic core. II. MAIN CONTENT Fig. 1 shows the configuration of the TFSPMM with hybrid magnetic core and stacking direction of the silicon steel sheet. The silicon steel sheets in hybrid magnetic core are stacked along several tangential directions, and each part of the silicon steel is of the same stack width. The remaining space on the circumference, which cannot be stacked by silicon steel sheet, will be filled by SMC. Fig. 2 (a) shows the position of the point on the yoke when the silicon steel sheet stack width is 21mm. Point 1 is on the end of the silicon steel sheet and point 5 is on the middle silicon steel sheet. Wss is the stack width of the silicon steel sheet. As the position of the point is farther from the end, the amplitude of By decreases. By is the component of magnetic density in the stacking direction of silicon steel sheet. According to Faraday's law of electromagnetic induction, eddy current loss will be generated on the silicon steel sheet affected By. The larger the amplitude of By, the larger the eddy current loss will be. Therefore, it is considered to reduce the stack width and increase the number of blocks with appropriate stack width. Fig. 2 (b) shows the By of point 1 with different stack widths and different silicon steel sheet blocks. 4 parts 12mm means number of silicon steel blocks is 4 and the stack width is 12mm and 8 parts 12mm means number of silicon steel blocks is 8 and stack width is 12mm shown in Fig. 1 (a). As can be seen, with the reduction of stack width, the amplitude of Bydecreases. When the stack width is 12mm, the amplitude of By of 8 silicon steel sheets is smaller than that of 4 silicon steel sheets. Compared with the silicon steel sheet located in the radial direction, the larger the stack width is, the larger By is on the silicon steel sheet at the end. Fig. 2 (c) shows the core loss of TFSPMMs with different hybrid magnetic cores. The results show that the loss difference is not significant, because the material is not assigned electrical conductivity, and the software does not calculate the eddy current loss at the end. If the silicon steel is set as a solid in the calculation, the solid loss calculated by the finite element method is shown in Fig. 2 (d). In practice, large eddy currents are not generated in the axial section because there is an insulating layer between the silicon steel sheets. But in the software, lamination setting cannot simulate the insulation layer. Therefore, solid manual partitioning is needed to approximate the actual situation. III. CONCLUSION In this paper, core loss of TFSPMMs with hybrid magnetic core are compared and analyzed. The stacking direction and width of silicon steel sheet as well as the match with SMC are described. Although the silicon steel sheet and the flux path are parallel to each other, the magnetic flux is not strictly limited to the plane of the silicon steel sheet. The component of magnetic density in the stacking direction of silicon steel sheet varies with the stack width. As the stack width of the silicon steel sheet decrease, By of point 1 will decrease. Since By is the magnetic density component perpendicular to the silicon steel sheet, the smaller By is, the smaller the eddy current loss generated on the silicon steel sheet will be. When the number of silicon steel blocks is 4, the eddy current loss at the end decreases with the reduction of stack width, but the reduction of the proportion of silicon steel in the hybrid magnetic core will reduce the flux linkage and reduce the thrust. Therefore, when the stack width is 12mm, the number of silicon steel blocks is increased to 8, so as to increase the proportion of silicon steel in the hybrid magnetic core, ensuring small eddy current loss at the end and increased thrust.

**References:**

[1] R. Cao et al., “Quantitative comparison of linear flux-switching permanent magnet motor with linear induction motor for electromagnetic launch system,” IEEE Trans. Ind. Electron., vol. 65, no. 9, pp. 7569_7578, Sep. 2018. [2] C. Liu et al., “Performance evaluation of an axial flux claw pole machine with soft magnetic composite cores,” IEEE Trans. Appl. Supercond., vol. 28, no. 3, Apr. 2018, Art. no. 5202105.

**Images:**

![](https://assets.underline.io/uploads/markdown_image/1/image/9d355e97588dc96892a4bbf22b162f14.jpg)
Fig. 1 (a) Configuration of TFSPMM with hybrid magnetic core; (b) point 1 on stator yoke and stacking direction of silicon steel sheet.

