France



&amp;lt; I. INTRODUCTION &amp;gt;&lt;br/&gt;Recently, oscillatory actuators have been applied in various fields, including consumer electronics, entertainment, and industrial equipment. Especially, haptic devices that induce a virtual force are attracting attention. Haptic display systems using oscillatory actuators utilize the non-linearity of human perception, and it is possible to illusion a continuous unidirectional force by an asymmetric vibration [1]. Such oscillatory actuators are expected to become smaller and can be three-dimensionally (3-D) driven to display 3-D traction, for improving portability and functionality.&lt;br/&gt;Kato et al. developed a three-degree-of-freedom linear oscillatory actuator (3-DOFLOA) [2]. This actuator consists of four sets of coils has mutual interference of force for driving to each axis. Moreover, Heya et al. proposed a 3-DOFLOA that consists of three sets of coils and can be driven in three axes without interference [3]. However, the conventional 3-DOFLOAs need complex and large support mechanisms outside the magnetic circuit, which increase the size and weight of the device.&lt;br/&gt;Then, we propose a novel magnetic structure of a 3-DOFLOA in this paper. The magnetic pole and permanent magnet shapes are designed to integrate with the elastic support mechanism using mechanical springs. The proposed structure contributes to reduce the size and weight of the actuator. The force characteristics are clarified using a magnetic field analysis modeled with a 3-D finite element method (3-D FEM). The calculated results show that the proposed actuator can be independently driven for each axis.&lt;br/&gt;&lt;br/&gt;&amp;lt; II. PROPOSAL OF MAGNETIC STRUCTURE AND OPERATING PRINCIPLE &amp;gt;&lt;br/&gt;&amp;lt;&amp;lt; A. DESIGN &amp;gt;&amp;gt;&lt;br/&gt;Conceptual scheme of the magnetic structure is shown in Fig. 1(a). In the case of the conventional support mechanisms, mechanical springs are attached to the outside of the magnetic circuit sections. On the other hand, by adopting magnetic pole and permanent magnet shapes that allow the mover and stator to be connected by mechanical springs, it is possible to reduce the size without losing the function as a support mechanism. Fig. 1(b) shows the overview of the proposed a 3-DOFLOA. The inner mover has cylindrical permanent magnets for installing mechanical springs inside the stator. The stator consists of six sets of coils and the stator yoke. The coil is wound around the iron core with a hole of the same size as the permanent magnet. The mechanical springs made of non-magnetic material are used to support the mover. A total of six springs are arranged, two for each shaft, through the holes of the permanent magnet and the iron core. From this design, the proposed actuator can be integrated the support mechanism. The movable range is ± 0.5 mm to each axis.&lt;br/&gt;Fig. 1(c) illustrates the magnetic circuit. The magnetization direction of the permanent magnets arranged in the x-axis and y-axis directions is outward, and the magnetization direction of the permanent magnets arranged in the z-axis direction is inward. Therefore, the magnetic flux generated from the permanent magnets passes through the stator yoke and forms a closed magnetic circuit.&lt;br/&gt;&lt;br/&gt;&amp;lt;&amp;lt; B. OPERATING PRINCIPLE &amp;gt;&amp;gt;&lt;br/&gt;Fig. 1(d) shows the operating principle of the proposed actuator. When the Coil-X1 and Coil-X2 are excited, a magnetic pole is induced and the mover moves in the x-axis direction due to magnetic attraction and repulsion forces. Similarly, the mover can move in the y-axis and z-axis directions by the excitation of the Coil-Y1/2 and Coil-Z1/2.&lt;br/&gt;&lt;br/&gt;&amp;lt; III. STATIC FORCE ANALYSIS &amp;gt;&lt;br/&gt;The static force is calculated employing a magnetic field analysis using 3-D FEM. The inner yoke and stator yoke are made of SUY. The residual magnetic flux density of the permanent magnets is 1.3 T. The number of turns of each Coil-X1/2 and Coil-Y1/2 is 85 turns, and the number of turns of each Coil-Z1/2 is 68 turns. The input current is 0.3 A.&lt;br/&gt;The detent and excitation forces when the mover is moved in the x-axis, y-axis, and z-axis directions are shown in Fig. 2(a), (b), and (c), respectively. From Fig. 2(a), the magnetic spring constant in the x-axis direction due to the detent force is 2.10 N/mm, and the average excitation force in the x-axis is 0.23 N. From Fig. 2(b), the magnetic spring constant in the y-axis direction is 2.10 N/mm, and the average excitation force in the y-axis is 0.23 N. From Fig. 2(c), the magnetic spring constant in the z-axis direction is 2.25 N/mm, and the average excitation force in the z-axis is 0.22 N. The detent and excitation forces are generated only in the direction of the driving direction. In addition, the linear detent force characteristics are obtained. Therefore, it has to become a stable system at the origin by incorporating mechanical springs to cancel the detent force.&lt;br/&gt;From these analyzed results, it is confirmed that the proposed actuator can be independently driven in three axes, and the interference of forces between each axis is not observed.&lt;br/&gt;&lt;br/&gt;&amp;lt; IV. CONCLUSION &amp;gt;&lt;br/&gt;The novel small 3-DOFLOA was proposed in this paper. It can be driven three axes, and has not interference between each axis. The proposed actuator can contribute to the miniaturization of 3-D haptic devices.

**References:**

[1] T. Amemiya, and T. Maeda, Asymmetric oscillation distorts the perceived heaviness of handheld objects, &lt;i&gt;IEEE Trans. Haptics&lt;/i&gt;, vol. 1, no. 1, pp. 9–18, (2008)&lt;br/&gt;[2] M. Kato, and K. Hirata, Control of Three-Degree-of-Freedom Resonant Actuator Driven by Novel Vector Control, &lt;i&gt;TRANSPORTATION SYATEMS AND TECHNOLOGY&lt;/i&gt;, Vol. 4, No. 3, pp. 90-101, (2018)&lt;br/&gt;[3] A. Heya, K. Hirata, T. Matsushita, and Y. Kono, Design and Analysis of a Three-Degree-of-Freedom Linear Oscillatory Actuator, &lt;i&gt;IEEE Transactions on Magnetics&lt;/i&gt;, Vol. 56, No. 2, 6701404, (2020)

**Images:**

![](https://assets.underline.io/uploads/markdown_image/1/image/314c6e54fbcb7aa8d3feead54c1895b4.jpg)
Fig. 1. Proposed 3-DOFLOA (a) Conceptual scheme of the structure., (b) The overview of the proposed actuator., (c) Magnetic circuit., and (d) Operating principle.

![](https://assets.underline.io/uploads/markdown_image/1/image/db0a1cd9c36d06adb010609127b31c33.jpg)
Fig. 2. Analysis results of the static force (a) When the mover moves in the x-axis., (b) When the mover moves in the y-axis., and (c) When the mover moves in the z-axis.

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 Design and Analysis of a Three-Degree-of-Freedom Linear Oscillatory Actuator Integrated with Support Mechanism

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&lt; I. INTRODUCTION &gt; Recently, oscillatory actuators have been applied in various fields, including consumer electronics, entertainment, and industrial equipment. Especially, haptic devices that induce a virtual force are attracting attention. Haptic display systems using oscillatory actuators utilize the non-linearity of human perception, and it is possible to illusion a continuous unidirectional force by an asymmetric vibration [1]. Such oscillatory actuators are expected to become smaller and can be three-dimensionally (3-D) driven to display 3-D traction, for improving portability and functionality. Kato et al. developed a three-degree-of-freedom linear oscillatory actuator (3-DOFLOA) [2]. This actuator consists of four sets of coils has mutual interference of force for driving to each axis. Moreover, Heya et al. proposed a 3-DOFLOA that consists of three sets of coils and can be driven in three axes without interference [3]. However, the conventional 3-DOFLOAs need complex and large support mechanisms outside the magnetic circuit, which increase the size and weight of the device. Then, we propose a novel magnetic structure of a 3-DOFLOA in this paper. The magnetic pole and permanent magnet shapes are designed to integrate with the elastic support mechanism using mechanical springs. The proposed structure contributes to reduce the size and weight of the actuator. The force characteristics are clarified using a magnetic field analysis modeled with a 3-D finite element method (3-D FEM). The calculated results show that the proposed actuator can be independently driven for each axis. &lt; II. PROPOSAL OF MAGNETIC STRUCTURE AND OPERATING PRINCIPLE &gt; &lt;&lt; A. DESIGN &gt;&gt; Conceptual scheme of the magnetic structure is shown in Fig. 1(a). In the case of the conventional support mechanisms, mechanical springs are attached to the outside of the magnetic circuit sections. On the other hand, by adopting magnetic pole and permanent magnet shapes that allow the mover and stator to be connected by mechanical springs, it is possible to reduce the size without losing the function as a support mechanism. Fig. 1(b) shows the overview of the proposed a 3-DOFLOA. The inner mover has cylindrical permanent magnets for installing mechanical springs inside the stator. The stator consists of six sets of coils and the stator yoke. The coil is wound around the iron core with a hole of the same size as the permanent magnet. The mechanical springs made of non-magnetic material are used to support the mover. A total of six springs are arranged, two for each shaft, through the holes of the permanent magnet and the iron core. From this design, the proposed actuator can be integrated the support mechanism. The movable range is ± 0.5 mm to each axis. Fig. 1(c) illustrates the magnetic circuit. The magnetization direction of the permanent magnets arranged in the x-axis and y-axis directions is outward, and the magnetization direction of the permanent magnets arranged in the z-axis direction is inward. Therefore, the magnetic flux generated from the permanent magnets passes through the stator yoke and forms a closed magnetic circuit. &lt;&lt; B. OPERATING PRINCIPLE &gt;&gt; Fig. 1(d) shows the operating principle of the proposed actuator. When the Coil-X1 and Coil-X2 are excited, a magnetic pole is induced and the mover moves in the x-axis direction due to magnetic attraction and repulsion forces. Similarly, the mover can move in the y-axis and z-axis directions by the excitation of the Coil-Y1/2 and Coil-Z1/2. &lt; III. STATIC FORCE ANALYSIS &gt; The static force is calculated employing a magnetic field analysis using 3-D FEM. The inner yoke and stator yoke are made of SUY. The residual magnetic flux density of the permanent magnets is 1.3 T. The number of turns of each Coil-X1/2 and Coil-Y1/2 is 85 turns, and the number of turns of each Coil-Z1/2 is 68 turns. The input current is 0.3 A. The detent and excitation forces when the mover is moved in the x-axis, y-axis, and z-axis directions are shown in Fig. 2(a), (b), and (c), respectively. From Fig. 2(a), the magnetic spring constant in the x-axis direction due to the detent force is 2.10 N/mm, and the average excitation force in the x-axis is 0.23 N. From Fig. 2(b), the magnetic spring constant in the y-axis direction is 2.10 N/mm, and the average excitation force in the y-axis is 0.23 N. From Fig. 2(c), the magnetic spring constant in the z-axis direction is 2.25 N/mm, and the average excitation force in the z-axis is 0.22 N. The detent and excitation forces are generated only in the direction of the driving direction. In addition, the linear detent force characteristics are obtained. Therefore, it has to become a stable system at the origin by incorporating mechanical springs to cancel the detent force. From these analyzed results, it is confirmed that the proposed actuator can be independently driven in three axes, and the interference of forces between each axis is not observed. &lt; IV. CONCLUSION &gt; The novel small 3-DOFLOA was proposed in this paper. It can be driven three axes, and has not interference between each axis. The proposed actuator can contribute to the miniaturization of 3-D haptic devices.

**References:**

[1] T. Amemiya, and T. Maeda, Asymmetric oscillation distorts the perceived heaviness of handheld objects, IEEE Trans. Haptics, vol. 1, no. 1, pp. 9–18, (2008) [2] M. Kato, and K. Hirata, Control of Three-Degree-of-Freedom Resonant Actuator Driven by Novel Vector Control, TRANSPORTATION SYATEMS AND TECHNOLOGY, Vol. 4, No. 3, pp. 90-101, (2018) [3] A. Heya, K. Hirata, T. Matsushita, and Y. Kono, Design and Analysis of a Three-Degree-of-Freedom Linear Oscillatory Actuator, IEEE Transactions on Magnetics, Vol. 56, No. 2, 6701404, (2020)

**Images:**

![](https://assets.underline.io/uploads/markdown_image/1/image/314c6e54fbcb7aa8d3feead54c1895b4.jpg)
Fig. 1. Proposed 3-DOFLOA (a) Conceptual scheme of the structure., (b) The overview of the proposed actuator., (c) Magnetic circuit., and (d) Operating principle.

![](https://assets.underline.io/uploads/markdown_image/1/image/db0a1cd9c36d06adb010609127b31c33.jpg)
Fig. 2. Analysis results of the static force (a) When the mover moves in the x-axis., (b) When the mover moves in the y-axis., and (c) When the mover moves in the z-axis.

technical paper

I. Introduction Yokeless and segmented armature axial flux machine (YASA) is being considered for an advanced electric traction and propulsion in electric vehicles applications, since it has compact construction and high torque density. Compared with other axial flux machine, YASA could have higher winding fill factor and lower stator iron losses because of the absence of the stator yoke [1]. However, it is very difficult to design the stator cooling structure of YASA especially large power motor for electric vehicles as its stator is in the middle of two rotors. [2] gives a stator water-cooling structure for a 30KW YASA machine, in which fins are insert the space between stator modules with U-shape water-cooling pipe. But the manufacturing process is very complicated. Recently, heat pipe, as a small size but highly efficient passive cooling element, is attracting more and more attention [3]. Therefore, in this paper, a novel stator cooling structure combined flat heat pipe with housing water-cooling is presented. II. Stator Cooling Structure of YASA For an electric drive motor, the sealing is the primary consideration. As a result, open fan cooling can not be applied and the entire machine must be assembled in a housing. In addition, the weight of the motor should be as light as possible ensuring the mechanical strength. [4] introduces a stator mechanical construction of YASA with a excellent plastic material named polyetheretherketon (PEEK), which verifies the feasibility of stator support made by PEEK. Considering the low density, high thermostability, workability and nonelectroconductibility of PEEK, it is chosen as the material of stator carrier. In consideration of the fixation improvement of stator modules and the accuracy enhancement of rotor assembly, the ultimate cooling structure is shown in Fig. 1. It is obvious that the stator irons and concentrated windings are imbedded in two plates made of PEEK for full fixation. And the PEEK plates are bolted to the protruding middle part of the inner surface of the housing. A aluminum alloy ring is placed inside to secure the bearing to avoid unbalance air gap. What's more, several heat pipe is inserted in the gap in parallel between stator modules to enhance the heat transfer efficiency of stator to housing water-cooling. Considering the space limitation, the flat instead of round heat pipe is selected. And The thickness, height and length of flat heat pipe selected are 2mm (one third of the total gap width), 4mm and 85mm, respectively. For the reducing thermal contact resistance and improving mechanical strength, the remainder of stator support is impregnated with epoxy resin. The spiral type of housing cooling structure is adopted, because of its simple processing. III. CFD simulation of cooling structure Before simulation, the properties of epoxy resin and heat pipe should be identified. The selection principle of epoxy is high thermal conductivity, low thermal expansivity, low viscosity and high mechanical stability [5]. Finally, EP2000 from Lord is chosen after comparison. And the thermal conductivity of EP2000 is 1.9 W/mK. The heat pipe is the heat-transfer device which combines the principles of both thermal conductivity and phase transition to transfer heat between two solid interfaces effectively [6]. Its equivalent heat conductivity can reach hundreds of times more than copper, which is the metal with the best thermal properties. In simulation, a heat pipe is usually equivalent to a copper bar of high thermal conductivity [7]. But its performance should be determined in detail. Different size of heat pipes have been measured in laboratory under different conditions. And it is found that the performance of them depends on their size and the input power of heat source. With the increase of input power, the equivalent heat conductivity will get higher, and finally plateau. But the relationship between performance with size is still under study. In the most extreme case, the input power is 5W, the average temperature of a 100mm long, 2.5mm high and 0.8mm thick heat pipe is 308K. And the equivalent thermal conductivity is calculated as 14801W/mK. Therefore, the thermal conductivity of heat pipe in CFD is set as 10000W/mK tentatively. The stator temperature distribution of CFD simulation with heat pipes is shown in Fig. 2 (a), whereas the result without heat pipes is given in Fig. 2 (b). As is shown in Fig. 2, it is evident that the heat pipes accelerate the release of heat from the stator, and reduce the temperature of the stator significantly. The average temperature of stator reduces from 480K to 352K, which is 128K lower. IV. Conclusion In this paper, a novel stator cooling system for yokeless and segmented armature axial flux machine with flat heat pipe is developed. The proposed structure can increase the heat transfer efficiency of stator and improve the temperature of stator sufficiently. And the validity of this cooling structure is verified by CFD simulation. However, there are also many works to be done in the future. The thermal properties of heat pipes, especially the relationship between size parameters with thermal conductivity, should be identified. And the mechanical stability of above construction needs to be confirmed.