![](https://assets.underline.io/uploads/markdown_image/1/image/e16de32d628da9877cbdc10f448ee469.jpg)
Fig. 2 (a) point on stator yoke; (b) By at point 1 under different stack widths; (c) core loss of TFSPMMs with different hybrid magnetic cores; (d) solid loss of TFSPMMs with different hybrid magnetic cores.

 Core Loss Reduction of Tubular Flux-Switching Permanent Magnet Machine with Hybrid Magnetic Core

Tubular linear permanent magnet machines (TLPM) have gained increasing popularity over the past years due to their versatile nature. Whether used as actuators for linear motion applications or as generators for energy production applications, these machines can provide accurate and precise motion with relatively compact structures and simpler control systems [1]-[3]. Like any other mechanical application, TLPMs might suffer from different structural deformations during either operation or assembly. The moving part of the machine called the translator is the most vulnerable to such mechanical deformations. One commonly occurring deformation happens at the level of the translator shaft and is usually called a translator eccentricity where the translator shaft is no longer aligned perfectly with the central axis of the machines's cylindrical structure. Eccentricity can take different forms and has been studied previously by some authors [4]-[7]. In [4] [5], eccentricity analysis was done using analytical and 3D FEM modeling and the calculated results were compared and validated by experimental results. In [6], eccentricity effects were studied for a surface mounted TLPM structure using the 3D finite element method (3D FEM). The study was continued in [7] by analyzing the eccentricity effects on another TLPM structure, also using 3D FEM, and comparing the eccentricity effects on the two structures. The authors of [6] [7] noted the extensive amount of time and effort required to compute some of the 3D FEM models and generate the results. Hence, an alternative modeling technique for eccentricity analysis that can provide relatively good results with less computation time is needed. The aim of this paper is to present a quasi-3D modeling technique that can represent the 3D problem using the superposition of multiple 2D models. It will be shown that the proposed model takes much less computation time and effort and produce relatively accurate results in comparison to the 3D model. Translator eccentricity unbalances the radial electromagnetic forces acting on the translator by forming a non-uniform air gap between it and the primary. Static eccentricity, which is a common eccentricity form, is the state of having the translator axis displaced by a constant distance ε from the primary axis while both axes remain parallel to each other. This work only focuses on the static eccentricity form leaving other forms for future works. The proposed quasi-3D model is achieved by dividing the 3D structure in the tangential direction (θ direction) into a certain number of sections (see Fig. 1). Since the air gap is non-uniform around the translator in the tangential direction, there is a different air gap value e (Fig. 1) for each θ along the tangential direction where θ∈[0, 2π]. Hence, each section in the tangential direction is modeled as a standalone 2D linear machine having an air gap equal to the air gap value at the central angle of the corresponding section. Then, for each 2D linear machine model, the magnetic induction in the air gap is calculated where it will have two components x and y corresponding to the axial (z-direction) and normal (r-direction) cylindrical components of the 3D model respectively. The tangential component can be associated with a z component which is impossible to determine in a 2D model that has only x and y directions, and hence it will be neglected. After gathering the results from all the 2D modeled sections, using the Maxwell stress tensor the electromagnetic forces acting on the translator in the axial and radial directions are calculated. Both the 2D and 3D models used are finite element method based models, and for this work, they were constructed using the FLUX commercial software. The quasi-3D model will be used to study two TLPM structures having static eccentricity. The first one is a surface-mounted PM structure, the second one is an inserted PM flux focusing structure, and the two have similar dimensions. Both open circuit and on load operations will be studied, and the two cases of linear and non-linear ferromagnetic materials will be considered as well. Studying the different cases will provide an insight into the accuracy and efficiency of the proposed quasi-3D model. Fig. 2 shows a comparison between the quantities (cogging force and attraction force) calculated using the quasi-3D and the 3D FEM models for the surface mounted TLPM having an eccentricity value ε equal to 0.75mm (75% of the air gap value). For this comparison, the ferromagnetic materials were defined in their linear case with a constant μr = 10000. Also, the relative permeability of the PM was defined as μr = 1 and its magnetic remanence as Br = 1T. Noting that the attraction force computed is acting along the axis aligned with the angles having the minimum and maximum air gap values. The results shown in Fig. 2 are relatively accurate and acceptable especially that the solution of 65 time steps of the quasi-3D model with 32 tangential divisions took only around 15 minutes with much less construction effort, while it took around 3 days to solve 33 time steps of the 3D FEM model with much more construction and computation efforts. Computation time differences will be even more noticeable and remarkable when solving non-linear models that will be presented and compared in the full paper release.