**References:**

[1] A. Gerlando, G. Foglia and M. Iacchetti, IEEE Transactions on Energy Conversion, Vol. 31, no. 2, pp. 750-758, 2016. [2] J. Chang, Y. Fan and J. Wu, IEEE Transactions on Industrial Electronics, early access. [3] J. Huang, S. Naini and R. Miller, IEEE Transactions on Vehicular Technology, vol. 68, no. 5, pp. 4467-4478, 2019. [4] B. Zhang, T. Seidler and R. Dierken, IEEE Transactions on Industrial Electronics, vol. 63, no. 4, pp. 2062-2071, 2016. [5] S. Nategh, A. Boglietti and D. Barber, IEEE Transactions on Energy Conversion, vol. 35, no. 2, pp. 1026-1035, 2020. [6] A. Faghri, Heat Pipe Science and Technology, Global Digital Press, 2016. [7] G. Fang, Y. Wei and Z. Yan, Applied Tthermal Engineering, vol. 152, pp. 594-604, 2019.

**Images:**

![](https://assets.underline.io/uploads/markdown_image/1/image/3a896721b5bca598cb368406564fbfc1.jpg)
Fig. 1 Stator cooling structure of YASA with flat heat pipe

![](https://assets.underline.io/uploads/markdown_image/1/image/15c0ece6138def21a245c72bc21caa06.jpg)
Fig. 2 Temperature distribution of stator (a) cooling without heat pipe (b) cooling with heat pipe

 A novel stator cooling construction for yokeless and segmented armature axial flux machine with heat pipe

Interior permanent magnet synchronous machine (IPMSM) is considered the best candidate for various applications such as traction and electric vehicle applications [1] [2]. In some cases, the IPMSM must have a high amount of torque density, power density, and wide speed range [3]. Recently, researchers have tried to increase the torque density of the vernier motor [4]. However, the vernier motor suffers a low power factor because of a high reluctance due to a large amount of leakage flux in the airgap [5] [6]. In this draft, a new dual stator interior permanent magnet synchronous motor (DS-IPMSM) is developed, comprised of 72 slots and 8 pole-pairs. In the proposed design, a new inner stator is inserted using the unused space of the rotor. The slots of inner stator are shifted by half pole pitch to attain the maximum amount of torque and to aid the flux lines of an inner and outer stator with each other. Although the torque of the proposed design increases; however, due to a new stator, the amount of losses increases. Hence, instead of silicon steel (50A1000), a soft magnetic amorphous material (2605SA1) is used in the outer and inner stator to reduce the proposed design losses. Fig. 1 (a) shows the configuration of the proposed designs. To reduce the inner winding temperature, a cup rotor is presented, as shown in Fig. 1 (b). It shows that the inner stator is attached to the inner stator shaft to support it mechanically. The rotor is attached to the shaft in the form of a cup. Moreover, the bearing is used between the inner stator support shaft and output rotor shaft to align the rotor. Because of cup shape rotor configuration, the inner windings' heat is removed from one side of the motor. The slot area of the inner and outer stators is not the same because of its geometry. To reduce the inner winding temperature, the inner stator's current density is less than the outer stator. Therefore, the proposed design uses two inverters, as shown in Fig. 1 (c). The proposed DS-IPMSM using conventional steel sheets are presented by PD-1; however, the design using the amorphous material is presented by PD-2. Fig. 2 shows the analysis and comparison of the proposed and conventional designs. Fig. 2 (a) shows the back EMF of the conventional and proposed designs. Additionally, the back EMF of outer stator PD-1 and PD-2 is identical with the back EMF of a conventional single stator model. Fig. 2 (b) shows the dramatic improvement in the torque compared to the conventional model. In the end, the losses of PD-1, PD-2, and conventional design (CD) are compared. It shows that although the torque performance of the PD-1 and PD-2 increased by a huge amount as compared to conventional modal. Contrary, the losses are also increased by 25 % due to a second stator in the conventional design. The total losses in the conventional design are 556.3 W, and the losses in the PD-1 are 753.2 W. The efficiency of the conventional design is 98.45 %, whereas it is decreased to 98.39% due to an increase in motor losses. To cope with the problem of increased losses, the amorphous material is introduced to the stator; thus, decreasing the overall losses of the motor up to 493.5 W, resulting in increased efficiency of 98.90 %.

**References:**

[1] Toba and T. A. J. I. t. o. i. a. Lipo, "Generic torque-maximizing design methodology of surface permanent-magnet vernier machine," vol. 36, no. 6, pp. 1539-1546, 2000. [2] F. Zhao, T. A. Lipo, and B. Kwon, "Magnet flux focusing design of double stator permanent magnet vernier machine," in Proc. COMPUMAG, 2013. [3] C. Liu, K. Chau, J. Zhong, W. Li, and F. J. I. T. o. a. s. Li, "Quantitative comparison of double-stator permanent magnet vernier machines with and without HTS bulks," vol. 22, no. 3, pp. 5202405-5202405, 2011. [4] Y. Wang, M. Cheng, M. Chen, Y. Du, and K. J. I. E. p. a. Chau, "Design of high-torque-density double-stator permanent magnet brushless motors," vol. 5, no. 3, pp. 317-323, 2011. [5] Y. Gao, R. Qu, D. Li, H. Fang, J. Li, and W. J. I. T. o. M. Kong, "A novel dual-stator Vernier permanent magnet machine," vol. 53, no. 11, pp. 1-5, 2017. [6] Y. Shi and T. J. I. T. o. M. Ching, "Power Factor Analysis of Dual-Stator Permanent Magnet Vernier Motor with Consideration on Turn-Number Assignment of Inner and Outer Stator Windings," 2020.

**Images:**

![](https://assets.underline.io/uploads/markdown_image/1/image/c7c840fe8ec575c6432af4af42c1dc3c.jpg)
Fig. 1. (a) Geometry of the proposed and conventional designs; (b) Proposed dual stator cup type rotor; (c) Configuration of the inner and outer stator with the inverter.

![](https://assets.underline.io/uploads/markdown_image/1/image/98f1f27479c5fc2af1c76c2798071cee.jpg)
Fig. 2 (a) Back EMF of the conventional and proposed designs; (b) torque comparison; and (c) losses comparison.

 An Improved Dual-Stator Cup Rotor Interior Permanent Magnet Synchronous Machine

1. Introduction The electrification of a vehicle has been gaining momentum recently, and more and more vehicles are being equipped with many electric motors and their inverters. In particular, a series–parallel hybrid electric vehicle and coaxial helicopter have two motors in its drive train [1-2]. In order to reduce the space of the motors, we proposed a two-controllable-rotor motor [3]. The proposed motor has two axially-arranged rotors and a common stator. The two rotors can be independently controlled using 3- and 6-phase superimposed currents created by a 6-phase inverter. One rotor is driven by a 3-phase current and the other is driven by a 6-phase current. In the transportation applications, it is important to continue the drive operation under fault conditions. Therefore, a fault tolerant control for multiphase motors have been studied [4-5]. In this paper, we propose a fault-tolerant control for a two-controllable-rotor motor using a 6-phase inverter, and its effectiveness is verified using finite element analysis. 2. Structure and Operational Principle Fig. 1 shows general and sectional views of our proposed motor with two axially-arranged rotors. The proposed motor has a common stator and two rotors: 6- and 12-pole rotors. The common stator has 36 slots and 6-phase distributed winding from A-phase to F-phase, and the in-phase coils are connected in series. This motor has a high heat dissipation because the stator is located outside. The two rotors can be easily supported because the two rotors are axially located with each other, and the 6-pole rotor shaft is mounted in the 12-pole rotor shaft through bearings. The 6- and 12-pole rotors are driven as 3- and 6-phase permanent magnet synchronous motors, respectively, by using superimposed currents created by a 6-phase inverter. Furthermore, it is known that the 3- and 6-phase currents do not mathematically interfere with each other [6]. In the proposed motor shown in Fig. 1, the stator and rotors diameters are Φ110 mm and Φ64 mm, respectively, the radial air gap length is 1 mm, and the rotor, stator thickness, and axial air gap length are 30 mm, 62 mm, and 2 mm, respectively. The winding factors of the 6- and 12-pole rotors are 0.7 and 1.0, respectively. The Maximum current density is 10Arms/mm2. 3. Fault-Tolerant Control A multiphase motor can be used for a fault-tolerant operation due to its additional degrees of freedom. This is because, under fault conditions, the remaining healthy phases in a multiphase machine can be used to compensate the faults and to continue the drive operation [4-5]. A two-controllable-rotor motor needs at least 5 phases to independently control the both rotors [7]. Therefore, a two-controllable-rotor motor using a 6-phase inverter can be driven although one phase is in a fault condition. In this paper, we propose a fault-tolerant control for a two-controllable-rotor motor when one phase is in an open circuit. In order to investigate the torque ripple due to a superimposed current when one phase is in an open circuit, an electromagnetic field analysis using finite element method is conducted. Figs. 2 (a) and (b) show the drive current waveforms for the 6- and 12- pole rotors under normal condition, respectively. Figs. 2 (c) and (d) show the drive current waveforms for the 6- and 12- pole rotors in A-phase open condition, respectively. Figs. 2 (e) and (f) show the torque waveforms of the 6- and 12- pole rotors, respectively under the normal and fault conditions. From Figs. 2 (e) and (f), the torque ripple rate of both rotor torque under the fault condition due to the superimposed current hardly changes. Therefore, the both rotors can be independently controlled under the fault condition. 4. Conclusion In this paper, we propose a fault-tolerant control for a two-controllable-rotor motor using a 6-phase inverter. The torque ripple rate of both rotor torque due to a superimposed current when one phase in an open circuit hardly changes compared with a normal condition. Therefore, the both rotors can be independently controlled under the fault condition. In the final paper, a mathematical model of the fault-tolerant control is described in detail. Furthermore, the torque ripple and N-T characteristics in normal and fault conditions are compared using a prototype.