**References:**

[1] N. Bianchi, S. Bolognani, D. D. Corte and F. Tonel, "Tubular linear permanent magnet motors: an overall comparison," in IEEE Transactions on Industry Applications, vol. 39, no. 2, pp. 466-475 (2003) [2] J. Wang, W. Wang and K. Atallah, "A Linear Permanent-Magnet Motor for Active Vehicle Suspension," in IEEE Transactions on Vehicular Technology, vol. 60, no. 1, pp. 55-63 (2011) [3] L. Szabo, C. Oprea, I. Viorel and K. A. Biro, "Novel Permanent Magnet Tubular Linear Generator for Wave Energy Converters," IEEE International Electric Machines & Drives Conference, pp. 983-987 (2007) [4] F. Marignetti, P. Cancelliere, and M. Scarano, “Analytical formulation of the no-load magnetic field in IPM tubular machines with translator eccentricity,” in IECON 2006 - 32nd Annual Conference on IEEE Industrial Electronics, pp. 829–835 (2006) [5] L. Cappelli, Y. Coia, F. Marignetti and Z. Q. Zhu, "Analysis of Eccentricity in Permanent-Magnet Tubular Machines," in IEEE Transactions on Industrial Electronics, vol. 61, no. 5, pp. 2208-2216 (2014) [6] D. Sarr Lo, H. Lawali Ali, Y. Amara, G. Barakat, and F. Chabour, “3D Finite Element Analysis of Eccentricity in a Tubular Linear Permanent Magnet Machine,” in 2018 XIII International Conference on Electrical Machines (ICEM), Alexandroupoli, p. 1996–2001 (2018) [7] H. Diab, Y. Amara and G. Barakat, "Comparison of Two Tubular Linear Permanent Magnet Machines with Translator Eccentricity," 2020 International Conference on Electrical Machines (ICEM), Gothenburg, pp. 612-618 (2020)

**Images:**

![](https://assets.underline.io/uploads/markdown_image/1/image/2fedcffd641e7ac4c796e19d941f5391.jpg)
Fig.1 A 3D TLPM structure divided in the tangential direction with a 2D axial cross-section view with dimensions

![](https://assets.underline.io/uploads/markdown_image/1/image/094e73a6fefce9ab02b34bf22374e05b.jpg)
Fig.2 Comparison of the cogging and attraction forces generated by the quasi-3D and 3D FEM models

 Translator Eccentricity Analysis in Tubular Linear Machines using Quasi-3D Finite Element Method Modeling