**References:**

[1]H. A. Borhan, A. Vahidi, A. M. Phillips, American Control Conference, pp. 3970-3976 (2009) [2]J. Paulos, B. Caraher, M. Yim, IEEE ICRA, pp. 7011-7016 (2018) [3]H. Suzuki, K. Hirata, N. Niguchi, IEEE INTERMAG (2018) [4]S. Dwari, L. Parsa, IEEE Trans. Industrial Electronics, vol. 58, no. 2, pp. 476-485 (2011) [5]W. Zhang , X. Liang, F. Yu, IEEE Trans. Magnetics, vol. 54, no. 11 (2018) [6]N. Tan, M. Arimitsu, Y. Naruse, The International Power Electronics Conference, pp 1482-1488 (2005) [7]H. Suzuki, K. Hirata, N. Niguchi, IEEE ICEM, pp. 5-11 (2020)

**Images:**

![](https://assets.underline.io/uploads/markdown_image/1/image/9522b5cce994fe503fc306c95d036b41.jpg)
Fig. 1 Our proposed two-controllable-rotor motor.

![](https://assets.underline.io/uploads/markdown_image/1/image/5de7dc7b5fcf2c594b799e8516a4529e.jpg)
Fig. 2 Current and torque waveforms.

 Fault-Tolerant Control for a Six-Phase Two-Controllable-Rotor Motor

Offshore wind has gained traction worldwide as an alternative to fossil-fuel based power plants. Coastal US states have started incentivizing the development of offshore wind farms, with at least 24,135 MW currently planned for deployment. To make offshore wind more cost-competitive, major turbine OEMs like GE, Siemens, and MHI Vestas have been steadily increasing the power produced per turbine from 1.6 MW in 2000 to 12 MW in 2020. Increasing the turbine power reduces balance-of-plant costs, increases electromagnetic efficiency, and boosts capacity factor, leading to lower levelized cost of electricity (LCOE) for power purchasers [1]. However, even at 12 MW per turbine, the cost of shipping, installing, and maintaining thousands of wind turbines makes the price of electricity prohibitively expensive. Traditional permanent-magnet turbine technologies make it pointless to increase the power rating beyond 15 MW since there size and weight are significantly increases their installation cost. Low temperature superconducting (SC) generators were explored for a long time for high-power application where volume and wight requirements are strict [2]-[3]. Recent technology development in lowe temperature and high temperature SC cables renewed SC machines interest in wind turbine applications [3]-[4]. Achieving ambitious goals of harvesting 24,135 MW of offshore wind energy hinges on the successful development of superconducting generator technology. Superconducting generators can produce extremely high specific power and run at higher efficiency than commercial, permanent-magnet generators because they can produce up to 10X more magnetic flux density which reduces the weight significantly. This shows that with 10X power per turbine is possible at same PM generator specification. To demonstrate these benefits, a 12 MW superconducting wind turbine has been investigated for 3 different generator topologies: fully active, hybrid and fully passive. 12 MW power was chosen instead of the theoretical limit of 120MW to shows that partial SC machines can outperform the best commercial 12 MW turbines while maintaining a 10T flux densities which have already been demonstrated in MRI applications. Low temperature NB3Sn [2] cables and high temperature VIPER [3] cables are analyzed for field windings with Litz wire armature windings. A multi objective optimization is performed on machine designs to achieve less than 60 Ton machine weight while reaching greater than 98% efficiency. Two studies are conducted at outer diameter (OD) of 4.2 m and 7.2 m to analyze the reduction in LCOE due to transportation and installation cost. Fig. 1 shows the obtained Pareto-front for motor active weight Vs motor efficiency at 7.2 m OD. Results shows that targeted weight and efficiency can be achieved at ultra-high-field designs. The best designs have 10-12 T peak field at field windings and 3-5 T peak field at airgap. The optimization performed here equally weighted the mass and efficiency of the generator to reach the final design. However, LCOE is sensitive to superconductor length than the mass and efficiency-related cost penalties. This will be considered to further refine design optimization study presented here. The thermal design studies will also be refined using temperature-dependent thermal conductivities ranging all the way from room temperature to the proposed 4K operating temperature. Along the same lines, there is an outstanding need to identify cryocoolers which can reach 4K while satisfying the heat load requirements from ambient conduction and radiation. Upon completion of the electromagnetic, thermal, and mechanical design, the weight estimates for the generator will be refined to include both the generator and its supporting systems for comparison with the weight of commercial turbines. This is an important research study which also translates to a concrete value proposition for the feasibility of SC offshore wind turbines. An integration scheme was explored by mounting the proposed generator inside of a commercial turbine with the associated aerodynamic, mechanical, electrical, and thermal systems needed for operation in an offshore wind farm. Fig. 2 shows the cross-section and full turbine structure and the detail winding arrangements of a selected 7.2 m OD design. The proposed turbine would output 200% of the rated power for a commercial turbine of the same size. As it can be seen in the figure, the volume of the turbine nacelle is not fully occupied, leaving additional room to increase the turbine power even further with larger-diameter SC machines. In conclusion, this paper will produce electromagnetic, thermal, mechanical and cost analysis to support the feasibility of using ultra-high-field superconducting generators for offshore wind applications.

**References:**

[1]. A. B. Abrahamsen et al., IEEE Transactions on Applied Superconductivity, vol. 28, no. 4, (2018). [2]. S. Minnich et al., IEEE Transactions on Magnetics, vol. 15, no. 1, (1979). [3]. H. Southall and C. Oberly, IEEE Transactions on Magnetics, vol. 15, no. 1, (1979). [4]. M. Hsieh, C. Lin, I. Lin, IEEE Transactions on Magnetics, vol. 49, no. 5, (2013). [7]. D. Liu, H. Polinder, A. B. Abrahamsen, IEEE Transactions on Applied Superconductivity, vol. 27, no. 5, (2017). [5]. A. Godeke, B. ten Haken, H. H. J. ten Kate, Supercond. Sci. Technol., vol. 19, no. 10, (2006). [6]. Z S Hartwig, R. F Vieira, B. N Sorbom, Supercond. Sci. Technol., Volume 33, no. 11, (2020).

**Images:**

![](https://assets.underline.io/uploads/markdown_image/1/image/73e82ab6cb09198c881fd3fc117abb5a.jpg)
Fig. 1 Optimal parteo-front of weight Vs efficency

![](https://assets.underline.io/uploads/markdown_image/1/image/9c13a68e8f53ee1f479b897746d39ef5.jpg)
Detailed mechanical design of proposed SC machine inside a commerciall turbine and detail winding arrangement

 Ultra High-Field, High-Efficiency Superconducting Machines for Offshore Wind Turbines

This digest introduces a novel BL-WRSM having a unique two-layer stator winding configuration that surpasses the performance of the BL-WRSMs presented in [1] - [5]. Moreover, the new two-layer stator winding is such that it utilizes all the slot space available, and it can produce fundamental MMF components for torque generation and sub-harmonic MMF components for field excitation. As it can be seen from Fig. 1, two sets of series-connected stator windings, ABC and XYZ, are wound in a two-layer arrangement. Although the bottom layer’s slots are entirely covered by the ABC winding, the top layer consists of both ABC and XYZ winding, where every other slot has XYZ winding conductors in the whole periphery of the top layer, which is shown in Fig.2. By this arrangement, a single current source inverter can drive the machine with a maximum number of slots utilization. It is important to state that the rotor of the machine is also comprised of two separate windings itself and they are: the field winding and the harmonic winding. While the field winding is wound in an 8-pole manner, the rotor harmonic winding is a 2-pole winding that is employed for inducing the sub-harmonic MMF generated by the XYZ winding of the stator. Both the rotor windings are connected via a diode bridge rectifier circuit mounted on the periphery of the rotor. To extend the understanding, the fundamental MMF component is produced by winding ABC, and it has an 8-pole arrangement with the frequency of ωm, where ωm is the frequency in the mechanical space domain. Whereas in the case of winding XYZ with its 2-pole configuration, it delivers the sub-harmonic MMF component with the frequency of (ωm)/4. Thus, the two components rotate at a different speed. In the rotor, on the other hand, the harmonic winding is placed on the rotor such that it aligns magnetically with the stator's XYZ winding, and both are wound in a 2-pole manner. Adjacent to the harmonic winding, which is connected through the diode rectifier, field winding having an 8-pole arrangement aligns magnetically with winding ABC having 8 poles. Given all that, the novel BL-WRSM achieves its brushless operation by inducing the sub-harmonic MMF to the rotor harmonic winding and by magnetically locking the stator fundamental MMF with the rotor field winding. The main manuscript of this topic will cover an extensive literature review, structure of the machine, winding configuration, and working principle of the proposed BL-WRSM. Along with these, a 2-D finite element analysis (2-D FEA) has been performed to test the validity of the machine, which is also partially reported in this digest. An 8-pole, 48-slot BL-WRSM with a 3-phase, 2-layer distributed stator winding is designed to validate the proposed topology. Based on the stator winding structure, both ABC and XYZ windings are placed throughout the 360-mechanical degree of the stator, and the rated power of the machine is considered 1281 W having a constant speed of 900 rpm. To examine the effectiveness of the proposed machine, its performance is compared with the conventional WRSM used in [3] and the single inverter BL-WRSM in [4]. For the proposed machine, however, all design parameters are kept exactly similar to the BL-WRSM presented in [4], which includes, outer diameters, air-gap length, stack length, the current density in the stator, and rotor winding to name but a few. While all the machines are supplied with an RMS current of 4.5 A, the rotor field current is found in the proposed BL-WRSM (21.61 A) is remarkably higher than that of the BL-WRSM (9.8 A) in [4]. The reason for the higher current is 2-pole sub-harmonic component and the proposed 2-layer stator winding arrangement. As a result, the 2-D FEA shows that the average torque of the new machine is found 12.03 Nm that is approximately 65% more than the average torque production of the machine in [4]. Despite the performance enhancement in torque generation, the torque ripple is found 34%. Moreover, the flux density and flux lines are somewhat evenly distributed throughout the stator rotor frame for the novel BL-WRSM, unlike in the case of BL-WRSM in [4], where the distribution was found uneven because of the under-utilized stator slot space. It also shows the proposed BL-WRSM works under the flux density of 2 T. The focal point of this work is to develop a brushless wound rotor synchronous machine. To study the performance and feasibility the new machine has substantiated with the 2-D finite element method. Although operation wise it resembles the BL-WRSMs, performance-wise it surpasses many crucial aspects, and it demonstrates some certain advantages over the newly developed BL-WRSMs in [2] to [5]. Hence, the proposed machine could be an alternative solution considering all the attractive features provided, and they are listed as follows, 1. Single inverter brushless operation. 2. Utilizes the maximum allowable slot capacity. 3. No need for implying complex current control to mitigate unbalanced current circulation in the stator. 4. No need for imposing a complex unbalanced-cooling system. 5. Greater torque generation. With all these advantages, it can certainly be a competitive alternative not only to the BL-WRSMs but also with the state of the art PMSMs those are used in contemporary electrified transportation application. Moreover, the proposed topology could potentially adopt the design presented in [6] to [9], and used in converter topologies presented in [10], [11]. These designs will extensively be covered in the main manuscript.

**References:**

1. F. Yao, D. Sun, L. Sun, and T. A. Lipo, ‘‘Dual third-harmonic-current excitation principle of a brushless synchronous machine based on double three-phase armature windings,’’ in Proc. 22nd Int. Conf. Electr. Mach. Syst. (ICEMS), Harbin, China, Aug. 2019, pp. 1–4. 2. Q. Ali, T. A. Lipo and B. Kwon, "Design and Analysis of a Novel Brushless Wound Rotor Synchronous Machine," in IEEE Transactions on Magnetics, vol. 51, no. 11, pp. 1-4, Nov. 2015, Art no. 8109804. 3. Ali, Q., Atiq, S., Lipo, T. A., & Kwon, B. (2016). PM Assisted, Brushless Wound Rotor Synchronous Machine. Journal of Magnetics, 21(3), 399-404. 4. Hussain, A., & Kwon, B. (2017). A new brushless wound rotor synchronous machine using a special stator winding arrangement. Electrical Engineering, 100(3), 1797-1804. 5. Hussain, A., Atiq, S., & Kwon, B. (2018). Optimal Design and Experimental Verification of Wound Rotor Synchronous Machine Using Subharmonic Excitation for Brushless Operation. Energies, 11(3), 554. 6. M. Ayub, S. Atiq, Q. Ali, A. Hussain and B. Kwon, "Dual-Mode Wound Rotor Synchronous Machine for Variable Speed Applications," in IEEE Access, vol. 8, pp. 115812-115822, 2020 7. G. Jawad, Q. Ali, T. A. Lipo and B. Kwon, "Novel Brushless Wound Rotor Synchronous Machine With Zero-Sequence Third-Harmonic Field Excitation," in IEEE Transactions on Magnetics, vol. 52, no. 7, pp. 1-4, July 2016, Art no. 8106104. 8. M. Ayub, A. Hussain, G. Jawad and B. Kwon, "Brushless Operation of a Wound-Field Synchronous Machine Using a Novel Winding Scheme," in IEEE Transactions on Magnetics, vol. 55, no. 6, pp. 1-4, June 2019, Art no. 8201104. 9. A. Hussain, S. Atiq and B. Kwon, "Consequent-Pole Hybrid Brushless Wound-Rotor Synchronous Machine," in IEEE Transactions on Magnetics, vol. 54, no. 11, pp. 1-5, Nov. 2018, Art no. 8206205. 10. Rafin, S. M., Lipo, T. A., & Kwon, B. (2014). A novel topology for a voltage source inverter with reduced transistor count and utilizing naturally commutated thyristors with simple commutation. 2014 International Symposium on Power Electronics, Electrical Drives, Automation and Motion. doi:10.1109/speedam.2014.6872057 11. Rafin, S. M., Lipo, T. A., & Kwon, B. (2015). Performance analysis of the three transistor voltage source inverter using different PWM techniques. 2015 9th International Conference on Power Electronics and ECCE Asia (ICPE-ECCE Asia).

**Images:**

![](https://assets.underline.io/uploads/markdown_image/1/image/c19b2bcf2e6a20f6d7c1978211333363.jpg)
Fig.1

![](https://assets.underline.io/uploads/markdown_image/1/image/592f59ac8e917d6f80aaf42a3d4fabcd.jpg)
Fig.2

 A sub-harmonically excited novel brushless wound rotor synchronous machine using a new two-layer winding topology.