I. INTRODUCTION Ferromagnetic materials have complex nonlinear hysteresis characteristics, which play a decisive role in simulating electrical equipment's magnetic properties, such as the power loss and the magnetic remanence of the core in transformers and motors[1]. Therefore, the accurate and efficient hysteresis model is fundamental to improving the power equipment's operational efficiency. As a fast and precise hysteresis model, the Preisach model is widely used in analyzing the magnetic properties of the electromagnetic device coupled with the FEM method. However, some problems remain to exist in the current Preisach model parameter identification methods, such as the convergence speed and the accuracy need to be further improved[2,3]. To develop a more efficient parameter identification method of the Preisach model, a hybrid algorithm combining conjugate gradient method (CGM) and velocity-controlled particle swarm optimization (VCPSO) algorithm is proposed in this paper: Firstly, a one-dimensional magnetic property test system is built to measure the hysteresis loop data of B30P150 steel sheet under quasi-static conditions. Secondly, the parameters of the Preisach model with closed-form Everett function are identified based on the proposed hybrid optimizing algorithm. Finally, in the full paper, the proposed method's accuracy is discussed by comparing the convergence speed and calculation accuracy of the proposed hybrid algorithm with other single intelligent algorithms. II. THE IMPLEMENT AND IDENTIFICATION OF PREISACH MODEL A. The Preisach model based on the closed-form Everett function The traditional Preisach model is implemented by the double integral of the distribution function in the Preisach plane, which is time-consuming and inefficient. To overcome the complex parameters identification process of the traditional Preisach model, the distribution function is simplified to the product of two single-valued functions, and a numerical Preisach model based on the closed-form Everett function is established. The improved model only needs a series of basic operations and a few parameters to be identified, which implement efficiency is much improved. B. Parameters identification of the Preisach model by combing the CGM and VCPSO algorithm A hybrid intelligent algorithm based on the conjugate gradient method (CGM) and a velocity-controlled particle swarm optimization (VCPSO) algorithm is proposed to overcome the low convergence speed and easily fall into a locally optimal solution. The identification process of a hybrid algorithm is shown in Fig 1. Firstly, according to the fast optimization of the VCPSO algorithm, finding the optimal solution range of parameters. Secondly, the searching range of VCPSO is taken as the initial parameter of the conjugate gradient method. Then the optimal solution is determined by local search. Finally, the optimal parameters are obtained by CGM. The measured and simulated hysteresis loop are compared after the identification process. The efficiency of the hybrid algorithm with other single optimal identification algorithms will be discussed in the full paper. III. RESULT AND DISCUSSION The simulated and measured hysteresis loops of the B30P150 silicon steel sheet are shown in Fig 2. It can be found that the model parameters identification results can well simulate the hysteresis loop of the B30P150 silicon steel sheet. The error is due to the approximate substitution of the Everett function, and increasing the number of relevant terms of the Everett function can improve the Model accuracy. What’s more, considering the reversible and irreversible magnetization, the simulation accuracy at the model's turning point can also be improved. IV. CONCLUSION In this paper, a hybrid optimization algorithm combining a velocity-controlled particle swarm optimization algorithm and the conjugate gradient method is proposed to identify the parameters of the Preisach model. The hybrid algorithm is an effective means for parameter identification of the Preisach model and contributes to the design and upgrading of electrical equipment.

**References:**

[1]Szabo Z, Fuezi J,Journal of magnetism and magnetic materials, 406(may):251-258(2016). [2] N. Duan et al., IEEE Trans. Magn. 51, 7300904 (2015) [3]Stoleriu L, Andrei P, Journal of Applied Physics, 103(7p.3):132-134(2018).

**Images:**

![](https://assets.underline.io/uploads/markdown_image/1/image/fb4bc67340641cc2bfad0149fde867f8.jpg)
Fig. 1 Identification process based on the hybrid algorithm of CGM and VCPSO

![](https://assets.underline.io/uploads/markdown_image/1/image/21705d8230b4c6fc54b147f747d1730c.jpg)
Fig. 2 Comparison between the measured and the simulated hysteresis loops

 Parameter identification of Preisach model based on the conjugate gradient method and velocity-controlled particle swarm optimization