I. Introduction. With the merits of broad speed range and high efficiency at high-speed region, variable flux machine (VFM) is gaining more and more attentions in the wide-speed applications, such as electric vehicle, railway traction and numerical control machine tool. With the high coercive-force (HCF) and low coercive-force (LCF) permanent magnets (PMs) both utilized, VFMs with hybrid PMs are widely investigated. According to the combination way of hybrid PMs, VFM can be divided into series scheme and parallel scheme. Parallel VFMs are proven with relatively small demagnetization current and wide speed-adjusting range. But the HCF PM of parallel VFM has demagnetization effect on LCF PM, which will reduce the stability of LCF PM. While the LCF PM of series VFM is relatively stable, but the demagnetization current of series scheme is always large. This digest proposes a novel series-parallel VFM, which improves the stability of LCF PM by applying segmented PMs. The structure of the machine is introduced firstly, and then the magnetization-regulation principle is analyzed. The influences of V-shape pole parameters on electromagnetic torque and speed adjusting range of the machine are investigated. II. Machine Structure and Magnetization-Regulation Principle The structure of the proposed series-parallel VFM with segmented PMs is shown in Fig.1, which is evolved from traditional V-shape PM synchronous machine (PMSM). The magnetic pole of the machine is composed of NdFeB and AlNiCo, and each side of the V-shape PM is divided into three parallel segments along the long-side direction of PM slot. The inner and outer PM segments are composed of series NdFeB and AlNiCo, and the middle PM segment is pure AlNiCo. Compared with the traditional parallel VFM, the proposed VFM has higher stability at load state, because of the NdFeB PM placed at the outer part of V-shape pole, where is more likely to be demagnetized by q-axis armature reaction. In addition, the series segments composed of NdFeB and AlNiCo has weaker demagnetization effect on middle AlNiCo than pure NdFeB, which can also improve the stability of the magnetic pole. The no-load flux distributions of the proposed VFM at forward and reverse magnetization state are shown in Fig. 1. It is illustrated that the flux provided by inner and outer PM segments are shorted by middle AlNiCo at reverse magnetization state, and the fundamental amplitude of no-load back electromotive force (EMF) is decreased from 73.00V to 5.94V, which proves the wide speed-adjusting range of the proposed VFM. III. Research on Electromagnetic Torque and Speed-Adjusting Range Electromagnetic torque and speed-adjusting range are important characteristics of VFM, and thus influence laws of V-shape pole parameters on them are investigated. Since the change of PM dimensions will influence the magnetization result of middle AlNiCo when the same forward magnetization current is applied, the magnetization states of middle AlNiCo in VFMs with different PM dimensions at fully forward magnetization state are different. Therefore, the influences of V-shape pole parameters are different from traditional interior PMSM. According to the relationship between NdFeB thickness and the average flux density of series segments at no-load state, it can come to the conclusion that the equivalent remanence of the series segment can be adjusted by changing NdFeB thickness with the series-segment thickness unchanged. Afterwards, the influence of middle-AlNiCo positon on electromagnetic torque and back EMF is investigated. It is found that moving the middle AlNiCo closer to the air gap can decrease the total harmonic distortion (THD) of back EMF at reverse magnetization state, and almost has no impact on electromagnetic torque and back EMF at forward magnetization state. The influence of PM ratio is analyzed, and it is indicated that VFM with larger PM ratio has wider speed-adjusting range, but the average torque is relatively low. With PM ratio kept unchanged, the average torque increases firstly, and then decreases with the increase of PM-pole width. Moreover, the difference between d- and q-axis inductances is found having the same variation law with the average torque. The reason is that changing PM-pole width will change the distance between poles, which will change the d- and q-axis magnetic circuits distribution of the machine. When the thickness of the PM pole is increased, the average torque is increased firstly, and then decreased, while the torque ripple is increased continuously. The efficiency maps of the proposed VFM at forward and reverse magnetization states are shown in Fig. 2. It is illustrated that when VFM is reversely magnetized, the maximum speed of the machine is increased, and the efficiency at high-speed region is increased. The maximum torque of the machine at high speed region is also increased. IV. Conclusions A novel series-parallel VFM with segmented hybrid PMs is proposed, and the selection principles for PM parameters of the machine is proposed. The speed-adjusting ability of the machine is elaborated by analyzing the magnetization-regulation progress, and the speed-adjusting performance is revealed by efficiency maps at different magnetization states.

**References:**

[1] H. Yang, H. Zheng, and H. Lin, in Proc. IEMDC, CA, USA, pp. 2259-2264 (2019) [2] M. Wang, C. Tong, and P. Zheng, IEEE Trans. Magn., vol. 55, no. 7, pp. 1-7 (2019) [3] G. Qiao, M. Wang, and F. Liu, IEEE Trans. Magn., doi: 10.1109/TMAG.2020.3013624 (2020)

**Images:**

![](https://assets.underline.io/uploads/markdown_image/1/image/65d9adf6d6d4c391b17a14c0637170b0.jpg)
Fig. 1. Structure of the proposed VFM and no-load flux distributions at forward and reverse magnetization states.

![](https://assets.underline.io/uploads/markdown_image/1/image/ac4bb1f3709e11f0f381ec109389cfa3.jpg)
Fig. 2. Efficiency maps of the proposed VFM at forward and reverse magnetization states.

 Design and Analysis of a Novel Series-Parallel Variable Flux Machine with Segmented Hybrid Permanent Magnets

A synchronous machine can have both mutual torque due to the permanent magnet (PM) excitation and reluctance torque due to the rotor saliency. According to the ratio of these two torque components, the machine can be classified as (i) the salient permanent magnet synchronous machine (PMSM) in which the mutual torque is dominant or (ii) the permanent magnet-assisted synchronous reluctance machine (PMA-SynRM) in which the reluctance torque is dominant. Both categories can be called permanent magnet synchronous reluctance machines (PMSynRM). Of course, the PMSM with no rotor saliency and the SynRM with no PM assistance can be regarded as special PMSynRM which have no reluctance torque or mutual torque, respectively. In the salient PMSM, there are many interior permanent magnet (IPM) rotor structures, such as the V-shaped and the Delta-shaped structures. The magnets not only produce magnetic excitation but also act as flux barriers to make the rotor magnetically salient. Basically, the Delta-shaped structure brings better rotor saliency than the V-shaped structure, as it has more layers of flux barriers. On the other hand, in the PMA-SynRM, there are even more layers (usually 3~5 layers) of flux barriers to further enhance the rotor saliency, but the magnets inserted in these flux barriers have inefficient utility due to the severe leakage in the circumferential bridges of the rotor and the series structure of all the magnets in each pole. In the conventional PMSynRM, the rotor is identical in the axial direction, but in the radial direction, from the rotor center to the airgap, there are iron core, flux barriers and magnets which are all merged, therefore, this kind of rotor can be called radially merged rotor structure (RMRS). A characteristic of the RMRS is that the PM excitation is along the d-axis of the rotor, whilst the rotor saliency is symmetric about the d-axis as well as the q-axis. In the RMRS machines, referring to Fig. 1(a), the current vector should be located along the q-axis to maximize the mutual torque per ampere, and ideally located in the middle between the q-axis and d-axis to maximize the reluctance torque per ampere. In other words, it is not achievable to maximize both torque components simultaneously, and thus the current vector has to be located at a compromised phase angle (see Fig. 1(a)). In this paper, the d-axis is re-defined as that about which the rotor saliency is symmetric, and the q-axis is away from the d-axis by 90 electric degrees. If the PM excitation vector is shifted from the conventional position (i.e., the d-axis) towards the q-axis by 45 electric degrees, as shown in Fig. 1(b), the back EMF vector will then be located in the middle between the q-axis and d-axis. In this case, the current vector can be simply placed at the same phase angle as the EMF vector, so that both the mutual torque and the reluctance torque can be maximized simultaneously. Fig. 1. Vectors in permanent magnet synchronous reluctance machines. (a) for conventional RMRS machines, (b) for proposed ACRS machines. An asymmetric rotor structure with extra flux barriers is proposed in [1], where the flux barriers force the PM flux to shift from the original d-axis towards the q-axis. However, it is very difficult to make the shifting angle as large as 45 electric degrees. Therefore, the current vector still has to be located at a compromised phase angle. In order to better shift the PM flux vector, it is proposed to use two types of rotor segments [2], one is the conventional surface-mounted permanent magnet (SPM) rotor, and the other is the conventional SynRM rotor without permanent magnet assistance. As shown in Fig. 2, these two types of rotor segments are axially combined on the shaft. Hence, this is called axially combined rotor structure (ACRS). Clearly, any shifting angle of the PM flux vector from the d-axis can be achieved. Fig. 2. Schematic of axially combined rotor structure (ACRS). In the full presentation, pros and cons of the ACRS machines will be detailed. For example, when using the same stator, the SPM rotor and the SynRM rotor can be optimized individually. Moreover, the SPM rotor has less leakage than the IPM rotor, hence the permanent magnets can be utilized more efficiently. And, the SynRM rotor contains no permanent magnet, so that the circumferential bridges suffer from less stress due to the centrifugal force, thus the bridges can be thinner and the rotor saliency can be better. Furthermore, in the full presentation, different criteria will be described to determine the shifting angle and the axial length ratio of the two rotor segments, so as to maximize the two torque components simultaneously as well as the machine torque density, or to enhance the power factor for increasing the voltage utilization, or to make the torque ripple on the two rotor segments cancel each other and consequently reduce the overall torque ripple. Moreover, each of the SPM and SynRM rotors can be divided into several segments, so that the axial combination will be more flexible though the 3D effect will be severer.

**References:**

[1] W. Zhao, et al, IEEE Trans. Mag., Vol. 50, p.8104504 (2014). [2] J.X.Shen, et al, Chinese Patent No.202020519211.7.

**Images:**

![](https://assets.underline.io/uploads/markdown_image/1/image/5bd4d96288e9d536a0fbdbabf0e993b7.jpg)
Fig. 1. Vectors in permanent magnet synchronous reluctance machines. (a) for conventional RMRS machines, (b) for proposed ACRS machines.

![](https://assets.underline.io/uploads/markdown_image/1/image/ea2d35a1aff9207903d24a604a746a36.jpg)
Fig. 2. Schematic of axially combined rotor structure (ACRS).

 Permanent Magnet Synchronous Reluctance Machines with Axially Combined Rotor Structure INVITED

INTRODUCTION: Recently an increasing demand for high-speed permanent-magnet synchronous motor (HSPMSM) emerges in the high-power density, high-efficient and limited-installation space applications, such as electric vehicles and aircraft [1,2]. However, high-power density inevitably results in high-loss density and limited heat dissipation surface, which will pose great difficulty in motor cooling. Besides, a high-strength sleeve is usually necessary to guarantee the mechanical strength of the surface-mounted HSPMSM rotor. But additional loss will be generated in the conductive sleeve, further increasing the difficulties on cooling. The overheating of magnets, which will cause the irreversible demagnetization, is regarded as one of the biggest problems in the HSPMSMs [3]. The slotless HSPMSM was thought has less rotor eddy-current loss, since without high-frequency space harmonics of the magnetic field in the airgap caused by stator slotting [4]. Nevertheless, it's well known that the flux density of the airgap is usually lower. But the power density is closely related with it. So as to counteract the shortage, higher electrical load of the winding can be regarded as a feasible approach, but which in turn increases the copper loss. Both the efficiency and heating of HSPMSM are determined by the losses. And the losses mainly consists of copper loss, iron loss, rotor eddy-current loss and friction loss, all of which are closely correlates with dimensions [5, 6]. Therefore, it is essential to find the optimal design to achieve the competitive power density and minimum losses in the limited space for the slotless HSPMSM. This paper introduces a slotless toroidal-winding HSPMSM with non-magnetic fillers in stator for fuel cell air compressor application. A multi-objective optimization process was discussed and carried out to find the best compromise between power density and efficiency for a 15 kW, 120 000 rpm HSPMSM. Based on optimization results, the optimal design scheme was selected and compared with a slotted scheme. BASIC TOPOLOGY AND PRELIMINARY DESIGN CHOICES: The configuration of the proposed HSPMSM was shown in Fig. 1. Slotless configuration was chosen in order to reduce the electric losses in the sleeve and magnet. The toroidal winding was adopted. Some slots are machined on the outside edge of stator core and the part of toroidal windings outside of the stator core are embedded in these slots. So the leakage flux will increase in the outside slots and the inductance of windings are accordingly increased. Moreover, the toroidal windings is tightly surround the stator core, so the fixation of slotless windings is very convenient. Litz wires was employed to alleviate the eddy-current losses of windings, due to skin effect and proximity effect. So as to achieve the good insulation between different phases of windings, some non-magnetic insulating fillers are desired to be mounted between them. The inside edge of stator core are machined with some notches, and the fillers are inserted to them with epoxy resin glue. After impregnation processing, the stator becomes a strong overall structure. The fillers are also benefical to withstand the tangential electromagnetic force of the winding, consequently protecting the insulation of magnet wires. In addition, the fillers can help transfer the heat of windings to iron lamination. Boron nitride ceramics was selected as the candidate of non-magnetic fillers material, because of the good electrical insulation and thermal conductivity. MULTI-OBJECTIVE OPTIMISATION: Power density and efficiency were selected as the objective functions to be maximized. The key design variables and constraint variables were described and discussed in detail, as well as the multi-objective modeling method based on genetic algorithm and finite-element analysis of electromagnetic field. OPTIMISATION RESULTS AND DISCUSSION: In order to achieve the optimum design for the studied HSPMSM, 2246 evaluations has been performed and about 15 offspring populations was generated until the optimization converged. The optimization results were provided in Fig. 2, where all the feasible designs explored are represented by colored points. Besides, the solid black line is the Pareto front. The best designs are those that located on the Pareto front or closest to it. It can be clearly seen that the two optimization objectives are conflicting, which means that an increase in power density is generally associated with a reduction in efficiency, and vice versa. A large number of design schemes were generated in the optimization process. Weighing a tradeoff between power density and efficiency, the final scheme was selected and the candidate point was also indicated in Fig. 2. As a comparison, a common slotted-configuration HSPMSM was also designed. The performance of the final scheme was analyzed and compared with the slotted scheme. The axil length of windings for the slotless HSPMSM is 24.8% lower than the slotted one, which is benefical to improve the critical speeds of the rotor. The self-inductance calculated was higher than the slotted scheme. Besides, the slotless scheme has higher copper loss. But thanks to the removement of stator teeth, the iron loss was significantly decreased. So the total stator loss are basically the same for the two scheme. Finally, the slotless schemes exhibits higher efficiency and power density than the slotted scheme.

**References:**

[1] D. Hong, B. Woo, and J. Lee, and D. Koo, IEEE Trans on Magn., vol. 48, no. 2, pp. 871-874 (2012). [2] N. Uzhegov, A. Smirnov, and C. H. Park, IEEE Trans. Ind. Electron., vol. 64, no. 11, pp. 8427-8436 (2017). [3]D. Gerada, A. Mebarki, and N. L. Brown, IEEE Trans. Ind. Electron., vol. 61, no. 6, pp. 2946-2959 (2014). [4] F. Luise, A. Tessarolo, and F. Agnolet, 2014 International Symposium on Power Electronics, 18-20 June 2014, pp.1237-1244. [5] G. H. Jang, J. H. Ahn, and B. O. Kim, IEEE Trans on Magn., vol. 54, no. 11, pp. 1-6, (2018). [6] W. D. Zhao, X. J. Wang, and C. Gerada, IEEE Trans on Magn., vol. 54, pp. 1-5(2016).

**Images:**

![](https://assets.underline.io/uploads/markdown_image/1/image/7247dcf6fb939140744ef7646c10419b.jpg)
Configuration of the proposed HSPMM

![](https://assets.underline.io/uploads/markdown_image/1/image/85da787ea993b78ea704d13d3d8363da.jpg)
Fig.2 Optimization results

 Design and Optimization of a Slotless High-speed Permanent-Magnet Synchronous Motor with Non-magnetic Fillers in Stator

High-speed permanent magnet synchronous machines (PMSM) are widely used. However, there are challenges in the machine design [1-3]. To protect the magnets against the centrifugal force, a retaining ring is often employed on the rotor with interference fit, applying a strong compressive stress on the magnets. When the machine runs at high speed, the centrifugal force further enlarges the compressive stress on the magnets. On the other hand, high-speed machines have a disadvantage of high loss density which results in high temperature rise [1] [4]. Since different materials on the rotor have different thermal expansion, the high temperature rise produces more stress on the magnets. Therefore, the magnets in the high-speed PMSM suffer from high stress which may be up to 90 MPa, as well as high temperature such as 130 °C or higher. It has been well studied that the property of the permanent magnets varies (usually deteriorates) at high temperature, and such variation of property is usually considered in the machine design. However, it is hardly reported in existing literatures how the magnet property changes under high compressive stress. Therefore, an experimental study on this issue is carried out, by measuring the magnet property under various stress and temperature. A test rig is designed and built, and 7 types of NdFeB and 2 types of SmCo magnets are tested. Fig. 1 shows the test results of the NdFeB 42SH magnet at 24°C. When the compressive stress is applied on the magnet, its remanence increases slightly and then stays almost constant, as the curve of test-1 shows. However, when the pressure is released, it is found that the magnet remanence decreases. Then, compressive stress is applied again, and the remanence increases more remarkably, as shown as the curve of test-2. If the pressure is released again, the remanence returns with no extra decrease. Pressure is applied and released repeatedly, and the magnet property variation also repeats (see the curves of test-3, 4, and 5). Moreover, the same property variation trend is observed for the other NdFeB magnets and the SmCo magnets, i.e., the magnet property under the second application of compressive stress is lower than that under the first application of stress, but, the property will not become worse when the stress is applied more times. Therefore, it can be said that the magnet property is stabilized by the compressive stress. Fig. 2 shows the stabilized property variation versus the compressive stress for the other NdFeB magnets at room temperature. Each grade of NdFeB magnet has its own significance of property variation, and so is the case of the SmCo magnets. Similarly, the property variation versus the compressive stress at other temperature is also measured, demonstrating the negative thermal coefficient of the magnet remanence. In high-speed PMSM, the magnets suffer from significant compressive stress and temperature rise, therefore, their property of remanence is different from the data given by the manufacturer, as the manufacturer’s data does not take into account the stress influence. Clearly, the stress influence is negligible in regular-speed machines, but should be considered in the high-speed machine designs. Acknowledgement This work was supported by the Natural Science Foundation of China under the grants