I. Permanent magnet (PM) synchronous motors have become excellent alternatives for asynchronous motors as well as DC motors in numerous applications in recent years. Therefore, torque density and efficiency improvements in development of PM synchronous motors has become a significant issue. In this study, various PM synchronous motor alternatives are proposed with both non-oriented and grain-oriented silicon steel material in the same stator in order to increase the power density and the efficiency. Detailed performance comparison using 2D/3D FEA is completed and a feasible stator designs manufactured. The proposed design is also compared with a conventional PM motor with a non-oriented stator. As a result of the investigations, it is observed that the proposed motor with both materials in the stator have higher torque density and efficiency. II. INTRODUCTION In this study, detailed investigation on the stator design with both non-oriented and oriented electrical steel is conducted so as to increase the performance of conventional permanent magnet synchronous motors. Different motors with the same stator dimensions and the same rotor are examined with six different cases as in Fig.1:Case-1 (Entire stator made of non-oriented silicon steel), Case-2 (Entire stator made of grain oriented silicon steel), Case-3 (Entire stator made of grain oriented silicon steel with 1/3rd segments), Case-4 (Stator yoke made of non-oriented silicon steel and the conventional teeth made of grain-oriented steel), Case-5 (Stator yoke made of non-oriented silicon steel and the Screw shaped teeth made of grain-oriented steel),Case-6 (Stator yoke made of non-oriented silicon steel and the wedge shaped teeth made of grain-oriented steel). Detailed FEA simulations of all of the 6casesare performed and compared. Detailed performance tests of the promising prototype motors are performed. Torque output, losses and efficiencies of the motors are obtained and compared in order to illustrate the benefits of the proposed solution. III. FEA of PROPOSED SKEWED MODELS AND EXPERIMENTAL VERIFICATION Detailed FEA is performed for both proposed and reference PM motors. Since the flux density in stator teeth is higher than that of stator yoke, stator teeth are modeled and manufactured with grain-oriented material. The model structure of the two motors is shown in Fig 2. Detailed analyses including cogging variations, back EMF, losses, torque-speed curves and output torque of both motors are performed. Flux densities and flux lines obtained by both simulations are obtained. A prototype motor is manufactured and tested to show the promising benefits. It is seen that the torque and losses shows the potential performance benefits of the proposed motor. III. CONCLUSION In this paper, the design of innovative PM motors with two different materials in the stator is proposed. Promising alternatives are investigated and summarized. Detailed FEA results are presented. The motor with promising oriented silicon steel material is manufactured and tested. Detailed test results will be provided in the final version of the paper. Based on the initial results of the analyses, the output torque is increased by 1.88% and the stator tooth losses is decreased by 51.1%. The test results and comparisons will also be given in the last version of the paper.

**References:**

V. References: [1] Stefano Cicalé, Luciano Albini, Francesco Parasiliti, Marco Villani, Design of a Permanent Magnet Synchronous Motor with Grain Oriented Electrical Steel for Direct-Drive Elevators, IEEE, 2012, 978-1-4673-0142-8 [2] Nuwantha Fernando, Fuad HaninMagnetic, Materials for Electrical Machine Design and Future Research Directions: A Review, IEEE, 2017, 978-1-5090-4281-4 [3] Damian Kowal, Peter Sergeant, Luc Dupré, Alex Van den Bossche, Comparison of Nonoriented and Grain-Oriented Material in an Axial Flux Permanent-Magnet Machine, IEEE, 2010, 0018-9464 [4] Andreas Lindner, Ingo Hahn, Investigation of a large air-gap E-core flux-switching machine with arbitrary magnet shape and grain-oriented material, IEEE, 2016, 978-1-5090-2538-1

**Images:**
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 Design of Innovative Radial Flux Permanent Magnet Motor Alternatives with Non-Oriented and Grain-Oriented Electrical Steel for Servo Applications



Additive manufacturing technology carries its potential to develop high-performance electrical machines enabling direct printing of complex shapes and incorporating the complex sequence of feedstock materials in a single build to optimize the demanding physical parameters such as low weight, cost, energy loss, and high mechanical strength and corrosion resistance. We printed Fe-3%Si stator core using selective laser melting and machined it down to thin laminates to assemble a prototype brushless permanent magnet electric motor. The magnetic and electrical properties of the thin heat-treated laminates and back emf of the electrical motors are studied. The thin laminates exhibit maximum relative permeability of 7494 with an approximate 0.8 T applied field and core-loss of 0.473 watts/lb at 60 Hz. We will report in detail about the magnetic and mechanical properties of Fe3Si soft magnets and their influence on the electrical motor. The designed machine is analyzed using finite element analysis and compared with measured electrical parameters.