**References:**

[1] J. Shen, et al, CES Transactions on Electrical Machines and Systems, Vol.2, p.23 (2018). [2] D. Gerada, et al, IEEE Trans. Industrial Electronics, Vol.61, p.2946 (2014). [3] F. Zhang, et al, IEEE Transportation Electrification Conference and Expo, Busan, Korea, p.843 (2016). [4] W. Li, et al, IEEE Trans. Industrial Electronics, Vol.61, p.3056 (2014).

**Images:**

![](https://assets.underline.io/uploads/markdown_image/1/image/4d497c5f4dde87ca9f08c6197100f287.jpg)
Fig.1 Remanence variation of NdFeB 42SH magnet versus compressive stress at room temperature.

![](https://assets.underline.io/uploads/markdown_image/1/image/e5371582c44082cee9d7f1184e82052c.jpg)
Fig.2 Stabilized remanence variation of different NdFeB magnets versus compressive stress at room temperature.

 Experimental Study on Remanence Variation of Permanent Magnets for High-Speed Machines

—The eddy current loss in the permanent magnet (PM) and the sleeve of high-speed surface-mounted permanent magnet synchronous motors (HS-SPMSMs) increases the temperature rise of the rotor, which increases the risk of irreversible demagnetization of the PM. In order to reduce the rotor eddy current loss (RECL) of HS-SPMSM, this paper analyzes the suppression effect of four different segmented rotor structures (axially segmented sleeve, circumferentially segmented PM, axially segmented PM, PM and sleeve are simultaneously segmented) on the RECL of HS-SPMSM. On this basis, the applicability of the rotor segmented structure to the RECL with different sleeve materials is further analyzed. Finally, the C-core experiment was used to verify the influence of different segmented rotor structures on the RECL. I. Introduction High-speed permanent magnet synchronous motors (HS-PMSMs) have the characteristics of high-power density, but they also have the characteristics of high loss density and high temperature rise [1]. The high temperature of the rotor not only affects the performance of the motor, but also increases the risk of irreversible demagnetization of the PMs [2]. In order to suppress the RECL of the PMSM, the PM is usually divided into pieces [3-4]. This method is usually applied to conventional speed motors without sleeves. For HS-SPMSM, which is widely used in HS-PMSM, the RECL is mainly concentrated in the PM and sleeve [5], and the segmentation of the PM will affect the eddy current loss in the sleeve. However, there is little research on the effect of permanent magnet segmentation on the RECL of HS-SPMSM with sleeve structure. In this paper, four segmented rotor structures are established, including axial segmented sleeve, circumferential segmented PM, axial segmented PM, PM and sleeve segmented simultaneously. This paper uses different segmented rotor structures with three commonly used sleeve materials (carbon fiber, titanium alloy, stainless steel), and analyzes the influence of different segmented rotor structures on the RECL of HS-SPMSM. Finally, the C-core experiment is used to verify the influence of different rotor structures on RECL. II. Segmented Rotor Structure and Analysis A 15 kW, 30000 rpm HS-PMSM is taken as an example to analyze the RECL under different sleeve structures. When the motor is running at the rated speed, the four segmented rotor structures are analyzed for their suppression effect on RECL. The four segmented rotor structures are shown in Fig. 1. This paper analyzes the RECL of HS-PMSM with several commonly used sleeve materials (carbon fiber, titanium alloy, stainless steel), and studies the influence of segmented rotor structure on the RECL of HS-PMSM by using the four structures shown in Fig. 1. According to the calculation results, for the high conductivity sleeve material, when the number of axial segments reaches a certain degree, it is helpful to reduce the RECL, while for the carbon fiber sleeve, the axial segmentations have little effect on the RECL. When the PM is segmented axially or circumferentially, the RECL can be significantly reduced for the carbon fiber sleeve, and the effect of axial segmentation is better than that of circumferential segmentation. However, for the sleeve made of high conductivity material, the shielding effect on the eddy current magnetic field is weakened due to the segmentation of the permanent magnet, which makes the sleeve produce more eddy current loss and increases the total RECL. If the rotor structure shown in Fig. 1 (d) is used, the RECL can be reduced for all three sleeve materials, but the number of segments of PMs and sleeves needs to be reasonably configured. III.Experimental Verification In order to verify the influence of various sleeve materials on rotor eddy current loss and the suppression effect of sleeve axial segment structure on rotor eddy current loss. This paper verifies the influence of different sleeve structures on the eddy current loss of the rotor through the C-core experiment. The C-core experimental platform and different sleeve materials are shown in Fig. 2. IV.CONCLUSIONS This paper compares and analyzes the RECL of HS-SPMSM with different rotor structures. The RECL generated by the high-conductivity material sleeve is relatively high, but it can shield the eddy current loss of the permanent magnet. For high-conductivity sleeves, the axial segmentation can reduce the eddy current loss of the sleeve by increasing the equivalent resistance of the sleeve and further suppress the RECL, when using PM segments, the sleeve will generate more eddy current loss, resulting in increased RECL. However, for carbon fiber sleeves, circumferentially or axially segmented PMs can help reduce the RECL, but the axial segmentation of the sleeve has almost no effect on the RECL. For the three sleeve materials, segmenting the sleeve and the PM at the same time will reduce the RECL, but the number of segmentations of the sleeve and the PM should be reasonably configured.

**References:**

 [1] C. Chen, Y. Wang and J. Wei, "Research on method for reducing eddy current loss of magnet in high-speed permanent magnet synchronous motor," 2019 22nd International Conference on Electrical Machines and Systems (ICEMS), Harbin, China, 2019, pp. 1-4. [2] J. Dong, Y. Huang, and L. Jin, "Thermal Optimization of a High-Speed Permanent Magnet Motor," IEEE Transactions on Magnetics, vol. 50, no. 2, pp. 749-752, Feb. 2014. [3] J. Shen, H. Hao, and M. Jin, "Reduction of Rotor Eddy Current Loss in High-Speed PM Brushless Machines by Grooving Retaining Sleeve," IEEE Transactions on Magnetics, vol. 49, no. 7, pp. 3973-3976, July 2013. [4] K. Yamazaki, M. Shina, and Y. Kanou, "Effect of Eddy Current Loss Reduction by Segmentation of Magnets in Synchronous Motors: Difference Between Interior and Surface Types," IEEE Transactions on Magnetics, vol. 45, no. 10, pp. 4756-4759, Oct. 2009. [5] H. Jun, J. Lee, and H. Lee, "Study on the Optimal Rotor Retaining Sleeve Structure for the Reduction of Eddy-Current Loss in High-Speed SPMSM," IEEE Transactions on Magnetics, vol. 51, no. 3, pp. 1-4, March 2015.

**Images:**

![](https://assets.underline.io/uploads/markdown_image/1/image/3f1f9570ee1a8851a342459ca9f76158.jpg)
Fig. 1. Various rotor structures. (a) Axial segmented sleeve. (b) Circumferential segmented PM. (c) Axial segmented PM. (d) PM and sleeve are segmented simultaneously.

![](https://assets.underline.io/uploads/markdown_image/1/image/a804f5a1b6871890c6cf1460dfa099d5.jpg)
Fig. 2. Experimental verification. (a) C-core loss test experiment platform. (b) Sleeve materials.

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