**References:**



**Images:**





 Additively manufactured Fe-3Si stator for high-performance electrical motor

This paper proposed a novel hybrid axial magnetic bearing (HAMB) for a vertical system, which can produce a unidirectional magnetic force. The principle of the new HAMB is analyzed. When the rotor is suspended in the equilibrium position, the bias force offsets the gravity of the rotor and the current is zero, which can reduce the copper loss. According to the magnetic field distribution, an equivalent magnetic circuit is established. Based on the analytical model, the curve of force-current, curve of force-displacement and bearing capacity are calculated. The results of analytical are consistent with the ones of the finite element method (FEM). Index Terms—Hybrid, axial magnetic bearing, magnetic circuit, finite element. I. Introduction Magnetic bearings (MBs) are widely used in high-speed rotating machinery, such as compressors, flywheel energy storages for the merits of contactless property and long life[1]. To reduce copper loss, hybrid MB has become the research feature combing the characteristics of active MB and passive MB [2]-[4]. Several axial MB are significant in vertical levitation rotor systems. A novel hybrid axial MB is proposed in [5], which reduces the reluctance of the control path and power loss of axial MB and through constructing the second air gap. In [6], an improved magnetic circuit model is introduced for a novel axial hybrid MB, which provides a reference for the MB design. In [7], an equivalent magnetic circuit with the conformal mapping method and Laplace’s equation is developed for axial MB with multiple air gaps. The results of analysis are consistent with the ones of finite element analysis. In this paper, a novel hybrid axial magnetic bearing (HRMB) is proposed for the vertical system, which produces a unidirectional electromagnetic force to replace the traditional bidirectional force. There is no need for the current to control the MB when the rotor is in the equilibrium position, which reduces the copper loss greatly. Then, the equivalent magnetic circuit method is adopted to analyze the performances of the HRMB. Ultimately, the accuracy of the analytical model is validated by the finite element method (FEM). II. Principle of Hybrid Axial Magnetic Bearing A. Principle of HAMB The structure of HAMB is shown in Fig.1. The blue solid line represents the bias flux, and the red dotted line represents the control flux. B. Magnetic circuit model The equivalent magnetic circuit models of the bias flux and control flux are established. III. Analysis of HRMB The variation of force with current and displacement can be obtained in Fig.2. The analytical results are in agreement with the results of FEM. The bias force and maximal force are calculated and compared with the finite element method. The margin of error is within 10%. IV. Conclusion A novel hybrid axial magnetic bearing is proposed in this paper and the principle of HAMB is analyzed. The structure has low copper loss when the rotor is stable in the suspension position. According to magnetic field distribution, an equivalent magnetic circuit is established. The curve of force-current, curve of force-displacement and bearing capacity are calculated by analytical and FEM respectively. The accuracy of the analytical model is confirmed by FEM.

**References:**

[1] J. X. Zhou, S. Q. Zheng, B. C. Han, et al. “Effects of Notch Filters on Imbalance Rejection With Heteropolar and Homopolar Magnetic Bearings in a 30-kW 60000-r/min Motor.” IEEE Trans. Ind. Electron., vol. 64, no. 10, pp. 8033-8041, Oct. 2017. [2] K. D. Bachovchin, J. F. Hoburg, and R. F. Post, “Magnetic fields and forces in permanent magnet levitated bearings,” IEEE Trans. Magn., vol. 48, no. 7, pp. 2112–2120, Jul. 2012. [3] W. Zhang and H. Zhu, “Improved model and experiment for AC-DC three-degree-of-freedom hybrid magnetic bearing,” IEEE Trans. Magn., vol. 49, no. 11, pp. 5554–5565, Nov. 2013. [4] B. C. Han, S. Q. Zheng, X. Wang, et al. “Integral design and analysis of passive magnetic bearing and active radial magnetic bearing for agile satellite application,” IEEE Trans. Magn., vol. 48, no. 6, pp. 1959–1966, Jun. 2012. [5] J. C. Fang, J. J. Sun, Y. L. Xu, et al. “A New Structure for Permanent-Magnet-Biased Axial Hybrid Magnetic Bearings.” IEEE Trans. Magn., vol. 45, no. 12, pp. 5319–5325, Dec. 2009. [6] X. X. Liu, J. Y. Dong, Y. Du, et al. “Design and Static Performance Analysis of a Novel Axial Hybrid Magnetic Bearing.” IEEE Trans. Magn., vol. 50, no. 11, pp. 1–4, Nov. 2014. [7] K. Wang, D. Wang, H. Y. Lin, et al. “Analytical Modeling of Permanent Magnet Biased Axial Magnetic Bearing With Multiple Air Gaps.” IEEE Trans. Magn., vol. 50, no. 11, pp. 1–4, Nov. 2014.

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